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Full text of "The August 1, 1975 Oroville earthquake investigations"

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The August 1, 1975 
Oroville Earthquake 

Investigations 






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ON THE COVER: 

Aerial View of Oroville Facilities 



Department of 
Water Resources 

Bulletin 203-78 



The August 1, 1975 
Oroville Earthquake 
Investigations 



February 1979 



Huey D. Johnson 

Secretary for Resources 



The Resources 
Agency 



Edmund G. Brown Jr. 

Governor 



State of 
California 



Ronald B. Robie 

Director 



Department of 
Water Resources 



Digitized by the Internet Archive 

in 2010 with funding from 
University of California Libraries 



http://www.archive.org/details/august11975orovi20378cali 



FOREWORD 



The epicenter of the 1975 earthquake near Oroville, California 
was close to the Oroville-Thermalito features of the California State Water 
Project. Therefore, the Department of Water Resources initiated structural 
reanalyses of the Project facilities and seismological and geological 
investigations . 

The performance of the Oroville-Thermalito facilities during the 
August 1975 earthquake sequence, reported in Bulletin 203 (April 1977), 
demonstrated their ability to withstand that seismic loading. No structural 
damage occurred. The only damage was to a few of the secondary facilities; 
this damage was only superficial. 

The Department conducted intensive investigations to determine : 

1. Geologic and tectonic conditions 

2. Fault mechanism and orientation 

3. Crustal movements 

4. Public safety as it relates to the Department's facilities. 

The Department established a Special Consulting Board for the 
Oroville Earthquake, consisting of nine experts in the fields of seismology, 
geology and dam design to review the Department's investigations and make 
recommendations . 

Bulletin 203-78 presents a detailed reanalysis of the Department's 
facilities and results of the detailed seismological and geological investi- 
gations. On the basis of determinations from the investigations completed 
to date, the Department concludes that the Oroville facilities do not pose 
a threat to public safety. 

The reanalyses of Thermalito Forebay, Afterbay and Power Plant 
Headworks, and the Bidwell Canyon Saddle Dam are still in progress, with 
publication planned by mid 1979. 




Ronald B. Robie, Director 
Department of Water Resources 
The Resources Agency 
State of California 



state of California 

Edmund G. Brown Jr. , Governor 

The Resources Agency 

Huey D. Johnson, Secretary for Resources 

DEPARTMENT OF WATER RESOURCES 
Ronald B. Robie, Director 

Charles R. Shoemaker Gerald H. Meral Robert W. James 

Acting Deputy Director Deputy Director Deputy Director 

Jack B. Johnston 
Acting Assistant Director 

Division of Operations and Maintenance 

Howard H. Eastin Division Chief 

Lawrence A, Mullnix Chief, Water Engineering Office 

Clifford V. Lucas Chief, Civil Maintenance Branch 

Philip F. Johns Chief, Oroville Field Division 



Authors 

Chapter I 
John R. Campbell Project Surveillance Section 

Chapter III 
Paul W, Morrison, Jr Earthquake Engineering 

Chapter IV 
John R. Campbell Project Surveillance Section 

Chapter IX 
J. P. Cedarholm Civil Maintenance Section 

Division of Design and Construction 

Gordon W, Dukleth Division Chief 

Keith G. Barrett Chief, Design Office 

Ernest C. James Chief, Civil Design Branch 



Division of Design and Construction (Continued) 

John W. Marlette Chief, Project Geology Branch 

William M. Verigin Chief, Dams and Canals Section 

Donald C. Steinwert Chief, Structures Section 

Authors 

Chapter II 

John W. Marlette Chief, Project Geology Branch 

Robert J. Akers Project Geology Branch 

Kenneth A. Cole Project Geology Branch 

Richard D. McJunkin Project Geology Branch 

Chapter V 

William D. Hammond Dams and Canals Section 

Leslie F. Harder Dams and Canals Section 

Chapter VI 

Samuel J. Linn, Jr Structural Section 

Edgar R. Najera Structural Section 

Chapter VII 

Samuel J. Linn, Jr. Structural Section 

Edgar R. Najera Structural Section 

Chapter VIII 

Arnold E, Eskel Structural Section 

Samuel J. Linn, Jr Structural Section 

Edited by 
Earl G. Bingham, Reports Administration 



CONTENTS 

Page 

Foreword ■'■■'■-'■ 



IV 



Organization 

CHAPTER I. INTRODUCTION 



Purpose 1 

Description of the Oroville Facilities 1 

The Investigating Organization 3 

Reanlysis of Project Structures for Earthquake Safety 4 

Summary of Conclusions and Recommendations 5 

Geological Investigations (Chapter II) 5 

Seismology (Chapter III) 6 

Vertical and Horizontal Geodesy (Chapter IV) 6 

Oroville Dam, Evaluation of Seismic Stability (Chapter V) 7 

Oroville Dam, Flood Control Outlet Structure (Chapter VI) 7 

Thermalito Diversion Dam (Chapter VII) 7 

Reappraisal of Secondary Structures (Chapter VIII) 7 

Fish Barrier Dam 7 

Edward Hyatt Powerplant 7 

Thermalito Powerplant 7 

Miscellaneous Structures 8 

Bridges 8 

Switchyards 8 

Contingency Plan for Seismic Emergencies (Chapter IX) 8 

Department's Findings 8 

Uncompleted Reports 8 

-*:*>■ Safety Review Requirements 8 

Report of the Special Consulting Board for the Oroville Earthquake 9 

Review by the Division of Mines and Geology 13 



CHAPTER II. GEOLOGIC INVESTIGATIONS 



Purpose of Investigation 15 

Previous Work 15 

Scope of Investigation 17 

Seismic History 19 

The 1975 Earthquake Series 21 

Ground Cracking 23 

Ground Elevation Changes 26 

Area Lineaments 27 

Geologic Setting 30 

Geographic Location 30 

Geologic Framework 30 

Descriptive Geology 33 

Bedrock Series Rocks 33 

Melange 34 

Previous Investigations and Age 34 

Contact Relationships 34 

Lithologic Description 35 



Arc Rocks 38 

Previous Investigations and Age 38 

Contact Relationships 38 

Lithologic Description 38 

Monte de Oro Formation 40 

Previous Investigations and Age 40 

Contact Relationships 40 

Lithologic Description 41 

Smartville Ophiolite 41 

Previous Investigations and Age 41 

Contact Relationships 42 

Litholigic Description 42 

Intrusive Rocks 47 

Previous Investigations and Age ; 47 

Contact Relationships 47 

Lithologic Description 48 

Origin of Sierra Nevada Plutons 49 

Superjacent Series Rocks 50 

Chico Formation 50 

Previous Investigations and Age 50 

Contact Relationships 50 

Lithologic Description 50 

lone Formation 51 

Previous Investigations and Age 51 

Contact Relationships 51 

Lithologic Description - lone Formation Undifferentiated . . 51 

Lithologic Description - Auriferous Gravel ... 51 

Lithologic Description - Oroville Tuff 52 

Love joy Formation 53 

Previous Investigations and Age 53 

Contact Relationships 53 

Lithologic Description 54 

Tuscan Formation 54 

Previous Investigations and Age 54 

Contact Relationships 55 

Lithologic Description 55 

Late Cenozoic Gravels 56 

Previous Investigations 56 

Contact Relationships 56 

Lithologic Description 57 

Quaternary Landslides 57 

Structural Geology 61 

Faults 61 

Mesozoic Faults - Northern Foothills 65 

Mesozoic Faults - Project Area 67 

Swain Ravine, Paynes Peak and Prairie Creek 

Lineament/Fault Zones 67 

Oregon Gulch Fault 68 

Monte de Oro Fault 69 

Unnamed Faults . 69 

Glover Ridge Fault 69 

Cenozoic Fault Movement 70 

Swain Ravine Lineament Fault Zone 71 



vii 



Prairie Creek Lineament Fault Zone 74 

Paynes Peak Lineament Fault Zone 7^ 

Thermalito Powerplant Foundation Faults 75 

Chico Lineament 75 

Soda Springs Lineament 77 

Web Hollow Lineament 77 

Paradise-Magalia Lineament 77 

Summary 79 

Mesozoic Folds 79 

Cenozoic Folding 79 

Summary of Geologic History 80 

Causes of the Oroville Earthquake 82 

^^^servoir- Induced Seismicity 83 

Potential Hazards to State Water Facilities 86 

Groirnd Shaking 86 

Fault Displacement 86 

Regional Changes in Ground Elevation 87 

Potential Hazard to Specific Facilities 87 

Oroville Dam and Saddle Dams 87 

Thermalito Forebay and Afterbay 87 

Thermalito Powerplant 88 

Other Structures 88 

Summary and Conclusions 88 

References Cited ..... 90 

Addenda: Department of Water Resources Exploration Trench Logs 103 



CHAPTER III. SEISMOLOGY 

Introduction 123 

Data 123 

Results 124 

Discussion 124 

Conclusions 124 

References Cited 139 

CHAPTER IV. VERTICAL AND HORIZONTAL GEODESY 

Vertical Crustal Movements 141 

Introduction 141 

Precise Survey Programs 141 

September 1967 141 

July-September 1968 142 

October-November 1969 142 

August-September 1975 142 

January-April 1976 142 

September-November 1976 142 

September-November 1977 142 

Precise Survey Adjustment 143 

Free Adjustment 143 

Spur Lines 144 

Elevation Differential Isograms 146 

General 146 

September 1967-October 1969 146 



October 1969-August 1975 146 

September 1967-October 1977 146 

August 1975-October 1976 146 

October 1976-October 1977 153 

August 1975-October 1977 153 

Elevation Differential Along Lines 153 

General 153 

Avocado 153 

Bald Rock 153 

Bidwell 153 

Bidwell Canyon Saddle Dam 157 

Canyon Drive 157 

Cleveland Hill 157 

Dam 157 

Dunstone 157 

Feather Falls 157 

Foothill 157 

Miners Ranch 157 

Mission Olive 157 

Morris 167 

Olive 167 

Oro-Bangor 167 

Oroville 167 

Potter 167 

Richvale 167 

Thompson Flat 167 

Wyn-Miners Ranch 167 

103 167 

Oroville Dam Crest Differential Settlement 167 

General 167 

Commentary 181 

Conclusions 181 

Horizontal Earth Movements 181 

Introduction 181 

Horizontal Geodetic Control and Triangulation Programs 181 

September 1967 181 

April 1968 183 

August- Sept ember 1975 183 

Computations and Analyses 183 

Commentary 185 

Conclusions 185 



CHAPTER V. OROVILLE DAM: EVALUATION OF SEISMIC STABILITY 

Acknowledgments 187 

1. Introduction 188 

Background 189 

Commentary 190 

Summary of Findings 190 

Conclusions 191 



2. Description of Embankment Materials and Dynamic Instrumentation. . . 192 

Embankment Materials 192 

Dynamic Instrumentation 194 

Original System 194 

Upgraded System 195 

3. Recorded Embankment Response to the 1975 Earthquakes 197 

General 197 

Recorded Events 197 

August 1, 1975 197 

August 5, 1975 199 

September 27, 1975 199 

Observed Natural Period 201 

4. Analysis of Static Stresses by Finite Element Method 203 

General 203 

Material Properties 204 

Static Stress Analysis 205 

Seepage Forces 205 

5. Determination of Dynamic Shear Modulus and Damping Values for 

Embankment Shell Material 208 

General 208 

Cyclic Triaxial Test 209 

Analysis of Recorded Embankment Response During the 1975 

Earthquakes 211 

General 211 

Natural Period for Two-Dimensional Analysis 213 

Shear Strain for Two-Dimensional Analysis 213 

Shear Modulus Reduction Factor 213 

K„ vs. Natural Period 214 

RiSP of K 214 

Comparison of Observed and Computed Crest Motions 214 

Embankment Response Model 216 

August 1 Event 216 

September 27 Event 217 

Dynamic Properties Adopted for Gravel Shell 217 

6. Reanalysis Earthquake 221 



7. Analysis of Dynamic Stresses for the Reanalysis Earthquake 229 

Methods of Response Computation 229 

Acceleration Response of Dam to Design Earthquake 229 

Input Variables and Computed Shear Stresses 231 

Influence of Shear Modulus of Shell Material 232 

Influence of Shear Modulus of Core Material 233 

Computer Programs LUSH and QUAD4 234 

Influence of Poisson's Ratio 237 

Influence of Embankment Section 239 

Combined Influence of Variables 241 

Three-Dimensional Effect 241 



Cyclic Shear Strength 244 

Cyclic Strength Test Program 244 

Sample Gradations and Density 244 

Modeling Embankment Shell Gradation 244 

Relationship of Test Sample Density to Field Density . . . 244 

Summary of Test Procedures 249 

30 cm Diameter Samples 249 

7.1 cm Diameter Samples 249 

Results of Cyclic Triaxial Tests 250 

Investigation of Sample Behavior of Dense Sands in Static and 

Cyclic Triaxial Tests 251 

Objective 251 

Program and Procedures 254 

Static Tests on Monterey Sand 256 

Cyclic Tests on Monterey Sand 260 

Cyclic Tests on Oroville Sand 262 

Analysis of Test Results 268 

Extension Strain 268 

Necking Behavior 268 

Sample "Tension" 268 

Cyclic Strength Interpretations Considered 276 

Strength Interpretation I 276 

Strength Interpretation II 276 

Evaluation of Performance 281 

General Considerations 281 

Method of Evaluation 281 

Failure Planes 281 

Equivalent Regular Stress Time History 282 

Cases Analyzed and Assumptions 282 

Case a 284 

Case b 284 

Case c 284 

Case d 285 

Comparison of Cases 285 

Predicted Behavior - Best Judgment Case 285 

Shell K g^ 285 

Cyclic sStar Strength 285 

Three-Dimensional Effect 285 

Drainage 285 

Predicted Behavior 286 

Estimated Displacements for Conservative Assumptions 287 

References , 288 



CHAPTER VI. SEISMIC ANALYSIS OF THE OROVILLE DAM FLOOD 
CONTROL OUTLET STRUCTURE 

Commentary 291 

Conclusion 294 

Earthquake Analysis of the Oroville Dam Flood Control Outlet 

Structure, June 1977, by Edward L. Wilson, Frederick E. Peterson, 

and Ashraf Habibullah 299 



CHAPTER VII. SEISMIC ANALYSIS OF THE 
THERMALITO DIVERSION DAM 

Commentary 351 

Earthquake Response Analysis of Thermalito Diversion Dam By 

Anil K. Chopra 355 



CHAPTER VIII. REAPPRAISAL OF 
SECONDARY STRUCTURES 

Introduction 389 

Fish Barrier Dam 389 

Description 389 

Original Seismic Analysis 389 

Recommendation for Seismic Reanalysis 395 

Power and Pumping Plant Facilities 395 

Edward Hyatt Powerplant 399 

Conclusion 399 

Thermalito Powerplant 399 

Conclusion 395 

Miscellaneous Structures 402 

Oroville Operations and Maintenance Center 402 

Oroville Dam 402 

Thermalito Forebay and Afterbay 402 

Feather River Fish Hatchery 402 

Conclusion 403 

Bridges 403 

Conclusion 403 

Swtichyard Structures and Apparatus 403 

Conclusion 404 



CHAPTER IX. CONTINGENCY PLAN FOR SEISMIC EMERGENCIES 

Organization and Responsibilities 405 

Division Policy 405 

Division Plan of Operation 405 

Oroville Field Division Command Post 405 

Operational Command Post 406 

Operational Facilities 406 

Operational Plan 406 

Security Command Post 406 

Security Plan 406 

Procedures for Reacting to Seismic Events 406 

Detection 406 

Earthquake Magnitudes and Epicenters 406 

Criteria for Notification 407 

Response 407 

Inspection of Project Facilities Following an Earthquake ...... 408 

Rapid Response Inspection Plan 408 

Follow-up Inspection Plan 408 

Returning Facilities and Equipment to Full Operational Status 408 



xii 



Regulating Features 409 

Nonregulating Features 409 

List of Operating Criteria for Regulating Lake Oroville 414 

Decision Making Criteria for Operating Features Which Can Be 

Regulated 414 

Palermo Outlet 414 

Oroville Dam Spillway 414 

River Outlet Valves 414 

Edward Hyatt Intake 415 

Edward Hyatt Powerplant 415 

Critical Conditions for Features Which Cannot be Regulated 416 

Oroville Dam 416 

Bidwell Canyon Saddle Dam 416 

Parish Camp Saddle Dam 417 

Oroville Dam Spillway 417 

Edward Hyatt Intake and Penstock 417 

Palermo Intake and Outlet 417 

River Outlet Valve Chamber 417 

List of Operating Criteria for Regulating Thermalito Diversion Pool . . . 417 

Decision Making Criteria for Operating Regulating Features 417 

Thermalito Diversion Dam 417 

Critical Conditions for Features Which Cannot Be Regulated 418 

Thermalito Diversion Dam 418 

Thermalito Power Canal Headworks 418 

List of Operating Criteria for Regulating Thermalito Forebay Reservoir 

and Power Canal 418 

Decision-Making Criteria for Operating Regulating Features 418 

Thermalito Intake Structure 418 

Thermalito Powerplant 418 

Critical Conditions for Features Which Cannot Be Regulated 419 

Thermalito Forebay Dam 419 

Thermalito Intake Structure 419 

Thermalito Power Canal (Cut Section) 420 

Thermalito Power Canal (Fill Section) 420 

List of Operating Criteria for Regulating Thermalito Afterbay Reservoir . 420 

Decision-Making Criteria for Operating Regulating Features 420 

Thermalito Afterbay River Outlet 420 

Sutter-Butte Outlet 420 

PG&E Outlet 420 

Western Canal and Richvale Outlets 420 

Thermalito Afterbay Dam Ground Water Pumping System 420 

Critical Condition for Features Which Cannot Be Regulated 421 

Thermalito Afterbay Dam 421 

Thermalito Power House Structure 421 

Thermalito Afterbay River Outlet 421 

Sutter-Butte Outlets 421 

PG&E Outlet 421 

Western-Richvale Outlets 421 

Commentary 421 

Conclusion 421 



TABLES 

No . Page 

1 Exploration Trenches in Foothill Belt — Oroville to Auburn Area ... 62 

2 Summary of Geologic Events 80 

3 Earthquake Epicenters, June 1975-December 1975 133 

4 Earthquake Epicenters, January 1976-May 1978 137 

5 Value of Stress-Strain Parameters for Analysis of Oroville Dam . . . 204 
6. Static Stress Comparison 205 

FIGITRES 
CHAPTER I 

1 Oroville Dam 1 

2 Location Map, Oroville Facilities 2 

3 Thermalito Diversion Dam 3 

4 Thermalito Forebay Dam 4 

5 Thermalito Afterbay Dam 4 

CHAPTER II 

6 Location map of six quadrangle study area 16 

7 Historic earthquakes within a 100 km (62 mi) radius of Oroville . . 18 

8 Aftershock locations of the Oroville earthquake 20 

9 Locations of the Cleveland Hill and Mission Olive crack zones and 

sites of Department of Water Resources exploration trenches ... 22 

10 Ground cracking that resulted from the August 1, 1975, Oroville 

earthquake 23 

11 Close-up view of a ground crack on southwest slope of Cleveland 

Hill 23 

12 Locations of ground cracking from the Oroville earthquake and major 

lineaments in the southern study area 24 

13 Aerial view of northern limit of ground cracking 25 



Figure No . Page 

14 Changes in ground elevations around Lake Oroville, August 1975 to 

October 1976 28 

15 Changes in crest elevations of Bidwell Canyon Saddle Dam between 

November 1967 and October 1977 29 

16 Lineaments and faults in the northwestern Sierran foothills .... 31 

17 Natural geologic provinces of California with field area location . 32 

18 Small-scale parasitic isoclinal fold within melange metasedimentary 

rock 35 

19 Relict bedding cross cut by foliation in melange metasedimentary 

rock 35 

20 Sheared volcaniclastic metaconglomerate in melange metasedimentary 

rock 36 

21 Exotic marble block in melange 36 

22 Sample of olistostromal marble-phyllite collected in melange .... 37 

23 Relict bedding in arc metasedimentary rock 39 

24 Arc complex metavolcanic tuff breccia 39 

25 Arc complex relict pillow and flow lavas cut by fault 39 

26 Relict bedding and cross-bedding in arc tuff breccia and tuffaceous 

metasedimentary rock 40 

27 Igneous stratigraphy of Standard Oceanic Crust and Smartville 

ophiolite 43 

28 Well developed metavolcanic Smartville pillows 43 

29 Well developed metavolcanic Smartville pillows 44 

30 Sheared metavolcanic Smartville pillows 44 

31 Metavolcanic Smartville sheeted dikes 45 

32 Gabbroic screen rock in Smartville metavolcanic sheeted dikes ... 45 

33 Granophyric screen rock in Smartville metavolcanic sheeted dikes . . 46 

34 Mesozoic time scale with corresponding intrusive epochs in the 

Sierra Nevada region 47 

35 Metavolcanic xenoliths within Swedes Flat plutonic rock 48 

36 View of Bald Rock exhibiting surface exposure and exfoliation 

that is typical of the Sierra Nevada complex 49 



Figure No . Page 

37 lone Formation auriferous gravel with intercalated Oroville tuff 

(Mehrten Formation-?) 52 

38 Lovejoy Formation basalt disconformably overlying lone Formation . . 53 

39 Lovejoy Formation basalt on North Table Mountain 54 

40 Young erosional surface cut into Tuscan Formation that is adjacent 

to older and structurally higher erosional surface cut into 
mesozoic metamorphic rock 

41 Tuscan Formation volcanic conglomerate, cross-bedded sand and laharic 

mudflow breccia 56 

42 Late Cenozoic gravel and cross-bedded sand (Red Bluff Formation-?) . 57 

43 Late Cenozoic gravel (Red Bluff Formation-?) unconformably overlying 

Oroville tuff (Mehrten Formation-?) along the Feather River .... 58 

44 Landslide in lone Formation 59 

45 Stringtown Mountain landslides 59 

46 Prehistoric landslide north slope of Bloomer Hill 60 

47 Major lineaments in the northwestern Sierran foothills showing 

exploration localities with faulting assessments for each site . . 61 

48 Foothills Fault System of the western Sierra Nevada, California . . 66 

49 Aerial view of Glover Ridge (klippe) and traced location of 

Glover Ridge Fault 70 

50 Aerial northeast view of the Cleveland Hill Fault along the 

western side of Cleveland Hill 71 

51 Aerial view of Chico monocline 75 

52 Normal fault in Tuscan Formation 76 

53 Normal fault in Tuscan Formation 76 

54 Vertical aerial view of eastern fracture zone developed in Tuscan 

Formation 77 

55 Map with cross section oriented perpendicular to suspected fault 

through Magalia Reservoir 78 

56 Comparison of foreshock-af tershock patterns for the Oroville 

earthquake and Mogi's "Type II" (reservoir-induced earthquakes). . 84 

57 Water level history of Lake Oroville from initial filling to 

September 1978 84 



CHAPTER III 

Figure No . Page 

58 DWR-USGS Oroville Sensitive Seismographic Network 125 

59 Oroville Foreshocks, Mainshock, and Aftershocks; June 1, 1975- 

December 31, 1975 126 

60 1975 Oroville Earthquake Hypocenters (North Vertical Cross 

Section) 127 

61 1975 Oroville Earthquake Hypocenters (Middle Vertical Cross 

Section) 127 

62 1975 Oroville Earthquake Hypocenters (South Vertical Cross 

Section) 128 

63 Oroville Earthquake Epicenters (January 1, 1976-May 31, 1978) . . . 129 

64 Oroville Earthquake Hypocenters, 1976-May 31, 1978 (North Vertical 

Cross Section) 130 

65 Oroville Earthquake Hypocenters, 1976-May 31, 1978 (Middle Vertical 

Cross Section) 130 

66 Oroville Earthquake Hypocenters, 1976-May 31, 1978 (South Vertical 

Cross Section) 131 

67 Oroville Earthquake Hypocenters, August 2, 1975-December 31, 1975 

(Middle Vertical Cross Section) 131 

68 Oroville Sequence, Number of Aftershocks/Month, Water Surface 

Elevation (August 1975-June 1978) 132 

CHAPTER IV 

Lake Oroville Water Surface Elevation 142 

70 Oroville Area Level Lines (1977) 143 

71 Precise Level Net for Study of Lake Oroville - 1967 144 

72 Precise Level Net for Study of the Oroville Earthquake - 1977 . . . 145 

73 Elevation Differential Isogram — September 1967-October 1969 .... 147 

74 Elevation Differential Isogram — October 1969-August 1975 148 

75 Elevation Differential Isogram — September 1967-October 1977 .... 149 

76 Elevation Differential Isogram — August 1975-October 1976 150 

77 Elevation Differential Isogram — October 1976-October 1977 151 



Figure No . Page 

78 Elevation Differential Isogram — August 1975- October 1977 152 

79 Avocado Elevation Differentials 154 

80 Bald Rock Elevation Differentials 155 

81 Bidwell Elevation Differentials 156 

82 Bidwell Canyon Saddle Dam Elevation Differentials 158 

83 Canyon Drive Elevation Differentials 159 

84 Cleveland Hill Elevation Differentials 160 

85 Dam Elevation Differentials 161 

86 Dunstone Elevation Differentials 162 

87 Feather Falls Elevation Differentials 163 

88 Foothill Elevation Differentials 164 

89 Miners Ranch Elevation Differentials 165 

90 Mission Olive Elevation Differentials 166 

91 Morris Elevation Differentials 168 

92 Olive Elevation Differentials 169 

93 Oro-Bangor Elevation Differentials 170 

94 Oroville Elevation Differentials 171 

95 Potter Elevation Differentials 172 

96 Richvale Elevation Differentials (1 of 2) 173 

97 Richvale Elevation Differentials (2 of 2) 174 

98 Thompson Flat Elevation Differentials 175 

99 Wyn-Miners Ranch Elevation Differentials 176 

100 103 Elevation Differentials (1 of 3) 177 

101 103 Elevation Differentials (2 of 3) 178 

102 103 Elevation Differentials (3 of 3) 179 

103 Oroville Dam Crest Differential Settlement (References to the 

Abutments) 180 

104 Horizontal Geodetic Control and Triangulation Net (1967-1975) . . . 182 



xviii 



CHAPTER V 

Figure No . Page 

105 Location Map 188 

106 Oroville Maximum Section 192 

107 Average Gradation Curves of Oroville Dam Materials 193 

108 Oroville Dam Embankment, Original Dynamic Instrumentation 194 

109 Oroville Dam Embankment, Present Dynamic Instrumentation (December, 

1978) 196 

110 Acceleration Records, Main Event of August 1, 1975 198 

111 Acceleration Records with Corrected Time Scales, August 1, 1975 . . 200 

112 Acceleration Records, Event of August 5, 1975 201 

113 Acceleration Records, Event of September 27, 1975 202 

114 Acceleration Response Spectra for Crest Motions, Event of 

August 1, 1975 203 

115 Finite Element Mesh, Maximum Section Oroville Dam . . .' 204 

116 Contours of Effective Maximum Principal Stress in Oroville 

Dam, Full Reservoir 206 

117 Contours of Effective Minimum Principal Stress in Oroville 

Dam, Full Reservoir 206 

118 Contours of Maximum Shear Stress in Oroville Dam, Full Reservoir . . 207 

119 Orientation of Principal Stresses 207 

120 In-Situ Shear Moduli for Saturated Clays 208 

121 Sample Gradation for Cyclic Triaxial Tests 209 

122 Modulus Determinations for Gravelly Soils 210 

123 Comparison of Damping Ratios for Gravelly Soils and Sands 211 

124 Section on Long Chord of Dam Axis 212 

125 Comparison of Natural Periods for Two-Dimensional and Three- 

Dimensional Embankment in Triangular Canyon 212 

126 Shear Modulus Reduction Curve for Embankment Soils 213 

127 Static Shear Strength Envelopes for Core Material 214 



Figure No . Page 

128 K2iiiax ^^ • Natural Period 214 

129 Maximum Accelerations Computed from 3D and Plane Strain Analyses 

Using Base Motions from Taft Record (after Makdisi) 215 

130 Damping Ratios for Embankment Soils 217 

131 Comparison of Acceleration Time Histories, August 1 Main Shock . . . 218 

132 Comparison of Displacement Time Histories and Acceleration Response 

Spectra for Crest Motions, August 1 Main Shock 219 

133 Comparison of Acceleration Time Histories and Response Spectra, 

September 27 Aftershocks 220 

134 Lineaments, Faults and Recorded Epicenters Around Oroville 222 

135 Location of Faults in Relation to Oroville Dam 223 

136 Relationship of Oroville Dam to Assumed Northward Extension of 

Fault 224 

137 Earthquake Ground Motion Characteristics 226 

138 Reanalysis Earthquake 227 

139 Response Spectra for the Reanalysis Earthquake 228 

140 Acceleration Response to Reanalysis Earthquake 230 

141 Influence of Shear Modulus of Shell Material on Computed Maximum 

Horizontal Dynamic Shear Stresses 232 

142 Influence of Shear Modulus of Core on Computed Maximum Horizontal 

Dynamic Shear Stresses 234 

143 Comparison of Horizontal Dynamic Shear Stress Time Histories from 

LUSH and QUAD4 235 

144 Comparison of Horizontal Dynamic Shear Stress Time Histories from 

LUSH and QUAD4 236 

145 Comparison of Computed Maximum Horizontal Dynamic Shear Stresses 

by Computer Programs LUSH and QUAD4 237 

146 Model Embankment for Determining Influence of Poisson's Ratio on 

Dynamic Shear Stresses 238 

147 Influence of Poisson's Ratio on Computed Dynamic Shear Stresses . . 239 

148 Comparison of Computed Maximum Horizontal Dynamic Shear Stresses 

for Different Embankment Sections 240 



Figure No . Page 

149 Comparison of Maximum Horizontal Shear Stresses Determined from 

3D and Plane Strain Analyses Using Base Motions from Taft 

Record (after Makdisi) 242 

150 Estimated Three-Dimensional Effect on Computed Maximum Horizontal 

Dynamic Shear Stresses 243 

151 Field and Modeled Oroville Gravel Gradations 245 

152 Final Statistical Analysis - Zone 3, Percent Compaction 246 

153 Field Control Tests - Zone 3 247 

154 Maximum Density Tests - Zone 3 248 

155 Cyclic Triaxial Test Records for Modeled Oroville Gravel 252 

156 Cyclic Triaxial Test Records for Modeled Oroville Gravel 253 

157 Cyclic Triaxial Test Records for Modeled Oroville Gravel 254 

158 Monterey "0" Sand and Oroville Sand Gradations 255 

159 Typical Static Triaxial Compression Test Results for Monterey 

"0" Sand 257 

160 Typical Static Triaxial Extension Test Results for Monterey "0" 

Sand 258 

161 Summary of Static Triaxial Test Results for Dense 

Monterey "0" Sand 259 

162 Cyclic Triaxial Strain Envelopes for Monterey "0" Sand 260 

163 Static and Cyclic Triaxial Test Results for Dense Monterey "0" Sand . 261 

164 Cyclic Triaxial Test Results for Monterey "0" Sand 262 

165 Cyclic Triaxial Test Records for Monterey "O" Sand 263 

166 Shear Plane Development during Final Stage of Necking for Monterey 

"0" Sand 264 

167 Cyclic Triaxial Test Records for Monterey "0" Sand 265 

168 Cyclic Triaxial Test Records for Oroville Sand 266 

169 Cyclic Triaxial Test Records for Oroville Sand 267 

170 Cyclic Triaxial Test Records for Modeled Oroville Gravel 269 

171 Cyclic Traixial Test Records for Modeled Oroville Gravel 269 

172 Cyclic Triaxial Test Records for Modeled Oroville Gravel 270 



Figure No . ° ■ 

173 Extension/Compression Cycle for Monterey "0" Sand Cyclic 

Triaxial Test 271 

174 Compression/Extension Cycle for Monterey "O" Sand Cyclic Triaxial 

Test 273 

175 Comparison of Shaking Table and Cyclic Triaxial Test Results 

for 5 Cycles 274 

176 Comparison of Shaking Table and Cyclic Triaxial Test Results 

for 10 Cycles 275 

177 Cyclic Strength Envelopes for Strength Interpretation I - Static 

and Cyclic Test Results 278 

178 Typical Extrapolations of Isotropically-Consolidated Cyclic 

Triaxial Tests on Modeled Oroville Gravel 279 

179 Cyclic Strength Envelopes for Strength Interpretation II - 

Extrapolated Cyclic Test Results 280 

180 Representative Relationship Between t/t and Number of Cycles 

Required to Cause Liquefaction (Seed et al, 1975) 283 

181 Computed Compressive Strain Potentials in Upstream Shell - 

Percent 286 



CHAPTER VI 

182 Oroville Dam Flood Control Outlet Structure Plan and Elevation . . . 292 

183 Elevation and Sections 293 

184 Maximum Tensile Stresses at time 7.76 sec 295 

185 Maximum Tensile Stresses at time 8.46 sec 296 

186 Maximum Tensile Stresses at time 7.76 sec. in steel 297 

CHAPTER VII 

187 Plan and Elevation 352 

188 Typical Sections 353 

CHAPTER VIII 

189 Location Map, Edward Hyatt Powerplant Facilities 390 

190 Fish Barrier Dam 391 

191 Fish Barrier Dam, Plan and Elevation 392 



Figure No . Page 

192 Fish Barrier Dam, Section and Details 393 

193 Location Map, Thermalito Powerplant^ Forebay^ and Afterbay 394 

194 Transverse Section, Units Nos. 3, 4, 5 and 6 396 

195 Generator Room, Plan - Elevation 252.0 397 

196 Overall view of Edward Hyatt Powerplant Intake Structure 398 

197 Transverse Section, Unit No . 1 400 

198 Switchyard and General Floor 401 

199 230-KV Power Circuit Breakers 404 

CHAPTER IX 

200 Schematic Diagram of Oroville Complex 410 

201 Schematic Diagram of Oroville Dam and Vicinity 411 

202 Schematic Diagram of Thermalito Forebay and Vicinity 412 

203 Schematic Diagram of Thermalito Afterbay and Vicinity 413 

APPENDIXES 

A Reports Prepared by the Special Consulting Board and Responses by 

by the Department of Water Resources 423 

B Acceleration Time Histories and Response Spectra for the August 1, 
1975 and September 27, 1975 Recorded Motions on Dam Crest and 
Bedrock, in Upstream-Dovmstream Direction. (Figs. B-1 through B-8) 457 

C Static Stresses from Static Finite Element Analysis. Figs. C-1 

through C-9) 467 

D Time Histories and Response Spectra for Reanalysis Earthquake. 

(Figs. D-1 through D-6) 483 

E Results of Dynamic Finite Element Analyses for Reanalysis 

Earthquake 489 

Maximum Section ~ Shell K2 max = 350, 200, 130 

Element Stresses and Strains (Figs. E-1 through E-18) . . 490 
Shear Stress Time Histories (Figs. E-19 through E-39) . . 500 

Section 2 — Shell Kn „^^ - 130 - LUSH and QUAD4 
I max ^ 

Element Shear Stresses and Strains (Figs. E-40 

through E-45) 521 

xxiii 



Acceleration Time Histories (Fig. E-46) 525 

Section 3 — Shell K2 ^g,^ = 130 - LUSH and QUAD4 

Element Shear Stresses and Strains (Figs. E-47 

through E-52) 526 

Acceleration Time Histories (Fig. E-53) 530 

Model Embankment — Shell K2 ^ax ^ ^^^ 

Effect of Poisson's Ratio on Stresses (Figs. E-54 

through E-55) 531 

F Embankment Response Model 533 

G Cyclic Triaxial Test Summaries of Modeled Oroville Gravel 

Tests. (Figs. G-1 through G-68) 541 

H Extrapolation of Isotropically-Consolidated Cyclic Triaxial Tests 

for Strength Interpretation II. (Figs. H-1 through H-25) 613 

I Cyclic Triaxial Test Results for Modeled Oroville Gravel Using 

Strength Interpretation II. (Figs. I-l through I-IO) 629 

J Procedure for Interpreting Cyclic Triaxial Test Data to Determine 

Cyclic Shear Stress on Potential Failure Plane 641 

K Cyclic Triaxial Test Results for Modeled Oroville Gravel Using 

Strength Interpretation I. (Figs. K-1 through K-IO) 649 

L (Available on request later in 1979) 

M Embankment Strain Potentials. (Figs. M-1 through M-4) 663 

PLATES 
(inside rear cover) 

Plate 1. Geology of Lake Oroville Area, Butte County, California 

Plate 2. Oroville Area Level Net Benchmark Locations 



CONVERSION FACTORS 
Metric to Customary System of Measurement 



Quantity 

Length 



Volume 



Flow 



Mass 

Velocity 

Power 

Pressure 

Specific 
capacity 

Concentration 

Electrical 
conductivity 

Temperature 



Metric Unit 

millimetres (mm) 

centimetres (cm) for snow depth 

metres (m) 

kilometres (km) 

square millimetres (mm^) 

square metres (m^) 

hectares (ha) 

square kilometres (km^) 

litres (I) 

megalitres 

cubic metres (m^) 

cubic metres (m^) 

cubic metres (m^) 

cubic dekametres (dam-^ ) 

cubic hectometres (hm-^) 

cubic kilometres (km-^) 

cubic metres per second (m-^/s) 

litres per minute (l/min) 

litres per day (I/day) 

megalitres per day (Ml/day) 

cubic metres per day (m^/day) 

kilograms (kg) 

tonne (t) 

metres per second (m/s) 

kilowatts (kW) 

kilopascals (kPa) 

kilopascals (kPa) 

litres per minute per 
metre drawdown 

milligrams per litre (mg/l) 

microsiemens per 
centimetre (//.S/cm) 

degrees Celsius ("C) 



Multiply by 


To get customary equivalent 


0.03937 


inches (in) 


0.3937 


inches (in) 


3.2808 


feet (ft) 


0.62139 


miles (m) 


0.00155 


square inches (in2) 


10.764 


square feet (ft2) 


2.4710 


acres (ac) 


0.3861 


square miles (mi^) 


0.26417 


gallons (gal) 


0.26417 


million gallons (10^ gal) 


35.315 


cubic feet (ft^) 


1.308 


cubic yards (yd^) 


0.0008107 


acre-feet (ac-ft) 


0.8107 


acre-feet (ac-ft) 


0.8107 


thousands of acre-feet 


0.8107 


millions of acre-feet 


35.315 


cubic feet per second (ft^/s) 


0.26417 


gallons per minute (gal/min) 


0.26417 


gallons per day (gal/day) 


0.26417 


million gallons per day (mgd) 


0.0008107 


acre-feet per day 


2.2046 


pounds (lb) 


1.1023 


tons (short, 2,000 lb) 


3.2808 


feet per second (ft/s) 


1 .3405 


horsepower (hp) 


0.145054 


pounds per square inch (psi) 


0.33456 


feet head of water 


0.08052 


gallons per minute per 




foot drawdown 


1.0 


parts per million 


1.0 


micromho per centimetre 


1 .8 y "C) + 32 


degree Fahrenheit ("F) 



CHAPTER I 

INTRODUCTION 



Oroville Dam (Figure 1) is situated in 
the foothills of the Sierra Nevada above 
the Sacramento Valley. The dam is 8 kilo- 
metres (5 miles) east of the City of 
Oroville and about 209 kilometres (130 
miles) northeast of San Francisco. 

On August 1, 1975, at 1320 hours (1:20 
p.m.) PDT, an earthquake of Richter Scale 
magnitude 5.7 occurred about 12 kilo- 
metres (7.5 miles) southwest of Oroville 
Dam. During the main event and the many 
aftershocks that followed, the Oroville 
facilities continued operating without 
interruption except for about a 45 minute 
shutdown of power generation. 

The Oroville earthquake sequence began 
with a number of foreshocks on June 28, 
1975. Then, on August 1, twenty-nine 
foreshocks occurred within 5 hours of 
the main shock. The largest of these had 
a magnitude of 4.8. Many aftershocks, 
the largest of which had a magnitude of 
5.1, occurred throughout August, and 
scattered shocks continued. 



Purpose 

Intensive investigations originating from 
the August 1, 1975, Oroville earthquake, 
were conducted to determine : 

1. Geologic and tectonic conditions. 

2. Fault mechanism and orientation. 

3. Crustal movements. 

4. Public safety as it relates to the 
Department's facilities. 

The results of these investigations are 
presented in detail in the following 
chapters . 

Description of the Oroville Facilities 

Oroville Dam and its appurtenances, along 
with the Thermalito facilities (Figure 2) , 
comprise a multiple purpose project, 
which includes water conservation, power 
generation, flood control, recreation. 




nc_'p^ 



Figure 1 . Orovi 1 le Dam 



GENERAL 
LOCATION 




% 



WEST ena»<CH 



H 



T T 

LAKE OROVILLE 
ELEV. 900' 



THERUALITO POWER CANAL y 




WESTERN PACIFIC R R 



YON CREEK BRIDGE 




BIDWELL BAR BRIDGE 



FE4THER RIVER 
Fish hstcheRY^ 

AND FISH 
BARRIER DAM 



•River outlet r 

TER' BUTTE CANAL OUTLET 



m 



VERLOOK V'^ eiDwELL CANYON 

\^2 



^ENTERPRISE BRIDGE 



''SflDOLE 0AM 



■/ 



kilometres 

5 
J I I I I 



-OROVILLE EARTHQUAKE 
EPICENTER M = 5.7 
AUGUST I, 1975 



Figure 2. Location Map, Oroville Facilities 



and fish and wildlife enhancement. The 
lake stores winter and spring runoff, 
which is released into the Feather River 
as necessary to supply project needs and 
commitments . The pumped-storage capa- 
bility of the Oroville facilities permits 
maximum use of peaking capabilities and 
increases the value of power produced by 
the releases. 

Water releases from Edward Hyatt Power- 
plant are largely diverted from the 
Feather .River at the Thermalito Diversion 



Dam (Figure 3) , a concrete gravity struc- 
ture with a radial gated crest section. 
These diversions pass through the Therma- 
lito Power Canal and Thermalito Forebay 
(Figure 4) , through Thermalito Powerplant, 
and into Thermalito Afterbay. The Therma- 
lito Diversion Pool, Power Canal, and 
Forebay have a coimion water surface to 
accommodate flow reversals for the pumped- 
storage operation. Thermalito Afterbay 
(Figure 5) stores the plant discharges 
for the pumped-storage or conventional 
operation and reregulates flow for \ini- 



form return to the Feather River. 

Migrating salmon and steelhead blocked 
by the Oroville Complex are diverted 
from the river into the Feather River 
Fish Hatchery at the Fish Barrier Dam, 
located 0.8 kilometre (0.5 mile) down- 
stream from the Thermalito Diversion Dam. 

The Investigating Organization 

On August 8, 1975, the Department of 
Water Resources convened its Consulting 
Board for Earthquake Analysis to review 
the general post-earthquake situation 
and the preliminary data assembled. On 
September 11 and 12, 1975, a Special 
Consulting Board for the Oroville Earth- 
quake, composed of the members of the 
Consulting Board for Earthquake Analysis 
and additional engineering consultants 
in the field of design and construction 
of dam and reservoir projects, was con- 
vened by the Department to review the 
Department's programs for data collec- 
tion and evaluation of structural 
seismic safety. 



The Special Consulting Board for the 
Oroville Earthquake consisted of the 
following members : 

George W. Housner, Chairman 

Clarence R. Allen 

John A. Blume 

Bruce A. Bolt 

Wallace L. Chadwick 

Thomas M. Leps 

Alan L. O'Neill 

Philip C. Rutledge 

H. Bolton Seed 

With recommendations from the Special 
Consulting Board, several earthquake- 
related investigations were undertaken 
by the Department. These include: 

1. Geologic studies and mapping of 
the epicentral area and the causa- 
tive fault. 

2. Seismological studies dealing with 
the earthquake sequence and fault 
plane resolution. 




Figure 3- Thermalito Diversion Dan 







Figure k. Thermal! to Forebay Dam 

3. Determination of the seismic safety 
of project structures under loading 
of the Reanalysis Earthquake 
selected for this area. 

4. Surveys of the area. 

5. Evaluation of the existing seismic 
instrumentation monitoring system. 

The Special Consulting Board was recon- 
vened on November 22 and 23, 1976, to 



review the progress of the investiga- 
tions and provide recommendations. On 
September 28 and 29, 1978, the Special 
Consulting Board was convened to review 
those completed reports on the investiga- 
tions and reanalysis of the Oroville 
facilities prior to their being published 
in this bulletin. Progress of the 
uncompleted reports was also reviewed. 

To meet regular State and Federal safety 
review requirements for the Oroville- 
Thermalito facilities, the Special Con- 
sulting Board was requested to be the 
Department's consultant for these 
requirements . 

Reports prepared by the Special Consul- 
ting Boards and Department ' s responses 
are included in Appendix A. 

Reanalysis of Project Structures for 
Earthquake Safety 

The performance of State Water Project 
facilities during the August 1975 earth- 
quake sequence demonstrated their ability 
to withstand this seismic loading. Only 
minor superficial damage was sustained 
by some of the secondary structures. 
The Department's Bulletin 203 (April 
1977) , Performance of the Oroville Dam 
and Related Facilities During the 
August 1, 1975 Earthquake , documents 
the performance of the Oroville complex 




Figure 5- Thermal ito Afterbay Dam 



during the main event and succeeding 
aftershocks. 

The Special Consulting Board reviewed 
the seismic environment of the Oroville 
facilities and recommended an earthquake 
motion be developed for reevaluation of 
the Oroville-Thermalito structures that 
has a magnitude of 6.5 producing a peak 
acceleration of 0.6g. Detailed develop- 
ment and characteristics of this 
Reanalysis Earthquake are described in 
Chapter V , 

A program for dynamic structural analy- 
sis of critical structures was imple- 
mented in cooperation with Professors 
H. Bolton Seed, A. K. Chopra, and Edward 
L. Wilson of the University of California, 
Berkeley. The critical structures to be 
reanalyzed are: Oroville Dam, Oroville 
Dam Spillway (flood control outlet 
structure) , Thermalito Diversion Dam, 
Thermalito Powerplant Headworks, Therma- 
lito Forebay Dam, and Thermalito Afterbay 
Dam. Dynamic strengths of the material 
in Oroville Dam were determined by large- 
scale laboratory testing conducted by 
the University of California. Explora- 
tion and soils testing for the Thermalito 
Forebay and Afterbay Dams were conducted 
by the Department of Water Resources. 
Summary of Conclusions 
and Recommendations 

The following are conclusions and recom- 
mendations from the chapters in this 
Bulletin. 

Geological Investigations (Chapter II) 

1. The August 1, 1975, Oroville earth- 
quake was accompanied by movement 
on the previously unrecognized 
Cleveland Hill Fault. A linear 
zone of discontinuous ground crack- 
ing developed along the fault about 
7 kilometres (4.3 miles) east of 
the main shock epicenter. 

2. Initial length of ground rupture 
on the Cleveland Hill Fault was 
about 1.6 kilometres (1.0 mile). 
Over a period of about 12 months 



the ground cracking extended pro- 
gressively to the north, reaching 
a total length of 8.5 kilometres 
(5. 3 miles) . 

3. Offset along the fault was greatest 
in the southern segment, where the 
original cracking occurred. Offset 
increased with time; movement 
amounted to about 50 millimetres 

(2 inches) vertical displacement 
and 25 millimetres (1 inch) hori- 
zontal extension. 

4. The Cleveland Hill Fault was not 
encountered by trenching or geo- 
physical investigation north of 
Mt. Ida Road. Aftershock hypo- 
centers projected up a calculated 
fault plane indicate the fault at 
the ground surface trends into 
Bidwell Canyon and that it may pass 
beneath Oroville Dam at depth. 

5. Trenching across the Cleveland Hill 
Fault by Department of Water 
Resources and others provides 
evidence for multiple small fault 
displacements during the past 
100,000 years. These displace- 
ments would likely have produced 
earthquakes similar to the 1975 
Oroville event. 

5. Three major lineament- fault zones, 
the Paynes Peak, Swain Ravine, and 
Prairie Creek, have been delineated 
in the area by geologic studies . 
These lineament-fault zones are 
complex bands of discontinuous, inter- 
twined, steeply dipping faults which 
were formed during Mesozoic or ear- 
lier time under the influence of a 
different tectonic stress regime 
that exists today. The Cleveland 
Hill Fault is within the Swain 
Ravine Lineament fault zone. 

7. Most Cenozoic fault movements in 

the Sierran foothill belt are caused 
by east-west extensional stresses 
reactivating pre-existing Paleozoic 
and Mesozoic faults such as those 
comprising the lineament-fault zone. 



8. Historic (Cenozoic) faulting and 
historic earthquake records in the 
foothill region demonstrate that 
the current and long-range level of 
seismic activity is one of low- to 
moderate-magnitude earthquakes at 
relatively long recurrence intervals, 
occasionally resulting in minor 
ground rupture and offset. 

9. Nothing was seen in this geologic 
study to indicate that earthquakes 
greater than Richter Magnitude 6.5 
should be expected in the Oroville 
area. 

10. Maximum offset that should be anti- 
cipated from another Oroville-type 
earthquake is estimated to be 50 
millimetres (2 inches) of vertical 
displacement and 25 millimetre (1 
inch) horizontal extension. For a 
somewhat larger event displacement 
might be several times larger than 
these values, along north-south 
trending faults . 



H- The evidence available does not 
indicate a causal relationship 
between Lake Oroville and the earth- 
quake, but the possibility cannot 
be eliminated conclusively at this 
time. 

Seismology (Chapter III) 

Since August 1, 1975, a correlation is 
not indicated between the Lake Oroville 
water surface variations and the rate 
of occurrence of Oroville aftershocks. 

Within the boundary of the aftershock 
zone north of 39 26 'N latitude, vertical 
cross-sectional plots indicate that the 
Cleveland Hill Fault is a single, well 
defined break, dipping to the west at 
about 60 with the horizontal and with a 
near north-south strike. Vertical cross- 
sectional plots south of 39 26 'N indicate 
that the fault breaks along more than one 
plane. 



Vertical and Horizontal Geodesy 
(Chapter IV) 

Vertical Crustal Movements 

The following conclusions are based on 
free adjustment holding the elevation of 
OM-27 fixed (1967 USC&GS adjustment) and 
therefore all elevations differentials 
are relative to OM-27. 

1. Based on the preearthquake datum of 
1967, the greatest elevation differ- 
ential was only 63 millimetres (0.207 
foot) on line Olive during the ten- 
year epoch (1967-1977) . 

2. The August 1, 1975, Oroville earth- 
quake is associated with minor sub- 
sidence in the Oroville area, mainly 
south and southwest of Lake Oroville. 

3. Most of the subsidence associated 
with the August 1, 1975, Oroville 
earthquake was measured between late 
August 1975 to October 1976. 

4. The elevation differentials show 
movement of the fault zone that 
passes through the level lines 
Cleveland Hill and Mission Olive 
(ground cracking was evident before 
the lines were established) . A fault 
zone may pass through the survey line 
Miners Ranch south of Lake Oroville; 
however, no ground cracking was 
found there. 

5. Minor subsidence of less than 25 
millimetres (0.082 foot) has been 
measured adjacent to Oroville Dam 
and Lake between 1967 to 1977 due to 
all causes. 

Horizontal Crustal Movements 

1. All computed horizontal movements are 
minor and in many cases within the 
accuracy of the existing surveys and 
computations . 

2. The August 1, 1975, Oroville earth- 
quake did not cause sufficiently 
large horizontal movements that could 



be reliably measured and calculated 
within the Lake Oroville Monitoring 
Network . 

Oroville Dam Evaluation of Seismic 
Stability (Chapter V) 

1. The seismic stability of Oroville 

Dam was investigated for the Reanal- 
ysis Earthquake of Richter Magnitude 
6.5, at a hypocentral distance of 
5 kilometres (3 miles) from the dam, 
and producing the following ground 
motion characteristics at the base 
of the dam: 



maximum acceleration 
predominant period 
duration 

acceleration time 
history 



0.6g 

0.4 seconds 
20 seconds 
modified Pa- 
coima plus 
modified Taft 



2. 



It was concluded that this ground 
shaking was more severe than any 
future shaking likely to affect the 
dam. 

Using "best judgment" choices for 
input soil properties and conditions, 
relatively small embankment deforma- 
tions were estimated by the seismic 
evaluation procedures . It is con- 
cluded that Oroville Dam would per- 
form satisfactorily if subjected to 
the Reanalysis Earthquake. 



Oroville Dam Flood Control Outlet 
Structure (Chapter VI ) 

The investigations performed indicate 
that when the Oroville Flood Control 
Outlet Structure is subjected to the 
Reanalysis Earthquake ground motion it 
is stable, and that expected compressive 
and tensile stresses are within the 
allowable limits established for the 
structure . 



Based on results of dynamic analyses 
and available data for concrete strength, 
it is concluded that Thermalito Diver- 
sion Dam should be able to resist the 
stresses expected during the earthquake 
(Reanalysis Earthquake) ground motion 
specified by the State Department of 
Water Resources. 

Reappraisal of Secondary Structures 
(Chapter VIII) 

Fish Barrier Dam - A review of the orig- 
inal design of this dam indicated maxi- 
mum compressive and tensile stresses of 
approximately 120 psi and 10 psi respec- 
tively. A check of stability using a 
pseudostatic analysis and seismic coeffi- 
cients of 0.25 and 0.6 indicated a shear 
friction factor of safety in excess of 9. 
Based on this finding, no additional 
seismic analysis is recommended for the 
Fish Barrier Dam. 

Edward Hyatt Powerplant - The powerhouse 
substructure has been reviewed using a 
comparative pseudostatic analysis of 
previously designed powerhouse substruc- 
tures. Based on this comparison, it has 
been determined that this substructure 
would be capable of resisting the forces 
induced by a 0.25g peak ground accelera- 
tion; therefore, no modifications are 
required. 

Modifications will be made to improve 
the seismic resistance of powerhouse 
superstructure components as necessary. 

Thermalito Powerplant - The powerhouse 
substructure has been reviewed using a 
comparative pseudostatic analysis of 
previously designed powerhouse substruc- 
tures. Based on this comparison, it has 
been determined that this substructure 
would be capable of resisting the forces 
induced by a 0.25g peak ground accelera- 
tion; therefore, no modifications are 
required . 



Thermalito Diversion Dam (Chapter VII ) 
Conclusion by Dr. A. K. Chopra: 



Modifications will be made to improve 
the seismic resistance of powerhouse 
superstructure components as necessary. 



Miscellaneous Structures - Damage that 
may occur to the miscellaneous structures 
are not considered to be a threat to 
public safety and property. For the 
purpose of the seismic reevaluation 
these structures are classified as 
noncritical. 

Bridges - Bridge components that will 
not sustain the forces generated by a 
0.25g peak ground acceleration will be 
modified to strengthen their seismic 
resistance. 

Switchyards - Based on the considera- 
tion that failure of electrical equip- 
ment in the Edward Hyatt or Thermalito 
Powerplant switchyards does not pose a 
threat to public safety or property, the 
switchyards are classified as noncritical 
elements of the Oroville Complex. 

Contingency Plan for Seismic Emergencies 
(Chapter IX ) 

The contingency plan is attentive to 
established Division Policy; it provides 
for detection, notification, and response 
to seismic events. The plan also includes 
a list of operational facilities and 
features along with criteria that must be 
met before returning to preearthquake 
operating status . 

Department's Findings 

Based on the preceding conclusions from 

the investigations completed to date, 

the Department concludes that these facil- 



ities do not pose a threat to public 
safety. 

Uncompleted Reports 

The reanalysis of the Thermalito Power- 
plant Headworks and the Thermalito Fore- 
bay and Afterbay Dams has not been 
completed. Dr. A. K. Chopra (the 
Department's consultant) is currently 
reanalyzing the Thermalito Powerplant 
Headworks. The Department is continuing 
with the reanalysis of the Thermalito 
Forebay and Afterbay Dams. Investiga- 
tions of the Bidwell Canyon Saddle Dam 
and the effect of fault movements with 
respect to the Oroville-Thermalito facil- 
ities will also be completed. These 
reports are planned for completion early 
in 1979 and publication by mid-1979. 

Safety Review Requirements 

The completion of the investigations and 
reanalysis of the Oroville facilities 
with concurrence by the Special Consul- 
ting Board fulfills the following two 
safety requirements: 

1. The five-year Federal Energy Regu- 
latory Commission's Part 12 Safety 
Inspection Report. 

2. The five-year Department of Water 
Resources, Division of Safety of 
Dams ' safety review under the regu- 
lations of the California Administra- 
tive Code, Title 23, Article 4, 
Sections 340-343. 



Report of the Special Consulting Board 
For the Oroville Earthquake 



November 15, 1978 
(see next page) 



15 November 1978 



Report of the Special Consulting Board for the Oroville Earthquake 



Mr. Howard H. Eastin, Chief 

Division of Operations and Maintenance 

Department of Water Resources 

P.O. Box 388 

Sacramento, California, 95802 



At meetings on September 28 and 29, 1978, of the Special Consulting 
Board for the Oroville Earthquake, staff members of the Department of Water 
Resources made presentations relative to the Department's draft of its 
report, "August 1, 1975, Oroville Earthquake Investigation - Bulletin 
203-78." The Board has also reviewed chapters of the draft prior to the 
meeting. At the conclusion of the meetings, the Board was asked to respond 
to the following questions relating to the content of the Oroville Earthquake 
Investigation report. Our responses are presented below. 

Question No. 1 . Does the Board concur with the conclusions and the 
recommendations set forth in the Summary of Conclusions and Recommendations? 

Response . The Board has reviewed the draft of the chapters of Bulletin 
203-78 and has heard the presentations of the staff members, and has reviewed 
the "final Draft Chapter 1" which contains the Summary of Conclusions and 
Recommendations. 

The Board concurs with the conclusions and the recommendations set forth 
in the Summary of Conclusions and Recommendations in the October 24, 1978, 
"final draft" of Chapter 1. 



Question No. 2 . Does the Board agree that the Oroville Division's 
critical structures (except Thermal i to Forebay and Thermal i to Afterbay Dams) 
would perform adequately with respect to public safety during the adopted 
earthquake ground motions? 

Response . The Board agrees that the critical structures (except 
Thermal ito Forebay and Afterbay Dams which have not yet been analyzed) would 
perform adequately with respect to public safety if subjected to the adopted 
earthquake ground motions. 

Question No. 3 . Does the Board have any comments on the studies completed 
to date for the seismic stability of Thermal ito Forebay and Afterbay Dams? 

Response . The Board does not have any comments except to urge completion 
of the studies at an early date. 

Question No. 4 . The Department intends to publish the results of the 
Thermal ito Power Plant - Headworks, Seismic Evaluation, and Thermal ito Forebay 
and Afterbay, Evaluation of Seismic Stability, reports next year, possibly as 
Bulletin No. 203-79. Does the Board consider another meeting necessary or 
could the Board's review, comments and report be handled by correspondence? 

Response . The Board feels that another meeting would be desirable at 
which staff members would present the results of the analyses and respond to 
questions from the Board. 

Question No. 5 . Does the Board have any other comments or recommenda- 
tions to make at this time? 



10 



-3- 



Response . 

a. The Board recommends that in Chapter 2-Contingency Plan for Seismic 

1/ 2/ 

Emergencie^the criteria for notification given on page 5 "5e revised to read 

base 
that O.lg recorded at the e^est of Oroville Dam will replace the 3.0 Richter 

Scale criteria for notification, and 0.15g will replace the 4.0 Richter Scale 

criteria. It is also recommended that Annual Earthquake Drills be held so 

that personnel will be prepared to act in the event of an actual earthquake. 



3./ 

b. With reference to Chapter 5, the Board recommends that DWR request 

NCAA to perform a precise level survey from a known, stable benchmark to 
benchmark OM 27, as this benchmark is a key element in the geodetic survey 
at Oroville Dam. 

II 

c. With reference to Chapter 5, the Board recommends that resurveys 

be made at 5 year intervals, or after significant earthquakes, to monitor 
crustal movements that might have taken place. 



The Board was favorably impressed by the investigations and analyses 
carried out by the Department, and commends the Department for the 
diligence and thoroughness of its work. 



1_/ Chapter IX in this bulletin. 
2/ See page ^7 in this bulletin. 
.2/ Chapter IV in this bulletin. 



11 




Respectfully Submitted, 



George W. Ho\lsi>er 





C. R. Alren 




J^ 



^ .^ .^ r'^/- 



John A. Blume 



rhu^^ Qy/^^J^ 



Bruce A. Bolt 



sMi^ 



T. M. Laps 




Alan L. O'Neill 



Philip a. Rutledge / 



^. (i^^Uo^ <^^^'<^ 



H. Bolton -S€ed »" 



^ \ r -, «„ — I ) i u^,^,,-;'^ i , 1-' — ' ' 



"Wallace h, (JhaawicK 
12 



REVIEW BY THE DIVISION OF MINES AND GEOLOGY 



The Department requested the Department of Conservation, 
Division of Mines and Geology to review the final draft of 
Bulletin 203-78 

The response from the Division of Mines and Geology 
on Bulletin 203-78 follows. 



13 



State of California The Resources Agency 

Memorandum 

To : Clifford V. Lucas, Chief Civil Maintenance Branch Date: December 22, 1978 
Division of Operations and Maintenance 
Department of Water Resources 
]h]6 9th Street 
Sacramento, California 9581^* 



From Department of Conservation 

Division of Mines and Geology 
1416 -9th Street, Sacromento 95814 



Subject: Review of DWR Bulletin 203-78 Final Draft 



This is our response to your request of December 11 that the Division of 
Mines and Geology provide review of final draft of DWR Bulletin 203-78. 
Our commentary is limited to one observation in Chapter IV, Sei smology . 1/ 

2/ 
The conclusion (page 5)~ that there is no correlation between water levels 

and the rate of earthquakes does not consider the strengths of the earthquakes, 

but is based on a count of events regardless of magnitude. When the strengths 

of the earthquakes are considered, the resulting pattern of seismic strain 

released might be related to water levels. Most of the seismic strain 

appears to be released following episodes of filling, and very little strain 

release appears to occur during the actual filling. 



We appreciate the opportunity to comment on this useful publication. 



nes F. Davis 
State Geologist 



cc: Pri sci 1 la C . Grew 





1_/ Chapter III in this bulletin, 
2/ Page 12^4- in this bulletin. 



14 



CHAPTER II 



GEOLOGIC INVESTIGATIONS 



On August 1, 1975, a strong earthquake 
occurred near Oroville, California. The 
earthquake sequence began June 28, 1978, 
with the occurrence of several fore- 
shocks; the largest of these foreshocks 
was magnitude 3.8. From July 8 through 
July 31, only five foreshocks occurred, 
giving the appearance that earthquake 
activity was ceasing. Then on August 1, 
twenty-nine foreshocks, the largest of 
which was magnitude 4.8, occurred within 
five hours prior to the magnitude 5.7 
main shock at 1320 hours Pacific day- 
light time. 

The hypocenter for the main shock of the 
Oroville earthquake series was approxi- 
mately 1 km (0.6 mi) east-northeast of 
Palermo at a depth of 8.8 km (5.5 mi). 
Fault movement ruptured the ground 
approximately 7 km (4 mi) east of 
Palermo. This ground rupture is called 
the Cleveland Hill Fault. 

Prior to the 1975 earthquake, seismic 
hazard was not regarded as being great 
in the Oroville area. It was recognized 
that earthquakes do occur in the sur- 
rounding region, and that the largest 
recorded earthquake was a magnitude 5.7 
in 1940 north of what is now Lake 
Oroville. Fault movements were not con- 
sidered likely to occur along faults in 
the area. Because fault movement occur- 
red where no fault was suspected before 
the Oroville earthquake, it was obvious 
that existing geologic information did 
not identify potentially active faults 
in the Oroville area. Therefore, geo- 
logic investigations were started 
immediately. 

PURPOSE OF THE INVESTIGATION 

The purposes of geologic investigations 
were threefold: (1) to understand the 
geologic and tectonic conditions which 
caused the Oroville earthquake; (2) to 



evaluate any potential hazards; and 
(3) try to determine if Oroville 
Reservoir caused the earthquake. 

The original thrust of investigation was 
to determine if the Cleveland Hill Fault, 
movement along which caused the 1975 
earthquake, was an old or new fault, and 
if it extended northward to endanger 
Department facilities in the Oroville 
area. The investigations later were ex- 
panded to cover the surrounding area. 

PREVIOUS WORK 

The area geology was first mapped in the 
late 1800 's by Waldemar Lindgren and 
Harry Turner. Their work was published 
as the Bidwell Bar folio (Becker and 
others, 1898) and the Smartsville folio 
(Lindgren and Turner, 1895) of the U. S. 
Geological Survey Atlas of the United 
States series. These works include the 
Bangor, Oroville Dam, and Berry Creek 
quadrangles which comprise half of our 
study area (Figure 6) . 

The northern Sierra foothills were again 
the center of detailed mapping in the 
1950 's for studies of the Merrimac plu- 
ton (Hietanen, 1951) and the Bidwell Bar 
area (Compton, 1955). In 1955, Robert 
Creely completed a doctoral mapping 
thesis of the Oroville 15-minute quad- 
rangle, which includes the Hamlin Canyon 
Cherokee, Shippee, and Oroville 
7-1/2 minute quadrangles. Creely 's 
work was published in 1965 as Bulletin 
184 by the California Division of Mines 
and Geology. 

Areas of the Oroville, Oroville Dam, 
Cherokee and Berry Creek quadrangles 
were mapped in detail during the late 
1950 's and early 1960 's by geologists 
from the Department of Water Resources 
for studies associated with the con- 
strutruction of Oroville Dam and related 



15 



APPROXIMATE LIMIT OF 
RECONNAISSANCE MAPPING 



Lake 
A I manor 




Figure 6. Location map of six-quadrangle study area, 



16 



facilities. Mapping to the south of 
the area by other individuals and 
agencies includes preliminary work in 
the Bangor quadrangle (Quintin Aune, 
unpub. data), graduate theses in the 
Smartville area (Buer, 1978; Costas 
Xenophontos, in progress), a regional 
study by the U. S. Army Corps of 
Engineers (1977) for the proposed Parks 
Bar Dam on the Yuba River, and fault 
studies made by Woodward-Clyde 
Consultants for both the U. S. Army 
Corps of Engineers and Pacific Gas and 
Electric Company. Additionally, 
Woodward-Clyde Consultants (1977) con- 
ducted a regional fault investigation 
in the northern foothill belt for the 
U. S. Bureau of Reclamation's Auburn 
Dam project on the American River. 

Numerous other references contribute to 
a general understanding of structural 
and stratigraphic relationships in the 
Oroville area. Some of these include 
the Geologic Map of California, Westwood 
Sheet (Lydon and others, 1960), Geology 
of the Richardson Springs quadrangle 
(Burnett and others, 1969), and work by 
Hietanen (1973a, 1976, 1977). An under- 
standing of the importance of faulting 
and some knowledge of its history and 
nature is provided in the works of 
Clark (1960, 1964, 1976), Cebull (1972), 
Duf field and Sharp (1975), and 
Schwieckert and Cowan (1975). Models 
and evidence for plate tectonic evolu- 
tion of the area are adopted and modi- 
fied after Hamilton (1969) , Moores 
(1972), Schweickert and Cowan (1975), 
and Schweickert (1976). 

SCOPE OF THE INVESTIGATION 

Several types of imagery were used to 
find major structural trends and linea- 
ments prior to field mapping. These 
included satellite imagery, high- 
altitude black-and-white and infrared 
photographs , radar imagery (SLAR) , and 
low-altitude black-and-white and color 
photographs. 

Satellite imagery was from the ERTS 
(now LANDSAT) program of the U. S. 



Government (NASA) . Radar imagery was 
obtained from Woodward-Clyde Consultants. 
High- and low-altitude photography is 
comprised of four sets flown for the 
Department of Water Resources, including 
a set of low-altitude low-sun-angle 
photographs, and one set from the 
U. S. Forest Service. 

Detailed geologic mapping was done in 
the Palermo, Bangor, Oroville, Oroville 
Dam, Cherokee and Berry Creek 7-1/2 min- 
ute quadrangles (Figure 6). Reconnais- 
sance geologic mapping was done in the 
remainder of the study area, extending 
northwest some 70 km (43 mi) from 
Oroville. 

Detailed mapping was generally done on 
a 7-1/2 minute quadrangle base. Where 
more detail was desired, mapping was 
done on quadrangles enlarged to 1:12,000. 
Areas mapped in detail were covered on 
foot; samples of rock were collected, 
photographed or sketched where appropri- 
ate, and mapped. Most of the rock units 
were sampled and petrographic analyses 
were made by Costas Xenophontos of the 
University of California at Davis. 

Reconnaissance mapping was done on 
7-1/2 minute topographic base maps wher- 
ever possible. Much of the northern 
study area was mapped on preliminary 
topographic maps received from the U. S. 
Geological Survey. Areas along photo 
lineaments and faults indicated on the 
State geologic maps were checked in de- 
tail. In other areas we relied on pre- 
vious works and field reconnaissance. 

Mapping began in May 1976, and continued 
until May 1978. A total of 55 work 
months went into field work and report 
writing. The majority of mapping was 
done by Department of Water Resources 
geologists; three geology graduate stu- 
dents from the University of California 
at Davis assisted during July and 
August 1977. 

Subsurface information was used whenever 
and wherever possible to map rock units 
and, especially, to aid in tracing faults 



17 



J 

r 






• •••••• A^ A 

• • • • /f %• A • 3«$ 




10 20 30 40 




U ---* 



SACRAMENTO 



MAGNITUDE 

5.0 - 6.0 

4.0 - 4.9 

• 3.0 - 3.9 

• 2.0 - 2.9 

A NOT KNOWN 

5 (number denotes multiple epicenter) 



Lake 
Tahoe 

39°-v 

vX 



Figure 7- Historic earthquakes within a 100-kilometre (62-mile) radius 
of Orovi lie. 



18 



beyond surface exposures. Subsurface 
information was derived from exploratory 
trenches, road cuts, railroad cuts and 
utility trenches; existing tunnel logs, 
well logs, test boring logs, mining re- 
ports and Department of Water Resources 
design and construction reports were 
also reviewed. Geophysical methods were 
used to try to trace the Cleveland Hill 
Fault beyond the northernmost ground 
cracks. 

Exploratory trenches in this study in- 
cluded 17 by the Department of Water 
Resources in the Oroville-Bangor area 
(trench logs, page 103 ff.) and several 
others by Pacific Gas and Electric 
Company, U. S. Army Corps of Engineers 
and the U. S. Bureau of Reclamation. 
The latter agencies were studying fault- 
ing in the Sierran foothills from Sonora 
to Oroville. Woodward-Clyde Consultants 
were involved as consultants to these 
three agencies. 

Trenches by the Department of Water 
Resources and other agencies were pri- 
marily used to explore suspected faults. 
Many were cut across surface cracks 
shortly after the August 1, 1975, earth- 
quake. Seismic and resistivity surveys 
were used by Department of Water 
Resources geologists, assisted by Elgar 
Stevens of the Department of Transporta- 
tion, in attempts to trace faulting asso- 
ciated with surface cracking. 

Tunnel logs reviewed include tunnels 1 
through 5, Western Pacific Railroad 
relocation by the Department of Water 
Resources. Logs for Miners Ranch and 
Kelly Ridge tunnels by Bechtel Corpora- 
tion for Oroville-Wyandotte Irrigation 
District were also reviewed. 

The area around Oroville Reservoir and 
south to Wyandotte was surveyed immedi- 
ately after the August 1, 1975, earth- 
quake by teams from the Department of 
Water Resources and the U. S. Geologi- 
cal Survey. It was hoped that precise 
leveling surveys across the Cleveland 
Hill Fault would indicate the sense of 
movement on this structure. 



The Department of Water Resources survey 
team ran additional first-order leveling 
surveys at regular intervals. Changes 
in elevation were contoured by computer 
and compared with geological and seis- 
mological data collected by the Depart- 
ment of Water Resources staff. 

Most precise leveling surveys around 
Oroville post-date the August 1975, 
earthquake. A comparison of August 1975, 
first-order leveling data with third- 
order leveling data compiled in the 
1940' s would be the closest approximation 
of crustal movements from the Oroville 
earthquake. 

SEISMIC HISTORY 

Epicenters for earthquakes occurring in 
the Oroville area prior to 1934 are esti- 
mated from newspaper accounts and reports 
by local residents. After 1934, epicen- 
ters in the area have been instrumentally 
determined. Historic seismicity in the 
Oroville area from 1851 to 1975 is shown in 
Figure 7 with a few recent earthquakes, 
providing more reliable data, indicated 
by their date above the epicentral 
location. 

Numerous low- to moderate-magnitude 
earthquakes have occurred in the northern 
Sierra Nevada in historic time (Townley 
and Allen, 1939; Wood and Heck, 1951). 
The most significant event affecting the 
Oroville area occurred near Virginia 
City, Nevada on December 27, 1869. An- 
other earthquake shook the Oroville re- 
gion on January 24, 1875, and is believed 
to have originated from movement of the 
Mohawk Valley Fault (Wolfe, 1967) loca- 
ted approximately 70 km (43 mi) to the 
northeast; reinterpretation of this 
earthquake suggests it was located south- 
west of Oroville (Paul Morrison, person, 
commun., 1978). 

An earthquake on February 8, 1940, cen- 
tered 54 km (34 mi) north of Oroville, 
is comparable in magnitude to the 1975 
Oroville event. Seismic monitoring was 



19 




Figure 8. Aftershock locations of the Oroville earthquake. 
The envelope encloses 90 percent of the epicenters. 



20 



not good in 1940 and data on the earth- 
quake are poor. A recent reassessment 
of these data shifted the epicentral 
location 40 km (25 mi) and changed the 
estimated magnitude from 6.0 to 5.7 
(Morrison, 1974, p. 8). The epicenter 
does not fall on any known fault, and 
no fault plane solution is obtainable 
from the seismic data. 

The May 24, 1966, "Chico" earthquake 
(Figure 7) is unique for this area be- 
cause adequate data were obtained to 
calculate a fault plane solution. The 
timing of this earthquake coincided with 
a crustal determination experiment using 
a subsurface explosion off the Northern 
California coast. This magnitude 4.6 
earthquake had a focal depth of 21 km 
(13 mi) and provided a fault plane solu- 
tion for a N30W strike and a 65 degree 
northeast dip, with dominant right- 
lateral movement (Lomnitz and Bolt, 
1967). 

THE 1975 EARTHQUAKE SERIES 

Foreshocks of the Oroville earthquake 
began June 28, 1975. Fourteen foreshocks 
with magnitudes from 2.1 to 4.7 occurred 
prior to the main shock at 1320 hours 
Pacific daylight time, on August 1, 1975. 

The main shock had a Richter magnitude 
of 5.7 and was centered about 1 km 
(0.6 mi) northeast of Palermo and 12 km 
(7.5 mi) south of Oroville Dam. The 
focal depth was 8.8 km (5.5 mi). 

Oroville earthquake aftershocks, includ- 
ing more than 5,600 recorded events, 
were still occurring in June 1978. Fre- 
quency of aftershocks has decreased 
steadily, from over 700 per day in early 
August 1975, to ten per month in 
June 1978. 

The region of aftershocks increased in 
size for several months after the 
August 1, main shock (Lester and others, 
1975); the most rapid increase was dur- 
ing the first week of this time period. 



Surface area envelopes of aftershock 
epicenters are ellipsoidal with the elon- 
gated axis oriented north-south. This 
pattern matches the north-south trend of 
the Cleveland Hill Fault in the subsur- 
face. Initially the aftershock zone 
expanded to the east, with the shocks 
occurring at shallower depths. Then 
the zone expanded both to the north and 
to the south, with the final expansion 
being to the north. An envelope con- 
taining the majority of aftershocks is 
shown in Figure 8. Aftershock epicen- 
ters have occurred as far north as 
Oroville Dam, and a few have occurred 
east of the Cleveland Hill Fault rupture. 

Depths of aftershocks were 8 to 9 km (5 
to 6 mi) on the west and became shal- 
lower toward the east . A fault plane 
solution shows the Cleveland Hill Fault 
strikes N25W to N5E (Beck, 1976; Clark 
and others, 1976; Morrison and others, 
1976), dips 60 degrees west and passes 
5 km (3 mi) beneath Oroville Dam (Lahr 
and others, 1976). First motion fault 
analyses indicate a dip-slip movement 
with the west side down; this movement 
indicates normal faulting and tensional 
deformation. 

Property damage caused by the earthquake 
consisted of fallen plaster, toppled 
chimneys, broken windows, items thrown 
from shelves, etc. Major damage included 
homes shifted off their foundations and 
an older brick home that was damaged be- 
yond repair. Damage to State facilities 
was very minor and consisted mostly of 
non-structural plaster cracks in build- 
ings and some settlement and cracking 
of uncompacted fill embankments around 
Thermalito Afterbay. 

Several changes in ground water occurred 
after the earthquake: (1) a few wells 
and springs went dry and others tempor- 
arily increased their flow, and (2) new 
springs appeared where none had been 
prior to the earthquake. 



21 




Figure 9. Locations of the Cleveland Hill and Mission Olive crack zones and 
sites of Department of Water Resources exploration trenches 



22 



GROUND CRACKING 

Investigations were made of reported 
ground cracking during the first few 
days after the earthquake. Most cracks 
observed were lurch or settlement indu- 
ced. Two areas of cracking did emerge 
as worthy of further investigation. 
These two zones are the "Palermo crack 
zone", trending northwest through the 
epicentral area, and cracking along 
Cleveland Hill Fault. 

The "Palermo crack zone", the western- 
most crack zone, consisted of discontin- 
uous cracks at a number of locations. 
Collectively, the crack sites are align- 
ed in a linear northwest trend. The 
crack zone lines up with the Prairie 
Creek Lineament to the southeast . The 
entire crack zone extended from 3 km 
(1.9 mi) south of Palermo to just south 
of Oroville for a total length of about 
9 km (5.6 mi). Additional cracking, 
discovered later during the investiga- 
tion, appeared to have occurred a con- 
siderable time after the main earthquake. 

The linear trend of the Palermo crack 
zone gave rise to the suspicion that the 



disturbance occurred along a fault zone, 
subsequent investigations did not reveal 
faulting but it is not really clear 
whether or not a subsurface fault extends 
northwest along the trend of the Prairie 
Creek Lineament. 

Cracking near Cleveland Hill (Figure 9) 
was first detected on August 6, 1975. 
Cracking occurred in an en echelon pat- 
tern and formed a discontinuous zone 1 
to 7 m (3 to 22 ft) wide and 1.6 km 
(1 mi) long. In 6 weeks some cracks 
widened to 40 mm (1.6 in) and showed up 
to 30 mm (1.2 in) of downdrop on the 
west side (Figures 10 and 11) . The 
cracks became more pronounced with time 
and some were still visible 3 years after 
the main shock. 

Early in October 1975, two approxi- 
mately parallel crack zones were found 
north of Cleveland Hill. These cracks, 
named the Mission Olive crack zone 
(Figure 12), are 400 m (1,300 ft) apart 
and extend discontinuously for about 
2 km (1.25 mi) to the north. The Mission 
Olive crack zone initially was seen in 
the asphalt paving on Mission Olive Road 
and Foothill Boulevard (Figure 9). The 

, ^.< 




Figure 10. Ground cracking that resulted 
from the August 1, 1975, Oroville 
earthquake. 



Figure 11. Close-up view of a ground 
crack on southwest slope of Cleveland 
Hill. 



23 




Figure 12. Locations of ground cracking from the Oroville earthquake and 
major photo lineaments in the southern study area 



24 



cracks definitely were not present in 
August 1975, when all paving between 
Cleveland Hill and the reservoir was 
inspected. 

In February 1976, another 182 m (600 ft) 
of cracking was found in a pasture 3 km 
(1.9 mi) south of Bidwell Canyon Saddle 
Dam. This cracking was not present in 
October 1975, when the site was selected 
for an exploration trench on the Swain 
Ravine Lineament by Department of Water 
Resources geologists. In August 1976, 
another 116 m (380 ft) of cracking was 
found north of the pasture in an olive 
orchard. The crack zone now reached 
nearly to Mt. Ida Road which is 2.1 km 
(1.3 mi) south of Bidwell Canyon Saddle 
Dam (Figure 13) . 

In February 1977, an additional short 
segment of ground cracking was found 
along the Swain Ravine Lineament 1.8 km 
(1.1 mi) south of the southern end of the 



Cleveland Hill cracks. This is the only 
instance of crack zone development pro- 
gressing southward. Periodic field 
checks after discovery of the last crack 
zone have shown no new cracking to the 
north or south. 

Initial cracking from the Oroville 
earthquake occurred only along the Swain 
Ravine Lineament west of Cleveland Hill. 
During the 12-month interval following 
the main shock, cracking propagated sev- 
eral kilometres northward. Timing of 
northward extensions of cracking is not 
precisely known and therefore cannot be 
related to specific aftershocks; exten- 
sion of cracking is probably related to 
continuing readjustment on the fault. 
Overall extent of cracking is 8.5 km 
(5.3 mi) and the collective length of 
individual zones is 5.3 km (3.3 mi). 
Displacement on individual cracks was 
most pronounced in areas of initial 
cracking. 



OroviHe Dam 



Saddl® Oam» 




Figure 13- Aerial north view of northern limit of ground cracking (Cleveland 
Hill Fault) just west of Wyandotte Miners Ranch Road. These cracks project 
toward the Bidwell Canyon Saddle Dam. 



25 



In addition to the progressive increase 
in the areal extent of cracking with 
time, displacements along the cracking 
also increased as time passed. The 
increase in displacement was discernible 
by eye, particularly in the first few 
months following the earthquake. The 
most pronounced visual impact of dis- 
placement increase was near Cleveland 
Hill where ground cracking was first 
discovered. Here a noticeable vertical 
displacement (down to the west) developed 
with time and the cracks became wider. 

The U. S. Geological Survey installed 
five arrays of survey monuments along 
the southern crack zone near Cleveland 
Hill to measure fault displacement. 
Initial surveys on some of these arrays 
were made as early as August 12, 1975, 
11 days after the earthquake. Maximum 
displacements measured by April 1976, 
(Philip Harsh, person, commun., 1978) 
were 34.1 mm (1.3 in) extensional sepa- 
ration and 28.8 mm (1.1 in) vertical 
displacement for the August 12, 1975, 
to April 4, 1976, period, with a com- 
puted dip-slip component of 44.6 mm 
(1.8 in). Eight millimetres (0.3 in) 
of right lateral displacemnt also were 
measured. Displacement measured in 
April 1976, on the other four arrays 
were markedly smaller, ranging down to 
a minimum of 4.3 mm (0.2 in) vertical 
displacement and 0.6 mm (0.02 in) exten- 
sional separation for the October 1975, 
to April 1976, period. 

The U. S. Geological Survey surveyed the 
arrays again in February 1978. Only 
dip-slip results from the 1978 survey 
have been reported to DWR, but the data 
does show movement continued. The 
44.6 mm (1.8 in) component of dip-slip 
reported in 1976, increased to 57.6 mm 
(2.3 in). 

An additional array of survey monuments 
was installed north of the U. S. Geolo- 
gical Survey array, by the Department 
of Water Resources, to measure changes 
in ground elevation across fault rupture 
discovered in October 1975. The DWR 
monuments were first releveled in 



October 1975, and last measured in 
October 1977. Maximum change in eleva- 
tion across the fault zone by 
October 1977, (See Figure 90, Chapter IV) 
was 50 mm (2 in) ,with the east side of 
the fault zone going up 10 mm (0.4 in), 
and the west side going down 40 mm 
(1.6 in). 

Total fault displacement is not known, 
because the increment of movement between 
the time the earthquake occurred and the 
time the survey monuments were installed 
is not accounted for. However, compar- 
ing what was seen in the field with the 
survey data, gives the impression that 1 
the survey data probably accounts for 1 
most of the fault displacement. 

The fault displacement cannot be evalu- 
ated precisely and, in fact, appears 
to vary from place to place. Maximum 
width of cracking measured over 40 mm 
(1.6 in), but this may be in part due to 
crumbling away of the crack edges. Most 
of the cracking was 25 mm (1 in), or less, 
in width. The collective data suggests 
vertical movements of about 50 mm (2 in) 
and horizontal extension of about 25 mm 
(1 in) . It is expected another earth- 
quake of similar magnitude would pro- 
duce similar displacements . 



GROUND ELEVATION CHANGES 

Immediately after the August 1, 1975, 
earthquake. Department of Water Resources 
survey teams conducted first-order lev- 
eling traverses on the pre-existing 
Oroville project survey network. Points 
and lines were added to the network to 
provide additional elevations where 
needed. These lines were releveled at 
approximate six-month intervals through 
the next 26 months. For calculating 
elevation changes, a benchmark near the 
eastern edge of the project area was 
considered to be stable and all other 
points were adjusted accordingly. 

Contours of elevation change for vari- 
ous time intervals following the earth- 
quake were calculated and plotted by 



26 



computer. These plots were overlaid on 
geologic maps and analyzed for relation- 
ships to geologic structure and litho- 
logic distribution. The plot for the 
period of August 1975, to October 1976, 
is shown in Figure 14. Since the inter- 
vals represented are all after August 1, 
1975, the changes plotted do not include 
any changes which may have occurred con- 
currently with the main earthquake or 
early aftershocks. 

In general, the settlement contour maps 
show an abrupt lowering of the ground 
surface west of a north-trending zone 
that coincides approximately with the 
Cleveland Hill Fault. Maximum total dif- 
ferential elevation change across this 
zone, for the period of August 1975, 
to October 1976, was 42 mm (1.7 in). 
The amount of settlement decreases to 
the north; survey control is lacking to 
determine how far this trend continues 
to the south. Elevation changes in the 
northern portion of the area are of les- 
ser magnitude, generally less than 
25 mm (1 in) , and are not readily rela- 
ted to local geologic structure. 

Repeated surveys show elevation changes, 
including increases, decreases and rever- 
sals in direction, continuing throughout 
the area — but at a decreasing rate. 
The continuing elevation changes suggest 
that readjustment along the fault zone 
was occurring 26 months after the main 
shock. 

Because initial studies indicated the 
northward progression of ground cracking 
was trending toward Bidwell Canyon 
Saddle Dam, periodic surveys were made 
along the crest of the dam to monitor 
changes in elevation. Results of these 
surveys are shown in Figure 15. The 
level line along the dam crest was sur- 
veyed four times since the earthquake. 
These surveys show the dam went down 
about 15 mm (0.6 in) more on the west 
than on the east. Most of the eleva- 
tion change occurred at the west end in 
the vicinity of shearing in the founda- 
tion. The marked lowering of elevations 
in the vicinity of Monuments BB-2 and 



BB-8 are mostly due to embankment set- 
tlement at places of maximum embankment 
height. 

Although not conclusive, these data 
strongly suggest that the down-to-the- 
west pattern of crustal movements seen 
at Cleveland Hill continues, albeit 
somewhat diminished, northward to Bidwell 
Canyon Saddle Dam. 

Four shears exposed in the foundation 
excavation for the saddle dam are shown 
on Figure 15. Most of the elevation 
change appears to be in the vicinity of 
the largest shear zone near the west end 
of the dam. This suggests the shearing 
at the west end of the dam is a zone 
along which the elevation changes are 
occurring. If so, the zone along which 
the Cleveland Hill Fault displacements 
occurred to the south, may pass through 
the west end of the Bidwell Canyon 
Saddle Dam. 

AREA LINEAlffiNTS 

A photo lineament is defined as, "Any 
line, on an aerial photograph, that is 
structurally controlled, including any 
alignment of separate photographic images 
such as stream beds, trees, or bushes 
that are so controlled. The term is 
widely applied to lines representing 
beds, lithologic horizons, mineral band- 
ings, veins, faults, joints, unconformi- 
ties, and rock boundaries" (Allum, 1966, 
p. 31). Because lineaments are often 
fault related, they are useful indicators 
of possible faults. Consequently, one 
of the early tasks for this project was 
to plot regional lineaments on topogra- 
phic maps and then inspect them in the 
field. 

The study of area lineaments began in 
May 1976, with a field reconnaissance of 
major photo lineaments north of Oroville 
Reservoir. A more comprehensive study, 
including detailed mapping and field 
observations, was done in September and 
October 1977. A similar study was con- 
ducted in 1976-77 by the U. S. Army 
Corps of Engineers as a part of their 



27 




3 KILOMETRES 



CONTOUR INTERVAL • 9 MILLIMETRES 



Figure \k. Changes in ground elevations around Lake Oroville, August 1975, to 
October 1976 



28 




29 



Marysville Lake project. Their work 
covered an area from Oroville to 9.6 km 
(6 mi) south of the Yuba River (U. S. 
Army Corps of Engineers, 1977). 
Woodward-Clyde Consultants also per- 
formed a major study of lineaments in 
the foothill belt as part of a seismic 
investigation for the Auburn Dam 
project. 

The purpose of the lineament study was 
to determine which lineaments are faults. 
Methods of investigation included aerial 
reconnaissance, photo interpretation, 
field inspection and mapping; topography, 
springs, continuity of rock units, fault 
gouge or any other features which would 
suggest or refute faulting were investi- 
gated. Previous geologic maps were 
field checked and incorporated in the 
study. 

Lineaments are categorized according to 
degree of certainty for a fault origin. 
Three categories are identified: 
(1) lineaments where field evidence con- 
tradicts a fault control and suggests 
origins are due to other causes (indi- 
cated as photo lineaments, dotted lines 
in Figure 16); (2) lineaments with no 
direct evidence of faulting, but are 
still thought to be faults (probable 
fault, dashed lines in Figure 16); and 
(3) lineaments which are faults (solid 
lines in Figure 16) . 

Three major photo lineaments are within 
the study area. These include the 
Paynes Peak, Swain Ravine and Prairie 
Creek Lineaments. Surface and subsur- 
face data have been analyzed by 
California Department of Water Resources, 
Woodward-Clyde Consultants, Pacific Gas 
and Electric Company, U. S. Army Corps 
of Engineers, U. S. Bureau of Reclamation 
and U. S. Geological Survey in attempts 
to evaluate the past history of activity 
along these fault zones. 

The major fault lineaments are in places 
defined by broad discontinuous linear 
valleys and aligned sections of sheared 
rock. The lineaments are not continuous 
but have gaps where nothing is evident 



either on imagery or on the ground. The 
Cleveland Hill faulting started in one 
of these lineament gaps. The Swain 
Ravine Lineament merges with the Prairie 
Creek Lineament approximately 10 km 
(6 mi) north of the Bear River. In turn, 
the Prairie Creek Lineament extends 
southward and is truncated west of 
Auburn by the Rocklin pluton. 

GEOLOGIC SETTING 

Geographic Location 

The Oroville study area is within the 
Sierra Nevada of California (Figure 17). 
The foothills of the Sierra Nevada sepa- 
rate the Sierran uplands from the rela- 
tively flat Sacramento and San Joaquin 
Valleys (the Great Valley) . The study 
area comprises the northernmost part of 
these foothills. 

Immediately north of, and included in 
the reconnaissance study, is the southern- 
most portion of the Cascade Range. 
Across the northern end of the Great 
Valley, northwest of the study area, are 
the Klamath Mountains which share many 
structural and lithologic characteris- 
tics with the Sierran foothills. 

Geologic Framework 

The Sierra Nevada province is an up- 
lifted block of Mesozoic plutonic and 
metamorphic rock bounded by normal 
faults on the east and tilted to the 
west. The eastern side is very steep 
with pronounced fault topography, while 
the western side is of gentler relief. 

The regional geologic fabric of Northern 
California is oriented north-northwest. 
Generally following this fabric is the 
Sierra Nevada crest, boundary faults, 
rock fabric, and foothills faults. The 
axis of the Great Valley, the crest of 
the Coast Ranges, and major faults in 
the Coast Ranges (e.g., San Andreas) 
also conform to this regional trend. 

The foothills of the Sierra Nevada are 
underlain by Paleozoic to Mesozoic 



30 




SCALE OF MILES 



SCALE OF KILOMETRES 



YubO City^ 

Fiqure 16. Lineaments and faults in the northwestern Sierran foothills 



31 



; 



I 



\ KLAMATH \ \ MODOC '^ ^ 

> MOUNTAINS , \ PLATEAU i ' 

; ' ^ 

\ / 

\ ^' ( 

/ CASCADE' \ 
V RANGE 



CHEROKEE '' 


-^ BERRY 
CREEK 


OROVILLE 


OROVILLE 
DAM 


PALERMO 




BANGOR 




39° 22 30 

Quadrangles comprising the 
geologic nnap. 



DESERT 



^ .o-\ 



O, 



-C< 



<>\ 



l'/^/ ^ o.^ 



^ej-^!. 



S ^'9 






Figure 17. Natural geologic provinces of California with field area location 



32 



metamorphosed sedimentary and volcanic 
rocks, and plutons which are similar to, 
but generally smaller than, those form- 
ing the Sierra Nevada. The metamorphic 
fabric and trend of faults in the foot- 
hills are most commonly concordant with 
regional northwest trends; local varia- 
tions occur about the intrusive bodies. 

Overlying the metamorphic foothills 
bedrock is a thick sequence of unmeta- 
morphosed upper Mesozoic and Cenozoic 
sedimentary and volcanic rocks (Super- 
jacent Series rocks) and alluvium. 
These rocks are in most places undeform- 
ed and dip gently west. 

The northern foothills have recently 
been interpreted as remnants of Mesozoic 
subduction complexes consisting of me- 
lange, arc rocks and ophiolite (Moores, 
1972; Cady, 1975; Schweickert and Cowan, 
1975; Buer, 1977, 1978; this study); 
this model also has been applied to the 
Klamath Mountains and Coast Ranges 
(Davis, 1969; Hamilton, 1969). The sub- 
duction zone is thought to have migrated 
westward during Late Jurassic time to 
form melanges of the Coast Ranges 
(Hamilton and Myers, 1966; Hamilton, 
1969; Burchfiel and Davis, 1972, 1975). 

DESCRIPTIVE GEOLOGY 

Bedrock Series Rocks 

Global applications of sea-floor spread- 
ing, as proposed by Hess (1959, 1962), 
have conceived new explanations and 
evolutionary interpretations for sea 
floor and continental rocks. Bedrock 
suites in this investigation are mapped 
and described assuming plate tectonic 
modes of origin and using evolutionary 
names, rather than formational names. 
Previous formational names are referen- 
ced in describing lithologic groupings. 
Geologic mapping of individual rock 
suites established the present structu- 
ral configuration and provides explana- 
tions for plate tectonic development of 
the region; this includes the origin of 
lithologic suites and time-separated 



episodes of faulting. Lithologic suites 
in the area include: 

(1) Melange - This suite consists of 
chaotically mixed metasedimentary and 
metavolcanic rocks which include serpen- 
tine and exotic blocks of marble. Me- 
lange was formed at a convergent bound- 
ary that existed between late Mesozoic 
California (American plate) and Pacific 
sea floor (ancestral Farallon plate that 
is now subducted beneath American plate) . 
These rocks formed by subduction under- 
thrusting as an accretionary prism in 
the Benioff zone. 

(2) Arc rocks - Rocks in this suite 
are volcanic and volcaniclastic deriva- 
tives that formed as an island arc com- 
plex in the ocean adjacent to Mesozoic 
California. Island arcs, common in many 
of todays oceans, develop relatively 
close to convergent plate boundaries 
where subducted lithosphere melts as it 
descends into the earth. The melted 
rock rises because of a lighter specific 
gravity, and volcanic mountains form if 
magmas reach the surface. 

(3) Ophiolite - This lithologic suite 
is named Smartville ophiolite in the 
study area and includes metamorphosed 
mafic rocks (amphibolite) that have ori- 
gins peculiar to the sea floor. Ophioli- 
tic suites form at oceanic spreading 
centers (rifts) and require that several 
mafic and ultramafic rock types be pre- 
sent for a complete ophiolite sequence. 
These lithologies form layers of varying 
thickness in undisturbed ophiolite and 
include: (1) an overlying mantle (layer 
1) of marine sediment and chert; (2) an 
intrusive-extrusive complex (layer 2) of 
submarine-extruded pillow lava that was 
fed by, and grades downward into intru- 
sive sheeted dikes; (3) a quasi- 
stratiform intrusive complex (layer 3) 

of gabbro, cumulate gabbro and dunite; 
and (4) a tectonite basement (layer 4) 
consisting of harzburgite and minor 
dunite. Ophiolite exposed within con- 
tinental margins necessitates special 
processes (obduction) for emplacement 
from oceanic source areas. 



33 



Melange 

Previous Investigations and Age ; 
Melange rocks in the western Sierra 
Nevada were first mapped and named 
"Calaveras formation" by Turner (1893, 
p. 309) for prominent exposures in 
Calaveras County. The name "Calaveras 
formation" subsequently became a 
"...catchall for all Paleozoic rocks in 
the Sierra Nevada and hence has no 
stratigraphic significance. ..." 
(Taliaferro, 1943, p. 280); this excludes 
Silurian and Upper Carboniferous rocks 
in the Taylorsville region. Exposures 
of Calaveras rock have been studied in 
many investigations (Lindgren, 1900; 
Clark, 1964, 1976; Creely, 1965; 
Hietanen, 1973a, 1976, 1977), but no 
regional correlation of units has been 
achieved. 

The first detailed study of melange in 
the immediate area was by Creely (1965). 
He subdivided and described the Pentz 
Sandstone and Hodapp Members from the 
Calaveras Formation; however, most occur- 
rences of the formation were mapped as 
"undifferentiated Calaveras". 

Melange terrane northeast of the study 
area is subdivided by Hietanen (1973a) 
into the Calaveras, Horseshoe Bend, 
Duffey Dome and Franklin Canyon Forma- 
tions. In later works, Hietanen (1976, 
1977) further mapped, subdivided and 
described the Horseshoe Bend Formation 
(Berry Creek quadrangle) and noted that 
these rocks are physically continuous 
with Creely 's Calaveras Formation to the 
east. In Hietanen' s works no mention of 
melange was used to describe these com- 
plex rock suites. 

The first published description of 
melange in the Sierran foothills was by 
Moores (1972), who suggested that "rem- 
nants of subduction zones" may be present 
in foothill areas. Subsequent studies 
(Bateman and Clark, 1974; Duf field and 
Sharp, 1975; Schweickert and Wright, 
1975; Schweickert and Cowan, 1975) have 
noted the widespread presence of melange 
in the Sierran foothills. Most recent 



studies, recognizing the nature of these 
suites, have dropped the name "Calaveras 
Formation" and adopted the term "melange" 
to describe the rocks. 

The age of melange in the foothill belt 
is misinterpreted as being late 
Paleozoic by most earlier researchers 
(Turner, 1893, 1894, 1896; Lindgren, 
1900; Taliaferro, 1943, 1951; Clark, 
1964, 1976). Early ages were establish- 
ed by using Tethyan fossils (Douglas, 
1967) in limestone and marble bodies 
that crop out in melange matrix. These 
carbonate bodies are interpreted to be 
exotic blocks within melange. 

Fossils collected from exotic blocks in 
melange do not represent the age of 
melange formation, but rather the age of 
the exotic block. In the Klamath Mount- 
ains late Paleozoic fossils are in bodies 
of limestone and marble (Irwin, 1972; 
Irwin and Galanis, 1976); cherts in 
melange-type rocks that include these 
carbonate bodies yield Late Jurassic 
radiolarians indicating the rock suites 
are much younger than fossils in marble 
and limestone bodies suggest (Irwin and 
others, 1977) . 

An Upper Jurassic fossil (a pelecypod, 
Buchia Concentrica as identified by, and 
in possession of Ralph Imlay, U. S. Geo- 
logical Survey) was discovered in meta- 
sedimentary melange near Pentz, Califor- 
nia (Bob Treet, person, commun., 1978); 
the fossil specifically dates from middle 
Oxfordian to upper Kimmeridgian time. 
The Buchia fossil locality is in the 
southeast quarter of section 13 (T21N, 
R3E) on the Cherokee qixadrangle. It 
substantiates that melange (Calaveras 
Formation) in the northwestern Sierran 
foothills is much younger than previously 
suggested. The Upper Jurassic age of me- 
lange is the same as arc and ophiolitic 
rocks located to the south and indicates 
contemporaneous origins. 

Contact Relationships ; The southern 
margin of melange in the project area 



34 



is interpreted to be an obduction bound- 
ary with arc rocks in the Cherokee quad- 
rangle. The contact with arc and ophio- 
lite in the Berry Creek quadrangle is 
complex and not well exposed. The base 
of melange is not exposed in the study 
area. 

Melange in the study area is overthrust 
by obducted arc rocks. Allochthonous 
rocks underlie large portions of the 
northern study area and form more sur- 
face exposure than autochthonous me- 
lange. Uncertainty exists as to how 
much melange is overridden. 

Lithologic Description ; The melange 
complex in the study area includes meta- 
sedimentary rocks (argillites, schists, 
phyllites, meta-tuf faceous beds, relict 
pebble conglomerates, exotic marble 
blocks and chert), metavolcanic rocks 
(relict basalts, diabases and andesites) 
and serpentine. These rocks, mixed by 
tectonic and olistostromal processes, 
are isoclinally deformed into tight 



folds that dip vertically or steeply to 
the east (Figure 18) . Structural dis- 
continuity and intercalation of melange 
lithologies suggests the sequence origi- 
nated as an accretionary prism in a sub- 
duction zone; similar rocks and structu- 
ral relationships are described at many 
active and ancient subduction zones 
(Hsu, 1971; Blake and Jones, 1974; 
Gansser, 1974; Karig, 1974; Scholl and 
Marlow, 1974; Karig and Sharman, 1975; 
Dickinson, 1975). Accretion, using sev- 
eral different models (Burk, 1965; Dewey 
and Bird, 1970; Gilluly, 1972; von Huene, 
1972; Moore, 1973), is postulated as the 
process by which lower plate rocks be- 
come transferred (accreted) to the upper 
plate during subduction. 

Relict basalt, diabase and andesite 
flows (?) and sills (?) are the most 
abundant metavolcanic rocks incorporated 
in melange. They are locally intercal- 
ated within metasedimentary sequences. 
Contacts between metavolcanic rocks and 
metasedimentary rocks are poorly exposed. 




Figure 18. Small-scale parasitic isocl inal Figure 19. Relict bedding (parallel to 
fold within melange metasedimentary rock, pencil) cross cut by steeply east-dipping 
1.5 km (1 ml) southwest of the West foliation in melange metasedimentary rock, 
Branch Bridge. 1.5 km (1 mi) southwest of the West 

Branch Bridge. 



35 



Meta-basalts and -diabases include ura- 
litized amphibole and/or pyroxene, 
sodic plagioclase and secondary epidote. 
Accessory minerals are ilmenite or hema- 
tite, quartz and secondary chlorite. 

Argillites and metagraywack.es are dark 
when fresh and retain some of the ori- 
ginal sedimentary features; shearing 
has locally transposed depositional 
structures into the plane of foliation 
(Figure 19). Sodic plagioclase, quartz, 
epidote, muscovite, chlorite and traces 
of metallic minerals comprise argil- 
laceous rocks. 

Pebble metaconglomerates of volcani- 
clastic origin are locally common within 
melange matrix (Figure 20). These are 
in places exposed against deep-water 
slates. Shearing has stretched clasts, 
however, the volcanic origin of both 
clasts and matrix remains visible. 

Schistose and phyllitic rocks vary from 
dark- to light-green where fresh and 




Figure 20. Sheared volcaniclastic meta- 
conglomerate intercalated with black 
slates (not shown) in melange meta- 
sedlmentary rock. Location is 1.5 km 
(Imi) southwest of the West Branch 
Bridge. 



are various shades of buff if weathered. 
Syntectonic shearing has destroyed ori- 
ginal textures. 

Chert has limited exposure in melange 
terrane. Localized occurrences expose 
thin- to medium-bedded light-gray to 
white chert. Thin sections indicate 
chert is composed of 95 to 100 percent 
recrystallized quartz. Most chert in 
melange was clastically derived 
(Hietanen, 1977, p. 7). 

Marble in melange (Figure 21) is 
white to bluish-gray. Most occurrences 
of marble are exotic and do not repre- d 
sent original in situ deposition. Sam- j 
pies collected along Nelson Bar Road | 
(Cherokee quadrangle) exhibit marble , 
and phyllite incorporated to form a rock 
with foliation discordantly cross-cutting 
the contact between the two lithologies 
(Figure 22). This indicates marble was 
incorporated into fine-grained sediments 
prior to metamorphism. An absence of 
shear at the marble-phyllite contact 



MARBLE 




Figure 21. View north from the West 
Branch Bridge of exotic marble block 
in melange. 



36 




igure 22. Sample of ol i stostromal marble-phy 
Bar Road just east of Oroville Reservoir (Ch 
and phyllite is nearly perpendicular to the 
types. 

suggests gravity was the emplacement 
mechanism; therefore,' this deposit is 
an olistostrome. Olistostromal deposits 
and tectonic knockers are common in many 
melange deposits (Hsu, 1965, 1968; 
Raymond, 1977). 

The distribution of fossilif erous 
marble in the field area is restricted 
to a linear belt of exposures in West 
Branch Canyon. Marble exposures in other 
areas are scarce, non-fossiliferous and 
concordant with local bedding in meta- 
sedimentary melange. 

Light- to dark-green serpentine, as 
highly-sheared to unsheared rocks, forms 
elongate discontinous exposures that are 
concordant with the local foliation. 



1 1 i te collected in melange along Nelson 
erokee quadrangle). Foliation in marble 
unsheared contact between the two rock 

Contacts between serpentine and adja- 
cent rocks are poorly exposed. Plate 
tectonic models associate serpentine 
with subduction (Benioff) zones at con- 
vergent boundaries (Hamilton, 1969; 
Bailey and others, 1970; Bateman and 
Clark, 1974; Lapham and McKague, 1964; 
Coleman, 1977). Lockwood (1971, 1972) 
has suggested that serpentine can be 
clastic or deposited as olistostromes. 
However, the sheared and truncated na- 
ture of serpentine in the study area 
suggests it was tectonically emplaced 
as opposed to a depositional origin. 
Elongate serpentine exposures in the 
area are interpreted as locations of 
ancient shear zones within the subduc- 
tion complex. 



37 



Metavolcanic arc rocks are exposed in 
melange in the Cherokee and Berry Creek 
quadrangles. Contacts are not well ex- 
posed but most appear to be relatively 
flat-lying; these contacts conform to a 
model of arc rock overthrust on melange. 
It is possible that arc rocks were tec- 
tonically mixed into melange during sub- 
duction, however, it is also possible 
that the thrust plane has been folded 
and these exposures of arc rock are 
klippen. 



Arc Rocks 

Previous Investigations and Age ; Base- 
ment Series greenstones in the Sierran 
foothill belt were first mapped and des- 
cribed by Becker and others (1898), and 
Lindgien and Turner (1895). Greenstone 
descriptions from these studies are re- 
fined in several later works (Creely, 
1955, 1965; Bateman and Clark, 1974; 
Clark, 1976; Hietanen, 1977). Moores 
(1972) suggests part of the greenstone 
complex is an ancestral island arc. 
Subsequent investigations (Cady, 1975; 
Moores, 1975; Schweickert and Cowan, 
1975; this study) have subdivided Sierran 
foothill greenstones into members whose 
origins are explained using a plate tec- 
tonic framework. 

The most detailed academic study involv- 
ing the westernmost suite of volcanic 
and volcaniclastic rocks in the green- 
stone belt is by Creely (1955, 1965). 
He applied the name "Oregon City Forma- 
tion" to describe this metavolcanic se- 
quence that is now recognized as an arc 
complex. These rocks are dated Late 
Jurassic (Oxfordian to Kimmeridgian) by 
an ammonite identified as Perisphinctes 
by Professor S. W. Muller, Stanford 
University (Creely, 1965). 

Contact Relationships : A reverse fault 
forms the eastern contact of arc rocks 
with Smartville ophiolite in the Oroville 
area. Western margins of the arc com- 
plex are unconformably overlain by late 
Cenozoic Superjacent Series deposits 
and alluvium. Arc rocks are not exposed 



in contact with other Basement Series 
rocks west of the foothills. 

The arc complex ends abruptly in the 
southeastern Cherokee quadrangle. Field 
evidence at this location indicates arc 
rocks are thrust over melange by a late 
Mesozoic thrust fault. This fault is 
nearly flat-lying and probably represents 
an obduction suture. 

Arc lithologies in the Oroville area 
are physically continuous to the south 
with Browns Valley Ridge volcanic rocks 
located in the foothills east of 
Marysville. Further south, they are 
time and structurally correlative with 
the Copper Hill and Gopher Ridge volcanic 
sequences located north of the Mokelumne 
River (Duffield and Sharp, 1975). 

The base of the arc complex is not ex- 
posed in the study area. The thickest I 
sequence is exposed in Morris Ravine on 
the west limb of the Monte de Oro syn- 
cline and includes approximately 400 m 
(1315 ft) of section. 

Lithologic Description ; Exposures of 
fresh arc rock are dark- to light-green 
and extremely well indurated. Foliation 
and relict flow structure are poorly 
developed. 

Foliation is generally accentuated by 
weathering. It is not known if folia- 
tion and relict flow structure are con- 
cordant; metamorphism has transposed 
original structures in local metasedi- 
mentary rocks (Figure 23), and this 
characteristic probably exemplifies 
foliation in the arc sequence. 

The arc complex is formed by several 
fine- to coarse-grained lithologies that 
are intermediate to basic in composition. 
Arc lithologies include andesitic tuff- 
breccia, lapilli tuff agglomerate, and 
epiclastic derivatives of these rocks; 
tuff-breccia and monolithologic agglome- 
rates (Figure 24) are by far the most 
common. Relict flows, pillows (Figure 
25), and sills (?) of meta-andesite 
and -basalt are present but not common. 



38 








Figure 23. Arc metased imentary rock 
displaying relict bedding (dipping 
into photograph) that is nearly per- 
pendicular to steeply east-dipping 
foliation (subparallel to pencil). 



Figure 2A. Arc complex metavolcanic 
tuff breccia at The High Rocks, 
approximately 1 km (0.6 mi) south- 
east of Oregon City. Note pocket 
knife for scale. 



Figure 25 
4 km (2 
Note pen 




Pii-UDW 

I 



X' 



Arc complex relict pillow and flow lavas cut by fault, 
5 mi) northeast of Oregon City along the North Fork of 
cil (center-left) for scale. 




approximately 
Lake Orovi 1 le. 



39 



The fine-grained, metamorphosed nature 
of arc rocks makes field identification 
difficult; however, most of the present 
complex was probably water-lain. 

Tuffaceous rocks include hornblende 
and/or augite, sodic plagioclase (saus- 
suritized) and lithic fragments; these 
are set in a finer-grained tuff matrix 
which comprises the bulk of the forma- 
tion. Vesiculated lithic fragments are 
commonly angular and of one rock type, 
suggesting derivation from local erup- 
tive vents. Subrounded fragments of 
different rock types also are common 
and represent clastic water-lain se- 
quences (Figure 26). Vesicles in clasts 
vary from 2 to 8 mm (0.08 to 0.3 in) in 
diameter and are filled with secondary 
quartz. Common minerals include epidote, 
clinozoisite, chlorite, and pyrite. 
Much of the groundmass observed in 
thin section is cryptocrystalline and 
comprises unidentifiable alteration pro- 



ducts derived from secondary and meta- 
morphic mineral reactions. 

Monte de Oro Formation 

Previous Investigations and Age : The 
Monte de Oro Formation was named and 
first described by Turner (1896) . His 
work noted fossil debris throughout the 
formation which probably inspired some 
of the later studies. Subsequent inves- 
tigations (Fontaine, 1900; Knowlton, 
1910; Diller, 1908; Taliaferro, 1942; 
Creely, 1965) describe these rocks in 
great detail. Fossil flora, and to a 
lesser extent fauna, provided the earli- 
est evidence that rocks in the Oroville 
area have Jurassic ages. 

Contact Relationships ; Monte de Oro 
Formation in the study area represents 
the tightly folded axis of a syncline 
overturned to the west. Approximately 
375 m (1,230 ft) of Monte de Oro Forma- 
tion is stratigraphically exposed. 




Figure 26. Relict bedding and cross-bedding in arc tuff breccia and tuffaceous 
metasedimentary rock, 5 km (3 mi) southeast of Palermo. Darker area of outcrop 
(right-center) was wetted to accentuate structure. 



40 



The western contact of Monte de Ore 
Formation is depositionally conformable 
on arc rocks. Arc complex flows are 
intercalated with bedding in lower por- 
tions of Monte de Oro Formation and 
form a gradational contact between the 
two sequences; for this reason, Monte de 
Oro rocks are interpreted to be a sedi- 
mentary facies of the arc complex. The 
eastern margin of the formation is trun- 
cated against arc rocks by an east- 
dipping reverse fault. 

Lithologic Description ; Monte de Oro 
Formation in the Oroville area represents 
the only named exposures of these rocks 
in the Sierran foothills; however, simi- 
lar metasedimentary rocks are exposed 
in the Bangor quadrangle to the south. 
These rocks are in structural alignment 
with Monte de Oro rocks near Oroville 
and may have been deposited in a common 
Upper Jurassic environment. 

The Monte de Oro Formation is predomi- 
nantly slightly-sheared, well-indurated, 
dark sandstone, siltstone, conglomerate 
and poorly-developed slate. Argillaceous 
siltstone and sandstone (metagrajwacke) 
with poorly-developed interbedded slate 
constitute the bulk of the formation. 

Exposures of metasiltstone are dark- to 
olive-gray where fresh and weather to 
light-olive-buff. Metasiltstone, com- 
monly containing relict sandy and clayey 
sections, is moderately- to well-bedded 
and laterally continuous. Plant debris 
is locally abundant on metasiltstone 
bedding-plane cleavages. 

Monte de Oro metasandstone is poorly- 
bedded and includes graywacke and arkose. 
Relict sandstone beds have lenticular 
shapes and are laterally discontinuous. 
The bulk of this material occurs in 
lower portions of the formation and is 
believed to have been reworked from 
underlying arc rocks. 

Graywacke consists of subrounded, 
medium- to coarse-grained, poorly-sorted, 
feldspar, rock fragments and quartz or 
detrital chert. These constituents are 



cemented by clay and silica (?) in an 
argillaceous matrix. 

Arkose is fine-grained and consists 
predominantly of subrounded, moderately- 
to well-sorted feldspar. Clastic grains 
are set in a green chlorite (?) relict- 
silt matrix. 

Monte de Oro metaconglomerate is formed 
of subangular to rounded, pebble- to 
cobble-sized clasts set in an argil- 
laceous relict-sandstone matrix. These 
beds are lenticular and most abundant in 
relict sandstone sequences. 

Predominant clast types in metacon- 
glomerate are poorly sorted and include 
plagioclase- and quartz-rich porphyritic 
dacite (?), dark chert and black slate. 
Dark, fine-grained, indeterminate volcanic 
clasts are common but less abundant. 

It is significant that many of the 
clasts in Monte de Oro metaconglomerate 
are not derived from the underlying arc 
complex. Exotic volcanic clasts, as 
well as accompanying chert and slate, 
were probably derived from pre-arc ter- 
rane; these sources may include melange. 

Smartville Ophiolite 

Previous Investigations and Age : 
Studies by Lindgren and Turner (1895) 
and Becker and others (1898) provide 
early maps and descriptions of meta- 
volcanic greenstones in the northwestern 
Sierran foothills. The first detailed 
investigations of this greenstone suite 
were by Hietanen (1951), Compton (1955) 
and Creely (1955, 1965). More recent 
investigations (Moores, 1972, 1975; 
Cady, 1975; Schweickert and Cowan, 1975; 
Bond and others, 1977; Buer, 1977; 
Day, 1977; Hietanen, 1977) describe this 
greenstone sequence as dismembered 
ophiolite. 

Cady (1975) proposed the name Smartville 
ophiolite in his study. This name is 
adopted in our investigation. 

The age of Smartville ophiolite was 



Al 



originally suggested to be late Paleozoic 
by Creely (1965) from a comparison with 
Oregon City Formation (arc complex) that 
was dated by fossils. In later works, 
Cady (1975) and Hietanen (1977) consider 
these rocks to be Jurassic in age. 

The Smartville complex is interpreted 
to have formed by back-arc spreading 
(Schweickert and Cowan, 1975; Eldridge 
Moores, person, commun., 1977); similar 
spreading basins are active today in 
many areas of the Pacific Ocean 
(Hamilton, 1969; Karig, 1970, 1971a, 
1971b, 1972; Moberly, 1972; Churkin, 
1975; Karig and Sharman, 1975). Behind- 
the-arc spreading is suggested to have 
occurred in Callovian to Oxfordian time 
by Schweickert and Cowan (1975); however, 
their model has ophiolite originating 
prior to eruption of the Oxfordian age 
(Creely, 1965) Oregon City volcanic se- 
quence. Smartville ophiolite is now 
interpreted to have formed in late 
Oxfordian to early Kimmeridgian time 
which is younger, but in part coeval with 
development of the arc complex. 

A fault separates arc rocks and ophi- 
olite in the study area, therefore, di- 
rect evidence is lacking to substantiate 
whether arc rocks are intruded by source 
magmas from a spreading interarc basin, 
or if the arc complex is built upon 
ophiolite. Field evidence is inconclus- 
ive and consists of: (1) Arc litholo- 
gies on Bloomer Hill, in the Berry Creek 
quadrangle, overlie ophiolite; however, 
poor contact exposures prevent a deter- 
mination of whether the contact is depo- 
sitional or fault controlled. (2) In 
the foothills east of Marysville, meta- 
basaltic dikes similar to those in 
ophiolite appear to intrude arc litholo- 
gies of the Browns Valley Ridge volcanic 
sequence (Costas Xenophontos, person, 
commun., 1978). (3) A few hundred metres 
west of the California Highway 20 bridge 
over the Yuba River, Smartville pillow 
basalt is conformably overlain by argil- 
lite and arc-derived (Koll Buer, person, 
commun., 1978) tuff and pyroxene ande- 
site tuff breccia. 



Contact Relationships ; The western 
margin of Smartville ophiolite in the 
study area is a near-vertical fault. 
The fault is not regional in extent; arc 
and ophiolite sequences are conformable 
along the Yuba River south of the area. 

Sierran plutons truncate Smartville 
ophiolite on the east. Intrusive rocks 
entered ophiolite in directions subparal- 
lel to regional foliation. 

The northern margin of the Smartville 
belt is poorly exposed; abundant meta- 
volcanic rock in melange further inhibits 
locating and interpreting the nature of 
the ophiolite-melange contact. The con- 
tact, although not mapped, is interpre- 
ted to be an obduction boundary. 

Lithologic Description ; Smartville 
terrane is a dismembered complex and 
does not contain all of the rock types 
and structural levels characteristic of 
ophiolite sequences as described by sev- 
eral researchers (Moores and Vine, 1971; 
Moores and Jackson, 1974; Coleman and 
Irwin, 1974; Williams and Stevens, 1974; 
Coleman, 1977). Common lithologies and 
structural layers that characterize the 
Smartville complex (Figure 27) include; 
(1) metasedimentary rock of layer 1 
ophiolite; (2) layer 2 meta-basaltic 
and -diabasic pillows, pillow breccia, 
dikes and sills over a complex of meta- 
basaltic and -diabasic dikes and sheeted 
dikes with felsic and gabbroic screen 
rocks; and (3) upper layer 3 gabbroic 
intrusions. Layer 2 pillows, dikes and 
sheeted dikes, with or without screen 
rocks, are the most common ophiolite 
members in the area. Layer 3 gabbroic 
intrusions, common in many ophiolites 
(Cass and Smewing, 1973; Jackson and 
others, 1975; Tysdal and others, 1977), 
are scarce in Smartville terrane. 

Individual pillows (Figure 28) have sub- 
spheroidal to lobate shapes and are 
usually poorly preserved. Well- 
developed pillows of the Smartville com- 
plex are exposed south of the project 
area along the Yuba River (Figure 29). 



42 



OPHIOLITE STRATIGRAPHY 

STANDARD OCEANIC CRUST SMARTVILLE OPHIOLITE 



SEDIMENTS 



PILLOW BASALTS 



LAYER 2 i[,2 MASSIVE BASALT AND 
>° DIABASE SHEETED 
5J DIKES AND SILLS 



LAYER 3^0. 



CUMULATE GA8BR0 



CUMULATE PYROXENITE 



CUMULATE DUNITE 



LAYER 4 tz HARZBURGITE WITH 
§2 MINOR DUNITE 




?FLYSCH + TUFF t CHERT 

^PILLOWS + FLOWS + 
SILLS 1 BRECCIA 



SHEETED DIKES 



MAFIC DIKES 

FELSIC AND GABBROIC 
SCREENS 

DIKE-GABBRO TRANS- 
ITION SOME PLAGIO- 
GRANITE AND DIORITE 



Figure 27- igneous stroligropny of Stondord Oceanic Crust with member thicknesses (after Moores 
and Jockson, 1974) ond Smortville ophiolite. Note that sections shown are unmetamorphosed. 
In most ophiolites, pillow basalts ond the sheeted dike complex are metamorphosed to greenschist 
or omphibolite facies with olmost total serpentinization of the cumulate and tectonized 
ultromofic rocks. An extensive shear zone commonly seporotes the cumulate and tectonized 
ultramofic rocks 




Figure 28. Well -developed metavolcanic 
Smartville pillows 1.5 km (1 mi) south 
west of Bangor. 



Synkinematically sheared pillows com- 
monly yield phyllonitic rock (Figure 30) 
In shear zones all original rock struc- 
ture is transposed and forms a cata- 
clastic foliation. 

Metabasalt forms most pillows and is 
dark-green to gray-green when fresh. 
Pyroxene, albite, epidote, and pyrite 
are the only minerals identifiable in 
hand specimen. Relict vesicles, filled 
with secondary quartz and epidote, are 
in places abundant. Quartz and epidote 
also fill discontinous veinlets in these 
rocks and the cores of some pillows. 

Sparse hyaloclastite or aquagene tuff 
forms selvages aroimd individual pillows. 
Chert has been described as abundant in 
some pillow basalts (Bailey and others, 
1964) but is not common in Smartville 
rocks. 



43 



-,J. 



.^ 






:/ 



/ 



J. ^ 

Figur 
of 
ind 
and 




e 29. 
the Ca 
i vidua 
di ps 



Well developed metavolcanic Smartville pillows at the south abutment 
lifornia Highway 20 bridge crossing of the Yuba River. Tails on 
I pillows indicate the section is right-side-up (to top of photograph) 
steeply west. 

Pillowed basalt grades downward into 
meta-diabasic and -basaltic dikes and 
sheeted dikes (Figure 31) . Dikes com- 
prise the greatest volumne of Smartville 
ophiolite in the study area. Models 
describing ophiolite (Moores and Vine, 
1971; Moores and Jackson, 1974) identify 
lower layer 1 as a structural level where 
meta-basaltic and -diabasic dikes and/or 
sills intrude pillows. This relationship 
is rare in the project area. 

Gabbroic and felsic screen rocks, indi- 
cating a deeper level of the ophiolite 
complex (Moores and Vine, 1971; Moores 
and Jackson, 1974), are locally abundant 
(Figures 32 and 33). Felsic screen rocks 
include quartz diorite, granophyric kera- 
tophre and trondhjemite, and represent 
differentiates from late-stage crystal- 
lization of sub-akaline magmas (Coleman, 
1971, 1977). Hyaloclastite screens are 
enclosed in dikes along California 
Highway 162 just south of Canyon Creek 
Bridge. These screens are phyllitic. 




Figure 30. Sheared metavolcanic Smart- 
ville pillows in North Honcut Creek 
stream bed near bridge crossing of 
the Oro-Bangor Highway. Shearing 
renders the outcrop appearance of a 
phyllonite. Note pencil (center- 
right) for scale. 



44 




1 A >* i 

Figure 31. Steeply east-dipping metavol can ic Smartville sheeted dikes along Rocky 
Honcut Creek, approximately 1 km (0.6 mi) west of Oro-Bangor Highway bridge 
crossing. 




Figure 32. Grabbroic screen rock in Smartville metavolcanic sheeted dikes along 
Olive Highway just east of Quincy Place (Oroville quadrangle). 



45 




Figure 33. Granophyric screen rock i 
along Rocky Honcut Creek, approxima 
Highway bridge crossing, 
fine- to medium-grained and contrast 
greatly with the darker meta-diabasic 
dikes. Tuffaceous screens indicate dikes 
intruded to very shallow levels of the 
ophiolite, and probably fed pillows on 
the ancestral sea floor. 

Dikes and sheeted dikes vary from 30 to 
100 cm (12 to 39 in) in thickness and 
have continuous trends where exposed; 
discontinuous dikes do occur, but are 
rare. Foliation is concordant with con- 
tact margins. Average strikes are 
N5-25W and inclinations dip steeply east 
at angles greater than 65 degrees; west 
dips occur locally but are not consid- 
ered representative for the dike complex. 

Contacts between sheeted dikes and be- 
tween dikes and screens are sharp and in 
places have chilled margins. 

Major minerals in metavolcanic ophio- 
lite are granoblastic clinopyroxene, 
albite, chlorite, epidote, clinozoisite, 
actinolite, tremolite and opaques (py- 



n Smartville metavolcanic sheeted dikes 
tely 1 km (0.6 mi) west of Oro-Bangor 

rite and chalcopyrite) . Unidentifiable 
cryptocrystalline metamorphic and hydro- 
thermal alteration products form a 
groundmass for these minerals. 

Gabbroic Smartville ophiolite in the 
study area is exposed as local dikes, 
plugs and stocks; gabbro is regionally 
limited in the complex. A gabbroic 
stock, with extremely complex intrusive 
contacts, is exposed west and northwest 
of Stringtown Mountain, in Woodman 
Ravine. 

Gabbroic rock in Woodman Ravine is com- 
posed of coarse-grained plagioclase and 
cummulate pyroxene (uralitized) . Dikes 
with well-developed chill margins cut 
gabbro in this area. Dikes contain gab- 
broic xenoliths which become less abun- 
dant to the south. Relationships of 
these mafic rocks are further compli- 
cated by intrusion of the Swedes Flat 
pluton. 



46 



Another gabbro (norite) intrusion under- 
lies a small portion of the North Fork 
Feather River canyon west-southwest of 
Bloomer Hill. Lake Oroville inundates 
much of the gabbroic surface area. 

Intrusive Rocks 

Previous Investigations and Age : 
Several previous investigators have 
mapped Sierran plutons in the study area 
(Becker and others, 1898; Hietanen, 1951, 
1973b, 1976, 1977; Compton, 1955; 
Evernden and Kistler, 1970; Bateman and 
Clark, 1974; Clark, 1976). Plutonic 
terrane mapped for this project includes 
only western margins of these earlier 
regional studies. 

Absolute ages of intrusive rocks in the 
study area are established by potassium- 
argon dating. Analyzed samples yield 
discordant hornblende and biotite ages 
and indicate that the dated plutons 
have experienced post-intrusive reheat- 
ing with subsequent degassing of argon. 

The Bald Rock pluton yields discordant 
ages of 131 and 126 million years on 
hornblende and biotite respectively 
(Evernden and Kistler, 1970). Two dis- 
cordant age dates for the Merrimac plu- 
ton, using the same minerals respec- 
tively, are 129 and 131 million years 
(Gromme and others, 1967) and 132 and 
129 million years (Evernden and Kistler, 
1970). Dated locations of these plutons 
are not within project boundaries; how- 
ever, granitic rocks are physically con- 
tinuous from these locations into the 
study area. 

Ages for the Bald Rock and Merrimac 
plutons indicate that emplacements were 
during Jura-Cretaceous time and syn- 
chronous with late stages of the Yoseraite 
intrusive epoch (Figure 34) . Late 
Jurassic to Early Cretaceous ages for 
the Sierra Nevada intrusive complex are 
suggested by several earlier researchers 
(Knoph, 1918, 1929; Erwin, 1934; Mayo, 
1934, 1935) without the aid of radio- 
metric dating. 

Contact Relationships : The Bald Rock 



and Swedes Flat plutons intrude 
Smartville ophiolite along the east- 
central and southeast margin of the meta- 
morphic complex. In most places the 
plutons entered ophiolite subparallel to 
the pre-existing regional foliation. 
Xenoliths, from a few centimetres to 
several tens of metres in diameter, are 
in places locally abundant in plutonic 



Age 

(my.) 


Sys- 
tern 


Se- 

ries 


Intrusive epoch 


70- 
80- 
90- 
100- 
110- 
120- 
130- 
140- 
150- 
160- 
170- 
180- 
190- 
200- 
210- 
220- 
230- 


r> 
o 

LJ 
O 

< 

1- 

cc 
u 


o. 
a. 

3 




Cothedrol Ronge 




5 
o 


Huntington Lake 




<_) 

(/) 

V) 

< 
en 

3 


Q. 
Cl 

r) 

■o 

? 
-J 


Yosem ite 




Inyo Mountains 




o 

V) 

to 
< 

IT 

1- 


Q. 
Q. 

r> 

-iHi- 

O 

_l 


Lee Vining 





(after Horland am' others, 1964), Data 
modified after Evernden and Kistler (1970). 

Figure },h . Mesozoic time scale with 
corresponding intrusive epochs in the 
Sierra Nevada region 



47 



rock near intrusive contacts (Figure 35) ; 
apophyses are locally present near intru- 
sive margins. 

The Merrimac pluton intrudes melange 
along the northeastern margin of the 
area. This pluton intruded subparallel 
with the regional foliation in melange 
country rock. 



Thermal low-shear metamorphism forms 
aureoles in country rock surrounding 
Sierran plutons. Country rock around 
the Bald Rock and Swedes Flat plutons 
is thermally recrystallized by amphibo- 
lite facies metamorphism in a 1 to 3 km 
(0.6 to 1.8 mi) wide aureole (Compton, 
1955). A contact aureole around the 
Merrimac pluton, also of amphibolite 
facies metamorphism and up to 4 km 
(2.5 mi) wide (Hietanen, 1977), is dev- 
developed in melange. 

Lithologic Description ; Plutons in the 
area have textures, mineralogies and 
ages that are typical of the Sierran 



intrusive complex. Rock types forming 
plutons include tonalite, granodiorite 
and quartz monzonite. Trondhjemite com- 
monly forms the central portion of 
local plutons. Aplitic and pegmatitic 
dikes, representing late-stage intrusive 
rocks, are in places abundant near in- 
trusive margins. 

The Merrimac pluton is primarily 
medium- to coarse- grained granodiorite. 
Mineralogy of the pluton includes zoned 
and unzoned plagioclase (An„^ to An,_) , 
quartz, potassium feldspar and ferro- 
magnesian minerals (biotite and horn- 
blende) . Accessory trace minerals in- 
clude apatite, epidote, muscovite, 
sphene and zircon. 

The Bald Rock pluton, a well-foliated 
compos it ionally-zoned intrusion, is a 
mixture of medium- to coarse-grained 
granodiorite, tonalite and trondhjemite. 
Tonalite and granodiorite are concen- 
trated in outer margins of the pluton; 
trondhjemite forms the core of the 
complex. Stoping, assimilation and sub- 




Figure'35. Metavolcanic xenoliths within Swedes Flat plutonic rock in Woodman 
Ravine, 6 km (3-5 mi) east of Oroville Dam. 



48 




Figure 36. View west-northwest of Bald 
Berry Creek, Rocks exhibit surface ex 
of the Sierran batholithic complex. 

sequent contamination is responsible for 
compositional layering in the Bald Rock 
pluton (Compton, 1955) . Tonalite and 
granodiorite include quartz, plagioclase 
(An„e to An_„), microcline, hornblende, 
biotite, ana accessory metallic minerals. 
Common minerals in trondhjemite are 
plagioclase (An„ to An„„), quartz, potas- 
sium feldspar and muscovite; ferromag- 
nesian minerals are rare. 

Flow structure is well developed in the 
Bald Rocl<. pluton. Flow banding dips 
steeply eastward and is defined by a 
planar parallelism of biotite, horn- 
blende and, to a lesser extent, plagio- 
clase. It is most strongly observable 
near intrusive margins where mafic 
minerals are concentrated. Flow layer- 
ing toward the center of the pluton 
maintains an easterly dip and is more 
concentric than along its margins 
(Compton, 1955). 



Rock (foreground) 6 km (3.8 mi) east of 
posure and exfoliation that is typical 

The Swedes Flat pluton is predominantly 
tonalite and granodiorite. Gabbro and 
diorite are present in subordinate 
amounts at the north and south ends of 
the pluton. Granophyric rock, as dikes 
and inclusion-charged masses, is abun- 
dant along the western margin of the 
pluton. Common minerals in Swedes Flat 
tonalite and granodiorite include saus- 
suritized plagioclase (An2Q to An^^), 
alkali feldspar, quartz, hornblende and 
biotite. Common accessory trace miner- 
als are epidote, apatite and, in 
places, sphene. 

Origin of Sierra Nevada Plutons ; Tona- 
lite and monzonite plutons in the study 
area (Figure 36) are similar in appear- 
ance, mineralogy and mode of origin to 
those comprising the Sierra Nevada 
batholith. Plate tectonic models devel- 
oped during the late 1960 's and early 
1970's provide new interpretations for 



49 



the large-scale origin of plutonic 
complexes. In both oceanic and 
continental-margin settings, voluminous 
calc-alkaline magmas are formed above 
Benioff zones 150 to 500 km (93 to 
124 mi) from the trench axis (Dickinson 
and Hatherton, 1967; Dickinson, 1968) 
and provide a tectonic model for Sierran 
plutonism. 

An east-dipping Benioff zone was adja- 
cent to the western coast of North 
America during much of Phanerozoic time 
(Hamilton, 1969; Burchfiel and Davis, 
1972, 1975); inclination of the subduc- 
tion zone is substantiated by potassium- 
silicon ratios in Mesozoic granitic 
rocks of California that increase east- 
ward (Moore, 1959; Bateman and others, 
1963; Dickinson, 1969) with a correspond- 
ing depth to the ancestral Benioff zone. 
Eastward subduction and partial melting 
of lithosphere at depth generated magmas 
(plutons) that rose to shallower struc- 
tural levels beneath Mesozoic California. 
The calc-alkaline plutons were tension- 
ally faulted, uplifted, and unroofed in 
Cenozoic time. These processes are cur- 
rently active and have erosionally 
removed more than 8 km (5 mi) of roof 
rock (Bateman and Wahrhaftig, 1966) to 
expose the plutons. 

Superjacent Series Rocks 

Chico Formation 

Previous Investigations and Age ; Sand- 
stone, shale and conglomerate of the 
Chico Formation were first described and 
named by Gabb (1869, p. 129). Diller 
and Stanton (1894) used the term "Chico 
group" in their study of these rocks; 
they considered all Cretaceous deposits 
in California part of the "Shasta-Chico 
series". Stanton (1896, p. 1,013) for- 
mally suggested the name "Chico group" 
to describe type-locality exposures 
along Chico Creek. Subsequent workers 
(Turner, 1896; Bryan, 1923; Brewer, 1930; 
Anderson, 1933; Taff and others, 1940; 
Popenoe, 1943; Creely, 1955, 1965) have 
described members and index fossils 
that characterize Chico Formation. 



Fossils in the Chico Formation are 
abundant and provide accurate strati- 
graphic control. Fossils indicate that 
the age of the Chico Formation is Upper 
Cretaceous (Taff and others, 1940; 
Creely, 1965). 

Contact Relationships : Basal contacts 
of Chico rocks are described as angu- 
larly unconformable in the Sierran foot- 
hills (Taff and others, 1940). The base 
of Chico Formation is not exposed in the 
study area, therefore, total thickness 
of the formation is uncertain. The 
thickest sequence in the study area in- 
cludes 20 m (65 ft) of section. 



Upper portions of Chico Formation in 
the project area are eroded and uncon- 
formably overlain by the Tertiary lone I 
and Tuscan Formations. West- and 
southwest-dipping strata in rocks above 
and below the erosional surface are 
slightly discordant and actually form 
a disconformity between Cretaceous and 
Tertiary rocks. 

Lithologic Description ; Cretaceous I 
marine sedimentary rocks, representing 
arc-trench gap deposits (Dickenson, 
1969), are regionally exposed at margins 
of the Central Valley of California and 
represent the base of Superjacent Series 
deposition. Chico Formation is the old- 
est Superjacent Series formation in the 
project area. 

Chico Formation in the study area is 
predominantly a fine- to medium-grained, 
fossil-rich, friable sandstone (arkose); 
siltstone and pebble to cobble conglomer- 
atic lenses occur locally. Fresh Chico 
Formation is light- to dark-buff to 
dark-gray; weathered exposures have 
orangish hues and are lighter in color 
than fresh rocks. Bedding, including 
abundant cross-beds, is thin to thick 
and well-defined. 

Arkosic beds of the Chico Formation are 
moderately- to well-sorted and poorly 
cemented by calcite and clay. Indi- 
vidual clastic grains, forming arkosic 



50 



beds, are angular to subangular. Com- 
position of grains includes quartz, 
feldspar (plagioclase and potassium 
feldspar) and rock fragments (primarily 
metamorphic clasts); accessory ferro- 
magnesian minerals include biotite, 
hornblende, epidote, clinozoisite and 
muscovite. 

Pebble- to cobble-sized clasts are well- 
rounded to sub-rounded and locally form 
interbeds in finer-grained sediments. 
These clasts include light to dark chert, 
quartzite, altered plutonic rocks and 
metavolcanic rocks. Conglomeratic beds, 
commonly containing an abundance of shell 
debris, are usually well-indurated by 
calcite cement. 

lone Formation 

Previous Investigations and Age ; The 
lone Formation was named and first des- 
cribed by Lindgren (1894, p. 3) who 
assigned exposures near lone, California 
as the type locality. Early investiga- 
tions of Tertiary sandstone near Oroville 
were by Lindgren (1911) and Dickerson 
(1916) . Subsequent detailed studies are 
by Allen (1929) and Creely (1965). 

The age of lone Formation is substan- 
tiated by fossil fauna and flora col- 
lected by many earlier researchers. 
These fossils indicate that lone depo- 
sition occurred in Middle Eocene time. 

Contact Relationships : The lone Forma- 
tion rests unconformably on the under- 
lying formations. lone deposits dip 
gently west and southwest and overlie 
arc, melange and the Chico Formation in 
the study area. 

Upper sequences of the lone Formation 
include auriferous gravel and tuffaceous 
sediment and are conformably overlain by 
Lovejoy basalt. Basalt extrusion was 
during late stages of lone aggradation; 
therefore, the unconformity formed by 
basalt at the top of the lone Formation 
is a matter of convention. 

Auriferous gravel and Oroville tuff 



(Mehrten Formation-?) are gradational 
in upper portions of the lone Formation. 
Auriferous gravel and tuffaceous sedi- 
ment, transported and deposited by lone 
fluvial processess, are mapped as forma- 
tional members in this study. Creely 
(1965) mapped quartz-rich sequences as 
"auriferous gravels" and tuffaceous rock 
as "Mehrten (?) Formation". Mehrten 
Formation in the Stanislaus drainage is 
dated by Dalrymple (1964) at 8.8 to 9.3 
million years while tuffaceous deposits 
on South Table Mountain are pre-Lovejoy 
basalt (23 million years old) and 
older than Dalrymple 's dated Mehrten 
Formation. 

Tuffaceous beds are locally exposed 
through the Oroville area. These tuffs 
do not expose basal contacts and are 
overlain by late Cenozoic gravels. Such 
contact relationships provide no strati- 
graphic correlation with tuffs exposed 
on North and South Table Mountains which 
are topographically higher. 

Lithologic Description - lone Formation 
Undifferentiated : White to yellowish- 
white, medium- to fine-grained, silty- 
clayey sandstone constitutes the great- 
est percentage of the lone Formation in 
the study area; intercalated in sand- 
stone are subordinate amounts of silt- 
stone, shale, conglomerate and minor 
quantities of lignitic coal. Conglome- 
ratic beds and pebble stringers are in 
most places composed of well-rounded 
quartz and chert pebbles. Bedding in 
sandstone is thick to thin and 
moderately- to poorly-defined; cross- 
bedding is common and best observed in 
cut slopes. 

Most sandstone is friable, argillace- 
ous and cemented by interstitial silt 
and clay. Individual sand grains are 
angular to subangular and composed of 
quartz, plagioclase, potassium feldspar 
and rock fragments. Trace amounts of 
heavy minerals include hematite, magne- 
tite, epidote, zircon, hornblende, tour- 
maline and clinozoisite. 

Lithologic Description - Auriferous 



51 



Gravel ; Auriferous (gold bearing) gra- 
vel contains high percentages of white 
quartz-rich sand and gravel. In the 
Oroville area this gravel has a maximum 
thickness of 100 m (330 ft) and is ex- 
posed by numerous hydraulic mines cut 
into side slopes of North and South Table 
Mountains (Figure 37) . 

Sand in gravel is medium- to coarse- 
grained, sub- to well-rounded and exhib- 
its fair sorting. Individual sandstone 
layers are thin- to thickly-bedded and 
manifested by slight variations in grain 
size, the presence of thin siltstone or 
pebble conglomerate lenses and thin 
mica-clay layers. 

Conglomeratic sections are composed of 
subrounded to well-rounded quartz peb- 
bles and cobbles. Clasts are loosely 
packed and set in a quartz-sand matrix. 
Individual conglomerate beds range from 



thin pebble stringers in sandstone to 
layers more than 1 m (3 ft) thick. 

Lithologic Description - Orovil le Tuff 
(Mehrten Formation-?) ; Tuffaceous de- 
posits include fine-grained clayey beds 
(relict ash); tuff clasts in coarse- 
grained, water-lain and cross-bedded 
deposits; white, fine-grained, sandy 
beds; and moderately- to well-cemented 
volcanic mudflow breccia. All of 
these rock types, including clay lay- 
era, which possibly represent an air- 
lain derivation, were eroded and trans- 
ported from sources to the east and 
north. 



Light colored and cross-bedded, sandy, 
tuffaceous sequences are locally exposed 
around Oroville and represent fluvial 
deposition. Whether separated tuffa- 
ceous outcrops represent rock- or time- 









I 



/* 



r 



■■«>*. 



Figure 37- lone Formation auriferous gravel 

(Mehrten Formation-?) in a hydraulic mining cut on the east side of 
Table Mountain. 




52 



stratigraphic horizons is vmcertain. 

Mudflow volcanic breccia is formed by 
angular to sub-angular, vesiculated and 
amygdaloidal rhyodacite clasts set in a 
reddish-brown, sandy-silty matrix. Brec- 
ciated clasts are not locally derived. 
Source areas of the lone Foinnation sug- 
gest the mudflow breccias were also de- 
rived from east and north of the area. 

Love joy Formation 

Previous Investigations and Age ; The 
basalt on Oroville Table Mountain was 
first mapped and named "older basalt" 
by Turner (1894) to differentiate the 
unit from younger flows in the area. 
This basalt was correlated with the 
Lovejoy Formation by Durrell (1959b, 
1966) which he considered, based on 
stratigraphic relationships, to be of 
Eocene age and derived from areas east 



of the present Sierra Nevada crest. 
Dalrymple (1964) radiometrically dated 
rocks above and below Lovejoy Formation 
and determined the age of basalt to be 
Early Miocene; his oldest date, 23 mil- 
lion years, was obtained from a tuff 
bed below Lovejoy Formation on South 
Table Mountain and should be a maximum 
age for basalt in this area. 

Contact Relationships ; Lovejoy Forma- 
tion in the study area disconformably 
overlies lone Formation and Oroville 
tuff (Mehrten Formation-?) (Figure 38); 
basalt rests unconformably on ophiolite 
and melange in two localized areas but 
this relationship is not common. Lower 
contacts of basalt are nearly planar and 
dip 2 to 3 degrees west-southwest. Upper 
and lower planar contacts indicate that 
the basalt has experienced little defor- 
mation during regional westward tilting 
and provide control for post-extrusive 
(late Cenozoic) faulting. 




Figure 38. View south of Lovejoy Formation basalt disconformably overlying 
lone Formation sedimentary rock in hydraulic cut face of the Cherokee Mine. 



53 




Figure 39. View east from upper reaches of Morris Ravine of Lovejoy Formation 
basalt on North Table Mountain. Basalt at this location has a minimum thick- 
ness of 75 m {2hG ft) and rests d i sconformably on lone Formation. 



Lithologic Description ; Lovejoy Forma- 
tion forms the flat-topped mesas of 
North and South Table Mountains 
(Figure 39). Lovejoy basalt includes 
one or more sub-horizontal flows with a 
cummulative thickness in the Oroville 
area of less than 50 m (164 ft). 

Lovejoy basalt is dark-brown to black 
and forms blocky outcrops. Poorly- 
developed columnar jointing is common in 
upper parts of the formation. Lower 
parts of the formation are generally 
fragmented and locally include a basal 
conglomerate. Vesiculated basalt is 
more abundant near the base of the for- 
mation. Basalt mineralogy includes 
plagioclase (An,e to An , _) , olivine and 
traces of augite. Plagioclase microlites 
are abundant in some samples. A crystal- 
line to glassy matrix comprises the 



54 



greatest volume of basalt. . 

Tuscan Formation 

Previous Investigations and Age ; Rocks 
of the Tuscan Formation were first des- 
cribed by Whitney (1865). Diller (1892, 
1895) named the formation and described 
the type locality at Tuscan Springs in 
Tehama County. 

Anderson (1933) published a comprehen- 
sive paper on the Tuscan Formation. 
This work includes many detailed rock 
descriptions and a discussion on the 
development of Tuscan breccia. 

Recent studies of the Tuscan Formation 
are by Creely (1965) and Lydon (1968). 
Lydon's work is comprehensive and deals 
with the source areas for the rocks. 



Tuscan Formation is Late Pliocene in 
age (Lydon, 1968). A potassium-argon 
age of 3.3 million years (Evernden and 
others, 1964) is determined for the 
Nomlaki Tuff member of the formation. 

Contact Relationships ; The Tuscan For- 
mation unconformably overlies melange 
in the study area; locally Tuscan rocks 
rest disconformably on Chico and lone 
Formations. Basal contacts of Tuscan 
Formation indicate the depositional sur- 
face is relatively flat and dips slightly 
to the southwest. This horizon trends 
below alluvium of the Sacramento Valley. 
Upper surfaces of Tuscan flows are rela- 
tively planar (Figure 40) and dip at low 
angles to the southwest. These flows 
are deeply incised by westerly flowing 
drainages. 



Lithologic Description ; The volcanic 
Tuscan Formation is composed of lahars, 
volcanic sand, conglomerate, tuff, tuff 
breccia, and intercalated andesite and 
basalt flows. These rocks when fresh 
are gray, purple, orange or brown. The 
maximum formational thickness in the 
study area is 180 m (590 ft) . 

Tuff breccia (lahar) forms about 75 per- 
cent of the formation. Clasts are 
basalt and andesite with basalt being 
predominant (Anderson, 1933). Flow brec- 
cias are unsorted and form irregular 
contacts with underlying rocks. The 
matrix of these rocks is well-indurated 
volcanic and tuffaceous sand. Interca- 
lated flow rocks, a minor component of 
the Tuscan Formation, are predominantly 
olivine basalt and pyroxene andesite. 



METAMORPHIC SURFACE 



TUSCAN SURFACE 




Ik'^vifcv 




Figure ^0. View north of lower and younger erosional surface on Upper Pliocene 
Tuscan Formation that is separated by the West Branch of the Feather River 
(not shown) from an older and structurally higher erosional surface cut into 
Mesozoic metamorphic rocks. Photograph taken from intersection of Highway 70 
and Messilla Valley Road (Cherokee quadrangle). 



55 



Tuff, tuffaceous sandstone and volcanic 
sandstone are locally intercalated with 
the flows and breccias. These units are 
composed of angular crystal and lithic 
volcanic fragments with andesitic to 
basaltic compositions. Sequences are 
well-bedded, well-sorted, and commonly 
cross-bedded (Figure 41) . Sediments 
are common at the western margins of the 
formation. Tuff breccia dominates the 
stratigraphically thicker eastern areas 
of Tuscan exposure. 

Late Cenozoic Gravels 

Previous Investigations : Late Cenozoic 
fluvial deposits of the Oroville area 
were first differentiated by Creely 
(1965). He assigned all older gravels 
in the area to the Pleistocene Red Bluff 



Formation. 

Recent mapping in the Bangor quadrangle 
(Quintin Aune, unpub. data) indicates 
there are several gravel units of vary- 
ing ages and source areas. Mapping for 
this study confirms the presence of mul- 
tiple gravels that probably are not time 
equivalent to the Red Bluff Formation. 
Therefore, they are named "late Cenozoic 
gravels" rather than Red Bluff Formation 
in this report. 

Contact Relationships ; Late Cenozoic 
gravels overlie both Basement Series 
rocks and Superjacent Series rocks in 
the project area. They are separated 
from basement rocks by an angular uncon- 
formity and from superjacent rocks by a 
disconf ormity . 




Figure 41. Tuscan Formation volcanic conclomerate, cross-bedded sand and 

laharic mudflow breccia along Sycamore Creek 3 km (2 mi) northeast of Chico. 



56 




Figure hi. Late Cenozoic gravel and cross-bedded sand (Red Bluff Formation-?] 
exposed in a railroad cut 0.5 km (0.3 rni) south of intersection of Baggett 
Palermo and Baggett Marysville Roads (Palermo quadrangle). 



Lithologic Description ; Late Cenozoic 
gravels in the study area have a maximum 
thickness of 30 m (100 ft) and are com- 
posed of poorly-sorted, rounded to sub- 
rounded, pebble- to boulder-sized clasts. 
These are weakly to moderately cemented 
by varying amounts of clay, silt and 
orange amorphous silica; cementation is 
weak where sandy and moderate in clayey 
sections. Clast types, in descending 
order of abundance, are metavolcanic 
rock (including ophiolite, arc and 
younger dike rocks), intrusive rocks, 
and fine-grained porphyritic volcanic 
and siliceous clasts (including quartz, 
quartzite and red and black chert). 
Imbricate pebbles indicate source areas 
are to the north and east. Well-sorted 
and cross-bedded, weakly-cemented sand, 
and thin-bedded, moderately indurated 
silt and clay comprise the gravel matrix. 



Sandy members of gravels are generally 
thinly- to moderately-bedded, lenticular 
and, in places, cross-bedded (Figure 42). 
Fine-grained sands and silty-clayey mem- 
bers, most common away from upland ter- 
ranes, probably represent flood plain 
deposits of the ancestral Feather River 
and associated tributaries. 

Clayey sections in gravel have minor 
occurrences in the study area. Clay in 
gravel sequences, probably reworked from 
underlying tuff units (Figure 43), repre- 
sents low energy deposition. 

Quaternary Landslides 

Our investigation indicates that large- 
scale landsliding is more common in the 
project area than suggested by earlier 
detailed investigations. Failures com- 



57 



monly occur from slopes underlain by 
lone Formation (Figure 44); this forma- 
tion is the least competent of study 
area rock types. 

North and South Table Mountains and the 
Campbell Hills are capped by Lovejoy 
basalt and have side-slopes underlain by 
gently west-dipping lone Formation; 
south- and west-facing slopes in these 
areas daylight lone bedding. Resistant 
cap rock overlying non-resistant lone 
Formation provides ideal conditions for 
large-scale landsliding. In this situa- 
tion, side-slopes are oversteepened by 
artificial support of the erosionally- 
resistant cap rock. A regional slope- 
stability study of the United States 
notes the western side-slopes of North 
and South Table Mountains at Oroville 
are highly susceptible to failure 
(Radbruch-Hall and others, 1976). 



Landslides were not mapped in detail 
on Table Mountain or Campbell Hill side- 
slopes for this study. The time requir- 
ed for mapping gravity-induced struc- 
tural complexities was not warranted for 
purposes of this investigation; there- 
fore, landslides probably underlie more 
area than is indicated by our geologic map. 

Numerous landslides occur along the 
Feather River and its major forks. Fail- 
ures in this area are within arc and 
ophiolitic lithologies. - The toe por- 
tions of these landslides occur near 
lower valley slopes and are now season- 
ally inundated by Lake Oroville. Land- 
slide movements are mostly prehistoric, 
however, several failures indicate re- 
cent activity. The largest recent land- 
slide is superimposed on an older fail- 
ure that moved from the northwest side 
of Stringtown Mountain (Figure 45). 




Figure ^43. Late Cenozoic gravel (Red Bluff Formation-?) unconformabl y overlying 
Oroville tuff (Mehrten Formation-?) along the Feather River 1.0 km (0.6 mi) west- 
southwest of Oroville. 



58 




Figure 44. View west of landslide in lone Formation. Note the vegetation stand 
in graben area of landslide. Location is in Campbell Hills just north of 
Thermalito Forebay by Highway 70 (center). 




Figure 45. Aerial southeast view of Stririgtown Mountain landslides. Note that 
the recent failure is superposed on a larger and older landslide. 



59 



\ 







i 






Figure 46. Aerial east-southeast view of a prehistoric landslide that is part 
of a much larger failure involving the entire north slope of Bloomer Hill into 
the North Fork of the Feather River. 



60 



The largest landslide in the project 
area, underlying the north slope of 
Bloomer Hill, is a failure of arc rock 
into the North Fork of the Feather River 
(Figure 46). The landslide moved north 
as a large slump of several individual 
failures. This landslide could have 
temporarily dammed the river. Arcuate 
scars of disturbed arc rock define the 
landslide boundaries which are best ob- 
served using high-altitude aerial 
photographs. 

STRUCTURAL GEOLOGY 

Faults 



Geologic evidence in the northern 
Sierran foothills suggests two periods of 
fault activity. The first episode of 
faulting was from compression before 
Late Jurassic time. This deformation 
occurred prior to the intrusion of local 
plutons (Nevadan orogeny-Yosemite intru- 
sive epoch) . The second period of 
faulting began in late Tertiary time and 
continues to the present. The late Ter- 
tiary to present tectonic regime is one 
of east-west extension which places older 
fault zones in tension. As a result, 
some recent movements have occurred along 
older Mesozoic faults. Other movements 
displace Tertiary rocks and have thus 
broken new ground, possibly from reacti- 
vation of underlying Mesozoic faults. 

Data on foothills faulting between 
Oroville and Sonora is derived largely 
from exploration trenches by Department 
of Water Resources and other agencies. 
Locations of exploration trenches on ma- 
jor lineaments are shown in Figure 47 
and findings are summarized in Table 1. 




BASE MAO 
WOODWARD 



MODIFIED AFTER / 

YOE CONSULTANTS f 



Figure Ul . Major lineaments in the northwestern 
Sierran foothills showing exploration localities 
with faulting assessments for each site. 



61 



TABLE 1 
EXPLORATION TRENCHES IN FOOTHILL BELT - OROVILLE TO AUBURN AREA 
(Exploration sites listed from north to south on given lineaments) 



Lineament , 
Agency, and 
Trench Number 

SWAIN RAVINE LINEAMENT 



Faulting 
Exposed 



Attitude 



Cenozoic 
Movement 



No 

No 
Yes 



No 
Yes 



No 
Yes 



N10-40W, 60-73SW 



N12W, 65SW 



NlOW, 69SW 



N20W. 70SW 



N15W, 65SW 



Yes Trench located on crack; faulting 

does not offset soil-bedrock inter- 
face. 

Yes Fault does not offset bedrock-soil 
interface . 

Yes Trench located on crack; faulting 

does not offset bedrock-soil inter- 
face. 

Yes Trench located on crack; faulting 

does not offset bedrock-soil inter- 
face . 

Yes Trench located on crack; faulting 
offsets gravel-soil contact 30 mm. 

? Faulting does not offset bedrock- 
soil contact. 



N9W. 54SW 



USCE (WCC) 
Grubbs 1 



PGandE (WCC) 
Grubbs 2 



PGandE (WCC) 
Sims 1 



USCE (WCC) 

Cleve. Hill 1 & 2 



PGandE (WCC) 

Lorraine 1 & 2 



Trench located on crack; faulting 
does not offset bedrock-soil inter- 
face . 

Trench located on East Mission 
Olive crack zone. 

Trench located on East Mission 
Olive crack zone. 



Trench located on West Mission 
Olive crack zone. 

Trenches located on northern end 
of Cleveland Hill Fault; bedrock 
fault with at least 3 episodes 
of displacement described. 

Trenches located on eastern splay 
of Cleveland Hill Fault. 



PGandE (WCC) 

Cleve. Hill 3 



USER (WCC) 

Orange Road 1-9 



PGandE 


WCC) 


Pace 


1-5 


USCE 




4F-1 




USCE 




4F-2 




4F-3 





N11-21E, 55-59SE; 
N18-20W, 80NE 



Yes 


NSW, 


41NE, 




N15W 


4 7NE 


No 




__ 


Yes 


N15W 


46NE 


Yes 


N2E, 


52SE, 




N22W 


48SW 


Yes 


N32W 


70NE 




N34W 


70SW 



Trenches located on Cleveland Hill 
Fault at southwest margin of 
Cleveland Hill. 

Trench located on Cleveland Hill 
Fault at southwest margin of 
Cleveland Hill. 

Faults located in trenches 2, 3, 4, 
5,6,9 

Faulting does not offset bedrock- 
soil interface. 

Faulting does not offset bedrock- 
soil interface. 



Faulting does not offset bedrock- 
soil interface. 

Faulting does not offset bedrock- 
soil interface. 

Faulting does not offset bedrock- 
soil interface. 



62 



TABLE 1 (Continued) 



Lineament, 
Agency, and 
Trench Number 

PAYNES PEAK LINEAMENT 

PGandE (WCC) 
Knapp No. 1 

PGandE (WCC) 

Burt No. 1 & 2 



Faulting 
Exposed 



Yes 



Yes 



Cenozoic 
Movement 



No 
No 



Faulting does not continue into 
overlying soils. 



PRAIRIE CREEK LINEAMENT 

DWR 18 No 

PGandE (WCC) 

O'Brien No 

PGand E (WCC) 

Wilson No. 1 & 2 Yes 



No 

No 



No faults exposed and trench not 
logged. 



Faulting does not continue into 
overlying soils. 



SPENCEVILLE LINEAMENT (Southern extension of Prairie Ck . Line.) 



US BR (WCC) 

Spenceville 1 Yes N32W, 63SW; 

(5 trenches) N55W, 67SW 

2 Yes 



Yes 
No 



Yes N15-50W, 70-75SW Yes 
No - - No 

No — No 



Faulting continues into overlying 

soils . 

Faulting does not continue into 

overlying soils. 

Faulting displaces paleo B. 



DEADKAN LINEAMENT (Southern extension of Spenceville and Prairie Ck . Line.) 



USER (WCC) 

Henriques & Wilson Yes N20W, 47SW 
(10 test pits) N20W, 55SW 



No fault assessment made because 
paleo B too scarse in local area 
for evaluation 



DEWITT LINEAMENT 

USBR (WCC) 

Hubbard Road 
(2 trenches) 



USBR (WCC) 
Bean Road 
(1 trench) 



)R (WCC) 
St. Joseph 
( 3 trenches ) 



N44-50W, 50-60SW Yes ( ? ) 



N60W, 65NE 



N38W, 60NE 



Faulting in paleo B but gravity 
also affects rocks making inter- 
pretations difficult; faulting 
classified (USBR criteria) 
"indeterminate" (active). 

Faulting does not continue upward 
into paleo B; faults classified 
(USBR criteria) "indeterminate" 
( inactive ) . 

Paleo B locally scarce, therefore, 
faulting at this locality classi- 
fied (USBR criteria) "indeterminate" 



MAIDU LINEAMENT 

USBR (WCC) 

Radio Tower (Located 

on E. splay of Maidu 

Line. ) 

(6 trenches) 
USBR (WCC) 

Maidu 

(2 trenches) 

(4 test pits) 



N55-60E, 30NW-90 ? 



NIOE, 45NW 



Paleo B and overlying soils 
locally scarce, therefore, faulting 
at this locality classified (USBR 
criteris) "indeterminate". 

Paleo B and overlying soils locally 
scarce, therefore, no fault 
assessment made for Late Cenozoic 
tectonics at this locality. 



63 



TABLE 1 (Continued) 



Lineament, 
Agency, and 
Trench Number 

USER (WCC) 
Maidu East 
(E splay of Maidu 
Line. ) 

( 5 trenches by WCC 
plus 25 trenches 
and backhoe pits 
by USER) 



Faulting 
Exposed 



Cenozoic 
Movement 



N13E, 77SE, 
N55W, 47SW, 
N3E, 67SE, 
N30W, 67SW, 
N20-25E, 82NW-90, 
N20-30E, 72-80NW 



PILOT HILL LINEAMENT 

USER (WCC) 

Pilot Hill 
( 3 test pits ) 

USER (WCC) 

Salmon Falls 
(4 test pits) 

SALT CREEK LINEAMENT 



No(?) 



Maximum vertical separation of 
Mehrten Fm. across fault zone is 
5.4 m (18 ft). Slickensides in 
soil with orientations similar t 
to bedrock faults and steps in 
colluvial base overlying bedrock 
fault traces indicate faulting 
is Cenozoic. To north a buried 
paleosol at least 100,000 years 
old is not cut by faulting; 
therefore, movements are too small 
to offset soils or fault displace- 
ments die out to north. Faulting 
confidence level is 2 on 0-10 
scale . 



Paleo E and overlying soils scarce, 
therefore, no fault assessment made 
for this locality. 

Thin shears exposed but no faults; 
lack of local paleo B for offset 
control. No fault assessment. 



USER (WCC) 
Salt Creek 
( 10 test pits) 

USER (WCC) 

Bayley House 
( 3 trenches ) 
( 12 test pits) 



RESCUE LINEAMENT 

USSR (WCC) 

Luneman Road 
( 3 trenches ) 



N54, 60SW 
N40W, 40NE 



USER (WCC) 
Knolls 
(1 trench) 



N20W, 50SW 



Paleo B locally lacking, therefore, 
no fault assessment made for this 
area . 

Ground water barriers define the 
lineament but are controlled by 
clay-rich weathering zones; no 
evidence of Cenozoic faulting noted. 



Faulting trends into overlying 
colluvium and terminates a 
paleo E with colluvium on east 
side of fault thicker than on 
west side; paleosol indicates 
0.55 m (1.8 ft) of down to east 
displacement. Fault in trench 
3 classified as "indeterminate" 
(active) by USER criteria; 
confidence level is 4 on scale 
0-10. 

Distinct lithologic blocks are 
bounded in places by clay seams 
that appear to juxtapose the 
blocks; basal contact is not 
obviously offset. Faulting 
classified by USER criteria as 
"indeterminate" (active). 
Confidence level is 2 on scale of 
0-10. 



64 



Mesozoic Faults - Northern Foothills 

Clark (1960) identified and named the 
Foothills Fault System (Figure 48). 
This system, bounded by the Melones 
Fault zone on the east and the Bear 
Mountains Fault zone on the west, is ' 
formed by numerous north to north- 
northwest trending preintrusive reverse 
faults (Clark, 1964, 1976). Major faults 
within this system can be identified by 
elongate bodies of serpentine, areas of 
structural and lithologic discontinuity 
and zones of intense and well-defined 
shear cleavage that dip steeply east. 

Subsequent to Clark's initial study, 
Mesozoic faults in the foothills were 
described and mapped in many geologic 
investigations. A few of these studies 
include works by Baird (1962), Burnett 
and Jennings (1962) , Bateman and others 
(1963), Clark (1964, 1976), Creely (1965), 
Cebull (1972) and Hietanen (1973a, 1976, 
1977). 

The Melones Fault zone, named by Clark 
(1960), strikes northwest along the east- 
ern margin of the Foothills Fault System 
in its type locality. The fault is de- 
fined by strongly sheared zones that, in 
places, incorporate serpentine and blocks 
of undeformed or less deformed rock. 
Shear cleavage within the zone is local- 
ly several hundred metres wide and dips 
vertically or steeply east (Clark, 1960, 
1964). 

The Melones Fault zone, best exposed 
south of the Cosumnes River, is the east- 
ern limit of the Foothills Fault System 
(Jennings, 1975). Clark (1960) noted 
that north of the American River the 
Melones Fault splits into several zones. 
The splay representing the Melones Fault 
zone in this area is defined as the 
boimdary between Paleozoic rocks to the 
east and Mesozoic rocks to the west 
(Clark, 1960, 1964; Duffield and Sharp, 
1975). 

The Bear Mountains Fault zone of Clark 
(1960) parallels the trend of the Melones 
Fault zone to the east and splits into 



several faults at its northern end near 
the Cosumnes River. The regional shear 
zone mapped by Burnett and Jennings 
(1962) to the southwest of the area may 
represent the northern extension of the 
Bear Mountains Fault zone. This fault 
zone averages a few hundred metres in 
width and dips vertically or steeply 
east (Clark, 1964). Net displacement 
across the system is unknown byt suggested 
to be large and probably represents sev- 
eral thousand metres of offset (Clark, 
1964). 

Origin of the Bear Mountains Fault, as 
with other faults in the Foothills Fault 
System, is the result of eastward under- 
thrusting during Farallon-American plate 
interactions in Late Jurassic time. East- 
ward underthrusting is suggested by some 
early researchers (Ferguson and Gannet, 
1932, p. 90; Knopf, 1929, p. 45-46); how- 
ever, a strike-slip motion, at least in 
part, is indicated by Clark (1960) and 
Cebull (1972). 

Mesozoic faults in the study area are 
considered to be part of the Foothills 
Fault System. These faults displace the 
late Oxfordian to early Kimmeridgian 
(Imlay, 1961, p. D8-D9) Monte de Oro 
Formation and are truncated by Sierran 
plutons of the Yosemite intrusive epoch. 
Radiometric dating of the plutons (Gromme 
and others, 1967; Evemden and Kistler, 
1970) yield a minimum age of about 
130 million years. These data suggest 
that Mesozoic faults developed during 
the Late Jurassic-Early Cretaceous 
Nevadan orogeny, about 130 million years 
ago. 

Foothills system faults were driven 
by a regional east-west compression and 
are synchronous with late stages of an 
epidote-albite-amphibole metamorphism. 
Compressive stresses and subsequent 
Foothills Fault System displacements were 
generated during subduction underthrust- 
\ ing and accretion of arc and ophiolitic 
^ rocks to Mesozoic California. This de- 
formation generated north-striking, 
steeply dipping faults, fold axes and 
slaty cleavage in rocks of the study 



65 




Data modified after Clark (I960). 
Figure h8. Foothills Fault System of the western Sierra Nevada, California 



66 



area. Deformation ceased when these 
rocks were firmly accreted (obducted) to 
the continent. Additionally, the west- 
erly migration of subduction was stabi- 
lized in areas of the present Coast 
Ranges terrane at this time. 

Mesozoic Faults - Project Area 

Mesozoic faults in the study area com- 
monly appear as photo lineaments. These 
lineaments, sharply defined in high- 
altitude photography, are commonly 
aligned with foliation and fold axes 
in foothill rocks. 

Field investigations of major lineaments 
indicate three are fault zones. Fault 
features include: (1) pronounced align- 
ment of ridges and valleys along linea- 
ment trends, (2) sheared rocks and 
numerous subordinate faults and shears 
subparallel with major lineaments, and 
(3) springs and seeps. 

Mesozoic fault movements were probably 
oblique, however, a large reverse compo- 
nent is indicated by many researchers 
(Hamilton, 1969; Schweickert and Cowan, 
1975; Clark, 1976; Russel, 1978; 
Standlee, 1978), Reverse, east-dipping 
Mesozoic faults are predictable in mo- 
dels of eastward subduction which was 
active at this time; the Glover Ridge 
thrust fault, an obduction suture, is 
an exception. 

Plate tectonic models for explaining 
the origin of Mesozoic foothills faults 
suggest large-scale movements. Displace- 
ments on larger foothills system faults 
such as the Melones Fault zone may ex- 
ceed several kilometres. 

In summary, the Foothills Fault System 
is a late Paleozoic to Late Jurassic 
feature. Compression driving these 
generation faults originated from epi- 
sodes of plate convergence and consump- 
tion along the western edge of Mesozoic 
North America, and produced major struc- 
tural elements of the Sierran foothills. 



Swain Ravine, Paynes Peak and Prairie 
Creek Lineament/Fault Zones ; The Paynes 
Peak and Swain Ravine Lineaments are the 
most striking photo lineaments in the 
area. Another prominent lineament, the 
Prairie Creek Lineament, projects into 
the study area from the south, west of 
the Swain Ravine Lineament . The trace 
of the Prairie Creek Lineament within 
the study area is not well defined. 

The Swain Ravine and Paynes Peak 
Lineaments trend approximately north- 
northwest and parallel each other in the 
southern field area. The two lineaments 
can be traced on the ground by aligned 
valleys, discontinuous areas of sheared 
rocks, springs and seeps, and are inter- 
preted to be Mesozoic fault zones. 

The Paynes Peak Lineament in the study 
area has a strong to moderate expression. 
It lies parallel to, and about 1.6 km 
(1 mi) east of the Swain Ravine Linea- 
ment. Surfically, the only conclusive 
fault features are exposed in Rocky 
Honcut Creek (Bangor quadrangle. 
Section 16, T18N, R5E) where a break in 
outcrops, a linear drainage and aligned 
springs define the fault trace. 

The northern extension of the Paynes 
Peak Lineament trends just east of 
Miners Ranch Reservoir and Bidwell Canyon 
Saddle Dam. Rock along the lineament is 
strongly sheared, however, field evi- 
dence for faulting north of Rocky Honcut 
Creek is poor. The Paynes Peak Lineament, 
as defined in this report, coincides 
with the eastern margin of the "regional 
shear zone" as mapped by the U. S. Army 
Corps of Engineers (1977, Plate V) and 
terminates about 1 km (0.6 mi) northeast 
of Bidwell Canyon Saddle Dam. 

South of the project area the Paynes 
Peak Lineament has a strong topographic 
expression. The lineament terminates 
north of Paynes Peak in the vicinity of 
Stone House (U. S. Army Corps of 
Engineers, 1977, Plate IV). 



67 



Surface and subsurface data along the 
Paynes Peak. Lineament were collected and 
analyzed by Pacific Gas and Electric 
Company and Woodward-Clyde Consultants. 
Three trenches, designated Knapp No. 1 
and Burt Nos. 1 and 2, were excavated 
between Bangor and the Yuba River by 
Pacific Gas and Electric Company. These 
trenches exposed bedrock faults that do 
not displace overlying soils. 

The Swain Ravine Lineament is the most 
significant lineament in the area because 
it coincides with Cleveland Hill Fault 
cracking that occurred during the 1975 
Oroville earthquake. No cracking along 
the lineament is reported south of Bangor; 
however, the lineament continues southward 
as a strong feature. Just south of the 
Yuba River, the lineament coincides with 
the eastern margin of a regional shear 
zone mapped by Burnett and Jennings (1962) 
and the U. S. Army Corps of Engineers 
(1977). A more thorough description of 
the Swain Ravine Lineament is included 
with the section on Cenozoic faults. 

The Prairie Creek Lineament has the 
least physical expression of the three 
major lineaments in the study area. The 
lineament, well-developed south of the 
Yuba River, is discontinuous to the 
north. It is prominent in the southeast 
corner of the Palermo quadrangle, but 
dies out to the north where the area is 
overlain by alluvium and late Cenozoic 
gravels. The length of the Prairie 
Creek Lineament from the study area to 
its southern end, where it is truncated 
by the Rocklin pluton, is approximately 
64 km (40 mi). 

Investigation of the Prairie Creek 
Lineament between Bangor and the Yuba 
River consisted of field mapping and a 
trench (DWR 18) in the Palermo quad- 
rangle by Department of Water Resources 
(this study) and three trenches by 
Pacific Gas and Electric Company. Two 
of the Pacific Gas and Electric Company 
trenches (Wilson, 1 and 2), located 
about 4.8 km (3 mi) northwest of Browns 
Valley, exposed bedrock faults which 
did not displace overlying soils. The 



68 



third trench (O'Brien) about 3.2 km 
(2 mi) northwest of Loma Rica exposed 
no fault "structures. 

In November 1977, Department of Water 
Resources personnel trenched the north- 
ern end of Prairie Creek Lineament about 
855 m (2,800 ft) north of Cox Lane 
(NEl/4 Section 27, T18N, R4E) . Aligned 
bedrock ridges and a small depression 
(sag pond-?) define the Prairie Creek 
Lineament in this area and suggest it is 
a fault. The trench exposed only strong- 
ly weathered, undisturbed bedrock and 
was backfilled without being logged. 
That our trench near Cox Lane and the 
O'Brien trench to the south did not ex- 
pose faulting may mean that the fault 
trends east or west of the exploratory 
trenches, or that it does not continue 
into this area. 

It is assumed the Swain Ravine, Paynes 
Peak and Prairie Creek Lineaments are 
complex zones of Mesozoic faulting. 
These fault zones are probably bands of 
small, discontinuous faults. Conclusive 
evidence in the way of trenches or clear- 
ly exposed faults is not available to 
j-iTove that this assumption holds true 
throughout the extent of the lineaments, 
however, other field evidence indicates 
it is a reasonable assumption. The 
sense of displacement along these linea- 
ments cannot be determined by local field 
relationships. 

Oregon Gulch Fault : The Oregon Gulch 
Fault juxtaposes arc rocks on the west 
against ophiolite on the east. The 
Fault was first mapped by Creely (1965) 
in his study of the area. Schweickert 
and Cowan (1975, p. 1,330) show the 
fault as a continuation of the Bear 
Mountains Fault zone. Our work indicates 
this fault is much smaller than the Bear 
Mountains Fault and that the two are 
probably not physically continuous. They 
are, however, interpreted to have formed 
during the same period of Mesozoic time. 

The Oregon Gulch Fault is traceable for 
approximately 29 km (16 mi) on a north- 
south trend through the central portion ^ 
of the area; the fault is obscured lo- 



cally along this trend by unconformably- 
overlying late Cenozoic terrestrial 
gravel and tuff. The fault zone, moder- 
ately defined where exposed, is less 
than 5 m (16 ft) wide and dips verti- 
cally or steeply east. A trench across 
the fault exposed juxtaposed arc and 
ophiolitic rocks, but the exact contact 
was not well-defined (see log. Trench 15 
on pages 118-119). 

Monte de Pro Fault : The Monte de Oro 
Fault, first mapped by Creely (1955), is 
exposed for 6 km (4 mi) in an approximate 
north-south trend through the central 
portion of the area. It is overlain in 
its north and south projections by Ter- 
tiary Superjacent Series rocks. The 
fault dips east and truncates Monte de 
Oro Formation against arc rocks on the 
east. Quartz is locally intruded into 
the fault zone and helps to define its 
location in poorly exposed areas. 

Unnamed Faults ; An unnamed fault, 
subparallel with previously described 
foothills faults, is exposed 1 km 
(0.6 mi) north of Oroville Dam. The 
fault is traceable in a north-south 
trend for 3 km (2 mi) just west of the 
North Fork of the Feather River. Gouge, 
aligned valleys and seeps define the 
trace. 

Small faults and shears are exposed 
locally through the area. These faults, 
traceable only for short distances, 
probably result from sjTnpathetic dis- 
placement and fracturing related with 
Foothills Fault System activity. 

Foothills system faults in melange at 
the north end of the area strike approxi- 
mately northwest and dip steeply to the 
northeast. A reason for the change in 
strike of the Foothills Fault System 
from north to northwest in this area is 
uncertain. 

Northern area faults can be differen- 
tiated into (1) faults that cut melange 
and (2) faults that are associated with 
serpentine. Faults associated with ser- 
pentine are suggested to represent rem- 
nant Benioff zones (Hamilton, 1969; 



Bailey and others, 1970; Bateman and 
Clark, 1974; Coleman, 1977). Occur- 
rences of serpentine were used in this 
study to identify the location of such 
faults. Faults that cut melange prob- 
ably formed at the same time as those 
associated with serpentine within the 
subduction complex accretionary prism. 

Accretionary prism models (after Karig 
and Sharman, 1975; Dickinson, 1975) 
indicate that large numbers of concor- 
dant reverse faults are formed in the 
accumulation and development of these 
regions. Poor exposure and large areas 
of similar rock type that are concordant 
with regional trends preclude identifi- 
cation and mapping of many pre-Cenozoic 
faults in melange terrane; therefore, 
the number of these faults shown on the 
geologic map (Plate 1) is probably fewer 
than those actually present. Extreme 
shearing is visible in many melange out- 
crops suggesting that faults in this 
terrane are numerous and conform to sub- 
duction complex models. Faulting within 
this complex incorporates and separates 
individual blocks of rock that contain 
smaller-scale faults concordant with 
regional trends. 

The amount of offset on faults in melange 
is not known. Plate tectonic models of 
continued underthrusting at convergent 
boundaries suggest cumulative displace- 
ments are large and may represent several 
kilometres . 

Glover Ridge Fault ; A major Mesozoic 
thrust fault, informally named the 
Glover Ridge Fault, is in the south- 
eastern Cherokee quadrangle. Glover 
Ridge is a klippe (Figure 49) separated 
from the main thrust sheet of arc rock 
by incised erosion in Vinton Gulch. The 
fault was mapped by Creely (1965) as an 
intrusive contact, however, clay gouge 
between arc rock and underlying melange 
indicates a fault, rather than an intru- 
sive contact. Approximately 23 km 
(14 mi) of fault trace is exposed in 
the area (Plate 1) . 

The Glover Ridge Fault trends into the 
southwestern Berry Creek quadrangle 



69 




Figure kS.^'^^fla] southeast view of G 
location of the Glover Ridge Fault, 
photograph. 

where its exact location is uncertain. 
The fault represents a thrust system 
that overrode melange; its presence 
should be suspected wherever arc or 
ophiolitic rocks overlie melange. Sev- 
eral exposures of isolated arc rock in 
the Big Bend area are in contact with 
melange metasedimentary rocks and may 
represent klippen or exotic blocks in- 
corporated in melange from partial sub- 
duction of the arc complex. 

Topographic configuration indicates the 
Glover Ridge Fault dips to the south- 
west. The contact is nearly flat-lying 
in exposed areas but in the subsurface 
to the south and west may be more steep- 
ly inclined. The Glover Ridge Fault is 
interpreted to represent a portion of an 
obduction suture along which arc and 
ophiolitic rocks were thrust over me- 
lange during accretion to continental 
terrane. 

Extent of the obduction suture is uncer- 
tain in the Sierran foothills because 
exposures of the contact are limited. 
A portion of possibly the same, or a 
contemporaneous thrust fault is exposed 



70 



over Ridge (klippe) and the traced 
West Branch Bridge is center-left in 

near the Bear River southeast of 
Marysville (Xenophontos and Bond, 1978). 

Another fault discordant to the Foothill 
Fault System is exposed near Box Hall 
Flat, approximately 5 km (3 mi) north of 
Oroville Dam. The fault, strongly pro- 
nounced in aerial photographs and poorly 
exposed on the ground, can be traced for 
approximately 3 km (2 mi) in a northeast- 
southwest direction. Crushed rock and 
gouge are exposed locally along the fault 
trace. Sense of slip and total offset 
on this fault are unknown because fault- 
ing is entirely within Smartville ophio- 
lite. Development of this fault is 
probably synchronous with obduction and 
therefore represents late-stage develop- 
ment of the Foothills Fault System. 

Cenozoic Fault Movement 

Cenozoic faulting in the Sierran foot- 
hills has been described by several geol- 
ogists (Lindgren, 1911; Ferguson and 
Gannet, 1932; Durrell, 1959a; Burnett 
and Jennings, 1962; Strand and Koenig, 
1965; Jennings, 1975, 1977; Alt and 
others, 1977). Most of the recognized 



Cenozoic faults are northwest-striking, 
east-dipping, high-angle, normal faults 
(summary in Alt and others, 1977, 
pp. 33-35). 

Cenozoic fault movements commonly occur 
along older, Mesozoic faults (Alt and 
others, 1977). Not all Mesozoic faults 
have experienced reactivation, but they 
must be considered avenues along which 
fault movements preferentially occur. 

Woodward-Clyde Consultants compiled data 
on 46 faults in the northern Sierra 
Nevada having evidence for probable late 
Cenozoic displacements (Alt and others, 
1977). Vertical displacements are unde- 
termined for 11 of these faults. For 
the remaining faults, vertical displace- 
ments are relatively small-scale, ranging 
from 0.6 m (2 ft) to about 180 m 
(600 ft). In the Oroville area, dis- 
placements range from 4.3 m (14 ft) on 
the Swain Ravine Lineament fault at the 
Orange Road trench site, about 17 km 
(11 mi) south of Oroville, to 46 m 
(150 ft) on a fault about, 27 km (17 mi) 
north of Oroville. 



Swain Ravine Lineament Fault Zone : 
Cenozoic fault movement in Basement 
Series rocks within the Oroville study 
area is recognized only in the Swain 
Ravine Lineament fault zone. This dis- 
placement was first noted after the 
August 1975, movement of the Cleveland 
Hill Fault. A striking feature of this 
movement is that it started in a linea- 
ment "gap" where there are no lineament 
features and little topographic evidence 
of faulting (Figure 50). 

The Cleveland Hill Fault can be traced 
on the ground and in exploration trenches 
to within 2.3 km (1.4 mi) of the Bidwell 
Canyon Saddle Dam. The Swain Ravine 
Lineament fault zone continues to the 
north and extends into Bidwell Canyon. 
Hypocenters of aftershocks extend 10 km 
(6 mi) north of the surface faulting and 
pass beneath Oroville Dam at a depth of 
about 5 km (3 mi) (Lahr and others, 1976), 

Surface investigations, trenching, and 
geophysical investigation failed to 



/ 



Surface evidence of Cenozoic faulting is 
rare. This is attributed to small dis- 
placements at sufficiently long recur- 
rence intervals to allow erosion to re- 
move evidence of fault movements. More- 
over, much of the study area is not 
covered by Cenozoic deposits, making 
determination of younger fault movements 
difficult. 

Within the area of detailed mapping, the 
Swain Ravine, Prairie Creek and Paynes 
Peak Lineaments were determined to be 
Mesozoic fault zones. North of the de- 
tail map area several lineaments are 
apparant on high-altitude infrared and 
.low-altitude black-and-white photographs. 
For this study the major northern linea- 
ments are informally named the Chico, 
Soda Springs, Web Hollow and Paradise- 
Magalia Lineaments. These lineaments 
were field checked to determine if they 
are faults. Figure 16 shows these linea- 
ments and associated faults. 



•/ 



r-f 





Figure 50. Aerial northwest view of the 
Cleveland Hill Fault along the western 
side of Cleveland Hill. Note that 
topographic expression for the fault 
is lacking. 



71 



reveal faulting beyond an olive grove 
south of Mt. Ida Road, about 2.3 km 
(1.4 mi) south of Bidwell Canyon Saddle 
Dam. However, the Swain Ravine Lineament 
fault zone appears to go northward into 
Bidwell Canyon. Faulting in the west 
end of Bidwell Canyon Saddle Dam foun- 
dation and a fault exposed in a water 
tunnel about 1 km (0.6 mi) west of the 
dam may be parts of a complex system of 
faults in the Swain Ravine Lineament 
fault zone. It can be conjectured that 
the fault system consists of a band of 
discontinuous, relatively short, small 
faults along which displacement could 
jump from fault to fault within the main 
zone. This kind of fault discontinuity 
could explain the inability to trace the 
Cleveland Hill Fault continuously far- 
ther north to link up with the faulting 
at Bidwell Canyon Saddle Dam. The west- 
ward tilt indicated by leveling surveys 
on the saddle dam further suggests that 
the fault system goes into the Bidwell 
Canyon arm of the reservoir. 

North of Lake Oroville a lineament con- 
tinues on through Canyon Creek, along 
the projected northerly trend of the 
Swain Ravine Lineament fault zone 
(Figure 16). However, geologic investi- 
gation failed to reveal faulting along 
the lineament. The markedly straight. 
North Fork of Lake Oroville also is sug- 
gestive of faulting, but no faulting 
could be found along the north side of 
the lake . Because the Swain Ravine 
Lineament fault zone is a strong feature 
of some size, it seems the zone should 
continue through the reservoir. However, 
field investigations did not prove this 
to be true. Therefore it is assumed the 
fault zone terminates in the reservoir. 

The fault movement along the Cleveland 
Hill Fault in the Swain Ravine Lineament 
system prompted a number of exploration 
trenches. These were mostly along the 
faulted portion, but some were on other 
parts of the Swain Ravine Lineament and 
also on the nearby Paynes Peak and 
Prairie Creek Lineaments. Trenching was 
done by Department of Water Resouces, 
U. S. Army Corps of Engineers, and by 



72 



Woodward-Clyde Consultants for'various 
clients. Purpose of the trenching was 
to investigate the nature of the faultin 
in the underlying bedrock and to deter- 
mine if previous Quaternary fault move- 
ment could be detected in soil profile 
overlying the bedrock. Although expo- 
sures of faults in the bedrock were 
usually obvious and easy to interpret, 
the exposures of the overlying soil pro- 
file were not so clear cut. Interpreta- 
tion of earlier^ fault displacements seen 
in the soil profiles were not necessaril 
concurred with by all that saw the expo- 
sures in the various trenches. For the 
purposes of this study we have accepted, 
without necessarily endorsing, the inter 
pretations made in trenches by others. 

For displacements of small magnitude, 
such as along the Cleveland Hill Fault, 
displacement of the soil profile in 
trenches is difficult to detect. We are 
not aware of any trench where such dis- 
placement from the August 1, 1975, earth 
quake could be detected, even though i 
the trenches were excavated across the , 
ground cracking. The cracking could 
usually be seen at least part way throug 
the soil profile and sometimes, clear 
through the usually shallow soil profile 
to bedrock. Although offset of features 
interpreted to be caused by earlier 
fault movements could be seen in one 
trench, other trenches nearby on the 
same fault might not show offsets, or 
show displacement of a different magni- 
tude. In short, trenching does not 
detect all of the previous fault move- 
ments with great clarity — instead very 
subtle features, subject to varying 
interpretation, are usually revealed. 
Earlier small fault movements can easily 
pass undetected. 

Of the 17 Department of Water Resources 
trenches, 15 were on the Swain Ravine 
Lineament fault zone. Eight of those 
trenches (A, B, D, 5, 8, 12, 13 and 16) 
were excavated across Cleveland Hill 
Fault ruptures, and each of these ex- 
posed a fault in bedrock beneath the 
ground rupture. Two other trenches (7A 
and 17) exposed bedrock faults, and the 



remaining five trenches on the Swain 
Ravine Lineament fault zone did not ex- 
pose any faults. 

Only one Department of Water Resources 
trench exposed convincing evidence of 
Quaternary movement on the Cleveland 
Hill Fault prior to the 1975 event. 
Trench DWR 8, across the Cleveland Hill 
faulting, exposed a contact between two 
alluvial units, at a depth of 3.3 m 
(10.8 ft), with about 30 mm (1.2 in) of 
apparent down-to-the-west vertical off- 
set on the fault. Soil features near 
the ground surface were not displaced, 
indicating movement occurred prior to 
the 1975 Oroville event. In their 
Cleveland Hill No. 1 trench, Woodward- 
Clyde Consultants report 46 cm (18 in) 
of apparent vertical offset which they 
interpret as having occurred in at least 
three, and perhaps more, separate events. 
This suggests that a 15 cm (6 in) move- 
ment would be the maximum displacement 
expected. 

Forty-two exploration trenches were ex- 
cavated by various investigators on the 
Swain Ravine Lineament fault zone between 
Bangor and Lake Oroville (Figure 47 
and Table 1) . Many of these trenches 
exposed west-dipping faults showing evi- 
dence of multiple, small-scale (less 
than 1 m) , normal Cenozoic movements. 
An exception to this occurs in the U. S. 
Bureau of Reclamation Orange Road 
trenches, where some faults were found 
to dip east, and one of these showed a 
4.3 m (14 ft) normal offset. Five 
trenches on the lineament just north of 
the Yuba River, by the Corps of Engineers 
(1977) exposed Mesozoic faults with no 
Cenozoic movement. 

Ages of movement on the Cleveland Hill 
Fault and at other places along the 
Swain Ravine Lineament fault zone are 
based on displaced soil horizons. The 
most prominent soil marker is a buried 
paleosol. Gene Begg (person, commun., 
1977) and Roy Shlemon (person, commun., 
1977) estimate the paleosol to be at 
least 100,000 years old, and soils 
overlying it are 50,000 to 70,000 years 



old. Similar paleosols in the Sierran 
foothills have also been estimated to 
be at least 100,000 years old by Swan 
and Hanson (1977, 1978). The U.S. Geo- 
logical Survey (1978, p. 43) has stated 
that soils overlying the paleosol may 
be younger - 10,000 to 25,000 years 
old - and that a fault which displaces 
the paleosol and not the overlying soil 
should be considered to be a minimum of 
10,000 years old. In summary, the part 
of the Cenozoic record available, the 
soil cover, indicates previous small 
fault movements within the last 10,000 
to 100,000 years. 

Where older soil horizons are offset 
more than successively younger ones, it 
is interpreted that the fault has moved 
several times during development of the 
soil profile. The amount of displace- 
ments seen, and the lack of surface fault 
features along the Cleveland Hill Fault 
suggest that small movements at relative- 
ly long recurrence intervals have occur- 
red over the past 100,000 years. 

An east-west channel deposit of late 
Cenozoic gravel crosses the Pajmes Peak 
and Swain Ravine Lineaments northwest of 
Bangor (Plate 1) , yet shows no field 
evidence of fault displacement. More- 
over, a small diorite plug across the 
Swain Ravine Lineament in the same area 
is not displaced, indicating it has not 
been offset since emplacement, about 
130 million years ago. 

The fact that fault displacements are 
seen in soils exposed by trenching, yet 
no field evidence can be seen for larger 
displacements of the late Cenozoic 
gravel channel deposits or pluton, sug- 
gests long-term cumulative displacement 
is too small to detect by normal field 
techniques. The apparently unbroken 
pluton in the Swain Ravine Lineament 
fault zone suggests undetectable dis- 
placement for about 130 million years. 
The same pluton relationship is seen 
on the Prairie Creek Lineament fault 
zone which is apparently truncated by 
the Rocklin pluton. 



73 



The evidence indicates a pattern of 
activity during the last 100,000 years 
of small, infrequent, vertical fault 
movements, along the Swain Ravine 
Lineament fault zone. Presumably, such 
displacements were produced by earth- 
quakes of about the same magnitude as the 
1975 Oroville earthquake. The average 
slip rate during the last 100,000 years 
may be faster than during most of the 
last 130 million years, otherwise the 
older rocks would be noticeably dis- 
placed. For example, cumulative fault 
displacement of 0.46 m (1.5 ft) in soil 
was interpreted from relationships in 
the Cleveland Hill No. 1 trench by 
Woodward-Clyde Consultants. This gives 
a maximum slip rate in the Cleveland 
Hill area of 0.46 m (1.5 ft) per 
100,000 years. Assuming the late 
Cenozoic gravels overlying the Swain 
Ravine Lineament fault zone are early 
Quaternary, or about two-million years 
old, then they should be offset about 
9 m (30 ft) if the same slip rate pre- 
vailed — a displacement large enough to 
be noticed in the field. Applying the 
same slip rate to the 130 million year- 
old pluton would produce an offset of 
about 600 m (1,960 ft). 

The Swain Ravine Lineament fault is 
the only one along which evidence of 
Quaternary fault displacements was 
found. However, these kind of fault 
displacements are subtle and difficult 
to detect, so it is possible, though 
not proven, that similar levels of fault 
activity have taken place along the 
Prairie Creek and Paynes Peak Lineament 
faults. South of the study area, the 
Swain Ravine Lineament fault merges with 
the Prairie Creek Lineament fault to 
form one system (Figure 47) . Trenches 
just south of where the two lineaments 
merge revealed what is interpreted to 
be Quaternary fault movements (Alt and 
others, 1977). The persistent evidence 
of Quaternary movements along the Swain 
Ravine and the merged Swain Ravine- 
Prairie Creek Lineament fault zones 
suggests this may be a preferred avenue 
along which small Quaternary fault move- 
ments occur. 



74 



Prairie Creek Lineament Fault Zone ; 
Several areas of localized ground crack- 
ing occurred during the Oroville earth- 
quake and form a crude alingment that ex- 
tends northwestward from the north end 
of the Prairie Creek Lineament for 10 km 
(6 mi) to Oroville. Further to the 
northwest, Lovejoy basalt on the 
Campbell and Sorensen Hills is discon- 
tinuous with basalt on North and South 
Table Mountains along a projection of 
the lineament. This trend, projected 
still further to the northwest, could 
coincide with faulting in Tuscan 
Formation along the Chico monocline. 
Evidence is lacking, however, these 
features may result from Cenozoic move- 
ment along an underlying Mesozoic bed- 
rock fault. 

Even though no evidence of Quaternary 
fault activity was seen along the 
Prairie Creek Lineament fault zone, the 
fact that Quaternary faulting has occur- 
red just south of where it merges with 
the Swain Ravine Lineament fault zone 
suggests Quaternary activity has, or 
could, take place along the Prairie 
Creek zone. This possibility is rein- 
forced by the occurrence of some small 
earthquakes along the lineament, and the 
yet unexplained system of cracks, the 
"Palermo crack zone", that developed 
along the northwest projection of the 
Prairie Creek Lineament fault zone dur- 
ing the Oroville earthquake. For these 
reasons, the Prairie Creek Lineament 
fault zone should be regarded as capa- 
ble of the same kind of activity seen 
along the Swain Ravine Lineament fault 
zone. If the Prairie Creek Lineament 
fault zone does continue northward as 
an old Mesozoic fault zone more or less 
concealed by younger rocks of the Super- 
jacent Series, it would be the longest 
of the lineament fault zones. 

Paynes Peak Lineament Fault Zone ; Of 
the three lineament fault zones in the 
Oroville area, Paynes Peak is the only 
one along which no evidence or suggestion 
of Quaternary activity is seen. Conse- 
quently, it is assumed fault activity is 



not as likely to occur along the Paynes 
Peak Lineament as along the other two 
zones. 



tion exposed on South and North Table 
Mountains, about 6 km (3.8 mi) to the 
northeast along projection of strike. 



Thermalito Powerplant Foundation Faults ; 
During construction, faults were exposed 
in the foundation excavation for 
Thermalito Powerplant. The zone consists 
of several interlaced faults striking 
from N30-44E, dipping steeply from 70NW 
to 80SE, and offsetting Miocene age 
Love joy Formation. The fault does not 
offset overlying gravels of late Cenozoic 
age. Apparent offsets are both normal 
and reverse, and apparent cumulative 
displacement across the zone is as much 
as 12 m (40 ft), with the southeast 
side moved downward relative to the 
northwest side. No evidence for this 
fault was observed in Love joy Forma- 



Chico Lineament ; The Chico Lineament 
coincides with the upper hinge of the 
Chico monocline (Figure 51) . A zone of 
tensional faults, generally small, lies 
along this hinge line in the area of 
the lineament. These faults cut Upper 
Pliocene Tuscan Formation. 

A comprehensive study of faulting along 
the Chico monocline was done by Burnett 
(1965) . Additional studies dealing 
with faults in Tuscan rocks are by 
Creely (1965) and Burnett and others 
(1969); field investigations were also 
carried out for this report. 






■•^#''^,:.'?5f-'?^-!^"-, •• 



•^rgfm^ 






M^ _*.v%5«- 



^f^. >•* ■ 






'■»y". -rf-ji^tnt i*ii III 



^ 

■m^!^ 



Figure 51. Aerial northwest view of the Chico monocline that is locally developed in 
Tuscan Formation. Note pattern of northwest-trending linear fractures defined by 
dark vegetation lines. Location is northeast of Chico. 



75 




Figure 52. Normal fault (attitude: N3W, 62NE) in Tuscan Formation exhibiting 30 cm 
(1 ft) of down-to-the-east displacement. Location is a roadcut exposure along Clark 
Road approximately 6 km (3-7 mi) south of Paradise. 




~~^»^ft^^. 



Figure 53. Normal fault (attitude N6W, 65SW) in Tuscan Formation exhibiting 70 
cm (2.2 ft) of down-to-the west displacement. Location is a roadcut exposure 
along Clark Road about 6 km (3.7 mi) south of Paradise. 



76 



Faults in Tuscan rocks trend north- 
northwest and dip either east or west . 
These are normal faults which usually 
have little vertical displacement 
(Figures 52 and 53) . Displacements are 
generally less than 1 m (3 ft); however, 
displacements of 18 m (60 ft) (Burnett, 
1965) and 31 m (100 ft) (Dave Harwood, 
person, commun., 1978) have been reported. 

Soda Springs Lineament ; The Soda 
Springs Lineament is about 16 km (10 mi) 
long and is composed of linear topogra- 
phic features and an area of springs. 
Field investigation along the lineament 
produced no evidence of faulting. Expo- 
sures of Tuscan rock along Ditch Creek 
(Sect. 9, T26N, R3E) are continuous 
across the lineament and provide good 
evidence against faulting. The springs 
are interpreted to be interflow seeps 
in the underlying, horizontally-bedded 
Tuscan Formation. 

Web Hollow Lineament : The Web Hollow 
Lineament is a strong north-south trend- 
ing feature at least 31 km (19 mi) long 
and has been suggested to be a fault 
(Quintin Aune, unpub. data; Alt and 
others, 1977). A field check of fault 
features described by Aune was made by 
Department of Water Resources geologists 
prior to this study; none of these fea- 
tures could be attributed to faulting 
(Mark McQuilkin, person, commun., 1977). 
During this study the entire Web Hollow 
Lineament was field checked and de- 
termined not to be a fault. 

A zone of very strong lineaments, paral- 
lel to, and just east of the northern 
end of the Web Hollow Lineament, can be 
seen in the Tuscan Formation on low- 
altitude photographs (Figure 54). These 
cut across the topography in a northwest 
trend. A check of these lineaments on 



the ground failed to substantiate the 
presence of faulting, however, photo- 
graphic features so strongly suggest 
fault control that they are interpreted 
to be faults. The lack of surficial 
fault features in the Tuscan Formation 
suggests these are Upper Pliocene or 
Pleistocene in age. 

Paradise-Magalia Lineament ; The 
Paradise-Magalia Lineament is about 
43 km (27 mi) long and trends north from 
Oroville to Magalia (Figure 16). Linear 
elements include a section of the 
Feather River at Oroville, the east edge 
of South Table Mountain, a drainage on 
North Table Mountain and a linear ridge 
on the west side of Magalia Reservoir. 




whl^^m '^tw" 







-^ 'IRON 
' Nipyi^TftlN 







Figure 5^. Vertical aerial view of fracture 
zone developed in Tuscan Formation near 
Iron Mountain on Deer Creek. Fractures 
appear as dark vegetation lines which 
transect topography. 



77 





o 
o 

oca 

OJO 



o 
oi 

in 



Rock 



VS 



ft^v^\^^'^■^ 






\:>\\^ 



Figure 55. Map with cross section oriented perpendicularly to suspected fault 
through Magi ia Reservoir. 



78 



This lineament was investigated through- 
out most of its length on the ground. 
Cenozoic fault features could only be 
found where it contacts Love joy Forma- 
tion on the south side of North Table 
Mountain. Here a small fault, less 
than 0.5 km (0.3 mi) long, cuts Lower 
Pliocene basalt. Another fault of the 
same size and attitude occurs just to 
the west — off the lineament. Poor 
exposures preclude determining a sense 
of displacement of either fault, and 
it is possible that rather than faults 
they are unusually prominent cooling 
joints . 

Old mining reports indicate faults occur 
in several mines along the Paradise- 
Magalia Lineament northwest of Magalia. 
A report by the California Department 
of Natural Resources (1930) cites dis- 
placements of up to 70 m (225 ft) in 
Tertiary channel deposits below Tuscan 
Formation. These faults do not cut 
overlying Tuscan Formation and must be 
pre-Pleistocene in age. Faults in the 
mines as well as a Mesozoic fault at 
Magalia Reservoir trend more westerly 
than the Paradise-Magalia Lineament and 
may not be related to it. 

Alt and others (1977, pp. 14-21) state 
that a prominent Cenozoic fault scarp 
occurs in Tuscan rocks along the Paradise- 
Magalia Lineament on the west side of 
Magalia Reservoir. Detailed mapping of 
the Tuscan-basement rock contact during 
this investigation showed no evidence 
of faulting. A cross-section based on 
this mapping, drawn perpendicular to 
this scarp, shows no displacement of the 
base of the Tuscan Formation (Figure 55) . 

Summary ; Cenozoic fault movements have 
occurred in the Oroville area on both 
pre-existing Mesozoic bedrock faults, 
such as the Cleveland Hill Fault, and on 
faults in Cenozoic rocks, such as those 
in the Tuscan and Love joy Formations. 
It is possible that the faults cutting 
Cenozoic rocks may also be along Mesozoic 
faults in bedrock concealed by the over- 
lying younger formations. 

Potential for future earthquakes and 
ground rupture in the Oroville area is 



considered to be greatest on the Swain 
Ravine Lineament fault zone, to the north 
and the south of that portion of fault 
that ruptured in the 1975 event. 

The fault responsible for the magni- 
tude 5.7 earthquake that occurred north- 
east of Chico, California, in 1940, was 
not found. 

Mesozoic Folds 

Smartville ophiolite in the Bangor 
quadrangle and melange in the Cherokee 
and Berry Creek quandrangles are iso- 
clinally folded. Exposures of relict 
pillows provide structural control for 
tops of section. Near Bangor, pillowed 
sequences dip east and are inverted; 
north of this location, pillowed rocks 
are right-side-up and dip east. Tops 
of these sections are oriented in oppo- 
site directions and suggest that large- 
scale folding deforms the ophiolite. 
Large-scale folding is locally confirmed 
by concordant small-scale folds inter- 
preted to be parasitic deformation. 

Mesozoic folds trend north-northwest 
and plunge south at angles less than 
35 degrees. Plunge directions conform 
with orientations described in several 
geologic studies of the foothills 
(Bateman and others, 1963; Clark, 1964, 
1976; Creely, 1965; Hietanen, 1973a; 
Duf field and Sharp, 1975). Axial planes 
of these folds, oriented subconcordant 
to metamorphic foliation, dip steeply 
east or are vertical. Localized folds 
occur in metamorphosed country rock 
adjacent to Sierran plutons (Compton, 
1955; Clark, 1964; Hietanen, 1973a). 

Cenozoic Folding 

Cenozoic folding does not affect study 
area rocks but is described at several 
locations in the nearby area. At Tuscan 
Springs, approximately 8 km (5 mi) north- 
east of Red Bluff, the Tuscan and Chico 
Formations are gently deformed into 
open folds (Anderson, 1933). Addition- 
ally, Hudson (1951, 1955) has described 
folded Superjacent Series rocks in the 
Mount Lincoln-Castle Peak area. 



79 



The Chico monocline, as described by 
Bryan (1923), Anderson (1933), Burnett 
(1965) and Burnett and others (1969), 
trends approximately N30W and forms the 
24 km (15 mi) straight eastern boundary 
of the Great Valley between Chico and 
Red Bluff (Figure 51) . The monocline 
steepens Tuscan Formation dips from 
2 degrees east of the fold, to 8 to 
10 degrees within the hinge area 
(Burnett and others, 1969). This is the 
best developed Cenozoic fold in the 
immediate area. 

Most Cenozoic folding occurred prior 
to deposition of the Pleistocene (?) 
Red Bluff Formation as these unconform- 
able gravels are not deformed (Anderson, 
1933). The Red Bluff Formation and late 



Cenozoic gravels in the study area are 
incised by actively-downcutting drainages 
which indicate the foothill region is 
experiencing uplift; some deformation 
must be associated with this uplift. 

SUMMARY OF GEOLOGIC HISTORY 

The geologic history and tectonic 
movement in the Oroville area is summar- 
ized in chronologic order of occurrence. 
The timing of events is substantiated 
by field data or referenced from inves- 
tigators who did particular studies con- 
cerning aspects of the geologic history. 
Absolute dates for events are used if 
available; relative time, based upon 
fossils, is used where radiometric dates 
are lacking. 



Table 2 
Summary of Geologic Events 



Event 



Date 



Cenozoic Time 



Oroville earthquake - small 
fault movement near Cleveland 
Hill along Swain Ravine 
Lineament fault zone 



1975 



2. Development of "A" and "B" 
soil profiles. Some small 
fault movements along Swain 
Ravine and merged Swain 
Ravine-Prairie Creek fault 
zones 



10,000-25,000 years before present 
(uses, 1978) 

50,000-70,000 years before present 
(Gene Begg and Roy Shlemon, person, 
commun. , 1978) 



Development of "Paleo B" soil 
profiles; small fault 
movements. 



4. Alluvium deposited 

5. Development of late Cenozoic 
faulting in the foothill belt 

6. Older gravels deposited 



100,000 years before present 
(Gene Begg and Roy Shlemon, 
person, commun., 1978) 
140,000 years old (USGS, 1978) 

Pleistocene - Recent 

Post-Pliocene (Alt and 
others, 1977) 

Pliocene-Pleistocene (?) 



80 



Event 



Date 



10. 



Tuscan Formation deposition 
began 

Sierran uplift and westerly 
tilting began 



Love joy Formation, basalt 
flows from eastern source 

Oroville tuffs (Mehrten 
Formation-?) deposited 



Pliocene, 3.3 million years 
(Lydon, 1968) 

Pliocene, 4-10 million years 
(Christensen, 1966; Wright, 1976; 
Hay, 1976) 

22.2 to 23.8 million years 
(Dalrymple, 1964) 

23.8 million years 
(Dalrymple, 1964) 



11. 



Auriferous gravels deposited 



12. lone Formation deposited 



Upper Eocene - Lower Oligocene 
(Durrel, 1966) 

Middle Eocene (Creely, 1965) to 
early Oligocene 



13. Early Sierran uplifts 



Cretaceous-early Tertiary 



Mesozoic Time 



Chico Formation deposited 



Upper Cretaceous (Taff and others, 
1940; Creely, 1965) 



Plutonism (Yosemite intrusive 
epoch) - intrusion is respon- 
sible for thermally metamor- 
phosing country rocks in the 
study area 



126-138 million years (Gromme and 
others, 1970; Evemden and 
Kistler, 1970) 



Mesozoic faults (Foothills 
Fault System) - formed by 
collision of arc and ophio- 
lite with continent 



Middle Kimmeridgian to late 
Tithonian 



4. Smartville ophiolite formed 

5. Monte de Oro Formation 
deposited 

6. Arc Rocks extruded 



7. Melange formed 



Oxfordian to Kimmeridgian 

Late Oxfordian to early 
Kimmeridgian (Imlay, 1961) 

Oxfordian (?) to early 
Kimmeridgian (Creely, 1965) 

Middle to Upper Jurassic 
(Kimmeridgian, Bob Treet, 
person, commun., 1978) 



81 



i 



The above tabulation of events indi- 
cates tectonic activity is spasmodic. 
Major activity was in Mesozoic time. 
The Sierran uplift is a major regional 
event, but the main influence on rocks 
of the western foothill belt around 
Oroville is to tilt Cretaceous and 
younger rocks gently to the west. Vol- 
canic activity about 22-24 million years 
ago deposited basalt flows (Lovejoy 
Formation) , lahars and water lain vol- 
caniclastic rocks of the Oroville tuff. 

In early Pleistocene time, fault move- 
ments began again in response to the 
most recent Sierran uplift, mostly along 
the old Mesozoic faults. Some younger 
rocks, mainly Tuscan Formation, also are 
faulted. In the Oroville study area, 
Cenozoic fault displacements are general- 
ly small. Elsewhere, Cenozoic displace- 
ments up to 197+ m (600 ft) are reported 
(Alt and others, 1977). 

The last 100,000 years has been a 
period during which small fault displace- 
ments on the order of centimetres occur- 
red at infrequent intervals along the 
older Mesozoic fault zones. 

CAUSES OF THE OROVILLE EARTHQUAKE 

Geologic investigations reveal that the 
Oroville earthquake is not unique in 
the seismic history of the northwestern 
Sierra Nevada foothill belt. The 1940 
magnitude 5.7 earthquake north of what 
is now Lake Oroville demonstrated that 
the region is capable of generating 
moderate-magnitude earthquakes. The 
trenching and other exploratory work 
revealed evidence for small displace- 
ments in soil during the last 
100,000 years or less, depending upon 
which interpretation is accepted for 
the age of soils displaced by faulting. 
Presumably these small fault movements 
also were accompanied by earthquakes 
similar to the 1940 and 1975 magnitude 
5.7 earthquakes that occurred in the 
Oroville region. 

The geologic investigations of the 
Oroville area revealed old fault zones 



in bedrock. These old fault zones formed 
during Mesozoic time — subduction at a 
plate boundary is postulated as the geo- 
logic model for genesis of the faults. 
Actually, the geologic model assumed for 
formation of the old fault system is not 
as critical as the fact that these older 
fault zones formed in a different tec- 
tonic regime than exists today. The im- 
print of earlier tectonism has left zones 
of weakness along which fault movements 
caused by the current Cenozoic tensional 
regime tend to occur. 



t 



The present tectonic pattern clearly 
seems to be one of general east-west 
extension. The behavior of the 
Cleveland Hill Fault, the opening of the 
fault cracks with time and the geodetic 
work in the Oroville area, all suggest 
normal fault movements resulting from 
east-west tension. A fault plane solu- 
tion of the August 1975, Oroville earth- 
quake series indicates normal dip-slip 
movement (Langston and Butler, 1976; 
Lester and others 1975), also indicating 
east-west tension. 

Recent leveling work done by Bennett and 
others (1977) suggests the Sierran block 
is still undergoing uplift. This uplift 
is postulated to be the cause of tension 
in the western foothills. Bennett (1978) 
continued examination of leveling data 
and was able to demonstrate noticeable 
changes in elevation across both the 
Bear Movintains and Melones Fault zones, 
again indicating crustal movements pre- 
fer to occur along the old fault zones. 

In summary, for at least the past 4 mil- 
lion years the foothill region probably 
has been in east-west tension. It can 
be speculated that this tension reduced 
frictional stresses along north-trending 
older fault systems, allowing small grav- 
itational adjustments and fault movements. 
These movements probably were accompanied 
by earthquakes comparable to the 1940 and 
1975 magnitude 5.7 earthquakes that oc- 
curred in the Oroville area. These fault 
movements commonly occur along the small- 
er individual faults within the older 
fault zone complexes. Possibly not all 



82 



movements occur along existing faults, 
but displacements revealed thus far sug- 
gest that most do. 

Reservoir-Induced Seismicity 

Reservoir-induced seismicity has been 
greatly studied in the last few decades. 
Several researchers (for example, Gupta 
and others, 1973; Rothe", 1973; Bozovic, 
1974) note the following characteristics 
of reservoir induced seismicity: 

1. Earthquake activity begins soon after 
initial impoundment. 

2. Large numbers of foreshocks occur 
over an extended time period before 
the main shock. 

3. Time versus frequency plots of fore- 
shocks and aftershocks differ from 
patterns of tectonically induced 
seismicity. 

4. Reservoir-induced earthquakes are 
shallow focus. 

5. Proximity of reservoir-induced seis- 
micity to the triggering dam or res- 
ervoir is usual. 

6. Relatively high "b" values for fore- 
shocks of reservoir-induced seismic 
events as established by the equation 
log N = a-bM where "b", the slope of 
the curve, is related to the propor- 
tion of large-to-small earthquakes, 
"N" is the cumulative frequency of 
magnitude "M" earthquakes, and "a" 

is a constant determined by area 
where data were gathered, time dura- 
tion and areal seismicity. 

7. A high ratio of the strongest magni- 
tude aftershock to the magnitude of 
the main seismic event. 

The most consistent characteristic of 
reservoir- induced seismic activity is 
the onset of the earthquakes soon after 
initial impoundment begins (Rothe, 1973). 
Some examples (after Packer and others, 
1977, Appendix A) are listed below. 



Reservoir 

Koyna 

Hsinfengkiang 

Hendrik Verwoed 

Boulder 

Contra 

Grandval 



Elapsed time from 
beginning impoundment 
to start of seismicity 

Immediately 

1 mo. 

6 mo. 

9 mo. 
10 mo. 
15 mo. 



In comparison, seismic activity at 
Oroville occurred more than eight years 
after water impoundment began. 

A second characteristic of reservoir- 
induced seismicity is that a large num- 
ber of foreshocks occur over a longer 
period of time than would be expected 
for tectonically-induced seismic events 
(Rothe, 1973; Bozovic, 1974). For ex- 
ample Kremasta Reservoir had 17 fore- 
shocks in 30 days. Koyna Reservoir 
had 90 foreshocks in 19 days, and 
Kariba Reservoir had 20 foreshocks in 
one day (Gupta and others, 1973). The J 
frequency of foreshocks at Oroville, 
21 events in 30 days prior to the main 
shock, is similar to the Kremasta data. 

Mogi (1963) discovered that the plot 
of time versus frequency for reservoir- 
induced foreshocks and aftershocks dif- 
fered from the pattern for tectonically- 
induced earthquakes. The reservoir- 
induced seismic pattern (Type II) 
includes a greater foreshock buildup and 
a longer period of aftershocks. 
Figure 56 shows Mogi's "Type II" 
(reservoir-induced) seismic pattern 
compared with the pattern of the Oroville 
earthquake series . 

Shallow focal depths for reservoir- 
induced earthquakes are also a common 
characteristic (Bozovic, 1974). Some 
examples of this characteristic are 
Monteynard Reservoir, 0.0 km (0.0 mi); 
Hendrik Verwoed Reservoir, 6 km 
(3.7 mi); Boulder Reservoir, 1.5 to 
9 km (0.9 to 5.6 mi). The main shock 
of the Oroville series had a depth of 
8.8 km (5.5 mi) and aftershock depths 
from 1.3 to 10.4 km (0.8 to 6.5 mi) 
(Akers and others, 1977). 



83 



FORESHOCKS 



TIME 
MOGl'S "type n" PATTERN 



11 



F'M'4 'm'j'j 

1975 
Time (6/28/75 to 4/30/76) 
PATTERN FOR THE OROVILLE SERIES 



u 



Figure 56. Comparison of foreshock-af tershock patterns for the Oroville earthquake 
and Mogi's "Type M" (reservoi r- i nduced) earthquakes. 



























A 




r\ 


\/\ 


JV 


A 


A 


n 




r^ 


r" 












\ 


/ 


\ 


~\ 


















1 




Vj 




j 














4 








1 














1 
1 
1 

1 









1967 19(8 



1976 1977 1979 



Figure 57. Water level history of Lake Oroville from initial filling to September 
1978 



84 



In most reservoir-induced seismic 
events the foci of the earthquakes are 
very close to, or directly under, the 
reservoir or dam. This characteristic 
is noted for earthquakes at Boulder, 
Monteynard, Grandval, Oued Fodda, 
Kariba, Kremasta, Marathon and Koyna 
reservoirs. The most distant earthquake 
was within 10 km (6.2 mi) of the Koyna 
Reservoir dam (Gupta and others, 1973). 
The main shock of the Oroville earth- 
quake series was more distant, approxi- 
mately 12 km (7.5 mi) from Oroville Dam. 

Relatively high "b" values for fore- 
shocks and aftershocks of reservoir- 

induced events are common (Gupta and 
others, 1973). A comparison of "b" 
values for accepted cases of reservoir- 
induced seismicity with those of the 
Oroville earthquake shows a significant 
difference. Note "b" value comparisons 
below (after Morison and others, 1976). 





"b" Values 




Reservoir 


Foreshocks 


Aftershocks 


Kariba 


1.1 


1.0 


Kremasta 


1.4 


1.1 


Koyna 


1.9 


1.3 


Oroville 


0.37 


0.61 



of 100 m (328 ft) in February 1968, and 
has never been lowered beneath that 
level. A plot of post-construction res- 
ervoir elevations with the date of the 
August 1, 1975, earthquake is shown in 
Figure 57. The 1975 change from low 
water in January to high water in late 
May represents an increase of 29 percent 
in 5 months; initial reservoir filling 
during the same time period in 1968 
generated an increase in storage of 
44 percent with no resulting seismicity. 

In comparing data from the Oroville 
earthquake with reported data from known 
reservoir-induced earthquakes, the 
Oroville event has some, but not all, of 
the characteristics attributed to 
reservoir-induced earthquakes. The long 
elapsed time between reservoir filling 
and the Oroville earthquake is a signi- 
ficant departure from what has occurred 
in generally accepted cases of reservoir- 
induced seismicity. The epicentral dis- 
tance from the reservoir is slightly 
large and the "b" values are signifi- 
cantly too small. A comparison of ac- 
cepted characteristics of reservoir- 
induced earthquakes with those of the 
Oroville earthquake, does not show whe- 
ther or not Lake Oroville caused the 
earthquake series. 



Another feature common to reservoir- 
induced seismicity is a high ratio, 
typically 0.8 to 0.9 (Gupta and others, 
1973), of the strongest aftershock mag- 
nitude to the magnitude of the main 
shock. The ratio for the Oroville 
series is 0.91. 

A relationship has been postulated 
between rate of reservoir loading, 
water depth and induced seismicity 
(Rothe, 1969). Induced earthquakes 
seem to be more commonly associated 
with reservoirs whose water depths are 
100 m (328 feet) or greater and during 
periods of rapid loading (generally 
the first reservoir filling) ; large 
water level fluctuations also induce 
seismicity (Carder, 1945). Lake 
Oroville^ approximately 200 m (656 ft) 
deep at maximum pool, reached a depth 



The picture is further clouded by the 
background seismicity which indicates an 
earthquake comparable to the 1975 
Oroville earthquake occurred before the 
reservoir existed. Small fault displace- 
ments in soil profiles exposed by trench- 
ing suggest Oroville type events have 
been occurring for the past 100,000 
years. In short, the Oroville earth- 
quake is compatible with the regional 
pattern of seismicity and, therefore, 
need not be related to Lake Oroville 
at all. 

The mechanisms customarily used to 
explain reservoir-induced earthquakes 
are (1) increase in stress caused by 
weight of water in the reservoir and 
(2) increase in pore pressure resulting 
from the increase in hydrostatic head 
imposed by the reservoir. The increase 



85 



in pore pressure mechanism is more favor- 
ed. Both theories assume stress condi- 
tions in the hypocentral area are in 
such delicate balance that only small 
incremental changes in stress will trig- 
ger an earthquake. 

The epicenter of the August 1, 1975, 
Oroville earthquake was 12 km (7.5 mi) 
from Lake Oroville. Although no computa- 
tions were made by the Department to 
estimate how much Lake Oroville changes 
stress in the hypocentral area, at that 
distance and depth change in stress 
should be very small. 

A reconnaissance survey of springs 
and wells in the area indicates ground 
water levels in the foothill area to the 
east and north of the epicentral area 
are generally higher than water in Lake 
Oroville. Assuming hydrostatic pressures 
in the hypocentral area are controlled 
by ground water levels, the addition of 
Lake Oroville in such a hydrostatic 
regime probably would have no affect on 
either pore pressure or degree of 
saturation. 

The evidence available does not indi- 
cate a causal relationship between Lake 
Oroville and the earthquake, but the 
possibility cannot be eliminated conclu- 
sively at this time. 

POTENTIAL HAZARDS TO STATE 
WATER FACILITIES 

Because the August 1, 1975, Oroville 
earthquake probably relieved much of the 
regional strain, it seems unlikely that 
a similar event will occur in the same 
place in the near future. Therefore, the 
next earthquake of comparable magnitude 
probably would occur north or south of 
the 1975 earthquake and its zone of 
aftershocks. Despite the improbability 
of another local earthquake, estimates 
of hazards to facilities are based on the 
assumption that a 1975 type earthquake 
will happen again in the same area. In 
other words, the most pessimistic or con- 
servative view was taken. 



Hazard posed by regional faults fall 
into three general categories, (1) haz- 
ards created by ground shaking, (2) 
hazards created by fault displacement, 
and (3) regional changes in ground 
elevation. 

Ground Shaking 

Nothing was seen during the course of 
geologic investigations that indicates 
local earthquakes would exceed the mag- I 
nitude 6.5 Reanalysis Earthquake recom- 
mended by the Special Consulting Board 
for the Oroville earthquake. 

The intuitive conclusion to be drawn 
from the geologic studies would be that 
the magnitude 5.7 earthquake of August 1, 
1975, is close to the strongest to be 
expected and that the magnitude 6.5 local 
earthquake assumed for reanalyses of 
structures is very conservative. Geo- 
logic studies suggest the Swain Ravine 
or Prairie Creek Lineament fault zones 
are the most likely source of future j 
strong earthquakes. 

Fault Displacement 

Earthquakes of the size likely to 
occur in the Oroville area may or may i 
not cause surface rupture. If surface 
fault displacement does occur it is most 
likely along the Swain Ravine Lineament 
fault zone or perhaps the Prairie Creek 
Lineament fault zone. However, in the 
east-west tensional environment which 
apparently prevails in the Oroville area, 
small displacement could occur along any 
of the older faults or shear zones. Such i 
an event along the minor faults is con- 
sidered possible, but improbable. 

Maximum displacements in the 1975 
earthquake were about 50 mm (2 in) verti- 
cal displacement and about 25 mm (1 in) 
horizontal separation. In a somewhat 
larger event, displacements might pos- 
sibly be several times larger than these 
values along north-south trending faults. 
Although the displacements along the 



86 



Cleveland Hill Fault took considerable 
time to reach a maximum in 1975, dis- 
placements may not always develop so 
slowly. Therefore, for purposes of 
analyses, instant displacement should be 
assumed. It appears that the only import- 
ant structure which could be subjected 
to such displacements on north-south 
trending faults would be the Bidwell 
Canyon Saddle Dam. 

Regional Changes in Ground Elevation 

Although regional changes in ground 
elevation were measurable at Oroville, 
the maximum change of 60 mm (2.5 in) is 
not enough to pose a hazard to State 
water facilities. It is not expected 
this magnitude of elevation change would 
be exceeded by any future earthquake. 
Therefore, elevation changes are not 
expected to pose a hazard to facilities. 

Potential Hazard 
to Specific Facilities 

Oroville Dam and Saddle Dams 

The Cleveland Hill Fault could not be 
traced north of Mt. Ida Road about 2.1 km 
(1.3 mi) south of Bidwell Canyon Saddle 
Dam. However, the Swain Ravine Lineament 
fault zone appears to go into the Bidwell 
Canyon area of the reservoir. A fault 
was mapped in the foundation of Bidwell 
Canyon Saddle Dam near the right abut- 
ment. It should be assumed this fault 
is capable of the maximum displacement 
cited under "Fault Displacement." Bid- 
well Canyon Saddle Dam is the only struc- 
ture at Oroville with this degree of 
exposure to the probability of future 
fault displacements. 

Numerous faults were mapped in the 
foundation of Oroville Dam. None of 
these appear significant and do not 
appear to be particularly related to the 
major north-trending Mesozoic fault zones. 
It is possible that some small displace- 
ments could occur along the faults in 
the dam foundation. It is considered 
improbable that this would happen. This 
aspect should be looked at as part of 
the reanalysis of the main dam. 



If the 60-degree westward dip of the 
Cleveland Hill Fault is assumed to con- 
tinue at the reservoir, then faulting in 
the Swain Ravine Lineament fault zone 
could dip under the dam, passing under- 
neath the dam at a depth of about 5 km 
(3 mi). It must be assumed that earth- 
quakes can occur right under the dam. 
However at such shallow depth, the 
earthquakes would be of small magnitude. 

Numerous landslides developed around 
Lake Oroville since the reservoir has 
been in operation and there is geologic 
evidence for a large number of older 
slides. When the steep slopes border- 
ing the reservoir are heavily saturated 
by winter rain, the area is landslide- 
prone and a strong shake during such a 
period could trigger landslides into 
the reservoir. While such failure could 
be dangerous to boaters on the lake, it 
is not anticipated the dams would be 
endangered due to the large amount of 
freeboard which was provided on the 
Oroville dams. This freeboard of 6.7 m 
(22 ft) is expected to contain any waves 
that might be generated by landslides. 
Large seiches did not develop in the 
reservoir during the 1975 earthquake, 
but if they should, the high freeboard 
is expected to contain seiche waves also. 

Thermalito Forebay and Afterbay 

Despite the absence of conclusive 
evidence, it is possible the Prairie 
Creek Lineament fault zone extends far- 
ther northwest on a trend roughly paral- 
leling Highway 70 between Thermalito 
Power Canal and Wicks Corners. A 
sequence of three earthquakes (magni- 
tude 2.8-3.0) occurred just east of 
this stretch of highway on December 12, 
1976. Therefore, it should be assumed 
that an earthquake comparable to the 
1975 Oroville earthquake could occur 
along the northwest projection of the 
Prairie Creek Lineament fault zone near 
the east end of Thermalito Forebay. 
Fault displacements if they were to 
occur, would probably be in the Power 
Canal in a section excavated below natu- 
ral ground and therefore would not pose 
great hazard. 



8—78786 



87 



Thermallto Powerplant 

Cenozoic faults underlie Thermallto 
Powerplant. As much as 12 m (40 ft) 
of apparent vertical displacement occurs 
along a system of faults that roughly 
parallel the longitudinal axis of the 
powerhouse. The Cenozoic fault activity 
suggests small displacements could occur 
again, particularly if a 1975 type earth- 
quake were to occur along the Prairie 
Creek Lineament fault zone. Such an 
occurrence is viewed as a possible, 
though improbable, event. 

Other Structures 

For the remainder of the Oroville 
facilities the main hazard would be from 
groimd shaking earthquakes might cause. 
A number of structures have faults in 
their foundations and, conceivably, 
small displacements could occur along 
these faults. It is considered improb- 
able that displacements would occur, and 
if they did, it does not seem the damage 
would be significant. Structures with 
faults in their foundations are listed 
below: 

Edward Hyatt Powerplant 
Oroville Dam Spillway 
Thermallto Diversion Dam 
Thermallto Power Canal 
Parish Camp Saddle Dam 

SUMMARY AND CONCLUSIONS 



1. The August 1, 1975, Oroville earth- 
quake was accompanied by movement on 
the previously unrecognized Cleveland 
Hill Fault. A linear zone of discon- 
tinuous ground cracking developed 
along the fault about 7 km (4.3 mi) 
east of the main shock epicenter. 

2. Initial length of ground rupture on 
the Cleveland Hill Fault was about 
1.6 km (1.0 mi). Over a period of 
about 12 months the ground cracking 
extended progressively to the north, 
reaching a total length of 8.5 km 
(5.3 mi). 



Offset along the fault was greatest 
in the southern segment, where the 
original cracking occurred. Offset 
increased with time; movement 
amounted to about 50 mm (2 in) verti- 
cal displacement and 25 mm (1 in) 
horizontal extension. 

The Cleveland Hill Fault was not en- 
countered by trenching or geophysical , 
investigation north of Mt . Ida Road. ' 
Aftershock hypocenters projected up 
a calculated fault plane indicate the 
fault at the ground surface trends 
into Bidwell Canyon and that it may 
pass beneath Oroville Dam at depth. 

Trenching across the Cleveland Hill 
Fault by Department of Water Resource; 
and others provides evidence for 
multiple small fault displacements 
during the past 100,000 years. These 
displacements would likely have pro- 
duced earthquakes similar to the 1975 
Oroville event. 

Three major lineament-fault zones, thi 
Paynes Peak, Swain Ravine, and PrairL 
Creek, have been delineated in the 
area by geologic studies. These 
lineament-fault zones are complex 
bands of discontinuous, intertwined, 
steeply dipping faults which were 
formed during Mesozoic or earlier 
time under the influence of a differ- 
ent tectonic stress regime than exists 
today. The Cleveland Hill Fault is I 
within the Swain Ravine Lineament 
fault zone. 

Most Cenozoic fault movements in the 
Sierran foothill belt are caused by 
east-west extensional stresses re- 
activating pre-existing Paleozoic 
and Mesozoic faults such as those 
comprising the lineament-fault 
zones. 

Historic (Cenozoic) faulting and 
historic earthquake records in the 
foothill region demonstrate that the 
current and long-range level of 
seismic activity is one of low- to 
moderate-magnitude earthquakes at 
relatively long recurrence intervals. 



88 



occasionally resulting in minor 
ground rupture and offset. 

9. Nothing was seen in this geologic 
study to indicate that earthquakes 
greater than Richter Magnitude 6.5 
should be expected in the Oroville 
area. 

10. Maximum offset that should be anti- 
cipated from another Oroville-type 
earthquake is estimated to be 50 mm 



11. 



(2 in) of vertical displacement and 
25 mm (1 in) horizontal extension. 
For a somewhat larger event dis- 
placement might be several times 
larger than these values along 
north-south trending faults. 

The evidence available does not 
indicate a causal relationship 
between Lake Oroville and the earth- 
quake, but the possibility cannot be 
eliminated conclusively at this time. 



89 



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Oroville January 1969 - December 1972." Calif. Dept. Water 
Res., Earthquake Engr . Memorandum n. 61, 60 p. 

Morrison, P. W. , Jr., Stump, B. W., and Uhrhammer, R. , 1976, "The 
Oroville Earthquake Sequence of August 1975." Seis. Soc. 
America Bull., v. 66, n. 4, p. 1065-1084. 

Popenoe, W. P., 1943, "Cretaceous, East Side Sacramento Valley, 

Shasta and Butte Counties, California." Am. Assoc. Pet. Geol 
Bull., v. 27, p. 306-312. 

Radbruch-Hall, D. H,, Colton, R. B., Davies, W. E., Skipp, B. A., 
Luchitta, I., and Varnes, D. J., 1976, Preliminary Landslide 
Overview Map of the Conterminous United States: U. S. Geol. 
Survey Misc. Field Studies MF-771, scale 1:7,500,000. 

Raymond, L. A., 1977, "Emplacement of Exotic Tectonic Blocks in 
the Franciscan Complex, Northern Diablo Range, California" 
(abs.). Geol. Soc. America Abstracts with Programs, v. 9, 
n. 4, p. 486. 

Rothe, J. P., 1969, "Earthquakes and Reservoir Loading." 
A- Proceedings , Fourth World Conference on Earthquake Engineering 
\ Chile , V. 1, p. A28-A38C. 

1973, "Summary: Geophysics Report," in Ackerman , W. C, White, 

G. T., and Worthington, E. B. ( eds . ) , Man-Made Lakes: Their 
Problems and Environmental Effects . Am. Geophys . Union, 
Geophysical Monograph 17, p. 441-454. 

Russel, L. R., 1978, "The Melones Fault Zone and the Tectonic 

Framework of the Western Sierra Nevada Between the Middle and 
South Forks of the American River, California" (abs.). Geol 
Soc. America Abstracts with Programs, v. 10, n. 3, p. 145. 

Scholl, D, W. , and Marlow, M. S., 1974, "Deposits in Magmatic Arc 
and Trench Systems: Sedimentary Sequence in Modern Pacific 
Trenches and the Deformed Circum-Pacif ic Eugeosyncline , " jji 
Dott, R. H., Jr., and Shaver, R. H. (eds.). Modern and Ancient 
Geosynclinal Sedimentation : Soc. Econ . Paleon. Mineral. Spec. 
Pub. 19, p. 193-211. 

Schweickert, R. A., 1976, "Early Mesozoic Rifting and Fragmentation 
of the Cordilleran Orogen in the Western USA." Nature, v. 260, 
p. 586-591. 

Schweickert, R. A., and Cowan, D. S., 1975, "Early Mesozoic Tectonic 
Evolution of the Western Sierra Nevada, California." Geol. 
Soc- America Bull., v. 86, p. 1329-1336. 



99 



Schweickert, R. A., and Wright, W. H. , 1975, "Preliminary Evidence 
of the Tectonic History of the Calaveras Formation of the 
Western Sierra Nevada, California" (abs,). Geol. Soc. America 
Abstracts with Programs, v. 7, n. 3, p. 371-372. 

Standlee, L. A., 1978, "Middle Paleozoic Ophiolite in the Melones 

Fault Zone, Northern Sierra Nevada, California" (abs.). Geol. 
Soc. America Abstracts with Programs, v. 10, n. 3, p. 148. 

Stanton, T. W. , 1896, "The Faunal Relations of the Eocene and Upper 
Cretaceous on the Pacific Coast." U. S. Geol. Survey 
Seventeenth Ann. Rept . , p. 1005-1060. 

Strand, R. G., and Koenig, J. B., 1965, Geologic Map of California, 
Olaf P. Jenkins Edition, Sacramento Sheet: Calif. Div. Mines 
and Geol., scale 1:250,000. 

Swan, F. H., Ill, and Hanson, K. L., 1977, "Quaternary Geology 
and Age Dating," jji Woodward-Clyde Consultants, Earthquake 
Evaluation Studies of the Auburn Dam Area : Woodward-Clyde 
Consultants, unpublished, v. 4, 83 p. 

Swan, F. H., Ill, and Hanson, K. L. , 1978, "Origin and Ages of Late 
Quaternary Deposits and Buried Paleosols in the Western Sierra 
Nevada Foothills, California" (abs.). Geol. Soc. America 
Abstracts with Programs, v. 10, n. 3, p. 149. 

Taff, J. A., Hanna, G. D., and Cross, C. M. , 1940, "Type Locality 

of the Cretaceous Chico Formation." Geol. Soc. America Bull., 
V. 51, p. 1311-1328. 

Taliaferro, N. L., 1942, "Geologic History and Correlation of the 
Jurassic of Southwestern Oregon and California." Geol. Soc. 
America Bull., v. 53, p. 71-112. 

1943, "Manganese Deposits of the Sierra Nevada, Their Genesis 

and Metamorphism, " iri Jenkins, O. P. (ed.). Manganese in 
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1951, "Geology of the San Francisco Bay Counties." Calif. Dept, 

Nat. Res,, Div. Mines Bull. 154, p. 117-150. 

Townley, S. D,, and Allen, M. W., 1939, "Descriptive Catalog of 

Earthquakes of the Pacific Coast of the United States 1769-1928, 
Seis. Soc. America Bull., v. 29, n. 1, 297 p. 

Turner, H. W. , 1893, "Some Recent Contributions to the Geology of 
California." Am. Geologist, v. 11, p. 307-495. 

1894, "The Rocks of the Sierra Nevada." U. S. Geol. Survey 

Fourteenth Ann. Rept., p. 435-495. 

1896, "Further Contributions to the Geology of the Sierra 

Nevada." U. S. Geol. Survey Seventeenth Ann. Rept., p. 521-762. 



100 



Tysdal, R. G. , Case, J. E., Wrinkler, G. R., and Clark, S. H. B. , 
1911 , "Sheeted Dikes, Gabbro and Pillow Basalt in Flysch of 
Coastal Southern Alaska." Geology , v. 6, n. 3, p. 377-383. 

U. S. Army Corps of Engineers, 1977, "Fault Evaluation Study, 
Marysville Lake Project, Parks Bar Alternate, Yuba River, 
California." U. S. A. C. of Eng . , Sacto. Dist., 25 p. 

U. S, Geological Survey Staff, 1978, "Technical Review of Earthquake 
Evaluation Studies of the Auburn Dam Area (Woodward-Clyde 
Consultants, 1977) — a report to the U. S. Bureau of Reclamation, 
U. S. Geol. Survey, 143 p. 

von Huene , R., 1972, "Structure of the Continental Margin and 

Tectonism at the Eastern Aleutian Trench." Geol. Soc . America 
Bull., V. 83, p. 3613-3626. 

Whitney, J. D., 1865, "Report of the Progress and Synopsis of the 
Field Work from 1860-1864." Geol. Survey of Calif., Geology, 
V. 1, 498 p. 

Williams, H. , and Stevens, R. K., 1974, "The Ancient Continental 
Margin of Eastern North America," _in Burk, C. A., and Drake, 

C. L. ( eds . ) The Geology of the Continental Margins : New York, 
Springer-Verlag, p. 781-796. 

Wolfe, J. E., 1967, "Earthquake Hazard Report (n. 28) for the State 
Water Project - Oroville Dam Site." California Dept . Water 
Res. , 9 p. 

Wood, H. O., and Heck, N. H., 1951, "Earthquake History of the 
United States, 1769-1950, Part II, Stronger Earthquakes of 
California and Western Nevada." U. S. Dept. of Commerce, 
Coast and Geodetic Sur., n. 609. 

Woodward-Clyde Consultants, 19 77, "Earthquake Evaluation Studies 

of the Auburn Dam Area," report prepared for the U. S. Bureau 
of Reclamation, 8 volumes. 

Wright, L., 1976, "Late Cenozoic Fault Patterns and Stress Fields 
in the Great Basin and Westward Displacement of the Sierra 
Nevada Block." Geology , v. 4, n. 8, p. 489-494. 

Xenophontos, C, and Bond, G. C, 1978, "Petrology, Sedimentation 
and Paleogeography of the Smartville Terrane (Jurrasic) - 
bearing on the Genesis of the Smartville Ophiolite," _in Howell, 

D. G., and McDougall, K. A. (eds.), Mesozoic Peleogeography 
of the Western United States : Soc. Econ . Paleon. Mineral., 
Pac. Sect., p. 291-302. 



101 



ADDENDA TO CHAPTER II 



Department of Water Resources 
Exploration Trench Logs 



9—78786 103 




MODERATELY 
FRACTURED ROCK 



STRONGLY 
FRACTURED ROCK 



STIFF RED CLAY 



STRONGLY FOLIATED ROCK 

TRENCH A 



BRECCIA 

TRENCH B 



SOFT BROWN CLAY 




FRACTURED ROCK 




Rock: Metovolcontc, dork groy to greenish-groy, massive to strongly foliated, moderately to strongly fractured, frocture surfaces 

commonly limonite stoined, grades from strongly weathered near soil contoct to moderately weatfiered ot deptfi. 
Breccio: Crushed frogments of metovolcanic rock, generally with some cloy matrix. 
Gouge: Plastic cloy, grades from reddish-brown near soil contact to greentsh-groy ot depth. 

Soil: Silty, reddish-brown, friobie, residual soil. 



Logs of Exploration Trenches Along Cleveland Hill Crack Zone 



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121 



CHAPTER III 
SEISMOLOGY 



Introduction 

Bulletin 203, April 1977, reported 
earthquakes , Richter magnitude 2 . 5 and 
greater, of the Oroville sequence through 
March 1976. This report will present 
the data recorded from June 1, 1975, 
through May 31, 1978, Richter magnitude 
1.0 and greater. 

Figure 58 shows the combined DWR-USGS 
telemetered seimographic network as it 
was variously composed throughout the 
sequence. Since about August 1976, when 
the USGS network was considerably 
reduced, the network has remained essen- 
tially unchanged. The table in Figure 58 
shows the DWR network as it was variously 
composed from June 1, 1975 to May 31, 
1978. 

Data 

Epicenters of earthquakes of magnitude 
1.0 and greater for 1975 and 1976-78 
are plotted in Figures 59 and 63, respec- 
tively. The hypocenters are listed on 
Tables 3 and 4 . The aftershock zone , as 
shown in Figures 59 and 63, is parti- 
tioned into north, middle, and south 
sections . The hypocenters from each 
section are projected onto a vertical 
east-west striking plane, shown in 
Figures 60 through 62 and 64 through 66. 
Figure 67 is a vertical cross-section of 
the middle section for events recorded 
between August 2 at 1400 GMT and 
December 31, 1975. 

Surface cracking was observed on Cleve- 
land Hill to the east of the aftershock 
zone soon after the mainshock. The ver- 
tical cross-sections show that the fault 
plane surfaces near the cracks. The 
fault was sxibsequently named the Cleve- 
land Hill Fault. The Department recorded 
network as recounted in the table on 
Figure 58, was composed of MGL, ORV, KPK, 
DOG, BUT and SUT, from July 1975 into 



January 1976. Note that OSTI and OHON 
were added in November 1975. From 
January 1976 onward, the network 
included additional stations PAM, and 
OCAM. 

Consequently, to make the Department 
network of August through December 1975 
more nearly equivalent to the network 
of 1976 onward, the U. S. Geological 
Survey in Menlo Park supplied the P-times 
for aftershocks from Woodward- Clyde 
portable stations WCl, WC2, WC3, and 
WC4 and USGS Stations OCAM, OHON, OSTI, 
OWYN and ORAT. Times from Woodward- 
Clyde stations were available for events 
recorded on August 2 at 1454 GMT through 
August 12 at 1945 GMT. Times were avail- 
able for the event recorded August 5 at 
0228 GMT through December 1975. 

Installation of the USGS network stations 
began soon after the August 1 mainshock, 
and recording was initially accomplished 
on magnetic tape beginning August 5, 1975. 
By August 16, 1975, a telemetry link to 
Menlo Park was established so that their 
stations were thereafter recorded on 16 
mm. film. Woodward -Clyde installed 
four portable stations in the aftershock 
area beginning August 2, 1975. 

Hypocenters were determined by the USGS 
Hypo 71 hypocenter program (1) * and 
the Byerly Crustal model. Station correc- 
tions were determined by averaging the 
hypocenter residuals of a number of 
well recorded events. 

Magnitudes were determined by two methods: 

1 . The duration method : Magnitude- 
duration curves were plotted for 
the Oroville and Jamestown (not 
shown in Figure 58) short-period 
vertical seismometers. The elapsed 
time from the initial seismic P-wave 



(1) Number refers to reference listed at the end of the Chapter. 



123 



onset to the time the maximum seismic 
trace amplitude falls consistently 
below an arbitrary level is plotted 
against "known" earthquake magnitudes. 
This provides a plot whereby the 
magnitudes of subsequent events are 
estimated (2) . 

2. The equivalent Wood-Anderson seis- 
mograph magnitude estimate is cal- 
culated from the Oroville east-west 
short-period seismometer response 
(3). 

Figure 68 is a time-plot of the Oroville 
water surface elevation and the number of 
aftershocks by month. 

Results 

Inspection of the epicentral plots 
Figures 59 and 63 indicates a north-south 
alignment of aftershocks confined within 
a rectangle about 15 kilometres north- 
south and 8 kilometres east-west. Inspec- 
tion of Table 3 shows that on August 2 
the main aftershock activity began to 
shift to the north and south sections of 
the apparent fault rupture. After 
December 31, 1975, about 52 percent of 
the aftershocks listed in Table 4 occur 
in the south section (Figure 63) and 
about 70 percent east of the 121.5 meri- 
dian. 

A 6CP angle to the west has been drawn 
through the vertical cross sectional plots 
in Figures 60, 61, 64, and 65 to indicate 
the approximate dip of the fault break in 
the north and middle portions of the rup- 
ture zone. In Figure 61 at least two 
alignments seem evident, one near 60° and 
another about 35°. Most of the hypo- 
centers in Figures 60 through 62 were 
determined before good depth control was 
available. Therefore, Figure 67 was 
plotted beginning with events on August 2 
when the first Woodward-Clyde station was 
available. A near 60° alignment is evi- 
dent in that figure. In Figure 62 no 
clear alignment is evident. 



Figure 66, as in Figure 62, shows no 
clear alignment of hypocenters and prob- 
ably indicates more than one rupture 
plane . 

Discussion 

As in the seven years before the August 1, 
1975, Oroville earthquake, subsequent 
Lake Oroville water surface fluctuations 
do not appear to affect nearby seismic 
activity. The rapid filling of the res- 
ervoir this past winter and spring (1977- 
1978) does not, at this time, appear to 
have influenced the slow decline of the 
seismicity rate in the aftershock zone. 

The near 60° dip of the Cleveland Hill 
fault plane evident in the north and 
middle cross-sectional plots is in agree- 
ment with that reported by others (Lahr, 
et al) (4) . Savage, et al (5) report 
that the results of the repeat level 
survey profiles across the aftershock 
zone before and after August 1 are con- 
sistent with about 0.36 m of normal slip 
on the fault plane delineated by the 
aftershock sequence . 

Conclusions 

Since August 1, 1975, a correlation is 
not indicated between the Lake Oroville 
water surface variations and the rate of 
occurrence of Oroville aftershocks. 

Within the boundary of the aftershock 
zone north of 39°26'N latitude, vertical 
cross-sectional plots indicate that the 
Cleveland Hill Fault is a single, well 
defined break, dipping to the west at 
about 60° and with a near north-south 
strike. Vertical cross sectional plots 
south of 39°26'N indicate that the fault 
breaks along more than one plane. 



Figures 64 and 65 show single alignments 
near 60 in the north and middle sections . 



124 



40 « 



STATIONS RECORDED BY DWR 
DURING THE OROVILLE SEQUENCE 



STATION 


DATE ON 


DATE OFF 


DRV 


1963 


PRESENT 


KPK 


1969 


PRESENT 


MGL 


1966 


PRESENT 


PAM 


76/1/20 


PRESENT 


BUT 


75/7/1 


PRESENT 


SUT 


75/7/1 


76/12/12 


DOG 


75/8/2 


76/8/6 


OGAM 


76/1/16 


76/9/1 


OHON 


75/11/5 


PRESENT 


OSUT 


76/9/1 


PRESENT 


OSTI 


75/11/5 


PRESENT 



n MGL 



D TELEMETERED STATION 
A PORTABLE STATION 

OSLO "O" PREFIX DESIGNATES USGS STATION 

THE USGS STATIONS WERE ESTABLISHED 
ON AUG. 6, 1975 OR SHORTLY THEREAFTER. 
UNDERLINE DESIGNATES STATIONS CON- 
TINUING IN OPERATION AFTER AUG. 1,1976 



QOSHP 



ABUT 



Thermolito 
After boy 



□ OSUT 



ASUT 




39' 



Figure 58. DWR-USGS Oroville Sensitive Sei sinograph i c Network 



125 








/ \r..w 






^j^ 






• K • • • 






^ 


~-o 


,.. 


1 




m; 






- 39030 

NORTH 


-,(^ 






~n 


^ ^ 


'ffyV^iJj- 


X + 




MIDDLE 



I "tlh^' 



"■^^ 



^^ 



m 



-m 



OWR TELEMETERED SEISMIC STATIONS 





LEGEND 




X 


1 i 


•^ < 


2 


+ 


2 < 


M < 


3/ 


CD 


3 < 


M < 


4 


A 


4 < 


M < 


5 


□ 


5 i 


N < 


6 



IQ 



SCBLE 1 : 192000 
M* 1.0 



Figure 59- Oroville Foreshocks, Mainshock, and Aftershocks; June 1, 1975" 
December 31 , 1975 



126 



NORTH VERTICAL X-SECTION 
EVENTS NORTH OF AND INCLUDING 39* 29.5' N 



lU 



CO 



DISTfiNCElKM) 
10 15 



u: " 

(E < 



t 10 




MS 1.0 

Figure 60. 1975 Oroville Earthquake Hypocenters (North Vertical Cross Section) 

MIDDLE VERTICAL X-SECTION 
EVENTS BETWEEN 39" 29.5' N AND INCLUDING 39" 26.0' N 



D[STfiNCE(KM) 
10 15 



< ^ 

(C < 



1 I 




Mil.O 

Figure 61. 1975 Oroville Earthquake Hypocenters (Middle Vertical Cross Section) 

127 



SOUTH VERTICAL X-SECTION 
EVENTS SOUTH OF 39" 26.0' N 



DrSTfiNCEtKM) 
ID 15 



2Q 



</>o 

I I 



25 



-I 1 1 1 1 1 1 1 1 1 1 1 I H-f- 1^. IW — I 1 n 1 1 1 1 1 1 1 1 1 

+ + + + ++1^*!% + 



1 ID 



+ ^ 


w^ 


+ 






+ 


X 







M2:|.0 

Figure 62. 1975 Oroville Earthquake Hypocenters (South Vertical Cross Section) 



128 








_EGEND 






X 


1 


i 


M 


< 


Z 


+ 


z 


i 


- 


< 


3 


o 


3 


^ 


M 


< 


4- 


A 


H 


^ 


M 


< 


5 


□ 


S 


^ 


H 


< 


g 



SCRLE 1 ••1920D0 



Figure 63. Oroville Earthquake Epicenters (January I, 1 976-May ^1, 1978) 



129 



NORTH VERTICAL X-SECTION 
EVENTS NORTH OF AND INCLUDING 39* 29.5' N 



-I 1 1 r- 



DLSTfiNCElKM) 
.0 15 



T 1 1 1 1 1 r 




LlI 

o ^ 

< ^ 

u. o 

q: < 

=3 a: 

en o 

I I 



S5 



-T 1 1 1 1 r 



i 10 



Q_ 
UJ 
O 



'» -I- 

+ J- + 



% 



60 



+ X 




M^l.O 

Figure 6A. Oroville Earthquake Hypocenters, 1976-May 31, 1978 
(North Vertical Cross Section) 

MIDDLE VERTICAL X-SECTION 
EVENTS BETWEEN 39° 29.5' N AND INCLUDING 39°26.0' N 



■si 



Q- 
UJ 
O 



10 



DISTfiNCElKM) 
10 15 



30' 



' ' ' ' ^ -it 



^ 



q: < 
I I 



^ 



/ 



5- 




25 



M^I.O 



Figure 65- Oroville Earthquake Hypocenters, 1976-May 31, 1978 (Middle Vertical 
Cross Section) 



130 



SOUTH VERTICAL X-SECTION 
EVENTS SOUTH OF 39° 26.0' N 



DISTfiNCE(KM) 



30' 



5^ 



-1 , 1 r-^, 1 1 . 1 1 1 r— . 1 1 1— JT F*— ''K 



+-H-X ^+ ^ 



-I 1 1 r 



10 



+ -^- 



A 






M^I.O 

Figure 66. Oroville Earthquake Hypocenters, 1976-May 31. 1978 (South Vertical 
Cross Section) 

MIDDLE VERTICAL X-SECTION 
EVENTS BETWEEN 39*'29.5'N AND INCLUDING 39*'26.0'N 



DISTANCE (KM) 
I 



t 10 



Q_ 

LiJ 

a 




M=I.O 



Figure 67- Oroville Earthquake Hypocenters. August 2, 1975-December 31. 1975 
(Middle Vertical Cross Section) 



• 
• 




1 






1 






s 


JUN. 


OROVILLE SEQUENCE 


_ 


MAT 


• 
• 
• 


NUMBER OF AFTER SHOCKS /MONTH 




- 


APR 1 


WATER SURFACE ELEVATION 




MAR 












FEB 


• 






» 


JAN 


. 








DEC 










NOV 










OCT 


. 








SEP 










AUG. 




• 






JUL K 


■ 


• 
• 










< 

JUN 




MAr. 


. 








APR 


. 








MAR. 










FEB. 


_ 






'—1 


JAN 


• 








1 




DEC 






NOV 






OCT 


• 
• 






r 






SEP 




AUG 


• 








JUL « 


• 
• 








- 


JUN ~ 
MAY 


• 










APR 


• 
• 










MAR 






FEB 


• 
• 










JAN 




DEC 


• 
• 
• 
• 


? 






NUMBER Of AFTER SHOCKS - 5600 


NOV 








OCT „ 








SEP 2 






1 1 




AUG 


i '^ 


si 


O 

o 


O C 

o c 


o 2 


S '^ 


o 


o 


o 



LAKE OROVILLE WATER 
SURFACE ELEVATION 



NUMBER OF AFTER SHOCKS/MONTH 



Figure 68. 
(August 



Oroville Sequence, Number of Aftershocks/Month, Water Surface Elevat 
1975-June 1978 ) 



ion, 



132 



TABLE 3. EARTHQUAKE EPICENTERS, JUNE 1975-DECEMBER 1975 



"ITMIN 15,0 KMS OF OROvILLE 



/1/75-1J/31/75 



LtTITUOE LONGITUDE 



AG OUAQR 



75/ 6/2B 
75/ 6/je 
75/ «/je 



75/ 9/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 9/ I 
75/ 9/ 1 
75/ 9/ 1 
75/ 9/ 1 



75/ 



75/ 8/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 8/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 9/ I 
75/ 8/ 1 
75/ 9/ I 
75/ 9/ 1 
75/ 9/ 1 
75/ 8/ 1 
75/ 8/ 1 
75/ 8/ 1 
75/ 8/ 1 



«I19|S3,* 

6139140.6 
U116ll9.f 
20i|01»0.0 
18119116.1 

61111 2.« 



10139152. • 
4117137.4 
151*5137.6 
161j7ll7.t 
17I26149.9 
18117122.5 
19l351t3.0 
201201 4.5 
2DI20I12.6 
20125112.6 
20129112.6 
20132139.6 



211 91 6.« 
211)0115.6 
21111132.1 
21116123.6 
2111912*. 5 
211201 7.6 
21121150.5 
21125158.8 
21129123.9 
21139159.* 
2115*135.5 
2115811*. 1 



221111 *.5 
2211*155.6 
221151*2.5 



39.tS2 
39.*70 
39. *5* 
39.«39 



39.««1 

39.**1 
39.**1 



39,»*7 
39.**5 
39. «3* 
39.*63 
39.*63 
39. *2* 
39,tt3 

39.**e 

39.*78 
39.«*3 
39.*10 
39.**0 
39.*38 
39.*36 
39.*55 
39.«79 
39.*57 
39.*S3 
39.*16 
39.*35 
39.*33 
39.**5 
39.*90 
39.«*7 
39.453 
39.*5o 



121.538 
121.5*0 
121.5*4 
121.540 
121.5** 
121.538 
121.536 
121.516 
121.5*2 
121.560 
121.525 
121.5*2 
121.5*7 
121.5*3 
121.577 
121.539 
121.53* 
121.534 
121.534 
121.534 
121.513 
121,513 
121,542 
121.517 
121.498 
121.501 
121.501 
121.541 
121.551 
121.512 
121.542 
121,516 
121,470 
121.541 
121.5*6 
121,509 
121.551 
121.529 
121.555 
121.536 
121,509 
121.512 
121.530 
121.571 
121.516 
121.502 
121.52* 
121.580 



2.1 C PALERMO 
2.6 D PALERMO 
1.9 C PALERMO 

1.0 C PALERMO 

2.2 B PALERMO 
2.8 C PALERMO 
2,5 C PALERMO 

2.1 C PALERMO 

2.5 C PALERMO 

3.6 B PALERMO 

4.7 8 PALERMO 
3.0 B PALERMO 

2.2 C PALERMO 
2.2 C PALERMO 

4.5 B PALERMO 
5.7 B PALERMO 

4.7 B PALERMO 

4.6 B PALERMO 
3.0 B PALERMO 
3.5 B PALERMO 
2.5 B PALERMO 
2.5 C PALERMO 
2.5 C BANGOR 
3.0 B PALERMO 

3.8 B PALERMO 

2.5 D PALERMO 

2.6 B PALERMO 

3.0 B PALERMO 
2.6 B PALERMO 
2.8 6 PALERMO 
2.8 BANGOR 
3.2 B PALERMO 

2.2 C PALERMO 
2,8 B PALERMO 

4.1 C PALERMO 

3.3 B PALERMO 

3.6 B PALERMO 

2.7 9 PALERMO 

2.2 PALERMO 
2.0 B PALERMO 

3.4 B PALERMO 

2.0 C PALERMO 

2.5 A PALERMO 

3.1 9 PALERMO 
2.1 PALERMO 

1.6 D PALERMO 



2.3 
2.3 

2.3 



3i2 

5.T 



75/ 8/ 
75/ 8/ 
75/ 8/ 



75/ 8/ 1 
75/ 9/ 1 
75/ 9/ 1 
75/ 8/ 1 
75/ 9/ 1 
75/ 8/ 1 



75/ 8/ 2 
75/ 8/ 2 
75/ 8/ 2 
75/ 9/ 2 
75/ 9/ 2 
75/ 8/ 2 
75/ 9/ 2 
75/ 9/ 2 
75/ 8/ 2 
75/ 9/ 2 
75/ 8/ 2 
75/ 9/ 2 
75/ 8/ 2 
75/ 8/ 2 
75/ 8/ 2 
75/ 8/ 2 
75/ 8/ 2 
75/ 8/ 2 
75/ 9/ 2 
75/ 8/ 2 
75/ 9/ 2 
75/ 9/ 2 



221231*3.6 
221261 *.* 
22136115.8 
22150120.3 
22I52IS6.6 
231 4112.6 
23li0l»3.6 
23114141.4 
23131132.6 
231*4140.8 
23150153.5 
231511*9.6 
gi 2I»».9 
01 3132,2 
01 91 59, 9 
0113129,1 
0123152,9 
01351 4,6 
01391 3,0 
0152148,3 
0158155,8 
1131129,6 
1141137,0 
1144112.4 
21271 4,9 
2131150,7 
2139149.1 
3115121.2 
31421 7,9 
3151134,7 
41 9113.4 
41191 6.1 
*1231I3,7 
tl*3l35.6 
*1*512*.3 
51331 5.6 
51*7112.5 
51*8138.* 
5157151.4 
6130113.1 
6131156.9 
9130133.1 
101 7155.6 
10111153.5 
10148159.9 
11151121.5 
11151150.6 



9.431 
9.436 
9.435 



9.425 121.526 
9.436 121.508 
121.486 
121.503 
121.551 
121,504 
121,550 
121,549 
121,541 
121.533 
121.545 
121.50* 
121. *79 
121.558 
121.50* 
121.512 
121.517 
121.501 
121.5*7 
121.528 
121,550 
121.510 
121.55* 
121.529 
121.503 
121.515 
121.509 
121. 479 
121.517 
121. *29 
121.521 
121.561 
121.57* 
121. *89 
121,503 
121.551 
121.599 
121.528 
121.505 
121. *99 
121.502 
121. t95 
121.546 
121.525 
121.495 



438 
.426 
9.473 



7 




3.2 




PALERMO 


1 
9 




2.5 
2.3 

2.7 
2.6 




PALERMO 
BANGOR 
PALERMO 
PALERMO 


8 




2.5 

2.5 




PALERMO 
PALERMO 


9 




2.0 




OROVILLE 


7 




2.5 
3.4 




PALERMO 
PALERMO 






2.4 




PALERMO 






2.2 




PALERMO 


5 




2.5 
2.4 




BANGOR 
PALERMO 


4 




2,3 




PALERMO 


i 

1 




2,6 

2.0 




PALERMO 
PALERMO 


i 




2,6 




PALERMO 


1 




2.2 

3.9 




PALERMO 
OROVILLE 


q 




2.6 




PALERMO 


s 




2.3 




PALERMO 






2.3 




PALERMO 


a 




2.2 




PALERMO 


I 
1 




2.0 
2.3 




PALERMO 
PALERMO 


! 




2.3 




PALERMO 






2.8 




BANGOR 






2.3 




PALERMO 


2 




2.2 




BANOOR 






2.0 




PALERMO 


T 




2,0 
2.0 






PALERMO 
PALERMO 


7 




2.9 


B 


BANGOR 


7 




2.^ 


B 


PALERMO 


3 




2,2 


D 


PALERMO 


i 




2.0 


D 


PALERMO 






2.2 


C 


PALERMO 


2 




2,2 


c 


PALERMO 


5 




2.0 


c 


BANGOR 


5 




3.2 


B 


PALERMO 






2,2 


C 


BANGOR 


7 




2.8 
3.1 


e 


PALERMO 
PALERMO 


: 
1 




3.3 

2.0 


B 

D 


BANOOR 
PALERMO 






3.4 


B 


PALERMO 



2t3 
*.3 
2.7 



S.3 
3.0 
i.t 
2*0 
2.« 
2.4 



Sit 
2.7 
2.3 



9.3 
3.2 

s.e 



2.T 
6.9 
*.9 



133 



TABLE 3. EARTHQUAKE EPICENTERS, JUNE 1 975-DECEMBER 1975 (Continued) 



PTHQUaKES MTHIN 15. D K^^S OF QPOjILLE MAJN ShOCK b/1/75-12/31/75 



TITUDE LONGITUDE Dr 



AC Q QUADRANGLE 



75/ 1/ 2 t »4 38 5 J9.»2» 121.504 6.3 3.J B PALEBMO 

;;, ly I 1 It 39 3 39!»35 Ul.50» 5.7 2-2 B P4LERM0 

;i \V,lt'M\\ ll-Ml ill:*?. 1.2 2.0 6 BJNOO« 

•»ey a/ 3 I>ii4l49.5 30.451 121.502 5.0 2.2 B P«LEI1H0 

«/ ?/ I 5 O " ! " tis 2 iSs 4.5 2.0 C PALERMO 

45 IV.IV. 111:15! ?:i !:i ? S? 

75/ '/ I 17.I4.29.1 39.482 121.489 6.5 4.3 S B4N00O 

75/ 8/ 2 1714)124,0 39,4»5 121.491 6.1 4.0 B B4N80B 

is;! lliiiiU:? U:U! lll:ni 1:5 I:? I «o 

?i; s; i isuii^s:? i?:tl? ll!:^? ^: : S 

75/ s/ ? 2CI59I55.6 39,449 121.493 B.7 5.2 6 B4NG0B 



3,.4 = » i^,.. = P .." -- " eANGOB 

39.439 121.495 2.8 2.3 S B4N60B J*^ 



21133156.3 

221351 8.5 39.415 121.463 

231 0123.8 -- "" ■-■ ■" 



8AN60B 5.3 

BANGOR *•* 



,o',»» ,21 497 i.7 2.2 e BANGOP 

mum lis iiis HI liiiis lii 

75/ »/ 3 21471 8.6 39.482 121.514 7.1 4.1 A PALEPMO ♦■» 

^i;?;^ '.i!; : :4l 111:5" ::9 1:3 b paleb«o j.t 

1 l!n:l IV:X 111:1?^ ^:? i:J I -"- H 

; ?!12:I 1?:UI lilitH ^:? 1:5 i i-?§S : 

:5;!t:I i?:t^^ 111:1^8 I:? I:S ? ^t^S:? : 

ii ? I lii?!'!?:: n::f^ 1I1::?J l:? I:5 S i» : 

51; 1; ? li\^^ IVMl WVMl r.8 I" B Zlll : 

^1 2 1?::!5 1I{:!^J I:^ 1:1 I ^ii^i^ : 

: b:;n:2 n::il lll:llt ?:5 ?:l S SS : 

;ii;: 11!,?:;;:! n:tsi 1I1:U! V.l l:? I bI^^oS" ?:J 



75/ 8/ 5 2128157.1 



121:497 6:2 3.3 « bANOOR '•' 

n^hl liliilSii i2:JU m-.m J:I 2:1 1 ?itl5B§ |:| 

75/ 8/ 5 151 3155.7 39.422 121.501 6.7 1.6 4 PALERMO 3'5 

75/8/5 20.44124.4 39.429 121.537 7.9 3.2 B PALERMO 1-3 



75/ 8/ I 23 57 lui 39 496 121.534 9.3 2.8 B PALERMO 

121.555 



75/ 8/ 6 3 50129:5 39:495 121.540 «.T 4.7 A PALERMO --- 

::. :. ! : =,»n., «..99 121.555 s.i 2.8 b palermo '•' 



121.529 S.2 2.4 A PALERMO 5.7 

121.531 ».0 2.0 A PALERMO 5.T 

121.491 5.3 1.9 B BANGOR »•* 

121.541 9.0 2.7 B PALERMO 5.J 

121.533 B.S 1.9 A OROVILLE '•» 



75/ 8/ 6 10135122.8 39.433 121.505 6.1 1.8 A PALERMO | 7 

75/ 8/ 6 131 3128.3 39.502 121.533 9.3 2.9 A OROVILLE »•' 

75/ 8/ 6 131 9130.9 39.503 121.523 9.0 2.5 A OROVILLE »•» 

75/ 8/ 6 6125147.2 39.406 121.487 5.7 3.1 A BANGOR J.T 

75/ 8/ 6 16141.51.8 39.495 121.537 9.1 3.6 A PALERMO »»0 

75/ 8/ 6 19142148.3 39.504 121.543 7.7 2.8 B OROVILLE J'l 

75/ 1/ t 2? ol32:e 39:410 121.522 9.4 3.0 A PALERMO J-J 

75/ 8/ 7 5.15.48.3 39.400 121.489 7.8 2.3 8 BANOOR »•« 

7I/ 9/ 7 14 20.45 1 39.509 121.541 9.8 2.9 A OROVILLE 7.6 

75/ 8/ 7 19.f2l55.8 39.504 121.534 6.2 2.4 A OROVILLE J.O 

75/ 8/ 7 26l!5.39:i 39,;il 121.482 4.3 2.6 , 9ANGOR »•» 

75/ 8/ 7 20.31.20.0 39.503 121.531 9.2 3.1 A OROVILLE *•» 

75/ 8/ 9 53.22.9 39.415 121,526 7.9 2.3 B PALERMO 3.0 

75/ ./ 9 7 50:0 39,496 121.524 8,5 4.9 8 PALERMO »•« 

7I; S/ I nls'ilS:? IV.hl UiMz ?:6 3.2 B PALER«0 7. 

75/ 9/ 8 15.501 8.4 39.468 121.486 6.1 2.6 B BANOOR J.l 

75/ 6/ 8 19. 3.27.3 39.410 121.498 7.« 3.1 A BANOOR *•• 

75/ a/ 9 7.38147.1 39.406 121.499 6.0 3.0 A BANOOR ♦•» 

75/ 8/ 9 12l?ll35:2 ^IslO 121.518 9:2 2.4 B OROVILLE 7.i 

75/ 8/ 9 20145131.0 39,413 121.489 5.1 2.6 4 BANOOR »•» 

"/ S/.i U.35.25.; 39.406 .21.498 4,3 2.2 fl flANGOR |.» 

7!/ 8/10 21.25.44,8 39,439 121,510 .1 2.1 C PALERMO Z-j 

75/ 8/11 2.40.14.5 39.4*0 121.46T J.O 3.0 4 BANGOR »•! 

75/ 8/11 6111.36.3 39,459 121,484 5.0 4.3 A BANGOR •• ' 

75/ 8/11 6.34.50.8 39,423 121.509 5.* 2.8 A RAL"MO |;» 

75/ 8/11 7141138.8 39,413 121.488 5.6 2.7 A BAN60R »•« 

A, 1/ 15 U 5 5 39 M6 121,536 9.0 3.6 A OROVILLE 7.2 

75/ 8/12 1. 5.21.4 39.411 121,492 4.5 2.3 A BANOOR *•» 

»/ IM ll29, 2.2 39:441 121.486 4.0 2.8 B BANGOR ».l 

75/ 8/1 2 11145.21,4 39,408 121,498 3.7 2.1 A BANGOR ••• 

75/ S/ll 1 56 11 9 39 ;59 12 .534 U.O 3.0 B "LERMO «•» 

A, IM 16.42. 8.7 39.519 121,529 8.4 2.9 B OROVILLE ».» 

75/ 8/12 19.45.12.1 39.439 [21.510 6.6 2.7 4 PALERMO 2»0 

75/ A/ifc «i 4111.4 39.415 121.496 4.5 2.2 B BANOOR *•• 

"/ S/16 lUel 9:2 59.t71 121.538 9.1 4.0 A PALERMO 3.3 



134 



TABLE 3. EARTHQUAKE EPICENTERS, JUNE 1 975-DECEMBER 1975 (Continued) 



E«»TmQJ«KFS •ITHIrj Ib.o KMS Of OrtOvRLE >'4IN ShOC« 6/1/75-1J/31 



YU »0 r)> HR "N SEC L«'ITUOE LONGITUDE DfPIH 



75/ 8/16 
75/ 8/16 
75/ 8/16 
75/ 8/18 
75/ 8/20 
75/ 8/Jl 
75/ 8/?3 
75/ 8/2* 
75/ 8/2« 
75/ 8/25 
75/ 8/25 
75/ 8'26 
75/ 8/29 
75/ 8/2? 
75/ 8/2« 
75/ S/31 
75/ 9/ 3 
75/ 9/ » 
75/ 9/ » 
75/ 9/ 4 
75/ 9/ 5 



75/ 9/ 7 
75/ 9/ 7 
75/ 9/ 7 



75/ 9/28 ^ 

75/ 9/28 ^ 

75/ 9/28 1" 

75/ 9/28 21 



75/ 9/30 17 

75/10/ 2 21 

75/10/ 2 22 

75/10/ 2 22 

75/10/ 3 1 

75/10/ 3 ? 



2312*. 3 
01 8.» 
9129.8 
56153.2 
30U6.3 
0138.8 
31153.0 
10137.2 
55157.2 
35111.0 
39lt3.9 
2TC43.3 

«5i3i.e 
>eii2.5 

6125.3 
3211».» 
20153.9 
171 1.9 
39125.6 
31»3.8 
1136.9 
371 5.0 
42139.0 
10154.3 
31130.1 
361U.9 



k9l58.5 
>51 9.0 
.6123.1 



22115.7 
51 2.0 

25137.4 

59150.6 
7114.7 
3136.3 
9137.0 

51111.6 
6133.9 

42125.6 



39.502 
39.495 
39.499 
39.406 
39.413 
39.424 
39.496 
39.485 
39.503 
39.484 
39.472 
39.411 
39.476 
39.464 
39.462 
39.406 
39.498 
39.412 
39.506 
39.411 
39.406 
39,406 
39.414 
39.405 
39.559 
39.427 
39.412 
39.519 
39.517 
39.521 
39.516 
39.511 
39.395 
39.409 
39.503 
39.411 
39.402 
39.434 
39.425 



39.410 
39.424 
39.418 
39.392 
39.504 
39.513 
39.525 
39.524 
39.516 
39.527 
39.518 
39.515 
39.509 
39,520 
39.515 
39.506 
39.523 
39.502 
39.415 
39.529 
39.529 
39.441 
39.504 
39.514 
39.406 
39.464 
39.466 
39.456 
39.416 
39.411 
39.511 
39.496 
39.410 



121.513 
121.506 
121.512 
121.506 
121.500 
121.472 
121.494 
121.498 
121.492 
121.496 
121.519 
121.545 
121,511 
121.529 
121.526 
121.500 
121.487 
121.543 
121.518 
121.497 
121.515 
121.516 
121.493 
121.514 

121.531 
121.563 
121.489 
121.526 
121.524 
121.526 
121.522 
121.491 
121.516 
121.521 
121.506 
121.500 
121.516 
121.506 
121.495 
121.526 
121.527 
121.509 
121.515 
121.505 
121.501 
121.495 
121.492 
121.529 
121.529 
121.525 
121.524 
121.522 
121.521 
121.525 
121.526 
121.525 
121.525 
121.516 
121.527 
121.526 
121.506 
121.520 
121.524 
121,512 
121.518 
121.505 
121.524 
121,490 
121,496 
121.491 
121.551 
121.541 
121.490 
121.500 
121.532 



HiO 


° 


0UA09ANGLE 


3.2 




OBOVILLE 


2.6 




PALERHO 


2.8 




PALE9M0 


2.9 




PALERMO 


2.9 




PALERMO 


2.6 
3.1 




BAN30R 
SANOOR 


3.3 




BANGOR 


2.1 




orovillE dam 


3.2 




BANGOR 


2.0 
2.7 
2.9 




PALERMO 
PALERMO 
PALERMO 


2.9 
2.9 




PALERMO 
PALERMO 


2.6 




BANGOR 


2.2 

3.0 




BANGOR 
PALERMO 


2.1 

2,0 




OROVILLE 
BANGOR 


3,2 

2.9 




PALERMO 
PALERMO 


2.5 




BANGOR 


2.5 




PALERMO 


2.3 


C 


OROVILLE 


2.0 


C 


PALERMO 


2.0 


A 


BANGOR 


3.2 


B 


OROVILLE 


2.5 


B 


OROVILLE 


3.5 
2.3 


B 
B 


OROVILLE 
OROVILLE 


3.5 


e 


OROVILLE DAM 


1.9 


C 


PALERMO 


2.7 


e 


PALERMO 


2.0 
1.6 
2.4 
2.4 


c 

B 


OROVILLE 
BANGOR 
PALERMO 
PALERMO 


2.0 

4.0 


6 
B 


BANGOR 
PALERMO 


1.7 


B 


PALERMO 


2.6 
1.9 


e 


PALERMO 
PALERMO 


3.2 


e 


PALERMO 


1.9 
2.1 


B 


PALERMO 
BANGOR 


2.5 


6 


OROVILLE 0A1 


4.6 

2.9 
1.3 
1.5 
1.6 
3.0 
3,0 


B 
6 


OROVILLE 
OROVILLE 
OROVILLE 
OROVILLE 
OROVILLE 
OROVILLE 
OROVILLE 


2.6 


B 


OROVILLE 


2.0 
1.5 


J 


OROVILLE 
OROVILLE 


1.6 
2.0 
1.7 


e 


OROVILLE 
OROVILLE 
OROVILLE 


2.2 

1.6 
3.4 
2.4 
1.3 
1-3 
1.9 
2.8 
2.8 
2,0 
1,6 


B 

8 

e 

8 

c 
e 
e 

B 
6 


PALERMO 

OROVILLE 

OROVILLE 

PALERMO 

OROVILLE 

OROVILLE 

PALERMO 

BANOOR 

BANGOR 

BANGOR 

PALERMO 


2,2 


B 


PALERMO 


1,5 
1.7 


e 


OROVILLE DAM 
PALERMO 


2.6 


B 


PALERMO 



*|4 

5.T 



3.i 
3.4 
4.4 



4.6 
4.2 
4.2 

4.6 



8.6 
6.4 
8.9 



75/10/ 4 


12128139.9 


39.513 


121.519 n. 


75/10/ 6 


9154142.5 


39.411 


121.548 1(. 


75/10/10 


7144147.4 


39.462 


121.492 3. 


75/10/11 


23154156.6 


39.522 


121.518 9. 


75/10/12 


4140153,6 


39.467 


121.485 i. 


75/10/12 


151 5135.2 


39.507 


121.527 7, 


75/10/12 


23124117.4 


39.400 


121.511 4. 


75/10/13 


14156137.0 


39.492 


121.511 3. 


75/10/13 


161 6151.2 


39.492 


121.515 5. 


75/10/13 


211301 7.1 


39.428 


121.466 2. 


75/10/14 


2144158.5 


39.506 


121.527 4. 


75/10/14 


91 11 5.7 


39.406 


121.512 ?. 


75/10/14 


211321 6.4 


39,471 


121.509 


75/10/16 


3120146.9 


39.400 


121.466 |. 


75/10/18 


13159142.2 


39.408 


121.509 7. 


75/10/20 


141411 .7 


39.512 


121.522 5. 


75/10/20 


151221 5.7 


39.506 


121.525 ?. 


75/10/21 


13126125,0 


39.414 


121.512 3. 


75/10/23 


201 7144.5 


39.475 


121.534 4. 


75/10/26 


6123119.3 


39.414 


121.496 


75/10/27 


211 2144.4 


39.506 


121.517 3. 


75/10/28 


3141116.0 


39.496 


121.510 3. 


75/10/26 


5113146.0 


39.525 


121.534 A, 



OROVILLE 

PALERMO 

BANGOR 

OROVILLE 

BANGOR 

OROVILlE 

PALERMO 

PALERMO 

PALERMO 

BANGOR 

OROVILLE 

P*LErmo 

PALERMO 

BANGOR 

PALERMO 

OROVILLE 

OROVILLE 

PALERMO 

PALERMO 

BANGOR 

OROVILLE 

PALERMO 

OROVILLE 



135 



TABLE 3. EARTHQUAKE EPICENTERS, JUNE 1 975-DECEMBER 1975 (Continued) 



TMOUIRES »ITHIN 15. g KMS OF 0«OuILLE >f«IN SHOCK 



ATITUOE LONGITUDE OfP 



AG QUAQtJANGLE 



75/10/31 

75/11/ 1 

75/11/ 3 

75/11/ 3 

75/11/ ♦ 

75/11/ 5 

75/11/ 5 

75/11/ 5 

75/11/ 7 

T5/11/ e 

75/11/ 9 

75/11/ 

75/11/ 

75/11/ 

75/11/ 

75/11/ 

75/11/ 

75/11/ 

75/11/ 

75/11/ 

75/11/. _ 

75/11/20 

75/ll/?3 

75/11/33 

75/11/?* 

75/11/25 

75/11/26 

75/11/26 

75/11/26 

75/11/26 

75/11/27 

75/11/30 

75/12/ 1 

75/12/ 1 

75/12/ 3 

75/12/ 3 

75/12/ 3 



75/12/ 5 
75/12/ 6 
75/12/ 7 
75/12/ i> 
75/12/10 
75/12/11 
75/12/12 
75/12/13 
75/12/19 
75/12/20 
75/12/21 
75/12/23 
75/12/23 
75/12/26 



51371»7.0 
19l30>3e.e 
231181 6.» 
23l5ei«2.5 



31351 1.6 
31*51 8.* 
* 122 13*. 3 

131*1116.'' 
51 0135.1 

1911711*. 6 
81»71 6.5 

131101*3.1 

12136116. e 
7132153.5 



211 



7.1 



,1139.6 
1113128.5 
121 812*. 2 
12151153.3 
13153153.7 
15159123.7 
23159120.6 
10123120.9 
7150129.5 
23126127.3 
1133112.0 
7112113.1 



9.5 



8152132.* 
13133115.1 
215*1*3. » 



2155119.6 
71 »120M 
111*12*. 8 



121. *93 
121. »95 
121.509 
121.507 
121.517 
121. *92 
121.521 
121. *89 
121.517 
121. »6* 
121. *91 
121.505 
121. *B7 
121. *90 
121.501 
121. *6* 
121.506 
121. *92 
121.526 
121. »87 
121.472 
121. *8* 

121.500 
121. *90 
121.521 
121.500 
121.521 
121.496 
121. *88 
121. »e5 
121.512 
121. *93 
39.*72 121. *e3 
39,»04 121.496 
39.403 121.500 
39.42* 121. »91 
39.398 121. *65 
39.396 121.475 
39.411 121.471 
39.408 121.501 
39.507 121.48* 
39.513 121.517 
39.497 121.524 
39.507 121.541 
39.457 121.510 
39.536 121.494 
39. 4*7 121.503 
39. *5* 121.532 
39.507 121.523 
121. *75 
121.516 
121.473 
121.511 



39.502 
39.427 
39.511 



BANGOR 
PALERMO 
BANGOR 
BANGOR 



PALERMO 

BANGOR 

OROVILLE 

BANGOR 

PALERMO 



B BANGOR 

A PALERMO 

C BANGOR 

B PALERMO 

A BANGOR 

B PALERMO 

B BANGOR 

A BANGOR 

B bANGOR 

A PALERMO 

C BANGOR 

B PALERMO 

C PALERMO 

B OROVILLE 

B BANGOR 

B BANGOR 

B OROVILLE DAM 

B PALERMO 

A OROVILLE DAM 

B BANGOR 

C BANGOR 

B PALERMO 

A BANGOR 

B BANGOR 

C BANGOR 



.E OAM 



39,523 
39.431 
39.41* 



75/12/27 51**150.3 39.*10 121.501 



BANGOR 

PALER» 

OROVIL 

OROVILLE 

PALERMO 

OROVILLE 

PALERMO 

OROVILLE DA 

PALERMO 

PALERMO 

OROVILLE 

BANGOR 

OROVILLE 

BANGOR 

PALERMO 



4. a 

*.9 
*>6 
3>T 
S.3 
S.2 
»•* 
S.2 
T.O 
3.9 
3.9 
T.2 
S.5 
*.9 
«.3 
*.5 
3.2 
6.S 
«.2 
».T 



*.a 

4.S 

s.* 

2.6 
4.9 
T.2 
A. 9 



8.6 
5.S 
5.2 



136 



TABLE k. EARTHQUAKE EPICENTERS, JANUARY 1 976-MAY I978 

EAKTmUUIKES ■ITmIU 15. J KMS OF OROvILLE M«IN ShOCK 1/1/76-5/31/78 



YB MD 01 


HR UN SEC 


LATITUDE LONGITUDE 


oep 


76/ ]/ 1 


16:58:3'. 3 


39.421 


2 


.494 




76/ 1/ ? 


10:22: ''.7 


39.454 


2 


.481 




76/ 1/ 5 


14:35:38.5 


39.432 


2 


.514 




76/ 1/ « 


10:4^:32.6 


39.410 


2 


.472 




76/ 1/ i. 


15:41137.7 


39.510 


2 


.528 




76/ 1/ 4 


15:so: 4.4 


39.498 


2 


.508 


1 1. 


76/ 1/ iJ 


17:54:26.0 


39.488 


2 


.487 




76/ 1/10 


2l: 5:23.7 


39.394 


2 


.479 




76/ 1/15 


23:3l:»5.3 


39.521 


2 


.473 




76/ 1/17 


7: 15:20.0 


39.424 


2 


.467 




76/ 1/18 


0:37l22.9 


39.417 


2 


.484 




76/ 1/23 


13:i3:i8.8 


39.413 


2 


.484 




76/ l/?6 


2: l:«2.6 


39.420 


i 


,466 




76/ 1/?« 


19:401 .7 


39.416 


2 


.479 




76/ i/?6 


21= 8:i4.o 


39.437 


2 


.468 




76/ 1/J8 


3:5211''. 5 


39.405 


2 


.517 




76/ i/je 


23:41:3*. 9 


39.39* 


2 


.502 




76/ J/ , 


IB: 7:56.7 


39.526 


2 


.557 




76/ ?/ ? 


19:ii:5b,4 


39.441 


2 


.486 




76/ ?/ ? 


21:43|53.7 


39.460 


2 


.503 


loi 


76/ ?/ 9 


9:56:47.9 


39.485 


2 


.»9o 




76/ ?/ t) 


ic; 9:33.5 


39.485 


2 


,485 




76/ ?/ 9 


11: 6:46.6 


39.484 


2 


.495 




76/ ?/ ■) 


13:33: 5.6 


39.503 


2 


.523 




76/ ?/ W 


13:42139.3 


39.492 


2 


.475 




76/ ?/ «» 


13:57:49.2 


39,495 


2 


.523 


z.. 


76/ ?/!<> 


lu: 3: '.» 


39.497 


2 


.513 


6, 


76/ ?/?3 


9:59:33.4 


39.481 


2 


.498 




76/ 3/ 6 


12: 6:30.4 


39,446 


2 


.499 


u , 


76/ 3/12 


5:38:26.8 


39.480 


2 


.501 




76/ 3/15 


6:52:31.7 


39.470 


2 


,466 


1. 


76/ 3/lS 


7;i4119.5 


39.409 


2 


.503 


1 . 


76/ 3/19 


2:i4:56.3 


39.541 


2 


,509 




76/ 3/?0 


Ij: 9:42.1 


39.409 


2 


.470 




76/ 3/21 


14: 3:26.1 


39.397 


2 


.539 


2, 


76/ 3/27 


15:51:42.7 


39.500 


2 


.484 




76/ J/24 


8:»2: 5.6 


39.499 


2 


.525 


8, 


76/ «/ 3 


22:49:32.2 


39,414 


2 


.480 


7, 


76/ u/ 4 


is: 4:31.6 


39.415 


2 


.489 


2. 


76/ J,/ ^ 


■5:41:22.4 


39.491 


2 


.5o7 




76/ 4/ 1 2 


7:43: 4.5 


39.410 


2 


.415 


q. 


76/ ./16 


17:ii:5i..7 


39.501 


2 


.497 


3. 




17:21:36.9 


39.498 


2 


.480 




76/ <./3„ 


20:5l:39.2 


39.428 


2 


.466 




76/ ^/ 7 


7:41143.4 


39.465 


2 


.469 




76/ S/ 7 


13:48: 3.3 


39.419 


2 


.491 


3. 


76/ './li 


1: 7118.1 


39.420 


2 


.490 


2. 


76/ 5/15 


15:36: 7.5 


39.418 


2 


.484 


4. 


76/ 5/17 


17:13: 6.1 


39.447 


2 


.491 


3. 


76/ s/ie 


9:46:46. « 


39.4'U 


2 


.465 




76/ 5/19 


1 :10: 3.1 


39.514 


2 


.486 


3. 


76/ 5/20 


7:27:34.5 


39.489 


2 


.491 


1. 


76/ 5/21 


18:41:42.7 


39,409 


2 


.514 




76/ 5/26 


3:55:25.3 


39.495 


2 


.527 


=1. 


76/ 5/31 


2:36: 1.1 


39.507 


2 


.523 


3. 


76/ 5/31 


51301 2.4 


39.412 


2 


.477 




76/ 5/31 


6:31143.0 


39.509 


2 


.518 


fi. 


76/ 6/ 6 


12:i4:16.7 


39.420 


2 


.489 


7. 


76/ 6/14 


6:37:28.2 


39.473 


2 


.529 


». 


76/ 6/14 


18:36: 6.2 


39,525 


2 


.525 


6. 


76/ 6/14 


23:30!2''-2 


39.473 


2 


.538 


8. 


76/ 6/l« 


23:51 :54.6 


39.470 


2 


.532 


7, 


76/ 6/25 


19:12:22.2 


39.413 


2 


.465 


3, 


76/ 6/26 


6: 3: .5 


39.407 


2 


.490 


4. 


76/ 6/29 


7: 3156.4 


39.525 


2 


.528 


7, 


76/ 7/ 1 


0:59:54.4 


39.511 


2 


.492 


6. 


76/ 7/ 6 


3155:17.5 


39.409 


2 


.527 


7, 


76/ 7/ 6 


7:59:33.9 


39.413 


2 


.519 


6. 


76/ 7/ 7 


3:43i«s.i 


39.544 


2 


.513 


6. 


76/ 7/11 


23:19:20.3 


39.403 


2 


.489 


1, 


76/ 7/20 


5:33: 9,7 


39.404 


2 


.484 


3. 


76/ 7/23 


16:31=15.7 


39.512 


2 


.525 


1". 


76/ 7/24 


12:57: 2.! 


39.482 


2 


.499 


s. 


76/ 7/31, 


17:27:41.5 


39.488 


2 


,528 


1 1. 


76/ n/16 


15:11157.9 


39.411 


2 


.494 


3. 


76/ B/16 


18:53:33.7 


39.409 


2 


.501 


3. 


76/ 6/16 


23:13:41.9 


39.403 


2 


.506 


3. 


76/ n/19 


5143149.5 


39.539 


2 


.507 


5. 


76/ S/l9 


8115: 4,5 


39.4SB 


2 


.473 


5. 


76/ 5/24 


7:i5i«e,5 


39.422 


2 


.497 


J , 


76/ e/31 


19129:35.2 


39.508 


2 


.543 


9. 


76/ 9/16 


22112:47.9 


39,499 


2 


.492 




76/10/21 


6l36:U'5 


39.389 


2 


.480 


H' 


76/10/22 


13:51:11.5 


39.407 


2 


.483 


4, 


76/11/10 


8:14:57.2 


39.412 


2 


.491 




76/11/23 


12:58: 6.4 


39.469 


2 


.488 




76/12/15 


1 :32i32.2 


39.408 


2 


.496 


4 , 


76/12/25 


15:58:24.4 


39.495 


2 


.488 




76/12/29 


3: 3:55.7 


39.416 


2 


.478 




77/ 1/ 9 


23:24:40,3 


39.487 


2 


.499 


4 . 


77/ 1/ 9 


23127144,1 


39.491 


12 


1.509 


5 


77/ 1/12 


3130113.7 


39.413 


2 


.493 


2, 


77/ 1/23 


11: 2115.0 


39,401 


2 


.486 


2. 


77/ 1/30 


6135:24.9 


39.440 


2 


.491 


4, 


77/ 3/ 2 


17:55:32-6 


39.407 
39.426 


2 


.493 


4' 


77/ 3/12 


0:46il5.5 


2 


.490 


''. 



2.1 


(J 
8 


QUADRANGLE 
8ANGGR 


1.2 


C 


BANGOR 


1.9 


B 


PALERMO 


1.1 


B 


BANGOR 


1.7 


B 


OROVILLE 


1.2 


C 


PALERMO 


2.1 


B 


BANGOR 


1.2 


C 


BANGOR 


1.9 


8 


OROVILLE DA 


2.8 


B 


BANGOR 


2.9 


B 


BANGOR 


1.6 


B 


BANGOR 


2.5 


6 


BANGOR 


3.2 


B 


BANGOR 


1*4 


B 


BANGOR 


2.3 


B 


p*lErmo 


2.3 


B 


PALERMO 


2.2 


C 


OROVILLE 


1.7 


C 


BANGOR 


1.3 





PALERMO 


1-i 


c 


BANGOR 


\.U 


c 


BANGOR 


1.8 


B 


BANGOR 


2.5 


B 


OROVILLE 


1.0 


B 


BANGOR 


1.9 


C 


PALERMO 


l.C 


B 


PALERMO 


1.0 
1 .1 


C 

B 


BANGOR 
BANGOR 


1.0 
1.0 


C 

B 


PALERMO 
BANGOR 


2.9 


B 


PALERMO 


1.7 


B 


OROVILLE 


2.0 


B 


BANGOR 


1.4 


B 


PALERMO 


2.5 


B 


OROVILLE DA 


2.2 


A 


PALERMO 


1.2 


S 


BANGOR 


2.2 


B 


BANGOR 


1.2 


B 


PALERMO 


1.2 


[) 


BANGOR 


2.9 


B 


OROVILLE DA 


1.2 
2.1 


B 


BANGOR 
BANGOR 


2.4 
1.7 


C 
8 


BANGOR 
BANGOR 


1.6 


B 


BANGOR 


2.2 


B 


BANGOR 


1.1 


B 


BANGOR 


2.5 


B 


BANGOR 


1.8 


B 


OROVILLE DA 


2.3 

1.3 


6 
C 


BANGOR 
PALERMO 


2.4 


B 


PALERMO 


2.6 


B 


OHOVILLE 


2.5 


b 


BANGOR 


2.4 


B 


OROVILLE 


1.2 


B 


BANGOR 


2-5 


B 


PALERMO 


2.1 


e 


OROvILLE 


3.4 


B 


PALERMO 


2.1 


e 


PALERMO 


2.6 


B 


BANGOR 


2.4 


B 


BANGOR 


2.0 


A 


OROVILLE 


2.6 
4.1 


4 


OROVILLE DA 
PALERMO 


2.2 


A 


PALERMO 


2.0 


B 


OROVILLE 


2.6 


9 


BANGOR 


2.3 


C 


b'ngor 


1.2 


C 


OROVILLE 


1.8 
1.0 


B 

B 


BANGOR 
PALERMO 


2.5 

1.0 


8 


BANGOR 
PALERMO 


2.9 


C 


PALERMO 


1.4 


B 


OROVILLE 


2.9 


A 


BANGOR 


2.6 


B 


BANGOR 


2.7 


B 


OROVILLE 


1.1 


D 


BANGOR 


2-3 


B 


BANGOR 


2.9 


e 


BANGOR 


1.0 


c 


BANGOR 


1.9 


c 


BANGOR 


1.5 


B 


BANGOR 


2.0 


B 


BANGOR 


2.9 


B 


BANGOR 


3.4 
2.3 


8 
B 


BANGOR 
PALERMO 


2.5 


A 


BANGOR 


1.0 


B 


BANGOR 


2.8 


A 


BANGOR 


2.6 


8 


BANGOR 


1.2 


e 


BANGOR 



4.1 
4.8 
2.0 
6.3 
7.7 
6.7 
6.6 
7.1 

10.3 
6.1 
5.1 
5.3 
6,3 
5.5 
5^7 
4.3 
5.7 
9.6 
4.2 
3.5 
i'i 
6.5 
5.8 
6.9 
7.6 
6.0 
6.5 
5.4 
3.1 
5.1 
6.7 
*.S 

11. '3 
6.6 
4.9 
7.9 



7.8 
4.4 
6.1 
4.4 
4.4 
5.0 
3.7 
6.8 
9.1 
6.5 
4.0 
6.0 
7.3 
5.9 
7.6 
4.6 
3-6 
9.3 
3.6 
3.2 
6.7 
5.3 
9.4 
8,6 
3.6 
3'.* 

11.6 
5.7 
6.0 
7.9 
5.5 
5.2 
4.8 
4.5 
4.8 

11.1 
5.6 
3.8 
7^5 
7.4 
7.;4 
5.8 
4.9 
5.0 
4.9 
7.3 

5;5 

6^0 
6.0 
4.7 
6.1 
3.7 



137 



TABLE h. EARTHQUAKE EPICENTERS, JANUARY 1976-MAY 1978 (Continued) 



KES »IThIN li.O KMS OF OHOVILUE MAIN ShOCK 1/1/76-5/31/78 



aTITUOE LONGITUDE OfPTH 



AG Q UUADRANGLE 



77/ 3/1-. 


2? 


37121.7 


3V 


401 


121 


77/ 4/15 


n 


39140.2 


39 


410 


121 


77/ »/?7 


4 


?6:29.2 


39 


404 


121 


77/ 4/JB 


16 


23113.0 


39. 


416 


121 


77/ 4/?6 


17 


45152.1 


39 


410 


121 


77/ 5/ » 


6 


niU.4 


39 


400 


121 


77/ 5/ 4 


6 


15136.3 


39 


401 


121 


77/ 5/ 4 


6 


S91 10.2 


39. 


400 


121 


77/ 5/ ^ 


7 


1=53.2 


39 


399 


121 


77/ 5/U 


16 


441 7.3 


39 


502 


121 


77/ 5''7 


17 


18126.9 


39 


405 


121 


77/ 5/ IE 


17 


20122.4 


39 


461 


121 


77/ 5/J5 


I'l 


16156.4 


39 


417 


121 


77/ 7/14 





39156.1 


39 


516 


121 


77/ 7/lf. 


9 


43120.7 


39 


405 


|21 


77/ 7/1? 





57123.6 


39 


421 


121 


77/ 7/20 


8 


461 16.6 


3' 


427 


121 


77/ 8/ t> 


10 


35l28.7 


39 


4I4 


121 


77/ S/30 


22 


57144.3 


39 


419 


121 


77/ 9/13 


4 


78126.5 


39 


403 


121 


77/ g/n 


6 


39149. H 


39 


403 


121 


77/10/ 3 


18 


22136.2 


39 


442 


121 


77/in/ie. 


10 


361 10.2 


39 


556 


121 


77/11/10 


20 


24142.2 


39 


410 


121 


77/11/J3 


8 


261 2.7 


39 


402 


121 


77/1?/ t. 


14 


14157.5 


39 


4U3 


l2l 


77/lJ/ 9 


4 


3111.5 


39 


403 


121 


77/1?/ 9 


9 


391 7.8 


39 


404 


121 


77/12/11 


8 


42'41.9 


39 


l^l 


121 


7'/l?/2' 


12 


331 a.' 


3' 


121 


78/1/4 


20 


56122.8 


3' 


410 


121 


78/ 1/31 


22 


9157.9 


3"' 


521 


121 


76/ 3/ 1 


13 


55114.7 


39 


454 


121 


78/ 3/?^ 


16 


20l30.6 


39 


467 


121 


76/ 4/ 2 


14 


46153.9 


39 


421 


121 


7B/ 5/ 6 


16 


48134. H 


39 


611 


121 


76/ 5/P9 


20 


?7i31.Q 


39 


40' 


121 



1.0 B BANGOR 
1.9 B BANGOR 



2 





a 


BANGOR 


1 


2 


B 


BANGOR 


2 


9 


B 


BANGOR 


2 


7 


e 


BANGOR 


3 


4 


a 


BANGOR 


1 


5 


c 


BANGOR 


2 





c 

c 


OROVILLE 
BANGOR 




7 


D 


BANGOR 




6 


B 


BANGOR 




8 
3 


6 
C 


OROUILLE 
BANGOR 




4 


C 


BANGOR 




4 


C 


BANGOR 


1 




1 


a 


BANGOR 
PALERMO 


2 


5 


B 


BANGOR 


2 


3 


a 


BANGOR 


2 


5 


B 


PALERMO 







a 


OROVILLE 




9 


B 


BANGOR 




2 


B 


BANGOR 







B 


BANGOR 




9 


a 


PALERMO 




9 


a 


BAmGor 





1 



a 
c 

B 


BANGOR 
BAnGOR 
BANGOR 









c 

c 
c 


OROVILLE 
PALERMO 
BANGOR 







D 


BANGOR 




2 




BIDWELL 







B 


BANGOR 



a>6 
4.9 
S.7 
5.3 
5.7 
6^4 
6.5 
6.6 
6.7 
8.3 
5.9 
7.0 
5.7 
8.6 
5.6 
7.1 
5.0 
4.6 
2.7 
8.6 
8.3 
1.2 

13. B 
4.6 
S.6 
6.9 
5.1 
6.0 
5.4 

12.1 
5.8 

10-1 
1.6 
5.9 
9.3 
9.0 
6.6 



138 



References 



1. Lee, W. H. K, , J. C. Lahr, Hypo 71 (revised). "A Computer Program 
for Determining Hypocenter, Magnitude, and First Motion Pattern 
of Local Earthquakes." USGS Open File Report 75-311. 

2. Lee, W. H. K. , R. E. Bennett and K. L. Meagher (1972). "A Method of 
Estimating Magnitude of Local Earthquakes from Signal Duration." 

U. S. Geological Survey Open File Report. 

3. Hofmann, R. B. , and R. W. Wylie (1964), "Proceedings of the VESIAC 
Conference on Seismic Event Magnitude Determination." Institute of 
Science and Technology, University of Michigan. 

4. Lahr, K. M. , J. C. Lahr, A. G. Sindh, C. G. Bufe and F. W. Winter 

(1976), "The August 1975 Oroville Earthquakes." BSSA, 66, 4, p. 1085- 
1099. 

5. Savage, J. C. , M. Lisowski, W. H. Prescott, and J. P. Church (1977) 
"Geodetic Measurements of Deformation Associated with the Oroville, 
California Earthquake." JGR, 82,11, p. 1667-1671. 



139 



CHAPTER IV 
VERTICAL AND HORIZONTAL GEODESY 



Vertical Crustal Movements 

Introduction 

Due to the August 1, 1975, Oroville 
earthquake (magnitude 5.7, main shock 
located about 12 kilometres southwest of 
Oroville Dam) , the Department began an 
intensive surveying program to reobserve 
previously established vertical and hori- 
zontal control networks to determine 
locations , magnitude of movement , and 
trends of the major faulting. 

A monitoring network to measure horizon- 
tal and vertical movement at Lake Oro- 
ville was established in 1967. In April 

1968, the horizontal network was remea- 
sured, and in late 1968 releveling of 
most of the network was completed. In 

1969, about half of the level network 
was releveled. 



Figure 69 shows a plot of the filling 
of Lake Oroville to Elevation 274.3 
metres (900 feet) starting in 1967 and 
the normal cycling of the lake. Also 
shown is the effect of the California 
Drought; beginning in 1976, Lake Oroville 
receded to its lowest water elevation 
since filling (198.1 metres (650 feet) 
in December 1977) . Between December 1977 
and June 1978, the lake refilled to 
within a few metres of full . 

The location of the Oroville area level 
lines (1977) are shown on Figure 70. 
The end points of the lines are not 
connected on this figure for purposes 
of identification. 

The following precise surveys were made 
in response to the August 1, 1975, 
Oroville earthquake: 



1975 August - September: 

1976 January - April : 

1976 September - November: 

1977 September - November: 

1978 August - September: 



Leveling and Horizontal Control . 

About half releveled. 

All level lines rerun. 

All level lines rerun. 

About 90 percent of total length to be rerun. 



Precise Survey Programs 

September 1967 . The Lake Oroville Moni- 
toring Network was the original program 
to monitor the area around Oroville Dam 
and Lake Oroville for movement caused by 
the filling of the lake. Figure 71 shows 
the level net for study of Lake Oroville - 
1967. 

The program consisted of establishing 106 
new bench marks as well as leveling an 
additional 139 bench marks for a total 
network of 94.15 kilometres (58.5 miles). 
The leveling accuracy was Class 1 first- 
orderi/, which means the closure error 



for each line may not exceed 3.0 milli- 
metres (0.010 foot) times the square 
root of the distance in kilometres. 

The leveling was run from an established 
U. S. Coast and Geodetic Survey (USC&GS) 
network in and near the City of Oroville. 
The leveling was extended easterly and 
northerly to diorite rock masses. The 
USC&GS established the elevations along 
the lines and terminal bench marks and 
the Department completed first-order 
leveling over the net interior. These 
elevations are used as the base refer- 
ence elevations. 



1/ Classification, Standards of Accuracy, and General Specifications of Geodetic 
Control Surveys, U. S. Department of Commerce (February 1974). 



141 



300 



250 




900 



-800 



700 



200 



-600 



500 



1967 



1968 ' 1969 ■ 1970 ' 1971 1972 1973 1974 1975 1976 1977 
Figure 69- Lake Oroville Water Surface Elevation 



1978 



July - September 1968 . Because the funds 
allotted to survey the entire 1967 net- 
work were inadequate. Line Olive leveling 
was omitted (12.1 kilometres, 7.5 miles), 
and the Line Bald Rock leveling was 
shortened by 9.7 kilometres (6.0 miles). 
This leveling was performed using second- 
order methods, single run, except where 
differences in elevations between bench 
marks previously established were in 
excess of first-order tolerances, then 
reruns were made for confirmation. A 
total of 57.6 kilometres (35.8 miles) 
were run in one direction between July 1 
and September 27. 

October - November 1969 . Twenty-six kilo- 
metres (16 miles) of Class 1 first-order 
leveling was conducted during October 
and November. 

August - September 1975. Because of the 



earthquake, 117 kilometres (73 miles) 
of Class 1 first-order levels were made 
between August 13 and September 12. 

January - April 1976 . Seventy-five kilo- 
metres (47 miles) of Class 1 first-order 
levels were made between January 19 and 
April 8. 

September - November 1976 . One hundred 
sixty-two kilometres (101 miles) of 
Class 1 first-order leveling was accomp- 
lished between September 15 and November 
11. 

September - November 1977 . This Class 1 
first-order survey releveled the 
September - November 1976 network of 162 
kilometres (101 miles) , between Septem- 
ber 13 and November 3. The level net, 
for study of Lake Oroville - 1977, is 
shown on Figure 72. 



142 



Precise Survey Adjustment 

Free Adjustment . Independent free adjust- 
ments for each epoch of leveling were 
made using the variation-of-parameters 
method of least squares. In free adjust- 
ments, the net is not constrained to fit 
previously established elevations. Only 
Bench Mark OM-27, Elevation 540.468 metres 



(1773.19 feet), (1967 USC&GS adjustment) 
is assumed to be stable at the fixed 
elevation, and all other elevations are 
adjusted in relation to it. Therefore, 
any comparison of the free adjusted 
elevation of a bench mark in one epoch 
to that of another epoch indicates 
apparent movement between two levelings. 




MINERS RANCH 

FEATHER FALLS 



BIDWELL CANYON SADDLE DAM 
fN-MINERS RANCH 



MISSION OLIVE 

CLEVELAND HILL 
ORO-BANGOR 



AVOCADO 



K(LOI«ETRE 

Figure 70. Orovi lie Area Level Lines (1977) 



143 




KILOMCTKE 



Figure 71. Precise Level Net for Study of Lake Oroville - 1967 



The level net used for the October 1977 
adjustment typifies the basic network as 
refined to that date (Figure 72) . 

Spur Lines . Six main spur lines are 
connected to the main net without benefit 
of closure back to the net; therefore, 
these lines are not adjusted and reflect 
only observed elevations. The lines are 
Feather Falls, Bidwell, Bald Rock, Rich- 



vale, Potter and Line 103. Also, sev- 
eral short spur lines are connected to 
the net. 

Line Feather Falls is the connecting 
link between the fixed Bench Mark OM-27 
and OM-20 (main connector to the level 
net) ; therefore. Feather Falls line is 
the actual observed elevations with no 
adjustments. 



144 



Lines Bidwell and Bald Rock are connected 
to the net at OM-20 and are observed 
elevations without any adjustments. The 
1977 elevation on the Bald Rock Terminal 
Bench Mark (L1092) is 22 millimetres 
(0.072 foot) lower than the established 
1967 elevation. This elevation differ- 
ence is within Class 2 first-order level- 
ing limits and, therefore, may not be 
indicative of a 22-millimetre (0.072-foot) 
subsidence. 



Line Richvale is also a spur line 
connected to the level net at the west 
side with excellent agreement between 
October 1977 and October 1976. The 
extreme west bench mark indicates 16 
millimetres (0.052 foot) of subsidence 
and the entire line varies between 
(1976-77) 10 and 20 millimetres (0.033 
and 0.066 foot) . 

After Line 103 leaves the level net, 
this spur line indicates the same type 




Figure 72. Precise Level Net for Study of the Oroville Earthquake - 1977 



145 



divergence between 1976 and 1977; that 
is about 20 millimetres (0.066 foot) 
lower than 1976, using observed eleva- 
tions. The 1976 data show uplift, and 
the 1977 data indicate subsidence. 

Potter shows approximately 10 millime- 
tres (0.033 foot) of uplift in 1977 com- 
pared to October 1976. The 1977 subsi- 
dence is more than 20 millimetres (0.066 
foot) compared to the reference date of 
September 1967. 

Elevation Differential Isograms 

General . The elevation differential 
isograms are hand-drawn representation 
lines of equal vertical elevation differ- 
ences for each epoch. By necessity, a 
certain amount of judgement is used in 
the determination of the contour lines . 
Generally, the contours developed from 
the spur lines are less credible because 
they are observed elevations. Therefore, 
these elevation-differential isograms 
(Figures 73 through 78) are limited in 
the area of the spur line, and care must 
be used in interpretation of the contours 
in these areas. The interpretation of 
these spur line contours was intention- 
ally limited by not developing contours 
to their extremities; however, all data 
for these spur lines are shown on the 
vertical elevation differential plots. 

September 1967 - October 1969 (Figure 73 ). 
This epoch shows elevation differentials 
during initial filling of Lake Oroville 
starting in October 1967 and reaching 
maximum lake elevation of 274.3 metres 
(900 feet) in July 1969. 

This isogram shows no subsidence south 
of Lake Oroville and only minor subsidence 
on the southeast side of the lake up to a 
maximum of only 20 millimetres (0.066 
foot) , based on spur line observed eleva- 
tions. This isogram is limited in extent 
because only 26 kilometres (16.2 miles) 
of the original 1967 net were releveled. 
It shows that only very minor subsidence 
occurred during this period. 



October 1969 - August 1975 (Figure 74 ) . 
This epoch is the result of (a) the 
normal lake cycling from 1969 to 1974, 

(b) the lower-than-normal cycle in 
winter of 1974 to elevation 228.6 metres 

(750 feet) , (c) refilling to maximum 
lake elevation in June 1975, and (d) the 
effects of the August 1, 1975, Oroville 
earthquake . 

This isogram is also limited in extent 
because of the short survey in 1969; how- 
ever, it does show that only minor sub- 
sidence was measured during August 1975 
after the main shock, with most of the 
subsidence occurring after this survey. 
The maximum 1975 subsidence contour is 
only 15 millimetres (0.049 foot) along 
the southeast side of the lake. 

September 1967 - October 1977 (Figure 75 ) 
The 1967-1977, ten-year epoch, encom- 
passes all measurable elevation differ- 
entials from all causes. They include 
the previous items plus the earthquake 
aftershock sequence, and the effect of 
the California drought, which resulted 
in Lake Oroville being drawn down to its 
lowest elevation of 198.1 metres (650 
feet) in October 1977. 

Generally, the subsidence adjacent to 
Lake Oroville and Dam is fairly uniform, 
ranging between 20-25 millimetres (0.066- 
0.082 foot). A significant subsidence 
area to the south and west of the dam 
indicates increased subsidence away from 
the lake and dam, especially in the 
southern direction, to a maximum of 60 
millimetres (0.20 foot). 

Throughout the area south of Lake Oroville 
the subsidence is quite predominant and 
may be attributed to the fault zone. 

August 1975 - October 1976 (Figure 76 ) . 
This epoch includes only the immediate 
aftershock sequence and decreasing lake 
elevation from 274.3 metres (900 feet) 
to 233.2 metres (765 feet). 



146 




+ UPLIFT 

- SUBSIDENCE 



OROVILLE EARTHQUAKE 
EPICENTER Ml=5.7 
AUGUST I, 1975 



NOTES: 

1. ALL CONTOURS ARE 
IN MILLIMETRES 

2. BENCHMARK OM-27 HELD 
FOR FREE ADJUSTMENT 



i 



Figure 73. Elevation Differential I sogram— September 1967-October 1969 



147 




Figure 7^. Elevation Differential I sograni--October 1969-August 1975 



148 




LEGEND 

+ UPLIFT 

- SUBSIDENCE 



OROVILLE EARTHQUAKE 
EPICENTER Ml=5.7 
AUGUST I, 1975 



NOTES: 

1. ALL CONTOURS ARE 
IN MILLIMETRES 

2. BENCHMARK OM-27 HELD 
FOR FREE ADJUSTMENT 



Figure 75. Elevation Differential I sogrann--September 1967-October 1977 



149 




+ UPLIFT 

- SUBSIDENCE 



OROVILLE EARTHQUAKE 
EPICENTER Ml=5 7 
AUGUST I, 1975 



NOTES: 

1. ALL CONTOURS ARE 
IN MILLIMETRES 

2. BENCHMARK OM-27 HELD 
FOR FREE ADJUSTMENT 



.1 .. . ,? 



Figure 76. Elevation Differential I sogram--August 1975-October 1976 



150 




OROVILLE EARTHQUAKE 
EPICENTER Ml=57 
AUGUST I, 1975 



NOTES: 

1. ALL CONTOURS ARE 
IN MILLIMETRES 

2. BENCHMARK OM-27 HELD 
FOR FREE ADJUSTMENT 



Figure 77. Elevation Differential I sogram--October 1976-October 1977 



151 




LEGEND 

+ UPLIFT 

- SUBSIDENCE 



OROVILLE EARTHQUAKE 
EPICENTER Ml=5 7 
AUGUST I, 1975 



NOTES: 

1. ALL CONTOURS ARE 
IN MILLIMETRES 

2. BENCHMARK OM-27 HELD 
FOR FREE ADJUSTMENT 



Figure 78. Elevation Differential I sogram--August 1975-October 1977 



152 



This epoch clearly shows subsidence to 
the west of the nearly north-south zero 
line. The magnitudes are small near the 
dam and lake; however, significant 
trends developed south of the lake. The 
contours south of the Lake show a north- 
south trending fault zone through lines 
Mission Olive and Cleveland Hill. The 
ground surface to the west shows a net 
subsidence of 40 millimetres (0.131 
foot) across this zone, with ground 
rupture present in this fault zone. 

October 1976 - October 1977 (Figure 77 ) . 
This epoch includes continued lowering 
of Lake Oroville — due to the drought — 
of approximately 35.1 metres (115 feet) 
and the declining aftershock sequence. 

North-south uplift between the two north- 
south zero lines is predominant for this 
epoch. Significant uplift between the 
two north-south 5-millimetre (0.016-foot) 
contours is well defined with two areas 
of 10-millimetre (0.033-foot) uplift. 

The fault zone through lines Cleveland 
Hill and Mission Olive is not clearly 
defined during this time period. How- 
ever, the area south of the dam defines 
that previous area, although it shows 
uplift of plus 5 millimetres (0.016 
foot) compared to the previous subsi- 
dence in this area. 

August 1975 - October 1977 (Figure 78 ) . 
This epoch includes the reduced lake 
elevation of approximately 76.2 metres 
(250 feet) and the entire aftershock 
sequence shortly after the August 1, 
1975, main shock. 

This time period clearly shows the fault 
zone through lines Cleveland Hill, 
Mission Olive, and just south of the 
lake. The north-south zero line separ- 
ates the subsidence to the west and the 
uplift to the east with the dam and lake 
in the subsidence area. The magnitude 
of movement adjacent to the dam and lake 
is very small and insignificant. The 
net subsidence across the fault zone is 
approximately 50 millimetres (0.164 foot) 



with lowering of ground surface to the 
west. 

Elevation Differential Along Lines 

General . The plots of the elevation 
differential for each of the lines are 
based on a free adjustment holding OM-27 
fixed. Spur lines are observed eleva- 
tions based on the adjusted junction 
bench mark elevation. 

The Oroville area level lines are shown 
on Figure 70 and the locations of the 
bench marks are shown on Plate 2 (inside 
rear cover) . The reference dates for 
the individual lines vary according to 
when the line was first established for 
monitoring of the Oroville area. Also 
shown on each figure is a plot of the 
approximate ground profile along the 
line for topographical referencing. 

The lines listed below are presented in 
alphabetical order along with comments 
concerning significant movements and 
anomalies. 

Avocado (Figure 79) (Reference Date 
February 1976) 

1. Possible southern extension of 
fault zone. 

Bald Rock (Figure 80) (Reference Date 
September 1967) 

1. Spur line, observed elevations 
only. 

2. The October 1977 plot is approxi- 
mately 20 millimetres (0.066 foot) 
below the August 1975 plot although 
within first-order survey error 
limits. 

Bidwell (Figure 81) (Reference Date 
September 1967) 

1, Spur line observed elevations 
only. 



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1. REFERENCE DATE 

2. BENCHMARK OM-2 
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3 SPUR LINE 
LEGEND 
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156 



2. Close agreement of plots, except 

that divergence starts at about 

OM-42 to OM-48 where this line 
connects to Bald Rock. 

Bidwell Canyon Saddle Dam (Figure 82) 
(Reference Date September 1967) 

1. Spur line from OM-33. 

2. Only minor subsidence is evident 
up to about 20 millimetres (0.066 
foot) . 

3. Normal embankment consolidation is 
shown at Moniaments 2 , 3 and 6 
through 10. 

Canyon Drive (Figure 83) (Reference Date 
August 1975) 

1. No significant subsidence after 
the February 1976 survey. 

2. Fifteen to 20 millimetres (0.049 
to 0.066 foot) subsidence mea- 
sured between August 1975 and 
February 1976. 

Cleveland Hill (Figure 84) (Reference 
Date October 1975) 

1. Spur line, observed elevations 
only. 

2 . Ground cracking occurred between 
N and M. 

3. Definite subsidence to the west of 
the ground cracking. 

4. Significant subsidence occurred at 
the ground rupture between Novem- 
ber 1975 and February 1976. 

Dam (Figure 85) (Reference Date 
September 1967) 

1. Subsidence fairly consistent at 
25 to 35 millimetres (0.082 - 
0.115 foot) . 

Duns tone (Figure 86) (Reference Date 
February 1976) 



1. Only minor variations of less 

than 10 millimetres (0.033 foot). 

Feather Falls (Figure 87) (Reference 
Date September 1967) 

1. Spur line fixed from Bench Mark 
OM-27. 

2. OM-27 is the fixed elevation bench 
mark for the entire level net and 
spur lines. 

3. Consistency of the adjacent bench 
marks (OM-26, H-925, H-80, and OM- 
25) shows a stable area for the 
fixed reference Bench Mark OM-27. 

4. Localized discontinuity between 
G1092 and OM25. 

Foothill (Figure 88) (Reference Date 
August 1975) 

1. Consistent elevation after Febru- 
ary 1976. 

2. Fifteen to 40 millimetres (0.049 - 
0.131 foot) of subsidence shown 
between August 1975 and February 
1976. 

Miners Ranch (Figure 89) (Reference Date 
September 1967) 

1. Anomaly occurs at OM-17 (17 milli- 
metres (0.056 foot) uplift October 
1976 to October 1977) , possibly 
disturbed by power pole installa- 
tion. 

2. Significant subsidence west of 
Q-925. 

Mission Olive (Figure 90) (Reference 
Date October 1975) 

1. Significant fault zone movement 
between 4RBR and 5RBR. Ground 
cracking observed between MO-4 
and MO-5, MO- 7 and MO-8. 

2 . Movement occurred between November 
1975 and February 1976. 



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3. Magnitude of fault zone is approx- 
imately 40 millimetres (0.131 
foot) . 

Morris (Figure 91) (Reference Date 
August 1968) 

1. Shows only minor uplift. 

Olive (Figure 92) (Reference Date Sep- 
tember 1967) 

1. Significant subsidence trend at 
OM-13 to a maximum of 63 milli- 
metres (0.207 foot). 

Oro-Bangor (Figure 93) (Reference Date 
August 1975) 

1. Almost stable line with the excep- 
tion of minor uplift between 
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tween 20 BWSRM and SBM-2 up to 
about 10 millimetres (0.033 foot) . 

Oroville (Figure 94) (Reference Date 
September 1967) 

1. Anomalies at OM-6, W-145 and OM-50 
indicate uplift from previous 
trends, probably localized condi- 
tions . 

2. A major portion of the line has 
subsided up to 40 millimetres 
(0.131 foot) since September 1967. 

Potter (Figure 95) (Reference Date 
September 1967) 



1. Spur line, observed elevations 
only. 

2. Almost perfect agreement between 
October 1976 and October 1977. 

3. Line subsidence of approximately 
10 to 20 millimetres (0.033 to 
0.066 foot) . 

Thompson Flat (Figure 98) (Reference 
Date September 1967) 

1. Spur line, observed elevations 
only. 

2. Subsidence trend range from 10 

to 20 millimetres (0.033 to 0.066 
foot) . 

Wyn-Miners Ranch (Figure 99) (Reference 
Date August 1975) 

1. Significant subsidence at the 

southern end of this line between 
A-234 and 5RBR ranging to 40 
millimetres (0.131 foot) entering 
into the fault zone . 

103 (Figures 100, 101, 102) (Reference 
Date 1957) 

1. Spur line, observed elevations 
only. 

2. Minor variation of the October 
1976 and 1977 surveys referenced 
to the 1957 datum but all are 
within the error limits. 



1. Spur line, observed elevations 
only. 

2. Consistent pattern with some 
variation. 

3. The 1977 range of settlement is 
approximately 20 millimetres 
(0.066 foot) . 

Richvale (Figures 96 and 97) (Reference 
Date August 1975) 



Oroville Dam Crest Differential Settle- 
ment (Figure 103) (Reference Date 
April 1969) (Referenced to Abutments) 

General . Figure 103 is included only to 
show the relationship of the earthquake 
to the consolidation rate of the Oroville 
Dam embankment. The graph shows that 
the consolidation rate increased after 
the July 1975 survey due to the August 1 
earthquake; however, the same settlement 
pattern continues. The lake elevations 



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at the time of the surveys are tabulated 
on the plot. 

Commentary 

The Department is requesting the National 
Oceanic and Atmospheric Administration 
(NOAAO to relevel key bench marks, espe- 
cially OM27, when they rerun their first 
order network in the Oroville area. 

The annual releveling frequency of the 
Oroville area network will be decreased 
to a longer interval not to exceed five 
years . 

Conclusions 



The following conclusions are based on 
free adjustment holding the elevation of 
OM-27 fixed (1967 USC&GS adjustment) and 
therefore, all elevation differentials 
are relative to OM-27. 

1. Based on the preearthquake datum of 
1967, the greatest elevation differ- 
ential was only 63 millimetres 
0.207 foot) on line Olive during the 
ten-year epoch (1967-1977) , 

2. The August 1, 1975, Oroville earth- 
quake is associated with minor sub- 
sidence in the Oroville area, mainly 
south and southwest of Lake Oroville. 

3. Most of the subsidence associated 
with the August 1, 1975, Oroville 
earthquake was measured between late 
August 1975 to October 1976. 

4. The elevation differentials show 
movement of the fault zone that 
passes through the level lines 
Cleveland Hill and Mission Olive 
(ground cracking was evident before 
the lines were established) . A 
fault zone may pass through the level 
lines Miners Ranch south of Lake 
Oroville; however, no ground crack- 
ing was found there . 

5. Minor subsidence of less than 25 
millimetres (0.082 foot) has been 



measured adjacent to Oroville Dam 
and Lake between 1967 to 1977 due 
to all causes. 

Horizontal Earth Movements 

Introduction 

During 1967, the Lake Oroville horizontal 
monitor network was established to iden- 
tify movement that might be associated 
with filling of Lake Oroville and changes 
that could occur in the event of major 
tectonic activity. Figure 104 shows 
the Horizontal Geodetic Control and 
Triangulation Net, 1967-1975, as refined 
to this latter date. 

Horizontal Geodetic Control and Triangu- 
lation Programs 

September 1967 . The 1967 original hori- 
zontal monitoring. Class 1 first-order 
triangulation, included four base lines 
measured by a geodimeter. A total of 
14 stations were established. Ten con- 
crete piers, with stainless-steel instru- 
ment adapters cast into them, were con- 
structed. A permanent tower 6.1 meter 
(20 feet) high, constructed with 103- 
millimetre (4-inch) galvanized pipe with 
a stainless-steel instrument adapter, 
was built on Kelly Ridge because of 
restrictions regarding the cutting of 
trees and brush set forth by the Division 
of Beaches and Parks. The remaining 
stations were those of the U. S. Coast 
and Geodetic Survey. Metal instrument 
stands were erected over them and bolted 
to rocks to avoid disturbing the station 
marks. Triangulation station Bald Rock 
is located in the same diorite rock mass 
as the northernmost bench mark. 

Observations were made at night and con- 
sisted of 16 sets of directions taken 
with a first-order theodolite. 

The observation at Gaub did not have 
pointings to Oroville that would have 
extablished an azimuth at the westerly 
extreme of the net and did not have any 
observations to or from Bald Rock to fix 



181 




C^L 






OROVILLE EARTHQUAKE 
EPICENTER Ml- 5 7 
AUGUST I, 1975 



I .5 I 



I .5 I 

L_LJ I 

KILOMETRE 



Figure 104. Horizontal Geodetic Control and Tri angul at ion Net (1967~1975) 



182 



any orientation at the northeasterly 
extreme, nor was an astronomical azimuth 
observed from any station. 

Seven lines were measured with a Model 
2A Geodimeter providing the scale for 
the net. Three lines were measured on 
two successive nights; the remaining 
four were single observations. The 2A 
Geodimeter was considered a first-order 
base-line instrument. 

April 1968 . The second observation pro- 
gram of the horizontal net at Lake Oro- 
ville was conducted during the period 
from April 1 through April 12, 1968. A 
small change was made in the original 
triangulation net. Observations from 
the original station Spillway became 
blocked by a fence; therefore, it was 
reestablished in the other spillway 
abutment. A property owner requested 
the removal of station Potter. Hence, 
Potter 2 was established on State prop- 
erty. Observations were made to and 
from the original station Potter before 
it was removed. 

Observations were made at night and con- 
sisted of 16 sets of directions (32 
pointings) with a first-order theodolite. 
Three of the lines measured in 1976 were 
remeasured. 

Comparison at that time between the ini- 
tial observation in 1967 and that in 
April 1968 resulted in no detectable 
horizontal movement. 

August - September 1975 . During August 
and September 1975, after the August 1, 
1975, Oroville earthquake, the Lake 
Oroville horizontal monitor net was 
reobserved. Night-time observations 
from the 16 existing stations were made 
with a Wild T-3 Theodolite, using the 
16-set criteria at each station. 

Several stations have been replaced 
since the original 1967 work: 

Potter 1 replaced by Potter 2 in 1968 
Spillway replaced by Spillway 3 in 1968 



Kelly 2 replaced by Kelly 2 Eccentric 
in 1972 

Computations and Analyses 

In general, the comparison of observed 
angles for 1967, 1968, and 1975 shows 
little change. The observations at 
Spillway 3 (1968 vs. 1975) indicate a 
five-second difference; however, an exam- 
ination of the triangle closures revealed 
that the 1968 index pointing from Spill- 
way 3 to ODPT was in error , and this 
line was deleted from the recomputation 
of the 1968 work. 

The lack of any direct azimuth orienta- 
tion for the 1967 observation was cor- 
rected in 1968. The original computa- 
tions show the Department's 1968 position 
for Bald Rock instead of the U. S. Coast 
and Geodetic Survey (USC&GS) published 
value. This 1968 value determination 
had been used in both the original 1967 
and 1968 computations. Difficulties were 
encountered when the 1968 repositioning 
of USC&GS second-order station Bald Rock 
was checked. Apparently, computations 
had progressed from Bald Rock through a 
fully-observed traverse to USC&GS first- 
order station Gaub. This closure error 
was somehow transferred back to Bald 
Rock and a new position established. 

Using the original 1967/68 field data, 
this traverse was recomputed with a 
closure error on the Bald Rock USC&GS 
published position of 24 millimetres 
(-0.08 foot) north, and 9 millimetres 
(-0.03) foot) east. The standard devia- 
tion of the direction work was 0.44 
second and the length error ratio was 
one part in 377,640. This closure 
error was much less than the 88 milli- 
metres (-0.29 foot) north and 49 milli- 
metres (+0.16 foot) east, indicated in 
the original 1968 computations. Since 
it appeared that the USC&GS published 
position of Bald Rock was actually com- 
patible with the other first- and second- 
order stations in the area, all work was 
recomputed on this basis. 

Length measurements were used in the 
computations to check the position of 



H— 78786 



183 



Bald Rock, but only one length (Line 2 
to Bald Rock) has been used for the 1967/ 
196B and 1975 complete net computations . 
All computation and recomputation was by 
the Department of Transportation "Cosmos" 
Computer Program, which is a least 
squares adjustment by variation of geo- 
graphic coordinates. Probabilities as 
computed in the program were : 

1967 Monitor Net - 0.45 seconds standard 
deviation for directions - lengths 
1 part in 467,049. 

1968 "Bald Rock" Traverse - 0.53 seconds 
standard deviation for directions - 
lengths 1 part in 391,512 

1968 Monitor Net - 0.74 seconds standard 
deviation for directions - lengths 
1 part in 282,029 

1975 Monitor Net - 0.78 seconds standard 
deviation for directions - lengths 
1 part in 267,022 

Although the errors and adjustments 
appear quite small, the computed dis- 
tances have not matched measured lengths 



as well as could be expected. The com- 
puted lengths are plus or minus 50 
millimetres (0.164 foot) long, as com- 
pared to measured lengths. During the 
1975 observations, the U. S. Geological 
Survey measured two lines in the Oroville 
Monitor Net, incidental to other work 
they were involved with. The measured 
length, Kelly 2 Eccentric to Cameron, is 
103 millimetres (0.338 foot) shorter 
than the computed length. The measured 
length, Kelly 2 Eccentric to Line 2, is 
46 millimetres (0.151 foot) shorter than 
the computed length. Five short lines 
were measured with the MA/100 telluro- 
meter by the Department during February 
1976; the Line Kelly 2 Eccentric to 
Line 2 was within 5 millimetres (0.016 
foot) of the uses measured distance. 
The other four lines were about 50 milli- 
metres (0.164 foot) shorter than the 
computed distances. The Line Kelly 2 
Eccentric to Cameron is beyond the MA/100 
range . 

Comparative coordinate position changes, 
in millimetres, from the 1967 observation 
are: 



1968 



1975 







mm 


feet 


mm 


feet 


Loafer 


North 


-37 


-0.12 


-37 


-0.12 




East 


-12 


-0.04 


+85 


+0.28 


Island 


North 


-37 


-0.12 


+18 


+0.06 




East 


+ 21 


+0.07 


+ 85 


+0.28 


Reed 2 


North 


- 9 


-0.03 


-43 


-0.14 




East 


+15 


+0.05 


+ 9 


+0.03 


Spill 3 


North 


_ 


_ 


- 6 


-0.02 




East 


- 


- 


- 3 


-0.01 


OPDT 


North 


-12 


-0.04 


-49 


-0.16 




East 


- 9 


+0.03 


+ 9 


+ 0.03 


Morris 


North 


- 6 


-0.18 


-15 


-0.05 




East 


+ 3 


+0.01 


+ 3 


+0.01 


Cameron 


North 


-55 


-0.18 


+ 30 


+0.10 




East 


+ 9 


+ 0.03 


+94 


+0.31 



184 



1968 



1975 







mm 


feet 


mm 


feet 


Intake 


North 


-27 


-0.09 


-34 


-0.11 




East 


- 3 


-0.01 


+27 


+0.09 


Line 2 


North 


-21 


-0.07 


-12 


-0.04 




East 


+18 


+0.06 


+ 9 


+0.03 


Potter 2 


North 


- 


- 


+ 6 


+0.02 




East 


- 


- 









The computed movements of the stations 
of this net are insignificant; and, in 
many cases, within the accuracy of the 
surveys. It appears that there is a 
scale problem in the network, but the 
effect on comparative position differ- 
ences would be slight. Results of the 
1975 survey indicate a north and east 
expansion of the easterly portion of the 
net. However, the measured lengths from 
Kelly 2 Eccentric to Loafer and Cameron 
as compared to computed lengths , do not 
support this indicated change . 

Commentary 

The 1975 observations and calculations 
were to be a duplicate of the 1967 and 
1968 work. However, all stations visible 
from each occupied station were observed, 
and all directions used in the computa- 
tions. Therefore, some of these observed 
lines form extremely poor figures and 
adversely affect the final station 
positions. 



Based on the comparison of preearthquake 
(1967) and postearthquake (1975) values, 
using all observations, the greatest 
computed movement occurred at station 
Cameron, with a movement of +98 milli- 
metres (+0.32 foot) northeasterly. 

The Oroville Horizontal Geodetic Control 
and Triangulation Net is being refined 
to include only strong figures. All the 
previous surveys will be recalculated 
on this basis for comparison to future 
surveys. 

Conclusions 

1. All computed horizontal movements 
are minor and in many cases within 
the accuracy of the existing surveys 
and computations . 

2. The August 1, 1975, Oroville earth- 
quake did not cause sufficiently 
large horizontal movements that 
could be reliably measured and 
calculated within the Lake Oroville 
Monitoring Network. 



185 



CHAPTER V 

OROVILLE DAM: 

EVALUATION OF SEISMIC STABILITY 

Acknowledgements for Chapter V 

For some time before the earthquakes of 1975, the Division of Safety of Dams 
had been working on static and dynamic analyses of Oroville Dam as part of their 
program for developing dynamic analysis capability. After the earthquakes, the 
Divisions of Operations and Maintenance and Design and Construction undertook the 
analysis of seismic preparedness and safety of the Oroville Complex including an 
evaluation of seismic stability of Oroville Dam. Much of the work already completed 
by Division of Safety of Dams was used in this evaluation, and they were requested 
to participate in the additional studies following the earthquake. 

John Vrymoed performed the static and most of the dynamic finite element 
analyses, and interpreted the acceleration records of the 1975 earthquakes. He 
Sso prokded advice on additional dynamic analyses. His office report. 'Dynamic 
Analysis of Oroville Dam." provided most of the material for several chapters and 
much of the additional information for this report. 

Bill Bennett planned the cyclic test program for Oroville gravel and carried 
out the first 20 tests. His report. "Evaluation of Sample Density for Triaxial 
Testing of Oroville Gravel," is the basis of the discussion of sample density in 
Section 8. 

Emil Calzascia made the modifications to the Pacoima and Taft acceleration 
records to develop the acceleration time history for the ^^f ^ly^^^^^^'^^jjj^^^^^f J„ 
did the filtering and corrections to acceleration records of the 1975 earthquakes to 
produce acceleration time histories and response spectra. 

Through many discussions. Rashid Ahmad, Emil Calzascia. Bill Bennett, and 
John Vrymoed of the Division of Safety of Dams contributed immeasurably to the under- 
standing of complex aspects of the analysis, and suggested methods for solving prob- 
lems associated with three-dimensional effects. 

Harry Kashiwada of the Soils Laboratory made many trips to Richmond to assist 
in conducting the cyclic triaxial tests, first with Bill Bennett and later with 
N. Banerjee. 

N. Banerjee took over the cyclic testing and completed the program, under the 
direction of Professor H. B. Seed. 

I The guidance and advice provided by Professor Seed during the studies is 
I especially appreciated. 



187 



INTRODUCTION 



Background 

Oroville Dam is situated on the Feather 
River in the foothills of the Sierra 
Nevada above the Sacramento Valley. The 



dam is 8 kilometres (5 miles) east of the 
City of Oroville and about 113 kilometres 
(70 miles) north of Sacramento (see 
Figure 105) . 



Oregon 




SAN FRANCISCO 



MILES 
Figure 105. 



It has a maximum embankmen 
235 metres (770 feet) and 
of 1 707 metres (5,600 fee 
gated spillway to the left 
The 61 000 000 cubic metre 
cubic yards) embankment is 
inclined impervious gravel 
founded on a concrete core 
sand-gravel-cobble transit 
shells upstream and downst 



t height of 
crest length 
t) from the 

abutment. 

(80,000,000 

made up of an 
-clay core 

block, with 
ions and 
ream. 



KILOMETRES 

Location Map 



Historically, there have been some mod- 
erately strong earthquakes in the Oro- 
ville region. However, when the dam was 
built in the early 1960 's, there were no 
known active faults within 32 kilometres 
(20 miles) of the dam. Design to resist 
earthquakes was a major consideration, 
and the best methods available at the 
time were used. Embankment slopes were 
analyzed by modified Swedish slip-circle, 



sliding-wedge, and infinite-slope meth- 
ods, with a 0. Ig horizontal acceleration 
force included to represent earthquake 
loading. The upstream slope of the em- 
bankment was investigated to find the 
critical conditions for stability with 
several reservoir levels. The minimum 
safety factor found was 1.2 for the up- 
stream slope, with the reservoir lowered 
90 metres (300 feet). 

In addition, a series of shake table 
tests, using a 1:400 scale embankment 
model, were conducted by Professor Seed 
at the University of California at 
Berkeley; he also performed analytical 
studies to calculate seismic coefficients 
for the dam for the El Centre earthquake 
(maximum acceleration = 0.25g), to esti- 
mate soil strengths that would exist in 
the gravel shell during an earthquake, 
and to determine safety factors for up- 
stream sliding wedges. Seismic coeffi- 
cients varied from O.lg to 0.25g, 
strengths (friction angle of gravel 
shell) varied from 42° to 38°, the in- 
clination of contact force between wedges 
varied from to 20°, and resulting safe- 
ty factors varied from 1.75 to 1. 

On August 1, 1975, an earthquake of Rich- 
ter Magnitude 5.7 occurred about 12 kilo- 
metres (7.5 miles) from the dam. The 
Oroville earthquake series began with a 
number of foreshocks on June 28, 1975. 
Then on August 1, twenty-nine foreshocks 
occurred within 5 hours of the main 
shock. The largest of these foreshocks 
had a magnitude of 4.7. Many after- 
shocks, with magnitudes up to 5.1, oc- 
curred throughout August, and scattered 
shocks continued for many months. 

I 
The embankment performed well in all the 
shocks of the Oroville earthquake seq- 
uence, which produced accelerations at 
the base of the dam of about O.lg on 
three different occasions. Instrument- 
ation results indicated maximum permanent 
displacements of about 25 millimetres 
(1 inch) , pore-pressure rise in the core 
of 15 metres (50 feet) , and maximum tran- 
sitory pore-pressure response in the up- 
stream transition of 83 kilopascals (0.8 



TSF) . Performance of all Oroville area 
water-project structures is detailed in 
DWR Bulletin No. 203, "Performance of the 
Oroville Dam and Related Facilities Dur- 
ing the August 1, 1975 Earthquake". 

Even though the embankment performed well 
in the 1975 Oroville earthquake, it be- 
came apparent that active faults were 
quite close to the dam. The question be- 
came: What earthquake is now appropriate 
for analysis of Oroville Dam, and how 
will the dam perform in that earthquake? 
To answer this question, the Department 
of Water Resources began the comprehen- 
sive investigation described in this 
report. 

The dam performance was to be evaluated 
by the latest state-of-the-art procedures, 
which included cyclic-strength testing of 
gravels, studies of the observed embank- 
ment response to ascertain the in-place 
shear modulus of the gravel shells, and 
static and dynamic finite-element-method 
analyses to determine stresses in the em- 
bankment. To assist in the evaluations, 
the Department convened a special consult- 
ing board of foremost specialists in geo- 
logy, seismology, dynamic analysis, and 
practical dam design. This board has 
provided guidance in completing the stud- 
ies discussed in this report and has re- 
viewed the findings. 

Commentary 

1. It is generally accepted that very 
dense cohesionless soils will not de- 
velop liquefaction flow slides. The 
cyclic triaxial tests on Zone 3 grav- 
els provide additional support for 
this concept. Pore water pressures 
might rise momentarily to the value 
of the confining pressure on any load- 
ing cycle, but would then drop quickly 
as the sample strained. In order for 
a flow slide to be possible, pore 
pressure would have to remain high as 
strain progressed. 

In evaluating embankment performance, 
liquefaction flow slides were not con- 
sidered possible. The objective was 



189 



to make the best estimate possible of 
the extent of deformtions that would 
be caused by earthquake shaking. 

2. An apparent discrepancy has been noted 
by the profession between strains in 
laboratory test samples and strains 
calculated by dynamic analyses. For 
dynamic stresses developed in an em- 
bankment during strong earthquake 
shaking, calculated linear elastic 
shear strains may approach 1 percent. 
At the same stresses, laboratory sam- 
ples could reach 5- to 10-percent 
shear strain. Developers of dynamic 
analysis procedures generally contend 
that calculated stresses are correct 
even though the strains may be 
incorrect. 

In situations where initial static 
shear stresses are high, the strain 
in any one cycle may be about the same 
in a test sample as that calculated 
for a field element. Sample strain 
accumulates incrementally in one di- 
rection until the accumulated strain 
is 5 to 10 percent. On any one cycle, 
shear strain may not be much greater 
than 1 percent. 

However in other situations, where 
initial static shear stress is small- 
er, the strain on any one cycle gen- 
erally reaches several percent. 

3. There has not been any method devel- 
oped and verified for calculating em- 
bankment deformations caused by earth- 
quake shaking, other than the rough 
estimate of average shear strain po- 
tential times height, made for Upper 
San Fernando Dam. This dam developed 
very high strain potentials, and prob- 
ably a liquefied interior zone. It is 
not at all clear that the same method 
would apply to a case with much small- 
er strain potentials and no liquefied 
zone. For example, it is commonly 
accepted that a dam which develops 
compressive strain potentials of less 
than 5 percent will not suffer signi- 
ficant deformations. But, if the des- 
cribed method is used, displacements 



of many feet would be calculated for 
a high dam with average compressive 
strain potential of only 2 or 3 
percent. 

4. The investigation was limited to the 
upstream shell whose strength might be 
reduced during earthquake shaking be- 
cause of saturation and possible lack 
of drainage. The downstream shell is 
essentially dry and would presumably 
retain full drained strength and 
therefore would develop smaller 
strains. The core is a compacted clay 
-gravel, a type of material found to 
perform well in strong earthquake shak 
ing. On the basis of these consider- 
ations, the maximum deformations would 
be expected in the upstream shell. 

Summary of Findings 

1. Oroville Dam is in a narrow canyon 
relative to the height of the dam, 
which complicates the problem of anal- 
yzing earthquake response. Abutment 
restraint has a significant effect on 
natural period, accelerations, dis- 
placements, and stresses. Two- 
dimensional methods of dynamic anal- 
ysis will not give correct values for 
these response factors. However, a 
two-dimensional analysis can be forced 
to give the correct period and crest 
accelerations (when they are known 
from crest acceleration records) by 
deliberately using an incorrect 
(pseudo) modulus for the embankment 
soil. An extension of this approach 
was used to take into account the 
abutment restraint effects on shear 
stresses generated by the Reanalysis 
Earthquake. A basic assumption was 
that the same pseudo shear modulus 
which gave the correct response in the 
recorded earthquake will also give the 
correct response in a stronger earth- 
quake. The effect of abutment re- 
straint is to reduce shear stresses 
significantly in the upper part of 
the embankment . 



190 



2. The cyclic strength of dense cohesion- 
less soil is difficult to assess, par- 
ticularly at consolidation stresses 
lower than the critical confining 
pressure. Cyclic triaxial tests were 
carried out on dense samples of 
Monterey "0" sand and Oroville sand at 
low isotropic consolidation stresses, 
with special attention given to ob- 
serving sample behavior. The results 
indicate that dense samples strain as 
uniformly as do loose samples, and 
that higher strains are produced when 
higher cyclic stresses are applied. 
These observations hold true even for 
cyclic stresses higher than the con- 
solidation stress. However, it was 
found that strain occurs only in the 
extension direction for dense sands. 

Studies by others on dense Monterey 
"0" sand also showed that triaxial 
sample strain is all in the extension 
direction, but that triaxial stress- 
strain behavior still correlates with 
shaking-table stress-strain behavior. 
These tests all used cyclic triaxial 
stresses less than the consolidation 
stress. 

On the basis of all these studies, the 
cyclic triaxial test is considered to 
be as applicable for evaluating cyclic 
strain behavior at consolidation 
stresses lower than critical confining 
pressure as at consolidation stresses 
above critical confining pressure. 
However, further studies are needed 
at higher cyclic stresses to extend 
the correlation between cyclic triax- 
ial and shaking-table tests. 

The cyclic triaxial tests on Oroville 
gravel were used to determine cyclic 
shear strength envelopes. For some 
tests, curves of strain vs. number of 
cycles were conservatively extrapolat- 
ed, because cyclic load dropped or 
apparent sample necking occurred early 
in the test. 

3. Many of the analysis conditions and 
soil properties could not be deter- 
mined precisely. Ranges of support- 
able choices and values were defined 



by testing, analysis of observed per- 
formance, and comparison with other 
published data. Embankment displace- 
ments were estimated for two cases de- 
fined as follows : 

a) Best Judgment Case — For each ana- 
lysis condition and soil property, 
use the value within the defined 
range that is best supported by 
available evidence and judgment. 

b) Conservative Case — For each ana- 
lysis condition and soil property, 
use the end of the defined range 
that produces the higher estimated 
displacement. 

Predicted behavior of the dam, based on 
the "best judgment case," is that no 
slides or large movements will develop; 
but permanent displacements on the order 
of a metre could develop at the surface 
of the upstream slope. This predicted 
behavior is considered conservative in 
many respects, and the possibility of 
greater displacements is considered 
remote. 

Displacements were also estimated for 
the "conservative case," which is consi- 
dered the extreme behavior that could be 
postulated from the defined ranges of 
soil properties and conditions. The sur- 
face of the upstream slope might undergo 
displacements of 10 metres (33 feet) be- 
tween the two berms, slumping near the 
upper berm, and bulging near the lower 
berm. Although uncomfortably large, 
these movements would not threaten the 
safety of the dam. Remember, this is 
not the predicted behavior, but the ex- 
treme limit that could be postulated if 
all soil properties and conditions were 
more adverse than the best judgment 
choices. 

Conclusions 

1. The seismic stability of Oroville Dam 
was investigated for the Reanalysis 
Earthquake of Richter Magnitude 6.5, 
at a hypocentral distance of 5 kilo- 
metres (3 miles) from the dam, and 
producing the following ground motion 



191 



characteristics at the base of the 
dam: 



maximum acceleration 
predominant period 
duration 

acceleration time 
history 



0.6 g 

0.4 seconds 
20 seconds 
modified Pacoima 

plus modified 

Taft 



It was concluded that this ground shal 
ing was more severe than any future 
shaking likely to affect the dam. 

2, Using "best judgment" choices for in- 
put soil properties and conditions, 
relatively small embankment deforma- 
tions were estimated by the seismic 
evaluation procedures. It is con- 
cluded that Oroville Dam would perfon 
satisfactorily if subjected to the 
Reanalysis Earthquake. 



2. DESCRIPTION OF EMBANKMENT MATERIALS 
AND DYNAMIC INSTRUMENTATION 



Embankment Materials 

Materials comprising the various zones 
of the dam considered in the analyses 
are shown on Figure 106. Gradation 



curves for these materials are shown on 
Figure 107. 



CO 

^ 300 

I- 
liJ 



O 100 

I- 
< 
> 

LU r, 



N.W. S. ELEV. 274.3m(900ft) 



CREST ELEV. 281.0m (922ft) 




ZONE I a 4 - 6 900 000 CUBIC METRES (9,000,000 CUBIC YARDS) 

IMPERVIOUS 
ZONE 2- 7 260 000 CUBIC METRES (9,500,000 CUBIC YARDS) 

TRANSITION 
ZONE 3- 46 710 000 CUBIC METRES (61, 100,000 CUBIC YARDS) 

PERVIOUS 
RIPRAP- 315 700 CUBIC M E TRES (413,000 CUBIC YARDS) 

CONCRETE - 222 500 CUBIC METRES (291,000 CUBIC YARDS) 

Figure 106. Oroville Dam Maximum Section 



192 



200 
100 



80 



100 



50 



30 



U.S. STANDARD SIEVE SIZES 
16 8 4 3/8" 3/4" I-I/2' 



' 




1 


1 






1 






y 


' 


/ 


















/ 


/ 


/ 


/ 
















/ 


/ 


/ 


:/ 
















^ 


/ 


/ 


/ y 


















/ 


/ 

















,el^ 




^ 




A 


/ 










zoD 


r^ 








/ 


A 










^ 






JO 


fioni^ 


^ 


y 














J 


ri 


\^£^ 


'Shelll 


/ 












-f 




1 


1 






1 






1 







0.1 



0.5 1.0 5.0 10.0 

GRAIN SIZE IN MILLIMETERS 



50.0 100.0 



Figure 107. Average Gradation Curves of Oroville Dam Materials 



The materials used in each zone and the 
compaction methods were: 

Zone 1 — Impervious core consisting of a 
well-graded mixture of clays, silts, 
sands, gravels, and cobbles to 8 centi- 
metre (3- inch) maximum size. Compaction 
was in 25-centimetre (10-inch) lifts by 
90.7-tonne (100-ton) pneumatic rollers. 
Average in-place dry density achieved 
was 2 2A3 kilograms per cubic metre (140 
pounds per cubic foot) at 8.0 percent 
moisture (average 100 percent compac- 
tion, DWR standard 20,000 ft-lbs per 
cubic foot). 



Zone 2 — Transition consisting of a well- 
graded mixture of silts, sands, gravels, 
cobbles, and boulders to 38-centimetre 
(15-inch) maximum size (6-percent limit 
on minus No. 200 U. S. Standard sieve). 
Compaction was in 38-centimetre (15- 
inch) lifts by smooth-drum vibratory 
rollers. Average in-place dry density 
achieved was 2 419 kilograms per cubic 
metre (151 pounds per cubic foot) at 
3.9 percent moisture (average 99 per- 
cent compaction, DWR standard vibratory 
maximum density test). 

Zone 3 — Shell of predominantly sands. 



193 



gravels, cobbles, and boulders to 61- 
centimetre (24-inch) maximum size (up to 
25 percent minus No. 4 U. S. standard 
sieve sizes permitted) . Compaction was 
in 61-centimetre (24-inch) lifts by 
smooth-drum vibratory rollers. Average 
in-place dry density achieved was 2 355 
kilograms per cubic metre (147 poimds 
per cubic foot) at 3.1-percent moisture 
(average 99 percent compaction, DWR 
standard vibratory maximum density test). 

Zone 4 — Buffer zone designed to compress, 
contains between 15 and 45 percent pass- 
ing No. 200 U. S. standard sieve with 20- 
centimetre (8- inch) maximum size. Com- 
paction was in 38-centimetre (15-inch) 
lifts by a smooth-drum vibratory roller. 
Average density was 1 666 kilograms per 
cubic metre (104 pounds per cubic foot). 
(Average 82-percent compaction, DWR 
standard 20,000 ft-lb per cubic foot.) 



Dynamic Instrumentation | 

The embedded dynamic instrumentation sys- 
tem at Oroville Dam has been operating 
on a limited basis since the August 1975 
earthquake. Since then the system has 
deteriorated to a point requiring a com- 
pletely new present "state-of-the-art 
system" in order to obtain reliable, con- 
sistent dynamic records. Following is a 
description of the original system and 
the upgraded system. 

Original System 

The originally installed djmamic instru- 
mentation system at Oroville is inoper- 
able. This system included four force- 
balance type accelerometers, 6 pore pres- 
sure sensors, and 15 soil-stress cells, 
installed at the maximum section (Statioi 
53 + 05) as shown: on Figure 108. 



• PORE PRESSURE CELLS 
■ ACCELEROMETERS 
"^ SOIL STRESS CELLS 



N.W.S. ELEV. 274.3m 
(900 ft.) 



CREST ELEV. 281.0m (922ft.) 




ELEV. 45.7m ( 150 ft.) 



FEET 



Figure I08. Oroville Dam Embankment, Original Dynamic Instrumentation 



194 



Iwo accelerometers were located in the 
embankment, one at the crest, and one 
in an abutment near the toe of the dam. 
rhe exact locations are as follows: 

No. A-1 Beneath the crest at Elevation 
I 207.3 metres (680 feet). 

No. A-2 Beneath the crest at Elevation 
244.1 metres (801 feet). 

No. A-3 Downstream toe, on rock at 
Elevation 45.7 metres (150 
feet) . 

No. A-4 On the crest at downstream 
edge Elevation 281.0 metres 
(922 feet). 

These instruments measured accelerations 
.along three orthogonal axes: Vertical, 
upstream-downstream (N46°E) and cross 
canyon. In cooperation with the U. S. 
Geologic Survey, (USGS) , three strong- 
motion accelerographs were placed at the 
site. One was located at the crest in 
the same vault with A-4, one in the core 
block gallery, and one on rock at Eleva- 
tion 341.4 metres (1,120 feet) about 1.6 
kilometres (1 mile) northwest of the dam 
(Seismograph Station ORV) . The core- 
block and crest instruments were orient- 
ed as described above. The seismograph 
station instrument was oriented with one 
of the horizontal axes at N37°E. With 
the exception of the core-block unit, 
all strong-motion instruments were oper- 
able during the 1975 earthquake activity. 

All six dynamic pore-pressure cells in- 
stalled in the upstream shell and transi- 
tion zones showed a response during one 
event or another of the August earthquake 
series. The five groups of stress cells 
were located in the downstream shell. 
Each cell group measures stresses verti- 
cally and at 45 degrees to vertical in 
the upstream and downstream direction. 
Each cell has two transducers; one mea- 
sures both static and dynamic stresses 
(CEC) , and the other measures static 
stress only (MAIHAK) . Cell Numbers 1, 
2, 5, 6, 7, 10, 11, 12, and 14 were op- 
erable during the 1975 earthquake 
activity. 



Upgraded System 

Following the August 1, 1975 Oroville 
earthquake, the special consulting board 
recommended improvements to the seismic- 
data-acquisition system at Oroville. In 
March, 1977, the system was upgraded by 
adding new strong-motion accelerographs 
at two stations on the dam crest, in the 
grout-gallery adits on each abutment, 
and in the core block. These five in- 
struments were all connected to a trig- 
ger at the toe of the dam and to a time- 
signal receiver (WWVB) . In December 
1978, the system was further upgraded 
by replacing failed accelerometers and 
pore-pressure signal-conditioning equip- 
ment, and by connecting all but two sen- 
sors to a digital recorder located in 
the Area Control Center. However, the 
dynamic soil-stress cells, which were 
rendered inoperable by a lightning 
strike at the dam in September 1976, 
were not replaced. 

The following is a detailed description 
of the upgraded system as of December 
1978 (Figure 109): 

1. Installed four new SMA-IA strong- 
motion accelerographs in the two in- 
strument vaults on the crest, in the 
left grout gallery portal, and in the 
toe seepage vault. These replaced 
existing SMA-1 units. These units 
will provide film record of accelera- 
tion at the unit and digital record 
in the Area Control Center. (The two 
existing SMA-1 units in the right 
grout-gallery portal and core block 
will provide film record at the unit 
only) . 

2. Installed three new FBA-3 force- 
balance accelerometers. Two of these, 
at the toe seepage vault and at crest 
Station 53 instrument vault, replaced 
failed units and provide redundancy 
with the SMA-IA records, which are 

on separate power supply. The third 
FBA-3 was installed in instrument 
house T on the downstream slope of 
the dam at midheight. This unit is 
a replacement for two existing FBA 
units buried in the embankment. The 



195 



buried units were at the limit of 
their life expectancy and were giving 
questionable readings. All the FBA-3 
units will provide digital record of 
accelerations in the Area Control 
Center. 

3. Installed two EFM-1 earthquake force 
monitors in the Area Control Center. 
They are connected to the SMA-IA units 
in the toe seepage vault and at crest 
Station 53. They will display the 
maximum acceleration experienced 



6 - • PORE PRESSURE CELLS 
3 - ■ ACCELEROMETERS 

6- A STRONG MOTION ACCELEROGRAPH 

(SMA-I OR SMA-IA) 



since the last reset. 

4. Installed new power-supply and cali- 
bration - signal conditioning equip- 
ment for the six pore-pressure cells. 

5. Installed new DDS-1105 digital record 
er in the Area Control Center, and 
connected it to four new SMA-IA, thre 
new FBA-3, and six pore-pressure cell. 
All units are connected to a common 
trigger. A common time base (WWVB) 
will be recorded on all records. 




CREST 



AREA CONTROL 
CENTER 



N.W.S. ELEV. 274.3 m 
(900 ft.) 



CREST ELEV. 281.0m (922ft.) 




ELEV. 45.7m (150 ft.) 



FEET 



METRES 



Figure 109. Oroville Dam Embankment, Present Dynamic Instrumentation 
(December, 1978) 



196 



3. RECORDED EMBANKMENT RESPONSE TO THE 1975 EARTHQUAKE 



I General 

If complete and clear records had been 
obtained for the three or four larger 
shocks of 1975, a rare chance would have 
been available to test the mathematical 
models used for dynamic analysis by com- 
paring the computed response with the 
observed response of the embankment. 
Unfortunately, the recording system was 
beset with problems and failures, and 
only partial records were obtained for 
the strongest shocks. One complete, 
clear set of records was obtained - for 
the September 27 aftershock. 



The main use made of the records was to 
estimate the natural period of the dam. 
Secondarily, computed and recorded crest 
motions were compared for the August 1 
and September 27 events (Section 5). 
These comparisons were complicated by the 
three-dimensional effect of the canyon. 
Recorded dynamic pore pressures in the 
upstream shell and transition zones were 
not significantly large. All dynamic 
normal stresses were small. 

Embankment response was evaluated for the 
following events : 







Epicenter 




Distance 




Seismic 




Lat. 


Richter 


from dam 


Depth 


Event 




Long. 


Magnitude 


(Km/mi) 


(Km/mi) 


Aug. 1 




39° 26-33' 


5.7 


11/7 


9/5.5 


(main sho( 


:k) 


121° 31-71' 








Aug. 5 




39° 28-73' 
121° 31-46' 


4.7 


7/4 


9/5.5 


Sept. 27 




39° 30-65' 
121° 32-69' 


4.6 


3.5/2 


5.5/3.5 



Many other foreshocks and aftershocks 
were recorded but were not used in these 
analyses. 

For the August 1 and August 5 events, 
there were gaps in the records during 
the strongest shaking. However even if 
records had been obtained during this 
interval, they could not have been deci- 
phered because of the overlap of adja- 
cent records (Figures FllO and F112) . 
The aftershock of September 27, produced 
the only complete, clear records; how- 
ever, the acceleration was of lower amp- 
litude and higher frequency than the 
first two. 



Recorded Events 

August 1, 1975 

The DWR accelerometers were triggered by 
a minor foreshock and were still record- 
ing when the main shock occurred. With 
the arrival of the large accelerations 
of the main shock, other instruments 
(pore pressure and stress cells) were 
triggered, resulting in an overload and 
a temporary loss of power. This loss of 
power caused all of the instruments to 
stop recording for most of the duration 
of the strong motion. After several sec- 
onds, the back-up power source was acti- 
vated and all of the instruments started 
to record again. The record is shown in 
Figure 110. 



197 



RECORDER NO. I 



r- Awi.i ■■"■ »m,-i^mi liuw 




''''■^;.ZZ #> -^ii^ 



ELEV. 680, VERTICAL AiiM^IV^^^j^^^i^iiiKHiW 
ELEV. 801, VERTICAL 

ELEV. 580, TRANSVERSE t f ^ 0h >ii>f**i'»ii^mim 

ELEV. 801, TRANSVERSE ||^jp(tV^^.|Wliyii^»ii|«in 



RECORDER NO. 2 



TOE - UP AND 
DOWNSTREAM 



CREST - UP AND 
DOWNS"!'REAM 



fM\~^' 



TOE — VERTICAL ►■V.V thiMim mtti mMm 



CEST-VERTICAL ^f^^^^\',iiKf\,^^*f'ly*t^^ 
TOE - TRANSVERSE V-«<V; ••.-<•- 



ii 111 ' -/i 



CREST - TRANVERSE 





3T J/4iHMifeM8^?ifeli ^ 



jy^^rVM^'*^^''*'*'*'! ^ ^ 4 T Aj/%l^^^f^^ 



I 2 S 4 5 

I ! 1 1 I t 

TIME IN SECONDS 



O iq 



2g 



VERTICAL SCALE 

(ACCELERATION) 



MAGNITUDE 5.7 



Figure 110. Acceleration Records, Main Event of August I, 1975 



198 



Examination of aftershock records on re- 
corder No. 2 showed the space between 
two events to be about 1.5 centimetres 
(0.6 inch), the same length as the gap 
in the August 1 record. After August 8, 
the speed of the recorders was increased 
2-1/2 times. From then on, the space 
between events was about 2.5 centimetres 
(1 inch). Therefore, it was presumed 
that the gap in the August 1 records 
represented the distance the accelero- 
meter drum rolled after power had been 
cut off, and the time gap could not be 
indicated correctly by the time scale 
on the chart. 

The power failure was reenacted to find 
out how much time elapsed between main 
power cutoff and activation of the back- 
up power source. It was determined that 
the generator, which is the source for 
the back-up power supply, needed a mini- 
mum of 5 to 6 seconds to start and sup- 
ply power to the recorders once the main 
power supply was cut off. Therefore, 
the time gap in the main event record 
was set at 6 seconds. 

Recordings of stress for the August 1 
event were also marred by a gap due to 
the power loss. Before the gap, a maxi- 
mum vertical normal stress of 159 kilo- 
pascals (23 psi) was recorded by cell 
No. 5. Pore pressure cell No. 1 regis- 
tered a maximum pressure increase of 90 
kilopascals (13 psi), which was dissipat- 
ed during the 6-second gap. Pore- 
pressure cells 4, 5, and 6 also showed 
minor fluctuations, on the order of 14 
to 34 kilopascals (2 to 5 psi) . 

To gain an insight into what occurred 
during the time represented by missing 
portions of the DWR acceleration rec- 
ords, USGS recordings of accelerations 
at the seismic station and at the crest 
of Oroville Dam were obtained and com- 
pared with the corresponding DWR records. 
Unfortunately, the first few seconds of 
the USGS crest record were lost, as not- 
ed in California Division of Mines and 
Geology Special Report 124, "Oroville, 
California Earthquake, 1 August, 1975". 
However, the last portions of the DWR 



and USGS crest records are nearly ident- 
ical. Any differences are due to base- 
line and instrument corrections per- 
formed on the USGS record. The DWR rec- 
ord was not corrected. The last portion 
of the two crest records (following the 
gap) can be lined up as shown in Figure 
111. This leaves 2-1/2 seconds where 
the record is missing from both the USGS 
and DWR instruments. 

The record at the USGS seismograph sta- 
tion, 1.6 kilometres (1 mile) NW of Oro- 
ville Dam, was positioned so that its 
two highest peaks line up with the two 
high peaks recorded on the DWR base ac- 
celerometer. This positioning of the 
USGS base record shows that the strong 
base motion had essentially ceased by 
the start of the USGS recorded crest mo- 
tion. The USGS seismograph station and 
dam crest records were digitized for use 
in making analyses by computers. 

August 5, 1975 

As can be seen on Figure 112, the DWR 
record again has a vital part of the 
event missing and hence could not be used 
in any subsequent analysis. It can be 
seen, however, that like the August 1 
recorded motions, the dam is freely oscil- 
lating while the amplitudes of the accel- 
erations of the crest are decreasing in 
a typical decay curve patteim. This 
again occurs with the amplitudes of the 
recorded base motion being negligible. 

The USGS does not have records of the 
August 5 event. 

September 27, 1975 

The seismic event of September 27, 1975, 
(Magnitude 4.6) was recorded in its en- 
tirety on the DWR accelerometers. Figure 
113. These records were digitized for 
use in subsequent analyses. The digi- 
tized records were processed using the 
routine coii5)uter processing methods for 
strong-motion accelerograms developed at 
Cal Tech. Some changes, however, were 
made in this standard processing tech- 
nique. The instrimient correction was not 



15—78786 



199 



DAM CREST, DWR A-4 
N 46° E 




5.0 TOO 15,0 

TIME IN SECONDS 
Figure 111. Acceleration Records with Corrected Time Scales 
August 1, 1975 



200 



$^ 



TOE-UP AND DOWNSTREAM 



CREST-UP AND DOWNSTREAM 



'■'"'' ^" ' ^' * aJ^'^/;*^,/V^/V*vvA/^'^^w^ 



TOE - VERTICAL 







CREST-VERTICAL 
TOE - TRANSVERSE 

|%l^^li^^ W^ ^> ' i^^^l^ ^ y||||f^ ^ lH>(>^lWl|l Ml^ 

■K , ^ fl i CREST-TRANVERSE 



«w« 



7 5 10.0 

TIME IN SECONDS 



VERTICAL SCALE (ACCELERATION ) 



19:50 PST MAGNITUDE 4.7 



Figure 112. Acceleration Records, Event of August 5, 1975 



performed because the accelerometers are 
of a force-balance type. It was assumed 
that the instrument response was unaf- 
fected throughout the frequency range of 
interest . 

The records of the base and crest mo- 
tions (horizontal and vertical) were 
baseline corrected and put through an 
Ormsby filter to obtain equally spaced 
acceleration points between 1.4 and 48 
hertz. This filter bandwidth deviates 
from the standard filter used at Cal 
Tech, because of the high frequency con- 
tent, low amplitude, and short duration 
of the records. Acceleration-time his- 
tories plotted from the digitized rec- 
ords, along with corresponding response 
spectra, are in Appendix B. 

The USGS has no records of any seismic 
events for September 27, 1975. 



A maximum vertical normal stress of 62 
kilopascals (9 psi) was recorded by cell 
No. 5. Using methods which will be de- 
scribed later, the vertical normal stress 
computed for cell No. 5 location was 41 
to 62 kilopascals (6 to 9 psi). 

Observed Natural Period 

For both the August 1 and August 5 events, 
the dam continued to vibrate after the 
earthquake had stopped. As shown in Fig- 
ures 111 and 112, after the base accele- 
rations had dropped to less than O.Olg, 
long period crest accelerations continued 
for several seconds, starting at an amp- 
litude of about O.lg and decreasing in a 
typical decay curve pattern for free 
vibration. 

For the August 1 record, 5 or 6 success- 



201 



in ,Lw. u , TOE- UP AND DOWNSTREAM 

■ ' ' I'Mi! M i f| I I ^1 . (I . « . I*, /*, CREST-UP AND DOWNSTREAM 



niMli 



■Mi^v,H^J^m 





TOE- VERTICAL 



CREST-VERTICAL 



TOE-TRANSVERSE 



CREST-TRANVERSE 






TIME IN SECONDS 
0.2g 0.3g 



VERTICAL SCALE (ACCELERATION) 



14:35 PST MAGNITUDE 4.6 



Figure 113. Acceleration Records, Event of September 27, 1975 



ive cycles have periods close to 0.8 
seconds. For the August 5 record, there 
are 3 or 4 cycles in the decay curve with 
a period of about 0.7 seconds. Accelera- 
tion response spectra for the August 1 
USGS crest record show a predominant 
period of 0.8 seconds (Figure 114). 
Response spectra were not calculated for 
the August 5 event, because the extreme 
overlap of adjacent records made accele- 
rations indistinguishable for the strong 
motion portions. The September 27 event 
was not used for estimating period be- 



cause it did not develop a clear decay 
curve pattern for free vibration. 

Since these observed periods are for free 
vibration conditions, they are the natu- 
ral periods of the dam; and since the 
fundamental period is known to be domi- 
nant in an earth dam, the observed per- 
iods are presumed to be the fundamental 
periods. Thus the fundamental natural 
period is determined to be 0.8 seconds 
for the intensity of shaking produced by 
the August 1 main shock. 



202 



o 



I- 
< 
cr 

LU 

_l 
UJ 

u 
o 
< 



0.80I- 



0.60 



0.40 



0.20 - 



AUG. I, 1975 
U.S.G.S. RECORD 
AT DAM CREST 



NOTE: 



BASED ON RECORDED 
ACCELERATIONS DURING 
FREE VIBRATION ONLY 




0.50 1.00 1.50 

PERIOD IN SECONDS 



2.00 



Figure ]\h. Acceleration Response Spectra for Crest 
Motions, Event of August 1, 1975 



ANALYSIS OF STATIC STRESSES BY FINITE ELEMENT METHOD 



General 

The behavior of an embankment dam sub- 
jected to dynamic loading by an earth- 
quake is significantly influenced by the 
stress condition existing in the embank- 
ment prior to the earthquake. Current 
methods of analysis for evaluating the 
seismic stability and permanent deforma- 
tions require knowledge of the static 
stress distribution for the maximum sec- 
tion of Oroville Dam. These static 
stresses can best be calculated by the 
finite element method, which permits the 
evaluation of stresses and deformations 
in an embankment through a series of 
steps or increments to simulate construc- 
tion and reservoir filling. The follow- 
ing sequence was used in this analysis: 



1. Construction of the core block in 
four layers. 

2. Construction of the cofferdam, up- 
stream of the core block, in 14 
layers. 

3. Construction of the remaining embank- 
ment in 27 layers. 

4. Application of water load in four 
stages, simulating filling of the 
reservoir. 

The finite-element representation of 
Oroville Dam is shown on Figure 115. 
This mesh, used in the static and dyna- 
mic analyses, contains 564 elements and 
585 nodes. 



203 



564 ELEMENTS 
585 NODES 




Figure 115- Finite Element Mesh, Maximum Section Oroville Dam 



Material Properties 

The success of finite-element analyses 
to model the behavior of an earth dam 
depends in a large part on how well the 
nonlinear response of soil and rock ma- 
terials under load can be described ana- 
lytically. Because of the good compari- 
son between observed and computed settle- 
ments in a previous analysis by Kulhawy 



and Duncan (1970), the same stress-straii 
parameters were used in this analysis. 
The difference between the parameters fo; 
the transition (Zone 2) and shell (Zone 
3) materials, shown in Table 5 is neglig- 
ible. Therefore, Zone 3 parameters were 
used for both Zones 2 and 3 in all the 
finite-element-method analyses. 



Table 5 



Values of Stress-Strain Parameters 

for Analysis of Oroville Dam 

(From Kulhawy and Duncan) 



Parameter 



Values Used in Analyses 



Symbol 



Shell 



Transi- 
tion 



Core 



Soft , 
Clay^^ 



Concret 



Unit weight (lb/ ft ) 

2 
Cohesion (tons/ft ) 

Friction angle (degrees) 

Modulus number 

Modulus exponent 

Failure ratio 

Poisson's 

ratio 

parameters 



150 

43.5 
3780 
0.19 
0.76 
0.43 
0.19 
14.8 



150 

43.5 
3350 
0.19 
0.76 
0.43 
0.19 
14.8 



150 
1.32-^/ 
25.1^/ 



345 
0.76 
0.88 
0.30 
-0.05 
3.83 



a/ 

— Zone of soft clay at upstream end of core block. 

hi — ? 

— c and i for (a, +0,) <50 tsf; (c = 10.2 tons/ft ^ 

^ for (a + a ) >50 

c/ Z 

— Tensile strength of concrete = 14 tons/ft (200 psi) . 



= 4»; 
tsf) 



125 
0.3 
13.0 

150 

1.0 

0.9 

0.49 





162 
216^/ 





137,500 



1.0 

0.15 







204 



Static Stress Analysis 

Computer program ISBILD was used to car- 
ry out the static-stress analysis. This 
program is similar to the one used in 
the earlier analysis of Oroville Dam by 
Kulhawy and Duncan. The major differ- 
ence is the type of element used. Kul- 
hawy and Duncan used a quadrilateral ele- 
ment divided into two triangles. Within 
each triangle the strains vary linearly; 
then, for compatibility reasons, the 
strain along the sides of the quadrila- 
teral element is kept constant. 

Program ISBILD uses a quadrilateral in- 
compatible isoparametric element. This 
means that in addition to 8 regular de- 
grees of freedom at 4 nodes, the element 
has 4 internal degrees of freedom to im- 
prove its bending behavior. These addi- 
tional nodes of displacement, in general, 
make the elements incompatible at the 
interelement boundaries. 
Seepage Forces 

Reservoir effects are simulated by consi- 



dering the water load in two parts: 
total stress forces and water pressure 
forces. To account for the effects of 
the seepage forces in the core, piezo- 
meter readings were used as input to the 
computer program NODALFOR (developed by 
Division of Safety of Dams). This pro- 
gram uses the water pressures at nodes 
and computes the forces at the sides of 
elements due to these pressures. The 
sum of these side forces is the result- 
ant water force on the element. Result- 
ant water forces are then distributed 
to element nodes in proportion to the 
contributing area of each node. The 
values distributed to a node from adja- 
cent elements are added to yield the net 
water force at the node. This net water 
force is added to the total soil force 
(based on saturated unit weight) at the 
node to get the effective soil force. 

Table 6 shows the comparison between the 
measured and calculated static stress 
values . 



Table 6 





Static Stress Comparison 




Direction 
of Stress 


Compressive Stress (tsf) 


Cell No.* 


Maihak Cell | FEM Analysis 



1 

2 

3 

4 

5 

6 

7 

8 

9 

10 

11 

12 

13 

14 

15 



13.8 
29.5 
14.4 



30, 
15. 
22, 



11.0 

7.9 
11.5 
25.4 
18.0 
10.0 




16.6 
28.2 
11.1 
23.0 
25.6 
14.2 
21.0 
24.5 
20.8 
18.9 
12.1 
20.2 
17.0 
9.1 
17.1 



*Note: All but three of the 
static stress cells 
were functioning during 
the 1975 earthquakes. 



205 



There is good agreement between measured 
and calculated vertical stresses. In- 
clined stresses, compression toward the 
downstream toe, also show good agreement. 
However, the two operable cells measur- 
ing compression toward the upstream toe 
measure only about half the calculated 
stresses. 

The computed static stresses in the 



shells are compared with results by 
Nobari and Duncan (1972) in Figures 116 
through 119. 

Stress comparisons are not valid in the 
core because Nobari and Duncan present 
total stresses, where the stresses 
computed in this study are effective 
stresses. Plots of stresses in all ele- 
ments are included in Appendix C. 




Figure 116. Contours of Effective Maximum Principal Stress in Oroville Dan 
Ful 1 Reservoi r 



CONTOURS ARE IN fsf 
DEPARTMENT OF WATER 
RESOURCES RESULTS 
NOBARI AND DUNCAN RESULT 




Figure 117. Contours of Effective Minimum Principal Stress in Oroville Dam 
Ful 1 Reservoi r ' 



206 



CONTOURS ARE IN T. S. F. 




Figure 118. Contours of Maximum Shear Stress in Oroville Dam, Full Reservoir 




DEPARTMENT OF WATER RESOURCES RESULTS _i_NOBARI AND DUNCAN RESULTS 

I 

NOT TO SCALE 



Figure 119. Orientation of Principal Stresses 



207 



5. DETERMINATION OF DYNAMIC SHEAR MODULUS AND 
DAMPING VALUES FOR EMBANKMENT SHELL MATERIAL 



General 

Dynamic shear modulus values at low 
strains can be measured in the field us- 
ing geophysical methods or in the labor- 
atory using vibration tests. Values at 
higher strains can be measured in the 
laboratory using cyclic shear tests. If 
recorded motions during an earthquake 
are available, calculations can be made 
to determine the modulus corresponding 
to those recorded motions. Damping is 
usually measured in the laboratory dur- 
ing the same test used to measure the 
modulus. 

Measurement of dynamic shear modulus and 
damping for the Oroville Dam shell mate- 
rial is a difficult task: Field mea- 
surements of shear wave velocity would 
require deep borings in the gravel and 
cobbles; undisturbed samples for labora- 
tory tests would be next to impossible; 
remolded samples for laboratory tests 
cannot reproduce field conditions of 
particle size, stress, time of loading, 
or variability of grading and compaction 
30000 



10000 



3000 



1000 



in the shell. In spite of the difficul-' 
ties, a considerable effort was made to 
determine the modulus and damping becaus( 
they are the most important parameters 
controlling the response of the dam to 
an earthquake . 

Studies by Seed and Idriss (1970) have 
shown that the dynamic shear modulus of 
granular soils can be related to the 
effective mean normal stress as follows: 

G = 1000 K2 (o'm)-'-''^ 
G = dynamic shear modulus in psf 
K„ = a parameter relating G and a'm 
am = effective mean normal stress 
in psf 

K„ is a function of strain level and 



void ratio. K. 
K„, is obtaine 

(io-^%). 



i"*!^' 



the maximum value of 
low shear strain 



For clays. Seed and Idriss found that 
shear modulus could be related to static 
undrained shear strength and dynamic 
shear strain as shown in Figure 120. 




300 



100 



10"* 



10-3 



10-2 10-' I 

SHEAR STRAIN % 
AFTER SEED AND IDRISS (1970) 

Figure 120. In-Situ Shear Moduli for Saturated Clays 



10 



208 



It was presumed that the gravel shells 
would dominate the response behavior 
during earthquake shaking because they 
occupy about 90 percent of the embank- 
ment. Therefore, the testing and anal- 
ysis for shear modulus were done only 
for the gravel shell material. The 
core modulus and damping were assumed 
equal to published values for clays. 

Two methods were used to determine the 
modulus of the gravel shell material : 
cyclic triaxial tests on remolded sam- 
ples, and analysis of recorded embank- 
ment response during the Oroville earth- 
quakes of 1975. Damping values for the 
gravel material were estimated from the 
cyclic triaxial tests. 



Cyclic Triaxial Tests 

Laboratory test data to define the dyna- 
mic properties of gravel material are 
very limited. Presently available cyc- 
lic test equipment can test specimens 
up to 30 centimetres (12 inches) dia- 
meter. The average gradation of the 
Oroville shell material has a maximum 
particle size of 15 centimetres (6 
inches) and would require a specimen 
diameter of 90 centimetres (36 inches). 

Comparison studies by Becker, (1972) 
have shown the same static strength for 
samples with modeled gradation of 5 cen- 
timetre (2-inch) and samples with field 
gradation of 15 centimetre (6-inch) 



U.S. STANDARD SIEVE SIZES 



"200 '100 
100 



90 



'50 "30 



'16 



3/8" 3/4" I 1/2" 



70 



50 



40 



liJ 

O 

(T 30 

kJ 

Q. 



20 



1 




1 


1 






1 


1 




/ 


1 / 


/ 




















/ 


// 


f 


















p 




// 








MODELED GRADATION FOR 




/ 




f/ 








CYCLIC TRIAXIAL TESTS 

1 1 1 




^ 


1 


/ 




















/ 


// 
























/ 








AVERAGE FOR 24 TEST PITS 
IN EMBANKMENT I964-I96€ 

1 1 1 








1 

r 






5 




















I 


7 






















r 
























li 
























ft 




















/ 




1 




















/ 




// 




















/ 


/ 


? 


















/ 




A 


■-— PRC 


1 
)JECT AVERAGE 












/ 


y^ 




1 i 1 J 














^^ 


:::.=^_FIELD GRADATIONS 














^ 


OROVILLE DAM, ZONE 3 


























1 


<^^ 


1 


1 






1 


1 




i 


1 





0.1 



0.5 



1.0 5.0 10. 

GRAIN SIZE IN MILLIMETERS 



50.0 100.0 



Figure 121. Sample Gradation for Cyclic Triaxial Tests 



209 



maximum particle size. It seems reason- 
able to extend this kind of modeling to 
cyclic testing. 

A general study was undertaken by Wong 
(1973) to determine the cyclic strength, 
dynamic shear modulus, and damping of 
gravels. Strain controlled cyclic tri- 
axial tests of 30 centimetre (12-inch) 
diameter samples of modeled Oroville 
gravel gradation were part of that 
study. Samples tested were composed of 
Oroville gravel with 5 centimetre (2- 
inch) maximum particle size and a grad- 



ation curve parallel to the average 
shell grading (Figure 121). 

The sample density used was about 2 430 
kilograms per cubic metre (152 pounds 
per cubic foot) . Average density of 
Zone 3 is 2 360 kilograms per cubic 
metre (148 pcf ) . All samples were iso- 
tropically consolidated with a pressure 
of 196 kilopascals (4,096 pounds per 
square foot). 

Figure 122 shows the results of these 
cyclic tests compared with the results 




10"^ 10" 
SHEAR STRAIN % 



•0-- FROM FIELD SHEAR WAVE VELOCITY MEASUREMENTS, 
SEED AND IDRISS 1970. 

• TRIAXIAL TESTS ON MODELED OROVILLE GRAVEL BY WONG 

MAXIMUM PARTICLE SIZE = 2 INCHES 

SAMPLE DIAMETER = 12 INCHES 

0-3C = 410 psf 

RELATIVE DENSITY = 100% 

SET UP DENSITY = 152 pcf 

N = 5 CYCLES 

■ FROM DYNAMIC ANALYSIS OF EMBANKMENT FOR 1975 

RECORDED EARTH QUAKES ( PERI OD = ONE SECOND) 

Figure 122. Modulus Determinations for Gravelly Soils 



210 




» DATA FOR OROVILLE MODEL GRADATION, THIS INVESTIGATION 

AVERAGE VALUE FOR SANDS, FROM SEED AND IDRISS, 1970 

— UPPER AND LOWER BOUNDS FOR SANDS 



Figure 123. Comparison of Damping Ratios for Gravelly Soils and Sands 



reported by Seed and Idriss. The Seed- 

Idriss curves are based on field shear 

wave velocity measurements for K„ , 
„ /„ , . 2max 
and the average K»/K„ reduction curve 
r J IT • 2^,2max , 
for sands. Usxng the same reduction 

curve to fit the triaxial test data 

gives a K„ value of 130. 

Figure 123 shows the damping results 
compared with the range of values for 
sands reported by Seed and Idriss. The 
test data points are very close to the 
average curve for sands. 

Analysis of Recorded Embankment Response 
During the 1975 Earthquakes 

General 



Even though there were gaps in the rec- 
ords of the stronger shocks, the 1975 
Oroville earthquakes afforded an oppor- 
tunity for analyzing embankment response 
to determine the dynamic shear modulus 
of the embankment materials. Accelera- 
tion time histories were recorded at 



both the base and the crest of the dam 
for several events. The August 1 main 
shock provided the best definition of 
the natural period of the embankment - 
the key to the analysis. 

The natural period is dependent upon 
stiffness (shear modulus) and mass dis- 
tribution. Knowing the mass distribu- 
tion and the period allows calculation 
of the shear modulus. Knowing the shear 
modulus and embankment strain allows 
calculation of shear modulus parameter 

K„ 
2max 

Computer program QUAD4 was used in the 
analyses. It determines the natural 
period by solution of the eigenvalue 
problem: 

[K] =w^ [M] (5-2) 
[K] = system stiffness matrix 
[M] = system mass matrix 
w = natural circular frequency 

The stiffness matrix is developed from 



211 



UJ 300 1- 



CHORD LENGTH = 5150 ft. 



CREST ELEV. 281.0m (922ft.) 
WITHOUT CAMBER 




O.G.L 



ESTIMATED DAM 
FOUNDATION STRIPPING LINE 



30*00 40*00 50*00 

STATIONS (100 ft.) 



DIVERSION TUNNEL NO. I 

DIVERSION TUNNEL NO. 2 - 

POWER PLANT ACCESS TUNNEL-i 

PALERMO OUTLET WORKS 



Figure \2h. Section on Long Chord of Dam Axis 



the shear modulus values for all the ele- 
ments in the maximum section FEM mesh 
— and is therefore a function of shear 
modulus parameter K„ . 

The Oroville embankment is located in a 
triangular canyon and has a crest length 
to maximum height ratio of approximately 
7. A longitudinal profile is shown in 
Figure 124. 

A three-dimensional (3D) analysis of the 
dam to determine K„ of the gravel 
shell material is not possible, because 
present computer capabilities are inade- 
quate for dams as large as Oroville. 
The solution to this problem is to find 
the appropriate values of natural period 
and shear strain to use in a two- 
dimensional (2D) analysis. 

Makdisi (1976) has derived a relation- 
ship between the natural periods comput- 
ed by 3D and 2D analyses for a 30-metre 
(100-foot) high dam with slopes of 2:1 
and a constant shear modulus throughout. 
Assuming this relationship is valid for 
the much higher Oroville Dam with vari- 
able shear modulus as defined by equa- 
tion 5-1, the 2D period can be computed 
corresponding to the observed period in 
the August 1, 1975 earthquake. 

The same value of shear modulus, G, was 



used in the 2D and 3D analyses in deriv- 
ing the natural period relationship. 
Since the parameter to be calculated, 
G , would have to be the same for 2D 
and 3D, then G/G and consequently the 
shear strain are also the same for the 
2D and 3D cases. Thus the same strain 
and modulus reduction factor must be 
used for the 2D and 3D analyses when 
applying the natural period relationship 



1 



TRIANGULAR CANYON 



H = I00 FEET 
SLOPES = 2:| 
Vs = 500 fps 
AFTER MAKDISI, 1976 




OROVILLE DAM L/H=7 
I 35 



Figure 125. Comparison of Natural Period; 
for Two- Di mens ional and Three-Dimens ioni 
Embankment in Triangular Canyon 



212 



It is important to note that this rela- 
tionship would not give the correct per- 
iod for a very long (2D) dam subjected 
to the August 1, 1975 earthquake shaking. 
But it does give a correct relationship 
among period, strain, and K„ , allow- 
ing the calculation of K„ 

2max 

Natural Period for Two-Dimensional 
Analysis 

Makdisi's correlation of 2D and 3D per- 
iods only goes to an L/H of 6. Extra- 
polating to an L/H of 7 gives a period 
ratio of 1.25 to 1.35 (Figure 125). The 
natural period of Oroville Dam Embank- 
ment in the August 1, 1975 earthquake 
was estimated to be 0.8 seconds (see 
page ) . The natural period range to 
be used in the 2D analysis is 1.0 to 
1.1 seconds. 

Shear Strain for Two-Dimensional 
Analysis 

The maximum displacement computed from 
the August 1 recorded (USGS) crest mo- 
tion is 1.5 centimetres (0.6 inch). The 
assumption is made that the peak shear 

1.0 



strain (y max) is constant throughout 
the dam and is equal to maximum crest 
displacement (D) divided by embankment 
height (H) . 

Y max = I X 100% = 0.007% 

A ratio of average strain to peak strain 
of 0.6 has been used in this study. 
Thus the average shear strain is about 
0.004 percent for the actual (3D) dam, 
which is also the value to use in the 
Two Dimensional Analysis. 

Shear Modulus Reduction Factor 

The average shear modulus reduction 
curve for sands shown in Figure 126 was 
used for the gravel shells. At a shear 
strain level of 0.004 percent, the modu- 
lus reduction factor is 0.86. This val- 
ue agrees well with the results of a 
dynamic response analysis for the Aug- 
ust 1, 1975 earthquake, which will be 
discussed later in this chapter. The 
range of reduction factors in that anal- 
ysis was generally only 0.78 to 0.88. 
An average value of 0.85 was used in the 

calculation of K„ 

2max 



2 
^ .4 



^ 


*^ 


^ 


^ 














c 


ORE- 


K 


' ^ 


\i 


^ 


















N 

) 


V 




^^ 














N 


s^ 


V 


-SHE 






LL 


* ■i.)i. 












\ 


^ 


^ 


fc=* 


^==t\ 



I0'2 10"' 

SHEAR STRAIN % 



-D- AVERAGE REDUCTION CURVE SANDS, FROM SEED AND IDRISS 

-O- AVERAGE REDUCTION CURVE CLAYS, FROM SEED AND I DR ISS 

-z^ REDUCTION CURVE CLAYS, Su = 8000 psf FROM SEED 

(PRIVATE COMMUNICATION) 



Figure 126, 
Soi Is 



Shear Modulus Reduction Curve for Embankment 



213 



(J) 

LlI 



20 



< — 



^ 


^ 


A 


s 







80 



20 40 60 

NORMAL STRESS ( tsf ) 

Figure 127. Static Shear Strength Envelopes for Core 
Material 



^2max 



vs. Natural Period 



Computer program QUAD4 was used to cal- 
culate natural periods for the follow- 
ing input properties and conditions : 

A. Finite element mesh of Maximum Sec- 
tion (Figure 115) 

B. Static stresses determined from 
static FEM analysis 

C. Shell and Core Material Densities 
used in static FEM analysis 

D. Core shear modulus - three trial 

values of G /S - 1100, 2200, 
4400 "'^^ " 

- CU shear strength envelope 
(Figure 127) 

E. Shell K^jj^^^ _ jQ^^ ^^^^^ values - 

100, 200, 300, 400 

F. Shear Modulus Reduction Factor for 
both shell and core - 0.85 

The results are plotted in Figure 128 

as a family of curves of K„ vs. nat- 

. , ^ , . 2max . . - 
ural period. Each curve is for a dif- 
ferent value of core modulus. Varying 
the core modulus by a factor of four 
changes the shell K- by 20 percent 
to 25 percent. 

Range of K„ - The range of K„ is 

135 to 210 for a natural period from 1.0 
second to 1.1 second and core Gmax/Su of 
1100 to 4400. 



Comparison of Observed and Computed 
Crest Motions 

Ideally, the value found for shell 
^2max ^°"-'-'^ ^^ used in a three- 
dimensional dynamic analysis using the 
recorded base motion. A comparison of 
the computed crest motion with recorded 
crest motion would then be a direct 
check on the mathematical model and 
material properties used. However, a 
three-dimensional analysis is not pres- 
ently possible for Oroville Dam, and a 



"1 1 1 1 1 1 1 1 I I r 




= 85 

I I 

▲ Gm»x /S„ = 1100 

• Gm«k /S^ - 2200 

■ G„„ /Su = 4400 

FROM CU STRENGTH ENVELOPE 
I I I 



NATURAL PERIOD (SECONDS) 



Figure 128. Ko vs. Natural Period 
^ 2max 



214 



two-dimensional analysis woulci give a 
different response acceleration as shown 
by Makdisi (Figure 129) . His study was 
done for a 30 metre (100 feet) high dam, 
Taft earthquake, and L/H =3. In some 
locations particularly at the crest, the 
difference is large — up to 100 percent. 



This is because the restraint or stiff- 
ening effect of the abutments is ignored 
in a plane strain analysis. Therefore, 
the two-dimensional analysis needs to be 
modified to account for the abutment 
restraint. 



L/H = 3 

Vg = 500 fpS 





0.98g 


p -- 0.10 






0.54g 




o 
o 


0.66g ^^\ 0.53^3-D 

40g" •^^v^.29^ Plane Strain 






0.229 , 0.22g ^^~^^ 
O.I6g ' 0.2lg 


O.2I9 
*^v^22g 



///mw<^. 



0.2g 
MID-SECTION 



0.39g 





0.53g 










0.32g 




0.3g 






10 


0.43g 




*"~-^.36g 






O.I9g 
0.2! g 


1 


O.I9g^^^^^ 
*0.2lg 


O.ISg 
\^2g 






/U^//A^\.\ 












0.2g 








QUARTER SECTION 





Figure 129. Maximum Accelerations Computed from 3D and Plane Strain 
Analyses Using Base Motions from Taft Record (After Makdisi) 



215 



Embankment Response Model 

The Embankment Response Model is a two- 
dimensional analysis with a modified 

Ko (pseudo K„ ) value that gives 

2max '^ , ^ 2max . , , ° ^, „ 
a computed natural perxod equal to the 

observed period (3D) for the actual em- 
bankment during the August 1, 1975 
earthquake. It is an implied assumption 
that the same adjustment which accounts 
for the three-dimensional effects on 
natural period will also account for the 
three-dimensional effects on acceleration 
and strain. The pseudo K^ value is 
higher than the K™ represeucing sheat 
modulus because the stiffening or re- 
straint effects of the abutments are 
included. 

By definition, the model applies to the 
August 1, 1975 earthquake. This same 
model may apply to other, stronger. 



earthquakes, but the 1975 earthquake 
series did not provide enough informa- 
tion to test this question. 

The value of pseudo K^ corresponding 
to the observed natural period of 0.8 
sec. is 285 to 370 (Figure 128). 

August 1 Event 

The acceleration record at the toe of 
the dam was not usable because of the 
six-second gap. The bedrock record froit 
seismograph station DRV was used, even 
though it is a mile from the dam, 275 
metres (900 feet) higher in elevation, 
and oriented 9 degrees different than 
the dam instruments. Where both records 
are present, accelerations are similar 
for seismograph station DRV and the toe 
of the dam. Other input was as follows; 



Computer Program LUSH 

Highest Frequency Used 8 Hertz 

Shell - Pseudo K^ 350 

- Average Modulus Reduction Curve for Sands 
(Figure 126) 

- Average Damping Curve for Sands (Figure 130) 

Core - G /S 1750 

- Shear Strength Envelope UU 

- Higher Modulus Reduction Curve for Clays (Figure 126) 

- Average Damping Curve for Clays (Figure 130) 



Poisson's Ratio 



0.3 



The comparison between the computed and 
observed crest accelerations is shown in 
Figure 131 along with the input bedrock 
motion. Comparisons of displacement 
time history and acceleration response 
spectra are shown in Figure 132. The 
shapes and magnitudes of the computed 
patterns are generally similar to those 
observed. The response spectra of the 
computed crest motion show the dam to 
oscillate with two distinct periods. 
The first period of 0.15 seconds is not 



evident in the response spectrum of the 
observed crest motion, probably because 
the period of 0.15 second corresponds 
to the forced vibrations during base 
shaking, which is the missing portion 
of the crest record. The second period 
shown in the response spectrum of the 
computed crest motion occurs at 0. 75 
seconds, slightly different from the 
period of 0.8 seconds for the recorded 
motion. This may mean that 350 is too 
high a choice for pseudo K„ , or that 



216 



35 



30 



15 



10 





















/' 


















A 


Y 














t 


X 


/ 












SHEL 


L ' 

r 


/ 


/ 
















/ 


/ 


/ 

•—CORE 

1 










X 


^ 


I f 


r 








C-l ( 


f=^ 


F^ 


r^ 


> 












.} > 





10" 



10-^ I0-' 

SHEAR STRAIN % 



-O- AVERAGE DAMPING CURVE SANDS, FROM SEED AND IDRISS 
-O- AVERAGE DAMPING CURVE CLAYS, FROM SEED AND IDRISS 

Figure 130. Damping Ratios for Embankment Soils 



the frequency content was different for 
the bedrock motions at the base of the 
dam and at seismograph station ORV. 

September 27 Event 

The bedrock motion of the September 27 
event recorded at the dam toe was used 
as input to the dynamic finite element 
model. Other input was the same as for 
analysis of the August 1 event, except 
the highest frequency used for LUSH was 
16 Hertz instead of 8 Hertz. 

Response was computed at the crest and 
at 74 metre (240 foot) depth. Compari- 
sons of computed and observed accelera- 
tion time histories and response spectra 
are shown in Figure 133. 

Accelerations at the crest are much 
smaller than in the August 1 main shock, 
and the duration of shaking is much 
shorter. The response is all at high 



frequency — 2 to 10 hertz. Either the 
shaking is not strong enough to excite 
the dam into definitive free vibration 
motion, or the fundamental period is 
much smaller for the September 27 event. 

Dynamic Properties Adopted for the 
Gravel Shell 

The shear modulus parameter, K„ , was 
determined for the gravel shell ^y 
cyclic triaxial tests and analysis of 
embankment response to the August 1, 
1975 earthquake. Both methods have 
serious limitations including: 

1. Remolded samples cannot faithfully 
model variations in gradation and 
compaction in the embankment. 

2. The triaxial test does not correctly 
model the stresses in the embankment. 

3. Shear strain is assumed constant 



217 



throughout the dam in the analysis 
of observed response. 

4. Makdisi's correlation for natural 
periods was developed for a 30 metre 
(100 foot) high dam with constant 
shear modulus and daiQping throughout. 
This correlation was assumed applic- 
able to Oroville Dam with a height of 
230 metres (750 feet) and a shear 
modulus that varies throughout. 

The two different methods gave different 



answers. However, these results bracket 
published values for dense gravels 
(Figure 122). Therefore, it was decided 
to use two values, 130 and 205, to rep- 
resent the range. 

Because the damping results determined 
in the cyclic triaxial tests agreed so 
well with the published average values 
for sands, it was also decided to use an 
approximation of the average damping 
curve in computing dynamic stresses 
generated by the Reanalysis Earthquake. 



COMPUTED RESPONSE AT CREST 
(NODE 4) 

SHELL PSEUDO Kjmax^^SO 
CORE Gmax/ Sy ^ '"^50 




A/Vr^ ^ \jJ\A/\p^ ' A.y^'\j\ 



STARTING TIME CHOSEN TO GIVE BEST 

MATCH OF OBSERVED ANDCOMPUTED RESPONSE 




'|VAy^\K^"VV^\''^'\^v ^ " ' .V ..-. . vV - . .^ ..- vJ\ . 



INPUT AT BASE 
(USGS) 

SEISMOGRAPH STATION ORV 
( I MILE FROM DAM ) 



O-OO 1.00 2.00 3.00 q.M 5.0O S.OO 7.00 8.0O 9.00 10.00 11.00 12. TO 

TIME [SECONDS! 

Figure 131. Comparison of Acceleration Time Histories, August 1 Main Shock 



218 




FROM OBSERVED CREST ACCELERATIONS 
FROM COMPUTED CREST ACCELERATIONS 

COMPUTER PROGRAM LUSH 
SHELL PSEUDO Kjmax-SSO 
CORE Gmax/Su=I750 



CURVE FOR OBSERVED DISPLACEMENTS LOCATED 

TO GIVE BEST MATCH WITH COMPUTED DISPLACEMENTS 



ACCELERATION RESPONSE 
SPECTRA FOR CREST MOTIONS 



ACCELERATION RESPONSE SPECTRA FOR 
BASE MOTION USED IN LUSH ANALYSIS 



OBSERVED 



COMPUTED 

5 % DAMPING 




'^.oo 



5% DAMPING 




'^.oo 



Figure 132. Comparison of Displacement Time Histories and Acceleration Response 
Spectra for Crest Motions, August 1 Main Shock 



219 



COMPUTED 



0.10- 




ACCELERATiON RESPONSE SPECTRA 



»- 1.0 2.0 5.0 4.0 

c TIME IN SECONDC 

ui 



OBSERVED 




-O.IOt 



O.IO 



o 



OB SERVED 
COMPUTED 




1.00 

PERIOD IN SECONDS 



!.0 2.0 3.0 4.0 

TIME IN SECONDS 

NODE 4 ICREST) ELEVATION 922 FEET 



COMPUTED 



ACCELERATION RESPONSE SPECTRA 



^ -.rMj\J^Jy\^ 



-o.;o| 



1.0 2.0 3.0 4.0 

TIME IN SECONDS 



OBSERVED 




Z 0.40- 

O 

I- 

< , 

Q^ 0.20- 

\il 

_l 

liJ 

o 

< 



OBSERVED 
COMPU'^ED 



5% DAMPING 



0.50 



1.00 



PERIOD IN SECONDS 



220 



1.0 2.0 3.0 4 

TIME IN SECONDS 

NODE 88 ELEVATION 680 FEET 
Figure 133. Comparison of Acceleration Time Histories and Response Spectra, 
September 27 Aftershocks 



REANALYSIS EARTHQUAKE 



On August 1, 1975, a Magnitude 5.7 earth- 
quake occurred approximately 12 kilo- 
metres (7.5 miles) southwest of Oroville 
Dam. The associated surface cracking, 
traced to within 5 kilometres (3.1 miles) 
of the dam, revealed a previously uniden- 
tified "active" fault (see Figure 134) . 
For a more detailed discussion, refer to 



Chapters II, III, AND IV of this bulle- 
tin, which describe geological and 
seismological studies as well as vert- 
ical and horizontal geodesy. 

Historically, other local events include 
the following earthquakes: 



Richter Magnitude 

5.7 
4.7 
4.7 
4.7 



Date 

February 8, 1940 
May 24, 1966 
April 29, 1968 
August 1, 1975 



Location from Dam 

50 km (31 miles) north 

37 km (23 miles) northwest 

48 km (30 miles) west 

14 km (9 miles) southwest 



In addition, other known faults and 
maximum credible earthqxiakes are as 
follows : 

Fault Richter Mag. 

San Andreas 8.5 

Honey Lake 7 . 5 

Mohawk Valley 
Bear Mountains - Melones 



Distance from Dam 

195 km (122 miles) 

117 km (73 miles) 

72 km (45 miles) 

58 km (36 miles) 



See Figure 135 for fault locations. 

Based on the hypocenter locations of the 
August 1 main shock and the subsequent 
aftershocks, the causative fault was de- 
fined as dipping to the west from the 
ground surface cracking. This fault 
system is presumed to extend northward 
beyond the limit of identified surface 
cracking. Thus, as illustrated in 
Figure 136, at depth it would pass 
directly under Oroville Dam. 

The main shock hypocenter was about 9 
kilometres (5.5 miles) deep; the after- 
shock hypocenters were 3 to 8 kilometres 
(2 to 5 miles) deep. It is assumed for 
purposes of developing Reanalysis Earth- 
quake motions, that for an earthquake 
larger than magnitude 5.7, the hypocen- 
ter would be 5 kilometres (3 miles) from 
the base of the dam. 



In view of the 1975 earthquake activity, 
the Consulting Board for Earthquake 
Analysis and the Special Consulting 
Board for the Oroville Earthquake rec- 
ommended the following: 

A. "In view of the developments, it is 
appropriate to consider that earth- 
quakes ranging up to magnitude 6.5 
may occur within a few miles of the 
dam site." 

B. "The Board considers that an appro- 
priate earthquake motion for reeval- 
uation of structures critical to pub- 
lic safety in the Oroville-Thermalito 
complex would be one producing a peak 
acceleration of 0.6g and having char- 
acteristics similar to those developed 
near Pacoima dam during the San 



221 




Photo lineoment 

•• ■ Probable Fault 

— V~^ Fault, dip indicoted 
if known 

• Epicenters 



s 

l_l_L 



1 1 1 


MILES 
5 

1 


10 



KILOMETRES 



Figure 13^. Lineaments, Faults and Recorded Epicenters Around Oroville 



222 




FAULT 

INFERRED FAULT 



I I I I — I — u. 



MILES KILOMETRES 

Figure 135. Location of Faults in Relation to Oroville Dam 



223 



SURFACE CRACKINGn 
^^^'^ OROVILLE DAM — -^'"^^ >*L^^^ | 






/ i 






/ 






/ 




+ 

4- 


/ 


+ 


4 4 

+ 


. / 


+ 

4 
+ 


+ 

+ 


/ + + 


4 

4 4 + 


4 

4- 


/ 


4^ -% 

4 

4 


/ 


FAULT PLANE ASSUMED 


/^ 


TO GO THROUGH SUR- 


HYPOCENTER FOR MAIN SHOCK 


r^ 


FACE CRACKING AND 
HYPOCENTER OF 


DATE, AUG-I , 1975, TIME 13^20 PDT 


/ 


MAIN SHOCK 


MAGNITUDE 5.7, DEPTH 8.8 ""^--^ 




KILOMETRES ^ ^ 


\ 




/ 







CROSS SECTION AT MAIN SHOCK HYPOCENTER 
PROJECTED NORTHWARD 10 KILOMETRES 

4- DENOTES ESTIMATED HYPOCENTER FOR EARLY AFTERSHOCKS, 
AUGUST I THRU 7, 1975 



I 1/2 
I I I I 1_ 



_1 I 



MILES KILOMETRES 

SECTION BEARING EAST-WEST 

Figure 136. Relationship of Oroville Dam to Assumed Northward Extension of 
Faul t 



224 



Fernando earthquake of February 9, 
1971. The time-history of such a 
motion should be obtained from a 
modified form of the Pacoima dam 
record, as discussed in the "Report 
of the Consulting Board for Earth- 
quake Analysis" dated May 22, 1973. 
The actual time-history could be the 
same as that forwarded to Mr. Jansen 
by Clarence R. Allen with his letter 
of January 16, 1974, except that the 
duration of shaking should be limit- 
ed to the first 20 seconds of the 
record provided, and all ordinates of 
the record should be multiplied by a 
suitable scaling factor to give a 
peak acceleration of 0.6g. 

"In addition the structures should 
be checked for the motions produced 
by the following earthquakes : 



(a) a magnitude 8.5 earthquake oc- 
curring at a distance of 161 
kilometres (100 miles) 

(b) a magnitude 7.25 earthquake oc- 
curring at a distance of 56 kilo- 
metres (35 miles) 

It is unlikely that these latter two 
earthquakes will produce conditions 
more critical than the motion dis- 
cussed in detail above, but the check 
should be made to verify that this is 
so. Design earthquakes for noncrit- 
ical structures can be less severe 
in intensity than those discussed 
above, and the Board will defer this 
recommendation vintil the evaluation 
of critical structures is completed." 

Ground motion characteristics are esti- 
mated for the recommended earthquakes as 
follows (Figure 137) : 



Magnitude 



Distance 
Km/Mi 



Peak 
Acceleration 



Predominant 

Period 

Sec. 



Duration 
(a > .05g) 
Sec. 



6.5 

7.25 

8.5 



5/3 

56/35 

161/100 



0.6g 

0.15g 

0.05g 



0.29 

0.4 

0.8 



20 

23 

3 



Based on these characteristics, the 
ground acceleration for the nearby event 
of magnitude 6.5 exceeds that from the 
others; and the duration is generally as 
great or greater; therefore, the 7.25 
and 8.5 magnitudes will not be considered 
further in the analysis. 

The acceleration time history shown in 
Figure 138 is essentially the one recom- 
mended by the consulting board. The 
accelerogram was derived by scaling the 
Pacoima S16E record down by 0.6/1.17 and 
adding the Taft record scaled up by 
0.3/. 15. The first 2.6 seconds of the 
Taft record were dropped and the joining 
made at time 11.2 seconds of the Pacoima. 
Accelerations in the Taft portion are 
about 30 percent higher by this proce- 
dure than by scaling all ordinates of 
the time history provided by Clarence R. 
Allen, to give a peak acceleration of 



0.6g. However, the Taft portion peaks 
are still small in comparison to the 
Pacoima peaks, and do not produce signi- 
ficant stresses in the embankment. The 
resulting Reanalysis Earthquake has the 
following characteristics: 



Richter Magnitude 



6.5 



distance from energy 5 kilometres 

source to dam (3 miles) 

maximum acceleration 0.6g 

predominant period 0.4 seconds 

duration 



acceleration time 
history 



20 seconds 

modified Pacoima 
plus modified 
Taft 



Figures 138 and 139 show the accelera- 
tion time history and response spectra. 



225 



40 



60 



KILOMETRES 

80 100 120 



140 





.8 


rv 




1 


. r 


Z 




o 


.6 


1- 




< 




(T 


fi 


UJ 




_J 




UJ 


4 


tJ 




o 




< 


.3 


2 




-) 




s 


. 2 


X 




< 


. 1 



. 1 


1 1 


1 


1 1 


1 1 


1 


























A \N 












\ \ Xv'^ 




















\4f \£rr<^ 


;::::>..^ 








X2^ 




^-^^ 









MAXIMUM ACCELERATION 

AFTER SCHNABEL AND SEED "ACCELERATIONS IN ROCK FOR EARTH- 
QUAKES IN THE WESTERN UNITED STATES" BULLETIN SEI S MO LOG ICAL 
SOCIETY OF AMERICA, VOLUME 63,1973 




PREDOMINANT PERIOD 

AFTER SEED, IDRISS AND KIEFER, "CHARACTERISTICS OF ROCK MOTIONS 
DURING EARTHQUAKES" JOURNAL, SMFE, SEPT., 1969. 



C/5 

a 


40 


H 




O 

o 


35 


UJ 




C/1 


30 



25 



< 


20 


:d 




Q 


15 


Q 




UJ 
t- 


10 


UJ 




^ 




o 


b 


< 




cc 




CD 






\V 












"^§S| 


^^M = 8 










V 1 


^ 




a> 0.0 


5g 




^^ 


V 


f > 2 h 


z 




x^. 


\\ 


\ 








w 


\ 


\^- 










\N 


^ 






1 


>.-,^^^^ 


==:,J2II 


_i ^ \ — 


- ^ - 


=__::^= 



20 40 60 80 100 

DISTANCE FROM ENERGY SOURCE-MILES 



120 



BRACKETED DURATION 

AFTER BOLT "DURATION OF STRONG GROUND MOTION" PROCEEDINGS 
FIFTH WORLD CONFERENCE ON EARTHQUAKE ENGI N EER I NG , ROME, I 974. 
Figure 137. Earthquake Ground Motion Characteristics 



226 




8.00 lO.OO 

TIME (SECONDS) 



Figure I38. Reanalysis Earthquake 



227 




2. so 3.00 





Figure 139- Response Spectra for the 
Reanalysis Earthquake 



228 



J 



7. ANALYSIS OF DYNAMIC STRESSES FOR THE REANALYSIS EARTHQUAKE 



Methods of Response Computation 

The response of a multiple-degree-of- 
freedom structure may be determined by 
solution of the following set of 
equations: 

[M] (u) + [C] {ui + [K] fu} = F(t) 
[''] = lEass matrix for the structure 
[C] damping matrix for the structure 
[K] = stiffness matrix for the 
structure 
u, u, u = nodal accelerations, veloc- 
ities, and displacements 
F(t) = earthquake load vector. 

Two of the most commonly used programs 
in the United States for solving these 
equations are QUAD4 (Idriss, et al, 
1973) and LUSH (Lysmer, et al, 1974). 
Both of these programs are currently in 
use in the Department of Water Resources 
for computing the seismic response of 
finite-element models of embankment dams. 
Some modifications have been made to them 
so the LUSH will generate stress time 
histories, as well as acceleration and 
displacement time histories, and QUAD4 
will take input in the same format as 
LUSH. 

Both programs use the equivalent linear 
method to account for the nonlinearity 
in the soil shear modulus and damping 
ratio. Every element in the structure 
is assigned an independent value of 
damping ratio and shear modulus, depend- 
ing upon the average shear strain anti- 
cipated during the earthquake. These 
properties remain constant during the 
shaking. After the response has been 
computed, the average shear strain and 
corresponding soil properties for every 
element are evaluated. If the differ- 
ence between the assumed and computed 
soil properties is less than a given 
tolerance, the solution is assumed con- 
verged. The average shear strain is 
computed as a fixed fraction of the 
maximum shear strain experienced during 



the shaking. 

QUAD4 solves the equations of motion by 
a direct integration method. Integration 
may be carried out by either the linear 
acceleration technique or Wilson's Theta 
Method. Rayleigh damping is used which 
filters out the structure's response in 
the higher frequency range. 

LUSH uses a complex number formulation 
of the elastic moduli and a method of 
complex response which assumes that the 
input motion is harmonic. This formula- 
tion allows viscous damping to be intro- 
duced in the construction of the stiff- 
ness matrix. The program was developed 
to analyze the response of high-frequency 
structures, such as nuclear power plants, 
and has the advantage of providing a more 
accurate response and a faster solution 
time. 

Acceleration Response of Dam to 
Reanalysis Earthquake 

Acceleration time histories were comput- 
ed at four elevations in the embankment 
for the following conditions: 



Coii5)uter Program 

Shell K„ 

2max 


QUAD4 
130 


Core Shear Modulus 




G /Su 
max 


2200 


Undrained Strength 
Envelope 


CU 



Average Shear Modulus 
Reduction Curves and 
Damping Curves 

Poisson's Ratio 



0.3 



Figure 140 illustrates how the motions 
are modified in progressing upward 
through the dam. 



229 



4.0 



TIME IN SECONDS 

8.0 12.0 16.0 20.0 



0.50 



0.50 

0.25 



z 

9. 0.25 

H 
< 
(£ 

1^0.25 

UJ 

^ 

< 

■d 0.25 



Z 

o 

N 

S 0.25 

O 

^ 



0.25 



0.50 




W« 



*f^fltgri 



IhE 



iwtf 



^ 



i 



uimE 



VATtON ?.QQ '^EIX 



,Y$1 



':?^'^:f 



^7 



P^™#' 



int 




mm_ 



TFTfrofK 



» 



#te 



8.0 12.0 16.0 20.0 
TIME IN SECONDS 




Figure 140. Acceleration Response to Reanalysis 

Earthquake 



230 



Input Variables and Computed Shear 
Stresses 

Comparative studies were made to eval- 
uate the influence of several variables 
on computed shear stresses. Two comput- 
er programs, LUSH and QUAD4, were used 
for the dynamic stress analyses. Both 
programs are in general use for comput- 
ing dynamic shear stresses of 
embankments. 

Shear stresses were calculated for two 
values of K^ for the shell — 130 and 
205. These values were considered to 
represent a reasonable range for the 
highly compacted gravels. The 130 value 
was determined by cyclic triaxial tests 
on 30 centimetre (12-inch) diameter sam- 
ples of gravelly sand with 5 centimetre 
(2-inch) maximum particle size. The 205 
value was chosen to represent the upper 



limit of the range estimated by a two- 
dimensional analysis for observed crest 
motions in the August 1, 1975 earthquake. 

Two of the soil properties used in the 
analysis, namely shear modulus of the 
core and Poisson's ratio, were assumed 
based on published information for other 
soils. Therefore, stresses were calcu- 
lated for a range of values of these 
properties to determine their influence 
on shear stresses. 

Although most of the study was concen- 
trated on the maximum section of the dam, 
shorter abutment sections could be more 
critical. Consequently, dynamic shear 
stresses were also calculated for dam 
sections 100 metres (330 feet) high and 
64 metres (210 feet) high. 

The following table summarizes the dif- 
ferent values used in the comparison 
studies : 



Variable 

Shell Shear Modulus 

Paramater K„ 
Core Shear Modulus 

Parameter G /S 

Shear Strength Envelope 

Modulus Reduction Curves 
Computer Program 



Poisson's Ratio 
Embankment Section 



Values Used in Comparison Studies 



130 

Low 

2200 

CU 

Low* 

QUAD4 

(Wilson's Theta Method) 
(.02 second time step) 

0.3 
750 ft. 



205 

High 

1120 

UU 

High* 

LUSH 

(Highest frequency = 
8 Hertz) 



0.45 
330 ft. 



.3/. 49** 
210 ft. 



*The modulus reduction curves and damping curves used are shown in Figures 20 and 

24. The low-modulus reduction curve for clays is the average curve by Seed and 
Idriss. It is generally for low values of Su. The high-modulus reduction curve 

for clays was provided by Professor Seed to John Vrymoed. It is for an S of 

8000 psf. It was used for core elements with S greater than 6000 psf. 

**0.3 in unsaturated downstream zone 
0.49 in submerged upstream zone 



17—78786 



231 



Appendix E includes plots of maximum 

dynamic horizontal shear stress and 

shear strain for the several conditions 

studied, shear stress time histories for 

K_ of 205 and 130, and acceleration 

timl^istories for K. of 205 and 130. 
/max 

The following embankment properties and 
conditions were held constant for all 
analyses : 

Acceleration time history for Reanalysis 
Earthquake 

Shell 

- moist density 150 pcf 

- saturated density 153 pcf 

- average modulus reduction curve for 

sands 

- average damping curve for sands 

Core 

- saturated density 153 pcf 

- average damping curve for clays 

Core Block 

- density 150 pcf 

- constant shear modulus 187,200,000 

psf 

- constant damping 

Static stresses from static finite ele- 
ment method analyses. 



Influence of Shear Modulus of Shell 
Material 

Response analyses were conducted for twc 
values of K„ ^ , 130 and 205, as dis- 
cussed previously (page 214). The fol- 
lowing embankment properties and condi- 
tions were used for both analyses: 

Computer Program LUSH 

Maximum Section 

Poisson's Ratio 0.3 

Core - G /S = 1120 
max u 

- UU static shear strength 

envelope 

- Higher shear modulus reduction 

curve for clays 

Table III shows the average maximum shea 
strains throughout the embankment. 





Table III 


Shell 
2max 


Average Maximum 
Shear Strain (%) 


130 
205 


0.18 
0.09 



Average shear strain = 0. 6 x maximum 
shear strain. 

Iterations continued until computed G 
was within 10 percent of trial G in 
nearly all elements. 



OROVILLE DAM - MAXIMUM SECTION 

REANALYSIS EARTHQUAKE— MAXIMUM ACCELERATION =0.6g 

LUSH DYNAMIC RESPONSE ANALYSIS 



CORE 



Su 



1120 (HIGHER CORE MODULUS 



Increasing the modulus from 130 to 205 
increases the shear stresses by 20-70 
percent in the lower portion of the up- 
stream shell as shown in Figure 141. 
Unfortunately, comparisons were not pos- 
sible in the upper portion of the dam 

because LUSH stresses for K^ of 130 

, . 2max 
were xncorrect in this area Xsee section 

on computer programs). 



MAXIMUM AT,, (K, 



^205) 



MAXIMUM AT C K, 



= 130) 



STRESSES FOR K 



2 MAX 



INCORRECT IN UPPER 300 FEET 




Figure 1^1. Influence of Shear Modulus of Shell Material on Computed Maximum 
Horizontal Dynamic Shear Stresses 



232 



Influence of Shear Modulus of Core 
Material 

A comparison of maximum dynamic shear 
stresses was made for two sets of core 
modulus input data. Both analyses were 
made using the following embankment pro- 
perties and conditions: 



Computer Program LUSH 
Maximum Section 
Shell K = 130 
Poisson s Ratio =0.3 

The two sets of core modulus parameters 
were as follows : 



Core Material 
Parameter 


Figure 
120 


High Core Modulus 

1120 
(Lower bound value 
reported by Seed & 
Idriss, 1977) 


Low Core Modulus 


G /S 
max u 


2200 
(Average value 
reported by Seed 
Idriss, 1977) 


Undrained 

Strength 

Envelope 


127 


Unsaturated 
UU 
(End of construction 
condition) 


Saturated 
CU 
(End of embankment 
consolidation 
condition) 



Normal Stress 



Modulus 

Reduction 

Curve 



116 



126 



l/2(a' lc+ a'3c) 
(From static FEM 

analysis this study) 



Higher curve for most 

elements because 

S = 6000 psf or more 



l/2(a'ic +a'3c) 
(From static FEM 
analysis by Nobari 
and Duncan, 1972)* 

Lower curve for all 
elements because S 
generally less than 
6000 psf 



*Normal stresses by Nobari and Duncan were used because their finite element mesh 
was finer in the core and transition zones than the mesh used in the present study, 
and the stresses were therefore better defined. However, after the analysis was 
completed it was discovered that these were total stresses, not effective stresses. 
Their figures 67 and 68 incorrectly defined the plotted contours as effective 
stresses. The net result is" that the low core modulus is somewhat higher than 
intended, but still four times less than the high core modulus. 



Of the listed parameters, the undrained 
strength envelope and modulus reduction 
curve were most significant. The CU 
strength envelope gave Su values about 
half as great as the UU envelope, gener- 
ally less than 8000 psf. This, in turn 
caused the lower reduction curve to be 
used which has about four times as much 
reduction in modulus as the higher curve 
at the strain levels in question. The 
net effect of all four parameters was 
that the low core modulus values were 



about one fourth the high core modulus 
values . 

A comparison of maximum dynamic horizon- 
tal shear stresses for the two core modu- 
lus conditions is shown in Figure 1A2 for 
the bottom 135 metres (450 feet) of em- 
bankment. A comparison for the top 90 
metres (300 feet) is not shown because 
the LUSH stresses are incorrect for 
these upper elements (see next section) . 



233 



STRESSES BY LUSH FOR K 



2 MAX 



OF 



130 INCORRECT IN UPPER 300 FEET 



REANALYSIS EARTHQUAKE 
COMPUTER PROGRAM LUSH 
SHELL K2MAX = 130 
POISSON'S RATIO = 0.3 




NOTE 



HIGHER CORE MODULUS ABOUT 
4 X LOWER CORE MODULUS 



Ar,y WITH LOWER CORE MODULUS 



ATxy 

Figure ]k2. Influence of Shear Modul 
Dynamic Shear Stresses 

Reducing the core modulus lowers the 
core stresses, and raises the stresses 
somewhat in the shells. The downstream 
shell has the greatest increase — up to 
40 percent. Most of the upstream shell 
has very little stress increase. A 
small zone adjacent to the base and the 
core has a 20-percent increase. And a 
narrow zone adjacent to the surface of 
the upstream slope actually has a 20- 
percent decrease. 

Computer Programs LUSH and QUAD4 

In order to resolve a question of unus- 
ual shear stress time histories by LUSH, 
and for general comparison, computed 
stresses by LUSH and QUAD4 were compared. 

Questionable shear stress time history 
patterns were found mainly in the upper 
90 metres (300 ft) of the maximum sec- 
tion for the LUSH analysis with K„ of 
130, for both the high and low core 
shear modulus values. There were signi- 
ficant shear stresses at time zero and 
large-amplitude, long-period stress 
fluctuations thereafter. A typical pat- 
tern is shown in Figure 143. (These un- 
usual patterns were not found in any 
elements for the LUSH analysis using 

K_ of 205). 
zmax 

The induced stresses should be zero at 
time zero, and should be insignificant 
for the first two seconds because the 
input accelerations are insignificant 
for the first two seconds. Furthermore, 



WITH HIGHER CORE MODULUS 

US of Core on Computed Maximum Horizontal 

these questionable patterns bear no re- 
semblance to other stress time histories 
throughout the embankment, which start 
at zero and have patterns consistent 
with the input acceleration time history 

Dynamic stresses were computed by QUAD4 
for the same input used for LUSH: 

Maximum section 
Shell K^^^ = 130 
Core, low shear modulus 
Poisson's ratio =0.3 

All time histories produced by QUAD4 had 
patterns consistent with that of the in- 
put earthquake motion. Shear stress pat- 
terns by LUSH and QUAD4 are the same in 
the lower part of the embankment, and 
are distinctly different in the upper 
90 metres, as illustrated by Figures 
143 and 144. It is concluded that the 
LUSH stresses in the upper 90 metres are 
incorrect for the analyses using a 

K^ of 130. 
2max 

The LUSH program was checked by analyz- 
ing a sample problem. It gave the pres- 
cribed results. Cause of the abnormal 
behavior is still undetermined. 

Figure 145 compares the maximum dynamic 
horizontal shear stresses for LUSH and 
QUAD4. In the lower part of the dam, 
where the LUSH stresses are valid, the 
two programs give about the same stress- 
es except in a zone adjacent to the up- 
stream slope, where LUSH stresses are 



234 





0. < 


i 


0> 


> 


* <::^ 




K 


^ 


°Z / 


/ 


<t u < 




= 2 


) 


O UJ <f 




_l ^ 


~J 




:3 





235 




2 "- 

a: ul 







236 



I 



about 20 percent higher. In the upper 
90 metres, particularly on the upstream 
side, the LUSH stresses are as much as 
60 percent greater than the QUADA 

STRESSES BY LUSH FOR Kjmax OF 
130 INCORRECT IN UPPER 300 FEET 



ATxy BY LUSH 



stresses. No conclusions may be drawn 
in the upper elements other than that 
f-e LUSH stresses are in error. 



REANALYSIS EARTHQUAKE 



LOWER CORE MODULUS 
POISSON'S RATIO = 0.3 




Figure ]k5. Comparison of Computed Maximum Horizontal Dynamic Shear Stresses 
by Computer Programs LUSH and QUADA 



It is recommended that in general use 
of program LUSH, if stresses appear to 
be out of line for any reason, the time 
histories should be checked to see if 
high stresses are occurring at time zero 
and for the period of time before input 
accelerations are significant. 

Influence of Poisson's Ratio 

A Poisson's ratio value of 0.3 was used 
for all the dynamic analyses of the Oro- 
ville embankment. On later reflection 
it was realized that a value of 0.5 is 
more appropriate for saturated soils 
during undrained loading. 

The influence of Poisson's ratio was 
examined using the following input: 

Model embankment (Figure 146) 

Computer program LUSH 

Shell K„ = 130 

2max „ , 
Static stresses = Y H (soil 

density x depth 

of overburden) 



Three Poisson's ratio conditions were 
analyzed : 

1. 0.3 for entire embankment 

2. 0.45 for the entire embankment 

3. 0.49 for saturated zone (upstream 
half) and 0.3 for unsaturated 
zone (downstream half) 

Figure 147 shows the influence of Pois- 
son's ratio on dynamic shear stresses. 
The higher Poisson's ratios generally 
cause about 10 percent higher dynamic 
horizontal shear stresses in the central 
portion of the embankment, but hardly 
any difference at the base or crest. 
Comparisons of horizontal and vertical 
dynamic normal stresses are included in 
Appendix E. Although the comparisons 
were not made on the Oroville maximum 
section directly, it is considered rea- 
sonable to generalize them for applica- 
tion to the Oroville section. 



237 



PHREATIC LINE 



AREA BELOW 
A PHREATIC LINE 



54 ELEMENTS 
63 NODES 




MODEL EMBANKMENT 



STUDIES PERFORMED = 

3 ANALYSES CONDUCTED WITH INDENTICAL INPUT VARIABLES EXCEPT POISSON'S RATIO (V) 

POISSON 'S RATIO CONDITIONS ANALYZED 

I). POISSON'S RATIO = 0.30 FOR ENTIRE EMBANKMENT SECTION 
2). POISSON'S RATIO = 0.45 FOR ENTIRE EMBANKMENT SECTION 
3). POISSON'S RATIO = 0.49 FOR ELEMENTS BELOW PHREATIC LINE 
POISSON'S RATIO = 0.30 FOR ELEMENTS ABOVE PHREATIC LINE 

METHOD OF ANALYSIS = LUSH PROGRAM 

HIGHEST FREQUENCY USED IN ANALYSIS = 8.0 HERTZ 

EFFECTIVE STRAIN RATIO = 0.60 

INPUT MOTION = REANALYSIS EARTHQUAKE 

MAXIMUM ACCELERATION = 0.6g 

AVERAGE MODULUS REDUCTION CURVE FOR SANDS 

AVERAGE DAMPING CURVE FOR SANDS 

STRESS CONDITIONS ASSUMED FOR STUDY 

H = HEIGHT OF SOIL ABOVE ELEMENT CENTROID 
y = DENSITY, 153 pcf FOR ELEMENTS ABOVE PHREATIC LINE 
91 pcf FOR ELEMENTS BELOW PHREATIC LINE 
= EFFECTIVE VERTICAL NORMAL STRESS I psf ) = / * H 
= EFFECTIVE HORIZONTAL NORMAL STRESS ( psf ) = Kq^ct^' 
-- COEFFICIENT OF EARTH PRESSURE AT REST = 0.60 



K„ 



0- ■+20-, 



o-^' = EFFECTIVE MEAN NORMAL STRESS! psf ) = — " — 3 ^— 

K2MAX = SHEAR MODULUS PARAM ETER = 130 _i_ 

Gmax = MAXIMUM SHEAR MODULUS AT LOW STRAINS ( psf ) = 1000* Kg^AX * ^""'"'^^ 

Figure 1^6. Model Embankment for Determining Influence of Poisson's Ratio on 
Dynamic Shear Stresses 



238 



A r xy iV =0.49/0.30) 
ATxy (Z/=0.30) 




REANALYSIS EARTHQUAKE 
COMPUTER PROGRAM LUSH 



K 



2 MAX 



= 130 



Figure l'*?- Influence of Poisson's Ratio on Computed Dynamic 
Shear Stresses 



Influence of Embankment Section 

Most dynamic analyses of dams are done 
for the maximum section. However, other 
sections could be more critical. To co- 
ver the range of possibilities, two 
other sections were analyzed using the 
same input properties and conditions as 
used for the maximum section: 

Computer program QUAD4 
Poisson's ratio =0.3 



Shell K, = 130 

Core G ^"'f^^ = 2200 

CU shear strength envelope 
Average modulus reduction curve 
for clays 

Finite-element meshes for the two sec- 
tions were the upper rows of elements of 
the maximum section mesh. Natural per- 
iods and maximum crest accelerations for 
the three sections are: 



Elements 

Height, metres/feet 

First mode natural 
period, sec. 

Maximum horizontal crest 
acceleration, g 

Maximum vertical crest 
acceleration, g 



Maximum Section 
564 
225/750 

1.98 

0.64 

0.38 



Section 2 

128 
100/330 

1.22 

0.75 

0.51 



Section 3 
66 
64/210 

.83 

0.79 



0.66 



239 



Figure 148 shows comparisons of the two 
shorter sections to the maximum section 
for maximum dynamic horizontal shear 
stresses. With respect to the maximum 
section, both short sections develop 



less stress in the outer parts of the 
shells, and about the same stress in 
the center; on the average, most of the 
upstream shell develops about the same 
stresses. 




SECTION 3 (64 METRES HIGH) 




SECTION 2 (lOOMETRES HIGH) 



REANALYSIS EARTHQUAKE 
COMPUTER PROGRAM QUAD 4 
SHELL K2MAX = 130 

Figure 1^8. Comparison of Computed Maximum Horizontal Dynamic Shear 
Stresses for Different Embankment Sections 



240 



Combined Influence of Variables 

The submerged upstream gravel shell is 
the area of concern for seismic stabil- 
ity. The following table summarizes 
the influence on stresses in that area 
by the variables studied: 



Variable (Figure) 



Shell K„ (35) 

2max 



Core Shear 
Modulus 



(36) 



Range 
205 vs. 130 
Lower vs Higher 



Influence on Dynamic 
Horizontal Shear Stresses 
in Upstream Shell 
(Percent Difference) 

+20 to +70* 

to -20* 



Computer (39) 

Program 

Poisson's (41) 

Ratio 

Embankment (42) 
Height 



LUSH vs. QUAD4 



0.45 vs. 0.3 



100 metre vs. 
230 metre 



to +20* 



to +5 



to ilO 



*Not defined in upper 90 metres because LUSH stresses with K„ = 130 
^ . \,, . 2max 

are incorrect in this area. 



For the dynamic shear stresses in the up- 
stream shell, the influence of Core 
Shear Modulus, Computer Program, Pois- 
son's ratio, and Embankment Height are 
relatively minor over the range of val- 
ues studied. However, for the combina- 
tion of Program LUSH and higher shear 
modulus in the core, the 20 percent dif- 
ferences are cumulative in a shallow 
zone along the upstream slope. This is 
the combination used to calculate stress- 
es for evaluation of embankment perform- 
ance and will give the highest stresses 
in this zone. 

The major influence is shear modulus of 

the shell. Increasing shell K„ by a 
£ ^ c t n \. 2max -^ 

factor of 1.6 causes shear stresses to 

increase by 1.2 to 1.7. The choice of 
^2max ^^^^^^ ^^^ range 130-205 will be 
a major determinant of computed embank- 
ment performance. The conservative ap- 
proach would be to use a value around 
200 - near the upper end of the range. 



Three-Dimensional Effect 

Jlomparisons of stresses by two- 
dimensional (plane strain) and three- 
dimensional analyses are shown in Figure 
149 (after Makdisi, 1976) for a dam with: 

H = 30 metres (100 ft) 

L/H = 3 

Slopes =2:1 

V = 150 metres per second (500 feet 

per second) 
Damping = 10% 
Earthquake = Taft (first 15 seconds) 

a = 0.2g 

max ° 

For the maximum section, the three- 
dimensional analysis gives stresses at 
the crest, base, and slope areas that 
are 50 percent to 100 percent lower than 
those for the plane strain analysis. In 
the central portion, stresses are the 
same for 2D and 3D analyses. 



241 




L/H =3 




00 % 



MID-SECTION 



1 




-50^'^^^. 


"^f\ 




o 
If) 




- 90%^. 


z-?^^^ 




' 




1 


^—80' 


"^^^^"-^ 



AFTER MAKDISI (1976) 



QUARTER SECTION 

Figure 1^49. Comparison of Maximum Horizontal Shear Stresses Determined 
from 3D and Plane Strain Analyses Using Base Motions from Taft Record 



For the quarter section, stresses by 2D 
and 3D analyses differ by less than 20 
percent for most locations, but near the 



crest the 3D stresses are twice as high 
as 2D stresses. 



242 



For Oroville Dam with L/H = 7, smaller 
differences would be expected between 
2D stresses and 3D stresses. 

One way to estimate the 3D stresses for 
the Oroville maximum section is to use 
the embankment response model defined 
in Section 5. The implied assumption 
here is that the same pseudo K„ value 
(350) which accounted for abutment re- 
straint effects in the August 1, 1975 
earthquake will also correctly account 
for abutment restraint effects in the 
Reanalysis Earthquake. Then the model 
will give the correct period and accele- 
ration response, and the correct comput- 
ed linear elastic shear strains in the 
maximum section of the actual (3D) 
embankment. 

The effect of abutment restraint is to 
reduce the strains of the maximum sec- 
tion by developing shear stresses on its 
sides. Thus all the inertia forces ao 
not have to be borne by shear stresses 
on the tops and bottoms of elements. It 
is these top and bottom stresses that 
are usually of concern in relation to 
embankment behavior. These element top 
and bottom shear stresses can be calcul- 
ated by multiplying the shear modulus 
of the material by the 3D shear strains 

obtained from the embankment response 
model . 



The embankment response (3D) was calcul- 
ated for the following conditions: 

Computer Program LUSH 

Maximum Frequency Used 10 Hertz 
Maximum Section 
Shell - Pseudo K„ = 350 

- Average Modulus Reduction Curve 
for Sands 

- Average Damping Curve for Sands 
Core - G^g^/S = 1750 

- Shear Strength Envelope UU 

- Higher Modulus Reduction Curve 

- Average Damping Curve for Clays 
Poisson's Ratio 0.3 

Appendix E includes 2D shear stresses for 
pseudo K2niax ~ 350; Appendix F explains 
the method used to calculate 3D stresses. 



Figure 150 shows the ratio of 2D to 3D 
stresses in the upstream shell computed 
for shell K^ ^^ of 205. In the lower 
interior part^of the upstream shell, 2D 
and 3D stresses are the same; in the 
central part and upstream of the coffer- 
dam core, 2D stresses are 25 percent 
higher; in the upper part, 2D stresses 
are over 50 percent higher. These re- 
sults are similar to Makdisi's in the 
upper part of the dam, but the differ- 
ences between 2D and 3D stresses is much 
less than those (Makdisi's) along the 
slope and base. 



OROVILLE DAM - MAXIMUM SECTION 

REANALYSIS EARTHQUAKE - MAXIMUM ACCELERATION 

LUSH DYNAMIC RESPONSE ANALYSIS 

PSEUDO K2MAX = 350 

PSEUDO Gmax ' Su = 1750 

SHELL KjMAX 

CORE Gmax > ^n = M20 



0.6 g 




MAXIMUM AT , FOR PLANE STRAIN CONDITIONS MAXIMUM Ar,„ ( K, 



MAXIMUM Ar,„ FOR PSEUDO -3D CONDITIONS 



205) 350 



MAXIMUM Ar,y (K2MAX=350) 

Figure 150. Estimated Three-Dimens ional Effect on Computed Maximum Hori-zontal 
Dynamic Shear Stresses 



243 



8. CYCLIC SHEAR STRENGTH 



Cyclic Strength Test Program 

Cyclic strength testing was carried out 
only for the saturated upstream gravel 
shell and transition. The downstream 
shell is not saturated and would be much 
stronger — with cyclic strength essen- 
tially equal to the static effective 
strength. The two clayey gravel cores 
occupy a relatively small proportion of 
the embankment. Also, studies have 
shown that compacted clayey embankments 
perform well during earthquakes (Seed, 
Makdisi, & DeAlba, 1978). 



The cyclic triaxial testing was carried 
out under the direction of Professor 
H. B. Seed at the University of Califor- 
nia, Richmond Field Station. The prograi 
consisted of about 90 cyclic triaxial 
tests on 30-cm (12-inch) diameter sample: 
to measure the cyclic strength of the 
gravel; and about 12 cyclic triaxial 
tests on 7.1-cm (2.8-inch) diameter sam- 
ples to determine the effects of aging. 
Tests of the larger samples included the 
following consolidation conditions to 
represent a wide range of locations with- 
in the upstream shell: 



Minor Principal 
Consolidation Stress 

196 kilopascals (4100 psf) 

784 kilopascals (16,400 psf) 

1370 kilopascals (28,700 psf) 

2550 kilopascals (53,300 psf) 



Consolidation 
Stress Ratio 
K = g' Ic/ g' 3c 



1, 1.5 & 2 

1 

1, 1.5 & 2 

1. 1.5 & 2 



Sample Gradations and Density 

Modeling Embankment Shell Gradation 

Zone 3 materials are sands, gravels, 
and cobbles up to 15 cm (6 in.) size. 
Figure 151 shows the average gradation. 
Even for gradations coarser than the 
average, maximum particle size rarely 
exceeds 22.5 cm (9 in.) with only a 
small percent of material larger than 
15 cm (6 in.). 

Testing of smaller size particles to 
represent full-scale material has been 
done for many years. Lowe (1964) was 
the first to use a model gradation 
parallel to the field gradation. This 
modeling method has since been used by 
others, notably Marachi et al (1969), 
Becker et al (1972), and Wong (1973). 
Marachi and Becker did static shear 
tests on full-size field gradations of 
Oroville Zone 3 material as well as on 



the modeled gradation. When compared 
at equal relative densities, the fric- 
tion angle for the two parallel grad- 
ations was the same (maximum difference 
was one degree) . 

The same modeled gradation has been 
used in this study for the cyclic tri- 
axial tests on gravels. Additionally, 
a second modeled gradation was used for 
cyclic triaxial tests of smaller sam- 
ples (Figure 151). The ratio of sample 
diameter to maximum particle size is 
six. 

Relationship of Test Sample Density to 
Field Density 

The objective was to prepare test sam- 
ples at the same percent compaction as 
achieved for the Zone 3 shell material 
compacted into the embankment. Figure 
152 shows the statistical distribution 
of percent compaction in the shell. 



I 



244 



U.S. STANDARD SIEVE SIZES 




0.5 1.0 5.0 10.0 

GRAIN SIZE IN MILLIMETRES 



100.0 



Figure 151. Field and Modeled Oroville Gravel Gradations 



Average = 99% 

Standard Deviation = 4% 

Significant Range = 90% - 110% 

Figure 153 shows that most of the gra- 
vel shell contained between 5 percent 
and 20 percent minus No. 4, with an 
average of 14 percent; and that for 
this average gradation, the average 
compaction achieved was 100 percent at 
a dry density of 2 387 kilograms per 
cubic metre (149 pcf). 

Maximum density tests were carried out 
on samples of minus 5-cm (2- in.) mater- 
ial (modeled gradation) to be used for 



the cyclic triaxial tests. The same 
equipment and procedures were used for 
these tests as for the control tests 
run during construction. Test proce- 
dures are described in the Oroville Dam 
Embankment Materials Report. In sum- 
mary, the test consists of vibrating a 
500 kilogram (110-lb.) saturated sam- 
ple in a 67. 5-cm (27-in.) diameter 
mold for five minutes with a 13.8- 
kilopascal (2-psi) surcharge. Vibra- 
tion is with a foundry type air-driven 
vibrator at a frequency not less than 
7,000 VPM. During vibration the mold 
is lifted off the floor. 



245 



100 



^ 75 
IxJ '^ 



50 



25 



mm 




I II I II I M I M I 



X = 99.0 
a = 4.2 
N = 841 




84 86 



92 94 96 98 100 102 104 
PERCENT COMPACTION 



06 !08 110 



OROVILLE DAM 
Figure 152. Final Statistical Analysis - Zone 3, Percent Compaction 



246 



PER CENT PASSING NO. 4 SIEVE 
10 20 30 40 



160 



150 



140 - 



130 




50^ 



■ AVERAGE OF MAXIMUM 
DENSITY TESTS 



/'AVERAGE OF FIELD 
/ DENSITY TESTS 



_1 


Q 


u. 


a: 




< 


U- 


> 


o 




10 


o 


u 


m 


5 


z> 


z> 


u 


_J 




o 


o 



20 



10 20 30 
PER CENT PASSING NO. 
Figure' 153. Field Control Tests - Zone 3 



Figure 154 shows the results of these 
tests, along with construction control 
test results of samples with 14 percent 
minus No. 4. For the six tests on mod- 
eled gradation samples, maximum density 
appears to be a function of vibration 
frequency. However, the construction 
control tests show a general scatter 
with no apparent correlation. 

This difference is understandable. For 
the tests on modeled gradation material, 
the gradation and specific gravity were 
exactly the same for all tests; and the 
tests extended over a period of two 
weeks during which vibrator character- 
istics would not be expected to change 
drastically. By contrast, the construc- 
tion control tests were run over a 



period of several years on materials 
with varying gradations and specific 
gravity. 

The frequency of vibration generally 
varied from 8,000 to 12,000 VPM, aver- 
aging 10,000 VPM during construction 
control testing. Therefore, it seems 
appropriate to use a frequency of 
10,000 VPM for maximum density testing 
of the modeled gradation material. An 
average density of 2 390 kilograms per 
cubic metre (149 pcf) was used for set- 
ting up cyclic triaxial test samples. 
The range was 2 360 to 2 410 kilograms 
per cubic metre (147 to 151 pcf) and 
the standard deviation was 18 kilograms 
per cubic metre (1.1 pcf). Both maxi- 
mum density at about 10,000 VPM and 



247 



160 



150 



140 



If) 

■2. 

LU 130 



X I 50 

< 



RANGE OF FREQUENCIES USED 
IN CONSTRUCTION CONTROL TESTING 



8000 TO 12000 RPM 



130 



> MAXIMUM DENSITY AT 

9700 RPM , 155 pcf 
AVERAGE MAXIMUM 



DENSITY 150 pcf 



MODELED GRADATION 
2INCH MAXIMUM, 30%-*4 
TESTED AT BRYTE LAB. 



AVERAGE TEST -... 

FREQUENCY 9700 RPM 



^ 



I 



ONE TEST AT BRYTE LAB 



•t 



AVERAGE MAXIMUM 



DENSITY 148 pcf 



APPROX AVERAGE GRADATION 
6" MAXIMUM, l4%-*4 TESTED 
AT OROVILLE LAB DURING CONST. 
J \ \ I 



7000 



8000 



9000 10000 1 1000 

FREQUENCY (RPM) 



12000 



13000 



27 INCH MOLD 

AIR VIBRATOR 

1100 POUND SAMPLE 

2 psi SURCHARGE 

5 MINUTE VIBRATION 

Figure 15^. Maximum Density Tests - Zone 3 



average maximum density for 7,000 to 
10,000 VPM are considered in comparing 



sample compaction with field compaction 
in the following table : 



248 



Compaction 
Conditions 

Average Maximum Density 
for All Frequencies 

Average Placed Density 

Average Percent Compaction 



All Zone 3 
Material 



99% 



Material With 
Average Gradation 
in Zone 3 
Embankment 



148 pcf 

149 pcf 
101% 



Cyclic 
Triaxial 
Samples 



150 pcf 
149 pcf 
99% 



Average Maximum Density 
for Frequencies From 
9.500 to 10.000 VPM 

Average Placed Density 

Average Percent Compaction 



150 pcf 
149 pcf 
99% 



155 pcf 
149 pcf 
96% 



Range of Percent 
Compaction 



91% - 107% 
(+2 standard 
deviations) 



95% - 100% 



The average percent compaction of the 
cyclic triaxial test samples is slight- 
ly (2 percent to 3 percent) less than 
the average percent compaction of Zone 
3 material with average gradation; and 
also slightly less (0 to 3 percent) 
than the average percent compaction of 
all Zone 3 material. 

Summary of Test Procedures 

30 cm Diameter Samples 

The sample was manufactured in 10 lay- 
ers in a membrane-lined mold on the tri- 
axial test base. Material for each 
layer was weighed, mixed, placed in the 
mold, and vibrated for four minutes by 
a vibrating weight placed on top of the 
layer. The top cap was placed on the 
sample, a vacuum applied, the mold re- 
moved, and a second membrane installed. 
The chamber was assembled on the base, 
filled with water, and pressurized. 
The sample was soaked by flowing water 
upward through it, and then backpres- 
sured until pore pressure parameter B 
was 0.9 or higher. Back pressure for 



most tests was 392 kilopascals (8,200 
psf). After consolidation was com- 
plete, cyclic loading was applied at a 
frequency of one cycle per minute, be- 
ginning with the compression half. 
Loads were set to be sinusoidal and 
syimnetrical about the static axial 
deviator stress. The slow frequency 
was necessary for the hydraulic load- 
ing system to maintain scheduled loads 
as the sample underwent large displace- 
ments. Continuous records were made 
of Ipad, displacement and pore pressure. 

7.1 cm Samples 

Procedures were basically the same as 
for the larger samples. Exceptions 
were: 

1. Material was not mixed individually 
for each layer. 

2 . Sample was made in five layers ra- 
ther than ten . 

3. Half of the samples were allowed to 
consolidate for two months before 
testing. 



249 



4. Loading frequency was one cycle per 
second. 

Results of Cyclic Triaxial Tests 

The testing report (Appendix L) , which 
is now being prepared, will be provided 
on request when it becomes available. 

Copies of the test records and associ- 
ated data, plots of cyclic stress vs. 
number of cycles for specified strains, 
and cyclic strength envelopes, all for 
tests of 30-cm (12-in.) samples, have 
been provided to DWR. Summaries of the 
test records are included in Appendix G. 
The cyclic strength enveopes are dis- 
cussed in the next section. 

Detailed results of the aging tests on 
7.1-cm C2.8-in.) samples have not been 
provided yet. A verbal report was made 
that aging did not produce any strength 
gain. 

Most of the tests were successful, but 
some records show deficiencies. Typ- 
ical test records for successful tests 
are shown in Figures 155 and 156. The 
deficiencies were mainly found on tests 
with isotropic consolidation (K = 1) 
and include the following: 

1. Loading was more than 10 percent 
asjTmnetric and drifted or jumped dur- 
ing the test. In addition, the load- 
ing amplitudes often attenuated quite 
severely with succeeding cycles. 
About half of the samples consoli- 
dated isotropically at 785 kilo- 
pascals (16,400 pounds per square 
foot) suffered from severe attenu- 
ation. An example of such load 
attenuation is shown in Figure 157. 

2. Load jumps and unusual pore pressure 
jumps rendered about the first 20 
cyclic tests questionable. It was 
unfortunate that most of the 
isotropically-consolidated tests at 
an initial confining pressure of 196 
kilopascals (4,100 psf) were in this 
group. Exan^sles of such tests are 
shown in Appendix G. 



3. Many samples which developed signi- 
ficant strain produced symptoms of 
necking. 

4. Many samples which developed only 

a small amount of strain were tested 
for only a relatively few number of 
cycles. For example, except for two 
tests which experienced severe load 
attenuation, all of the samples con- 
solidated isotropically to 784 kilo- 
pascals (16,400 psf) were stressed 
only to 12 cycles or less. 

5. After the testing program was com- 
pleted, the load calibration was 
found to be different from the des- 
ignated value. The actual applied 
loads were approximately 15 percent 
higher than those recorded. The 
correction not only changes the val- 
ues of the cyclic stresses but also 
alters the values of the anisotropic 
consolidation stress ratios, because 
the hydraulic actuator was used to 
increase the major principal consol- 
idation stress to a higher value 
than the minor principal consolida- 
tion stress. There were some check 
tests performed in an attempt to as- 
certain when the calibration actual- 
ly deviated from its designated 
value. On the basis of these tests, 
this point was found to be about 
halfway through the testing program. 
However, the 15 percent difference 
would be virtually impossible to 
find by check tests, because the 
variations of cyclic test results 
are at least that large. It is also 
quite reasonable to assume that the 
calibration error was present through 
out the testing program. 

Several corrections must be made to the 
cyclic triaxial stresses on the rest 
records : 

1. The C correction is made because 
the triaxial test does not duplicate 
the stresses present in an actual 
soil element. A C value of 1.0 

X" 

is usually us'ed for anisotropically- 
consolidated samples; a value of 



250 



about 0.6 is used for isotropically- 
consolidated samples. 

2. Axial stresses should be multiplied 
by 1.15 to correct for the load cal- 
ibration error. For the anisotrop- 
ically consolidated samples this 
will also give slightly higher con- 
solidation stress ratios. K values 

c 
of 1. 5 and 2 should be changed to 

1.57 and 2.15 respectively. 

3. Membrane Strength Correction: This 
correction is used to account for 
the fact that the membranes surround- 
ing the sample carry some of the ap- 
plied axial load. The correction is 
a function of the induced strain and 
is shown in Appendix L (available 
later in 1979) • This correction be- 
comes significant only at the lowest 
effective consolidation stress used 
in this study. 

4. Membrane Compliance Correction - 
During consolidation, the membrane 
penetrates into spaces between part- 
icles around the surface of the sam- 
ple. During cyclic loading, pore 
pressure increases are a controlling 
factor in sample behavior. The sys- 
tem is intended to be undrained, 
keeping the pore volume constant. 
However, if the membrane penetration 
decreases slightly, the pore volume 
expands slightly and the pore pres- 
sure increase is less than if the 
voltmie was kept constant. Verbal 
reports are that cyclic stresses 
should be multiplied by 0.9 to ac- 
count for membrane compliance. 

Investigation of Sample Behavior of 

Dense Sands in Static and Cyclic 

Triaxial Tests 

Objective 

Many limitations of the cyclic triaxial 
test have been pointed out by Seed and 
Peacock (1970) and by Seed (1976) and 
include : 



1. The cyclic triaxial test does not 
reproduce the correct initial stress 
conditions within the ground. 

2. There is a 90° rotation of the direc- 
tion of the major principal stress 
during the extension and compression 
halves of the loading cycle. 

3. The intermediate principal stress 
does not have the same relative val- 
ue during the two halves of the load- 
ing cycle . 

4. Unless special precautions are 
taken, friction may develop between 
the san5>le and the end caps, which 
will cause stress concentrations. 

5. During the extension half of the 
stress cycle, necking may develop 
and invalidate the test data beyond 
this point in the test. 

These limitations are legitimate, but 
the test is used despite them, because 
triaxial test results have been success- 
fully related to other cyclic shear 
tests such as the cyclic simple shear 
by an appropriate correction factor. 

However, consideration of these limit- 
ations played an important role in the 
interpretation of the cyclic triaxial 
tests carried out for the modeled Oro- 
ville gravel samples. The cyclic test 
results for the isotropically consoli- 
dated samples produced many questions 
concerning the development of sample 
strain. These samples produced strain 
almost totally in the extension direc- 
tion. There was debate over whether 
some samples developed only a limited 
amount of strain during testing. It 
seemed that many samples required cyc- 
lic "tension" stresses 
(o dp/2a '3c">0. 5) to produce signifi- 
cant strain levels. Other samples ex- 
hibited so much extension strain that 
necking was suspected. A single cause 
for this behavior was difficult to iso- 
late due in part to the severe load 
attenuation during many of these tests. 



251 



It was not clear whether the previously 
mentioned test limitations were the 
cause of the described behavior, or whe- 
ther the test results were valid (as 
valid as triaxial tests that don't pro- 
duce this observed strain behavior) . 
Of particular concern was resolving 
whether these test results could be 
used to evaluate the performance of 
the embankment during earthquake 
shaking. 

Since available information on the be- 
havior of very dense samples is limit- 
ed, it was decided to conduct a labora- 
tory investigation to examine more 
closely the behavior of dense samples 
and to document the results photograph- 
ically. The specific objective was to 
answer these three questions: 



Is the development of primarily ex- 
tension strain a valid result for 
isotropically consolidated dense 
samples, or is it a product of an 
erroneous test condition such as 
sample necking? 

Did the isotropically consolidated 
cyclic tests on the modeled Oroville 
gravel develop necking problems? 

What happens to an isotropically 
consolidated sample when the cyclic 
stress ratio (a , /2a'_ ) exceeds 
0.5 (tension) and does this consti- 
tute a valid test condition? 



ts^o 




-8200'- 



TEST NO. 72 
Vd = 149 5pcf 
Kj =1.0 
CTj'c = 16400 pcf 





Figure 155. Cyclic Triaxial Test Records for Modeled Oroville Gravel 



252 






TEST NO 45 

y^ = 149 2 pcf 

Kj = 2.0 

CTj'c = 28700 psf 



TEST NO 56 
Y^ = 149 5 pcf 
Kj =15 
o-3'e= 53300 psf 






Figure 156. Cyclic Triaxial Test Records for Modeled Oroville Gravel 



253 



55^ 




tlWtWtWi Ittlltlllllllt 




\l 



?ii 



^ 



^'fy 



^ 





r r ^ ^ f\ 


' \ ,0 ,5 i ! 20 



f" 




11 




1 


1 


n '^ 


1 

J 

1 


1/ 




^ 


J 




b 


J 

1 


lil 


95 










100 





Figure 157. Cyclic Triaxial Tes 
Program and Procedures 

Approximately 60 static and cyclic tri- 
axial tests were carried out on dense 
samples with particular attention to 
observing sample behavior and relating 
it to strain and pore pressure charac- 
teristics. The tests were conducted 
at the Department of Water Resources' 
Laboratory at Bryte and at the Depart- 
ment of Transportation Laboratory in 
Sacramento. The two materials tested 
were Monterey "0" sand and the minus 
1.27-cm (1/2-inch) portion of the mod- 
eled Oroville gravel, designated Oro- 
ville sand. The gradations of these 
two test materials are shown in Figure 
158. Films, slides, and photographs 
were taken during the tests to document 
sample behavior and aid in the inter- 
pretation of the test measurements. 

Monterey "0" sand was used to investi- 
gate the behavior of dense, cohesion- 
less samples that were constructed as 
uniformly as possible and tested under 
ideal conditions. The behavior of 
these uniform samples could then be 



TEST MO 76 
Tj " 149 7 pet 
Ke • 1.0 
0-3'^' 16400 pif 

t Records for Modeled Oroville Gravel 

compared to the behavior of samples 
that simulated the sample characteris- 
tics of the modeled Oroville gravel. 

The Monterey "0" sand was chosen be- 
cause it has a uniform gradation and 
has been used extensively in previous 
investigations. These samples were 
prepared by pluvial compaction through 
air in an attempt at producing the most 
uniforto sample possible. In addition, 
the samples were capped with "friction- 
less" Incite platens lubricated with 
silicon grease and covered with a cir- 
cular piece of rubber membrane. 

The Oroville sand material was used to 
represent the sample characteristics of 
the modeled gravel samples. Both types 
of samples had gradation curves parallel 
to the average field gradation and had 
the same ratio of sample diameter to 
maximum grain size. Both types of sam- 
ples would also be constructed using 
the same preparation technique (high- 
frequency vibration in layers). No 
special precautions were taken to mini- 
mize the friction at the end platens. 



254 



U.S. STANDARD SIEVE SIZES 



100 



90 



^200 '^lOO *50 *30 *I6 *8 ** 



4 3/8" 3/4" I 1/2" 



6 12 



I 70 



LlI 



60 



50 



40 



O 

tr 

^ 30 

Q. 



20 



10 



1 




1 


/ ' 






1 1/ 

f 




1 


1 








\ 


1 1 

MONTEREY 


/ 
1 












U 


SAND 


S 












t\ 


1 






















1 
1 


















\ 






a 


^^~~-~ ( 


)ROVI 


_LE S/ 


\NDS 










\ 




> 

/ 












































L 


/ 
/ 




















^ 


f 




















1 


y 


1 


1 






1 


1 




1 


1 





0-5 



1.0 5.0 10.0 

GRAIN SIZE IN MILLIMETRES 



50.0 100.0 



Figure 158. Monterey "0" Sand and Oroville Sand Gradations 



255 



All of the samples constructed were 
12.7 centimetres (5 inches) high and 
6.4 centimetres (2.5 inches) in dia- 
meter. Monterey "0" sand samples were 
confined by a 0.3-millimetre (0.012- 
inch) rubber membrane and the Oroville 
sand samples by either a 0.3- or 0.6- 
millimetre (0.012- or 0.025-inch) rub- 
ber membrane. Most of the Monterey "0" 
sand samples had dry densities ranging 
from 1 698 to 1 714 kilograms per cubic 
metre (106 to 107 pounds per cubic 
foot). A few additional static test 
samples, however, were constructed with 
dry densities of about 1 569 to 1 586 
kilograms per cubic metre (98 to 99 
pounds per cubic foot). 

The Oroville sand samples were prepared 
in five 2.5-centimetre (1-inch) layers 
and had dry densities of approximately 
2 275 kilograms per cubic metre (142 
pounds per cubic foot). The relative 
density of the very dense Monterey "0" 
sand samples was estimated to be 95 to 
100 percent. The relative density of 
the Oroville gravel samples was esti- 
mated to be about 85 to 90 percent. 

Saturation details are similar to those 
used by Mulilis et al (1975). Carbon 
dioxide gas is first passed through the 
sample to replace the air within the 
voids and cell lines. Carbon dioxide 
is used because its solubility in water 
is much greater than that of air. After 
the carbon dioxide stage, de-aired wa- 
ter is slowly introduced into the sam- 
ple from the bottom. The water moves 
into the sample at a very slow rate, 
filling most of the voids in the sam- 
ple. After passing water through the 
sample, back pressure is applied to 
dissolve any remaining gas bubbles. 
The degree of saturation is checked by 
measuring the pore pressure parameter 
B. Almost all samples tested had B 
values of 0.90 or greater and back- 
pressure values equal to 393 kilopas- 
cals (8,200 pounds per square foot). 
All samples tested were consolidated 
isotropically to an effective consoli- 
dation pressure of 196 kilopascals 
(4,100 psf). 



Static Tests on Monterey "0" Sand 

Static tests were conducted only for 
the Monterey "0" sand samples. Consol- 
idated undrained tests were carried out 
for medium and very dense samples in 
both the compression and extension di- 
rections, with a strain rate of 0.03 
percent per minute. 

Typical stress and pore pressure devel- 
opment with strain are shown on Figures 
159 and 160. The following observa- 
tions may be made : 

1. For all of the tests conducted, the 
pore water pressure increases at low 
strains. After a certain strain val- 
ue is reached, the pore pressure be- 
gins to decrease. This drop in pore 
pressure continues until the sample 
fails. The pore pressure develop- 
ment with strain is nearly identical 
in both the extension and compression 
directions. 

2. The maximum stress and the slope of 
the stress-strain curves are much 
greater for the compression tests 
than for the extension tests, thus 
indicating the relative weakness of 
the extension direction. 

3. At a given axial stress, the lower 
density samples strain farther than 
do the higher density samples. This 
is true for both the extension and 
compression directions. 

4. The difference in the slopes of the 
stress-strain curves in the extension 
and compression directions is more 
pronounced in the samples of higher 
density. 

Figure 161 summarizes eight stress- 
strain curves for the static tests con- 
ducted on the very dense samples of the 
Monterey "0" sand. The compression modu- 
lus is about five times the extension 
modulus. 



256 



50000 



40000 



-| r 



AXIAL STRAIN (%) 

4 5 6 7 8 



-| T 



▲ 



— r 
▲ 



i-- 106.8 pcf 
d 



IS) 
UJ 

^ 30000 


i 


en 




O 


A 


- 20000 


- 


UJ 

O 


▲ 




▲ 


10000 


'a • 




r 



Y'- 99.4 pcf 
d 



J L 



J L 



+ 2000 



UJ 
QL 

■n 

UJ 

a: 

Q. 
UJ 

o 

Q. 




STATIC TRIAXIAL COMPRESSION 
Kc = 1.0 
a3c' = 4100 psf 
Ub = 8200 psf 



-4000 



-8000 



d'^-- 99.4 pcf 



-^ ^j = 106.8 pcf 



J L 



3 4 5 6 7 

AXIAL STRAIN (%) 



10 II 



Figure 159- Typica 
Monterey "0" Sand 



1 Static Triaxial Compression Test Results for 













AXIAL STRAIN (%) 










10000 




1 


2 


3 4 


5 6 7 8 


9 


10 II 






1 


1 


1 1 

A ^ 


1 - . , 


1 


1 




in 
a. 


8000 








A 


/j = 106.6 pcf 








b 

CO 

in 

LiJ 

tr 

1— 

CO 

(E 


6 000 


- 




A 
A 


• 
• 


. • • • 

• 


• 


• 


- 


1- 
< 

> 


IT 


ff , " JO.O^ [Jul 


4 000 


- 






f 








- 


Q 


2000 


4 


A 
1 


1 


9 

1 1 


1 1 1 1 


1 


1 


- 












STATIC TRIAXIAL EXTENSION 


















Kc = 1.0 








(/) 


+ 2000 


- 








(T2,c ^ 4100 psf 
Ub = 8200 psf 






- 


a 





4 


P^ 


!^ 












3 




▲ 


^ 


^ 










3 
CO 
CO 
UJ 
IE 
0. 

UJ 


-4000 


- 




A 
A 


•• 

1 1 


/j =98.84 pcf 

•• • • • 

106.6 pcf 


• 


• 


- 


q: 
o 


-8000 


_ 


1 


1 


A A A A 

■ III 


1 


1 1 


- 



3 4 5 6 7 8 

AXIAL STRAIN (7o) 



10 II 



Figure 160. Typical Static Triaxial Extension Test Results for 
Monterey "0" Sand 



258 



16000 


- 




1 

• 
♦ 




1 


1 1 


1 


14000 


- 












- 






♦ 










12000 


- 


♦ 










- 


10000 


- 


i 








• 


- 


8000 




O 

♦ 






• 


V 




6000 


♦ 




• 


V 






4000 


-o 




DA 








- 


2000 


♦ 

3 


DA 

• 

V 


• 
V 

1 


V 


1 


1 1 


1 



2 3 4 

AXIAL STRAIN (%) 



Kc=l 


.0, 


Cr3C= 


4,100 psf 


STATIC COMPRESSION TESTS 






STATIC EXTENSION TESTS 


# TEST NO. A , S^d = 106.5 pcf 






V TEST NO.B, ^d = 106.9 pcf 


■ TEST NO. F , if d = 106.8 pcf 






• TEST NO.E, Vd = 106.6 pcf 


♦ TEST NO. N , V d = 106.6 pcf 






n TEST NO.CD.i^d = 107.1 pcf 


O TEST NO. CB, tfj = 106.4 pcf 






A TEST NO.CE, Vd = 107.1 pcf 



Figure I6l. Summary of Static Triaxial Test Results for Dense 
Monterey"0" Sand. 



259 



Cyclic Tests on Monterey "0" Sand 

Cyclic stress ratios (a, /2a\ ) vary- 
ing from 0.3 to 2.2 were used m the 
testing of very dense Monterey "0" sand 
samples. Results of these tests showed 
that the axial strain developed almost 
totally in the extension direction. 
This was true despite the fact that the 
samples were carefully prepared and com- 
posed of uniformly graded sand. This 
behavior is illustrated in Figure 162, 
which presents the cyclic strain enve- 
lopes for several tests. Observations, 
photographs and movies made during test- 
ing show that the asymetric strain is 
not caused by sample necking. Most sam- 
ples retained a uniform cylindrical 
shape until the final two or three cyc- 
les, when distinct failure occurred. 



The strain levels (1/2 peak-to-peak) 
developed in ten cycles for the tests 
shown in Figure 162 plotted against cyc- 
lic deviator stress ( o dp) in Figure 
163. Also shown are the average static 
stress-strain curves for both the com- 
pression and extension directions. The 
cyclic test results plot along the sta- 
tic extension stress-strain curve. 

After the first few cycles, the rate of 
strain increase is quite gradual for all 
of the cyclic triaxial tests depicted in 
Figure 162. This was true despite the 
fact that the peak pore water pressure 
values are at, or close to, the initial 
confining pressure and that the cyclic 
stress ratios were extremely high. The 
highest strain level reached was only 
+5 percent. For each test, despite the 




70 80 90 100 

NUMBER OF CYCLES 



Figure 162. Cyclic Triaxial Strain Envelopes for Monterey "0" Sand 



260 







T 1 T 1 

• 


- 


^ — AVERAGE OF STATIC 




COMPRESSION TESTS 


- 




\. + AXIAL STRAIN 
/^ FOR 10 CYCLES OF 


- / 




/ CYCLIC TEST 


- 




/ 


. 


> 


(^^« 




J 


^~~^AVERAGE OF STATIC 


- 


;i 


EXTENSION TESTS 


/ 

f 


• 


Kj = 10 

CTj^' = 4100 psf 

1 



AXIAL STRAIN (%) 

Figure 163. Static and Cyclic 
Triaxial Test Results for Dense 
Monterey "0" Sand 
value of the stress level, the sample 
reached a point where the rate of strain 
increase was so low that it would have 
required tens or hundreds more cycles 
to develop an additional one percent of 
axial strain. This behavior has been 
described as limiting strain by Mulilis 
et al (1975) and DeAlba et al (1975), 
who also tested samples of Monterey "0" 
sand. Mulilis performed his tests us- 
ing cyclic triaxial equipment, whereas 
DeAlba used large-scale simple shear 
(shaking table) apparatus. Both stud- 
ies found a general increase in the 
limiting strain as the relative density 
decreased. 

In Figure 164, the cyclic test results 
are presented as the number of stress 
cycles required to produce a specified 
amount of strain for different stress 
levels. Extremely high cyclic stress 
ratios (a 11 o\ ) were required to 
cause significant strains in relatively 



low numbers of cycles. These cyclic 
stress ratios ranged as high as 2.2 for 
these tests. However, tests having cyc- 
lic stress ratios greater than 0.5 have 
often been classified as having erron- 
eous tensile stresses. Many refer- 
ences, including the U. S. Bureau of 
Reclamation (1976) and Seed et al 
(1975), state that exceeding the 0.5 
cyclic stress ratio boundary can cause 
the sample cap to lift off, which may 
result in the sample failing premature- 
ly by necking near the top. 

The tests carried out for the dense 
samples of Monterey "0" sand showed 
that large cyclic stress ratios (great- 
er than 0.5) do not necessarily produce 
cap lift off and necking at the top of 
the sample. The large stress ratios 
produced uniform strains throughout the 
length of the sample, the same as for 
samples tested at much lower stresses. 
Figure 165 shows portions of the cyclic 
tesj: records for samples tested with 
cyclic stress ratios of 0.27 and 1.0. 
The only difference that may be ob- 
served in the strain development in the 
two tests is that the larger stress 
ratio produces a higher level of strain 
in the first few cycles. This charac- 
teristic was generally true for all the 
cyclic tests conducted on the Monterey 
"0" sand. The larger the' applied 
stress, the higher the strain level be- 
came in the first few cycles. 

It should be noted that all of t.he cyc- 
lic tests were continued until the 
samples eventually necked. The necking 
developed despite the fact that these 
samples were carefully prepared and 
tested. However, this was not caused 
by the cap lifting off and causing a 
neck at the top of the sample. Necking 
developed at different locations for 
different samples anywhere from the bot- 
tom of the sample to the top. In addi- 
tion, most Monterey "0" sand samples 
only indented slightly before the devel- 
opment of a shear plane. A typical 
shear plane is shown in Figure 166. 
After the development of a shear plane, 
the samples quickly necked completely. 



261 





18000 

( 


) 


1 


1 


1 1 1 


1 1 1 1 

\ 


1 1 1 1 1 1 

\ ° 


III 1 


- 


— 


16000 


_ 










\ % 




- 
















\ Vl* 






Q. 














\ ^< 




















\ ^>k 








14000^ 


_ 










V, 'Z, 




_ 


+1 


( 


\ 










\^\ 








12000 


\ 


\ 












_ 


LiJ 






\ 














a: 






X 














\- 






N 


\ 












CD 


10000 


- 




\ 










- 


(T 








\ 


>. 










O 










>y^ 










< 


8000 


- 






^ 


sO O 






- 


> 
UJ 
Q 












O^ 








6000 


— 












— 


O 

_l 

o 














^^\^o 


^^/> 
o^^^^ 




>- 
o 


4000 


- 


Kc = 


1.0 






o 


^^^o 


- 








^,- 


106 - 


- 107 


pcf 






2000 


- 


<^3c 


' = 4100 p; 


f 




o 


- 








1 


1 


1 1 1 


1 1 1 1 


1 1 1 1 1 1 


III 1 





100 



13 10 30 

NUMBER OF CYCLES 
Figure ]Sk. Cyclic Triaxial Test Results for Monterey "0" Sand 



300 



The development of this necking behavior 
is illustrated in Figure 167. 

Cyclic Tests on Oroville Sand . Ten 
cyclic triaxial tests were carried out 
for the Oroville sand samples. The 
cyclic stress ratios ranged from 0.3 
to 1.0 and the samples were cycled un- 
til they necked. 

These samples also developed predominant 
extension strain although the magnitudes 
were slightly higher than for the Mont- 
erey "0" sand samples. In addition, ob- 
servations and photographs reveal that 



the necking was different for these 
samples than for the Monterey "0" sand 
samples. Instead of developing a shear 
plane in the final stage of necking, the 
Oroville sand samples developed strain 
concentrations and indentation in the 
middle portions of the samples. These 
necks seemed to always be concentrated 
in one of the middle layers. In gener- 
al, these samples did not remain as uni- 
form as the Monterey "0" sand samples 
prior to the final stages of necking. 
Figures 168 and 169 illustrate this 
behavior. 



262 



Hi UJ 




^ en 




if) 




^ III 




s? 






CYCLES 



tr 


4100 


lU UJ 




H cc 




<=>"S 





? V) CL 




CO 1 




U UJ 3 


-4100 


cr (r 




o Q- 






-8200 




TEST NO 12 
Kc = 1.0 
D'jc=4IOO psf 
Yd = 106.6 pcf 
Odp/2CTje=0.27 



— (_ to O. 1|^ ' 

-" < Lij I n k 

o _ q: Q. , 

"Jii'^^ looooJfe 




TEST NO. 15 
K = 1.0 
Csc^ 4100 psf 
ti- 106.2 pcf 
Odp^2a,e=I.O 



Figure I65. Cyclic Triaxial Test Records for Monterey "0" Sand 



19—78786 



263 




Figure 166. Shear Plane Development during Final Stage of Necking for 
Monterey "0" Sand 



264 



a: -4 000- 



Ttnrf 



185 190 

EXTENSION PORE CHARACTERISTIC 

PRESSURE PEAKS PORE PRESSURE 

INCREASE IN DROP PEAK 





CHARACTERISTIC , 
LOAD PEAK- 
STRAIN INDICATES 

HIGHER STRESS STRAIN SHARPty DEPARTS 

'-. FROM THE AXIS 



< - 1^'- ~--^ 



BEFORE TEST 



•CYCLE 188 F 



^CYCLE 192 E 



^CYCL E 193 E 





TEST NO. 14 {i^= 106.9 pcf, K^= 1.0, cr^^' = 4100 psf) 
Figure 167. Cyclic Triaxial Test Records for Monterey "0" Sand 



265 



a: 
o o 

o — 
> > 

O UJ 
Q 




(/) 


^-4000 


(/) 


a. 


Ijj 




tr 


'q. 



< 

cc 



4000- 
-10- 

I 
uJ 0- 

5- 



mm\. 



P- 



m 



TEST NO, 16 
Xd = 138.2 pcf 

CTjc - 4100 psf 



BEFORE TEST 



* CYCLE 2E 



* CYCLE 5E 




J 


SB 


1 










"l^ 




Figure 168. Cyclic Triaxial Test Records for Oroville Sand 



266 



UJ 4100- 



'-■i. .: :::V-^'~r-T:' I- T ■■■!- 



CHARACTERISTIC 
PORE PRESSURE 
PEAK 



CHARACTERISTIC 

LOAD PEAK 



O o. 
< b 

> 6000- 



STRAIN INDICATES 
HIGHER STRESS -^ 



STRAIN SHARPLY 
DAPARTS FROM 
THE AXIS 



BEFORE TEST 



# CYCLE 14 E 



)(■ CYCLE 26 E 




m 
m 




} 




^ Mm"^ 


i 


IJU 


1 


'« jEji 


o 


r' 


1 "^^^^ -m= 




1 - 



TEST NO. 25 ( Kq = 142.9 pcf, Kc = 1.0, (r ^^ ^100 psf) 
Figure 169. Cyclic Triaxial Test Records for Oroville Sand 



267 



Analysis of Test Results 

Extension Strain 

The results of the cyclic tests show 
that predominant extension strain is 
not unusual for isotropically- 
consolidated samples of dense, cohesion- 
less material. Visual observations show 
that this behavior is not a result of 
necking. 



Analysis of cyclic triaxial test rec- 
ords produced by Mulilis et al (1975) 
and other testing programs reveals that 
the extension strain is consistently 
greater than the compression strain. 
This effect increases with increasing 
density so that very dense samples 
strain almost totally in the extension 
direction. 

The asymmetry could possibly be ex- 
plained by the inherent limitations of 
the test. The stress conditions do not 
have the same relative values during the 
extension and compression halves of the 
stress cycle. Samples of higher dens- 
ities require higher cyclic stress ra- 
tios to cause significant strain levels. 
With higher cyclic stress ratios, the 
stress conditions in the two halves of 
the stress cycle become more asymmetric. 
This would explain why the extension 
strain becomes more pronounced than the 
compression strain for higher densities. 
Although the extension direction is 
weaker than the compression direction, 
an average of the two strains produced 
seems to be appropriate because it has 
been successfully related to cyclic sim- 
ple shear conditions (see Figures 175 
and 176). 

Necking Behavior 

The cause of necking is theorized to be 
non-uniformities and stress concentra- 
tions with the sample. As cycling con- 
tinues, the sample strains and will 
eventually develop a stress concentra- 
tion until all the axial strain occurs 
primarily in one location and the sam- 



ple necks. As the uniformity of the 
sample increases, a higher number of 
cycles is required to cause necking. 
This would explain why the more uniform 
Monterey "0" sand samples held together 
better than the Oroville sand samples. 

Necking can sometimes be detected in 
the test records alone. This is be- 
cause drastic necking leaves charac- 
teristic readings in the pore pressure, 
strain, and loading measurements. 
These characteristic readings are illus- 
trated for both materials in Figures 
167 through 169 and include: 

1. A sharp increase in the extension 
strain. 

2. The strain goes significantly into 
extension during compression loading. 

3. The pore water pressure drop during 
extension loading increases. 

4. Pore water pressure and axial load 
records develop characteristic shapes 
during the final stage of necking, 
when the sample separates. 

These necking symptoms develop only dur- 
ing drastic necking. The samples may 
develop necks of smaller magnitudes 
without producing these symptoms. With- 
out producing detectable symptoms in 
the test records, severe necking has 
been observed in the samples as far 
back as 12 cycles before complete separ- 
ation. Symptoms of drastic necking have 
also been found in some of the test rec- 
ords for the modeled Oroville gravel 
samples. This leads to the conclusion 
that some of the modeled Oroville grav- 
el samples developed drastic necking. 
Examples of the test records where 
drastic necking has been found are shown 
in Figures 170 through 172. 

Sample "Tension" 

Many sand samples were tested well be- 
yond the "tension" boundary of 0.5 cyc- 
lic stress ratio, but still behaved like 
samples tested at lower stress ratios. 



268 




CHARACTERISTIC PORE 
PRESSURE PEAK 



CHARACTERISTIC 
LOAD PEAK 




TEST NO 37 
y^ - 148.0 pc.f. 
K^ = I 

= 28700 p s f 



STRAIN SHARPLY DEPARTS 
FROM THE AXIS 




Figure 170. Cyclic Triaxial Test Records for Modeled Oroville Gravel 



53300 
UI 

3 In 

" °- 26650 




CHARACTERISTIC PORE 
PRESSURE PEAK 



CHARACTERISTIC 
LOAD PEAK- 




TEST NO 65 
/■<) = 148.6 pcf 
Kj =1 
o-jj' = 53,300 



SHARP INCREASE IN 
EXTENSION STRAIN- 




51- 

Figure 171 



STRAIN SHARPLY DEPARTS 
FROM THE AXIS 



Cyclic Triaxial Test Records for Modeled Oroville Gravel 



269 





hi 






q: 




o 


h- 


Q. 


_l 


co 


1 


a 






> 


tr 




u 


o 

1- 
< 


b 



53 300 1- 



26650 



-26 650 



-53300>- 



-4l00i- 



4 lOOL- 




EXTENSION PORE 
PRESSURE PEAKS 
BEGIN TO DECREASE 



CHARACTERISTIC 
PORE PRESSURE 
PEAK 

CHARACTERISTIC 
LOAD PEAK 




TEST NO. 69 
Ya = 148.9 pcf 
Kc = 10 
o-jg' = 53 000 psf 



5"- 



STRAIN SHARPLY 
DEPARTS FROM 
THE AXIS 




Figure 172. Cyclic Triaxial Test Records for Modeled Oroville Gravel 



This Is because the designation of a 
cyclic stress ratio of 0.5 as the bound- 
ary for sample "tension" and cap lift- 
off has little meaning. This definition 
was probably developed assuming that 
when the cyclic stress ratio was greater 



than 0.5, the extension stress would be 
greater than the effective confining 
pressure and the sample cap would have 
to lift off. This would be a total 
stress definition. Soil behavior, how- 
ever, is controlled by effective stresses 



270 



The idea of a constant "tension" bound- 
ary throughout a cyclic triaxial test 
is incorrect. During a cyclic triaxial 
test, the residual pore water pressure 
at the end of each complete stress cycle 
tends to increase with each applied 
cycle. As the residual pore pressure 
approaches the chamber pressure, the 
effective confining pressure is reduced. 
The cyclic load, however, remains con- 
stant. Thus, if an isotropically con- 
solidated sample is cycled long enough 
to approach initial liquefaction, it 
experiences a "tension" condition re- 
gardless of the cyclic stress ratio 
being applied. 

The question that must be addressed is 
why does the sample hold together during 
"tension" and how does this relate to 
actual soil behavior during earthquake 
loading . 

First it should be noted that the static 
extension test produced normal uniform 
sample behavior up to a stress ratio 
(a(jp/2a'o ) of 1.0; and could have gone 
higher if a higher back-pressure had 
been used. 

In Figures 167 through 172, which show 
cyclic test results, the pore water 
pressure develops into a repetitive 
steady-state pattern after the first 
few cycles. Examination of the steady- 
state pore pressure patterns presented 
reveals that, as the cyclic stress 
curve crosses the zero axis, the pore 
pressure approaches the chamber pres- 
sure, and the effective confining pres- 
sure drops to virtually zero. At this 
time, the sample begins to strain quite 
rapidly. As the sample strains, the 
pore pressure begins to drop and the 
sample strain begins to level off. Most 
of the strain develops at relatively low 
percentages of the applied stress. This 
behavior is the same for both extension 
and compression halves of the stress 
cycle. The main differences between the 
two halves of the stress cycle during 
this steady-state pore pressure pattern 
are in the magnitudes of the pore pres- 
sure drop and the amounts of axial 




-15000' \— 




Kc = 10 

>^d = 106.7 pcf 

o-j'j.: 4100 psf 

Figure 173. Extension/Compression Cycle 
for Monterey "0" Sand Cyclic Triaxial 
Test 



271 



strain. For cyclic triaxial tests on 
dense isotropically consolidated sam- 
ples, the axial strain is concentrated 
in the extension direction, and the pore 
pressure drop in the extension direction 
is approximately four times the drop in 
the compression direction. 

The pore pressure drop is what holds the 
sample together. Without the drop in 
pore pressure, the sample would experi- 
ence unlimited strain in either direc- 
tion of loading. The drop in pore pres- 
sure has often been explained by the 
tendency of the sample to dilate. How- 
ever, the drop in pore pressure during 
the extension half of the stress cycle 
could possibly be caused by an erroneous 
feature of the cyclic triaxial test. 
If the cap lifted off, the resistance 
to extension loading would be a result 
of suction on the water alone and not 
represent actual sample behavior. Neck- 
ing might not result; but the test would 
no longer represent a shearing test. It 
is very important to note that, if this 
behavior exists, it exists for every 
isotropically consolidated cyclic tri- 
axial test that approaches initial 
liquefaction. 

Although the possibility of this erron- 
eous "suction" behavior exists, it is 
not believed responsible for the behav- 
ior of the sample. Instead, it is pre- 
sumed that the extension half of the 
stress cycle is actually analogous to 
lateral compression. This idea is sup- 
ported by the fact that the extension 
and compression halves of the stress 
cycle yield similar patterns of pore 
pressure change. In Figures 173 and 
174 are detailed plots of single stress 
cycles for two cyclic triaxial tests on 
Isotropically-consolidated samples of 
dense Monterey "0" sand. One cycle be- 
gins with compression and the other be- 
gins with extension. It may be seen 
that a pore pressure rise occurs during 
the initial loading in either direction. 
Then, after the sample has experienced 
axial strain, the pore pressure drops. 
Although the magnitude of the drops are 



different for the extension and compres- 
sion directions, the general behavior 
is the same. This same behavior is 
shown in Figures 159 and 160, which de- 
pict static extension and compression 
test results. Every time axial stress 
is applied in either direction, in sta- 
tic or cyclic loading, the pore water 
pressure rises first and then drops 
with increasing strain. 

The cyclic triaxial test behavior can 
also be related to static test behavior 
by the development of strain. Results 
of studies by Mulilis, et al (1975), 
DeAlba, et al (1975), Seed and Lee 
(1966), and many others show that cyclic 
triaxial tests develop higher strains 
for samples composed of lower densities. 
Examination of the static test results 
presented in Figures 159 and 160 re- 
veals that, to produce the same amount 
of pore pressure drop, samples of lower 
densities require much more strain. 

The testing system that has been consi- 
dered the best measure of the deforma- 
tion potential of isotropically consoli- 
dated samples is the large-scale simple 
shear (shaking table) device. Compari- 
sons between shaking-table and cyclic 
triaxial test results carried out for 
isotropically consolidated samples of 
Monterey "0" sand are depicted in 
Figures 175 and 176. The shaking-table 
results are consistently weaker than the 
results of the cyclic triaxial test. 
The ratio of the two strengths ranges 
between 0.5 and 0.6 for the conditions 
depicted. This range of cyclic strength 
ratios is consistent with the theoret- 
ical range of 0.55 to 0.70 developed by 
Seed and Peacock (1970) for the C cor- 
rection needed to account for the diff- 
erence in stress conditions. 

The similar pore pressure tendencies 
exhibited in both the static and cyclic 
triaxial tests, the observed effects of 
sample density on both static and cyc- 
lic triaxial sample strains, similarity 
of sample behavior at cyclic stress ra- 
tios ( Oj /2a' n ) above and below 0.5, 
dp Jc 



272 



15000 



en 




en 


10000 


LlI 




a: 




h- 




(f) 


5000 


tr*- 




oi^ 




H- I 




< 1 





> ^ 




U b 






-5000 


O 




-I 




o 




>- 


-10000 


o 





-15000 




5000 t— 



cn 
to 

UJ 

cr 

Q. 
UJ ' 

I- 3 
<: 



-5 000 



-10000 




-10 




Kc --I.0 

}fd = 106.2 pcf 

0-3 p' =4100 psf 

Figure 1 7^+ . Compression/Extension Cycle for Monterey 
Sand Cyclic Triaxial Test 



273 



INITIAL LIQUEFACTION 



5% SHEAR STRAIN 



Cr = 0.62 0.54 




I 
O 0.5- 



T 1 1 1 1 1 1 1 r 

10 30 50 70 90 

10% SHEAR STRAIN 



0.4- 



w 0.3- 



0.2- 



0. I- 



Cr = 0.53 








15% SHEAR STRAIN 






Cr= 0.51 


0.5- 






0.4- 




1 


0.3- 




/t . 






/ 


/ 






/ 


/ 


0.2- 




/ 
/ / 
/ / 


/ 


1 - 




/ y 






/ 






^- 


rill 1 1 1 1 



Figure 1 
for 5 



RELATIVE DENSITY % 
NOTES = 

A = TRIAXIAL TEST (o-rip/2a-3c' ) - MULILIS ET AL (1975) 
• = SHAKING TABLE TEST {T\,^/(t^) - DE ALBA ET AL (I975) 

75. Comparison for Shaking Table and Cyclic Triaxial Test Results 
Cycles 



274 



INITIAL LIQUEFACTION 



5% SHEAR STRAIN 



Cr=0.63 




lo 

I 

o 



q: 



30 50 70 

10% SHEAR STRAIN 



0.5 



LlJ 0.4 



0.3 



0.2 



0.1 



c 


r=0.57 


/ 


/^ 1 1 1 1 1 1 1 1 



0.4 - 



0.3 



0.2 - 



0.1 - 




10 30 50 70 90 

15% SHEAR STRAIN 



0.5 


Cr = 0.53 
1 


0.4 


- 


0.3 


_ 






i 
/ 


^/ 


0.2 


/ 


y 




/ ,^ 1 




/ y 




0.1 


/ y^ 






/ ^^ 




^^ 1 1 1 1 1 1 1 1 



30 50 70 90 10 30 50 70 

RELATIVE DENSITY (%) 



A = TRIAXIAL TEST ( Qjp / aCTjc ) - M ULI LI S ET AL (1975) 

• = SHAKING TABLE TEST {Z^^/Q'^) -DE ALBA ET AL ( 1975) 

Cr= (Fhv/CTo') / (CJdp/aj'c) 

Figure 176. Comparison of Shaking Table and Cycle Triaxial Test Results for 
10 Cycles 



275 



and the comparisons between the cyclic 
triaxial test and shaking-table results 
all indicate that the cyclic triaxial 
test can be used to estimate the deform- 
ation potential of a soil during cyclic 
loading. In addition, the development 
of a shear plane during cyclic loading 
indicates that a shearing behavior is 
indeed taking place. Since the question 
of sample "tension" eventually occurs 
in every sample, it would seem that cyc- 
lic stress ratios greater than 0.5 are 
just as valid as lower stress ratios. 

Cyclic Strength Interpretations 
Considered 

Because of the uncertainties generated 
by the cyclic triaxial tests performed 
for the modeled Oroville gravel samples, 
two different strength interpretations 
were considered. The two interpreta- 
tions are contrasted by different judg- 
ments concerning strain development. 
Strength interpretation I was based up- 
on the observation that the isotropical- 
ly consolidated gravel samples did not 
seem to develop significant strain 
levels at low confining pressures and 
that the static strength might be appro- 
priate to use at these consolidation 
stresses. Strength interpretation II 
was developed using the results of the 
laboratory investigation of dense sands 
to interpret and extrapolate the test 
results of the modeled Oroville gravel 
samples. 

Strength Interpretation I 

The results of the cyclic triaxial tests 
for the modeled Oroville gravel are sum- 
marized in Appendix K. The assumptions 
used in defining the strain levels, load 
levels, corrections, and other para- 
meters will be included in the final 
report on testing (Appendix L) , which 
will be available on request when 
completed. 

The results of these tests were convert- 
ed into the cyclic strength envelopes 
shown in Figure 177. The strengths are 
designated as the shear stress required 



to cause a specified strain in six cyc- 
les. A C correction of 0.6 for tests 
with K^ = 1, and a (ji'of 41.5° were used 
in the conversion. 

Static undrained strength results were 
used to define the cyclic strengths for 
alpha values of and 0.1 at low consol- 
idation pressures because: 

1. The samples that were isotropically 
consolidated at 196 and 784 kilopas- 
cals (41,000 and 16,400 psf) devel- 
oped only a limited amount of strain 
— generally less than +5 percent. 

2, The critical confining pressure is 
about 800 kilopascals (about 17,000 
psf) based on static triaxial tests 
done by the U. S. Army Corps of 
Engineers in 1964. 

At consolidation pressures lower than 
the critical confining pressure, static 
test samples have very high strength 
because negative pore pressure develops. 
The basic assumption of interpretation 
I is that negative pore pressure will 
also develop in cyclic tests, and there- 
fore the cyclic strength has to be as 
high as static strength. This assump- 
tion seems to be verified by the limit- 
ed strains that developed in cyclic 
tests at low consolidation pressures. 

Strength Interpretation II 

The assumption for interpretation I is 
probably valid for liquefaction consid- 
erations, i.e., at confining pressures 
less than critical, a cohesionless ma- 
terial probably will not liquefy if sub- 
jected to cyclic stresses lower than the 
static shear strength. However, lique- 
faction is not a concern for the Oro- 
ville gravels. The cyclic triaxial 
samples never showed any tendencies to 
develop sudden unlimited strains. The 
objective of the cylic testing was to 
define strain behavior as related to 
consolidation stress conditions and 
cyclic stress levels. 



276 



As shown in Figures 159 and 160, posi- 
tive pore pressure develops at low 
strain levels in static tests, even for 
very dense samples and low consolida- 
tion pressure. Furthermore, in the 
shaking table tests reported by DeAlba, 
positive pore pressures developed and 
eventually reached the value of the 
overburden pressure. In both cases, 
the pore pressure then decreased during 
further increase in strain. The pore 
pressure drop is the mechanism which 
prevents liquefaction. However, the 
sample can strain to an extent consis- 
tent with the effective stresses that 
develop. 

Cyclic triaxial test behavior was comp- 
licated by the severe load attenuation 
and necking problems encountered in the 
testing program. For example, many of 
the samples that did not develop large 
amoimts of strain had either significant 
load attenuation or were not tested to 
large numbers of cycles. Tests at the 
higher consolidation pressures that did 
develop large strains also exhibited 
symptons of necking behavior. 

If the samples consolidated at the lower 
consolidation pressures were tested at 
higher stresses and numbers of cycles, 
higher cyclic strain levels would have 
been produced. Based on the tests of 
dense Monterey "0" sand, cyclic strain 
would be expected to increase propor- 
tionately with an increase in cyclic 
stress. This would not lead to a sud- 
den jump in shear strength envelope as 
is the case for interpretation I. Thus, 
it is probable that the actual cyclic 
strength values at alphas of 0.0 and 0.1 
at the lower confining pressures are not 
as high as the static strength values 
presented in Figure 177. 

The second strength interpretation was 
developed from the cylclic triaxial test 



records. It was assumed that the in- 
crease in strength due to the calibra- 
tion error canceled out the reduction 
in strength due to the membrane correc- 
tions. However, the consolidation 
stress ratios (K ) were corrected for 
the calibration change. This assumption 
was judged to be conservative. 

Axial strain was defined as follows: 
cumulative peak compressive strain for 
anisotropically consolidated samples; 
one-half peak to peak strain for iso- 
tropically consolidated samples. For 
tests with load attenuation or necking, 
strain curves were extrapolated to high- 
er cycles based on the early portions 
of the tests (prenecking or preattenu- 
ation). Typical extrapolations of 
strain are shown in Figure 178. 

The conservatism shown in the figure was 
used to account for the severe load at- 
tenuation and necking behavior experi- 
enced in the isotropically consolidated 
cyclic tests. The strain extrapolations 
for the remaining tests are shown in 
Appendix H. Presented in Appendix I are 
the cyclic-stress vs. number-of-cycles 
curves developed for this second strength 
interpretation . 

The resulting cyclic strength envelopes 
for 5- and 10-percent strain in 10 cyc- 
les are shown in Figure 179. The proce- 
dures used in developing these cyclic 
strength envelopes are illustrated in 
Appendix J. AC correction of 0.63 
(for K =1.0 tests), and a *'of 44° 
were used in the conversion. It should 
be noted that this second strength in- 
terpretation is judged to be conserva- 
tive, because cyclic strain envelopes 
have a tendency to level off as cycling 
continues. A straight line extrapola- 
tion, therefore, can be considered 
relatively conservative. 



277 



2.5 7o RESIDUAL AXIAL STRAIN IN 6 CYCLES 



10 











, _ -, r r- < 





. 








_^_^ 


^^j^^^^]Lj^ — 


" 










^=06,^ 









^::::^^ 




H 


^=a£jji. 























5 10 15 20 25 30 35 40 

NORMAL STRESS ON FAILURE PLANE DURING CONSOLIDATION O^c^^q /<=m2 



_i 

Q. 

LlI 



o> 6 



5% RESIDUAL AXIAL STRAIN IN 6 CYCLES 















_-— J 





A 


) 


\ 




^'-^ 


^^^ 


^\ -— ' 


,^'-— ^ 






\ 




-"""6: 


, q7 , Cr 










\ ^ 


<;^ 




^^=00,0132- 


-"■■"^ 






::^ 














.*'^^^^" 


\^--^ 















10 



15 



20 



25 



30 



NORMAL STRESS ON FAILURE PLANE DURING CONSOLi DATiON , 0",c, K- /cm^ 



35 40 



10% RESIDUAL AXIAL STRAIN IN 6 CYCLES 



I- 



























c/-:^^ 




br^ 














— ■ — 






\ 


^ — ' 








\ 


v^^ 


^.---^ 




^^oo^ 


.36- — ' 


— 




^ 


^ 




<^ 






^ 


-^^^ 


\. 




^ 








,<^ 

































5 10 15 20 25 30 35 40 

NORMAL STRESS ON FAILURE PLANE DURING CONSOLl DATION , CTfc.Kq /cm^ 

Figure 177. Cyclic Strength Envelopes for Strength Interpretation I - 
Static and Cyclic Test Results 



278 



-20 



-15 - 



,-10- 



^2.5 



'10 



= 6 
= 9 
= 27 




AMPLE PRESUMED 
NECKED 






TEST NO. II 
K, = 1.0 



-4- 



-+- 



ID 15 20 

NUMBER OF CYCLES 



25 



PREDOMINANT AXIAL CYCLIC STRESS a: ± 10900 psf 
FOR FIRST CYCLES 



10 



-20 




TEST NO. 76 

Kc =1.0 

CTj^' = 16400 psf 



50 75 100 

NUMBER OF CYCLES 



125 



PREDOMINANT AXIAL CYCLIC ST RESS =» ± 7000 psf 
FOR FIRST CYCLES 



10 



Figure 178. Typical Extrapolations of I sotropi cal 1 y -Consol idated Cycl Ic 
Triaxial Tests on Modeled Oroville Gravel 



279 



•20000 



1 1 1 1 1 r 

5% COMPRESSIVE STRAIN IN 10 CYCLES -{ 




20000 40000 60000 

NORMAL STRESS ON FAILURE PLANE 



DURING CONSOLIDATION 



( psf ) 



20000 



CO -^ 




ir> a. 


16 000 


iij s. 




^< 




en ' 




UJ 


12000 


en 2 




< < 




UJ _) 




I a. 




<n 


8000 


UJ 




u fr 




"i => 




o -J 


4000 




20000 40000 60000 

NORMAL STRESS ON FAILURE PLANE 
DURING CONSOLIDATION, o-,^ ( psf ) 

Figure 179. Cyclic Strength Envelopes for Strength 
Interpretation II - Extrapolated Cyclic Test Results 



280 



9. EVALUATION OF PERFORMANCE 



General Considerations 



Method of Evaluation 



So far, this chapter has dealt mostly 
with the concerted efforts to determine 
the input properties and conditions for 
a complete dynamic evaluation of embank- 
ment performance during the Reanalysis 
Earthquake. A fairly wide range of in- 
terpretations was possible for dynamic 
shear modulus and cyclic shear strength. 
Other properties were found to have 
only a minor effect on computed stress- 
es. Thus, even though the procedures 
used are sophisticated — the current 
state-of-the-art — the inability to 
define all input properties closely 
limits the confidence level of the 
results. 

This is not meant to imply that the 
dynamic analysis procedures are infer- 
ior to other methods of evaluating seis- 
mic performance of dams. It is only 
meant to emphasize that a fairly wide 
range of answers will often be found, 
even with the most diligent efforts to 
measure or otherwise determine dynamic 
material properties and other conditions 
affecting embankment behavior. Gener- 
ally, other methods of analysis suffer 
the same limitations as dynamic analysis 
procedures, plus additional shortcomings 
of their own. Carefully documented ob- 
servations of dam performance during 
strong earthquakes are needed. 

Meanwhile, each engineer tries to apply 
the lessons learned from the few dams 
shaken by moderately strong earthquakes. 
Unfortunately, there is not always 
agreement among dam design engineers on 
just what the lessons are from a given 
set of observations. One area of gener- 
al agreement is that evaluation of seis- 
mic stability of dams requires the exer- 
cise of sound judgment by engineers ex- 
perienced in dam design. 



The seismic stability analysis of an 
earth dam involves four major steps: 

1. Determine the stresses induced into 
the soil in the field — both static 
and dynamic. 

2. Simulate as closely as possible these 
stresses on samples of similar soil 
in the laboratory and observe the 
behavior. 

3. Extrapolate back from the laboratory 
to the field to estimate probable be- 
havior of the actual earth dam. 

4. Compare the predicted performance of 
the dam with established criteria for 
acceptability. 

The static and dynamic stress analyses 
and the laboratory testing have already 
been discussed in Sections 4, 7 and 8. 
The remainder of this section deals 
with steps 3 and 4. 

Two assumptions are needed to relate 
test sample stresses to field stresses 
— that the failure planes are known 
in both cases, and that the irregular 
field stress time history can be repre- 
sented by an equivalent uniform stress 
time history. 

Failure Planes 

It is assumed that horizontal planes in 
field elements should be related to 
failure planes in test samples. For 
consolidation conditions, normal and 
shear stresses must be the same on field 
horizontal planes and sample failure 
planes. For cyclic loading conditions, 
cyclic shear stresses on these planes 
are compared in order to relate field 



281 



behavior to test sample behavior. 
This comparison is usually in the form 
of a safety factor or strain potential 
for each field element. Either form is 
a measure of the amount of strain that 
a test sample would develop if subjected 
to the same stress history as the field 
element. 

For triaxial test samples consolidated 
anistropically, the failure plane is 
assumed to be inclined at an angle of 
4) = 45 + (|)'/2 degrees from the major 
principal plane. For samples consoli- 
dated isotropically, ^ is assumed to be 
45 degrees, and the correction factor 
C is used to account for the inability 
of the triaxial test to duplicate all 
the field stresses correctly. 

Equivalent Regular Stress Time History 

Two procedures have been developed for 
converting an irregular shear-stress 
time history to an equivalent regular 
shear-stress time history (Lee and Chan, 
1972, and Seed et al, 1975). Both pro- 
cedures were derived from essentially 
the same basic assumption — that the 
irregular, and equivalent regular, 
stress patterns would produce the same 
accumulated strain. The amount of 
strain produced by each cycle is related 
to the stress level of the cycle and as- 
sumed to be independent of its location 



within the time history. However, stud- 
ies by Harder (1977) show that computed 
equivalent regular stresses may vary by 
as much as 30 percent, depending on the 
location of the higher peaks within the 
irregular pattern. 

Weighting curves are used to determine 
the relative contribution of each cycle. 
The cylic stress (t ) vs. number of cyc- 
les (N) curves from the cyclic tests 
are used as weighting curves. Each lo- 
cation (element) within the embankment 
requires a different T vs. N curve 
specifically for the consolidation 
stress conditions of that location, 
thus requiring many interpolations from 
the test curves. Also, triaxial test 
curves have different shapes than do 
simple shear test curves. 

Variations in choice of procedures, 
weighting curve interpolations, and 
test method can lead to different re- 
sults. To remedy this situation. Seed 
et al^ (1975), have presented a universal 
weighting curve based on large-scale 
shaking-table tests on sand. 

This curve, shown in Figure 180, was 
used in the calculations of equivalent 
uniform shear-stress time histories. 
Two different combinations of regular 
stress level and number of cycles at 
this level were computed : 

Combination No. 

12 1 



Ratio of regular stress to peak 
irregular stress 



0.65 



0.5 



Number of regular stress cycles 6 

Combination No. 2 was used in the performance evaluations. 



10 



No attempt was made in this study to 
account for the location of the larger 
stress peaks within the time history. 

Cases Analyzed and Assumptions 

One assumption made at the beginning of 
this study was that only the submerged 
upstream shell would be of concern. 



The downstream shell is unsaturated and 
is assumed to have essentially full sta- 
tic drained strength, which is much 
greater than the cyclic undrained 
strength of the upstream shell. Equal- 
ly pertinent, observations of embankment 
performance in earthquakes, theory, and 
judgment all lead to the conclusion that 
a well-compacted, dry rockfill or gravel 



282 



1. STRESS LEVEL, T/Tmax. 

2. NUMBER OF CYCLES REQUIRED TO CAUSE 
LIQUEFACTION, N. 

3. STRESS LEVEL WITH SAFETY FACTOR OF L5 ON rmax(%) 

4 LABORATORY TEST VALUES REDUCED TO ACHIEVE A 
SAFETY FACTOR OF L5. 

5 ORIGINAL LABORATORY TEST VALUES FOR CYCLIC SIMPLE 
■ SHEAR TESTS T -SHEAR STRESS , r ^ax. MAXIMUM 

SHEAR STRESS. 



- 100 




® 



500 1000 



Figure l80. Representative 

Required to Cause Liquefaction (Seed et al , 1975) 



Relationship Between t/t and Number of Cycles 



283 



embankment will perform well in a strong 
earthquake. 

The core is a well-compacted, clay- 
gravel material. It should perform as 
well as compacted clay embankments that 
have withstood strong earthquake shaking 
without any detrimental effects (Seed 
et al, 1978). 

Many of the input properties and anal- 
ysis conditions could vary over a wide 
range without significantly affecting 
the predicted behavior. Some, such as 
material density, are well defined. 
However, there are four items which can 
vary over a wide range, which have a 
major effect on the predicted behavior, 
and on which there are differences in 
opinion as to the best defined or most 
reasonable value: 

1. Dynamic shear modulus of shell 
material. 

2. Cyclic shear strength of upstream 
shell material. 

3. Abutment restraint (3D) effects on 
dynamic shear stresses. 

4. Degree of drainage in the upstream 
shell during earthquake shaking. 

The influence of items 1 and 3 on stress- 
es has already been examined in Section 
7. On item 2, two different strength 
interpretations have been discussed in 
Section 8. Item 4, drainage, has not 
been discussed previously, but it was 
observed in Section 3 that in the Aug- 
ust 1 earthquake, pore pressure in the 
upstream transition rose 90 kilopascals 
(13 psi) and then dissipated during the 
six second gap in the record. 

Strain potentials in the upstream shell 
were computed for four cases involving 
different assumptions for these proper- 
ties and conditions. Other properties 
and conditions were either well defined 
or were chosen from the conservative end 
of the defined range. (Conservative 
means that the value chosen produces the 



highest strain potential of any value 
in the range.) The rationale for each 
case is supportable, and to a large de- 
gree is a matter of judgment, or of 
philosophy in dealing with level of 
risk. Strain potential contours for 
all four cases are shown in Appendix M. 

Case a 

Shell K = 165 

Strength Interpretation II (lower), 
but with consideration of effect 
of conservatism in data extra- 
polations and of possible strength- 
ening effect of seismic history. 

Abutment restraint (3D) effects 
included. 

Drainage effects considered 
qualitatively. 

In the authors' judgment, these values 
and conditions are the most supportable 
choices based on considerations of the 
data and evidence developed in this 
study, and many other studies over the 
last few years. Case a is called the 
"best judgment case." 

Case b 

Shell K ^ = 205 

Strength interpretation II (lower) 

Abutment restraint (3D) effects NOT 

included 
NO drainage 

Each of these choices is at the end of 
the defined range which produces the 
higher estimated displacement. Case b 
is called the "conservative case." 

Case c 



Shell K^^ ^ = 205 

Strength™ Interpretation I (higher) 

Abutment restraint (3D) effects NOT 

included 
NO drainage 

The choices for the first two items 
represent the viewpoint of some of the 
many engineers who contributed to this 
study. For the last two items, the 



284 



usual conservative assumptions were 
used. 

Case d 

Shell K X " -^-^^ 

Strength interpretation II (lower) 

Abutment restraint (3D) effects NOT 

included 
NO drainage 

For the last two items the usual conser- 
vative assumptions were used. The first 
two items are from cyclic triaxial tests 
of 30-centimetre (12-inch) remolded sam- 
ples. Several studies have shown that 
dynamic strength and shear modulus are 
higher in situ than in remolded samples. 
The assumption is that the strength and 
modulus from the tests, although both 
too low, give about the same strain 
potentials as the correct in situ 
strength and modulus would give. 

Comparison of Cases 

Cases c and d result in slightly higher 
strain potentials than case a; case b 
strain potentials are substantially 
higher than any of the other three. 
Therefore only cases a and b will be 
considered further in assessing be- 
havior of the dam. 

Predicted Behavior - Best Judgment Case 

The best judgment choices of input 
values and conditions have already been 
described. The reasoning for these 
choices is presented here. 

The predicted behavior is the assessment 
of permanent displacements that the re- 
analysis earthquake would cause in the 
upstream shell. 



Shell K, 



2max 



Two extrapolations were made to extend 
the natural period ratio curve to L/H of 
7. The more consistent one gives a per- 
iod ratio of 1.35 (Figure 125). The 
corresponding 2D period is 1.1 seconds, 
and range of K- is 135 to 165 
(Figure 128). ^'"^'' 



Cyclic Shear Strength 

There is probably some conservatism in 
the use of remolded samples to represent 
the behavior of in-place materials — 
even though the two-month aging tests 
did not indicate a gain in strength. 
For example, the recent past seismic 
history probably strengthened the gravel 
shell material against possible future 
earthquakes. 

Also, there is some conservatism in the 
extrapolation of the cyclic strain en- 
velopes for Strength Interpretation II. 

Three-Dimensional Effect 

Both Makdisi's work and analyses of this 
study indicate that lower shear stresses 
will develop for three-dimensional con- 
ditions than for two-dimensional condi- 
tions — particularly in the area of 
higher strain potentials in the upper 
part of the shell. Makdisi's results 
for a narrower canyon (L/H = 3) than 
Oroville give to 75 percent lower 
stresses. Estimates by this study for 
the actual Oroville Canyon (L/H = 7) 
give about 20 percent lower stresses. 
The San Fernando Dams should have had 
a much smaller or negligible 3D effect 
because the L/H were 12 and 13. There- 
fore, it is assumed that the 3D effect 
is not "built in" to the procedure 
which was developed and tested against 
the observed earthquake performance of 
the San Fernando Dams. 

Drainage 

All the strength evaluations are based 
on the assumption of completely un- 
drained conditions in the upstream 
shell. However, drainage does take 
place as indicated by the rapid dissi- 
pation of the pore pressure developed 
in the August 1, 1975 earthquake. At 
cell No. 1, a cyclic pore pressure of 
90 kilopascals (13 psi) developed early 
and dissipated during the six-second gap 
in the records. This cell is located in 
the upstream transition zone near the 
core, and indicates that the gravel shell 



285 



and transition do experience some degree 
of drainage during earthquake shaking, 
even in interior locations. Drainage 
relief of pore pressures presumably 
would be greater at locations closer to 
the surface of the slope, where the 
strain potentials tend to be higher. 

Predicted Behavior 

Stresses were calculated for a shell 
^2max °^ ^^^ corresponding to the 1.1 

PSEUDO THREE DIMENSIONAL ANALYSI 

PSEUDO Kg MAX =350 

PSEUDO CORE Gmax/Su = 1750 

SHELL K2MAX = 165 

CORE G /S = 2200 



second natural period, and for the three 
dimensional effect (abutment restraint). 
As shown in Figure 181, the resulting 
compressive strain potentials are less 
than 5 percent except for a small zone 
in the middle of the upstream shell. 
This would generally be regarded by dam 
design engineers as indicating accept- 
able behavior involving only minor 
displacements . 




PREDICTED FOR BEST JUDGMENT CASE 
NOTES ■■ 

REANALYSIS EARTHQUAKE 
COMPUTER PROGRAM LUSH 

CYCLIC STRENGTH INTERPRETATION H - EXTRAPOLATED CYCLIC 
TRIAXIAL TEST RESULTS 

UNDRAINED CONDITIONS 

Figure l8l. Computed Compressive Strain Potentials in Upstream Shell - Percent 



Calculations were not made for the ef- 
fect of higher strength or drainage be- 
cause information is not complete enough 
to quantify these factors. However, 
the strain potentials would be even 
lower, and could be described as less 
than 5 percent everywhere in the up- 
stream shell. 

An interesting question here is what 
displacements would result from the 
method of calculation used at Upper San 
Fernando Dam. If average compressive 
strain is assumed to be 2 percent, then 
average shear strain is only 3 percent, 
and a 91-metre (300-feet) high section 
would produce a surface displacement 



of 2.7 metres (9 feet), which is not 
what most engineers would think of as 
minor displacement. 

It may well be that the method of cal- 
culating displacement applies to cases 
with high strain potentials and lique- 
faction, as at Upper San Fernando Dam, 
but does not apply to cases with low 
strain potentials and no liquefaction. 
It may also be that dams of great 
height would experience substantial dis- 
placements corresponding to low strain 
levels, and that not enough experience 
is available with earthquake performance 
to realize it. It is likely a little 
of both. 



286 



From all the considerations discussed, 
it is concluded that compressive strain 
potentials will be small — less than 
5 percent — and that no slides or 
"large" movements will develop. It is 
not so clear just how "large" might be 
the displacments associated with the 
predicted strain potentials. Because of 
the great height of the dam, it is con- 
sidered conceivable that permanent dis- 
placements on the order of a metre could 
develop at the surface of the upstream 
slope as a result of the small shear 
strains within the upstream shell. 

Estimated Displacements for Conservative 
Assumptions 

The predicted behavior for the best 
judgment case is considered conservative 
in many respects, and the possibility of 
greater displacements is considered re- 
mote. Nevertheless, it is worthwhile 
to check this remote possibility to see 
how bad the situation would be if soil 
properties and conditions proved to be 
more adverse than the best judgment 
choices. The four pertinent input val- 
ues and conditions for the conservative 
case were: 



- horizontal displacement of the 

surface by a few tens of feet 
in the interval between the two 
berms. 

- slumping of the shell material 

near the upper berm. 

- bulging of the shell material 

near the lower berm. 

The displacement and slumping would be 
limited to the upstream shell material. 
Slumping would not be expected to extend 
upslope to the crest (judgment based on 
extent of slumping at Lower San Fernando 
Dam) . The compacted gravel in the up- 
stream shell would be as strong and per- 
form as well after deformation as 
before. 

Based on the behavior of triaxial test 
samples, there is no concern over sudden 
massive shear slides or liquefaction 
flow slides. Movement would occur only 
in short-duration increments and only 
during the highest peaks of earthquake 
acceleration, the several increments 
accumulating to a total of perhaps 
10 metres. 



Shell K^^ = 205 

Strength Interpretation II (lower) 

Abutment restraint (3D) effects 

NOT included 
No drainage 

The strain potential pattern and the 
method of estimating displacements are 
included in Appendix M. The extreme of 
deformations of the upstream slope 
might be as follows : 



Remember that this is the most extreme 
case that could be supported by the re- 
sults of the analysis — by adopting 
simultaneously the most conservative 
values for material properties and other 
input conditions. Although these post- 
ulated movements are uncomfortably 
large, they would not threaten the 
safety of the dam. 



287 



REFERENCES 

Becker, E., Chan, C. K. and Seed, H. B. "Strength and Deformation Characteristics 
of Rockfill Materials in Plane Strain and Triaxial Compression Tests". 
Report No. TE 72-3 to State of California, Department of Water Resources. 
1972 

Bolt, Bruce A. "Duration of Strong Ground Motion". Proceedings - Fifth World 
Conference on Earthquake Engineering. Rome, 1974. 

DeAlba, P., Chan, C. K. and Seed, H. B. "Determination of Soil Liquefaction 

Characteristics by Large-Scale Laboratory Tests". Report No. EERC 75-14. 
University of California, Berkeley. 1975 

Harder, Leslie F., Jr. "Liquefaction of Sand Under Irregular Loading Conditions". 
Thesis, submitted in partial satisfaction of requirements for the Degree of 
Master of Science in Engineering, Graduate Division, University of California, 
Davis. 1977. 

Idriss, I. M. , Lysmer, J., Hwang, R. and Seed, H. B. "A Computer Program for 

Evaluating the Seismic Response of Soil Structures by Variable Damping Finite 
Element Procedures". Report No. EERC 73-16. University of California, 
Berkeley. 1973. 

Kulhawy, F. H. and Duncan, J. M. "Nonlinear Finite Element Analysis of Stresses 
and Movements in Oroville Dam". Report No. TE 70-2 to State of California, 
Department of Water Resources. 1970. 

Lee, Kenneth L. and Kwok Chan. "Number of Equivalent Significant Cycles in Strong 
Motion Earthquakes". Proceedings, Conference on Microzonation for Safer Con- 
struction, Seattle. November 1970. 

Lowe, J. "Shear Strength of Coarse Embankment Dam Materials". Proceedings, 8th 
Congress on Large Dams, pp. 745-761. 1964. 

Lysmer, J., Udaka, T., Seed, H. B. and Hwang, R. "LUSH - A Computer Program for 
Complex Response Analysis of Soil-Structure Systems". Report No. EERC 74-4. 
University of California, Berkeley. 1974. 

Makdisi, F. I. "Performance and Analysis of Earth Dams During Strong Earthquakes". 
Dissertation, submitted in partial satisfaction of the requirements for the 
Degree of Doctor of Philosphy in Engineering, Graduate Division, University 
of California, Berkeley. 1976. 

Marachi, N. D., Chan, C. K. , Seed, H. B. and Duncan, J. M., "Strength and Deforma- 
tion Characteristics of Rockfill Materials". Report No. TE 69-5 to State of 
California Department of Water Resources. 1969. 

Mulilis, J. D., Chan, C. K. and Seed, H. B. "The Effects of Method of Sample 
Preparation on the Cyclic Stress-Strain Behavior of Sands". Report 
No. EERC 75-18. University of California, Berkeley, 1975. 



288 



REFERENCES (Continued) 

Nobari, E. S. and Duncan, J. M. "Effect of Reservoir Filling on Stresses and 

Movements in Earth and Rockfill Dams". Report No. TE-72-1. University of 
California, Berkeley. 1972. 

Schnabel, Per B. and Seed, H. Bolton. "Accelerations in Rock for Earthquakes in 
the Western United States". Bulletin, Seismological Society of America. 
Volume 63. 1973. 

Seed, H. Bolton. "Evaluation of Soil Liquefaction Effects on Level Ground During 

Earthquakes". State-of-the-art paper presented at Symposium on Soil Liquefac- 
tion, ASCE National Convention, Philadelphia. October 2, 1976. 

Seed, H. Bolton and Idriss, I. M. "Soil Moduli and Damping Factors for Dynamic 
Response Analysis". Report No. EERC 70-10. University of California, 
Berkeley. 1970. 

Seed, H. Bolton, Idriss, I. M. and Kiefer, Fred W. "Characteristics of Rock Motions 
During Earthquakes". Journal, SMFE. September 1969. 

Seed, H. B., Idriss, I. M. , Makdisi, F. and Banerjee, H. "Representation of 
Irregular Stress Time Histories by Equivalent Uniform Stress Series in 
Liquefaction Analyses". Report No. EERC 75-29. University of California, 
Berkeley. 1975. 

Seed, H. Bolton and Lee, Kenneth L. "Liquefaction of Saturated Sands During Cyclic 
Loading". Journal of the Soil Mechanics and Foundations Division, ASCE. 
Vol. 92, No. SM6, Proc. Paper 4972, pp. 105-134. November 1966. 

Seed, H. B., Lee, K. L., Idriss, I. M. and Makdisi, F. "Analysis of the Slides 

in the San Fernando Dams During the Earthquake of February 9, 1971". Report 
No. EERC 73-2. University of California, Berkeley. 1973. 

Seed, H. Bolton, Makdisi, Faiz, I., and DeAlba, Pedro "Performance of Earth Dams 
During Earthquakes". Journal of the Geotechnical Engineering Division, ASCE. 
Volume 104, No. GT7, Proc. Paper 13870, pp. 967-994. July 1978. 

Seed, H. Bolton and Peacock, W. H. "Applicability of Laboratory Test Procedures 

for Measuring Soil Liquefaction Characteristics Under Cyclic Loading". Report 
No. EERC 70-8. University of California, Berkeley. 1970. 

U. S. Bureau of Reclamation. "Dynamic Analysis of Embankment Dams". For Submission 
to the International Commission on Large Dams for Publications as a State-of- 
the Art Paper. 1976. 

Vrymoed, John, et al. "Dynamic Analysis of Oroville Dam". Final Draft of Office 

Report, Department of Water Resources, Division of Safety of Dams. June 1978. 

Wong, R. T. "Deformation Characteristics of Gravels and Gravelly Soils Under Cyclic 
Loading Conditions". Dissertation, submitted in partial satisfaction of the 
requirements for the Degree of Doctor of Philosophy in Engineering, Graduate 
Division, University of California, Berkeley. 1973. 

289 



CHAPTER VI 

SEISMIC ANALYSIS OF THE OROVILLE 

DAM FLOOD CONTROL OUTLET STRUCTURE 



Commentary 

As a result of the August 1, 1975 Oro- 
ville earthquake, of magnitude 5.7, the 
Department found it appropriate to reana- 
lyze the Oroville Dam Flood Control Out- 
let structure (Figures 182 and 183) 
using a stronger earthquake (magnitude 
6.5) and the latest techniques in seis- 
mic investigation. 

A seismic study, monitored by the Depart- 
ment, was conducted under a consulting 
agreement between Dr. Edward L. Wilson 
and the Department of Water Resources. 
The results were presented in the report, 
"Earthquake Analysis of the Oroville 
Dam Flood Control Outlet Structure", 
by Edward L. Wilson, Frederick E. 
Peterson, and Ashraf Habibullah. Their 
report is included as the final part of 
this chapter. 

The finite-element method and dynamic 
techniques were utilized in this study 
to perform a lineraly elastic three- 
dimensional analysis of the reinforced- 
concrete structure. The three- 
dimensional analysis was chosen because 
of the complexity of the structure. In 
this analysis, dead and hydrostatic 
loads were applied to the same finite- 
element model; the interaction of the 
reservoir was not included due to limi- 
tations of the present state of the art. 
However, to include the full participa- 
tion of the reservoir, a finite-element 
two-dimensional analysis with hydro- 
dynamic interaction was also performed. 
This resulted in stresses approximately 
20 percent higher. 

The modified Pacoima and Taft, 6.5 magni- 
tude earthquake accelerogram (see 



Chapter V) , was used in these analyses. 
Horizontal acceleration of the ground 
was applied parallel to the outlet 
centerline. This produced the largest 
stresses in the piers. 

The effects of a horizontal acceleration 
in the transverse direction were exam- 
ined by the Department. For this case 
the piers were assumed fixed at the 
breast wall and bottom slab. Displace- 
ment of the supports equal to the width 
of two contraction joints was assumed. 
Stresses produced were below allowable 
working stresses and therefore no 
further investigation for this case was 
necessary. 

Maximum stresses from Dr. Edward L. 
Wilson's report were utilized by the 
Department to perform a reinforced- 
concrete-theory analysis of the structure, 
which resulted in the tensile stresses 
shown in Figures 184 through 186. These 
are peak instantaneous stresses with the 
occurrence time shown at the bottom of 
each figure. The largest concrete ten- 
sile stress of any concern is 2172 kPa 
(315 psi) . It was obtained in an area 
of the piers where the reinforcing steel 
available is almost negligible (Figure 
185, Elevation 850, Station 12+68). 
However, experimental research and pro- 
totype observation have shown— that 
the dynamic tensile strength of mass 
concrete is at least 10 percent of its 
static compressive strength or, in this 
instance, 3792 kPa (550 psi); therefore, 
the peak stress of 2172 kPa (315 psi) 
at the downstream end of the piers is 
well within the tensile strength of the 
concrete. 

Maximum concrete tensile-stress values 



1/ Refer to text and references of Chapter VII ("Earthquake Response Analysis of 
Thermalito Diversion Dam" by Anil K. Chopra) . 



291 




292 




293 



values shown in Figure 184 are not con- 
sidered critical, because the rein- 
forcing steel available in that area is 
capable of resisting the total earth- 
quake tensile force without any contri- 
bution from the surrounding concrete 
(see Figure 186) . 

Although piers 1 and 10 were not investi- 
gated numerically by the Department, 
they are not expected to develop critical 
stresses because they carry half as much 
load as do the adjacent piers, and a 
significant portion of their height is 
in direct contact with the rock abutment. 

The Department also investigated the 
structure for stability against sliding 
through the shear-friction equation 



Q = 



CA + N tan . 
H 



The use of a 



cohesion value of 3447 kPa (500 psi) 
produced a shear-friction factor of 11.6, 
which is ample against sliding. 



Conclusion 

The investigations performed indicate 
that when the Oroville Flood Control 
Outlet Structure is subjected to the 
Reanalysis Earthquake ground motion^ it 
is stable, and that expected compressive 
and tensile stresses are within the allow 
able limits established for the structure 

Introduction to Figures 184 through 186 



1. Smaller tensile stresses and com- 
pressive stresses were intentionally 
left out due to their uncritical 
magnitudes. 

2. Although stresses in piers 5 and 6 
are somewhat different from those in 
piers 2, 3, 4, 7, 8 and 9, the small 
variation did not justify showing 
them independentaly. 

3. Figures 184 through 186 show the j 
elevation of a pier with an element 
mesh layout. Actual stresses are 
tabulated inside the elements. 



294 



920.17 




Maximum Tensile Stresses m 
Piers 2 thru 9 at time 7.76 sec. 

Top- Reinforcing Steel Stress ( psi ) 

Bot- Concrete Stress ( psi ) 

Figure IS'*. Maximum Tensile Stresses at Time 7.76 Seconds 



295 



920.17, g 

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14 
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p 
v> 


Maximum Tensile Stresses in c 
Piers 2 thru 9 at time 8.46 sec. ^ 







Top- Reinforcing Steel Stress (psi ) 




Bot- Concrete Stress (psi) 




Figure I85. Maximum Tensile Stresses at Time 8.^6 S 


econds 


296 

































920.17 




Maximum Tensile Stresses (ksi) in 
Piers 2 thru 9 in reinf. steel at 
time 7.76 sees, disregarding 
contribution from concrete 

Figure 186. Maximum Tensile Stresses at Time 7-76 Seconds in Steel 



297 



A REPORT TO THE 

DEPARTMENT OF WATER RESOURCES 

STATE OF CALIFORNIA 



on the 



Earthquake Analysis of the Oroville Dam 
Flood Control Outlet Structure 



by 



Edward L. Wilson 

Professor of Civil Engineering 

University of California, Berkeley 

Frederick E. Peterson, President 
Engineering/Analysis Corporation 
Berkeley, California 

Ashraf Habibullah, President 

Computers/Structures International 

Oakland, California 



June 1977 



299 



TABLE OF CONTENTS 

PAGE 

I. Introduction 1 

A. Background 1 

B. Basic Assumptions 1 

II. Three-Dimensional Analysis Without Hydrodynami c Interaction . 3 

A. Objective and Scope 3 

B. Analysis Procedure 6 

1. Idealization 6 

2. Analysis process 6 

3. Output of results 8 

C. Structural Model 8 

1. Finite element mesh 8 

2. Material properties 9 

3. Displacement boundary conditions 9 

4. Mass distribution 9 

D. Analysis Cases and Results 9 

1. Static Analysis 9 

a. applied loads 13 

b. results 15 

2. Analysis for mode shapes and frequencies 22 

3. Response spectrum analysis 22 

4. Time history response analysis 29 

a. solution parameters 29 

b. results 31 

III. Two-Dimensional Analysis With Hydrodynamic Interaction ... 44 

A. Analysis Without Interaction 44 

B. Analysis With Interaction 44 

C. Extension to Three Dimensional Analysis 46 

IV. Transverse Earthquake Analysis 47 

V. Vertical Earthquake Analysis 48 

VI. Final Remarks 49 



300 



I. INTRODUCTION 

A. Background 

The analysis presented in this report was conducted under a consult- 
ing agreement between the Department of Water Resources of the State of 
California and E. L. Wilson. The analysis was monitored by personnel of 
the Division of Design and Construction. 

The purpose of the report is to present an Earthquake Analysis of the 
Oroville Dam Flood Control Outlet Structure utilizing the finite element 
method and dynamic techniques suitable for this particular structure. It 
is not the objective of this report to assess the structural safety of the 
structure but to present the results (stresses and displacements)of a 
specific analyses based on the given loadings and stated assumptions. 

B. Basic Assumptions 

The structure to be analyzed is a complex three dimensional structure 
given in detail in drawings A-3B5-1 to A-3B5-7 which were supplied by 
Department of Water Resources. Since the three dimensional dynamic 
analysis of the complete series of monoliths is beyond the present state- 
of-the-art with respect to computer capabilites it was necessary to study 
the three dimensional dynamic behavior of a typical monolith only. For 
elastic behavior this idealization should introduce only minor errors-- 
less than 10 percent. 

The specified earthquake loading produces stresses and displacements 
which must be combined with the results of other load conditions if the 
total state of stress is to be evaluated. In order for the results of the 
various load conditions to be combined in a rational manner dead loads 
and static hydrostatic loads were applied to the same three dimensional 
finite element model and are presented in this report as separate results. 

301 



2 



Also, the upstream-downstream direction was assumed to be the 
critical direction with respect to the dynamic earthquake loading. A 
separate simplified analysis is presented to evaluate the significance 
of the earthquake loading in the transverse direction. In addition, 
potential vertical earthquake stresses must be considered separately. 

At the present time an exact three dimensional analysis of a dam- 
reservoir system is not possible. Therefore, it was necessary to estimate 
this effect by examining this effect on a two-dimensional structure with 
similar dynamic properties. 

Only linearly elastic behavior was considered. After the results 
from the various analysis are combined it will be necessary to apply 
experience and engineering judgment in order to estimate the significance 
of nonlinear behavior and to estimate the realistic behavior of the 
structure under the specified earthquake. 



302 



II. THREE DIMENSIONAL ANALYSIS WITHOUT HYDRODYNAMIC INTERACTION 

A. Objective and Scope 

This section describes the three dimensional static and dynamic 
analysis of a typical section of the mass concrete spillway structure. 
The objective of the analysis is to compute stresses in the concrete due 
to seismic excitation. 

A critical exposure condition for the spillway structure occurs when 
seismic input acts parallel to the direction of flow. For this case, 
horizontal motion is induced in the breast wall which must be resisted 
by overturning (i.e., vertical stresses) in the piers. From the propor- 
tions of the structure, it is reasonable to approximate the horizontal 
translation of the foundation as a rigid body; i.e., at any instant, the 
acceleration at the base of all piers is identical. The ends of the 
spillway structure terminate in substantial embankments; thus, it is 
reasonable to neglect horizontal motion transverse to the direction of 
flow for an earthquake acting in the direction of flow. 

A typical section of the spillway structure is bounded between two 
vertical planes as shown schematically in Figure II. 1. One plane bisects 
the pier and the other coincides with a mid-plane of the breast wall. 
The X, Y, Z reference system shown in the figure is chosen as follows. 
X is horizontal and transverse to the direction of flow. Y is vertical 
and directed up. Z is horizontal, parallel to flow and directed upstream. 
A plan view of this same section is shown in Figure II. 2. Assuming that 
the X-translation is zero on both vertical planes and that the same Z- 
direction ground input acts at the base of all piers, an analysis of the 
section shown in Figure II.l will predict a response typical of that for 
the entire structure. 

303 




FIGURE II. 1 OROVILLE SPILLWAY STRUCTURE 
SCHEMATIC OF TYPICAL SECTION 



304 



>x 




u w 



305 



B. Analysis Procedure 

The purpose of the analysis is to predict seismic stresses in the 
concrete structure. For comparison, static stresses under operating 
conditions are also calculated. 

1 . Idealization 

The typical section isolated in Figure II. 1 is idealized as a three 
dimensional elastic continuum. The structure considered is the mass 
concrete only. Non-structural masses are lumped at various locations in 
the structure; e.g., the gate mass is lumped at the trunnion. Added mass 
due to structure/reservoir interaction is not considered in the three 
dimensional analysis. 

An elevation view of a typical section is shown in Figure II. 3. 
Ground acceleration is applied through the base of the pier and is 
directed in the global Z-direction. The structure is free to deform in 
the Y,Z plane, but is restrained against displacement normal to the 
vertical planes: X = ( pier) and X = 11.292' ( breast wall). 

2. Analysis Process 

The three dimensional stress analysis was performed using the linear 
elastic finite element computer code known as EASE2. Usage and theoretical 
basis for EASE2 are described in the reports: 

(1) "EAC/EASE2 - Elastic Analysis for Structural 
Engineering", User Information Manual, 
Control Data Corporation, Cybernet Services, 
Publication No. 84002700, Revision B (6-15-76). 

(2) "EAC/EASE2 - Dynamic Analysis for Structural 
Engineering", User Information Manual, 
Control Data Corporation, Cybernet Services, 
Publication No. 84000040, Revision A 
(6-15-76). 

The EASE2 analysis was performed in the sequence outlined below: 

(1) prepare a finite element mesh consisting 
of the three dimensional, eight-node solid 
elements; 



306 



\^<^)(b.e>a' h»— 



i(h.i^' 




&Ge>.ib' 



3\^.(bO' 



&\\.e>o' 



FIGURE II. 3 OROVILLE SPILLWAY STRUCTURE 
ELEVATION OF TYPICAL SECTION 



307 



(2) lump non-structural masses at corresponding 
locations in the model; 

(3) calculate static loads due to gravity, 
hydrostatic pressure and trunion prestress 
and analyze the structure for operating 
condition static stresses; 

(4) analyze the structure for mode shapes and 
frequencies; 

(5) knowing the structure frequencies, calculate 
the spectral response due to the earthquake 
time history (modified Pacoima and Taft, 20 
seconds duration); 

(6) to locate regions of critical stress, perform 
a response spectrum analysis of the structure 
assuming Z-direction input at the base of 
the pier; 

(7) requesting output at locations of high 
stress, calculate the dynamic response of 
the structure due to Z-direction ground 
input acting at the base of the pier. 

3. Output of Results 

Complete output from a three dimensional time history analysis is 
enormous. To reduce print-out to manageable proportion, it is necessary 
to limit output to those quantities of primary concern. For the spillway 
structure, this primary quantity is principal tension in the concrete. 
The static and response spectrum analyses produce complete spatial 
distributions of stress and thus can be used to locate regions of high 
stress. Output from the seismic response analysis is limited to principal 
tension plotted versus time, one plot for each location of potentially 
critical stress. From the plots, the times at which maximum stresses 
occur can be determined. Finally, complete distributions of principal 
tension are displayed at those times at which maxima were found to occur. 

C. Structural Model 



1. Finite Element Mesh 



308 



The section of Figure II.l is modeled with 285 eight-node solid 
elements. The mesh is built up of four vertical layers of elements as 
shown in Figure 11,4, Figure 11,5 plots the +X face of elements l-to-177 
which represent the pier, and Figure II. 6 is an enlargement of one of the 
three breast wall element layers. 

2, Material Properties 

The modulus of elasticity and Poisson's ratio of the concrete are 
taken as 5000 ksi and 0,2, respectively. The weight density of the 
concrete is 160 pcf. 

3, Displacement Boundary Conditions 

The base of the pier is fixed. Transverse X-displacement is zero 
for all nodes in the pier mid-plane and breast wall mid-plane. These 
boundary conditons allow free Y,Z translation of the structure at all 
nodes above the base, and X- translation is free at all nodes except 
those on the vertical boundary planes, 

4, Mass Distribution 

The total mass of the structure model is the sum of the mass due 
to element volume (i.e,, structural mass) plus non-structural added mass. 
The total mass due to element volume was computed by EASE2 as 3906 , 
This is 6% lower than the working drawing estimate of 4153 (corrected 
to 160 pcf concrete). The total non-structural mass is 337 , Each non- 
structural mass item (e.g., inspection walkway, maintenance deck, etc.) 
has its mass distributed to the node or line of nodes nearest its actual 
location in the structure. 

D. Analysis Cases and Results 
1 . Static Analysis 
Three static loading conditions: 



309 




I 



/state: of CRLIFORNIfi/DEPRRTMEHT OF HRTER RFSOURCES/OROV ILLE ORM SFILLWRY/ 
/STRTIC AND OrtlRMIC THREE DldENSlONRL FINITE ELEtlENT RNRLYSIS/ 
ISOMETRIC VIEW OF TOTRL MODEL -- NODE AND ELEMENT NUMBERS SUPPRESSED 



EflC/CSI/ERSEZ/E2FL0r 



VIEW NUMBER I 



FRAME NUMBER 1 



RUN DATE lO/M/76 



FIGURE II. 4 OROVILLE SPILLWAY STRUCTURE 
3/D FINITE ELEMENT MODEL 



310 



I 



n 




/STRTE OF CRLIFORNIR/DEFRRTMENT OF WPTER RESOURCES/OROV ILLE DRM SPILLWAY/ 
/STRTIC AND OYNRtllC THREE DIMENSIONAL FINITE ELEMENT RNRLYSIS/ 
/PROJECTION OF THE STRUCTURf ON THE Y-Z PLANE/ 



EAC/CS1/EASE2/EZPL0T 



VIEW NUMBER Z 



FRAME NUMBER 1 



RUN DATE 10/15/76 



FIGURE II. 5 



OROVILLE SPILLWAY STRUCTURE 
PIER ELEMENT LAYER 



311 




/SfRTE OF CRLlFORNlfl/OEFRRrnENT OF HRTER RESOURCES/OROV [LLE ORH SFILLHRY/ 
/STATIC RNO OTNRfllC THREE DIMENSIONRL FINITE ELEMENT RNBLYSIS/ 
/LRTER OF ELEMENTS BETHEEN i=4 AND IrS/ 



ERC/CSI/EPSE2/E2FL0T 



VI EH NUMBER I 



FRPME NUMBER Z 



FIGURE II. 6 



OROVILLE SPILLWAY STRUCTURE 
BREAST WALL ELEMENT LAYER 



312 



T3 

(1) gravity; 

(2) hydrostatic pressure (closed gate, water surface elevation 
of 900 feet); 

(3) trunnion anchorage prestress 

are analyzed individually and in combination. The combined loading case 
is called the "operating condition", 
a. Applied Load s 

Gravity loading is a 1-g static acceleration applied in the minus 

k k 

Y-direction. The total gravity loading is 4243 ; this includes 3906 

due to structure plus 337 due to non-structural mass items. 

Hydrostatic loads are computed assuming that the water surface is 

at an elevation of 900.00 feet. The loading is separated into two major 

parts: 

(1) pressure on the exposed surface of the mass concrete structure; 
and, 

(2) pressure on the radial gate. 

The structure is loaded by applying pressure normal to the face of all 
exposed solid elements. Integrating the pressure distribution applied 
on the finite element model, we obtain a total horizontal force of 1377 
(-Z) and a total vertical force of 752*^ (+Y); see Figure II. 7. Actual 
uplift acting on the elliptical underside of the breast wall was calculated 
as 757 . The horizontal and vertical resultants due to hydrostatic 
pressure on the gate are 1296'^ (-Z) and SSo"^ (+Y), respectively. These 
forces act on the pier at the trunnion. 

A self-equilibrating load set simulating trunnion anchorage prestress 
was applied to the pier to offset the hydrostatic gate load. Twenty four 
(24) tendons (in 1/2 of the pier) at 142.6'^ per tendon results in a total 
prestress force of 3422 . The line of action of the opposing prestress 
forces was aligned through the average trajectory of the tendon array. 



313 




'i^OO. OO -^^- / 



1^77'< 



e>\^.(2>o' 




nb 



^K^\f\V. (SAte 



FIGURE II. 7 HYDROSTATIC LOAD RESULTANTS 



314 



15 



b. Results 

Static analysis results are described in this section. Stresses due 
to gravity alone and stresses for the operating condition (i.e., gravity 
+ hydrostatic + prestress) are presented. 

The spatial distribution of stress in the concrete is displayed as 
follows. A Y,Z projection of the 177 solid elements used to model the 
pier is shown in Figure 11.8(A). Figure 11.8(B) is an elevation view 
of the same 177 elements transformed to an integer coordinate system in 
which all element faces have the same area. If element centroidal 
stresses were printed at their respective physical (Y,Z) locations, the 
result would be unreadable. Values printed in the integer system (Figure 
11.8(B)) form a regular array as maybe seen in Table II. 1. This table 
lists the vertical (Y) stress in psi calculated in each pier element for 
the gravity loading condition; negative values indicate compression. 
Maximum vertical compression is 300 psi in element 14 which is located 
in the heel of the pier. Tables II. 2 and II. 3 are distribtuions of 
minimum and maximum principal stress, respectively, due to gravity only. 
Note that the vertical and minimum principal stresses in the vacinity 
of the heel are nearly the same; this indicates that the minimum stress 
is approximately vertical in the heel region. Stresses in the breast 
wall due to gravity are low; the min/max principal stresses fall in the 
range -100 psi/10 psi. 

The principal stresses for the operating condition (sum of all 
static loads) are shown in Tables II. 4 and II. 5. One notes that principal 
stresses are not an exact sum at the principal stresses due to the indi- 
vidual load conditions; since theirdirections are not the same. However; 
near the heel where all load conditions tend to produce vertical stresses 



315 



— t 


— 1 


— \ 


1 


1 


t 1 


— \ — \- -i — \ — 1 — ^-^y 
















































\ 






































































































































































































' 


- 


'\ 


x' 






^ 


s 
















\ 


\ 


V 


\ 


\ 




N, 


.N 


n"' 














^ 


\ 






\ 




^ 


N 


X 
















X 


N 


N 


\ 


\ 




N 


\ 















W Eh 



S B ca 
M H X 

H CO 

> o 

Eh U 
Z Z El 

O M < 
H Z 

Eh D M 

!?:«9 



J rt o 

HSU 




W E 

W Eh 



< M 

> to 

M X 

►J K 

U 0. 



316 



17 



HEADING LINE THREE 
STATIC LOAD CASE NUMBER 
ELEMENT DISPLAY SET NUMBER 
SCALE FACTOR 
OUTPUT STRESS COMPONENT 



GRAVITY LOADS / VERTICAL STRESS / PSI UNITS 
II 
l> 

.IO00E*04) 
2) 



OROVILLE DAM SPILLWAY/ HALF OF TYPICAL BAY/ STATIC ANALYSIS RESULTS/ 
PIER ELEVATION VIEW/ ELEMENTS O0I-TO-IT7/ 
GRAVITY LOADS / VERTICAL STRESS / PSI 'UNITS 

SOLID ELEMENT CENTROIOAL STRESSES 



-n 


-13 


-12 


-12 


-12 


-11 


-11 


-10 


-9 


-8 








-16 


-16 


-16 


-18 


-20 


-18 


-17 


-14 


-12 


-9 








-21 


-22 


-22 


-23 


-26 


-24 


-21 


-17 


-13 


-9 








-27 


-27 


-28 


-30 


-32 


-29 


-26 


-21 


-15 


-5 








-32 


-32 


-33 


-36 


-42 


-40 


-36 


-30 


-21 


-7 


-25 






-39 


-38 


-38 


-40 


-47 


-46 


-43 


-36 


-23 


-19 


-3 






-*9 


-47 


-44 


-45 


-54 


-56 


-52 


-40 


-27 


16 


-53 


2 


-8 


-69 


-67 


-58 


-55 


-67 


-71 


-59 


-43 


-20 


-14 






-41 


-126 


-134 


-117 


-100 


-90 


-81 


-62 


-41 


-23 


22 






-46 


-226 


-239 


-217 


-174 


-122 


-88 


-66 


-45 


-18 


2 


-1 


-13 


-48 


-244 


-236 


-215 


-180 


-138 


-101 


-74 


-50 


-27 


-9 


-6 


-20 


-36 


-25A 


-240 


-216 


-184 


-147 


-112 


-83 


-58 


-36 


-21 


•16 


-22 


-26 


-265 


-247 


-219 


-188 


-154 


-120 


-91 


-66 


-45 


-32 


-26 


-23 


-23 


-281 


-252 


-219 


-190 


-159 


-127 


-98 


-74 


-54 


-41 


-33 


-27 


-18 -13 


-300 


-246 


-219 


-191 


-162 


-131 


-104 


-80 


-61 


-48 


-39 


-30 


-22 -10 



TABLE II. 1 



STATIC GRAVITY STRESSES 
VERTICAL (Y) COMPONENT 



317 



HEADING LINE THREE 
STATIC LOAD CASE NUMBER 
ELEMENT DISPLAY SET NUMBER 
SCALE FACTOR 
OUTPUT STRESS COMPONENT 



= <6RAvnY LOADS / MINIMUM PRINCIPAL STRESS / PS| UNITS 

= ( 1) 

= ( 1) 

= < .1000E*0<i) 

= < 7) 



OROVILLE DAM SPILLWAY/ HALF OF TYPICAL BAY/ STATIC ANALYSIS RESULTS/ 

PIER ELEVATION VIEW/ ELEMENTS OOI-TO-177/ 

ORAVITY LOADS / MINIMUM PRINCIPAL STRESS / PSI UNITS 

SOLID ELEMENT CENTROIOAL STRESSES 



-1« 


-13 


-13 


-13 


-13 


-12 


•11 


-10 


-10 


-10 








-16 


-16 


-17 


-19 


-21 


-19 


-17 


-14 


-12 


-11 








-21 


-22 


-22 


-24 


-27 


-24 


-21 


-18 


-13 


-10 








-27 


-27 


-28 


-30 


-33 


-30 


-26 


-22 


-15 


-5 








-32 


-32 


-33 


-36 


-43 


-40 


-37 


-31 


-22 


-12 


-32 






-39 


-38 


-38 


-40 


-48 


-47 


-44 


-37 


-28 


-26 


-9 






-49 


-47 


-44 


-45 


-54 


-57 


-53 


-43 


-31 


7 


-59 


-12 


-10 


-70 


-68 


-59 


-56 


-67 


-71 


-61 


-4 7 


-29 


-28 






-42 


-129 


-137 


-121 


-102 


-91 


-81 


-63 


-44 


-27 


-2 






-46 


-227 


-240 


-219 


-178 


-124 


-88 


-66 


-46 


-18 


-4 


-4 


-13 


-50 


-2*4 


-236 


-215 


-181 


-139 


-101 


-74 


-51 


-27 


-13 


-15 


-27 


-36 


-254 


-240 


-216 


-184 


-148 


-112 


-83 


-58 


-36 


-23 


-24 


-27 


-27 


-265 


-247 


-219 


-188 


-154 


-121 


-91 


-66 


-45 


-32 


-29 


-26 


-23 


-281 


-253 


-219 


-190 


-159 


-127 


-99 


-74 


-54 


-41 


-34 


-27 


-18 -13 


-304 


-247 


-220 


-192 


-163 


-133 


-105 


-81 


-62 


-48 


-39 


-30 


-22 -10 



TABLE II. 2 



STATIC GRAVITY STRESSES 
MINIMUM PRINCIPAL COMPONENT 



318 



19 



HEADING LINE THREE 
STATIC LOAD CASE NUMBER 
ELEMENT DISPLAY SET NUMBER 
SCALE FACTOR 
OUTPUT STRESS COMPONENT 



6RAVITY LOADS / MAXIMUM PRINCIPAL STRESS / PSI UNITS 
1) 
I) 

.1000E*04> 
9» 



OROVILLE DAM SPILLWAY/ HALF OF TYPICAL BAY/ STATIC ANALYSIS RESULTS/ 

PIER ELEVATION VIEW/ ELEMENTS 00I-TO-I77/ 

GRAVITY LOADS / MAXIMUM PRINCIPAL STRESS / PSI UNITS 

SOLID ELEMENT CENTROIOAL STRESSES 












1 


2 






























1 


3 


5 


4 


2 


1 


















1 


2 


4 


9 


7 


6 


4 


















1 


2 


5 


11 


11 


10 


9 




5 













1 


2 


5 


13 


12 


12 


11 




8 


1 














1 




13 


14 


15 


14 


16 


20 


12 



















11 


14 


15 


16 


15 


23 


I 


9 


7 




3 


5 


5 




6 


9 


11 


15 


23 


23 






-2 




-6 


-8 


-10 


-4 








5 


12 


21 


50 






1 




6 


-4 


-10 


-14 


-4 


1 





4 


13 


27 


30 


14 


5 




3 


5 


5 




1 













4 


10 


6 







2 






























2 


1 







-I 


-I 


-I 


-1 



























5 


3 


4 




3 


2 


2 


2 




1 


1 


1 


1 





33 


-27 


-24 


-20 


-17 


-14 


-11 


-8 


-7 


-5 


-4 


-3 


-2 


-1 



TABLE II. 3 STATIC GRAVITY STRESSES 

MAXIMUM PRINCIPAL COMPONENT 



319 



20 



■jr/inpir. LINF THPCr = (r,R4VlTYtHrOPOSTaTrC»PRF5TRESS^ MTNIMUM PPINCIPAL STRESS/ PSI UNIS 

STflTTC unnn rtsr fUnpFR = ( u) 

nP->VTLLt: nan '-.r-TLLWAY/ HALF OF TYPICAL BAV^ STATIC ANALYSIS PFSULTS/ 
PT€R EL'^"ATTOK' VIFVI/ FLE1FNTS fOl-Ta-177/ 
'■."AVITYvHYIDTSTATI'-trrrsTRESS/ MIMI1U" ^RTNCI^aL STR'SS/ PSI UMITS 

soLTn ^L'^''ENT rF^'■^P0I^aL stresses 



= ( 


1) 


: ( 


. 1 'Or- to<») 


= ( 


7) 



-1' 


-1' 


-12 


-1? 


-13 


-12 


-12 


-11 


-IG 


-8 










-lu 


-It; 


-17 


-iq 


-2] 


-2' 


-18 


-15 


-12 


-10 










-2? 


-?? 


-2' 


-21. 


-27 


-21. 


-21 


-18 


-15 


-11 










-T' 


-20 


-?7 


-2" 


-3u 


-c7 


-25 


-23 


-IQ 


-1.3 










_T(S 


-■»T 


-31 


-32 


-?(S 


-»7 


-38 


-35 


-■'u 


-15 


-33 








-UT 


-37 


-36 


-37 


-I.? 


-1*7 


-1.8 


-i»6 


-36 


-31. 


-23 








-U6 


-1*2 


-1.1 


-liU 


-57 


-62 


-61. 


-6 1 


-53 


-21 


-1.3 


-18 


-11 




-=56 


-51 


-U7 


-5f^ 


-8Q 


-83 


-9!. 


-8 2 


-72 


-67 






-ifS 




-1? 


-« " 


-72 


-Pf, 


-11" 


-1C3 


-99 


-1)2 


-11.2 


-iro 






-t.9 




129 


-121. 


-I2r 


-123 


-13r 


-115 


-107 


-135 


-265 


-168 


-131 


-27 


-55 




12' 


-IIU 


-lit 


-IL 5 


-in/ 


-135 


-11.7 


-21" 


-?C2 


-2?" 


-150 


-57 


-21. 




11'^ 


-ll-' 


-IIF 


-113 


-95 


-161 


-?d9 


-217 


-237 


-165 


-l-'7 


-75 


-31 




-■^7 


-116 


-127 


-IZIS 


-122 


-12? 


-297 


-21.8 


-12j 


-11.8 


-13^ 


-go 


-71. 




-•f" 


-12? 


-l?o 


-I'll 


-l(»tt 


-139 


-196 


-19 8 


-15-' 


-153 


-139 


-112 


-91. 


-96 


-r^ 


-12-' 


-xu; 


-15-7 


-156 


-146 


-11'-' 


-l-^l 


-190 


-165 


-11.6 


-12P 


-103 


-81. 



TABLE II. 4 STATIC OPERATING CONDITION STRESSES 
MINIMUM PRINCIPAL COMPONENT 



320 



21 



HF/\nTNr, LTME THOFF 
"STATIC LO/vn cftS? NDMnF"^ 
■^L^'I^NT niSPLAV SFT NUM3FR 
SCAL"^ FHCT03 
lOUTPUT ^TC'T^^; r^^'F'^NFHT 



GPflVTTY+HYDrnSTnTTC*PPFSTRFSS/ MAXIMUM PRINCIPAL ST'FSS/ PSI UNITS 
1.) 
1) 

. 1 C ? c ♦ 1 1. ) 
9) 



OPO'iTLl-E HAM <^PILLMAY/ HALF OF TYPICAL "AY/ STATIC ANALYSIS PFSIJLTS/ 

"1'"° ELEVATION vii^W/ FLEMFt'Tt^ jiJt-TO-177/ 

GRAVIT YmYnPOSTaiTC+PPESTPrSS/ maximum POTN^I^AL stress/ PST UNITS 

SOLin ^LE^^'NT CEMTpnirflL S''=!ESSES 



? 


T 


1 


1 


T 


1 


1 


- 


3 













T 


" 


C 


1 


7 


1 


1 


n 
















-2 


_? 


-2 


-2 


-1 


- 





n 


r 













-5 


-5 


-^ 


-5 


-3 


1 


n 


J 


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-7 


-7 


-6 


-fi 


-I4 


1 





p 


c 


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-1 








-9 


-n 


-6 


-f> 


-1. 


1 





- 


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-2 


1 








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-7 


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1 


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19 




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-9 


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-u 


-11 


-10 


3 


" 


•y 


• 7 


uq 






<t 




-^ 


-9 


-5 


-7 


-6 


2 





1* 


-9 


-6 


52 


17 


12 






^ 


1 


2 


1 





u 


-f> 


6 


-12 


2i» 


-5 


2 




-1 


" 


7 


12 


1" 


19 


-7 


7 


-' 


5 


n 


1 


2 




-1 


ti 


?? 


I4: 


75 


?li* 


-?7 


-■5 


3 


-1 





-1 


-1 






IT 


3LI 


■55 


12-< 


Z<^' 


205 


3 


t; 


u 


3 


2 


3 


1 


77 


lU 


2? 


1.6 


155 


221 


!!♦•" 


IS 


-1?- 


-15 


-13 


-9 


-6 


- 3 



TABLE II. 5 



STATIC OPERATING CONDITION STRESSES 
MAXIMUM PRINCIPAL COMPONENT 



321 



22 

a direct summation is a good approximation. Also, the stresses in the 
breast wall for this combined load case are in the range -100 psi to 10 
psi. 

2. Analysis for Mode Shapes and Frequencies 

The ten (10) lowest natural frequencies and mode shapes were determined 
using EASE2. Table II. 6 lists the first five (5) frequencies and their 
associated global (X,Y,Z) mass modal participation factors. By definition, 
the X-direction mass modal participation factor for the i-th mode is: 

ij;^ = (t)T M I 
^x ^1 = -X 

where ^. is the i-th mode eigenvector, M is the system mass matrix and 

I is a vector containing ones at X translational degrees of freedom 

and zeroes elsewhere. If \pl is small (in comparison to ip^ and i|jM , it 

X y z 

means that the i-th mode has practically no X component. From Table II. 6 
it is seen that the ^ are negligible from which we conclude that the 
mode shapes are essentially two-dimentional in the vertical Y,Z plane. 

Figures II. 9, 11.10 and 11,11 plot eigenvectors for modes 1, 2 and 3 
respectively. The undeformed structure is shown dashed in these plots. 
The second mode represents localized vibrations in the bent which supports 
the road bridge. Generally, however, the modes are combinations of shear 
and flexure involving the entire pier. The breast wall responds essentially 
as a rigid mass atop the pier cantilever. 

3. Response Spectrum Analysis 

The "modified Pacoima and Taft" acceleration time history (shown for 
20 seconds in Figure 11.12) is assumed to act at the base of the pier in 
the global Z-direction. The response spectrum for 5% damping for the 
modified Pacoima and Taft is plotted in Figure 11.13. Note that the 
spectral acceleration approaches the peak acceleration (0.6g) at the high 
frequency end. 



322 



23 



MODE 

NUMBER, 

i 


NATURAL 
FREQUENCY , 


MASS MODAL PARTICIPATION FACTORS 


^x 


4 


*l 


1 


10.07 hz 


.0005 


0.7057 


2.4645 


2 


22.60 hz 


-.0028 


0.8632 


0.1599 


3 


24.84 hz 


-.0058 


2.0934 


1.5704 


4 


28.14 hz 


-.0073 


1.6520 


-0.7618 


5 


36.39 hz 


-.0067 


-0.0664 


0.2202 



TABLE II. 6 OROVILLE SPILLWAY STRUCTURE 

NATURAL FREQUENCIES AND ASSOCIATED 
MASS MODAL PARTICIPATION FACTORS 
FOR THE LOWEST FIVE MODES 



323 



24 



vKvTx \ \ \ \ \ \ "^ 

\AXi\--^rr\\\ \ \ ^ ^ 
u\\\ \\\\\\\ ^ 

\ WWi^^vWVxv^ ^ 
mvM \ \\\-\-^Va^ - - . 

Ur4r-V\\ \ \ \ \ V\ \1 








\ \'\1\-V-V^ 


Nr\\vv~" "^^^ 


T- 1 


fc^X-.-^ 


VT\\\\, \ 


1 




r^trtTirTi lzl 


( 


1 
+ 


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+- 


f Jrr ^ 1 1 




1 


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, n 


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F ] 




1 [ 


K 








r\ 








/STqiE OF CRLIFORNIP/DEPPRTMENT OF URTER RES0URCE5/0R0V ILLE DfiM SPILLUflV/ 
/STRTIC RNO OYNRMIC THREE DIMENSIOHPL FINITE ELEMEHT RNRLYStS/ 
/PROJECTION OF THE STRUCTURE ON THE Y-Z PLRNE/ 


z, — 1 



EBC/CSI/ERSE2/E2PL0T 



VIEU NUMBER 1 



FRAME NUMBER 1 



RUN DOTE 10/30/76 



FIGURE 11,9 OROVILLE SPILLWAY STRUCTURE 

PLOT OF MODE SHAPE NUMBER ONE 



324 



25 




/STRTE OF COLIFGRNIfi/OEFflRTMENT OF URTER RESOURCES/OROV I LLE DflM SPILLURY/ 
/STRTIC AND DYNAMIC THREE OIMEHSIONRL FINITE ELEMENT RNRLYSIS/ 
/PROJECTION OF THE STRUCTURE ON THE Y-Z PLANE/ 



EAC/CSI/EASE2/E2PL0T 



VIEU NUMBER 1 



FRAME NUMBER 2 



RUN ORTE 10/30/76 



FIGURE 11,10 OROVILLE SPILLWAY STRUCTURE 

PLOT OF MODE SHAPE NUMBER TWO 



325 



26 




/STRTE OF CfiLIFORNin/DEPflRTMENT OF UfiTER RE50URCES/0R0V I LLE DRM SPILLUfir/ 
/SrfiTlC fIND UYNRMIC THREE DIMENSIONRL FINITE ELEMENT RNRLYSIS/ 
/PROJECTION OF THE STRUCTURE ON THE Y-7 PLRNE/ 



EflC/CSI/ERSE2/E2PL0T 



VIED NUMBER 1 



FRAME NUMBER 3 



RUN DATE 10/30/76 



FIGURE 11.11 



OROVILLE SPILLWAY STRUCTURE 
PLOT OF MODE SHAPE NUMBER THREE 



326 



27 




o 




h 




x 


R 


o 


^ 


H 


H 


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H 


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K 


§ 


W 




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p£ 


H 


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R 


H 




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s 


c; 


O < 


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h 


W 


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CJ 


O 


u 


f^ 


< s 



1 30DN 

/lldOti 99b/ 
NOliHnSNbdi- 



3—78786 



327 



/ST9TE OF CRLIFORNIfl/DEFRRTMENT OF UHTER RESOURCES/OROV ILLE OfiM SPILLURT/ 
/STATIC AND DTNflMIC THREE DIMENSIONAL FINITE ELEMENT ANALYSIS/ 
1- 



I- 0.90 




X-TRANSLRTION 

/ABS ACCEL/ 

NODE 1 



FREQUENCY (CYCLES/UNIT TIME) 



FIGURE 11,13 



RESPONSE SPECTRUM (5% DAMPING) FOR 
MODIFIED PACOIMA AND TAFT 



328 



29 



A response spectrum analysis using a square root of the sum of the 
squares combination of the lowest ten modes was perfomred with EASE2. 
Vertical (Y) stresses in the pier due to Z-direction base input are 
listed in Table II. 7. Mote that the critically stressed region is the 
pier heel; 457 psi is predicted in element 14 at the base. Stresses in 
the breast wall are considerably lower than those predicted in the pier. 
The maximum X-direction stress in the wall occurs at the crest and is 
estimated to be 32 psi . 

The maximum component of deflection calculated by EASE2 is 0.136 
inch (Z) at the crest. A hand estimate of the crest displacement shows 
that the first mode contributes over 95%; this means that the structure 
is responding primarily in its first mode. 

4. Time History Response Analysis 

From the response spectrum analysis we note the following: 

(1) the first mode is the principal contributor to the response 
of the structure; 

(2) stresses are largest in the pier heel region but are also 
significant in the vacinity of the pier toe; 

(3) the pier provides horizontal support to an essentially rigid, 
massive breast wall; stresses in the breast wall are low 
because it responds nearly rigidly. 

The essential characteristics of the time history analysis can be 
inferred from the results of the response spectrum solution. The results 
of both analyses should be practically identical because the principal 
contribution to overall response is contained in the lowest modes, 
particularly the first mode. 

a. solution parameters 

The forcing function is applied as Z-direction ground input acting 
at the base of the pier. The same "modified Pacoima and Taft" accelera- 
tion history used for the response spectrum analysis (Figure 11.12) is 



329 



HEADING LINE THREE 

RESPONSE CASE NUMBER 
ELEMENT DISPLAY SET NUMBER 
SCALE FACTOR 
OUTPUT STRESS COMPONENT 



= (VERTICAL (Y-DIRECTION) STRESS/ PSI UNITS/ 

= ( 1) 
= ( II 
= ( .lOOOE^O'.l 



= ( 



2) 



OROVILLE DAM SPILLWAY/ HALF OF TYPICAL BAY/ RESPONSE SPECTRUM ANALYSIS/ 
PIER ELEVATION VIEW/ ELEMENTS OOl-TO-177/ 
VERTICAL (Y-OIRECTION) STRESS/ PSI UNITS/ 

SOLID ELEMENT CENTROIOAL STRESSES 



29 


24 


21 


18 


14 


10 


6 


4 


7 


17 








27 


25 


23 


25 


26 


17 


8 


10 


25 


47 








36 


32 


30 


29 


32 


18 


7 


22 


47 


76 








*6 


39 


33 


32 


34 


17 


12 


35 


62 


108 








5<* 


44 


35 


32 


36 


16 


19 


47 


84 


84 


50 






62 


48 


37 


32 


37 


14 


26 


61 


85 


111 


188 






72 


57 


42 


33 


38 


15 


33 


65 


107 


179 


25 


6 


14 


96 


80 


56 


38 


40 


16 


41 


82 


141 


210 






23 


164 


155 


107 


59 


38 


21 


58 


108 


155 


205 






24 


293 


277 


199 


108 


42 


26 


69 


115 


156 


161 


82 


6 


29 


327 


278 


206 


125 


55 


27 


68 


112 


147 


153 


113 


63 


30 


SS*. 


289 


212 


134 


63 


27 


66 


108 


137 


143 


128 


114 


84 


385 


303 


216 


138 


68 


27 


63 


102 


128 


137 


135 


143 


200 


421 


310 


215 


139 


71 


28 


59 


96 


121 


133 


138 


161 


192 211 


457 


299 


214 


139 


73 


28 


55 


91 


116 


130 


143 


166 


183 162 



TABLE II. 7 



RESPONSE SPECTRUM ANALYSIS STRESSES 
VERTICAL (Y) COMPONENT 



330 



31 

also used in the time history response analysis. The structure is subjected 
to ground motion for the full 20 second duration. 

Even though the response is principally represented by the first mode, 
contributions for all ten (10) of the structure's lowest modes are included 
in the transient response analysis. Five (5) percent damping is assumed 
uniformly for all modes. The time interval at which output is displayed 
is limited to 1/5 of the period of the first mode or 0.02 second (i.e., 
20% of 1/10.07 hertz). Since the excitation is applied for 20 seconds, 
output is produced at 1000 time points. Using the predictions from the 
response spectrum analysis, the amount of time history output can be 
narrowed considerably; i.e., principal stresses in the pier are of major 
concern, and output is limited to these stress versus time histories. 
Complete spatial distributions of principal stresses in the pier are 
recovered only at those times at which maxima occur. 

b. results 

The histories of absolute Z-direction acceleration for four nodes 
along the crest of the pier were averaged and plotted versus time; the 
resulting acceleration time history is shown in Figure 11.14. Although 
the earthquake acts on the structure for 20 seconds, peak values were 
found to occur in the 7-to-9 second interval; consequently, only the 
first 10 seconds of response are shown. From Figure 11.14 it is seen 
that the peak crest acceleration is about 1.2 g's. 

Figures 11.15, 11.16, 11.17 and 11.18 are principal stress time 
histories covering the first 10 seconds of response. Figures 11.15 and 
11.16 show minimum and maximum principal stress, respectively, developed 
at the centroid of element 14; element 14 is located at the pier base, 
upstream face (i.e., at the pier "heel"). Figures 11.17 and 11.18 are 



331 



"-a 
> z 



l-QZ 
o: D 



i-'^, 


• 


— OUJ 




2 — a 




cezu 




oEa 




b-Z 








-iDin 




CC UJ 




UOK 




20 




u-a 








00 




lu — a: 








CZCEUJ 














"" ' " " 1 " M 1 , ,, 1 1 1 I..), ,. 




I SDb 

NBIiUy31333b 

33Hy3Ab 



332 



33 




fii anos 



333 



34 




334 



M 0I13S 

SS3^i£-Ccl 

CIDi'iN3a 



35 




£1 anss 

ss3yis-id 



335 



36,, 




336 



ST anos 
aioyiN33 



37 



plots of minimum and maximum stress, respectively, predicted at the e.g. 
of element 15 (at the pier "toe"). The peak values of principal stress 
predicted in the pier heel and toe regions are sumnarized in Table U.S. 
Note that the critical values occur at two times: 7.76 and 8.46 seconds. 
Since the earthquake can act in either the +Z or -Z direction the maximum 
seismic stresses (irrespective of sign) are 480 psi in the heel and 305 

psi in the toe. 

Tables 1 1. 9 and 11.10 list the minimum and maximum principal stresses, 
respectively, in all pier elements at time 7.76 seconds. Similarly, 
Tables 11.11 and 11.12 show the min/max stresses, respectively, developed 
in the pier at 8.46 seconds. Note that the peak stresses always occur 
either in the heel or toe. Table 11.13 shows the vertical (Y) stress 
distribution in the pier at time 7.76 seconds. By comparing Tables 11.10 
and 11.13 it is seen that the critical stresses in the pier are nearly 
vertical in the heel at 7.76 seconds (e.g., in element 14, 480 psi 
maximum principal versus 464 psi vertical). Also, from the time history 
analysis the maximum vertical stress in the pier heel (i.e.. element 14) 
was calculated to be 464 psi (see Table 11.13); this prediction agrees 
very well with the 457 psi value predicted by the response spectrum 
analysis (see Table II. 7). 



337 



w 






CJ 






3 






W W « 

Saw 






vo 


00 


w D g 


H 


iH 


fo o s 


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W H t) 

« fe 2 


H 


H 


H 


H 


w 


^ 


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u 


U 


U 


fe s 


0) 


Q) 


°^ 


w 


tn 


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vo 


vo 


S U 


r-- 


^ 


H U 


. 


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Eh O 


r- 


00 


^^ 






< 


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•H 


s 0. 


W 


W 


D H CO 


Qa 


04 


S U CO 






ys§ 


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s cu CO 






w 






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in 


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W t> pq 


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fe CJ S 






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H 


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w 




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u 


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0) 





o 2 


to 


tn 


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2; H S 


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rn 


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1 


1 


S 






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a 
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EH D 

Q S 
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w g 

pL| CO 

m H 
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^^ 
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K 
EH Q 

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CO g 

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CO O 

CO U 

M »< 

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w 


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338 



HF«nTi\r. {,}<\r THr-'F"" 

TT'IF ^Trn tiuMi-fp 

FIFMCf.T nISPL'^Y "^f T MiiMi-r^ 

•^r fi\ r F r.rjri'? 

JJffr (F riiifPiii 



39 



= t/MIiJI'-'Ll" PI'Ii'iClPAL STtJFSS/pt^I Um I TS/T TMFSTFP r^UMRFR 3>M^ / 



• HP) 
1) 



, 1 (innt ♦iii.) 



. / /'-nt ♦on 



fiDOVni F rfl'^ spii I wftY/ HALF OF TYPICAL t'Al/ [ Mfc HiSlOKY AMALYSIS/ 

PTFO FIFW^Tin.i VTM.V t^l.FMFNTS nni-TO-1/7/ 

/MlNir^iiM PRINFTPAI ^T'-'FSS/Psr Urj I TS/T 1 1^^ ST c P ^JUl■"^rP 3PH/ 

SOI in FLFMFrJT CF'TPO F nai. miJ|;<,scs 



^\ 



_^h -?r -71 

_S(i -bif -<^? 

-/b -88 -ni 

-<.in -116 -ji^g 

-117 -n? -n<s -HI 

-) 1^ -l*^? -1,S4 -1H6 

-Idl -i>Jn -PI? 10 -7 

-176 -?18 -?«7 

-U'i* -?/"j -loo 

-1 M -^16 -?'.? -lAO -24 

-lf.4 -rJOO -?17 -l^JO -107 

.p« _m -c.f, _]-),j -\h?. -191 -?ri8 -?01 -lf>7 

-pf, -i<) -Hf, -1?< -ic^.^ -1H3 -?->o -2CS -^06 -?47 

-II? -I"i0 -17S -loi -P(jJ -219 -^r,^ -305 

-11.1 _i^>, -\U -K" -?ill -?2^ -P;.} -i?-< 



-1 


-T 


-) 'I 


-?•) 


-;'5 


-,"•7 


-t; 


-1 1 


-?o 


-TS 


- '• 'J 


-4S 


-7 


-,r. 


-r>H 


-SI 


-SS 


-66 


-7 


-\(- 


-Tl 


-'■I 


-12 


-H'* 


-7 


-\f. 


-If 


-7S 


-m 


-101 


-R 


-17 


--<Q 


-R6 


-100 


-117 


-P 


-IR 


-A? 


-96 


-113 


-136 


-R 


-IP 


-/.S 


-105 


-1?A 


-W.R 


1 


-1 1 


-^o 


-11 1 


-)?a 


-IS? 


-1 


-R 


-uh 


-n.T 


-1^^ 


-lAH 


IS 


-?n 


-■\h 


-7T 


-110 


-1 v 



-1 

-7 



-?7 



-10? 



-1? 



TABLE II. 9 



MINIMUM PRINCIPAL STRESSES IN THE PIER 
AT TIME 7.76 SECONDS 



339 



HFAHINr. LIMF THOFF 



= (/l.'ArlMUM Pt^IUCTPftL STPFSS/P^I UNlTS/TTMFSTfP NUMHfR Iflfi/ 



TTi'F ^TF" riUMMFi' 

FLF'-'F^T nTSPL'lY' Si'T MiiMi; 

Sr^l.T FATTOP 

oiiTP'iT STPF9S ro'<i-^OM'"nT 

TI'IF C.F OUTPUT 



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n 



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OPOVTI.I.F "AM <;PTM ^'AY/ H/>I_F OK TYPICAL fiAY/ rl'-IE HISTORY ANALYSTS/ 

PIFP FLFVATIdrj VIi-W/ FLFMF^'TS 001-10-177/ 

/MAXT'-'ll" PPTMrlPA( STr-fSS/PSI UM ] TS/T I MF^^TF P NUMPFR 3R<^/ 

SOLIn FLFMFNT CFMTPniliAL STPFSSFS 
11 



3't 



30 


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l*. 


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l'^ 


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20 


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PO 


■>u 



3? 
3'? 



31 
33 
37 



6« 


*^0 


'W 


Uh 


78 


1?'^ 


1?7 


QA 


70 


QH 


??h 


2?') 


17^ 


1 l'^ 


P7 


?^7 


?'*3 


1 fiR 


IT^ 


104 


31? 


?h'=- 


204 


144 


no 


3f>'« 


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21*^ 


li^^S 


112 


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21^ 


ICS 


no 


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I'^l 


10"") 



11 


n 


10 


7 


22 


14 


14 


7 


?_Q 


22 


lb 


8 


32 


?3 


14 


5 


34 


2'5 


19 


9 


4 6 


37 


30 


34 


<^6 


46 


4l 


27 


64 


SO 


30 


8 


6fl 


44 


14 


2 


7? 


44 


24 


7 


7rf 


S4 


36 


24 


H2 


^1 


4H 


45 


«3 


64 


bJ 


44 


PO 


62 


S2 


46 


74 


•^s 


44 


38 



2 
1 




-1 
12 


10 


^7 


8 


3 


47 



20 




11 


17 


5 

» 


48 


12 


2 


4S 


26 


6 


36 


12 


-6 


24 


3 


7 


IS 


-3 


-q 



TABLE 11.10 



MAXIMUM PRINCIPAL STRESSES IN THE PIER 
AT TIME 7.76 SECONDS 



3A0 



41 



HF en TNT, LIMF THi-'ri 



(/MNI"!!' Pi'IlirH^AL '^T'^F'^'^/P'^I UMlTS/IIMf STFP NUMMtH 4?3/ 



oiiTPiir <;T-^rss ro'M'Pi.Fvr 

TTMf OF OUTPUT 



I) 



.U'Oiit 'C^l 



.Haf,nt + 01 ) 



O'^OVTLLF ■>'»''' tPTMKIY/ t.iMF OF TYPICAL P^Y/ T I MF HISTORY ftN/VLVSIS/ 

PIFU FIFVATIOri WTi^'V K.Pi'FMTS Oni-fn-177/ 

/MI'ITWIIU Pt?(NrliMI ST^TSS/MSI UM I T S/T I ^FSTF (' NUMfiFK A??/ 

SOLIO FLFMFMT Crt'TPOIr^AL ^.T^-'FSSfS 



-31 


-?0 


-IS 


-13 


-1? 


-IP 


-1? 


-10 


-7 


_r> 








-?r 


-I*^ 


-If. 


-pfN 


-P'i 


-?3 


-!'-> 


-1-4 


-7 


-2 








-?? 


-?1 


-?I 


-?5 


-34 


-?8 


-?1 


-14 


-7 


-1 








-2Q 


-Pf- 


-PU 


-?« 


-3V 


-Tfl 


-?1 


-12 


-4 


1 








-3^ 


-3" 


-?ft 


-pQ 


-a4 


-3S 


-?s 


-l-^ 


-7 


-9 


-p 






-4? 


-3 3 


-?7 


-30 


-i,Q 


-39 


-30 


-24 


-2H 


-23 


-7 






-S1 


-<4 


-31 


-33 


-S5 


-4f. 


-37 


-33 


-20 


T 


-32 


-13 


-IS 


-ftq 


-r,o 


-4 3 


--{Q 


-t^h 


-•^1 


-39 


-2? 


-2 


-fl 









\?.} 


-1)Q 


-P^ 


-^;q 


-70 


-S^ 


-32 


-n 





1 






-6 


^l'^^ 


-?10 


-l'^3 


-ijf, 


-7 4 


-■^J 


-3(1 


-14 


-1 


-1 


-4 


-U 


-4 


?A7 


-?17 


-1»^3 


-11-1 


-79 


--^7 


-37 


-23 


-IP 


-26 


-3S 


-R 


-2 


?7H 


-^M 


-1 7/. 


-1?1 


-M T 


-S9 


-42 


-33 


-32 


-1ft 


-33 


-?n 


-s 


31? 


-P^P 


-Ifl? 


-IP', 


-■*'^ 


-S9 


-au 


-3 7 


- <4 


-T3 


-26 


-9 


s 


3^1 


-?^\ 


-1«I 


-IP'S 


- J3 


-Sh 


-u "> 


-Jli 


-n 


--•7 


-17 


- 4 


-s 


loa 


-?'^1 


-17.J 


-1P3 


-7'y 


-SI 


-3'> 


-2>' 


-24 


-pri 


-9 


3 


7 



TABLE 11.11 



MINIMUM PRINCIPAL STRESSES IN THE PIER 
AT TIME 8.46 SECONDS 



341 



TJMF CTFP ri'lMPF" 

srii"^ Ff.rKip 

TTMF rr niiTPUT 

riPovTLLf I'AM '■.PTM '/AY/ MAi.F "F Ty^ir/vL nav/ iimf history ftMAUrsis/ 

pjfrp HFv/\TTO'l Wf-w/ ritfiFt'TS 0" I -Tl)-1 77/ 

/MrXT^MM ppT'Nirl'^/M ^T'HS^/P';! HMITS/TIMFST!- P MDMnFM '*?^/ 

SOI in FLFMFNT CF^'TP'^ln''L ST^Ji^SSFS 



(/MAXIMUM 


in^irjriP 


( a rM ) 




( II 




( .1 


^not*r»^) 


( '1 






-Mlf ♦01 ) 



irjrlP/M. ST^FSS/PSI ljr<jns/TIMFSTFP NUMHF« ^?3/ 



n 
1 1 



9 


?n 


?4 


?b 


P.l 


?h 


Pf> 










19 


13 


37 


'.^ 


47 


bX 


(^n 










?5 


47 


S3 


M 


fi9 


R^ 


9S 










PB 


'-.f^ 


'^'7 


7V) 


P9 


ICfS 


14R 










TO 


t~ S 


7ii 


H9 


10 4 


1?4 


1P4 


75 








T^ 


7 3 


R'; 


101 


1^2 


140 


14B 


170 








ri 


79 


nit 


n^ 


139 


lf<6 


193 


-S 


f. 


?. 




3'i 


"? 


ion 


1^3 


14H 


18H 


a'-.i 






9 




11 


P? 


ino 


1?J 


153 


192 


a's'^ 






Q 




?■( 


7 


93 


117 


147 


IBO 


?n? 


ISO 


IR 


?4 




!<=> 


Ut~: 


HP 


in? 


1 If-. 


lh>4 


17fl 


ISS 


fl6 


?n 




1 1 


->.'^ 


hi 


9 9 


W'* 


ISS 


KQ 


lh2 


133 


79 




'. 


?(- 


S9 


9<' 


Ic--^ 


147 


l*^! 


1^4 


l'-3 


19? 




? 


?1 


■^3 


nr. 


1 1 1 


140 


l'-4 


lf>l 


173 


1"7 


23f> 


_p 


1h 


'4 9 


M ■( 


I 1 I 


1 10 


K-n 


1^0 


17S 


1R9 


177 



TABLE 11.12 



MAXIMUM PRINCIPAL STRESSES IN THE PIER 
AT TIME 8.46 SECONDS 



3A2 



43 



HEADING LlNf THR^E = (/VERTICAL STRESS IN P lER/V-OIRECT lON/PSI UNITS/TIME STEP NUMBER 388/ 



TIME STEP NIIMHLP = ' 388) 



ELEMENT DISPLAY SET NUMBER = I 

SCALE EACTOP = ' .IOOOE'04. 

OUTPUT STRESS COMPONENT = ( 2) 

TIME OF OUTPUT = < .7760E»on 

OROVILLE DAM SPILL»*AY/ HALF OF TYPICAL BAY/ TIME HISTORY ANALYSIS/ 

PIFR FIFVATION VIEW/ ELEMENTS OOl-TO-177/ 

/iERTICAL STRESS IN P lER/Y-D IRECT lON/PSI UNITS/TIME STEP NUMBER 388/ 

SOLID ELEMENT CLNTROIDAL STRESSES 

17 K 1? 10 7 'i 2 -1 -5 -13 

15 u )3 15 1ft 10 2 -8 -20 -36 

20 18 18 \'i 21 11 -2 -'8 -38 -«>1 

2b 23 21 22 25 11 "7 -28 -52 -91 

3? 27 23 23 28 9 -U -39 -71 -71 -'.2 

-20 -52 -T* -96 -163 

-27 -57 -95 -159 1* "6 9 

30 35 -34 -73 -127 -191 -3 

42 28 -10 -51 -99 -U'. -197 -<• 

77 20 -17 -64 -107 -150 -154 -80 -4 -24 

□7 32 -19 -63 -106 -141 -148 -113 -68 -36 



39 31 25 24 
48 39 3U 26 



67 58 42 

120 120 H4 

221 223 159 

266 235 171 



31, 255 181 108 41 -16 -63 -103 -132 -140 -131 -125 -98 
362 277 189 113 46 -13 -6l -98 -124 -136 -140 -156 -230 

,15 ^9 -,0 -57 -94 -119 -133 -145 -179 -221 -244 



416 290 190 



464 



280 191 117 51 -7 -54 -90 -115 -133 -152 -185 -209 -188 



TABLE 11.13 VERTICAL (Y) STRESSES IN THE PIER 
AT TIME 7.76 SECONDS 



343 



III. TWO DIMENSIONAL ANALYSIS WITH HYDRODYNAMIC INTERACTION 

A. three dimensional dynamic analysis which includes hydrodynamic 
interaction is beyond present research development for a structure of 
this type. It is possible to estimate the magnitude of this effect 
from the analysis of a two dimensional model with similar dynamic 
properties. 

A. Analysis Without Interaction 

A two dimensional finite element model was selected with the same 
mesh idealization as used in three dimensional model of the pier (Figure 
II. 5). Normal concrete properties were used except the weight density 
was scaled so that the fundamental period of the two dimensional model 
was the same as the fundamental period of the three dimensional model 
without hydrodynamic interaction. From Table III.l one notes that the 
maximum stresses obtained from this two dimensional model are in 
excellent agreement with the three dimensional analysis. This confirms 
the results from the three dimensional analysis which indicated \iery low 
stresses in the breast wall. 

B. Analysis With Interaction 

The analysis of the two dimensional model with hydrodynamic inter- 
actions was accomplished using the following program: 

*EADHI - "A Computer Program for Earthquake Analysis of Gravity 
Dams Including Hydrodynamic Interaction," by P. Chakradarti and 
A. K. Chopra, May 1968, College of Engineering, U.C. Berkeley, 
California. 

From Table III.l the results indicate an increase in critical 
stresses at the toe and heel of approximately 20 percent due to hydro- 
dynamic interaction. 



344 



45 



2/D ANALYSIS 

W/ HYDRODYNAMIC 

INTERACTION 


B 

■H 
+J 

PI 
-H 

e 


vo 

00 

o 

LD 


r-- 

00 


0) 

e 

■H 
4-> 
\ 
X 

e 


1J3 

\ 

in 


00 

\ 

vo 
ro 


2/D ANALYSIS 
W/0 HYDRODYNAMIC 
INTERACTION 


e 

•H 
-P 

-H 


00 

H 


\ 

o 

n 


e 

•H 
+J 
\ 


\ 
en 


vo 
'I' 

00 

•\ 

VD 


3/D ANALYSIS 
W/0 HYDRODYNAMIC 
INTERACTION 


g 
•H 
+J 

•H 


00 

ro. 


\ 

in 
o 
n 


e 

•H 
-P 

e 


r- 
o 

00 


1^ 
»* 

00 
VD 

n 


o 

H 

u 

o 


EH 


in 

rH 
EH 

W 

S 

M W 

O J 

Eh W 



345 



46 
C. Extension to Three Dimensional Analysis 

A reasonable engineering approximate solution to the three dimensional 
hydrodynamic interaction is to apply the + 20 percent correction to the 
results obtained from the three-dimensional analysis without interaction. 
This is a suggested approach for a horizontal earthquake. The effects of 
hydrodynamic interaction due to a vertical earthquake have not been 
considered. 



346 



47 



IV. TRANSVERSE EARTHQUAKE ANALYSIS 

The monoliths of the flood control outlet structure are separated 
by contraction joints. The actual size of these joints must be measured 
in the field. The monoliths which do not contain gates are solid and 
are short relation to the two monoliths with gates. Therefore, it is 
reasonable to evaluate the stresses in the gate monoliths only due to 
a transverse earthquake. Because of the special geometry of these 
structures the maximum possible transverse displacement would be equal 
to the size of two contraction joints regardless of the magnitude of 
the earthquake. 

Furthermore, transverse displacement of the gate monoliths will 
cause bending in the piers between the foundation and the roof of the 
intake which is a distance of approximately 34 ft. A good estimate of 
bending developed in the pier can be calculated by assuming the pier 
to act as a fixed end beam subjected to a support displacement equal 
to the size of two construction joints. For a 5 ft. pier, the extreme 
fiber strain will be 0.0012A where A is the contraction joint displace- 
ment. Since these piers are reinforced a damage evaluation can be made 
from normal reinforced concrete theory. 



347 



V. VERTICAL EARTHQUAKE 

For a structure of this type the modes of vibration which have 
significant vertical components have a very high frequency—greater 
than 50 cycles per second. For this high frequency the structure moves 
as a rigid body at the earthquake acceleration. Therefore, a very good 
approximation of the maximum stresses due to vertical earthquakes can 
be calculated by multipling the gravity stresses by the maximum accelera- 
tion (given as a fraction of gravity). 

The vertical stresses due to gravity are given in Table II. 1. If 
the maximum acceleration due to a vertical earthquake was 30% of gravity 
the vertical stress at the heel will vary +90 psi or from -210 psi to 
-390 psi. 



348 



49 



VI. FINAL REMARKS 

This report summarizes the results of a three dimensional finite 
element analysis of the Oroville Dam Flood Control Outlet Structure. 
Hydrostatic, prestress, gravity and horizontal earthquake results are 
presented separately. In addition, approximate methods of analysis are 
given for hydrodynamic interaction and for transverse and vertical 
earthquake behavior. 



349 



CHAPTER VII 
SEISMIC ANALYSIS OF THE 
THERMALITO DIVERSION DAM 



Commentary 

As a result of the August 1, 1975 Oroville earthquake, of 
magnitude 5.7, the Department found it appropriate to reanalyze the 
Thermalito Diversion Dam (Figures 187 and 188) using a stronger earth- 
quake (see Chapter V) , and the latest techniques in seismic investigation. 

An earthquake study, monitored by the Department, was conducted 
under a consulting agreement with Dr. Anil K. Chopra. Dr. Chopra's 
results and conclusions were presented in his report, "Earthquake Response 
Analysis of Thermalito Diversion Dam". That report is presented in 
this chapter (beginning on page 355 ) . 

Dr. Chopra performed a dynamic two-dimensional response analysis 
of the dam using the finite-element method. The Department agrees with 
the methods used and conclusions presented in his report. 

The Department investigated the Dam for sliding, utilizing the 

shear-friction equation 

CA + N tan 
Q = . A cohesion value, C = 3447 

kPa (500 psi) , produced a shear-friction factor of approximately 5. This 
is considered a safe value against sliding. 



351 




Figure I87. Plan and Elevati 



352 




Figure 188. Typical Sections 



353 



EARTHQUAKE RESPONSE ANALYSIS OF THERMALITO DIVERSION DAM 

by 
Anil K. Chopra 



355 



EARTHQUAKE RESPONSE ANALYSIS OF THERMALITO DIVERSION DAM 

by 
Anil K, Chopra 

Introduction 

The Department of Water Resources, State of California, entered 
into an agreement with Dr. A. K, Chopra to "perform structural analy- 
sis using finite element techniques of Thermalito Diversion Dam". 
The agreement stipulated that "the analysis will employ the most 
suitable dynamic methods applicable to the specific structures". 

Results of this analysis for the groujid motion specified by 
the Department of Water Resources, State of California, are presented 
in this report. 

The scope of the work necessary for structural analysis of 
Thermalito Diversion Dam was first discussed and outlined in a 
preliminary meeting between Dr. Chopra and Denzil Carr of the 
State Department of \\later Resources. Subsequently, results from 
preliminary analyses were discussed at a meeting on September 7^ 
1976 between Dr. Chopra and Messrs. Sam Linn, Edgar Najera, and 
Vernon Persson of the State Department of Water Resources. The 
series of analyses and results required by the State Department of 
Water Resources were defined. In accordance with these requirements, 
a draft report was submitted on November 4, 1976. 

Messrs. Linn, Najera and Don Steinwert, State Department of Wate 



356 



Resources, coiiunented on the draft report at a meeting on February 25, 1977 with 
Dr. Chopra. The draft report has been revised and expanded to 

account for these comments, resulting in this final report. 

The draft report of November 4, 1976 was based on analyses of the dam for 
the ground motion originally recommended by the Special Consulting Board for the 
August 1, 1975 Oroville Earthquake (Fig. 1). The Board made supplementary 
recommendations regarding ground motions to be considered in their report of 
November 23, 1976. After a study of this supplementary recommendation, it was 
concluded in a report of March 4, 1977 (see Appendix) that, for analysis of 
Thermalito Diversion Dam and Oroville Dam Flood Control Outlet Structure, there is 
no need to supplement the ground motion originally recommended by the Board, for 
which the structures had already been analyzed. All results presented in the main 
body of this report are therefore based on analyses for the original ground motion. 

Finite Element Models 

The dynamic response analysis of gravity dams is done by the finite element 
method, assuming that monoliths act independently of each other and in a condition 
of plane stress; in practice a one-foot thick slice of the dam is analyzed. If 
the foundation material is significantly softer than the dam concrete, then it 
may have a significant effect on the dynamic behavior of the dam and must be 
included in the finite element idealization. In the case of Thermalito Diversion 
Dam, however, the properties of the foundation rock are such that it should 
have only insignificant influence on the dynamic response of the dam. Furthermore the 
degree to which the response is affected depends strongly on the depth of rock 
included in the finite element model, and where rock essentially similar to the 
rock under the dam and near the ground surface may be asumed to extend to great 
depths -- as in the present case — there is no rational basis for defining the 
limits of the finite element model. Consequently, in this investigation the 



357 



I 



concrete monoliths were assiamed to be supported on a rigid base, and the specified 
earthquake ground motion was applied at the base. It was concluded that no greater -; 
reliability in the dynamic response results could have been achieved by including ' 
an arbitrary layer of rock under the monoliths in the finite element model, even 
though it is common practice to include such a layer in performing purely static 
analysis. 

The finite element models defined for the analysis of monoliths 10, 12 and 
18 are shown in Figs. 2-5. They all employ the isoparametric quadrilateral 
element used in the SAP program developed by Professor E. L. Wilson at the 
University of California, Berkeley. They have graded meshes, with slender elements 
near the monolith faces to better define the stresses in those regions. In each 
case, the number of elements through the upstream-downstream direction and the 
number of rows of elements through the height, are considered to be sufficient 
to provide good definition of the stress, especially in the critical zones. 

Two different finite element models were used to represent monolith 12: 
The one shown in Fig. 2 which includes only the monolith itself was defined for 
purposes of preliminary analysis. The other shown in Fig. 3 includes an 
approximate two-dimensional model of the appurtenances: pier, bridge and radial 
gate. This model, although not appropriate for determining the details of 
dynamic response of the appurtenances themselves, is believed to be adequate 
to represent effects of the appurtenances on the dynamic response of the 
monolith, which is the main concern of this investigation. P 

The modulus of elasticity of the finite elements included in the model of 
Fig. 3 to represent the piers was set at 5/45 (=width of pier/width of monolith) 
of the value for concrete. The density of these elements was set in a slightly 
different ratio to include the weight of the radial gate and other equipment. 
For the top-most row of elements, the density was increased to include the 
weight of the bridge, but the modulus of elasticity was taken as that of the concrete 

358 



The static and dynamic analysis of all finite element models were performed 
by the computer program EADHI developed at the University of California, Berkeley. 
This program includes the effects of interaction between the dam and water, and 
of water compressibility. 

The properties of concrete were taken as those provided by the Department of 
Water Resources, State of California, for the 2-1/2 sack mix, which is used 
throughout the monoliths except for a small thickness near the exposed surfaces 
and galleries: 

6 

• Modulus of Elasticity = 5.1 x 10 psi 

• Poisson's Ratio = 0.17 

• Unit Weight = 155 pcf 

Earthquake Response Analysis 

Before proceeding with analysis of the dam, the computer program EADHI was 
extended to handle water level above the crest of a monolith. This capability 
is necessary to analyze overflow monoliths 10 and 12. 

Several preliminary results were generated to obtain an overall impression 
of the dynamic response of the dam. For this purpose the finite element model of 
Fig. 2 for monolith 12, excluding the appurtenances, was analyzed. 

The frequencies of the first three natural modes of vibration are shown 
in Table 1. It is apparent that the dam has rather high vibration frequencies. 

Stress analyses were performed considering the static loads assumed to be 
acting prior to occurrence of an earthquake. These include the dead weight of 
the monolith, and the hydrostatic pressure of the water when the reservoir is at 
the normal level (El. 225). 

The dynamic response of the finite element model of Fig. 2 to the specified 

ground motion (Fig. 1) assxamed to act in the upstream-downstream direction was 

determined. Only those modes with frequencies less than 30 Hz were considered 

359 



Table 1: Natural Frequencies of Vibration of 
Monolith 12 without Appurtenances 



Mode No. 


Natural Frequencies , cps 


Dam Only 


Dam with Water 
at EL. 225.0 


1 
2 
3 


14.5 
30.4 
34.8 


8.5 
29.1* 
28.9* 



The natural vibration modes of the dam without water are numbered 
according to standard convention: The natural mode having the lowest 
vibration frequency is called the first mode, that having the next 
higher frequency is the second mode, etc. Because hydrodynamic 
interaction effects depend on the frequency and shape of the vibration 
mode2, the vibration frequencies of the three modes of the dam are not 
in increasing order when effects of water are considered. However, 
this is of no consequence in the analysis, because all the modes which 
have significant contributions to the total response are included. 



360 



in the dynamic response analysis, because the earthquake motions are not defined 
accurately for higher frequencies. Analysis by the computer program EADHI leads 
to the time history of horizontal and vertical displacements at all nodal points 
of the finite element system and the time history of the three components of total 
stress — static plus dynamic — in all finite elements. Only a small portion of 
these results which is most pertinent to evaluating safety of the dam is included 
here. 

Fig. 6 shows the contours of "envelope" values of the maximum principal 
stress. These are peak values of maximimi principal stress — the most tensile 
stress — developed in each element at any time during the earthquake; they are 
not all concurrent values. The static stresses have been combined with the 
dynamic stresses (taking proper account of the tensorial nature of the stress 
components) so that these contours indicate the absolute magnitude of the tensile 
stresses that must be resisted by the monolith during the earthquake. Of major 
significance are the zones of tensile stress at the downstream face just above 
the bucket and at the upstream edge of the base. The latter zone is in part a 
consequence of the discontinuity between the concrete and the assumed rigid base. 
It is in part the result of a singularity in the mathematical model under study 
and that part could be removed by changing the mathematical model. However, this 
singularity is of little concern in this case because the maximum tensile stress 
is about 310 psi, which, as will be discussed later, is considerably below the 
tensile strength of concrete. 

The final results of critical stresses in Monoliths 10, 12 and 18 which are 
obtained from dynamic analysis of the response of finite element models of 
Figures 3-5 produced by the specified ground motion (Fig. 1) are presented next. 

Fig. 7 shows the contours of "envelope" values of the maximum principal 
stresses in Monolith 12. These are total stresses including those due to static 
loads. As mentioned earlier, the two-dimensional finite element model for the 

361 



appurtenances above the monolith is suitable for including their effects on 
stresses in the monolith, but is too crude for determining the response of the 
appurtenances themselves. Consequently stresses in the appurtenances are not 
presented. Of major significance in Fig. 7 are the zones of tensile stress on ■ 
the downstream face just above the bucket and the upstream edge of the base. It 
is of interest to note that these envelope values of stresses do not differ 
significantly from those computed without including the effects of appurtenances 
(Fig. 6) . 

In order to further examine the response, contours of the instantaneous 
maximum tensile stresses are presented in Figs. 8 and 9 at two instants of time: 
when the tensile stresses on the downstream face attain their peak value (7.35 
sees after beginning of the earthquake motion) and when the tensile stresses at 
the upstream edge of the base reach their peak value (t = 8.47 sees). It is 
apparent that at each of these time instants, stresses in a significant portion 
of the dam are compressive. 

Fig. 10 shows the contours of "envelope" values of the minimxam principal 
stresses — the most compressive stresses — developed in each element at any time 
during the earthquake, including both static and earthquake effects. 

The corresponding contours of "envelope" values of the maximum and minimum ' 
principal stresses for Monolith 10 are shown in Figs. 11 and 12 and those for 
Monolith 18 in Figs. 13 and 14. These are similar in general form to the results 
obtained for Monolith 12 but are significantly smaller in magnitude. These 
shorter monoliths, which have very high natural vibration frequencies, do not 
respond dynamically to any great extent. 

Summary of Results 

The principal features of these dynamic analysis results may be summarized 
as follows: I 



362 



1. The maximum compressive stresses (Figs. 10, 12 and 14) due to both static 
and dynamic effects were about 200, 425 and 350 psi respectively in Monoliths 
10, 12 and 18. These are well within the capacity of this concrete and 
constitute no cause for further consideration. 

2. The maximum tensile stresses due to both static and dynamic effects were about 
200, 310 and 225 psi respectively in Monoliths 10, 12 and 18. Later, these will 
be compared with the tensile strength of concrete. 

Tensile Strength of Concrete 

Although standard criteria for design of concrete dams do not allow tensile 
stresses of these magnitudes , evidence is available to support the conclusion that 

significant dynamic stresses in tension can be supported by sound concrete. 

3,4 
Experiments conducted in Japan showed that static tensile strength of concrete 

is about 8 to 9 percent of the static compressive strength, which is similar to 

the usual assumption of a 10 percent ratio. Moreover, under dynamic conditions, 

at loading rates to be expected in concrete gravity dams subjected to intense 

earthquake motions, these experiments showed that the concrete strengths are 

significantly -- up to 50 percent -- higher (both in tension and compression) than 

under static loading. Similarly, recent tests in the United States on mass concrete 

cores from three dams showed, on the average, a corresponding increase in tensile 

strength of 67 percent . 

Further evidence of the dynamic tensile strength of concrete was provided 

by the performance of Pacoima Dam during the San Fernando Earthquake of 1971. 

The ground motion experienced by this structure must have been very intense; 

accelerations exceeding Ig were recorded near the dam. Analyses carried out at 

University of California, Berkeley, of dynamic response to that motion indicated 

that the dam must have developed maximum tensile stresses in the order of 750 psi. 

Yet no evidence of cracking could be found on either face of the dam. 

363 



On the basis of both the laboratory test data and the experience at Pacoima 
Dam, it is reasonable to assume that the concrete in Thermalito Diversion Dam can 
resist tensile stresses of at least 10 percent of the static compressive strength, 
increased by 50 percent to account for the faster loading rates during vibration 
of the dam. 

Comparison of Analytical Results and Tensile Strength 

As mentioned eariler, the maximum compressive stresses predicted by 
analyses are well within the capacity of concrete and are therefore of no 
concern. 

The corresponding maxim\im tensile stresses in various parts of Monoliths 
10, 12 and 18 are summarized in Table 2. Also included are the static compressive 
strength values, provided by the State Department of Water Resources, and 
estimates for dynamic tensile strength, based on the preceeding section of this 
report, for the four concrete mixes employed in the dam (Fig. 15) . 

It is apparent that the tensile stresses predicted by analyses are less than 
one-half of the tensile strength of the concrete. The concrete is therefore 
capable of safely resisting these tensile stresses. 

The earthquake ground motion specified by the State Department of Water 
Resources, for which only a single horizontal component of motion was provided, 

is the excitation for which all the analyses presented above were carried out. 

2 
However, research studies have shown that the contributions of the vertical 

component of ground motion to the response of concrete gravity dams are 

significant. Even for vertical ground motion, hydrodynamic pressiires act in 

nearly the horizontal direction on a nearly vertical upstream face, thus causing 

lateral response. Although this additional lateral response can be quite 

significant for short dams, the available margin in tensile strength (see Table 2) 

should be sufficiently large to keep the total (due to horizontal 

364 



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365 



and vertical ground motion) tensile stresses within the available tensile strength. 

Conclusion 

Based on results of dynamic analyses and available data for concrete strength 
it is concluded that Thermalito Diversion Dam should be able to resist the stresses 
expected during the earthquake ground motion specified by the State Department of 
Water Resources. 

References 



1. Chakrabarti, P., and Chopra, A.K., "A Computer Program for Earthquake Analysis 
of Gravity Dams Including Hydrodynamic Interaction," Report No. EERC 73-7, 
Earthquake Engineering Research Center, University of California, Berkeley, 
May 1973. 

2. Chakrabarti, P., and Chopra, A.K. , "Earthquake Response of Gravity Dams Includii 
Reservoir Interaction Effects," Report No. EERC 72-6, Earthquake Engineering 
Research Center, University of California, Berkeley, December 1972. 

3. Hatano, T. , and Tsutsumi, H., "Dynamical Compressive Deformation and Failure of 
Concrete Under Earthquake Load," Report No. C-5904, Central Research Institute 
of Electric Power Industry, Tokyo, September 30, 1959. 

4. Hatano, T. , "Dynamical Behavior of Concrete Under Impulsive Tensile Load," 
Report No. C-6002, Central Research Institute of Electric Power Industry, 
Tokyo, November 5, 1960. 

5. Raphael, J.M., "The Nature of Mass Concrete in Dams," Douglas McHenry 
Symposium Volume, American Concrete Institute, Detroit, Michigan, 1977. 



366 




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TOTAL NUMBER OF 

NODAL POINTS =164 
FINITE ELEMENTS = 145 




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369 



EL. 233.0 



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TOTAL NUMBER OF 

NODAL POINTS = 114 
FINITE ELEMENTS = 



EL. 205.0 



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100 



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MAXIMUM TENSION 




Fig. 11 Envelope Values of Maximum Principal Stress (Static + Dynamic) ; 
Monolith 10 



377 



MAXIMUM COMPRESSION 




Fig. 12 Envelope Values of Minimum Principal Stress (Static + Dynamic) ; 
^^onolith 10 



378 



1 



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MAXIMUM TENSION 




Fig. 13 Envelope Values of Maximum Principal Stress (Static + Dynamic) 
Monolith 18 



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MAXIMUM COMPRESSION 




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Monolith 18 



380 



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27 



APPENDIX 



383 



28 



RE: Report (dated November 23, 1976) of the Special Consulting Board for the 

August 1, 1975 Oroville Earthquake to Mr. R. B. Robie, Director, Department 
of Water Resources, State of California. 



Board Recommendation : 

"The Board recommends that for critical structures with high fundamental 
frequencies, the previously recommended time-history of earthquake be supplemented 
by a time history meeting the high frequency (10 Hz or greater) requirements 
specified by the Nuclear Regulatory Commission in its Regulatory Guide No. 1.60, 
with the spectrum scaled to 0.4g at zero period." 

Response (Prepared by Anil K. Chopra) 

Two pseudo-acceleration response spectra, both for damping ratio of 5 percent, 
are presented in the attached figure: one for the earthquake motion previously 
recommended by the Board; and the other is the spectrum specified in the AEC (now 
NRC) Regulatory Guide No. 1.60, scaled to 0.4g at zero period. The natural periods 
of vibration of Thermalito Diversion Dam and Oroville Dam Flood Control Outlet 
Structure lie within the period range to 0.15 sec shown in the attached figure. 

It is apparent that in the range of vibration periods of interest there is 
little need to supplement the ground motion previously recommended by the Board, 
for which the structures have already been analyzed. 

If each spectrum was normalized with respect to its ordinate at zero period, 
in the range of periods of interest, ordinates of the normalized AEC Regulatory 
Guide No. 1.60 spectrum would be significantly larger than ordinates of the 
normalized spectrum for the ground motion previously recommended by the Board. 
However, the actual (without normalizing) spectra do not have the same 
relationship because the ordinates at zero period are different by a factor of 
50 percent :0.5g for tlie ground motion previously recommended by the Board, and 
0.4g for the AEC Regulatory Guide No. 1.60 spectrum. 



March 4, 1977 

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386 



30 



April 13, 1977 



Professor George W. Itousner 
Ittvlslon of Civil Ensineering 
and Applied Mathematics 
California Institute of Technology 
1201 East California Boiilevard 
Pasadena « CA 91109 

Dear Professor Houaneri 

In the November 23, 1976 report of the Special 

Consulting Board for the August 1, 1975 Oroville 

Earthqviaice, the Board recommended that the adopted 

time history of earthquake motion be supplemented 

by a time history of higher frequencies. 

Dr. Anil Chopra of UCB investigated the effect 

of this recommendation. His findings are enclosed. 

Please give us your comments at your earliest 
eonvenlence . 

Sincerely, 



ISonald C. Stelnwei^, Chief 
Structural Unit 
Sesign Branch 
Civlsion of Design and 
Construe ti on 

Eno. 

ENaJerarmrs 

bfb: H, H. Eastin w/attach 
R, B, Jansen w/attach 
K. 0, Barrett w/attach 
E, C. James w/attach 



387 



GEORGE W. HOUSNER 
I20t EAST CALIFORNIA BLVD. 
FASAOENA. CALIFORNIA 9II2S 



April 26, 1977 



Mr. Donald C. Steinwert 

Division of Design and Construction 

Department of Water Resources 

P. O. Box 338 

Sacramento, California 95802 

Dear Mr. Steinwert: 

This is in reply to your letter of April 13th, concerning 
Dr. Chopra's investigation on the OroviUe facilities. I am 
satisfied from Dr. Chopra's statement and the spectrum 
curves that he shows that further analysis need not be made 
with the ground motion specified by the Nuclear Regulatory 
Commission in its Regulatory Guide No. 1.60, with a spectrxim 
scaled to 0.4 g at 0. The ground motion originally recommended 
by the consiolting board is adequate. 



GWH:bb 




lEORGEW. HOUSNER 



j;^^-^;^--^ 



388 



CHAPTER VIII 
REAPPRAISAL OF SECONDARY STRUCTURES 



I Introduction 

As a result of the August 1, 1975 Oro- 
ville earthquake, the Department of 
Water Resources found it appropriate to 
reanalyze the major structures of the 
Oroville complex using both the "latest 
state-of-the-art" dynamic analysis and 
the Reanalysis Earthquake described in 
Chapter V. It was determined that non- 
critical structures could be reassessed 
using a lesser seismic force if a reanal- 
ysis was necessary. The Board would 
defer its recommendation as to the need 
for reappraisal of these secondary struc- 
tures until the evaluation of the 
critical structures is complete. 

After evaluating how a structural failure 
would affect project operation, possible 
loss of life and property, and the possi- 
bility of failure, the Department recom- 
mended what further seismic analysis 
is needed for the facilities in this 
report. The locations of these facil- 
ities are shown in Figures 189 and 193. 

Fish Barrier Dam 

Description 

The Fish Barrier Dam is a concrete 
gravity structure (Figures 190 through 
192) founded on generally fresh and hard 
rock consisting of meta-andesite and 
meta-conglomerate rocks. The dam con- 
sists of a central low-overpour section 
250 feet long, a high-overpour section 
on either side of the lower section with 
a total length of 54 metres (176 feet) , 
and a non-overpour section on the right 
abutment 53 metres (174 feet) long. The 
low-overpour section consists of a 
gravity section with a cantilevered 
reinforced-concrete crest apron extending 
2.7 metres (9 feet) downstream of the 
dam face, and two aeration piers. The 
other two sections are gravity concrete 
sections. The maximum structural height 



of the dam is 28 metres (91 feet) . A 
thorough inspection of the dam after the 
August 1, 1975 earthquake revealed no 
damage to this facility. 

Original Seismic Analysis 

The original seismic analysis consisted 
of a pseudostatic analysis; an accelera- 
tion coefficient of O.lg acting either 
downstream or upstream was used. The 
hydrodynamic force was determined using 
the Westergaard formula and an assumed 
natural period of 1 second for the 
structure. In addition to the earth- 
quake force, the pseudostatic analysis 
included all normal forces including 
river flows up to 50,000 cfs. Utilizing 
these loading conditions, the structure 
was analyzed at several levels for maxi- 
mum and minimum principal stresses, 
safety against sliding (f ) , and the 
shear friction factor of safety 



(s 



sf 



f V + r A 
H 



•) . In addition. 



the overturning safety factor was 
checked for the Federal Power Commission 
review in 1972. 

On the basis of the preceding analysis, 
the maximum principal compressive stress 
was approximately 120 psi, and maximum 
tensile stress was about 10 psi. Both 
are considerably lower than the allow- 
able stress established for this structure. 

The analysis for sliding on the founda- 
tion indicated that the allowable sliding 
factor (ratio of total horizontal forces 
to total vertical forces) of 0.7 is 
exceeded by as much as 57 percent. This 
is offset by the high values of the shear 
friction factor of safety for this struc- 
ture, in excess of 17 compared to minimum 
allowable for seismic loading of 3.25. 



389 



GENERAL 
LOCATION 




EDWARD HYATT 
POWERPLANT 
(UNDERGROUND) 



Fi gure II 



Location Map, Edward Hyatt Powerplant Facilities 



390 




Figure 190. Fish Barrier Dam 



391 




392 




393 




-i! SUTT&R-BUTTE^ ; ILARKIN Vi 
oil CAMAL OUTLET I ^ROAD j j 



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Figure 193. Location Map, Thermal ito Powerplant, Forebay, and Afterbay 



394 



Recommendation for Seismic Reanalysis 

The Department recommended using a pseu- 
dostatic analysis and a seismic coeffi- 
cient of 0,25 to reanalyze the secondary 
structures, in lieu of the rigorous 
dynamic analysis used for the critical 
structures. The chairman of the Earth- 
quake Consulting Board concurred with 
this recommendation. 

A quick check using the pseudostatic 
analysis and a seismic coefficient of 
0.25 and 0.6 indicated a minimum shear 
friction factor of safety in excess of 9. 

A brief study of the consequences of 
failure of this structure indicates the 
following: 

1. Possible loss of life would be 
limited to fishermen along the river 

_ at the time of the event. This 
f' would be further limited to those 

fishermen to close to the dam to be 

warned. 

2. Property damage would be minor and 
would be less than that caused by 
the Standard Project Flood. 

3. Loss of the dam would have little 
effect on operation of the project 
until repaired or replaced. Complete 
loss of the dam would have an effect 
on the operation of the fish 
facilities. 

On the basis of the preceding analysis , 
the Department has determined that no 
additional seismic analysis is recommended 
for the Fish Barrier Dam. 

Power and Pumping Plant Facilities 

Edward Hyatt Powerplant 

The Edward Hyatt Powerplant is an under- 
ground, hydroelectric, pumping-generating 
facility located on the Feather River 
approximately 5 miles northeast of the 
City of Oroville, Butte County. 

The powerhouse chamber, located in the 
left abutment near the axis of Oroville 



Dam (Figure 189) , was excavated in a 
metavolcanic rock formation that is pre- 
dominately amphibolite. The rock was 
fresh and three prominent joint sets 
imparted a certain blockiness to it, but 
the individual joints were generally 
tight . 

Since much of the powerhouse is placed 
against the rock, (Figure 194) , it can 
be assumed that it will experience peak 
ground acceleration (PGA) with negligible 
magnification. The powerhouse substruc- 
ture is a rigid, massive, reinforced 
concrete structure which should exper- 
ience little or no distress from the 
designated load factor. 

As a result of the DWR Earthquake Hazard 
Committee inspection of the facility, 
minor modifications consisting princi- 
pally of installations of additional 
holddown bolt anchorages and bracing 
were made to increase earthquake resis- 
tance of unit control centers, emergency 
equipment, spare parts storage shelves, 
CO2 cylinder racks , and numerous other 
items. Items in the power plant still 
to be investigated are the anchorages 
that fasten the precast wall panels to 
the columns; and the columns themselves, 
which rest on the generator floor, Ele- 
vation 252.0 (Figure 195) . The panels 
will most likely experience a higher 
response acceleration factor since their 
mode of vibration will be considerably 
different from that of the main struc- 
ture. A pseudostatic analysis using a 
peak ground acceleration of 0.25g will 
be used to investigate the powerhouse 
components . This work is scheduled to 
be completed during the 1978-79 fiscal 
year. 

It was determined that the intake struc- 
ture to the powerhouse (Figure 196) was 
structurally stable; however, additional 
anchorages were installed for a number 
of items, the most important of which 
are the crane trolleys on the shutter 
gantry crane. Holddowns are required 
for the trolleys when the crane is not 
in operation to keep them on the track 
during seismic events. 



395 




Hi: 



4F 

Is 
1r 



396 




Figure 196. Overall View of Edward Hyatt Powerplant Intake Structure 



398 



Conclusion 

The powerhouse substructure has been 
reviewed using a comparative pseudo- 
static analysis of previously designed 
powerhouse substructures. Based on 
this comparison, it has been determined 
that this substructure would be capable 
of resisting the forces induced by a 
0.25g peak ground acceleration, therefore 
no modifications are required. 

Modifications will be made to improve 
the seismic resistance of powerhouse 
superstructure components as necessary. 

Thermalito Powerplant 

Thermalito Powerplant is a pumping- 
generating facility located approximately 
4 miles west of the City of Oroville, 
Butte County (Figure 193) . The power 
plant substructure (Figure 197) is 
keyed into a basalt formation, and its 
foundation lies on an interflow mater- 
ial consisting of basalt breccia in a 
matrix of amorphous material. The 
plant substructure is a rigid, massive, 
reinforced concrete structure which 
should move with the basaltic rock 
formation and thus experience peak 
ground acceleration (0.25g) with little 
or no magnification. 



This facility was also inspected for 
earthquake hazards, and modifications 
similar to those at the Edward Hyatt 
Powerplant were made. 

Items still to be investigated are the 
rigid steel frames that form the super- 
structure, and the precast concrete 
panels and the anchorages that fasten 
them to the superstructure (Figure 198) . 
This work is scheduled to be completed 
during the 1978-79 fiscal year. Above 
elevation 165.0, the superstructure 
will vibrate in a lower mode of vibration 
than the substructure and therefore 
experience a somewhat higher response 
acceleration factor than the 0.25g 
assigned to the substructure. 

Conclusion 

The powerhouse substructure has been 
reviewed using a comparative pseudo- 
static analysis of previously designed 
powerhouse substructures. Based on 
this comparison, it has been determined 
that this substructure would be capable 
of resisting the forces induced by a 
0.25g peak ground acceleration, there- 
fore no modifications are required. 

Modifications will be made to improve 
the Sjeismic resistance of powerhouse 
superstructure components as necessary. 



399 




400 




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401 



Miscellaneous Structures 

The miscellaneous structures inspected by the DWR Earthquake 
Hazard Committee include those listed below: 

Oroville Operations and Maintenance Center 

Miscellaneous Structures 

1. Administration and Maintenance Center 

2. General Maintenance Headquarters Building 

3. Plant Maintenance Shops 

4. Mobile Equipment Repair Building 

5. General Maintenance Warehouse 

6. Vehicle Storage Building 

Oroville Dam Miscellaneous Structures 

1. Palermo Outlet Works Control House 

2. Instrument Vault 

3. Reservoir Gage Station 

Thermalito Forebay and Afterbay Miscellaneous Structures 

1. Heavy Equipment Building 

2. Western Canal and Richvale Canal Outlet Control Building 

3. PG&E Lateral Outlet Control Building 

4. Sutter Buttes Canal Outlet Control Building 

5. Feather River Outlet Control Building 

6. Feather River Outlet Control Station 

Feather River Fish Hatchery Miscellaneous Structure 

1. Maintenance Office Building 

2. Hatchery-Spawning Building 

3. Ultraviolet Treatment Building 

As a result of the earthquake hazard inspection, additional anchor- 
ages have been installed for much of the operational equipment in these 
facilities. 



402 



I Conclusion 

Damage that may occur to the miscellan- 
eous structures is not considered to be 
a threat to public safety and property. 
For the purpose of the seismic reevalua- 
tion, these structures are classified 
as noncritical. 

i Bridges 

Public bridges in the Oroville area were 
inspected by the Department of Transpor- 
tation following the August 1, 1975, 
Oroville earthquake. Fourteen bridges 
in the general area showed evidence of 
movement, minor damage, or both. 

If these bridges were to experience a 
peak ground acceleration of 0.25g, some 
of them may sustain greater damage . 
Therefore, all project- related bridges 
will be analyzed for such a loading 
during the 1978-79 fiscal year. 

Conclusion 

Bridge components that will not sustain 
the forces generated by a 0.25g peak 
ground acceleration will be modified to 
strengthen their seismic resistance. 

Switchyard Structures and Apparatus 

Switchyards play an important dual role 
in hydroelectric power systems . They 
collect and distribute generated or 
incoming power and protect power and 
pumping plants. Unlike many power plant 
features, switchyard electrical equip- 
ment exhibit only light damping charac- 
teristics, are fragile, and have very 
little ductility because much of the 
supporting systems are porcelain. In 
addition, many of the physical structures 
take on a "lollipop" type of mass dis- 
tribution which is conducive to high 
amplification factors during severe 
earthquake-induced ground movements. 

The two switchyards under investigation 
are at the Edward Hyatt and Thermal i to 
Powerplants . 



The Edward Hyatt Powerplant Switchyard 
is located at the downstream toe of the 
Oroville Dam on the south bank of the 
Feather River, approximately five miles 
northeast of the City of Oroville 
(Figure 189) . 

Thermalito Powerplant Switchyard "W" is 
located several hundred feet west of 
the power plant, approximately four 
miles west of Oroville (Figure 193) . 

Both switchyard facilities contain 
similar electrical apparatus; therefore, 
it will suffice to discuss weaknesses 
common to both of them. 

The types of equipment in the switch- 
yards are: 

1. Current transformers 

2. Potential transformers 

3. Disconnect switches 

4. Lightning arresters 

5. Line traps 

6 . Bus supports 



7. 



230 kV ATB-6 power circuit 
breakers 



Since no loss of life is expected due 
to failure of equipment within the yard, 
the switchyards have been classified as 
noncritical facilities. Accordingly, 
a peak ground acceleration of 0.25g has 
been assigned for this seismic 
evaluation . 

To determine dynamic characteristics of 
switchyard equipment, a testing program 
had previously been conducted at other 
Department switchyards containing similar 
power circuit breakers and structural 
support systems. Those tests had 
revealed critical frequencies and damping 
characteristics . 

However, since a loss of certain switch- 
yard equipment would create only 



403 



operational inconveniences and minor 
outages, investigation efforts were 
directed to the most earthquake prone 
and critical equipment in the switch- 
yards; the 230 kV ATB-6 power circuit 
breakers (Figure 199) . 

The fundamental frequency for the circuit 
breakers ranged from 2.8 to 3.5 hertz, 
depending on the direction of excitation 
applied, and damping values ranged from 
3.6 percent to 6.4 percent of critical. 
Dynamic magnification of about 7 was 
observed during the testing, which 
means at an acceleration input of 0.25g, 
the maximum response acceleration would 
be about 1.75g. Further studies indi- 
cated that the porcelain supports for 
the breaker heads cannot withstand 
loads of this magnitude without addi- 



tional damping or isolation. M 

For the proposed study a newly developed 
seismic shock-isolation system will be 
tested under one of the power circuit 
breakers to determine its capability to 
protect the breakers during severe 
ground movements. The testing program 
is scheduled to be conducted during the 
1978-79 fiscal year. 

Conclusion 

Based on the consideration that failure 
of electrical equipment in the Edward 
Hyatt or Thermalito Powerplant switch- 
yards does not pose a threat to public 
safety or property, the switchyards are 
classified as noncritical elements of 
the Oroville Complex. 




Figure 199. 230-KV Power Circuit Breakers 



404 



CHAPTER IX 
CONTINGENCY PLAN FOR SEISMIC EMERGENCIES 



rhe purpose of this chapter is to specify 
procedures to be followed by the Division 
of Operations and Maintenance in reacting 
to seismic events, and the process for 
returning equipment to pre-earthquake 
operating levels. 

The discussion outlines (1) organization 
and responsibilities for both Headquarters 
and Field Division; (2) procedures for 
reacting to seismic events, including 
notification and response; and (3) proce- 
dures returning equipment and facilities 
to full operational status, which estab- 
lishes the criteria for qualifying opera- 
tional readiness. 

Organization and Responsibilities 

Division Policy 

The policy of the Division of Operations 
and Maintenance during emergencies will 
be as follows: 

1. Management and operation of the State 
Water Project during emergencies will 
be in accordance with the contingency 
plan. 

2. The primary emphasis will be to pro- 
vide all possible support to line 
activities. 

3. The project facilities and control 
centers will continue to be operated 
in accordance with the Plan of 
Operation. 

4. Operation will be on as nearly a 
normal basis as possible. 

5. The Division Command Post in Sacra- 
mento, together with a Command Post 
in each of the five field divisions, 
if required, will be set up to coordi- 
nate activities. 

6. A declaration of emergency will be 
made by the Director upon recommenda- 
tion by the Division Engineer, if 



required for the Division to expe- 
dite contracts, service agreements, 
and purchasing processes. 

7. Instructions to cease operation of 
facilities, to operate additional 
facilities, or to change the plan 
of operation shall be made only by 
the Division Engineer or by the per- 
son in charge of the Division Com- 
mand Post, after consultation with 
the Director, except in emergencies 
where immediate action is required. 

Division Plan of Operation 

Should an emergency occur, the Operations 
Control Branch will prepare, distribute, 
and implement a Plan of Operation appro- 
priate to the situation. The Oroville 
power facilities will be kept in opera- 
tion by all feasible means . 

Where possible, the pumping and genera- 
ting units will be operated on a contin- 
uous basis to minimize manpower require- 
ments on start-up and shutdown, to 
lessen chances of shutdown. 

Priority will be given to continuing the 
operation at the Hyatt Powerplant should 
circumstances require discontinuing opera- 
tion of the Thermalito facilities. 

Oroville Field Division Command Post 

On determination of a seismic emergency, 
the Field Division Chief, upon direction 
or at this discretion, will establish a 
Command Post to be manned 24 hours per 
day to assure continuous communication 
with the Division Command Post. 

The chain of command will not change 
because of emergency conditions. While 
the urgency of the moment may necessi- 
tate shortcuts in lines of communications 
and activities, every effort will be 
made, as time permits, to backtrack and 
reestablish that chain. 



405 



Every effort shall be made to maintain 
the normal channel of communication 
between the field and headquarters for 
the dissemination of operating orders, 
condition reports, and forecasts. 

The Oroville Field Division has estab- 
lished two command posts; one is desig- 
nated "Operational", which is the 
Master Command Post, and the other is 
designated as the "Security Command 
Post". 

Operational Command Post . The Opera- 
tional Command Post will be the respon- 
sibility of the Operations Superinten- 
dent under the authority of the Field 
Division Chief. He will have the 
responsibility to meet the Plan of 
Operation for water and power deliveries. 

1. Operational Facilities . To the 
extent possible, the following 
facilities in the Oroville Field 
Division will remain operational 
for the duration of any emergency: 

Area Control Center 
Command Post - Operational 
Edward Hyatt Powerplant 
Thermalito Powerplant 
All Hydraulic Structures 

2. Operational Plan . The above facil- 
ities will be operated and main- 
tained on as nearly a normal basis 
as possible and in accordance with 
existing power and water delivery 
schedules. The Operations Super- 
intendent or his designee shall 
keep the Project Operations Con- 
trol Center advised of all conditions 
relative to the operation of the 
Project. 

If for any reason, communication links 
between the Oroville Operations Command 
Post and Project Operations Command 
Center are severed, the Operations 
Command Post shall maintain the same 
operating status that prevailed prior 
to the loss of communications. Any 
deviations from this procedure must be 



authorized by the Operations Superin- 
tendent or his designee. 

Security Command Post . The Security 
Command Post will be the responsibility 
of the Chief of the Civil Maintenance 
Section under the authority of the Field 
Division Chief. 

1. The Security Plan . The Security 
Command Post Manager will: 

(a) direct and monitor Project 
facility inspections of civil 
features; 

(b) recommend to the Field Divi- 
sion Chief operational devia- 
tions when warranted due to 
detectable threats to the 
integrity of Project structures; 
and 

(c) maintain a log of damage and 
establish priorities for repair 
work. 

The Security Plan involves two actions: 
Surveillance and Corrective Response. 

Surveillance is defined as efforts 
expanded toward knowing what is happen- 
ing project wide. Corrective response 
is defined as efforts expended toward 
maintaining integrity or restoration of 
facilities. 

Procedures for Reacting 
To Seismic Events 

Oroville Field Division 

Detection 

Seismic events are detected either by 
personal senses in the Field or by instn 
mentation response in the Earthquake 
Engineering Section in Sacramento. 

Earthquake Magnitude and Epicenters 

Accurate earthquake magnitude and hypo- 
center are determined and reported by 



406 



the Earthquake Engineering Section in 
Sacramento, normally well beyond the 
time required to help determine the need 
for Project inspections. Preliminary 
estimates of magnitudes above 4 . are 
made by the Project Operations Control 
Center and generally reported to the ACC 
within 30 minutes after the event. 
This information is helpful but not 
complete enough to determine the need 
for inspection. 

Criteria for Notification 

When an earthquake exceeding an estimated 
magnitude of 3.0 Richter Scale is felt or 
reported with 25 kilometres (15.5 miles) 
of the Oroville Seismic Reporting Sta- 
tion, the Area Control Center shall 
notify the Surveillance Unit Chief, the 
State Police Mobile Unit on duty, and 
the Project Operations Control Center. 

When an earthquake is felt or reported 
in excess of an estimated magnitude 4.0 
Richter Scale, within 50 kilometres 
(31 miles) of the reporting station, the 
Area Control Center shall additionally 
notify the Civil Maintenance Section 
Chief. 

Within the near future, peak accelera- 
tion values measured in the foundation 
and crest of Oroville Dam will be dis- 
played in the Area Control Center on 
the Data Acquisition Panel. At that 
time, O.lOg recorded at the base of 
Oroville Dam will replace the* 3.0 Richter 
Scale criteria for notification, and 
0.15g will replace the 4.0 Richter Scale 
criteria. 

Response 

1. When an earthquake estimated in 
excess of magnitude 3.0 Richter 
Scale occurs, or the acceleration 
values are in excess of O.lOg: 

(a) Water and Power Operations 
shall continue to operate 
under routine constraints. 



(b) The Surveillance Unit Chief 
shall determine that seismic 
instrumentation is functioning 
and ready to record a subse- 
quent event, and gather and 
record certain data in accor- 
dance with prescribed standing 
instructions. 

(c) The State Police mobile unit on 
duty will perform a general 
inspection of the Project in 
accordance with prescribed 
standing instructions. 

When an earthquake exceeding an 
estimated magnitude 4.5 Richter 
Scale occurs, or with the determina- 
tion that peak accelerations exceed- 
ing 0.15g have occurred at the base 
of Oroville Dam: 

(a) Water and Power Operations 
shall continue at the operating 
level resulting from a seismic 
event. For instance, if four 
units are on line when the event 
occurs and only two units remain 
on line , water and power oper- 
ations shall remain in that 
configuration until either: 

(1) standard operating proce- 
dures dictate a change, or (2) 
an inspection reveals that it 
is safe to continue or to 
increase/decrease operations. 

(b) The Security Command Post shall 
be activated by the Civil Main- 
tenance Section Chief at his 
discretion to an estimated mag- 
nitude 5.0, he may call the 
necessary personnel for an 
inspection. At an estimated 
magnitude 5.0, the complete 
Rapid Response inspection is 
mandatory. The ACC shall be 
advised when the Security Command 
Post is activated and when a 
Rapid Response Inspection has 
been initiated. 



407 



Inspection of Project Facilities 
Following an Earthquake 

Significant seismic events are presumed 
to precede additional seismic events. 
Until enough time has elapsed to perceive 
a decay in the frequency and magnitude 
of shocks, inspection and investigations 
will continue in order to assure that 
structures have not been weakened to a 
point approaching failure. 

Inspections are categorized two ways: 
Rapid Response and Follow-up or Final. 

1. The Rapid Response Inspection Plan 
is a method to immediately determine 
that problems may or may not be 
developing projectwide. It is a 
means for the Field Division Chief 
to determine whether corrective 
action should be initiated and to 
what degree follow-up inspections 
are required. 

(a) Operational Command Post . The 
Operations Section is obligated 
to comply with Operational Pro- 
cedures (OP series of instruc- 
tions) that specifically outline 
the inspection and safe operation 
of plant equipment. In addition, 
the Area Control Center has on 
its Status Display boards, indi- 
cators of water and power condi- 
tions of the overall Project 
Features. Rapid inspection and 
response therefore are well 
outlined. 

(b) Security Command Post . The area 
of responsibility of the Security 
Command Post is divided geograph- 
ically and assigned to three 
inspection teams. (See Schematic 
Diagram of Oroville Complex 
Figure 200) . 

(1) Oroville Dam and vicinity 

(2) Thermalito Forebay and 
vicinity 

(3) Thermalito Afterbay and 
vicinity 



Detailed instructions of routes to fol- 
low and items to inspect are provided 
for each team. In addition a training 
film has been prepared that defines the 
plan of inspection, shows the order of 
inspection, the types of damage to look 
for in each geographical area, and 
specifies the procedure to be used in 
reporting damage and conditions. 

The surveillance crew, within all these 
geographical areas, attends seismic and 
performance instrximentation. 

2. The follow-up Inspection Plan is 
developed as information becomes 
available. In some cases visual 
inspections may be adequate. Gen- 
erally, when accelerations of less 
than 0.15g are experienced in Oro- 
ville Dam, follow-up inspections 
will not be required. Damage reports 
will be filled out for all abnormal- 
ities to civil features discovered 
during inspections. Damage reports 
will serve to identify large or 
small problems and to initiate Job 
Requests or Job Orders as may be 
determined by review and verifica- 
tion of the Security Command Post 
Manager . 

Returning Facilities and Equipment 
To Full Operation Status 

The following pages are lists of all 
operational facilities and features in 
the Oroville Complex, starting with Lake 
Oroville, moving downstream, and ending 
with the Thermalito Afterbay ground water 
pumping system. 

The primary features of concern in the 
Lists of Operating Criteria are the four 
bodies of water that can be regulated in 
the Oroville Complex which are : 

Lake Oroville (Page 414) See Figure 201 
Thermalito Diversion 

(Page 417) see Figure 202 

Thermalito Forebay 

Reservoir (Page 418) See Figure 202 
Thermalito Afterbay 

Reservoir (Page 420) See Figure 203. 



408 



Regulating Feature 

The features that can be regulated have 
a normal operating criteria, and cui 
emergency operating condition. The 
emergency condition is dictated by the 
degree, or potential degree, of loss of 
structural integrity. The listed cri- 
teria is mandatory and must be met, 
either on an interim basis while investi- 
gations continue, or as assurances of 
integrity come in from either field or 
plant investigations . Many of the 
criteria conditions can be determined 
quickly, whereas others take time and 



testing for full assurance that they 
are validly met. 

Nonregulating Features 

Nonregulating features are those struc- 
tures that cannot be operated, such as 
Oroville Dam or the Thermalito power 
canal. Certain conditions as are 
described after each nonregulating 
feature constitute a declaration of true 
emergency for that feature. Water and 
power operations or both then become 
secondary to whatever response is 
required to protect lives or reduce 
further damage. 



409 



OROVILLE DAM S VICINITY 




SUTTER BUTTE OUTLET 



THERMALITO 
AFTERBAY 



Figure '200. Schematic Diagram of Oroville Complex 



410 



OROVILLE DAM AND VICINITY 























LAKE OROVILLE 






WATER FLOW 


PATH 1 


A 


^ L 






^ 








^ 






REGULATING 
FEATURES 






NONREGULATING 
FEATURES 




^ 




1 




^ 




1 

1 




r 




i 




i 




y 








1 




PALERMO 

OUTLET 




FLOOD 
CONTROL 
SPILLWAY 




RIVER OUTLET 
SPHERE VALVES 




HYATT INTAKE 
SHUTTERS 




OROVILLE 
DAM 


u 

1 


PARISH CAMP 
SADDLE DAM 


1 


r 










^r 










1 
1 

1 




EXPORT 






HOWELL 
BUNGER 
VALVES 




INTAKE 
GATES 




BIDWELL 

CANYON 

SADDLE DAM 


-L 


EMERGENCY 
SPILLWAY 


/ 






















/ 


^ 




HYATT 
POWER PLANT 




¥■ 














r 








^ 


r 




y 






TAP GUARD 
VALVES 




TURBINE 

SHUT-OFF 

VALVES 














^ 


r 










WICKET 
GATES 












iV 








w 






UNITS 






■ 










1 


f 




1 


f 




1 








CONDENSING 




GENERATING 




PUMP BACK 












^ 




A 


\. 








k 


THERMALITO DIVERSION 
POOL 














▼ 








i 


L 




L EG END 



Figure 201. Schematic Diagram of Oroville Dam and Vicinity 



411 



THERMALITO FOREBAY AND VICINITY 



RADIAL GATES 



BYPASS GATE 



CONDENSING 



THERMALITO DIVERSION 
POOL 



WATER FLOW PATH 



REGULATING 
FEATURES 



DIVERSION DAM 
RELEASE 
FEATURES 



HOWELL 
BUNGER 
VALVE 



FEATHER 
RIVER 



REGULATING 
FEATURES 



THERMALITO 
INTAKE 



THERMALITO 
POWER PLANT 



WICKET GATES 



GENERATING 



FISH 
HATCHERY 



NONREGULATING 
FEATURES 



THERMALITO 

DIVERSION 

DAM 



POWER CANAL 
RADIAL GATES 

I A 



POWER 
CANAL 



THERMALITO FOREBAY 
RESERVOIR 



NONREGULATING 
FEATURES 



THERMALITO 

INTAKE 
STRUCTURE 



POWER 
CANAL 



TL. 



THERMALITO 

FOREBAY 

DAM 



PUMP BACK 

*~ 



L EG END 

WATER FLOW PATH 



TAIL 
CHANNEL 



THERMALITO AFTERBAY 
RESERVOIR 



Figure 202. Schematic Diagram of Thermal i to Forebay and Vicinity 



412 



THERMALITO AFTERBAY AND VICINITY 



THERMALITO AFTERBAY 
RESERVOIR 

:*n 



WATER FLOW PATH 



REGULATING 
FEATURES 



RIVER 
OUTLET 
GATES 



SUTTER 
BUTTE 
GATES 



PGSE 
GATES 



FEATHER 
RIVER 



NONREGULATING 
FEATURES 



GROUND WATER 
PUMP SYSTEM 



WESTERN 

RICHVALE 

GATES 



THERMALITO 

AFTERBAY 

DAM 



LEGEND 
^ WATER FLOW PATH 

Figure 203- Schematic Diagram of Thermal ito Afterbay and Vicinity 



413 



LIST OF OPERATING CRITERIA FOR REGULATING LAKE OROVILLE 
DECISION MAKING CRITERIA FOR OPERATING FEATURES WHICH CAN BE REGULATED 

A. PALERMO OUTLET 

1. 30-inch butterfly valve {1 valve) 

a. Must be closed when fixed dispersion cone valve 
integrity is questionable. 

b. Must be closed when Palermo intake works integrity is 
questionable. 

c. Must be open to meet water delivery demand by Oroville 
Wyandotte Irrigation District. 

2. 12-inch fixed dispersion cone valve (1 valve) 

a. Must be closed when 30-inch butterfly valve integrity 
is questionable. 

b. Must be closed for Palermo Canal failure. 

c. Must be open to meet water delivery demand by 
Oroville Wyandotte Irrigation District. 

B. OROVILLE DAM SPILLWAY 

1. Radial gates (8 gates) 

a. Must be available for maintenance of the flood 
control reservation. 

b. Must be available for Lake Oroville regulation. 

C. RIVER OUTLET VALVES 

1. Spherical valves (2 valves) 

a. Must be closed when Howell Bunger valves integrity 
questionable. 

b. Must be closed when intake integrity is questionable. 

c. Must be open when Howell Bunger valves will be 
operated . 

d. Should be closed when Howell Bunger valves are not 
needed. 

2. Howell Bunger valves (2 valves) 

a. Must be closed when diversion tunnel No. 2 has 
severe blockage. 

b. Must be closed when spherical valve integrity is 
questionable. 

c. Must be opened when water delivery is unavailable 
thru Hyatt Powerplant. 



414 



D. EDWARD HYATT INTAKE 

1. Intake gates (2 gates) 

a. Must be closed during a powerhouse disaster. 

b. Must be closed for Penstock, turbine shutoff valve 
or tap guard valve damage. 

c. Must be open to meet Edward Hyatt Powerplant water 
demand . 

2. Shutters (26 shutters) 

a. Enough must be removed in order to comply with 
Shutter Submergence Criteria. 

b. Enough must be in place in order to control the 
temperature of water releases. 

c. Enough should be in place to provide protection 
from debris at lower levels. 

E. EDWARD HYATT POWERPLANT 

1. Turbine shutoff valves (6 valves) 

a. Must be closed if powerplant is flooded. 

b. Must be closed if penstock integrity is questionable. 

c. Must be open for unit operational demand. 

d. Should be open to control auto oscillation, 

e. Should be closed if tailrace tunnel is blocked. 

f . May be open if intake gates are closed. 

2. Tap guard valves (6 valves) 

a. Must be open for raw and cooling water supply to Edward 
Hyatt Powerplant units . 

b. Must be closed for failure of cooling water line 
below the valves. 

c. Should be closed in the event of powerplant flooding. 

3. Wicket gates (6 gate sets) 

a. Must be open if units are generating or pxomping. 

b. Must be closed for synchronous condensing. 

c. May be closed for unit shut down. 

d. May be closed for testing purposes of sediment bearing 
water appearing in the tailrace tunnel. 

4. Unit operation- synchronous condensing (6 units) 

a. Cannot be done if units are generating or pumping. 

b. May be done for power system stability (as requested 
by the power company.) 



415 



5. Unit operation-generating (6 units) 

a. Cannot be done if intake gates are closed. 

b. Cannot be done if powerplant disaster exists. 

c. Must not be done if unit integrity is questionable. 

d. Must not be done if an incompatible stage differential 
exists between Lake Oroville and Thermalito Diversion 
Pool. 

e. Cannot be done if units are pumping or synchronous con- 
densing. 

f. May be done to provide station service. (units 1 or 4). 

g. Must not be done if tailrace tunnels are blocked. 

h. Should not be done if switchyard or transmission line integrity 

are not OK. 
i. Should not be done if no power or water demands exist, 
j . Should be done if regulation of Lake Oroville is 

required. 

6. Unit operation-pump back (3 units) 

a. Cannot be done if intake gates are closed. 

b. Cannot be done if powerplant disaster exists. 

c. Must not be done if unit integrity is questionable. 

d. Must not be done if an incompatible stage differential 
exists between the Thermalito Diversion Pool and Lake 
Oroville. 

e. Cannot be done if units are synchronous condensing or 
generating. 

f. Must not be done if switchyard or transmission lines 
are not OK. 

g. Should not be done if tailrace tunnel is blocked, 
h. Should be done at the power company request. 

i. Should be done for water conservation. 

j. May be done for Diversion Pool regulation. 

II. CRITICAL CONDITIONS FOR FEATURES WHICH CANNOT BE REGULATED 

A. OROVILLE DAM 

1. Uncontrollable water born sediment passing thru embankment, 
groins, foxindation, grout gallery or tailrace tunnel. 

2. Significant slips or cracks in embankment. 

3. Crest settlement which could lead to overtopping. 

4. Significant vertical or horizontal displacement. 

5. Excessive increases in pore pressure. 

B. BIDWELL CANYON SADDLE DAM 

1. Uncontrollable water born sediment thru embankment, groins 
or foundation. 

2. Significant slips or cracks in embankment. 

3. Crest settlement which could lead to overtopping. 

4. Significant vertical or horizontal displacement. 



416 



C. PARISH CAMP SADDLE DAM 

1. Uncontrollable water born sediment thru embankment, groins 
or foundation. 

2. Significant slips or cracks in embankment. 

3. Crest settlement which could lead to overtopping. 

4. Significant vertical or horizontal displacement. 

D. OROVILLE DAM SPILLWAY 

1. Extraordinary damage to concrete monoliths or groins. 

2. Inoperable radial gates limit ability to maintain flood 
control reservation. 

E. EDWARD HYATT INTAKE AND PENSTOCK 

1. Extraordinary damage to structure, trash racks, emergency 
gates, or penstocks. 

F. PALERMO INTAKE AND OUTLET 

1. Loss of control of water thru outlet. 

G. RIVER OUTLET VALVE CHAMBER 

1. Loss of control of water thru outlet. 

2 . Extraordinary damage to tunnel plugs . 



LIST OF OPERATING CRITERIA FOR REGULATING 
THERMALITO DIVERSION POOL 

DECISION MAKING CRITERIA FOR OPERATING REGULATING FEATURES 

A. THERMALITO DIVERSION DAM 

1. Radial gates (1 4 gates) 

a. Must be open to pass flood control waters to 
Feather River. 

b. Should be operable to regulate waters in Thermalito 
Diversion Pool and Thermalito Forebay. 

c. Should be closed for Thermalito Powerplant operations. 

d. May be used to meet water delivery commitments in the 
Feather River. 

2. Howell Bunger valve (1 valve) 

a. Should be open to maintain minimum water release 
for stream flow maintenance in the Feather River. 

3. Fish hatchery valve (1 valve) 

a. Should be open to meet water demand from the 
Feather River Fish Hatchery. 



417 



4, Radial gates to power canal (3 gates) 

a. Must be closed to protect the Thermalito Power Canal 
and Thermalito Forebay during extreme flood control 
conditions in Thermalito Diversion Pool. 

b. Must be open to keep Thermalito Powerplant operational. 

c. May be closed if Thermalito Power Canal or Thermalito 
Forebay Dam integrity is questionable. 

II. CRITICAL CONDITIONS FOR FEATURES WHICH CANNOT BE REGULATED 

A. THERMALITO DIVERSION DAM 

1. If uncontrollable water born sediments are passing thru 
the groins or foundation. 

B. THERMALITO POWER CANAL HEADWORKS 

1. Extraordinary damage to radial gates. 



LIST OF OPERATING CRITERIA FOR REGULATING THERMALITO FOREBAY 
RESERVOIR AND POWER CANAL 

I. DECISION MAKING CRITERIA FOR OPERATING REGULATING FEATURES 

A. THERMALITO INTAKE STRUCTURE 

1. Bypass gate (1 gate) 

a. May be open when Thermalito Powerplant is not 
operational . 

b. May be used for Thermalito Afterbay regulation. 

c. May be used for Thermalito Forebay regulation. 

2. Fixed wheel gates (2 gates) 

a. Must be used for uncontrollable water thru units. 

b. Should be used for penstock rupture. 

B. THERMALITO POWERPLANT 

1. Wicket gates (4 gate sets) 

a. Must be open if units are generating or pumping. 

b. Must be closed for synchronous condensing. 

c. May be closed for unit shut down. 

2- Unit operation-synchronous condensing (4 units) 

a. Cannot be done if units are generating or pximping. 

b. May be done for power system stability (as requested 
by the power company) . 



418 



Unit operation-generating (4 units) 



Cannot be done if fixed wheel gates are closed. 

Cannot be done if a powerplant disaster exists. 

Must not be done if unit integrity is questionable. 

Must not be done if an incompatible stage differential 

exists between the Thermalito Forebay and the 

Thermalito Tail Channel. 

Cannot be done if units are pumping or synchronous 

condensing. 

Must not be done if tail channel is blocked. 

Should not be done if switchyard or transmission 

line integrity are not OK. 

Should not be done if there are no power or water 

demands . 

May be done for regulation of Thermalito Forebay or 

Thermalito Afterbay. 

May be done to provide station service. 



4. Unit operation-pumpback (3 units) 



b. 



Cannot be done if fixed wheel gates are closed. 

Cannot be done if powerplant disaster exists. 

Must not be done if unit integrity is questionable. 

Must not be done if an incompatible stage differential 

exists between the Thermalito Forebay and the 

Thermalito Tail Channel. 

Cannot be done if units are synchronous condensing 

or generating. 

Must not be done if switchyard or transmission 

lines are not OK. 

Should not be done if tail channel is blocked. 

Should be done at the power company request. 

Should be done for water conservation. 

May be done for Thermalito Afterbay or Thermalito 

Forebay regulation. 



II. CRITICAL CONDITIONS FOR FEATURES WHICH CANNOT BE REGULATED 

A. THERMALITO FOREBAY DAM 

1. If uncontrollable water born sediments are passing thru 
the embankment, groins or fovindation. 

2. Significant slips or cracks in embankment. 

3. Crest settlement which could lead to overtopping. 

4. Significant vertical or horizontal displacement. 

B. THERMALITO INTAKE STRUCTURE 

1. Uncontrollable water passing thru or under intake structure. 

2. Extraordinary damage to structure, trash racks or bypass 
gate. 

3. Extraordinary damage to end wall gravity dam. 



419 



C. THERMALITO POWER CANAL (Cut Section) 

1. Extraordinary damage to canal cut or lining. 

D. THERMALITO POWER CANAL (Fill Section) 

1. If uncontrollable water born sediments are passing thru 
the embankment, groins, or founiation 

2. Significant slips or cracks in embankment. 

3. Crest settlement which could lead to overtopping. 

4. Significant vertical or horizontal displacement. 

LIST OF OPERATING CRITERIA FOR REGULATING 
THERMALITO AFTERBAY RESERVOIR 

DECISION MAKING CRITERIA FOR OPERATING REGULATING FEATURES 

A. THERMALITO AFTERBAY RIVER OUTLET 
1. Radial gates (5 gates) 

a. Should be open to meet Feather River stream flow 
maintenance commitments. 

b. May be open to regulate Thermalito Afterbay 
Reservoir. 

B. SUTTER- BUTTE OUTLET 

1. Slide gates (4 gates) 

a. Must be open to meet Sutter-Butte Irrigation 
District water demands . 

C. PG&E OUTLET 

1. Slide gates (1 gate) 

a. Must be open to meet PG&E water demands. 

D. WESTERN CANAL AND RICHVALE OUTLETS 

1. Slide gates (8 gates) 

a. Must be open to meet Western Canal (PG&E) or 
Richvale Irrigation District water demands. 

E. THERMALITO AFTERBAY DAM GROUND WATER PUMPING SYSTEM 
1. Ground water pumps (15 pumps) 

a. Must be on to maintain ground water aquifer level. 

b. Must be off to prevent overdraft of ground water 
aquifer. 



420 



II. CRITICAL CONDITIONS FOR FEATURES WHICH CANNOT BE REGULATED 

A. THERMALITO AFTERBAY DAM 

1. If uncontrollable water born sediments are passing 
thru the embankment^ groins^ or foundation. 

2. Significant slips or cracks in embankment. 

3. Crest settlement which could lead to overtopping. 

4. Significant vertical or horizontal displacement. 

B. THERMALITO POWERHOUSE STRUCTURE 

1. Uncontrollable inflow into Thermalito Tail Channel. 

2. Extraordinary damage to Thermalito Powerhouse. 

C. THERMALITO AFTERBAY RIVER OUTLET 

1. Extraordinary damage to structure or radial gates. 

D. SUTTER- BUTTE OUTLET 

1. Extraordinary damage to structure. 

E. PG&E OUTLET 

1, Extraordinary damage to structure. 

F. WESTERN-RICHVALE OUTLETS 

1. Extraordinary damage to structure. 



Commentary 

Seismic emergencies at the Oroville Field 
Division are addressed through the pre- 
ceding contingency plan. 

Priorities for inspection and work are 
established on the following basis: 

1. Attention to those structures whose 
failure can lead to loss of life 
and significant property damage. 

2. Attention to those facilities whose 
operation can lead to further damage 
or failure to Project structures. 

3. Attention to those structures whose 
condition may lead to limiting the 



ability to meet power and water 
delivery commitments . 

4. Attention to those structures and 
facilities that are supportive to 
efficient operation but are not in 
themselves, critical to priorities 
1, 2, and 3. 

Conclusion 

The contingency plan is attentive to 
established Division Policy; it provides 
for detection, notification, and response 
to seismic events. The plan also 
includes a list of operational facilities 
and features along with criteria that 
must be met before returning to pre- 
earthquake operating status. 



421 



APPENDIX A 

REPORTS PREPARED BY THE SPECIAL CONSULTING BOARD 

AND RESPONSES BY THE DEPARTMENT OF WATER RESOURCES 



1. Reports of the Consulting Board for Earthquake 
Analysis, 11 August 1975 424 

2. Report of the Special Consulting Board for the 
Oroville Earthquake to Mr. R. B. Robie, 

12 September 1975 429 

3. Memorandum, "Proposed Department Activities 
in Response to Consulting Boards." from 
Robert W. James to Mr. Ronald B. Robie, 

October 30, 1975 436 

4. Report of the Special Consulting Board for the 
August 1, 1975 Earthquake to Mr. R. B. Robie, 

23 November 1976 444 

5. Memorandum, "Proposed Department Activities 
in Response to the Special Consulting Board 
Meeting of November 22 and 23, 1976," from 

Robert B. James to Ronald B. Robie, March 4, 1977 . . 452 



423 



11 August 1975 



REPORT OF THE CONSULTING BOARD FOR 
EARTHQUAKE ANALYSIS 



Messrs: Robert Jansen 

H. G. Dewey, Jr. 

Gentlemen : 

At a meeting on 8 August 1975 with the Consulting Board for 
Earthquake Analysis, staff members of the Department of Water 
Resources review^ed the instrumental and other data obtained in the 
vicinity of Oroville Dam as a result of the series of moderate 
earthquakes which have occurred in that region, and presented 
evaluations of the performance of the SWP facilities in response to 
the earthquake effects. At the conclusion of the briefing the Board 
was asked to respond to questions relating to the earthquakes and 
possible future events. Our responses are presented below. 

Question 1 

The designs of the SWP facilities in the Oroville area were predicated 
upon certain appraisals of probable future regional seismicity in the 
site vicinities. In view of the recent earthquake activity in the 
Oroville area are the original appraisals still valid? What adjustments, 
if any, should be made in those appraisals? 

la. Although the original appraisal by DWR staff that "the 
Oroville dam site is in an area of relatively light seismic activity" 
may have been justified by the data available at the time of design 
(1958), it should now be modified in the light of the recent earthquake 
activity in the Oroville area and of knowledge gained since 1958. In 
view of the developments, it is appropriate to consider that earthquakes 
ranging up to magnitude 6. 5 may occur within a few miles of the dam 
site. 



424 



Page 2 



Question 2 

What factors should be examined in determining if recent appearance 
of extended seismic activity is related to the 10-year existence of 
Oroville Reservoir in the area of activity? 

2a. If studies along these lines in other parts of the world 
are any indication, investigations of this relationship are likely to be 
quite difficult, and even inconclusive. Nevertheless, a number of 
sets of observations may throw some light on the matter and it is 
essential that they be made soon and as precisely as possible. 

a) The DWR should arrange to have repeat geodetic surveys 
made of past triangulation nets and, particularly, level lines in the 
region of the reservoir and recent earthquake activity. These surveys 
should be made as quickly as practicable and perhaps repeated after 

a few months. 

b) DWR should undertake a timed chemical explosion in a 
borehole near the epicenter of the main shock (August 1) while the 
majority of field seismographs are still operating in the earthquake 
area. Such an explosion is a proven way of calibrating the location 
of earthquake foci in the region against a seismic source with known 
position. In addition, the calibration explosion will calibrate the 
polarity of the seismograph responses thereby enabling more reliable 
fault-plane solutions to be computed. 

c) It is essential that surface fractures possibly associated 
with faulting at depth be carefully delineated and documented without 
delay. Low-altitude, low-sun-angle aerial photography may be of 
assistance in this effort. It is also critical that the local surficial 
geology be more completely mapped and better understood, if surface 
faulting has indeed occurred. Two critical questions are: (1) Have 
the earthquakes occurred along a pre-existing fault, particularly one 
with Quaternary displacement and (2) can the causative fault be traced 



425 



Page 3 

beyond the area of recent earthquakes into areas of future hazard? 
In all geological efforts, DWR personnel should coordinate their work 
with those of other groups studying the earthquake. 

d) Another line of investigation of any connection between the 
reservoir and recent earthquake activity depends upon computation 
of stresses in crustal models loaded by appropriate surface forces. 
Theoretical work along these lines should be supported and evaluated 
against the Oroville data. 



Question 3 

What does the board recommend in the way of immediate and future 
seismic data collection? Seismic data evaluation? 

a) All instrumental data recorded by DWR in connection with 
the Oroville earthquakes should promptly be put into usable form 
and published. The significance to the Department and to scientific 
and engineering communities cannot be overestimated, and care must 
be taken not only to adequately preserve the original records, but 
also to reproduce the data in suitably annotated graphic form; and 
when appropriate, records should also be digitized. As soon as the 
recorded data are in an easily understood form, it is requested that 
copies be provided to each Board member. 

b) DWR should establish a permanent telemetered seismic 
station near the epicenter of the August 1st main shock, and temporary 
stations should continue to be operated tn the area as long as significant 
aftershock activity lasts. The Department should continue to make 

sure that it has portable seismographic units available to move into 
critical areas of suspicious seismic activity in California, as it did 
several weeks prior to the main Oroville event. 

c) DWR should install additional strong -motion accelerographs 
in the vicinity of Oroville Dam. There should be two permanent 
accelerographs on the crest of the dam and one permanent accelerograph 



426 



Page 4 

on each abutment; and three temporary accelerographs should be 
installed in a triangular array in the epicentral region of the August 1975 
earthquakes, replacing the Caltech instruments. All these instruments 
should be equipped with radio time recording. The two AR 240 accelero- 
graphs presently located on the dam and abutment should be removed from 
their present locations, renovated and used elsewhere. 

d) DWR should review its procedures for reacting to the occurrence 
of an earthquake near a dam or other major facility, i. e. , to plan appro- 
priate actions for getting additional strong -motion accelerographs, and 
other instruments, in the field, checking operability of instruments, etc. 

e) The static and dynamic data from instruments in and about the 
dam should be processed and put into completely usable form, and then 
be used together with current, accepted analysis procedures to evaluate 
the dynamic properties of the dam and its materials. 

f) The experience gained frona the Oroville earthquakes in sensing 
and recording significant physical behavior should now be applied to all 
major DWR dams and facilities with a view to improving the collection of 
data. This should include installation of new instruments, improving 
existing instrumentation, and increasing the reliability of the instrument 
systems. 

g) DWR should specially review the seismic instrumentation program 
at Oroville with experts in instrumentation, in recording and processing 
data, making use of the latest knowledge and expertise to improve the system. 

h) A survey program of leveling and triangulation of the dann and 
adjacent area should be completed as soon as possible, consistent with 
accuracy control, etc. In addition, arrangements should be made in colla- 
boration with other appropriate agencies to resurvey a more general area and 
to tie the surveys together. Comparisons should be made with prior data to 
determine if there have been any changes of a differential or of an absolute 
nature in the region or the dam. The survey data from the dam and adjacent 
points should also be correlated with the measurements from movement 
devices in the dam. 



427 



Page 5 



C.R. ALLEN ^^ ^,1^/y^ 
3jk. BLUME 



B.A. BOLT 



/ G.W. HOUSj^ER 
/ 



H.B. SEEDf ^ 



428 



12 September 1975 



Report of the Special Consulting Board for the Oroville Earthquake to: 

Mr. R. B. Robie, Director 
Department of Water Resources 



At a meeting on September 11 and 12, 1975 with the Special Consulting 
Board for the Oroville Earthquake, staff members reviewed information 
relevant to the Oroville earthquake and the performance of Oroville dam 
and its facilities, and described the proposed seismic reevaluation of the 
dams, structures and equipment. At the conclusion of the meeting the 
Board was asked to respond to seven questions. Our responses are pre- 
sented below. 

Question 1 

At its meeting on August 8, 1975, the Consulting Board for 
Earthquake Analysis advised that the appraisals of the regional 
seismicity in the Oroville area be modified and recommended several 
specific actions as a part of that reappraisal. What comments and 
further suggestions does the Special Consulting Board on the Oroville 
Earthquake have with regard to the progress of the Department's im- 
plementation of those recommendations? 

The Department has responded commendably to the actions recom- 
mended on August 8, 1975. A few projects are not yet complete and 
these should be carried forward as quickly as feasible. These include 
particularly the calibration explosion and the detailed geological 
mapping. 

As well as the additional strong-motion instrumentation to be in- 
stalled at the dam and the improved recording capability there, the 
Department should establish water gage stations at several suitable 
positions around the reservoir. The purpose of the gages woiild be to 
determine if regional tilting of the crust \inder the reservoir, perhaps 
related to an impending earthquake, is occurring. 



429 



Report of the Special Consulting Board for the Oroville Earthquake 
page 2 

During the presentations it became apparent that some additional 
administrative attention needs to be given to the line of responsibility 
for continuous maintenance, for emergency operation during earthquakes! 
and modernization of monitoring equipment. In particular, it would 
seem best if the earthquake engineering group had the responsibility 
for ensuring the satisfactory performance of all seismographic instru- 
mentation and analysis. 

Question 2 

What comments does the Board have regarding the performance 
of the dams and other structures? 

The Board ^was presented v/ith extensive oral and written reports 
covering observations of various structures of the Oroville— Thermalito 
complex immediately following the Oroville earthquake, and for many of 
these structures during the interim since that earthquake. All data 
submitted indicate that the related structures performed satisfactorily 
without distress or damage, and as anticipated in the design. 

The Board commends the Department for its prompt inspection of 
project structures following the earthquake. 

Question 3 

What are the Board's views on the identification of the causative 
fault? Is it possible to identify the fatilt beyond the recent epicentral 
area? 

The Board feels that the causative fault zone has already been 
identified with reasonable confidence, although the zone does not neces- 
sarily comprise a single fracture surface. Both the seismological and 
geological studies strongly suggest failure by normal (extensional) 



430 



Report of the Special Consulting Board for the Oroville Earthquake 
page 3 

faulting on a zone trending roughly north, and dipping steeply west. 
In all liklihood, the causative fault zone extends farther to the north 
and south than the segment broken at depth during this series of 
earthquakes, but these extensions have not as yet been positively identi- 
fied. It is important that further work to identify these extensions be 
vigorously pursued- -by detailed geologic mapping, by continued seis- 
mic monitoring, by repeated geodetic surveys across the suspected area, 
and by further trenching of suspicious features. Particular attention 
should be given to understanding the lineaments identified from aerial 
imagery and to searching thoroughly for all possible exposures of faulted 
Quaternary strata. 

Question 4 

What are the Board's recommendations concerning the design 
earthquakes proposed for use in the seismic reanalyses of the Oroville- 
Thermalito structures? 

The Board considers that an appropriate earthquake motion for re- 
evaluation of structures critical to public safety in the Oroville— 
Thermalito complex would be one producing a peak acceleration of 
0.6 g and having characteristics similar to those developed near Pacoima 
dam during the San Fernando earthquake of February 9, 1971. The 
time-history of such a motion should be obtained from a modified form 
of the Pacoima dam record, as discussed in the Report of the Consulting 
Board for Earthquake Analysis dated May 22, 1973. The actual time- 
history could be the same as that forwarded to Mr. Jansen by Clarence 
R. Allen with his letter of January 16, 1974, except that the duration of 
shaking should be limited to the first 20 seconds of the record provided, 
and all ordinates of the record should be multiplied by a suitable scal- 
ing factor to give a peak acceleration of 0.6 g . 



431 



Report of the Special Consulting Board for the Oroville Earthquake 
page 4 

In addition the structures should be checked for the motions pro- 
duced by the following earthquakes: 

(a) a Magnitude 8.5 earthquake occurring at a distance of 100 
miles 

(b) a Magnitude 7.25 earthquake occurring at a distance of 35 
miles 

It is tinlikely that these latter two earthquakes will produce conditions 
more critical than the motion discussed in detail above, but the check] 
should be made to verify that this is so. Design earthquakes for non- 
critical structures can be less severe in intensity than those discussec 
above and the Board w^ill defer this recommendation until the evaluation 
of critical structures is completed. 

Question 5 

What are the Board's comments concerning the proposed progrcim 
for the seismic reanalyses of the Oroville-Thermalito structures? 

The Board concurs with the Department's concept of establishing 
criteria for the relative priority of reassessing the seismic safety of 
the various Oroville-Thermalito structures. It also concurs that those 
structures most critical in terms of public safety should be analysed by 
the best available dynamic methods. Among these structures the 
Board includes both Oroville dam and its Spillway. 

In regard to the Thermalito embankment dams, it is suggested 
that those two or three sections of the Forebay and Afterbay dams 
which appear to be least stable from fo\indation and/or dynamic response 
points of view be selected for detailed reevaluation using dynamic analy- 
ses procedures. "When these studies have been completed other embank- 
ment dams in the Oroville-Thermalito complex might well be reassessed 
by judgment without detailed analysis. 



432 



Report of the Special Consulting Board for the Oroville Earthquake 
page 5 

In regard to the many reinforced concrete structures in the complex, 
only those that can be shown to be critical to public safety would seem 
to justify the use of sophisticated dynamic analysis procedures, but all 
structures evaluated in the original design of the project should be 
checked for adequacy either by judgment procedures or by testing their 
adequacy under increased assximed earthquake loading. 

Question 6 

The Board is requested to provide further explanation and gui- 
dance concerning the Earthquake Analysis Board's recommendation {2d) 
to evaluate stresses in crustal models. 

The Board has no further recommendations at this time. Upon 
completion of the crustal stress analyses no^v being made by others, 
the Departanent may wish to review the question again. 

Question 7 

Does the Board have any other comments or recommendations? 

The Board offers the following suggestions, some of v/hich con- 
stitute reinforcement of procedures discussed previously or in-process, 

(a) The Department should take full advantage of data collected or 
developed by all other agencies, both public and private, con- 
cerned with the Oroville area seismicity. Cooperation with 
such agencies and exchange of data would ensure that all 
reliable data are made available toward the solution of the 
problem. 

(b) The Department should develop a detailed procedure for the 
proposed seismic stability evaluation of Oroville dam embank- 
ment, with particular definition of the steps planned for 



433 



Report of the Special Consulting Board for the Oroville Earthquake 
page 6 

determining the dynamic strength properties of the various 
embankment materials under eonfi-ning pressures of up to 500 
psi. 

(c) The Department should review procedures and contingency plans 
at all dams and major installations for returning equipment and 
facilities to full service after a shutdown due to an earthquake. 
Directives for return of eqmpment to preearthquake operating 
levels should be based upon full knowledge of project conditions 
in order to avoid premature start-up and potential extension 
damage. 

(d) In vie'w of recent press reports concerning the alleged likelihood 

of future large earthquakes near Oroville, the Special Board empha- 
sizes that the hypothetical maximum earthquake of Magnitude 6.5 
mentioned in the Earthquake Board's report of 1 1 August 1975 is 
considered to be a very \mlikely event and is intended to be used 
for safety review. Furthermore, it is our judgment that any 
earthquake significantly stronger than the Magnitude 5«i7 event 
of 1 August 1975 is improbable in the near future. 



434 



Respectfully Submitted, 

J.AA y 




C. R. Allen 




JohA A. Blume 



c pr.< >< ><^t/a?. 



Bruce A. Bolt 
Wallace L. Chadwick 



eorge W. Housner 
T. M. Lejfe 



Alan L. O'Neill 

lilip (3/ Rutledge / 

H. Bolton See'd' 



435 



£tate of California 

Memorandum 



The Resources Agenl| 



To : Mr. Ronald B. Robie 



Robert W. James 

From Department of Water Resources 



Date : OCT 3 1975 

File No.: 

Subject: Oroville Earthquakj 
of August 1975, Proposed 
Department Activities in 
Response to Consulting 
Boards 



As requested by your August 22, 1975 memorandum, presented below 
is a description of our program to implement the recommendations 
made by the Consulting Board for Earthquake Analysis in their 
August 8, 1975 report and by the Special Board for the Oroville 
Earthquake in their September 12, 1975 report. An estimate of 
cost for carrying out these activities is also included. This 
memorandum will also satisfy the requirements of Water Resources 
Eiigineering Memoreindum No. 23 by accounting for the actions taken 
to the Boards' conclusions and recommendations. 

The items below are listed by number as they appear in the Board 
reports. Copies of the reports are attached for easy reference. 

Consulting Board for Earthquake Analysis 



la. 



The level of seismic activity in the Oroville area will be 
reappraised. Accelerograms will be developed for earthquakes 
that are considered to be credible. Included will be a 
local earthquake with Magnitude 6.5 as recommended by the 
Board and a Magnitude 8+ on the San Andreas fault. Finite 
element analyses will be conducted on Oroville Dam using 
one or more of the strongest accelerograms. These analyses 
will be carried out by personnel of the Division of Safety 
of Dams who presently have the expertise. Dr. H. B. Seed 
will be retained throughout the study to provide guidance. 
Additional soils testing, under the direction of Dr. Seed, 
will be conducted at the Richmond Laboratory. The Division 
of Design and Construction will fund the entire study and 
will have overall responsibility for its completion and 
final report. Dr. Seed has been contacted, and he generally 
agrees to this approach. 

The accelerograms will be examined to determine if the seismic 
factors used for the design of other major structures in the 
Oroville-Thermalito Complex are still considered adequate. 
Structures will be reanalyzed as necessary. The manpower 
shown is tentative as it will depend upon early staff findings, 



Staff time 2.5 man years 
Laboratory soil testing 



$ 82,000 

20, 000 

$102,000 



A36 



Mr. Ronald B. Robie -2- OCT 30 1975 



Five survey parties, two from the Division of Operations 
and Maintenance, two from the Division of Land and Right 
of Way, and one crew from the U. S. Geological Survey 
were involved in the work in the Oroville area. Coordination 
was provided by Division of O&M, Chief of Precise Surveys. 

Principal objectives .that have been achieved include: (a) 
vertical suid horizontal surveys of Oroville Dam, (b) level 
rtins over previously established lines to attempt to deter- 
mine location of major faults, (c) a survey of the horizontal 
and vertical control network about Lake Oroville, (d) a survey 
of the epicentral area of the earthquake with a tie to 
established bench marks outside the affected area, (this 
work was done by U.S.G.S.), and (e) surveys of the smaller 
structures in the Oroville-Thermalito Complex. 

The U.S.G.S. survey party operated at its own expense. 

Department survey cost (including 

travel expenses) $ 55,000 

The calibration explosion is being coordinated with U.S.G.S. 
Because CWR equipment could be made available before U.S.G.S. 
equipment, the recommended shot was undertaken by this 
Department. The plan involves one drill hole with the 
explosion at a depth of about 300 feet. Drilling is now 
under way, and is expected to be completed before October 17. 
The hole will be loaded and shot as soon thereafter as 
possible. 

Aside from the shot recommended by the Consulting Board, 
the U.S.G.S. has proposed to cooperate in staging two 
additional shots - one in the Yuba River near Marysville 
and another in one of the northern arms of Lake Oroville. 
These shots would supplement the Department shot and would 
more precisely define the crustal structure in the Oroville 
region. 

Cost of DWR shot: drilling, materials 

staff time $ 40,000 

A search for the fault that caused the earthquake revealed 
a cracked zone along an old fault south of Wyandotte. It 
is believed this is the fault, or one of the faults, related 
to the recent seismic activity. Remote sensing imagery 
consisting of U-2 infrared color, ERTS satellite and side- 
looking airborne Radar (SLAR) imagery were obtained and 
used for a study of lineaments in the epicentral region. 



437 



Mr. Ronald B. Robie -3- OCT 3 1975 



Both vertical and low sun angle aerial photographs were 
obtained in the epicentral area and will be used both for 
geologic mapping of the epicentral region and for detection 
of features that might be faults . 

Geologic mapping is now progressing northward from the zone 
of surface cracking toward project facilities . Discussions 
have been held with Division of Mines and Geology to have 
them do some of the geologic mapping in the epicentral area 
at their cost. Discussions also have been held with PG&E, 
Woodward-Clyde Associates and U.S. Corps of Engineers on 
the problem of obtaining a better understanding of the 
regional tectonic framework of the western Sierra Nevada. 
Objectives of all these activities are: 

1. Identify the causative fault and determine its relation 
to project facilities. 

2. Obtain better knowledge of geology in epicentral area 
in order to decide if Oroville Lake was a contributing 
factor to the earthquake. 

3. Get a better understanding of the regional tectonics 

to better evaluate potential for future seismic activity. 

Cost: Aerial photos, imagery, etc. $ 7,000 
Trenching and other exploration 15,000 
Staff time 1.5 man years 6o,000 

$ 82,000 

2d. Further contact with Board members revealed that computation 
of stresses in crustal models had been undertaken at Cal Tech. 
and U. C. Berkeley. Only preliminary results were available. 
The Board members indicated they had no further recommenda- 
tions pending completion of these analyses. Upon their 
completion your staff will review the results with the Board 
to determine whether or not additional work is desirable. 

Cost: None at this time 

3a. Graphical presentations of various recorded seismic data were 
developed. Noteworthy accelerograms are being digitized. 
Some of this digitized data was ready by the time of the 
Special Board meeting. 

Cost of data preparation $ 3,000 



438 



Mr. Ronald B. Robie -4- 



3b. A permanent seismic station near the epicenter of the 
August 1st shoctc will be established. A site has been 
selected and right of entry acquired. 

Cost: Staff time, planning & design $ 4,000 
Equipment, materials & construction 4,000 

$ 8,000 

Three DWR portable sensitive seismographic units are presently 
installed and operating near Oroville, including one about 
3 miles from the main shock: epicenter. Two portable visual 
recorders are needed for portable units to aid in fast and 
precise determination of epicenters for calibrating a par- 
ticular area for accurate epicenter determination. 

Equipment cost: $ 8,000 

3c. Five SMA-1 strong-motion accelerographs have been ordered 
to replace and augment Oroville Dam strong-motion instru- 
mentation. The instruments will be installed one on each 
abutment, two on the crest and one in the core block. 

Six SMA-1 strong-motion instruments have been ordered to 
provide for emergency situations, such as the Oroville 
earthquake, to augment existing instrumentation on SWP 
structures. Three of these will be installed temporarily 
in a triangular array around the Oroville epi central area. 
In addition to the equipment cost below, there will be some 
additional cost associated with maintenance of the equipment. 

Equipment cost: $21,000 

3d. The Division of Operations and Maintenance's procedure for 
responding to significant or unusual seismic activity 
affecting SWP structures entails augmentation of existing 
instrumentation where needed. A check list of possible 
additional courses of action will be compiled for use in 
future earthquakes including review of instrument maintenance 
practices. 

Cost for augmentation of instrumentation 
at Oroville is covered under Item 3c. 

Cost for additional instruments for 

the remainder of the project $48,000 



439 



Mr. Ronald B. Robie -5- OCT 30 1975 



3e. Processing of the dynamic data is covered under Item 3a. 
Nondynamic data are being plotted on expanded scales for 
clarity and increased functionality. 

The properties of the materials in Oroville Dam that can 
be derived from acceleration and stress data recorded 
during the earthquakes will be evaluated. This work will 
be accomplished in similar manner to that covered under la. 
with funding by D&C, the work accomplished by Division of 
Safety of Dams' personnel, and Dr. Seed utilized in a 
consulting capacity. 

Cost: Staff time, nondynamic data 

processing $ 8,000 

Staff time, stress analyses 40,000 
Contract work. 10, 000 

$58,000 

3f. Improving seismic data collection at other SWP facilities 
is now in progress. New insights gained as a result of 
the Oroville earthquakes will be incorporated in a total 
system reevaluation. 

The costs involved with this item are included under 
Item 3d. or are presently otherwise budgeted for. 

3g. Department personnel will review Oroville instrumentation 

program auid modify to strengthen elements where deficiencies 
may exist. Of particular need is a real time base (WVTVB) 
and a noninterruptable power supply. Replacement of the 
four existing recorders for the dam dynamic instrumentation 
is under review. 

Estimated cost (including recorders) $25,000 

3h. Surveys discussed under this item are included with Item 2a. 

Special Consulting Board for the Oroville Earthquake 

1. The preceding comments on the report of the Consulting Board 
for Earthquake Analysis have generally outlined the program 
and progress on the calibration explosion, geologic mapping, 
and dynamic instrumentation supplementation. 

Additional lake stage recorders in the upper arms are planned 
for other operational purposes. It is believed that these 
recorders will serve the Board's intended purposes; however, 
a thorough evaluation will be made. 



440 



Mr. Ronald B. Robie 



_6- OCT 3 1975 



The Department will review its programs for both the 
maintenance of the instrumentation and for the processes 
for handling and evaluation of records. At the present, 
responsibility for these activities is vested in the 
Project Surveillance program with participation by both 
the Division of Design and Construction and the Division 
of Operations and Maintenance's Earthqualce Engineering 
Section. 

Estimated cost: Included under ongoing programs. 

2. Your staff agrees with the Board's conclusion and thanks 
them for their commendation. 

3. See response to Item 2c of the Consulting Board for Earthquake 
Analysis Report. 

4. The design earthqualce motion suggested by the Board, modified 
February 1971 Pacoima recording, will be used for analyzing 
structures in the Oroville-Thermalito Complex. In addition, 
safety of the structures will be evaluated for the other 
suggested events: Magnitude 8.5 at 100 miles and Magnitude 
7.25 at 35 miles. 

Due consideration will be given to the criticalness of each 
structure within the complex when evaluating the intensity 
of loading to be applied. 

Cost: (Included under other items) 

5. As stated under Question 4, evaluation of the criticalness 
of each structure will be made and appropriate loading 
criteria applied in the resmalyses for seismic safety. 
Oroville Dam and the spillway will, of course, be given 
maximum treatment. Suggested analyses for Thermalito 
Forebay and Afterbay Dams will be accomplished. 

Analyses of Oroville Dam (previously listed) 

Analyses of spillway & other 

structures (previously listed) 

Analyses of Thermalito Dams : 

Staff time $30,000 

Soils testing 40, OOP 

$70,000 

6. This subject is commented upon imder Item 2d of the report 
of Consulting Board for Earthquake Analysis, 



441 



Mr. Ronald B. Robie -7- OCT 3 19/5 



7a. It is our intent to utilize all data developed by others in 
the evaluation of seismic safety of the Oroville-Thermalito 
Complex. Similarly all data developed by the Department will 
be shared with those cooperating in the studies, in preliminary 
form as the studies develop, and in report form upon their 
completion. 

7b. Detailed procedures for analyses of Oroville Dam and the 
necessary soils testing are being developed. 

7c. The emergency plans for dams and the procedures for continuing 
operations of plants, or for return to full operations in 
event of shutdowns due to earthquakes, will be thoroughly 
suialyzed relative to completeness or adequacy of assessments 
of potential damages . 

7d. Your staff agrees with the Board's conclusion. No other 
comments are necessary. 

Estimated cost: Included under ongoing programs or 
under other items above. 

The total estimated cost, as listed above, for implementation 
of the Boards' recommendations is $520,000. For clarity the 
costs are summarized in the table below. 



Item 

Earthquake Analysis Board 

la. Reevaluate seismicity & design criteria 

2a. Surveys 

2b. Calibration explosion 

2c. Mapping 

2d. Crustal models 

3a. Seismic record processing 

3b. Seismic station 

Portable sensitive seismograph units 
3c. Strong motion accelerographs 
3d. Augmentation of instrumentation: 

Oroville 

Remainder of project 
3e. Evaluate properties of dam 
3f. Improve data collection 
3g. Review Oroville instrumentation program 
3h . Surveys 



442 



Cost 


Budgetin 


($1,000) 


Organizat 


102 


D&C 


55 


O&M 


40 


O&M 


82 


D&C 


- 


D&C 


3 


O&M 


8 


O&M 


8 


O&M 


21 


O&M 




O&M 


48 


O&M 


58 


D&C 


25 


O&M 


- 


O&M 



Mr. Ronald B. Robie 



OCT 3 1975 



Item 



Special Board 



Cost 
($1,000) 



Budgeting 
Organizatioi 



1. Implementation of Earthquake Board recommendations 

2. Structure performance 

3. Fault identification 

4. Design earthquake 

5. Seismic reanalyses 70 

6. Crustal models 

7. Other recommendations 



D&C 
D&C 



Cost: O&M Budget 
Cost: DSsC Budget 

Total Cost 



$208* 

312 
$520 



With your approval the program described above, responding to the 
recommendations of the Consulting Board for Earthquake Analysis 
and the Special Board for the Oroville Earthquake, will be 
implemented. 

APPROVED: 



/Director 

K />/7r 



Date 



Attachments 



♦Implementation of 04M related items will require a budget 
augmentation of $187,000. 



443 



23 November 1976 



Report of the Special Consulting Board 
for the August 1, 1975 Oroville Earthquake to: 



Mr. R. B. Robie, Director 
Department of Water Resources 



At a meeting on November 2Z and 23, 1976 with the Special Consulting 
Board for the Oroville Earthquake, DWR staff members reviewed the work 
being done by the Department on the seismic reevaluation of the Oroville 
and Thermalito dams, structures and equipment, as supported by the 
related geological, seismological and surveillance observations accumu- 
lated by DWR and associated agencies. At the conclusion of the meeting 
the Board was asked to respond to six questions. Our responses are 
presented belo'w: 

Question No. 1 . A considerable amount of work has been done along the 
western Sierra Nevada by various groups since the last Board meeting. 
Much of this work has been directed at trying to evaluate future seismicity. 
Has anything developed that would make the Board want to change the 
recommended earthquake motion for reevaluation of Oroville structures 
(report on September 12, 1975 meeting)? 

Response . Since the last naeeting of the Board, substantial investigations 
of past and potential future seismic activity along the western Sierra 
Nevada have been made by the Department of Water Resources, by 
Woodward Clyde Consultants, and by the U. S. Army Corps of Engineers. 
We are not aware that these investigations have produced any information 
to date which would cause the Board to change the earthquake motion it 
recommended in its response to Question No. 4 in the report of 



444 



IZ September, 1975 for seismic re-evaluation of the Oroville-Thermalito 
structures. However, a special supplementary motion, applicable only to 
high frequency structures and facilities, is discussed in answer to Question 
2 below. 

Question 2 . Does the Board have any comments or recommendations con- 
cerning the results or methods used in the seismic re-analysis of the 
critical structures completed to date? 

Response . The Board considers the methods being used thus far in seismic 
re -analysis of critical structures to be appropriate and, in general to 
represent the current state of the art. It is obvious that care is being taken 
to model the structures in a realistic manner and to consider the dynamic 
aspects of the problems at hand. In most cases, final results have not yet 
been obtained in the sense that calculated stresses and strains have not 
been compared with allowable values. This aspect of the work should be 
pursued with vigor. 

The matter of allowable tension in concrete should be resolved to the 
extent practicable at this time, with specific quotations from authoritative 
reference material. The appropriate extent of dynamic water loading of 
the 3D models needs to be resolved for the spillw^ay system. 

The Board recommends that, for critical structures with high funda- 
mental frequencies, the previously recommended time history of earthquake 
motion be supplemented by a time history meeting the high frequency 
(10 Hz or greater) requirements specified by the Nuclear Regulatory 
Commission in its Regulatory Guide No. 1. 60, with the spectrum scaled 
to 0. 4 g at zero period. 



445 



Care should be taken in analyses, and in evaluating the results of 
analyses, not to compound safety factors by using only the most critical 
results or conditions in a sequential fashion. 

Question 3. Does the Board have any conclusions regarding the possible 
relation between Lake Oroville and the Oroville Earthquake sequence? 
If not, does the Board have any recommendations or comments concerning 
gathering of additional data or making further analytical studies to enable 
reaching a conclusion in the future? 

Response . At its meeting on 1 1 August 1975, the Consulting Board for 
Earthquake Analysis Indicated that conclusions regarding any causal 
relation between Lake Oroville and the 1975 Oroville earthquake sequence 
would be difficult to reach. It was suggested that certain observations be 
made that might throw light on the matter. Even with more definitive 
seismological and geological information related to the sequence now 
available, however, it still appears that it is not possible to draw any 
firm inference on whether the earthquakes w^ere, or were not, triggered 
by the reservoir. 

It should be noted that the problem of association between large 
reservoirs and nearby earthquakes is now receiving considerable attention 
w^orldv/ide, and much research on the problem is now under-way in the 
United States and abroad. We recommend that the DWR take steps to 
keep informed of the results of this research, with a view to possible 
application to the Oroville situation and other DWR facilities. 

Question 4. Does the Board have any recommendations or comments 
concerning the draft copy of Bulletin 203 before it is published? 



446 



Response . Bulletin 203 will be useful as a documentation of the performance 
of the dam and related facilities during the August 1, 1975 earthquake and 
aftershock sequence, and as a vehicle for distributing the wealth of seis- 
mological and geological data gathered before and after the earthquake. In 
this regard, the Board believes Bulletin 203 should be limited to include 
only data describing the seismic events, related geological studies, and 
performance of the structures. Although there are numerous minor 
editorial comments which could be made, at this time the Board offers 
only the following specific recommendations: 

(a) A more appropriate title would be "Performance of the Oroville 
Dam and Related Facilities during the August 1, 1975 Earthquake. " 

(b) The purpose should be clearly defined in the beginning of the 
report. 

(c) The final draft deserves further editing to achieve uniform presen- 
tation of the findings and conclusions. 

(d) In reporting the factual observations and events, care should be 
taken to avoid the inference that the Department has made a 
definite conclusion regarding the relationship or lack of relationship 
of the reservoir to the earthquakes. 

(e) It is requested that the listing of the Board members on an 
introductory page of Bulletin 203 be deleted, inasmuch as the 
Board has not participated in preparing the report. It Is similarly 
recommended that the reports of the Board included in Appendix D 
be deleted. 

(f) The re-evaluation earthquake studies, recommended previously 
by the Board, apparently will not be completed before mid-1977. 
Hence, any conclusions and recommendations relating to such 



447 



studies would be premature at this time. It therefore would appear 
appropriate to issue a separate bulletin or report on this phase 
of the work in late 1977, as a follow-up to Bulletin 203. 

K the foregoing concept is adopted, it would seem desirable 
that Bulletin 203 include a specific list of all damages to the 
Oroville complex resulting from the Oroville Earthquake, together 
with a notation of the type and cost of repair work completed. 

Question 5. Does the Board have any recommendations for future geologic 
work? 

Response . The Board emphasizes the value of determining and attempting 
to understand the growth of surface faulting following the Oroville Earth- 
quake, and it urges that this work continue to be pursued vigorously. On a 
broader scale, it is important, to the long-term safety of the Oroville 
Project, that the geologic environment associated with the Oroville earth- 
quake be understood as well as is realistically possible. Two important 
questions are (1) what is the relationship of the surface faulting associated 
with the 1975 earthquake to the mapped surface geology, and (2) what can be 
said about possible future northw^ard extensions of the 1975 fault break? 
Answers to these questions will undoubtedly require additional trenching 
and additional detailed geologic mapping. Including areas north of the lake. 
Continuing efforts should be miade to relate local geology to geodetically 
observed deformation patterns. To do this effectively, the area must be 
re-surveyed for elevation changes at regular intervals, preferably semi- 
annually, for the next several years. 



448 



Question 6. Does the Board have any other comments or recommendations 
to make at this time? 

Response. The Board offers the foUowing comments and recommendations: 
(a) The Board would like to draw DWR's attention to the small but 

finite, likelihood of a future recurrence of an earthquake sequence 
similar to that of 1975 near to Oroville Dam and its associated 
facilities. Somewhat analogous seismological and geological con- 
ditions in other parts of the world make it not implausible that 
a possible repetition of the sequence may occur northward from the 
1975 events. Indications, if any, of the above development should 
be sought in future seismological, geological and geodetic 
monitoring, 
(b) The Board believes that there is urgency to complete the re-analyses 
of all of the dam elements in the Oroville- The rmalito complex at an 
early date, in order to determine whether any reinforcement may 
be required to assure ability of those structures to resist the effects 
of a 6. 5 magnitude local earthquake, 
(c) The surveillance attention being given to the project is commendable. 
The surveillance provides early detection of damage but time in 
which to mobilize effectively for major emergency repairs required 
by seismic damage to embankments would probably not be available. 
Therefore, inherent structural integrity must be the alternative. 
In particular, the most critical portions appear, at this time, to be 
some locations along the Thermalito Forebay and Afterbay Dams. 
Accordingly, it is recommended that locations of critical sections 
of these dams be determined on the basis of the existance of 



449 



low-density soils, particularly loose sands, in the foundations. 
Field sub- surface explorations, followed by analyses of these 
sections under the effects of the "re -evaluation earthquake, " 
should be carried out on an urgent basis and, where potential 
instability may be indicated, corrective designs should be 
developed and the construction accomplished as soon as possible. 



450 



J 



Respectfully Submitted, 



C. R. Allen 




John .W. Blume 



Bruce A. Bolt 



ffh-i£0u^yX\ y 



Wallace L. Chadwick 





eorg'e W. Housner 
T. M. Leps T 
Alan L. O'Neill 



'hi lip dj Rut: 



X^JuUjj-i^X. 



b 




Philip (\J Rut ledge 
H. Bolton Seed 



451 



Memorandum 



The Resources Agel 



To : Ronald B. Robie 



Dote : MAR 4 1977 



Robert W. James 
From Department of Water Resources 



File No.: 

Subject: Oroville Earthquat 
of August 1975, Proposed 
Department Activities in 
Response to the Special 
Consulting Board Meeting 
November 22 and 23, I976 



Presented below is a description of our program to implement the 
recommendations made by the Special Board for the Oroville Earthquake] 
of August 1975, in their report transmitted to us by letter dated ' 
December I5, 1976, for the meeting held November 22 and 23, I976. 
The recommendations generally concern completing analyses and work, 
initiated as a result of their recommendations in reports dated 
August 8, 1975 and September 12, 1975^ and outlined in my memorandum 
to you dated October 30, 1975. Our response to each item in the 
Board's latest report is listed below by number as they appear in 
their report. A copy of their report is attached for reference: 

1. No changes are required in the earthquake motions that are 
being used in our reevaluation of Oroville critical Structures 
except for high frequency structures and facilities. Our 
action concerning these structures is covered under Question 2. 

2. The staff agrees with the Board's comments that final results 
of the dynamic analysis be pursued with vigor and the final 
results be compared with allowable values. It is intended to 
proceed as rapidly as possible with the analysis. A funding 
augmentation of $300, 000 has been approved for the remainder 
of this fiscal year and it is now estimated that we will need 
$166,000 for 1977-78 fiscal year. The additional funds were 
needed because the scope of the investigation was expanded. 

The problem of allowable tension in concrete for dynamic or 
transient loads has been given considerable study in recent 
years as recent dynamic analyses of concrete dams have 
indicated larger tensile stresses than earlier design 
procedures. It is intended that we will determine what 
allowable tensile stresses can be used by a search of 
authoritative reference material to support our contention 
that dynamic tensile stresses indicated by the analysis 
are satisfactory. We will be investigating the extent 
of the dynamic water loading of the 3D models for the spillway 
system. 



452 



Ronald B. Robie 
Page 2 



Both the diversion dam and the spillway have fairly high 
fundamental frequencies, therefore we will investigate the 
structures for the higher frequency ground motions as 
recommended by the Board. 

We have attempted to evaluate the performance of the 
structures realistically and not compound safety factors. 
We have initially evaluated the structures conservatively 
and refined the analysis if the performaxice appeared to be 
questionable. We will continue to examine our results with 
this in mind. 

3. We plan to Iceep informed of the results of research in the 
association between large reservoirs and nearby earthquakes 
by studying written material as it is published and by 
observing performance of structures in California. Of 
particular interest will be New Melones Dam on its initial 
filling. At the present time no additional funds are needed. 

4. The report will be rewritten to include only data describing 
the seismic events, related geological studies, and performance 
of the structures. Conclusions from these studies will be in 
the final report after all studies are complete. 

(a) The title has been changed as suggested to "Performance 
of the Oroville Dam and Related Facilities during the 
August 1, 1975 Earthquake." 

(b) A statement on the purpose has been added to the foreword. 

(c) The Report Administration Section in the Division of 
Planning has edited the bulletin to achieve uniform 
presentation of the findings and conclusions. 

(d) The section discussing the relationship of the reservoir 

to the earthquake has been rewritten to avoid the inference, 
The conclusion on this subject will be in the final report. 

(e) The listing of the Board members has been deleted from the 
introductory page. The Board members are still listed in 
the text where it is discussing that a Special Board had 
been established. The Board's reports have been deleted 
from Bulletin 203. 

(f) The final report will be a "follow-up bulletin" and 
include the conclusions from the many studies now in 
progress, and the Board's reports. 

A section is being prepared to list the damages to the Oroville 
Project Facilities including the type and cost of repair work. 



453 



Ronald B. Robie 
Page 3 



5. We concur with the Board's recommendation that an understanding 
of the growth of surface faulting should be pursued vigorously 
and that it is important to the long-term safety of the Oroville 
project to understand the geologic environment associated with 
the Oroville earthqualce. The Project Geology Section has 
developed the following program for geologic investigation to 
comply with the recommendations of the Special Consulting Board 
for the Oroville earthquake: 

(a) Determine the extent of the fault thought to be responsible 
for the Oroville earthquake. It is particularly important 
to determine where the northern extension of the fault is 
in relation to the Oroville facilities. 

(b) Verify the nature, age of last movement, euid extent of the 
two faults previously mapped by others just west of Orovill 
Dam. 

(c) Do geologic mapping in the Palermo and Bangor quadrangles. 
Also do geologic mapping north of Lake Oroville. Do 
geologic mapping of Tertiary formations in the vicinity 
of the O&M Headquarters . 

(d) Investigate the Palermo Crack Zone-Prairie Creek lineament 1 
to see if it is a fault system that could pose a hazard ! 
to project facilities. Continue investigation "of the 
Paynes Peak, Swain Ravine, and Prairie Creek lineaments, 
and other suspicious lineaments, both north and south of 
Lake Oroville. This will involve extensive field studies 
including both geologic mapping and trenching. 

(e) Continue study of ground water levels in the epicentral 
area to determine interrelationship of local ground water 
systems with Lake Oroville. 

Target date for completion of the above program is July, 1978. In 
order to meet that target date, we anticipate that four DWR geologits 
will work full time on the program. Additional temporary assistance 
may be required to do some geologic mapping. We anticipate the 
additional assistance required for geologic mapping possibly might 
be done by graduate students during the summer of 1977 and possibly 
the summer of 1978, but this kind of arrangement has not been exploit 
yet with the universities . 

Estimated cost of the geological program is $l43,000 for 1976-77 anc 
$334,000 for 1977-78 fiscal years. It will be necessary to hire twc 
additional Junior Engineering Geologists to carry out the program. 



454 



Ronald B. Robie 
Page 4 



The Division of Operations and Maintenance plans to resurvey 
this area again during the summer of 1977 with its precise survey- 
crews plus a maps and survey crew if one is available to determine 
deformation patterns. Estimated cost for the resurvey is $8o,000. 

6. (a) We concur that we should be prepared for additional 
seismic events in this area. Monitoring of seismic 
activity in the Oroville area will continue under the 
Division of Operations and Maintenance Earthquake 
Engineering Program. Our current plans for geological 
and geodetic monitoring are covered under our response 
to Question 5. 

(b) We plan to have the re-analyses of all the dam elements 
in the Oroville-Thermalito complex completed next fiscal 
year, 1977-78. 

(c) We concur that the structural integrity of Thermalito 
Forebay ajid Afterbay Dams under severe earthqualce loading 
is uncertain. We are in the process of evaluating the 
stability under the recommended loading and expect to 
have these completed next fiscal year 1977-78. 

Attachment 

cc: H. H. Eastin 
G. W. Dulcleth 
J. W. Marlette 



455 



APPENDIX B 
ACCELERATION TIME HISTORIES AND RESPONSE 
SPECTRA FOR THE AUGUST 1, 1975 AND 
SEPTEMBER 27, 1975 RECORDED MOTIONS 
ON DAM CREST AND BEDROCK, IN UPSTREAM-DOWNSTREAM DIRECTION 
(FIGS. B-1 THROUGH B-8) 



457 




458 



fldCEliERflTIbN iRe^P 5PEt- 

:LLE ^DRM JCREST N4BE 8-1-15 USGS ! 

:ng + zL B^i.ioxUozi I : \ 




Figure B-2. Computed Acceleration Response Spectra for USGS August 1, 1975 
Crest Motion 



459 




Figure B-3. USGS August 1, 1975 Recorded Rock Motion 



460 



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1 frHtt-Hi riTl- pemOTl imoNoii [ITlHi-[frihF[TH'HB;hHfH:n 



Figure B-k. Computed Acceleration Response Spectra for USGS August I, 1975 
Pock Motion 



461 




462 




Figure B-6. Computed Acceleration Response Spectra for DWR September 27, 1975 
Crest Motion 



463 




464 




Figure B-8. Computed Acceleration Response Spectra for DWR September 27, 1975 
Base Motion 



465 



APPENDIX C 

STATIC STRESSES FROM STATIC FINITE ELEMENT ANALYSIS 
(FIGS. C-1 THROUGH C-8) 



467 



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469 



0;ROVILLE DAM -MAXIMUM SECTION 




470 




Figure C-3. Major Principal Stresses, O] (tsf) 



471 



4 



OROVILLE DAM - MAXIMUM SECTION 




472 




Figure C-4. Minor Principal Stresses, o^ (tsf) 



473 



oroviiIle dam -maximum SECTIOK 



I 

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475 



OROVILLE DAM - MAXIMUM SECTION 




476 




Figure C-6. Horizontal Normal Stresses, a^ (tsf) 



477 



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OROVILLE DAM - MAXIMUM SECTION 




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Figure C-7. Maximum Shear Stresses, T^ax (*^sf) 



479 



OROVILLE DAM -MAXIMUM SECTION 




480 





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Figure C-8. Horizontal Shear Stresses, x^^y (tsf) 



481 



APPENDIX D 
TIME HISTORIES AND RESPONSE SPECTRA 
FOR REANALYSIS EARTHQUAKE 
(FIGS. D-1 THROUGH D-6) 



483 





Figure D-1 Accelerogram 
for Reanalysis Earthquake 



Figure D-2. Computed Velocity Time Histor 
for Reanalysis Earthquake 




Figure D-3. Computed Displacement Time 
History for Reanalysis Earthquake 



484 




Figure D-^. Computed Acceleration Response Spectra for Reanalysis Earthquake 



485 




i 



Figure D-5. Computed Velocity Response Spectra for Reanalysis Earthquake 



486 




Figure D-6. Computed Displacement Response Spectra for Reanalysis Earthquake 



487 



APPENDIX E 

RESULTS OF DYNAMIC FINITE ELEMENT ANALYSES 

FOR RE ANALYSIS EARTHQUAKE 

MAXIMUM SECTION ~ SHELL K- , = 350, 200, 130 

2 max 

ELEMENT STRESSES AND STRAINS CFIGS. E-1 THROUGH E-18) 

SHEAR STRESS TIME HISTORIES (FIGS. E-19 THROUGH E-39) 

SECTION 2 ~ SHELL K- = 130 - LUSH AND QUAD-4 

2 max 

ELEMENT SHEAR STRESSES AND STRAINS (FIGS. E-40 

THROUGH E-45) 

ACCELERATION TIME HISTORIES (FIG. E-46) 

SECTION 3 — SHELL K„ = 130 - LUSH AND QUAD-4 

2 max 

ELEMENT SHEAR STRESSES AND STRAINS (FIGS. E-4 7 

THROUGH E-52) 

ACCELERATION TIME HISTORIES (FIG. E-5 3) 

MODEL EMBANKMENT — SHELL K„ =130 

2 max 

EFFECT OF POISSON'S RATIO ON STRESSES (FIGS. E-54 
THROUGH E-55) 



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OROVILLE DAM - MAXIMUM SECTION PROGRAM = QUAD 4 

K2MAX = 130 LOWER CORE MODULUS 


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OROVILLE DAM - MAXIMUM SECTION PROGRAM = QUAD 4 

K2MAX = 130 LOWER CORE MODULUS 
REANALYSIS EARTHQUAKE-AAAXIMUM ACCELERATION = 0.6g 


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NODE 4 



0.50 




4.0 8.0 12 16.0 

TIME IN SECOND 
LUSH RESPONSE ANALYSIS 



20.0 




-0.50 



4.0 8.0 12.0 16.0 

TIME IN SECOND 
QUAD 4 RESPONSE ANALYSIS 



20.0 




Figure E-'tS. Comparisons of Acceleration Responses between LUSH and (iLIAD4 

Analyses of Section 2 Using K„ =130 and Low Core Modulus 

^ 2 max 



525 



OROVILLE DAM - SECTION 3 

REANALYSIS EARTHQUAKE - MAXIMUM ACCELERATION 

LUSH DYNAMIC RESPONSE ANALYSIS 

SHELL K o = 130 



0.6 g 



CORE 



2 max 
G max -2200 (LOW CORE MODULUS) 




Figure E-^?. Maximum Horizontal Shear Stresses, x , from LUSH Analysis of 
Section 3 (tsf) ^^ 



OROVILLE DAM -SECTION 3 

REANALYSIS EARTHQUAKE - MAXIMUM ACCELERATION = 0.6 g 

LUSH DYNAMIC RESPONSE ANALYSIS 

SHELL K « = 130 



CORE 



2 mox 
G max 



2200 (LOW CORE MODULUS) 




Figure E-48. Maximum Shear Strains, Ymax, from LUSH Analysis of Section 3 

Using K„ =130 and Low Core Modulus (percent) 
i- max "^ 



526 





rf- 


T 



0) 3 

— -a 
«- o 
o 2: 

— L. 
DC O 



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10 


c*\ 






S- 




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II 


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Q. cn 
>- c 
I- — 



527 



OROVILLE DAM- SECTION 3 

REANALYSIS E ARTH OU AK E - M AX I MU M ACC ELER ATION = 0.6g 

QUAD 4 DYNAMIC RESPONSE ANALYSIS 

SHELL K2 mox = '30 



CORE 



2200 (LOW CORE MODULUS) 









St^.^M%^ 


"^ 






^^.19i^65i.95\1.08\. 
^^1.35 /1.55ib.90\ 1.91 "~~v^ 




j,^r,k 1.77 11.97/1.12 \ i.ee i|.>,.. 










..-^1.85 


1.62 12.51 J 1.21 2.07 


2.13--^ 




^-^.B& 


1.87 


2.14 ^.Hl/l. 103; 1 2.21 


2.51 "^ 




-<rT5 


1.37 


2.66 


2.73 *i5)/ 1.11 l/l 2.90 


J-58 /."SJ^ 


.^ 


1.16 


2.39 


2.80 


1.27/ 1.S5 /s8 


3.19 


3.21 


••96 ^^ 


^O;,i^0.98 1 


1.* 


2.10 


li-Vi yi.zx / i.ei^^fj^.n 


3.98 


3.11 


2.60 [ 1.39^^ 



Maximum Horizontal Shear Stresses, x , from QUAD^ Analysis of 
Section 3 Using K^ ^3^ = '30 and Low Core Modulus^ 



Figure E-50. 



OROVILLE DAM - SECTION 3 

REANALYSIS E ARTHQU AK E - M AXI MUM ACC ELE R AT ION = 0.6 g 

QUAD 4 DYNAMIC RESPONSE ANALYSIS 

SHELL K2 ^Q, =130 

G max 



CORE 



-- 2200 (LOW CORE MODULUS) 




Figure E-5I . Maximum Strains, y . from QUAD^ Analysis of Section 3 

Usinq K., =130 and Low Core Modulus 
^ 2 max 



528 







529 



NODE 4 



0.50 



-0.50 



O50 



0.50 



-0.50 



0.50 



-0.50 




tililiiiiliyiliiiiJil 

4.0 8.0 12.0 16.0 20.0 

TIME IN SECOND 
LUSH RESPONSE ANALYSIS 



4.0 8.0 12.0 16.0 20 

TIME IN SECOND 
QUAD 4 RESPONSE ANALYSIS 




Figure E-53. Comparisons of Acceleration Responses Between LUSH and QUAD^ 
Analyses of Section 3 Using K- =130 and Low Cora Modulus 



530 





V 


/1.3S 


-^ 


0.13 


1.21 


0.99\ 


o.w\ 


cry 


( V -- 


0.49/0,30 




o-y (V=0.30) 




0.T6 1 


0.S7 


8.09 


0.99 




0.60 


o.so 


O.SO 


2.3S 


l.OB 


0.99 


i.^\ 






X 


o.n 


a.9^ 


0.63 1 


0.19 


8.09 


1.11 


1.00 


l.Ok 


1.0l\^ 






^ 


1.18 


0.9? 


O.M 


0.91 1 


O.T 


J. 01 


I. IS 


1.01 


1.03 


1.01 


i.ooN^ 




y^.°D 


1.31 


0.91 


o.ex 


0.79 


i.ie yo.33 


i.as 


i.n 


l.OS 


1.03 


1.01 


0.97 


o.»\ 



MAXIMUM VERTICAL NORMAL STRESS (cTy ) 





yX.\Z 


y^^ 


V 


/1.25 




0.80 


0.85 


0.90\ 


0.95\ 


Ux 


I 1/ - 


u.ts 


/ U. 




^.5« 


(T, (V 


= 0.30 ) 




0.61 


0.67 


0.39 


0.90 




0.70 


0.68 


0.87 


0.93 


0.98 


0.94 


i.oiN^ 








o.as 


1.09 


1.17 


0.V8 


O.Sl 


0.94 


0.98 


1.00 


i.rox^ 






1.13 


1.86 


1.39 


1.S8 


O.SU 


0.71 


l.OM 


i.oe 


1.01 


1.01 


i.»\ 




yLa 


1. 67 


l.Sl 


1.63 


1.86 


8.« 1 


0.119 


1.06 


1.17 


1.07 


1.03 


l.Ol 


1.03 


sk^ 



MAXIMUM HORIZONTAL NORMAL STRESS ( cr^ ) 



Txy (V= 0.49/0.30) 




MAXIMUM HORIZONTAL SHEAR STRESS (T^y ) 



LUSH DYNAMIC RESPONSE ANALYSIS 

REANALYSIS EARTHQUAKE — MAXIMUM ACCELERATION =0.6g 



EMBANKMENT K 



2 MAX 



130 



Figure E-S'*. Effect of Polsson's Ratio on the Induced Dynamic Stresses in 
the Model Embankment 



531 





V 


y\.\a 


^X^ 


0.79 


o.w 


o.g7\^ 


1.06\ 


CTy 


(1/ 


= 0.45) 




0.92 


0.8E 


0.61 


0.76 


0-y (1/ 


-- 0.30) 




0.8S 


0.83 1 0.91 


0.56 


0.78 


0.89 


l.lli\^ 








^ 


0.94 


0.82 


0.87 


0.9* 


0.S3 


0.82 


0.86 


1.03 


1.13\^ 






yO& 


1.05 


0.86 


0.83 


0.91 


0.* 


9. 47 


0.90 


0.89 


0.94 


1.11 


i.oi\ 




y<yB3 


1. 11 


0.9S 


0.87 


o.et 


0.98 1.03 

1 1 


O.W 


0.99 


0.9J 


0.93 


0.9S 


1.20 


0.93^ 



MAXIMUM VERTICAL NORMAL STRESS (cTy) 




MAXIMUM HORIZONTAL NORMAL STRESS (crj 





V 




/ 


1.02 


0.98 


o.*\ 


0.37\^ 


r^y (T -- 0,45) 




y^Oi 


ya.in 


/^.99 




1.03 1 


1.06 


l.OB 


1.05 


Txy (1/ = 0.30) 




l.tN 


1.06 1 


1.11 


l.Il 


1.10 


1.08 


i.re\ 






yyi.'S 


1.08 


1.06 


l.OE 1 


1.13 


1.12 


1. 10 


1.07 


1.09 


1.12~\ 






yX.aa 


t.ll 


1.09 


1.06 


1.00 1 


1.02 


l.CK 


l.CK 


1.07 


1.09 


l.W 


l.lii\^ 




/^tCoe 


1.01 


1.01 


1. 00 


0.97 


0.92 


0.93 


0.93 


0.93 


0.95 


1.01 


1.03 


l.CH 


i.iJN.. 



MAXIMUM HORIZONTAL SHEAR STRESS (T*„) 



LUSH DYNAMIC RESPONSE ANALYSIS 

REANALYSIS EARTHQUAKE — MAXIMUM ACCELERATION =0.6g 



EMBANKMENT K 



2 MAX 



= 130 



Figure E-55. Effect of Poisson's Ration on the Induced Dynamic Stresses in 
the Model Embankment. 



532 



APPENDIX F 
EMBANKMENT RESPONSE MODEL 



533 



APPENDIX F 
EMBANKMENT RESPONSE MODEL 

As described previously in the main text, the 

Embankment Response Model is a two— dimensional plane strain 

analysis with a modified K„ (pseudo K„ ) value. This 

2 max '^ 2 max 

model was developed to account for the three— dimensional effect 
of the canyon on the dynamic response of the embankment. 

Oroville Dam is located in a triangular-shaped 
canyon and has a variable cross-section (Figure F— l). In a 
two— dimensional plane strain analysis, the length of the dam 
(z— axis) is assumed to be infinite and all of the stresses 
induced to resist movement are in the x-y plane (Figure F-2). 
However, the abutments impart a restraining effect which gives 
additional stiffness to the embankment. This additional stiffness 
results from stresses in the y— z and x— z planes. Stresses in 
these two planes are not accounted for in a two— dimensional plane 
strain analysis. 

In an attempt to simulate three— dimensional response, 

an artifically-high (pseudo) K„ was used to account for the 

'^ 2 max 

stresses in the y-z and x-z planes. As detailed in Section 5, 

a value of 350 was developed for the pseudo K„ value. This 

^ ^2 max 

value was determined from analyses of embankment response to the 
1975 Oroville Earthquakes. In extending this model for use with 
the Reanalysis Earthquake, it is assumed that the model can 
simulate three— dimensional embankment response to earthquakes 
of varying magnitude and frequency content. 



535 




Figure F-1. Three-dimensional Problem 




Figure F-2. Two-dimensional Plane Strain Representation 



536 



In applying the model, it is assiojtned that the model 

will simulate the three-dimensional response of the maximiim 

section of the dam in the x-y plane. This means that the 

accelerations and displacements would be approximated. However, 

the shear stresses in the x— y plane would not be correct. This 

is due to the fact that all of the dynamic stresses have essentially 

been Itmiped together into the x-y plane by using the two-dimensional 

plane strain analysis with a pseudo K„ value of 350. Since the 

2 max 

earthquake-induced shear stresses are of considerable importance 

in a dynamic analysis of an embankment, the stresses must be 

estimated in a different manner. 

The method which was adopted to estimate the shear 

stresses in the x-y plane resulting from a three— dimensional 

embankment response assumed that the Embankment Response Model 

approximated the correct shear strains in the x-y plane. Using 

these shear strains, and the best estimate of the actual K„ 

2 max 

value for the Oroville gravel, the horizontal shear stresses in 
the x-y plane were estimated. 

The procedure is detailed in the following equations: 

t = Y * G (1) 

where ^ = Shear Stress 

^ = Shear Strain 

G = Shear Modulus 

G = R^ * K^ * ( Q' ')^'^ * 1000 (2) 

d 2 max m 

where R = Shear Modulus Reduction Factor, Dependent 
Upon Shear Strain 

^T = Effective Mean Normal Stress in psf 

1' = ^ * R , * K„ * ( T )'^ * 1000 (3) 

d 2 max m 

537 



Usine the Embankment Response Model and the pseudo K„ 

^ ^ 2 max 

value of 350: 



^/^^ 



= y * R^ * 350 * (CT')" * 1000 (k) 
350 ^''350 



The actual shear stress in the x— y plane induced in 

a three— dimensional response would be computed using the actual 

value of K„ for the gravel. As discussed previously, the 
2 max ° f J ■: 

best estimate of the actual K„ value is about l65. A 

2 max 

comparison of the actual and pseudo K values plotted against 
shear strain is presented in Figure F— 3« The actual shear 
stress in the x— y plane would be defined by: 

Z = ^' ^ R^ * 165 * ( Cr')'^ * 1000 (5) 

^^30 ^^30 ^ 

Assuming that the actual (3D) shear strains in the x-y 
plane are approximated by the strains produced in the Embankment 
Response Model, 

y = ^ (6) 

''''3D =^20^3^ 

^=^3D ' "-V2D "^ *'''*' ^-'^ * "'°° *" 

Because the shear modulus reduction factor (R,) and the initial 

d 

effective mean normal stress ( ) are the same, equations k 

m 

and 7 may be combined to yield: 

t = i65 * X (8) 

^^30 350 ^YsD 

350 



538 



I 



This approach employs many assiimptions and has inherent 
limitations. However, it is considered to model the actual 
embankment response more accurately than the traditional plane 
strain analysis using actual material properties. 



539 




540 



APPENDIX G 

CYCLIC TRIAXIAL TEST SUMMARIES 

OF MODELED OROVILLE GRAVEL TESTS 

(FIGS. G-1 THROUGH G-6 8) 



541 



APPENDIX G 
CYCLIC TEST SUMMARIES 

In order to present the test behavior of the cyclic 
triaxial tests carried out for the modeled Oroville gravel, 
cyclic test summaries were prepared. These summaries show the 
peak values of cyclic deviator stress, pore water pressure 
increase, and axial strain, plotted against cyclic number. 

The test summaries are derived from the cyclic test 
records and show uncorrected test behavior. Before utilizing 
this information, corrections for membrane compliance, calibra- 
tion error, membrane strength, and consolidation conditions (C ) 
should be applied. 

Cyclic deviator stress peaks in the extension direction 
are considered negative and are labeled so. Cyclic deviator 
stress peaks in the compression direction are considered positive, 
The sign convention for axial strain is also defined as having 
compression being the positive direction. The strain peak 
envelopes, however, are labeled with either "extensive" or 
"compression" to identify the direction of the stress pulse 
when the strain peak occurred. 

The peak values of pore pressure increase were plotted 
by using the back pressure value as a zero point. A value above 
the back pressure was denoted as positive and a value below the 
back pressure was denoted as negative. Also shown in the pore 
water pressure summaries as a horizontal dashed line is the 
initial effective confining pressure. Pore pressure envelopes 



543 



rising above this line show either incorrect calibration or 
a change in the triaxial cell pressure. 

These summaries are only intended to illustrate the 
general behavior of the samples during testing. For a more 
detailed examination, the actual test records should be 
consulted. 



544 



-15000 



-10000- 



-5000- 



> b" 

a 
o 



5000- 



o 

u 10000 



15000 



EXTENSION 



COMPRESSION 



-'> 



5000 



(/) 
(/) 

u 

QC 

a. ^ 

• 

< 

UJ 
QC 
O 

a. 



5000 



y^ 



v^' 



-'> 



-5 



-'-^ 



EXTENSION 

COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



40 



-'-' 



80 



85 



Figure G-1 . Cyclic Test Envelopes for Test No, 6 (a' = 4,100 psf, K^ = 1.0) 

545 



15000 



-10000 



5000- 



5000 



G toooo 



15000 



EXTENSION 



COMPRESSION 



I 



^y 



5000 



5000 



._/ 



-rJ^ 



>.^/' 



-r-r 



-15 



-10- 



-5 



EXTENSION 



COMPRESSION 



20 30 

NUMBER OF CYCLES 



h- 



^> 



40 7 5 80 



Figure G-2. Cyclic Test Envelopes for Test No. 7 (a', = '^.lOO psf, K = 1.0) 

jC c 



546 



-ISOOOr 



-lOOOOh 

I 
I 



5000- 



< 

> b' 

Q 
U 



01- 



5000 1-^ 



10000- 



15000 



EXTENSION 



> / 
I / 

COMPRESSION 



•t 



-r> 



50OOf 



5000 




-15 



-lOl- 



<< 



10 



EXTENSION 



COMPRESSION 



^y 



10 



20 30 

NUMBER OF CYCLES 



40 



60 



65 



Figure G-3. Cyclic Test Envelopes for Test No. 10 (a' = 4,100 psf, K = 1.0) 

jQ C 



547 



-ISOOOr 



-10000- > 



5000- 



5000 



10000 



19000 



EXTENSION 



COMPRESSION 



5000 



a. 



5000 









1 


s 


\ 
\ 


1 




\ 


/ 






■H 






X LOWER 


PORE PRESSURE ENVELOPE BEYOND RECORDING LIMITS 



-5- 



EXTENSION 



COMPRESSION 



5 - 



20 30 

NUMBER OF CYCLES 



40 



Figure G-^. Cyclic Test Envelopes for Test No, 11 (a'^^ = 4,100 psf, K^ = l.O) 



5A8 



-ISOOOr 



-toooo- 



-5000- 



O S 






9000- 



o 

U 10000 



15000 



EXTENSION 



COMPRESSION 



50O0 




5000 



-15 



-10- 



JS 



- 












y 


/ 




^ 


" 


- 


- 


EXTENSION 


y 




N 


- 


- 


- 


- 


"compression 
















1 1 




1 I— 



K) 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-5. Cyclic Test Envelopes for Test No. 12 (o' = ^4,100 psf, K = l.O) 

jC c 



549 



-ISOOOr 



-100O0- 



5000- 



> b' 

Q 

U 

_l 

o 

>- 
u 



5000 



10000 



15000 



EXTENSION 



COMPRESSION 



5000 



-5000 




a5 



-15 



-10 



-5 



5- 



10 



EXTENSION 



\ 



COMPRESSION 

U 



*COMPRESSION STRAIN ENVELOPE 
BEYOND RECORDING LIMITS 



10 



20 50 

NUMBER OF CYCLES 



40 



50 



Figure G-6. Cyclic Test Envelopes for Test No. 13 (°', = ^,100 psf, K = 1.5) 

jO c 



550 



-15000 



^ -10000^ 

Vi 



5000- 



q: ^ 

< a 
> b' 
Q 

<J 

_l 
O 



01- 



soooi- 



10000- 



15000 



EXTENSION 



COMPRESSION 



5000 



-5000 




W o^ 



-15 



-10 



-5 



10 



EXTENSION 



COMPRESSION ~~ — — _^ 
__l I 111= L 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-7. Cyclic Test Envelopes for Test No. 14 i°' ^^ = 4,100 psf, K^ = 1.5) 



551 



-ISOOOr 



-10000- 



5000- 



Ui 

o 
u 



5000 



u 

u 10000 



15000 



EXTENSION 



COMPRESSION 



5000 



(A 
M 

UJ 
(T 

a. ^ 



-5000 




< 



-15 



-10- 



5- 



10 



."Zl-n.-:--- EXTENSION 



COMPRESSION 



10 



20 K> 

MUMKR OF CYCLES 



40 



50 



Figure G-8. Cyclic Test Envelopes for Test No. 15 (o' = 4,100 psf, K = 2.0) 

ic c 



552 



-ISOOOr 



-10000- 



-5000- 



> b' 

UJ 

o 
u 



5000 



10000 



15000 



EXTENSION 



COMPRESSION 



5000 



5000 







.'■"N 






/ 

/ 
1 




* 




* LOWER PORE PRESSURE ENVELOPE BEYOND RECORDING LIMITS 





< 
ct 

CO »^ 



-10- 



-5 



5- 



IQ 



EXTENSION 

y 



COMPRESSION 



10 



20 30 

^4UMBER OF CYCLES 



40 



50 



Figure G-9. Cyclic Test Envelopes for Test No. I6 (cf'^^ = 4,100 psf, K^. = l.O) 



553 



-ISOOOr 



-10000- 



-5000- 



> b" 

u 
o 

o 

-I 

>- 
o 



50O0- 



10000- 



15000 



EXTENSION 



COMPRESSION 



-N' 



5000 










a 


^ 


.^ 



-5000 


^^'^ 






-'i- 









-15 



-10 



-5 



5- 



10 



__EXTE_NSION 
COMPRESSION 



K) 



20 30 

NUMBER OF CYCLES 



40j 



-'-' 



50 55 



Figure G-10. Cyclic Test Envelopes for Test No. 17 ( «^ , = 4,100 psf, K = 1.5) 



554 



-iSOOOr 



-fOOOO- 



5000- 



O • 



> b" 



5000- 



10000- 



ISOOO 



r 



EXTENSION 



COMPRESSION 



5000 



5000 




z 
< 



-15 



-10- 



-5- 



5- 



10 



EXTENSION 












COMPRESSION 






I 


1 


L 1 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-11. Cyclic Test Envelopes for Test No. 18 io'^ = 4,100 psf, K^ = 2.0) 



555 



-i9000r 



toooo- 



-5000- 



5000 



10000 



l»000 



EXTENSION 



* 



COMPRESSION 
*■ CYCLIC STRESS NOT RECORDED FIRST 17 CYCLES 



UJ 

ec 

z> 
«/) 
i/> 

UJ 

tr 

a. ^ 

m 
< 



5000 



-50O0 







~ 


1 

1 y ^ 




\^^^ 







-15 



-10 



-5 



10 



■~" — ^—^ EXTENSION 



COMPRESSION 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-12. Cyclic Test Envelopes for Test No. 19 (o' ^ = 4,100 psf, K^ = 2.0) 



556 





-I500O 


to 


-10000 


V) 




UJ 




ac 




^- 




in 


-sooo 


cr 


^ 


o 


• 


»- 
< 





> 


b° 


u 




n 






5000 


o 




.J 




o 




>- 
u 


10000 



15000 



EXTENSION 



COMPRESSION 



-T-r 



5000 



-50O0 



/ 



/ 
/ 
/ X 

/ / 
/ 



-f-t ' 



-T-T 



tn «" 



-10 - 



-, _ -^ ~ ~~ ^ 



__EXTENSION 

compression" ~~ — -^ __ 

¥: Compression strain beyond recording limlls 
I I L. 



10 



40 



60 65 



20 30 

NUMBER OF CYCLES 
Figure G-13. Cyclic Test Envelopes for Test No. 20 (a'^^ = A, 100 psf, K^ = 2.0) 



557 



-I50CX) 



-lOOOO- 



-5000- 



< Q 
> b" 

Q 
O 



5CO0>- 



G 10000- 



19000 




5000 



-5000 




< 

*- -^ 

< 

X 



-15 



-10 



-5- 



10 





* COMPRESSION STRAIN 


ENVELOPE 


BEYOND 








RECORDING 


LIMITS 










- 


EXTENSION 














"~"---^ 


-v-^^ 










COMPRESSION 
1 






— ^'"x 

•^ 

1 ~~" ~~ 


* . 


-^^_ 



10 



40 



50 



20 30 

NUMBER OF CYCLES 

Figure G-14. Cyclic Test Envelopes for Test No. 22 (a' = ^,100 psf, K = 2.0) 



558 



ISOOCf- 



< c 

> b' 

u 

Q 



-IOOOO>- 



50001- 



0(- 



50O0^ 



G looooh 



EXTENSION 



COMPRESSION 



iSOOOl 




5CO0 



5000 



■^^ 



-15 



-lOh * COMPRESSION STRAIN ENVELOPE BEYOND 
RECORDING LIMITS 



° k - — . - 



5- 



— EXTENSION 



COMPRESSION 



>-r 



20 30 

NUMBER OF CYCLES 



40 



<Z 



¥^\ — • 



60 65 



Figure G-I5. Cyclic Test Envelopes for Test No. 23 (a', = 4,100 psf, K = 2.0) 



559 



-ISOOOr 



-10000- 



-5000- 



»000 



10000 



15000 



EXTENSION 



COMPRESSION 



-'> 



UJ 

cr 
(/> 

Ui 

< 

UJ 
(T 
O 
Q. 



5000 



-5000 



* PORE PRESSURE TRACES UNCERTAIN 



-r> 



-10 



-5 



10- 



X — — ^_ EXTENSION 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



-L. 



401 



-'-^ 



55 



60 



Figure G-16. Cyclic Test Envelopes for Test No. 24 (a' = i*,100 psf K = 2 O) 

3c "^ ' c 



560 



-19000 



-10000- 



5000- 



< o 

> b" 

u 
o 

o 

-J 
o 
>- 
u 



5000- 



10000- 



15000 



EXTENSION 



COMPRESSION 



5000 



z> 
I/) 

UJ 

a ^ 
a: <k 

UJ _ 



-5000 







y " 


/ 




1 




1 




f 




^ 




* LOWER PORE PRESSURE ENVELOPE BEYOND RECORDING LIMITS 





-10 


- 


-5 




EXTENSION 




f) 


^ ^^COMPRESSION 


in 


1 1 ""-r i 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-17. Cyclic Test Envelopes for Test No. 25 (a', = ^,100 psf, K = 2.0) 

jc c 



561 



-ISOOOr 



-10000- 



5000- 



> b" 

UJ 

a 
o 



5OO0- 



o 

o lOOOO 



I5000 



--< 



EXTENSION 



COMPRESSION 



-'> 



5000 



M 

UJ 

ir 



5000 






0<^ 



i^ 



-15 



-10- 



-5- 























J 


' 


- 










1 


"^ — 


1 


EXTENSION 

COMPRESSION 

1 


-=. 


— ■=5 
1 , 













10 




20 


30 




40| 



0.=T^ 



5- 



Figure G-18. 



10 

NUMBER OF CYCLES 
Cyclic Test Envelopes for Test No. 28 (a' 




- = 4,100 psf, K = 2.0) 
3c ' "^ ' c 



562 



-15000 



M 

hi 

»- 

CO 

> b' 

o 
o 



-10000- 



5O00- 



5000 



10000- 



I5O00 



---' 



EXTENSION 



COMPRESSION 



-%' 



50O0 



^' 



-5000 



/.- — 



-'J' 



aS 



-15 



-10- 



-5 



5- 



10 



-'V- 



- extension 
compression" 



K) 



40 



125 



130 



20 30 

NUMBER OF CYCLES 

Figure G-19: Cyclic Test Envelopes for Test No. 29 (a', = '♦,100 psf, K = 1.5) 



563 



-15000 



-10000- 



- 5000 ~ ^ _ _ _ 



> b" 

O 



5000 = 



o 10000 



15000 



.--' 



EXTENSION 



COMPRESSION 



-'^' 



5000 







/ 

( 



-5000 



f> 



-r-T 



-5 



5 - 



10 - 



>> 



EXTENSION 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



4CH 



135 



140 



Figure G-20. Cyclic Test Envelopes for Test No. 30 (a' = 4,100 psf, K = 2.0) 

3c c 



564 



-60000 



(O -40000 
(/) 

u 

(T 

J^ -20000- 



O S 

> kT 

o 



20000- 



O 40000 



600001 



EXTENSION 



COMPRESSION 



30000 



200OO 



10000 



-lOOOO 



/ 

i 

1 

1 

r 






,^-'' 


^^' 


-- 


^-^-^ 


^ 


1 

I 

/ 





M 



-15 



-10 - 



ac -5 

3? 



5- 



10, 



/ 
/ 


^ 


^ 




EXTENSION 


-~ 


- 


- 


- 


- 


— 


~~ 


-. 






-J 


1 


COMPRESSION 
1 1 


■^ 


^ 


_L 


- 


- 


- 


- 


— 






20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-21 . Cyclic Test Envelopes for Test No. 33 (a', = 28,700 psf, K =1.0) 



565 



-60000 



-40000 



-20000 






_ Q. 



20000 - 



40000 - 



60000 



EXTENSION 



COMPRESSION 



30000 



20000 - 



10000 



-10000 




-15 



-10 



-5- 



5- 



10, 



EXTENSION 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-22. Cyclic Test Envelopes for Test No. 3^ (o'-, = 28.700 psf, K = l.O) 

Jc c 



566 



-60000 



-40000 



a: 

U) -20000 



> b' 

"^ 20000 
O 



O 40000 



60000 



EXTENSION 



COMPRESSION 



30000 



20000 - 



10000 



-10000 




(/) o^ 



I 



-15 



-10 



-5- 



5- 



10. 



EXTENSION 



g I — . — -•— ' 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-23. Cyclic Test Envelopes for Test No. 35 (a' ^ =28,700 psf, K^ = l.O) 



i 

58—78786 



567 



-60000 



-40000- 



(A 

I;; -20000 



20000 



40000 



60000 



EXTENSION 



COMPRESSION 



90000 
UJ 

c 

Vi 

0^ 20000 

UJ 
(T 

/ 
(T <^ lOOOOM- 

I 

< 



-10000 



-15 



9 
IT 



- 1 


'extension 


/ 


- / 


/ 






COMPRESSION 


_ 1 1 1 1 



Ot^ 



5- 



10 



10 



20 30 

NUMBER OF CYCLES 



40 



Figure G-24, Cyclic Test Envelopes for Test No. 36 (a' = 28,700 psf, K = l.O) 

Jc t 



568 



-eoooor 



y, -40000J- 
vt 

UJ 

^ -20000 

IT « 

> b' 

UJ 

o 



- EXTENSION 



20O0O 



O 

O 40000 



60OO0 



- COMPRESSION 



30000 

UJ 

D 
CO 

«^ 20000 

UJ 
(T 

a ^ 

cr <* 10000 - 



-lOOOO 





~- 


_ / 

/ 




/ 
_/ 


1 


/ 


1 





-15 



-10 



EXTENSION 



/ / 

/ ^COMPRESSION 

/ 



10 20 30 40 

NUMBER OF CYCLES 



50 



Figure G-25. Cyclic Test Envelopes for Test No. 37 (a' = 28,700 psf, K = l.O) 

jC c 



569 



-6OO00 



-40000- 



;;; -zoooo- 



2CXXX)- 



40OO0- 



600001 



>-' 



EXTENSION 



COMPRESSION 



-'-' 



6CX>00 



40000 



20OO0 



UJ 

c 
o 

a 



0- 



-20000 



<- 



-15 



-10- 



-5- 

















- 












^ 








EXTENSION 








- 




( 


COMPRESSION 
1 


1 




.. 







10 


20 


30 




40 



NUMBER OF CYCLES 




330 



Figure G-26. Cyclic Test Envelopes for Test No. 38 ( a' = 53,300 psf, K =1.0) 

^C c 



570 



Vi 



-60000 



-40000J- 

ui 

<r 

M -20000- 
P • 



> k' 

UI 

o 

200001- 

_i 
u 

U 40000 



•0000^ 



EXTENSION 



COMPRESSION 



60000 




-20000 



-15 



Z 



I 

/EXTENSION 



'COMPRESSION 



20 iO 

NUMBER OF CYCLES 



40 



50 



Figure G-27. Cyclic Test Envelopes for Test No. 39 ( a' = 53,300 psf, K =1.0) 

jC c 



571 



-600C» 



y, -400CX>t- 
V) 

III 
<r 

vi -20000- 



p • 



UJ 

o 



20O00- 



u 

O 40000 



60000^ 



EXTENSION 



COMPRESSION 



60000 

a 

to 

<^ 40O00 

UJ 
K 
O. ^ 

M 

Jt «^20000 



-20000 



Qt— - 



-15 



-10- 



o:. 



5 - 



EXTENSION 






COMPRESSION 



20 X 

NUMBER OF CYCLES 



40 



50 



Figure G-28. Cyclic Test Envelopes for Test No. hO (a' = 53,300 psf, K =1.0) 

jC c 



572 



-15000 



-tOOOO- 



5000- 



O 



5000 = 



10000- 



15000 



EXTENSION 



COMPRESSION 



-'> 



5000 



0^ 



-5000 



-f-r 



jrJr 



a? 



-15 



-10- 



-5- 



10, 



10 



>> 



EXTENSION 



COMPRESSION 



20 30 

NUMBER OF CYCLES 



4CH 



^'-^ 



100 105 



Figure G-29. Cyclic Test Envelopes for Test No. k\ (a' = 4,100 psf, K^ = l.O) 



573 



-60000 



y, -40000 

Ui 

w -20000- 



O S 
< e 

> b' 
Ui 

o 



_ EXTENSION 



20000- 



40O00 



60000*. 



- COMPRESSION 



eocoo 



40000- 



^20000 



-20000 




-15 



I 

-10|- / 

/ 

/ 

/ 



5- 



10 



'EXTENSION 



/ 

f /compression 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-30. Cyclic Test Envelopes for Test No. kl (a' = 53,300 psf, K =1.0) 



574 



-60000 



■40000- 



-20000 



O * 



u 
a 

u 

-J 
u 
>- 
o 



20000 



40000 



6O0O0 



-'-' 



EXTENSION 



COMPRESSION 



-.'-r 



I 



30000 



20000 - 



«^ toooo- 



-10000 



-'-r 



-'-■' 



z 
cr 



-15 



-10- 



-5- 



5- 



10 













- 






4 


' 






EXTENSION 










COMPRESSION 


~'^~ 




- 


1 


i i 


1 







10 


20 30 


40* 





NUMBER OF CYCLES 




440 445 



Figure G-31 . Cyclic Test Envelopes for Test No. ^3 (a ' ^ = 28,700 psf, K * 2.0) 



575 



60000 



■40O0O- 



U 

I;; -20000 

IT „ 

> b' 

UJ 

20000 



O 40000 



60000 



EXTENSION 



COMPRESSION 



30000cr 




■10000 



z 

IT 

*- >5 
en «" 



-15 



-5 



10 



""*"— -.^ ^ — — — _^EXTEN 



EXTENSION 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-32. Cyclic Test Envelopes for Test No. kk (a* = 28,700 psf, K = 2.0) 



576 



-60000 



y, -40000f=^ — 

(/) 

U 

a: 

t5 -20OO0 

> b' 

UJ 



20000- 



40000 - 



60000 



EXTENSION 



COMPRESSION 



SOOOOcr 



20000 



o: * »OO00|L 



UJ 



-10000 




9 



-15 



-10 



-5 



10 



EXTENSION 



^COMPRESSION 
\ I 



to 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-33. Cyclic Test Envelopes for Test No. ^5 ( o' = 28,700 psf, K = 2.0) 



577 



-•oooor 



-40000- 



-20000- 



O S. 

UJ 

a 

soooo 



t> 40000 



60000 



EXTENSION 



COMPRESSION 




-10000 



9 

tr 



-10 



-5 



10 



EXTENSION 



\ 



\ 



\ 



N 



\ 



\ 



\ 

^^COMPRESSION 

J ii_l L 



10 



20 90 

NUMBER OF CYCLES 



40 



50 



Figure G-34. Cyclic Test Envelopes for Test No. ^6 (0' = 28,700 psf, K = 2.0) 

3c c 



578 



-ISOOOr 



y, -looooH 



-5000- 



(T 


^_ 


O 


• 




< 


a. 


> 


b' 


UJ 




a 






5000 


o 




_j 




o 




<_> 


10000 



15000 



EXTENSION 



COMPRESSION 



-'-r 



5000 



Or 



5000 



if-h 



/ , 



JTZ 



-15 



-10- 



-5 



5- 



10 



JrJr 



EXTENSION 



COMPRESSION 



20 30 

NUMBER OF CYCLES 



401 



350 355 



Figure G-35. Cyclic Test Envelopes for Test No. kl (a' = 4,100 psf, K = l.O) 

ic c 



579 



-60000 



•40000- 



-20000- 



20000- 



40000 - 



60000 



• EXTENSION 



COMPRESSION 




-10000 



z 
<r 



■10 



-5 - 



EXTENSION 



'^ COMPRESSION 
\ 
\ 



10 



20 30 

NUMBER OF CYCLES 



40 



Figure G-36. Cyclic Test Envelopes for Test No. ^8 (a' ^ = 28,700 psf, K^ = 2.0) 



580 



-ISOOOf 



-10000- 



-5000- 



9000 



10000 



19000 



EXTENSION 



.--■1 



COMPRESSION 



-'-' 



UJ 
(C 

(A 
(/) 

UJ 
IT 

a ^ 

• 
a: * 



9000 



9000 



-'Jr 



/ 



/ 



/ 



-'-r 



a? 



-15 



-10 



-5 



10. 



EXTENSION 



0^ 



COMPRESSION 



10 



401 



75 



80 



20 30 

NUMBER OF CYCLES 

Figure G-37. Cyclic Test Envelopes for Test No. 49 (a' ^ = 4,100 psf, K^ = 2.0) 



581 



-60OOO 



-400OO- 



M 
UJ 

(E 

^ -20000 



20000 



40000 - 



60000 



EXTENSION 



^''' 



COMPRESSION 



30000 



20000 



/ 



a: «^ 10000 H 



-10000 



''.' 



-10 



-5- 



5- 



10- 



15 























1 


* 




>^ 


1 


~ 


— 


__ EXTENSION 
COMPRESSION 

1 1 


~ 


^ 


~ 


^ 


< 







10 






20 30 










4<? 




NUMBER OF CYCLES 

Figure G-38. Cyclic Test Envelopes for Test No. 50 (a' =28,700 psf, K = 2.0) 



582 



-6O0O0 



y, -40000- 

UJ 

a: 

^ -20000}- 



20000- 



40000 



600001 



EXTENSION 



COMPRESSION 



60000 



40000 f 



'200O0 



-20000 




-10 



-5- 



10 



la 



N 



EXTENSION 



COMPRESSION 



•0 



40 



50 



20 30 

NUMBER OF CYCLES 

Figure G-39. Cyclic Test Envelopes for Test No. 51 (a' = 53,300 psf, K = 2.0) 



583 



-60000 



y, -40000J- 
</) 

u 
a: 

V> -20000- 



UJ 



20000- 



u 

U 40000 



eoooo^ 



EXTENSION 



COMPRESSION 



60000 



40000 



'20OO0 



-20000*- 



I 

Ir- 



-10 



-5 



10- 



r::_-^____EXTENsiON 

compression" ~~ -^"^^-=~-=-:r^-::_--__ 



X 



a. 



40 



50 



O K> 20 30 

NUMBER OF CYCLES 

Figure Q-kO. Cyclic Test Envelopes for Test No. 52 (a'3<- = 53,300 psf, K^ = 2.0) 



58A 



-75000 



y, -50000 

</) 

UJ 

a: 

V) -25000|- 



P • 



UJ 

o 

25000 
Sri 

-t 
O 

O 50000 
75000 



EXTENSION 



COMPRESSION 



•oooo 




-20000 



15 



N EXTENSION 



to- xCOMPRESSION 



10 



40 



50 



20 30 

NUMBEI^ OF CYCLES 

Figure G-41 . Cyclic Test Envelopes for Test No. 53 (0*3^ = 53,300 psf, K^ = 2.0) 



585 



-60000 



(^ -40000 



\r) -20000 



20000- 



40000 



60000^ 



EXTENSION 



>*' 



COMPRESSION 



-'-' 



60000 



40OO0- 



(£ '^ 20OO0 J,' 
i*J , I, 

- 



-20000 



>"> 



W o^ 



-5- 



10- 



^^-=i=~__ EXTENSION 



20 30 

NUMBER OF CYCLES 



40 



ISO 



165 



Figure G-42. Cyclic Test Envelopes for Test No. 55 (o ' , = 53,300 psf, K = 2.0) 

jc c 



586 



-60000 



-40000 



-20000- 



EXTENSION 



600001 



20000- 



40000 - — 1— — „Tr:.T. 



COMPRESSION 



60000 



y* 40000 



20OO0 



-20OO0 




3? 



-10 



-5 



15 



v"^x 

\ \ EXTENSION 

u \ ^ 
\ 
\ 

N 
\ 

COMPRESSION 



_L 



10 



-L 



-L 



40 



50 



20 30 

NUMBER OF CYCLES 

Figure G-43. Cyclic Test Envelopes for Test No. 56 (a ' = 53,300 psf, K =1.5) 



587 



-60000 



^ -40000 



M -20000- 

C 
O 



< o 

> k' 

UJ 

o 
y 
o 



20000- 



40000 



•0000^ 



EXTENSION 



COMPRESSION 



•oooo 



40000 



- / 



^20000 4 



-20000^ 



//" 



-10 



-5 



10- 



J. 



— C~.C--^ ^_EXTENSION 

COMPRESSION 



X 



O O 20 30 

NUMBER OF CYCLES 

Figure G-kk. Cyclic Test Envelopes for Test No. 57 (a' 



40 



50 



3j. = 53,300 psf, K^ = 1.5) 



588 



-15000 



-•0000- 



-sooo- 



5000- 



10000 



1 5000 



EXTENSION 



COMPRESSION 



5000 



5000 




z 
< 

*- :5 
Vi r 

m 
-J 
< 
X 
< 



-10 



-5 



10 



15 



EXTENSION 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-45. Cyclic Test Envelopes for Test No. 58 (a" = 4,100 psf, K = 2.0) 

jC c 



589 



-coooo 



y, -400001- 
(/) 

u 

^ -20000- 



p • 



u 
o 



20000- 



o 400OO 



eoooo^ 



EXTENSION 



COMPRESSION 



eoooo 



«^ 40000 



^20000 



-20OO0 




-lU 

-5 






EXTENSION 





^ ^^ 


-V 




\ \ 




\ 


b 


- \ 




\ 




\ 




\ 


10 


\ 




\ 


IS 


COMPRESSION 

1 1 I L_ . 



to 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure Q-kS. Cyclic Test Envelopes for Test No. 59 (a' = 53,300 psf, K =1.5) 

jC c 



590 



-600O0 



-40000- 



(/> 

Ui 

oc 

^ -20000 

< * 

2OO0O 



G 40000 



EXTENSION 



COMPRESSION 



60000^ 




-10000 



i 



EXTENSION 



10 



COMPRESSION 



X 



» 

Figure G-47. Cyclic Test Envelopes 



40 



50 



20 SO 

NUMBCR OF CYCLES 

for Test No. 60 (a ' ^^ = 28,700 psf, K^ = 1-5) 



591 



-•OOOOf 



y, -40000- 

Ui 

c 

I;; -20000 

o s. 

- * 

> k 



20000 



O 4O000 



•oooo 



EXTENSION 



COMPRESSION 




-10000 



10 



-5- 



9 



10- 



EXTENSION 



COMPRESSION "^ ^ 



10 



to 90 

NUMBCR OF CYCLES 



40 



50 



Figure Q-k8. Cyclic Test Envelopes for Test No. 61 (a' = 28,700 psf, K =1.5) 



592 



-60000 



-40000- 



-20000 



20000 



40000 - 



EXTENSION 



eoooo 



COMPRESSION 




-10000 



-10 



2 -= 



10 



|- / 
/ 
/ 



EXTENSION 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure Q-kS. Cyclic Test Envelopes for Test No. 62 (o ' 3^ = 28,700 psf, K^ - 1.5) 



593 



-600O0 



y, -40000h 
(/) 

ti -20000 
> b' 

UJ 

o 



20000- 



O 



O 40000 



6OOO0 



<,'-■' 



EXTENSION 



COMPRESSION 



-'^r 



30000 



in 

«« 20000 

UJ 
(T 

a. ^ 

ac * looool- / 






■10000 






-"-' 



z 

i 



10 



-'-' 



EXTENSION 



z: __::.. 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



-'-'' 



40" 



295 300 



Figure G-50. Cyclic Test Envelopes for Test No. 63 (a' = 28,700 psf, K = 1.5) 



594 



-60000 



-40000 



-20000- 



< e 
> h° 

a 



20000- 



40000- 



60000*- 



EXTENSION 



>-' 



COMPRESSION 



-'.' 



60000 



4O0O0 



'200O0 



-20000 



^-> 



/ / 
1/ 



<- 



t 



EXTENION 



COMPRESSION 



to 



40 



-'-' 



245 



250 



20 30 

NUMBER OF CYCLES 

Figure G-51. Cyclic Test Envelopes for Test No. 64 (a ' = 53,300 psf, K = 1.5) 

^c t 



595 



-60000 



-40000- 



V) 

UJ 

(rt -zoooof- 



< • 

> b^ 

u 
o 



20000- 



40000- 



eoooo^ 



EXTENSION 



COMPRESSION 



ftOOOO 



40000 



•20000 f 



-20000>- 




^ 



-10 


/extension 

/ 

/ 

/ 
/ 


-5 



/ 
/ 


■^^-^ 




compression 


5 

n 


1 1 1 1 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-52. Cyclic Test Envelopes for Test No. 65 (a ' , = 53,300 psf, K =1.0) 

^C c 



596 



-eoooo 



y, -40000 

UJ 



EXTENSION 
Jn -ZOOOOt- ■- 



> b' 

u 
o 



eooooL 



zooooh 

COMPRESSION 



40000 



60000 



40000 



'20O00 



-20000 




# 



■10 



/extension 



/ 



/ 



/ 



/ 



— ^ 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-53- Cyclic Test Envelopes for Test No. 66 (a ' ^^ = 53,300 psf, K^ = 1.0) 



597 



-6OO00 



y, -40000h 



V> -2000O- 



P • 



> kT 

Ui 

o 



2CXXX)- 



o 

O 40000 



600001 



EXTENSION 



COMPRESSION 




-20000 



-15 



/EXTENSION 



/ 'COMPRESSION 

1/ ' 



5- 



20 30 

NUMBER Of CYCLES 



40 



Figure G-S'*. Cyclic Test Envelopes for Test No. 67 (a ' = 53,300 psf, K =1.0) 



598 



-60000 



y, -40000 
to 

UJ 
(T 

^ -20000- 



O 5 
< c 

> b^ 
UJ 

o 



20O0O- 



40000 



60000L 



EXTENSION 



COMPRESSION 



60000 



40000- 



'20000 



-20000^ 




-15 



-10 



-5 



5- 



10 



EXTENSION 



/COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-55. Cyclic Test Envelopes for Test No. 68 (a ' ,^ = 53,300 psf, K = 



.0) 



599 



-60000 



^r, -40000 

UJ 
(T 

^ -20000- 



O S 

< c 

> b' 
UJ 

o 



20000- 



u 

O 40000 



60000^ 



EXTENSION 



- COMPRESSION 



60000 




-20O00 



3? 



-15 



- (EXTENSION 



COMPRESSION 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-56. Cyclic Test Envelopes for Test No. 69 (a' = 53,300 psf, K = l.O) 

jC c 



600 



-60000 



-40000- 



(/> 

UJ 

^ -20000 



2OOO0 



40000 



60000 



EXTENSION 



COMPRESSION 



30000 



^ 200O0 - 



< 



10000 4 






-10000 




-15 



-10 



-5 



/ 

/ 

/ EXTENSION 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-57. Cyclic Test Envelopes for Test No. 70 (a' = 28,700 psf, K^ = 1.0) 



601 



-60000 



-40000- 



^ -20000 - 

o £ 

< 

> b" 

UJ 

a 



20000 



40000 



60000 



EXTENSION 






COMPRESSION 






30000rr 



20000 



10000 



-10000 




9 



-10 



-5 



10- 



EXTENSION 

^ 
/ \ 



—^ \ 



* EXTENSION STRAIN BEYOND RECORDING 
LIMITS AFTER CYCLE NO. 4 



\ 



\ COMPRESSION 



K) 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-58. Cyclic Test Envelopes for Test No. 71 (o', = 28,700 psf, K = 1.5) 



3c 



602 



-ISOOOr 



(A -KXX)Of- 

UJ 

K 

t) -9000- 



s • 



O 



9000- 



u 

u fOQOO 



19000 



EXTENSION 



COMPRESSION 



toooo 




-20000* 



-15 



10 - 



-5 



& - 



EXTENSION 



COMPRESSION 



10 



40 



50 



to 90 

NUMBER OF CYCLES 

Figure G-59. Cyclic Test Envelopes for Test No. 72 (a ' ^ = 16,400 psf, K^ = l-O) 



603 



-15000 



(rt -10000 
to -SOOO 



< c 
> b' 

UJ 

o 



5000- 



o 

O 10000 



I5OO0 



EXTENSION 



COMPRESSION 



20000 



fn 10000 -f 



-lOOOO 



-20000 




S5 



-15 



-10 



-5 



10 



EXTENSION 



£^ 



COMPRESSION 



K) 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-60. Cyclic Test Envelopes for Test No. 73 (a ' , = 16,^00 psf, K =1.0) 

jc c 



604 



-15000 



-10000- 



</) 
V) 

Ui 

In -9000 

5 1 



9000 



10000 



19000 



\ EXTENSION 
\ 
\ 



.-^ COMPRESSION 




-20000 



-10 



/ EXTENSION 



COMPRESSIN 



10 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-61 . Cyclic Test Envelopes for Test No. Jk (a ' ^^ = 16,^400 psf, K^ - l-O) 



605 



w -10000 

UJ 

c 

^ -5000 

Is 

o 

5000 
(_> 

-i 
u 

U 10000 


-^ ^ EXTENSION 


_- COMPRESSION 



UJ 20000 
c 

(A 

2 lOOOO 

a. 



-10000 



-20O00 



-15 



-10 - 



10 



EXTENSION 



COMPRESSION 



5 - 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-62. Cyclic Test Envelopes for Test No. 75 (o ' = 16,^00 psf, K = l.O) 

ic c 



606 





-15000 


i/) 


-10000 


v> 




UJ 




IT 




(O 


-5000 


g 


•1 


< 




a 


> 


b^ 


UJ 




o 






5000 


u 




_) 




o 




o 


lOOOO 



I5OO0 



-"' 



- — ~ __ EXTENSION 



COMPRESSION 



-'-r 



, , 20000 
(T 

J2 lOOOO 
(T 

a. 



-KXXX) 



-20000^ 



/ ^ 



-15 



-I©-- 



W o- 



10 



-'-^ 



EXTENSION 



COMPRESSION 



20 30 

NUMBER OF CYCLES 



-'-' 



40l I 90 95 



Figure G-63. Cyclic Test Envelopes for Test No. 76 (a ' = 16,^00 psf, K = l.O) 



607 



o 



o 

























-10000 


- 










-5000 

i 


_/" "* ^_ 


EXTENSION 


.'*' 








> 










5000 


- 














COMPRESSION 


■ " "'f 1 


10000 


- 










15000 





















UJ 
IT 
13 
M 
(A 
Ui 
(T 

a. ^ 

• 

a: <^ 



5000 



0< 



5000 



^'> 



-'> 



-W- 



-5 



5 - 



10 























- 


















^ 


^ ~ 








EXTENSION 


















1 


COMPRESSION 

1 


1 


~ 


- 


- 


1 









10 


20 


30 








40i 



, ^ __ __ 




^4UMB€R OF CYCLES 

Figure G-Si*. Cyclic Test Envelopes for Test No. 77 (a ' = k ,\m psf, K =1.5) 

^C c 



608 



-i5000r 



-HOOOO- 



■5000- 



> b" 

u 
a 



5000 



o 

o 10000 



19000 



^ EXTENSION 



COMPRESSION 



5000 



D 
(/) 

to 

u 
cr 

a. 

IT 

U 

»- 
< 

a 
o 
a. 



-5000 




-10 



EXTENSION 

/ 
/ 



--^COMPRESSION 



K) 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-65. Cyclic Test Envelopes for Test No. 78 (0*3^. = ^,'00 psf, K^ - 1-5) 



609 



-ISOOOr 



-10000- 



-5000- 



tt « 
O S 



UJ 

o 



9000 



10000 



19000 



EXTENSION 



COMPRESSION 



UJ 

</> 
M 

UJ 



5000 



-5000 




a? 



-10 



-5- 



5- 



10- 



15 



■ ^ EXTENSION 



COMPRESSION 



10 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-66. Cyclic Test Envelopes for Test No. 79 (a' = A , 1 00 psf, K =1.5) 

jO c 



610 



-15000 



-10000- 



-50OO- 



sr ^ 

< o 

> b* 

o 
o 



5000 



o tOOOO 



15000 



EXTENSION 



COMPRESSION 



>> 



y ■>^^ ^ 



5000 



-5000 



-f-r 






-r-r 



^ .^- 



-15 



-10- 



10 



.^ __EXTENSI0N 



-'-- 



COMPRESSION 



1^ 



105 



Figure G-67. Cyc 



10 20 30 401 

NUMBER OF CYCLES 

lie Test Envelopes for Test No. 80 (a ' 3^ = ^JOO psf, K^ = 1.5) 



611 



-ISOOO 



w - toooo -- 

to -50001- 



5 1 



UJ 

o 



SOOO- 



o 

o KXXX) 



I9000 



EXTENSION 



COMPRESSION 



20000 




-20000"- 



-10 - 



< 
on 



10 



/ EXTENSION 



COMPRESSION 



K) 



20 30 

NUMBER OF CYCLES 



40 



50 



Figure G-68. Cyclic Test Envelopes for Test No. 81 (a' 



3c 



16,400 psf, K = 1.0) 



612 



APPENDIX H 
EXTRAPOLATION OF ISOTROPICALLY-CONSOLIDATED CYCLIC 
TRIAXIAL TESTS FOR STRENGTH INTERPRETATION II 
(FIGS. H-1 THROUGH H-25) 



613 



APPENDIX H 

EXTRAPOLATION OF ISOTROPICALLY-CONSOLIDATED 
CYCLIC TRIAXIAL TESTS FOR STRENGTH INTERPRETATION II 

As described previously, the isotropically-consolidated 
test records were extrapolated to higher strain levels because 
of load attenuation and necking problems. These extrapolations 
were made conservative and are presented in the following 
figures. It should be noted that, although the extrapolations 
were intended to account for testing discrepancies, the straight 
line extrapolations were rather arbitrary and other extrapolations 
equally valid are possible. This strength interpretation is 
judged to be conservative because cyclic strain envelopes have 
a tendency to level off as cycling continues. A straight line 
extrapolation, therefore, can be considered relatively conservative. 



615 



-20 



-15 - 



-10 
< 
t7 -5 



< 



5 - 



)0 



/ 








/ 
/ 


TEST NO. 6 






/ 


Kc = 10 






/ 
/ 


0"3c' = ^'00 


psf 




/ 








/ 
/ 


N2.5 =22 




- 


/ 


Ng =44 






N,o =86 




~ 


50 100 150 


200 


250 




NUMBER OF 


CYCLES 






PREDOMINANT AXIAL CYCLIC 


STRESS 2= + 


3400 psf 




FOR FIRST CYCLES 









-20 



-15 - 



— -\0[- 

< 

K -5 
CO 

_l 
< 

X 
< 



5 - 



10 



Figure H-1. Cyclic Triaxial Test No. 6 





N25= 96 TEST NO. 7 


- 


N5 = 200 K^ = 1.0 - 




N,Q=425 (TjJ = 4100 psf ^^^ 


















































— 


^^^— -•"'^ 






^ 






50 160 150 200 250 




NUMBER OF CYCLES 




PREDOMINANT AXIAL CYCLIC STRESS C!^ + 4 300 psf 




FOR FIRST CYCLES 



Figure H-2. Cyclic Triaxial Test No. 7 



616 



-20 



-15 



-10 - 



< 

t- -5 

CO 



X 

< 



5 - 



/ 








TEST NO. 12 
Kc = 10 




/ 


°"3c' " '^'^^ P^^ 




// ^2.5= 9 




- 


/X Ns =23 
^^ N,o =52 






1 1 


— 


V -^£^^20 40 GO- 


80 100 


— 


NUMBER OF 


CYCLES 




PREDOMINANT AXIAL CYCLIC 


STRESS* + 10900 psf 




FOR FIRST CYCLES 







-20 



-15- 



- -lOh 

< 

I- -5 



< 
X 

< 



5 - 



Figure H-3. Cyclic Triaxial Test No. 12 



- 


N2.5 = 
N5 = 
N|0 


3 
■9 
= 22 




^ 


^^ TEST NO. 
Kc =10 


16 




- 


'l 








-| 


.. 1 . 


^3c = '^'OO psf 
1 1 




- 






5 


~ 


10 


15 


20 


25 














NUMBER OF 


CYCLES 










PREDOMINANT 


AXIAL 


CYCLIC STRESSES ± 


7900 


psf 






FOR 


FIRST 


CYCLES 













Figure H-4. Cyclic Triaxial Test No. 16 



617 



-20 



-15 



-10 



N2.5 =415 



TEST NO. 4 1 

Kc = 10 

CTjj.' = 4100 psf 




100 150 200 

NUMBER OF CYCLES 

PREDOMINANT AXIAL CYCLIC STRESS ^ + 4700 psf 
FOR FIRST CYCLES 



-20 



-15 - 



— -10 - 



< 

1- -5 

CO 



5 - 



10 



Figure H-5. Cyclic Triaxial Test No. ^1 





N25= 190 TEST NO. 47 


- 


N5 = 370 Kc = 1.0 -^^"^ 




N|Q = 790 0-3^ = 4100 psf ^^^ 













^^0^ ^ 






































^ 


^.^^^^"""'^ 






r 1 1^^ 1 1 

100 . 200 300 \^ 400 500 




NUMBER OF CYCLES 




PREDOMINANT AXIAL CYCLIC STRESS ^ ± 2800 psf 




FOR FIRST CYCLES 



Figure H-6. Cyclic Triaxial Test No. k"] 



618 



-20 



-15 



-10 - 



< 

•- -5 



< 

X 

< 



5- 



10 





y 








y TEST NO. 


72 






/ 








/ K^ =1.0 








y^ 0-3^' = 16 400 psf 






y 








y 








/^ ^2.5= 3 

y N5 =8 






/ 


y N,o =16 




~ 


* — 




~~25 







NUMBER OF CYCLES 








PREDOMINANT AXIAL CYCLIC STRESS :^ + 


12000 psf 






FOR FIRST CYCLES 







-20 



-15 - 



— -10 - 

z 
< 

t- -5 

_J 
< 
X 



5 - 



10 



Figure H-7. Cyclic Triaxial Test No. 72 





J, — 






-»' 






^ 






N2.5 = 5 ^^^ 






N5=I2 
^0 = 25 




' 


^^ TEST NO. 73 
^^ K^ = 1.0 




<< 


^^jj_;::lj; 0-3^ =16 400 psf 

1 . . 4-- ..;;^i^- ) 1 1 


' 






— 




NUMBER OF CYCLES 






PREDOMINANT AXIAL CYCLIC STRESS :i: ± 8600 psf 
FOR FIRST CYCLES 





Figure H-8. Cyclic Triaxial Test No. 73 



619 



-20 



-15 - 



-10 - 



< 



< 

X 



5 - 



10 





/ 


















/ 








TEST 


NO, 


74 






/ 








^ : 


1.0 








/ 

/ 








°-3c 


= 16400 psf 






/ 


















y 








^2.5 
^5 


= 2 
= 4 






^ 


^ 






1 


1 


= 9 






■"-— . 


■ — ^ — 


1 
-10 





1 
-I5_ 


20_ 




25 








NUMBER 


OF 


CYCLES 










PREDOMINANT 


AXIAL 


CYCLIC 


STRESS 


~ + 


12 900 psf 






FOR FIRST CYCLES 















Figure H-9. Cyclic Triaxial Test No. lU 



-20 



-15 - 



-10 - 



H -5 



< 
X 

< 

51- 





/■■ ■ ■ — 
/ 
/ N2 5=''^ 


TEST NO. 76 




/ 
/ 
/ 




N5 = 25 
N,o = 5l 


K = 1.0 

0-3^ = 16 400 psf 


~ 


/ 
/ 
/ 
/^^^ 


— 


-— 





- 






1 1 


1 1 




2'5- 




50 75 


100 125 








NUMBER OF CYCLES 




PREDOMINANT 


AXIAL CYCLIC 


STRESS ^ + 7000 psf 




FOR FIRST 


CYCLES 







Figure H-10. Cyclic Triaxial Test No. 76 



620 



-20 



-15 - 



-10 - 



-5 



10 





/ 

/ 


- 


N2.5=8 / 
N5=ll / 


- 


/ TEST NO. 81 




/ Kc = 1.0 


- 


/ o-jp' = 16 400 psf 




5 10 15 20 25 




NUMBER OF CYCLES 




PREDOMINANT AXIAL CYCLIC STRESS :t + 10 600 psf 




FOR FIRST CYCLES 



Figure H-11. Cyclic Triaxial Test No. 8l 



621 



-20 



N =4 
2.5 



TEST NO. 33 

Kc = 10 

cr,. '= 28 7 00 psf 




10 



10 20 ^irrso.^-- 40 50 

NUMBER OF CYCLES 

PREDOMINANT AXIAL CYCLIC STRESS -; ± 17200 psf 
FOR FIRST CYCLES 



-20 



-15 - 



— -10 - 



< 
q: 
I- -5 



< 

X 



5 - 



Figure H-12. Cyclic Triaxial Test No. 33 





y' 




y 




N25=8 y^ TEST NO. 34 


~ 


N5 =17 ^y K^ = 1.0 




N|Q=36 ^y ^^ 0-3^'= 28 700 psf 




^ ^^ 


_ 


X ^^,,,0''^ — 




^ 




/ ^^^^"^"^ 




y^ ^^^^^^ 




y^^,*^*^^ 


^ 






1 — ca, :r J^^ ' ' ' 

10 20 30 40 50 




NUMBER OF CYCLES 




PREDOMINANT AXIAL CYCLIC STRESS ^ ± 13300 psf 




FOR FIRST CYCLES 



Figure H-13. Cyclic Triaxial Test No. 3^ 



622 



-20 



- 15 - 



^-lOh 

< 

tl -5 



< 

X 



5 - 



N2.5-24 

_ N5 =49 

""''^ — "^^ 1 1 


— ^ — 

TEST NO. 35 

K^ = 1.0 

0-3^ ' = 28 700 psf 

1 1 


20 4(5" "60 80 160 

NUMBER OF CYCLES 

PREDOMINANT AXIAL CYCLIC STRESS -± 11300 psf 
FOR FIRST CYCLES 



-20 



-15 - 



- -10 - 



< 
cc 

\- -5 
to 



5 - 



10 



Figure H-1^. Cyclic Triaxial Test No. 35 













N2.5=2 




/^ TEST NO. 36 




N5 =3 

N,o--6 




/ K,= 1.0 

"^ a = 28 700 psf 
3c 


^ 






— 


/. 


. — j — __■«- 




.1 1 1 1 




2 




4 6 8 10 
NUMBER OF CYCLES 




PREDOMINANT 


AXIAL CYCLIC STRESS ^ + 21500 psf 




FOR FIRST 


CYCLES 



Figure H-I5. Cyclic Triaxial Test No. 36 



623 



- zv 




y 




— 15 


N2.5 
- N5 


= 4 /^ TEST NO. 37 
= 9 1 •'^ K, = 1.0 






N,0 


= 17 / /-^ ^^3^ = 28700 psf 

/ y 




-10 


- 


/ y 


~ 


-5 



^ 


/y 

1 —^ 1 1 1 1 


- 




5 10 15 20 25 


~~~ 






NUMBER OF CYCLES 




5 


- 




- 




PREDOMINANT AXIAL CYCLIC STRESS :^ +15600 psf 




in 


FOR 


FIRST CYCLES 





Figure H-16. Cyclic Triaxial Test No. 37 



-20 



-15 - 



-10 - 



< 

I- -5 



< 

X 



5 - 





N5 =3 ^^ 


TEST NO. 70 
K^ = 1.0 




N,0 = 6 /^ 


0- '= 28 700 psf 


^ 


y 
1 — ^^-^ 1 1 ! — 


- 




2 4 - 6 — 8- 


1 Q — — 




NUMBER OF CYCLES 






PREDOMINANT AXIAL CYCLIC STRESS 
FOR FIRST CYCLES 


cc± 23300 psf 



Figure H-I7. Cyclic Triaxial Test No. 70 



624 



- 20 



-15 - 



-10 

z 

< 
cr 
I- -5 

_l 
< 

X 



5 - 



10 





TEST NO. 38 


- 


N„ . = >1000 ^c = '0 

o-jj = 53300 psf 


- 


^„___-..-...: 




100 200 300 400 500 




NUMBER OF CYCLES 




PREDOMINANT AXIAL CYCLIC STRESS ~ + 16100 psf 




FOR FIRST CYCLES 



Figure H-l8. Cyclic Triaxial Test No. 38 



-20 



-15 - 



^ -10- 



< 

a: 

I- -5 



< 

X 



5 - 





N25= 13 TEST NO. 39 


- 


N 5 = 25 Kj. = 1.0 




N|Q = 50 o-^^' = 533 00 psf 


- 






5 10 15 20 25 




NUMBER OF CYCLES 




PREDOMINANT AXIAL CYCLIC STRESS ::^ ±28500psf 




FOR FIRST CYCLES 



Figure H-19. Cyclic Triaxial Test No. 39 



625 



-20 



-15 



'2.5 



= 6 



TEST NO. 42 

Kj, = 1.0 

0-3^' = 53300 psf 




I L 



2 3 4 

NUMBER OF CYCLES 



S 



PREDOMINANT AXIAL CYCLIC STRESS :?:+ 37700 psf 
FOR FIRST CYCLES 



-20 



-15 - 



-10 - 



< 
cr 
^ -5 



5 - 



10 



Figure H-20. Cyclic Triaxial Test No. 42 





^2.5 = 3 


^^ 


- 


N5 =6 


^^^ 




N,0='2 


y ^^-^ 






y^ ^^^ TEST NO. 65 
X^^^ K, = 1.0 


^ 


1 


^^"^ o-jg' = 533 00 psf 




^ 


4 6 8 L°___ 

NUMBER OF CYCLES 




PREDOMINANT AXIAL CYCLIC STRESS :t ± 20700 psf 




FOR FIRST 


CYCLES 



Figure H-21 . Cyclic Triaxial Test No. 65 



626 



-20 



-15 - 



■10 - 



< 

I- -5 



< 
X 



5 - 



10 



N2,5 = 5 
N5 =12 
N.o --26 


^^ TEST NO. 66 




1 ^-^ K, = 1.0 
J^ 0-3^. = 53300 psf 




\ 1 1 1 


5 


^10 i5 2'o 25 




NUMBER OF CYCLES 


PREDOMINANT AXIAL CYCLIC STRESS :^ + 20 700 psf 


FOR FIRST 


CYCLES 



Figure H-22. Cyclic Triaxial Test No. 66 



z 
< 
tr 

\- 
co 

_j 
< 


-15 

-10 

-5 



5 

in 


^2.5= 
- N5 = 

N,0 = 


3 
6 
II 

1 


^ — 

/ ^^^ TEST NO. 67 
/ ^--^ Kc = 10 

/ ^^ o-jj." = 533 00 psf 


< 


2 4 6 8 10 
NUMBER OF CYCLES 

PREDOMINANT AXIAL CYCLIC STRESS c:t: + 30100 psf 
FOR FIRST CYCLES 



Figure H-23. Cyclic Triaxial Test No. 67 



627 



-20 



< 

•- -5 
if) 



< 

X 



5 - 



N2 5 = 50 


TEST NO, 68 


Ng =100 


Kc = 10 


N,o =200 


Q- ' = 53300 psf 




NUMBER OF CYCLES 

PREDOMINANT AXIAL CYCLIC STRESS ^+ 15800 psf 
FOR FIRST CYCLES 



-20 



Figure W-lk . Cyclic Triaxial Test No. 68 




TEST NO. 69 

Kc = 1.0 

(TjJ = 533 00 psf 



5 - 



10 



4_ 6 8 ^10 

NUMBER OF CYCLES 

PREDOMINANT AXIAL CYCLIC STRESS tt + 35400 psf 
FOR FIRST CYCLES 



Figure H-25. Cyclic Triaxial Test No. 69 



628 



APPENDIX I 

CYCLIC TRIAXIAL TEST RESULTS FOR MODELED OROVILLE 
GRAVEL USING STRENGTH INTERPRETATION II 



Figures I-l through I-IO are cyclic test results 
employing strength interpretation II. 



629 




-0 'SS3diS 80iVIA3a OHOAD 



630 




'ss3yis aoiviA3a onoAO 



631 



z 












■% 2 












Q- « ^ 












e S. iiJ 












o o o ^ _1 












055- s § ^ 












'^ (0 (0 CVJ 00 












Q 03 10 C\J OD _■ _ 












-1 z 












UJ - 












> ,'*' 












< 1- ^ b 












s^tr ^ ^„ _ 

" .- UJ = ^ 










^m 










7 


^. CVJ ^ ^ > <« UJ 

-1 2 ^ 1- UJ ir 










/ 


-1 Q UJ - S "^ => 










/ 


> S°"t - ^ 










/ 


§UJ2 tl UJ ^ 










/ 


O'^pS-J?'^ 3 










/-J 


-1 < ** i^ ^ 










/ 4^ 


-1 _J UJ 1- H UJ 










/ 7 


i^ UJ Q- — < a: 










/ / 


5 q: c/) ? CD u. 










/ / 














7 












m 


/, 










S5 


/ 1 


• 













1 


' 













1 j 












+1 


/•/ 












< , 


' / 














/« 


/ 








^5 




/ 


f 


^•B 







If) 


7 


/ 


/ 


^55s5vO 






+1 




1 


/ 
















</) m 






" 




/ i 




UJ ■ • • 






< 




' / 




3 CM 10 

< +1 +1 +1 

Q > 




55 


I 




r 




2 Q 




<\J 


^N. 




/ 




_ , , UJ < < < _ 




+1 


^s. 




/ 




^ 1- •" "> - 








/ 




< - - - 




< 




^•^^ / 


/ 




UJ -1 z z z 








^^/ 


/ 




1 ? < < < 
-■ ?: (E (T q: 










/ 




^ •- •- •- 
t {/) to «n 




k^ 




<J 


1 




H 




\ 


s. 








X -I _1 _1 






^N. 








UJ < < < 






^^V. 






















"^ X X X 






^ 


,. / 






< < < 








/ ^ 






3 C 


1 




f 


< 


r 


■^ r-i 






ii^ ^ 



isd "^ 'ss3yis aoiviA3a diioao 



632 




^sd ''"-D 'SS3yiS yOiVIABQ OnOAO 



633 



OROVILLE GRAVEL 

WELL GRADED 2 tNCH TO NO. 200 
RELATIVE DENSITY, D^ = 86 % 
SPECIMEN DIAMETER = 30.5 cm 
ikiixi A 1 irc err TU/c- 


z 

z 

a. Q. ill 

_l 

o o o 

'f 00 _ _ 

ro 
b 

n- ■" . y 










CONFINING PRESSU 
BACK PRESSURE u 
K 
FREQUENCY 


■ / 


j m 






q 
in 

< 


■ 

o 
o 

< 




'/ 








m 

CVJ 


< 


/ 


Q 
2 


AXIAL STRAIN, e^ = 2.5 % ^ 
AXIAL STRAIN, e^ = 5.0% • 
AXIAL STRAIN, c^ = 10.0 % ■ 




< <- 


tk^y^ 




o 

_l 



o. 
o 



dp 



^sd ^^ 'SS3aiS d0lVIA3a OHDAD 



634 




'ss3yis yoiviA3a onoxo 



635 



z 
e ^ S UJ 

'^ (O O lo cvj <n o 
Q oo lo m 00 _• _ 

u ^" " -"„ 

< 1- ,. b 




55 


OROVILLE 
WELL GRADED 2 IN 
RELATIVE DENSITY 
SPECIMEN DIAMETE 

INITIAL EFFECTIVE 
CONFINING PRESSU 

BACK PRESSURE i 

k 


>- 
o 

z 

LU 

o 

UJ 
(T 
li. 


1 • 

IK 


6 

o 
m 






/ / 


in 

CO 




^' 


// 


< 

^* ■ 

in o o 

Cvi lO o 

Q ,1 „ 1. 




^ 


/ 


e) - ' - 

UJ ? ? ? 

1 < < < 

-■0:0:0: 

1- 1- K 
w to trt 

_J -J _l 

< < < 
XXX 

< < < 






^ 





UJ 

_J 
o 
> 
o 



dp 



isd ^ 'ss3ais aoiviA3a onoxo 



636 




^sd % 'SSBaiS b01VIA3a diioao 



637 



z 


^ 




1 

OROVILLE GRAVEL 

WELL GRADED 2 INCH TO NO. 2 
RELATIVE DENSITY, Dr = 86 
SPECIMEN DIAMETER = 30. 
INITIAL EFFECTIVE 


w 9 b 

00 CO _ 

.a u 

3 it 
UJ 

K 

3 

UJ " 






BACK PR 
FREQUE 


/ , 
■/ ^° 

/ ° 
/ ° 




1 


/ • / 


55 
o 
m 




/ i 


/ ^ 


• 


4« ■ 

lO o o 
cJ in CJ 

Q ' 




• 
1 -^ 




S J* ^'^ < 
O - - - 
LJ 12 2 

, < <t < 

-1 CC q: (E 
•- 1- 1- 

(O CO CO 

_l _l -1 

< < < 

XXX 

< < < 




1 ^ 


^ 





638 



isd "jj 'ss3yis aoiviA3a diioad 



OROVILLE GRAVEL 

WELL GRADED 2 INCH TO NO. 200 
RELATIVE DENSITY, D^ =86% 
SPECIMEN DIAMETER = 30.5cm 

INITIAL EFFECTIVE 

CONFINING PRESSURE 0- ': 53300 psf 

BACK PRESSURE ub = 8200 psf 


. 

S 
\ 

Ul 

_l 
o 

9S ■ 

CVJ _ / 
o / 


55 
o 
6 




>- / 
<-> / 

z / 
UJ / 

3 / 
o / 

UJ / 
0^ ■ 

u. T 


*\ <J^ 






/ « 


9 

IT) 






/ /^ 


< 
(1/ 


D VALUES) 
= 2.5 % ▲ 
= 5.0% • 
= 10.0 % ■ 




M ^ ^ 




UJ i^ J* ^ < 
O < - - - 

o < < < 

-J OL CC Q. Q. 

< H 1- H 
(t (rt CO (/5 

X _l _l -I 
Ul < < < 

"" X X X 

< < < 




^ ^ ^ 







jsd ''Pjs 'ss3yis yoiviA3a oiioao 



639 



APPENDIX J 

PROCEDURE FOR INTERPRETING CYCLIC TRIAXIAL 

TEST DATA TO DETERMINE CYCLIC SHEAR STRESS 

ON POTENTIAL FAILURE PLANE 



641 



APPENDIX J 

PROCEDURE FOR INTERPRETING CYCLIC TRIAXIAL TEST 

DATA TO DETERMINE CYCLIC SHEAR STRESS ON 

POTENTIAL FAILURE PLANE 



Current procedures in evaluating the dynamic strength 
of an embankment utilize cyclic strength test data to assess 
the strain potential of each individual soil element. The 
dynamic strength of the soil would be dependent upon the 
stress conditions existing on the potential failure plane 
prior to seismic loading. The dynamic strength of a soil 
element would be determined by the performance of a laboratory 
specimen which duplicated the static shear and normal stresses 
on the failure plane. A fundamental assximption is that the 
horizontal planes within an embankment are the most critical 
from a viewpoint of seismic stability. To represent the static 
stress conditions on the failure plane in the field, the static 
vertical normal stress (^y) and the alpha (OC) value are used. 
The alpha value is defined as the ratio of the initial static 
shear stress, tT^ , divided by the static normal stress, ^f^r 
on the failure plane. 

The potential failure planes in a triaxial sample are 
assumed to be, dependent upon the consolidation stress ratio 
(K ) of the sample. For isotropically-consolidated samples 
(K =1.0) the failure planes are assumed to incline 45 from 
the horizontal. Failure planes for anisotropically-consolidated 
samples are assumed to incline 45 + 0'/2 to the horizontal. 
To obtain the current stress conditions required to cause a 
specified amount of failure in a particular number of cycles, 



6A3 



A p, Mohr's circle relationships must be employed. These 

relationships are used to determine both the initial static 

stresses and the superimposed cyclic stresses on the failure 

plane in the sample. The procedures for both isotropically- 

consolidated and anisotropically-consolidated triaxial samples 

are shown in Figures J-1 and J-2. 

As discussed previously in the main text, the cyclic 

stresses in the cyclic triaxial test must be modified by the 

C correction. This correction is assumed to be unity for 
r -^ 

consolidation stress ratios (K ) of 1.5 or greater. For 
isotropically-consolidated triaxial samples, C can range from 
0.5 to 1.0 depending upon the field K value. For the 
isotropically-consolidated cyclic triaxial tests carried 
out for the modeled Oroville gravels, a C value of 0.6 was 
used. 

A fairly large testing program was carried out for 
the modeled Oroville gravel. Because of this, sufficient data 
were generated to assess most static stress conditions within 
the embankment. The cyclic strength is plotted as the cyclic 
shear stress required to cause a particular failure criterion 
in a specified number of cycles for a range of consolidation 
stresses. The shear strength envelopes for five percent com- 
pressive strain in ten cycles is presented for illustration 
in Figure J-3. 



644 




" fc - 

A r, 



r 



0.0 
= Cr -^^ — 



Figure J-1. Procedure for Interpreting Cyclic Triaxial Test 
Data for Isotropical ly-Consol idated (Kj.=1.0) 



645 



r 




o-fc - -^ [(Kj + D-lK^-l) cos(l80-2e)l 

Tj^^ — ^ [(K^-l) SIN (180 - 2 9)1 
'fic (K^-l) SIN (180-29) 



+ I)-(K -I) COS (180- 29 



(K -1) + cr. 



] SIN 



(180-29) 



~ dp 



ATT, 



SIN (180 - 29) 



Figure J-2. Procedure for Interpreting Cyclic Triaxial Test 
Data for An i sotropical ly-Consol i dated (K^=1.0) 



646 



— 20000 - 



16000 - 

a. 

'12000 

LlI 

z 
< 
_l 
a. 8000 

UJ 



4000 



-1 1 1 1 1 r 

5% COMPRESSIVE STRAIN IN 10 CYCLES -\ 




20000 40000 60000 

NORMAL STRESS ON FAILURE PLANE 
DURING CONSOLIDATION ; cr^^ ( psf ) 

Figure J-3. Cyclic Strength Envelopes for Five Percent 
Compressive Strain in Ten Cycles 



647 



APPENDIX K 

CYCLIC TRI AXIAL TEST RESULTS FOR MODELED OROVILLE 
GRAVEL USING STRENGTH INTERPRETATION I 



Figures K-1 through K-10 are cyclic test results 
employing strength interpretation I. 



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659 



APPENDIX L 
CYCLIC TRI AXIAL TEST REPORT - TO BE AVAILABLE EARLY 1979 

Appendix L is the cyclic triaxial testing report 
concerning tests made for modeled Oroville Gravel samples. 
It will be made available in 1979 and will be supplied on 
request. 



661 



APPENDIX M 
EMBANKMENT STRAIN POTENTIALS 
CFIGS. M-1 THROUGH M-4) 



663 



APPENDIX M 
EMBANKMENT STRAIN POTENTIALS 

Figures M-1 and M-2 show the computed compressive 
strain potentials in the upstream shell for the "best judgment 
case" and the "conservative case". Figures M-3 and M-^ show 
two other cases which will not be considered further because 
they are within the range defined by Figures M-1 and M— 2. 

A rigorous or well— tested procedure has not been 
developed for relating actual displacements of an embankment 
to computed strain potentials. However, two rough correlations 
between strain of laboratory test samples and embankment 
deformations have been made — Otter Brook Dam for static 
loading, and Upper San Fernando Dam for earthquake shaking. 
In both cases, surface deformations of several feet were foiuid 
for locations corresponding to axial strain of test samples 
greater than 10 percent. These deformations were considered 
"excessive" . 

On the basis of these correlations, the zones within 
the 10 percent compressive strain potential contours in Figures 
M-1 and M-2 could be expected to develop excessive deformations. 
This would be the three small zones for the "conservative case" 
and no zones for the "best judgment case". 

The method used to calculate displacement of the Upper 
San Fernando Dam can be used as a rough indicator of the magnitude 
of displacements. The method is carried out by estimating 
deformation in critical zones of high strain potential. This 



665 



PSEUDO THREE DIMENSIONAL ANALYSIS 
PSEUDO K2MAX =^50 
PSEUDO CORE Gmax/Su 
SHELL K2MAX 
CORE G^^^/S 



1750 




CASE a; PREDICTED FOR BEST JUDGMENT CASE 
Figure M-l.i Predicted for Best Judgement Case 



ESTIMATED UPSLOPE LIMIT OF SLUMPING 

V 




TWO DIMENSIONAL ANALYSIS 
SHELL K2MAX = 205 
CORE Gmax /Su = 1120 



CASE b: POSSIBLE EXTREME FOR CONSERVATIVE CASE 
Figure M-2. Possible Extreme for Conservative Case. 



NOTES = 

REANALYSIS EARTHQUAKE 

COMPUTER PROGRAM LUSH 

CYCLIC STRENGTH INTERPRETATION H - EXTRAPOLATED CYCLIC 

TRIAXIAL TEST RESULTS 

UNDRAINED CONDITIONS 



666 



TWO DIMENSIONAL ANALYSIS - COM PUTE R PROGRAM LUSH 
REANALYSIS EARTHQUAKE 



SHELL K 



2MAX 
CORE Gmav /S„ = 



205 



120 



CYCLIC STRENGTH INTERPRETATION I 
UNDRAINED CONDITIONS 




CASE c 
Figure M-3. First Strength Interpretation 



TWO DIMENSIONAL AN A LYSIS - COMPUTER PROGRAM QUAD 4 

REANALYSIS EARTHQUAKE 

SHELL K2MAX = 130 

CORE Gmax /S, = 2200 

CYCLIC STRENGTH INTERPRETATION H 

UNDRAINED CONDITIONS 




CASE d 



Figure H-k. Second Strength Interpretation 



667 



procedure requires conversion of compression strain potential 
to shear strain potential. For saturated soils defonning at 
constant voliJine in plane strain conditions, the shear strain 
potential can be taken as 1.5 times the compressive strain 
potential. Since the elements developing lower strain potentials 
will tend to restrain the movement of elements of higher strain 
potentials, an appropriate estimate of the deformation in a zone 
would employ an average value of shear strain potential. By 
taking this average shear strain potential and multiplying it 
by the height of the critical zone, one obtains the relative 
horizontal displacement between the top and bottom of the zone. 

For the "best judgment case", distribution of compressive 
strain potentials has not been defined except that they are less 
than > percent essentially throughout the upstrecim shell. For 
illustration purposes, an average of 2 percent is assumed for 
compressive strain potential over a height of 91 metres (300 ft.). 
Horizontal displacement would then be calculated as 0.02 x 1.5 x 
91 = 2.7 metres (9 ft.). 

For the "conservative case", the average compressive 
strain potential within the 5 percent contours is about 8 percent, 
and the average height within this contour is 91 metres (300 ft.). 
Relative horizontal displacement between the surface of the slope 
and bottom of this contour would be calculated as .08 x 1.5 x 91 = 
11 metres (36 ft.). Because this method is only a rough indicator, 
the displacement can best be described as a few tens of feet, or 
in round numbers, 10 metres. 



668 



Overall behavior associated with the illustrated 
strain potentials might reasonably be as follows: 

— upstream displacement of the slope by a few 
tens of feet in the interval between the two 
berms . 

— slumping of the shell material near the upper 
berm. 

— bulging of the shell material near the lower 
berm. 

Displacement and slumping would be limited to the 
upstream shell material as indicated by the strain potential 
pattern. Slumping would not be expected to extend upslope 
beyond the k^ degree line shown in Figure M— 2 (judgment based 
on extent of slumping at Lower San Fernando Dam). The compacted 
gravel in the upstream shell would be as strong and perform as 
well after deformation as before. 



669 

6—950 2-79 IM 



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