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Xibran?

of the

'Clnivereiti? of Mieconain

i L

TIMBER FRAMING

HENRY D. DEWELL

Assoc M. Am. Soc C. E.

Structural Engineer for Alaska-Yukon-Pacific Exposition;

Chief Structural Engineer for Panama-Pacific

International Exposition

Published by the

DEWEY PUBLISHING COMPANY

San Francisco

1917

^ I

Dewey Publishing Company

MAR 11 1118

PREFACE

The material in the following chapters has appeai*ed,
in the main, as a series of articles in Western Engineer-
ing, In being arranged for publication in the present
form, this has been revised and enlarged. The matter
contained therein is the result of some eleven years* ex-
perience in timber-framing, during which time I have
been intimately connected with the design and the super-
intendence of construction of nearly two hundred million
board feet of timber, most of this being represented by
the structural features of two expositions.

In this work, I have found that the published record
of timber construction is meagre. Especially is this state-
ment true of details of design, and strength of timber
joints. I have searched through all available engineering
literature for the results of tests of timber joints and
fastenings, and have been disappointed in finding so few
recorded. To supplement these few tests, I have made
additional ones on various types of timber joints.

I have, in my own work, always tried to design the
particular structure, so that it would be eflEective and
efScient in action, and at the same time be simple and
direct for the carpenter to frame. These two conditions
are not always possible to obtain; their correlation, how-
ever, is always to be sought. With this end in view, I
have, whenever possible, followed my designs into the
field, observed the framing and erection of the structure,
its behavior under load, and the effect of fime and the
elements.

The results of this experience and study are presented
in the following pages. As explained in the introductory
chapter, this volume is in no sense a text-book, and does
not cover equally all phases of timber-construction. Its
many shortcomings are realized, but it is hoped that the
contents may be of some benefit to those who may have
occasion to design or construct timber-framing.

With some of the theories advanced, and conclusions
drawn, there may be differences of opinion. I am frank
to state that certain of these conclusions may have to be
modified in the light of future tests. It is always unwise
to attempt to extend the results of tests too far. How-
ever, until such further tests are made, it is imperative
that working-values be established for present use. The
best that can be done under those circumstances is to use
the most reasonable theory that can be found, utilizing
the available tests as a guide. This method is certainly
better than a blind guess, or a rule-of-thumb method.
As an illustration of the condition just mentioned, the
present method or methods of designing bolted joints
may be cited.

I wish to acknowledge the assistance of W. L. Huber,
E. L. Cope, and C. H. Munson. Especial acknowledge-
ment is due to A. W. Earl and T. F. Chace for aid in
making calculations, and for valuable criticism and sug-
gestions in the preparation of the text, and to Robert
S. Lewis of the University of Utah, for the material on
timber mine-structures contributed by him. I desire,
also, to express my appreciation of the courtesy of the
publishers of Engineering Record, Engineering News,
Engineering and Mining Journal, and to the American
Railway Engineering Association for permission to re-
produce material from their publications.

Henry D. Dewell.

San Francisco, May 1, 1917.

CONTENTS

Page
Chapter I 7

Introduction.

Chapteb II 11

Mill and Yard Specifications, General Grading Rules.

Chapter III 26

" Unit Working Stresses, Time Element as affecting the
strength of timber.

Chapter IV 39

Washers and Pins.

Compression on Surfaces Inclined to the Direction of

the Fibres, Resistance of Timber to Pressure from

Cylindrical Metal Pins, Joints Framed with Shear

Pins.

Chapter V 58

Spiked, Screwed and Bolted Joints.

Lateral Resistance of Spikes and Nails, Common

Wood-Screws, Lag-Screws, Bolts.

Chapter VI 90

End Joints.

Chapter VII 112

Intermediate Joints.

Chapter VIII 119

Tension and Compression Splices.

Chapter IX 139

Main Members of Trusses.

Compression Chords and Struts, Composite of Lami-
nated Compression Members, Curved Laminated
Truss-Chords, Timber Tension-Members, Tension-
Rods.

Chapter X 160

Bracing-Trusses, Details of Howe-Type Roof Truss,
Lattice Trusses, Truss Connections to Posts.

Chapter XI 184

Theory of Column-Action, Tests of Timber Columns.

CONTENTS

Page

Chapter XII 194

Column Splices and Girder Connections, Floor Gird-
ers and Joists, Joist Hangers, Mill Construction.

Chapteb XIII 209

Foundations.

Chapter XIV 223

Miscellaneous Structures.

Chapter XV 246

Wind Pressure and Wind Stresses, Working Drawings.

Chapter XVI 258

Specifications for Timber-Framing.

TIMBER FRAMING

CHAPTER I

Introduction

While timber as a structural material has been largely
supplanted by steel and concrete, especially in perman-
ent work, there are still many occasions where it may be
employed advantageously in bridges and buildings, and
other structures of a somewhat permanent nature. A
knowledge of the properties of timber, its capabilities,
and its limitations for use in construction, is therefore
an essential part of the education of a civil engineer.

The old-school bridge engineer was a past master in
the art of timber framing. Many^ of his structures, it is
true, were framed more by experience and judgment
than by considerations of theory and of computed
stresses, yet the number of timber railroad bridges still
giving service testifies to the soundness of his design.
The results of his experience have been handed down to
his successors and are represented today in the accepted
standards of the railway engineer's office. Outside of
this class of engineers, however, it may be truthfully
said that neither is the art of timber framing generally
understood, nor is the value of such knowledge appre-
ciated.

For the design of wooden buildings of exceptional
size or of unusual proportions, a structural engineer is
now generally retained; otherwise, the plans for fram-
ing are prepared in the architect's office. In the latter
event the work is usually given to an architectural
draftsman possessing little experience in actual con-
struction, and only a superficial and therefore fre-
quently dangerous knowledge of structural mechanics.
This practice results from the commonly accepted ideas

8 TIMBER FRAMING

that timber designing consists in computing the sizes of
beams and girders, or in solving the stresses in a roof
truss, and that, given the required sizes of the principal
structural members of a frame, the carpenter is fully
capable of designing the joints. This conception of the
scope of timber designing is erroneous. There is no
timber structure of an appreciable size which will not
justify a careful and intelligent study of the framing
details, not alone on the ground of safety, but also from
the consideration of economy. Important details should
not be left to the judgment of the contractor or car-
penter. With all respect for the ability of the experi-
enced carpenter, there is at times nothing so impractical
as a so-called 'practical man.' I have seen instance
after instance where it would seem that the carpenter
had gone out of his way to frame a joint in the weakest
possible manner.

Obviously, the method of finding the stresses in a
structure is the same whether the material be timber or
steel or concrete, and timber joints are as susceptible to
analysis for strength as are details in any other material.
The cause of the weak details so often seen in timber
trusses has been largely the failure on the part of the
designer to realize that the joints needed attention. As
a test for the display of ingenuity and as a problem to
develop one's knowledge of practical and efficient con-
struction, the design of an ordinary mill building in
timber and the superintendence of its framing and erec-
tion has few equals.

For an intelligent design in timber, a knowledge of
sawmill and timber-yard methods is essential. The dif-
ficulties of actual framing and erection must also be
anticipated and provided for; the designer must im-
agine himself in the carpenter's place and realize, for
example, what cuts will be most difficult to make and
what holes will be hard to bore ; in other words he must
foresee in what details careless work is most likely to
occur. The possibility of the timber being green and
the consequent shrinkage must be recognized, and if

TIMBER FRAMING 9

such shrinkage is detrimental to the strength of the
structure, means must be provided for tightening the
joints after the shrinkage has taken place. As possible
incipient causes of failure by shear, the checks due to
seasoning must not be neglected. In short, all the limi-
tations of the material must be fully realized.

In the case of structures of steel, the majority of the
details can be made in accordance with the standards of
present-day practice, fully treated in the text-books and
the handbooks of the steel companies and bridge shops.
In the realm of reinforced concrete design, certain stand-
ards for detailing are being formed rapidly. For
timber, however, there are no such standards, except
those for bridge and trestle-work generally followed by
the railroad engineer. Or, it may be said that in timber
construction, many details called standard can be justi-
fied by no consideration of efficiency. Even the stand-
ards used by the old-school bridge engineers cannot be
employed indiscriminately. Certain of these, while en-
tirely suitable for the woods obtainable in the Eastern
States, have been transferred to the West and applied
without modification to timbers with entirely different
properties from those for which the details were de-
signed. The most notable example of this practice is
the use of the standard cast or malleable-iron washer
with wood as soft as Douglas fir.

For both steel and concrete design and construction
there are many good text-books and standard specifica-
tions, but for timber framing, such as heavy building
and bridgework there are only a few text-books and, to
my knowledge, no standard specifications. Among the
few books dealing with this subject, Jacoby's * Elements
of Heavy Framing' and Howe's 'Simple Roof Trusses'
are notable for their excellence, and their contents
should be mastered by anyone interested in the design
of timber structures. It is with the view of supple-
menting these and other existing works, by bringing
into correlation the drafting-room design and the re-
quirements of the field, rather than covering the whole

10 TIMBER FRAMING

subject of structural design in timber that the present
treatise has been undertaken.

A general knowledge of structural design on the part
of the reader has been assumed and no attempt has been
made to cover the whole field of timber framing, but by
discussing the advantages and disadvantages of diflEerent
typical details, I have tried to point out the structural
limitations of the material and the difficulties which
arise during construction. Only by a thorough under-
standing of the many elements that enter into the de-
sign of joints and details in timber framing is it possible
to make the finished structure safe, efficient, and eco-
nomical.

A set of general specifications for timber-work is
given in the concluding chapter. These specifications
are intended primarily for buildings, but with certain
obvious modifications are applicable to any timber
structure. Since the greatest forests occur in the West,
these specifications apply particularly to Douglas fir,
but with different unit stresses they may be used for any
other timber. The properties of Douglas fir are not far
different from those of long leaf yellow pine, so that the
specifications may be used with but slight changes for
structures built of the latter timber. With these speci-
fications, I hope to establish to some extent certain work-
able standards for timber framing in general, and for
building construction in particular.

TIMBER FRAMING H

CHAPTER II
Mill and Yard Specifications

The strength of individual sticks of timber varies
greatly. For this reason the statement is sometimes
made that refinement in calculation of timber framing,
and even the " computation of stresses, is unnecessary.
However, the variation in strength of timbers classed
under any one grade is not so great but that definite
working stresses can be established with the certainty
of such stresses being safe.

It is of the utmost importance then that the designer
should be familiar with the probable qualities of the
timber of which his structure will be built. He must
know the allowable variation in size due to sawing, siz-
ing, and surfacing, also the allowable number and size
of the knots and other defects. For this reason, there
follow extracts from the * Standard Classification, Grad-
ing, and Dressing Rules for Douglas Fir, Spruce,
Cedar, and Western Hemlock Products' as adopted by
the West Coast Lumber Manufacturers Association.
These specifications while local to the Pacific Coast are
typical of the grading rules for any timber.

Greneral Grading Rules

1. All lumber is graded with special reference to its
suitability for the use intended.

2. With this in view each piece is considered and its
grade determined by its general character, including the
sum of all its defects.

3. What is known as *'Yard Lumber,'' such as di-
mension common boards, finish, etc., is graded from the
face side, which is the best side, except that lumber
which is dressed one side only is graded from the dressed
side.

12 TIMBER FRAMING

5. The defects in lumber are to be considered in eon-
nection with the size of the piece, and for this reason
wider and longer pieces will carry more defects than
smaller pieces in the same grade.

6. No arbitrary rules for the inspection of lumber
can be maintained with satisfaction. The variations
from any given rule are numerous and suggested by
practical common-sense, so nothing more definite than
the general features of different grades should be at-
tempted by rules of inspection.

7. Lumber must be accepted on grade ill the form in
which it was shipped. Any subsequent change in manu-
facture or mill- work will prohibit an inspection for the
adjustment of claims, except with the consent of all
parties interested.

8. A shipment of any grade must consist of a fair
average of that grade, and cannot be made up of an un-
fair proportion of the better or poorer pieces that would
pass at that grade. A shipment of mixed widths shall
contain a fair assortment of each width. A shipment of
mixed lengths shall contain a fair assortment of each
length.

9. Material not conforming to standard sizes shall be
governed by special contract.

11. The grade of all regular stock shall be deter-
mined by the number, character, and position of the
defects visible in any piece. The enumerated defects
herein described admissible in any grade are intended
to be descriptive of the coarsest piece such grades may
contain, but the average quality of the grade should be
midway between the highest and lowest pieces allowed in
the grade.

12. All dressed lumber shall be measured and sold
at the full size of rough material used in its manu-
facture.

13. All lumber one inch or less in thickness shall be
counted as one inch thick.

14. In determining the seriousness of the pitch
pocket as a defect both its width an(J length must be

TIMBER FRAMING 13

considered. The tighter the pocket the longer it may be.

15. Size and number of pockets admissible in any
piece must be left largely to the judgment of the grader
and a reasonable deviation from the number of pockets
specified in the rules will be permissible.

16. Pitch shakes are clearly defined openings be-
tween the grain of the wood, are either filled with granu-
lated pitch or not, but are in either case a serious defect,
and must not be admitted in any grade above No. 2
common.

17. A pitch streak is a well defined accumulation of
pitch at one point in the piece and when not sufficient to
develop a well-defined streak, or where fibre between
grains is not saturated with pitch, it shall not be con-
sidered a defect.

18. A small pitch streak shall be equivalent to not
over one-twelfth the width and one-sixth the length of
the piece wherein it is found.

19. A standard pitch streak shall be equivalent to not
over one-sixth the width and one-third the length of the
piece it is in.

20. Splits and checks shall be considered as to length
and directions.

21. Wane is bark or lack of wood on edges of lumber
from any cause.

22. Chipped-grain consists in part of the surface
being chipped or broken out in small particles below the
line of the cut, and as usually found should not be
classed as torn-grain and shall be considered a defect
only when it unfits the piece for the use intended.

23. Torn-grain consists in a part of the wood being
torn out in dressing. It occurs around knots and curly
places and is of four distinct characters, slight, medium,
heavy, deep.

24. Slight torn-grain should not exceed ^ in. deep,
medium, 1/12 in., and heavy i in. Any torn-grain
more than i in. shall be termed deep.

25. Loosened grain consists of a point of one grain
being torn loose from the next grain. It occurs on the

14 TIMBER FRAMING

heart side of the piece and is a serious defect, especially
in flooring.

26. In standard manufacture of factory flooring,
decking, or thick-dressed and matched stock and stock
grooved for splines, and for shiplap, the finished width
shall be ^ in. less over all than the count or measured
width of the rough material used in manufacturing and
the tongue and lap shall be measured to determine the
finished width.

27. Equivalent means equal, and in construing and
applying these rules, the defects allowed, whether speci-
fied or not, are understood to be equivalent in damaging
effect to those mentioned applying to stock under con-
sideration.

Defects

28. Eecognized defects are knots, knot-holes, splits,
checks, wane, rot, rotten streaks, pin and grub-worm
holes, dog and picaroon holes, pitch seams or shakes,
pitch pockets, chipped, torn and loose-grain, solid, pitch,
stained heart, sap-stain and imperfect manufacture.

Knots

29. Knots shall be classified as pin, small, standard
and large as to size; round and spike as to form; and
tight, loose, and rotten as to quality.

30. A pin knot is tight and not over i in. diam.

31. A small knot is tight and not over f in. diameter.

32. A standard knot is tight and not over IJ in.
diameter.

33. A large knot is tight and any size over 1^ in.
diameter.

34. A round knot is oval or circular in size.

35. A spike knot is one sawn in a lengthwise direc-
tion.

36. A tight knot or sound knot is one solid across itti
face, is as hard as the wood itself, and is so fixed by
growth or position that it will retain its place in the
piece.

TIMBER FRAMING 15

37. A loose knot is one not held firmly iu place by
growth or position.

38. A rotten knot is one not as hard as the wood

itself.

TIMBER FRAMING

Pig. 3. laboe khot.

Pig. 4. small spike knot.

39. The mean or average diameter of knots shall be
considered in applying or eonstniing the rules.

Pitch

40. Pitch pockets are openings between the grain of

TIMBER FRAMING

the wood, containing more or less pitch and surrounded
by sound grain wood.

Sap

41. Bright sap shall not be considered a defect in

TIMBER FRAMING

any of the grades, except as specially provided for in
the following rales.

42. Sap-stain shall not be considered a defect except
as herein provided.

43. Discoloration of heart-worfd or stained heart
must not be confounded with rot or rotten streaks. The

TIMBER FRAMING

FlO. 9, CLUSTER OF KNOTS.

Fia. 10. cLoacn

presence of rot is indicated by a decided softness of the
wood where it is discolored, or by small white spots re-
sembling pin-worm holes.

Standard Sizes

46. In the absence of a special agreement between
the buyer and seller for each order, all dressed lumber
is finished to the following sizes.

47. Flooring: 1 by 3 in., finished size ^J by 2J-in.
face; 1 by 4 in., finished size 4S by 3i-in. face; 1 by 6
in., finished size ^f by 5i-in. face ; 1^ by 3 in., finished

TIMBER FRAMING

Fig. 12. laboe t

Flo. 13. SMALL PITCH STREAK.

TIMBER FRAMING 21

size l^^ by 2i-in. face ; 1^ by 4 in., finished size l^V by
3J-in. face ; 1 J by 6 in., finished size 1^^ by S^-in. face ;
1 by 6 in. fiat-grain fiooring, finished size f by 5^ in.
Standard lengths are multiples of one foot.

53. Widths if dressed on one or both edges: 4 in.
to Si in. ; 5 in. to 4^ in. ; 6 in. to 5^ in. ; 8 in. to 7i in. ;

10 in. to 9i in. ; 12 in. to llj in. ; 14 in. to 13 in. ; 16 in.
to 15 in. Standard lengths are multiples of one foot.

59. Common boards, SIS, or shiplap to f inch.

60. Grooved roofing, f by 7J in., 9J or Hi in. face ;
i-in. groove, 1^ in. from each edge.

61. Shiplap and Dressed and Matched: 1 by 8 in.,
finished size f by 7-in. face; 1 by 10 in., finished size
I by 9-in. face ; 1 by 12 in., finished size | by 11 in. face.
Standard lengths are multiples of two feet. .

62. Dimension, SISIE, or S4S : 2 by 4 in. to If by
3f in. ; 2 by 6 in. to If by 5f in. ; 2 by 8 im to If by

11 in. ; 2 by 10 in. to If by 9^ in. ; 2 by 12 in. to If by
Hi in. ; 3 by 6 in. to 2^ by 5^ in. ; 3 by 8 in. to 2^ by
7i in. ; 3 by 10 in. to 2^ by 9^ in. ; 3 by 12 in. to 2i by
11^ inches.

63. Timbers, SlSlE, or S4S, 4 by 4 in. and larger,
i in. oflf each way. Standard lengths are multiples of
two feet unless otherwise specified.

64. All sizes in dimensions and timbers are subject
to natural shrinkage.

Fir Common

Boards and Shiplap and Dressed and Matched :

117. One-inch select common, 4 to 12 in.; shall be
square edged; will admit sound knots not over 1 in.
diameter in 4 in. and 6 in. and not over 1^ in. in 8 in.
to 12 in., but situated away from edge; medium-sized
pitch pockets and slight stain, but should be of a sound
strong character. Hemlock permitted in this grade.

118. Common : Will admit of any two of the follow-
ing, or their equivalent of combined defects : Wane i in.
deep on edge, 1 in. wide on face, extending not over one-
sixth of the length of the piece; knots not more in di-
ameter than one-third of the width of the piece ; stain ;

22 TIMBER FRAMING

torn grain; pitch streaks; pitch pockets; seasoning
checks; one straight split not longer than the width of
one piece or a limited number of worm-holes well scat-
tered. These boards should be firm and sound and suit-
able for use in ordinary construction without waste.
Hemlock permitted in this grade.

119. No. 3 common boards or sheathing : Will admit
of all stock below the grade of common that is suitable
for cheap sheathing and will allow: Coarse knots, knot-
holes, splits, rotten sap, and any number of grub or pin-
worm holes. Hemlock permitted in this grade.

Dimen^ioin

121. Common dimension: Generally speaking, this
stock must be suitable and of sufficient strength for all
ordinary construction purposes without waste. Will ad-
mit of coarser knots than 1-in. common, which in a 2 by
4-in. should not be larger than 2 in. Spike knots not
over two-thirds the width of the piece; wane not over
i in. deep on edges and 1 in. wide on face up to 2 by 6
in., and i in. deep on edge and 1^ in. wide on face on
2 by 8 and wider, extending not more than J the length
of the piece ; stain ; solid pitch ; pitch pockets ; seasoning
checks; one straight split, not more than the width of
the piece, 2 or 3 grub-worm hples, a limited number of
pin-worm holes and torn grain. Hemlock permitted in
this grade in 4 and 6-in. widths.

122. No. 2 common dimension: This grade must be
suitable for use in a cheaper class of construction than
common. Will allow coarse and unsound knots and
knot holes that do not unfit the piece for use intended,
rotten streaks, pitch seams, pitch pockets, a reasonable
amount of rotten sap and pin-worm holes, a few grub-
worm holes well scattered. It is understood that no
culls or stock that will not work without waste will be
allowed in this grade. Hemlock permitted in this grad^
in 4 and 6-in. widths.

Fir timbers

123. Selected common : 2 by 4 in. to 2 by 12 in. and

TIMBER FRAMING 23

3 by 4 in. to 4 by 6 in. shall be square-edged. Will ad-
mit any quantity of sound knots not over 1 in. diam., or
small pitch pockets not over 4 in. in length. Sizes larger
than 4 by 6 in. will admit sound knots not to exceed
1^ in. diameter; pitch pockets not to exceed 6 in. in
length.

124. Common : Bough timbers, 4 by 4 in. and larger,
shall not be more than i in. scant when green, or i in.
scant when SISIE or S4S, and be evenly manufactured
from sound stock and must be free from knots that will
materially weaken the piece.

125. Timbers lO'by 10 in. may have a 2 in. wane on
one comer, or its equivalent on two or more corners,
one-fourth the length of the piece. Other sizes may
have proportionate defects. Season checks and checks
extending not over one-eighth the length of the piece
admissible.

126. No. 2 common timbers: This is a grade of tim-
ber that will admit of large, loose, or rotten knots,
shakes, or rot that do not impair its utility for tem-
porary work. Hemlock and white fir will be allowed in
this grade.

Fir Car Material

149. Railroad ties: Shall be sound common lumber.

Fir Bridg^e-Stringers

150. Common: Shall be sound common lumber, free
from large unsound knots or knots in clusters, or other
defects that will materially unfit the piece for the pur-
pose intended.

151. Select common : Sap shall not show on any one
corner more than 10% of any side or edge measured
across the surface anywhere along the length of the
piece. Shall be free from shake, splits, or pitch pockets
over f in. wide or 5 in. long. Knots greater than 2 in.
diam. will not be permitted within one-fourth of the
depth of the stringer from any comer nor upon the edge
of the piece ; knots shall in no case exceed 3 in. diameter.

24 TIMBER FRAMING

Weotern Hemlock

214. Western hemlock is a wood well adapted to
many uses. It is strong, holds nails well and therefore
makes good framing lumber. It is hard and wears well
as flooring. It is easily dressed to a smooth surface, and
takes a fine polish, which, together with the beauty of
grain and color, makes a fine interior finish. Western
hemlock is entirely free from the 'wind shake' so com-
mon in the hemlock of the East. This lumber has been
sold in the East under various names, such as * Alaska
pine,' * Columbia pine,' 'gray fir,' 'Washington pine,'
etc., and has given good satisfaction.

215. In a general way the rules for grading fir and
spruce are applied to hemlock.

The preceding specifications apply to lumber as ship-
ped in carload lots to the various retailers. When
lumber is purchased from the retail lumber yard, it is
usually classified as 'No. 1 Common,' or 'Merchantable.'
The specifications governing this grade are as follows,
taken from the 'Domestic List No. 6 of the Pacific
Lumber Inspection Bureau.' Edition of 1912.

No. 1 Common

This grade shall consist of lengths 8 ft. and over
(except shorter lengths be ordered) of a quality suit-
able for ordinary constructional purposes. Will allow
small amount of wane, large sound knots, large pitch-
pockets, colored sap one-third the width and one-half
the thickness, slight variation in sawing, and slight
streak of solid heart-stain.

Defects to be considered in connection with the size
of the piece.

Discoloration through exposure to the elements or
season checks not exceeding in length one-half the
width of the piece shall not be deemed a defect exclud-
ing lumber from this grade, if otherwise conforming to
the grade of No. 1 Common.

No. 2 Common

This grade shall consist of lumber 6 ft. and over

TIMBER FRAMING 25

(except shorter lengths be ordered) having defects that
prevent it being graded as No. 1 Common, but must be
suitable for a cheaper class of construction than the
preceding grade.

Will admit large coarse knots, knot holes, and splits
that do not render the piece unfit for use ; colored sap,
or wane on corner leaving a fair nailing surface, worm-
holes, large pitch-pockets, and solid heart-stain one-
half the piece.

The quality of the material purchased under these
conditions, it is hardly necessary to state, depends
largely upon the standards of the local yard, and in
general lumber purchased as No. 1 Commpn will con-
tain a considerable amount of No. 2 Common.

The accompanying photographs, reproduced from the
1915 Manual of the American Railway Engineering
Association, Report of Special Committee on Grading
of Lumber, illustrate some of the defects in Douglas
fir timbers.

2f) TIMBER FRAMING

CHAPTER III

Unit Working Stresses

The allowable unit stresses to be used in any material
are always a matter of individual judgment in the end.
They should be decided from considerations of probable
quality of material, nature of the loading, that is,
whether live or dead, and if live, whether accompanied
by impact, whether a constant load or one occurring at
rare intervals ; the particular detail under cofisideration,
the purpose which the structure is to serve, the proba-
bilities of future increase in loading or modifications of
use, of the structure, and the character of the superin-
tendence. Theoretically, the last condition may not in-
fluence the design, since the engineer is not necessarily
responsible for the field inspection and must many times
in his design assume competent and conscientious super-
intendence. Practically, however, no reputable engi-
neer would use high working stresses did he know that
the field inspection would be of questionable amount and
quality.

For steel and concrete, the working stresses have been
quite definitely established both by tests and experience.
Of all structural materials, steel is the most uniform in
quality, skilled labor being employed at all stages of its
manufacture, fabrication and erection. In the case of
concrete, the quality and strength of the finished prod-
uct can be determined in general by selecting and pro-
portioning the ingredients and by careful workmanship.
Hence for different proportions of material, working
stresses are now quite definitely established. On the
other hand, sawed lumber is a finished product, and its
strength must be judged by its physical appearance
alone. The diversity of opinion as to the proper unit
stresses or design may be demonstrated by even a cur-

TIMBER FRAMING

sory examination of the building codes of the different
cities of the United States. This lack of similarity is
especially striking in the case of the specified working
stresses for timber in bending, these stresses varying by
150%.

However, in comparing the different unit working
stresses adopted by the various building codes, it must
be remembered that the specified loadings also vary,
and to make a true comparison all stresses must be re-
duced to the same load-base. The working unit stresses
given in building ordinances are generally rather low,
in accordance with the average low grade of timber used.

Probably the best guide in selecting stresses for timber,

^Z'/yrvyv^

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S/^s/*fp/f>is

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il

Co/^^^£'^^/€>/>f-

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6UX>

JBOO

i^JlOOOO

^90

I!V

270

J/O

S30

^JO

5^oo

fSOO

900

/200{f-i/60e0

£ii»>piii^/^^'

^SPO

/300

:t fiJOOOO

7BO

^ao

3oo

iSO

^Bo

seo

3300

JSOO

930

J^OOff- i/SOcO

5Aorftot^/^/ta

6600

IJOO

J 40OOOO

I70

Aac>

fX>

540

i7o

^oo

IJOO

890

JJOO(/-l/6Oc0

M^//ff/?r>«

4^00

BOO

J J30 000

4oo

JOO

J33

290

JfiO

3000

JOOO

;»?

Joooff-i/6o«i)

^iffruce

4300

JOOO

J 3 JO OOO

SOO

J70-

370

JdO

3200

JJOQ

3SO

JJOO^J- l/6Oe0

27oo

/^b>**at//9/7ifi

4300

eoo

JJ90 OOO

jao*

J30

230

JOO

J30

SOO

eoo

eooCJ'i/fOeO

TSr/Tfona'cy^

MOO

&00

J 2£0 000

aTo

/;»

Baj

JOO

220

Ssffoo

JOOO

7*7'

Jooo(J- 2 /So a)

Miss/' /it!iv>*»cA

oeoo

JJOO

J 480 000

630

JSO

£?o

JOO

440

aso

dsoo

J200

900

J£OOrJ'l/6Oa0

^Xsa>t«oaf

Sooo

Boo

eoo OOO

300

4oo

J90

33O0

900

eao

9ooCJ'.i/4)0^

3a^JkyOya>»ss

40OO

Boo

J JOOOOO

OOO

i2o

340

J70

3900

JJOO

3^0

JJOOff- i jl6Oc0

^at/CMHtxr

-^SOO

Soo

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470

230

£000

Boo

68o

Boaff-J/A>cl)

k//u^Oa^ iJTVO

JJOO

JJSOOOO

S40

2JO

270

JJO

B£d

430

3SOO

J300

BOO

J3OCVJ'l/6Oe0

Table I — unit stresses in structubajl. timber in pounds per square inch.

Adopted by the American Railway Engineering and Maintenance of Way Asso-
ciation upon recommendation of the Committee on Wooden Bridges and Trestles.

♦Partly air-dried.

? = length in inches.

(f = least side in inches.

The stresses are for a green condition of timber and are to be used without
increasing live-load stresses for impact.

The working stresses given in this table are intended for railroad bridges
and trestles. For highway bridges and trestles, the unit stresses may be in-
creased 25%. For buildings and similar structures in which the timber is pro-
tected from the weather, and practically free from impact, the unit stresses may
be increased 50%. To compute the deflection of a beam under long-continued
loading instead of that when the load is first applied, only 50% of the corre-
sponding modulus of elasticity given in the table is to be employed.

28 TIMBER FRAMING

is the Report of the Committee on Wooden Bridges and
Trestles of the American Railway Engineering Asso-
ciation. Table I gives a summary of their recommenda-
tions. H. S. Jacoby, himself a member of this com-
mittee, says :* * * These unit stresses are the result of an
extended study of all the full-sized tests of structural
timber available, as well as the unit stresses which have
been in use in designing wooden bridges and trestles
and have been demonstrated to be safe by extensive
experience. As indicated in the footnote (see Table I),
the values are based on a green condition of the timber,
but in a few cases where no data for green timber
were available, those for partly air-dry timber were
inserted.* * * The table contains no working unit
stresses for pure tension. Wood has a greater re-
sistance to tension than to any other kind of stress,
and it is found to be diflfijcult to break it in a true tensile
test. As there is more or less cross-grain, it is advisable
to use the same unit stress in designing tensile members
as for bending. ' '

The working stresses as given by this table (increased
50% for buildings) are the highest that should be used
for any structural timber. In fact they would seem
to be too high for the material commonly used for or-
dinary building construction. The specifications given
for determining the quality of timber for which these
stresses are recommended are much more strict than
those of the lumber manufacturers.

In 'Properties and Uses of Douglas Fir,'t there is
given the result of bending-tests on 175 green bridge-
stringers, purchased in the open market, 8 by 16-in.
cross-section, and graded according to various standard
specifications. It is interesting to note that only 54 of
these stringers fell within the No. 1 railroad grade of
the Specifications of the American Railway Engineering
and Maintenance of Way Association. Table 10 of this

♦'Structural Details/ pages 358, 360.

tForest Service Bulletin No. 88, U. S. Department of Agri
culture, p, 43-45.

TIMBER FRAMING 29

bulletin shows the modulus of rupture of the high 10%,
embracing five tests, to be 8468 lb. per square inch ; that
of the low 10%, embracing five tests, 5750 lb. per square
inch. The average modulus of rupture of all of these
tests was 7108 lb. per square inch. The corresponding
stresses at the elastic limit were 5800 lb., 3374 lb., and
4516 lb. per square inch.

When graded according to the specifications for No.
2 railroad, 67 stringers came within the limit. The high
10%, comprising eight stringers, had a modulus of rup-
ture of 7430 lb. per square inch and a fibre stress at the
elastic limit of 5006 lb. per square inch. Eight stringers
also constituted the low 10%, having fibre stresses of
4761 lb. and 3109 lb. per square inch at the ultimate
strength and elastic limits, while the general average
fibre stresses were 6116 lb. and 4057 lb. per square inch.

In a similar manner the stringers were graded accord-
ing to the export grading rules of the Pacific Coast
Lumber Manufacturer's Association, adopted 1903. In
the essentials, these are the same as the present specifi-
cations of the West Coast Lumber Manufacturer's Asso-
ciation. Thirty stringers fell within the grade of * mer-
chantable.' The high 10% had a modulus of rupture of
6353 lb., the low 10% 3710 lb., while the average modulus
was 4946 lb. per square inch. The stresses at the elastic
limits were 4597 lb., 2580 lb., and 3532 lb. per square
inch, respectively.

Taking the safe working fibre stress in bending for
timbers of the grade of Railroad No. 1 as 1800 lb. per
square inch, it is interesting to compute the correspond-
ing fibre stresses for timber of the grade of * merchant-
able' by the Pacific Coast export grading rules. On the
basis of the average ultimate strength this fibre stress

4946

would be j^ X 1800 or 1250 lb. per square inch. On

3532

the basis of the elastic limit the stress would be t^tt X

451d

1800 or 1410 lb. per square inch.

Similarly, assuming that a working stress of 1800 lb.
per square inch is satisfactory for timber of the grade

30 TIMBER FRAMING

of Railroad No. 2, the corresponding stresses for mer-

4946

chantable timber would be ^tt^X 1800 or 1450 lb. per

bllD

square inch, based on the respective moduli of rupture,
and ^^^ X 1800 or 1560 lb., per square inch, based on

the respective elastic limits.

On the basis of the above comparison it is believed
that 1500 lb. per square inch is the highest working unit-
stress for bending that should in general be allowed
for ordinary building construction with good inspec-
tion. Where the inspection is likely to be either poor or
else wholly lacking, it would seem that the value of
1200 lb. per square inch should not be exceeded for
timber in bending.

This statement is made for the following reasons:
first, generally poor grade of timber as furnished by
the local lumber-yard; second, undersize of timbers,
especially of joists, due to surfacing or resawing, since
any material over 1^ in. thick sells as 2-in. stock; and,
third, holes bored, or notches cut in joists to accommo-
date conduits and pipes, these holes or notches often
being placed in the worst possible position as regards the
strength of the joist.

It will be noted that the table of unit stresses of the
American Railway Engineering and Maintenance of Way
Association gives no value for the elastic limit of timber
in bending. While it is true that timber has not the
definite elastic limit of steel, yet there is a definite yield
point and no working stresses should be determined
without a consideration of this property. The elastic
limit is especially important in the case of bearing per-
pendicular to the fibres of the timber, and the allowable
stress for cross-bearing should be based on this limit-
ing resistance and not on the ultimate strength. The
folly of small washers of insufficient size, for rods or
bolts taking tension, has been mentioned before and will
be treated more fully in a succeeding article. There are
reproduced here (Fig. 14 and 15) two diagrams taken
from Forestry Bulletin No. 88, showing the variation

TIMBER FRAMING 31

in the ultimate stresses and the stresses at the elastic
limit of the Douglas fir bridge-stringers tested.

The working stresses of Table I and the results
quoted from Forest Service Bulletin No. 88 both repre-
sent values for green timber. The effect of seasoning
on timber is, in general, an increase in strength. For
example, Forest Service Bulletin No. 88 gives the re-
sults of bending tests on green and air-dried halves of
ten 8 by 16-in. by 32-ft. stringers. That is to say, ten
green stringers, 8 by 16 in., 32 ft. long, as nearly uni-
form in quality throughout their lengths as possible,
were selected. *' One-half of each 32-ft. piece was tested
in a green condition, and the other half tested after
air-seasoning. * * The average moisture content of the
air-seasoned material was 16.4%." The average ulti-
mate strength in bending of the green material was
5440 lb. per sq. in., while the same value for the air-
seasoned timber was 6740 lb. per sq. in., or an increase
of 24%. The corresponding values for the elastic limit
were 3740 and 5478 lb. per sq. in., showing an increase
in strength due to seasoning of 47%. Also, ''a number
of tests were made on various grades of Douglas fir
stringers seasoned from six to eight months ; the grades
select, merchantable, and seconds being those defined
in the export grading rules of the Pacific Coast Manu-
facturers Association adopted in 1903. In this group
of stringers the fibre-stress at elastic limit and the
modulus of rupture, in the case of select material, are
increased, respectively, 8% and 5% by seasoning, the
modulus of elasticity remaining practically unchanged.
In the merchantable material the increase in these func-
tions is respectively 19%, 33%, and 6%. In the sec-
onds the fibre-stress at elastic limit increased 6%,
while the modulus of rupture and modulus of elasticity
show, respectively, a decrease of 12% and 2%.'* And,
''The failures in seasoned Douglas fir stringers and
car-sills were similar to those in green material, ex-
cept that failures in horizontal shear were more com-
mon."

32 TIMBER FRAMING

** Failure in horizontal shear is more common in
seasoned than in green timbers, because the net areas
resisting shear along the neutral plane is often consid-
erably decreased by checks. It seldom occurs in weak,
low-grade material, which fact is doubtless due to the
dowelling-pin action of the knots invariably associated
with low-grade timbers/'

''This summary of failures, as well as that for green
material, . . . indicates conclusively that in general
the point of greatest weakness in Douglas fir beams is
the part subjected to the highest stresses in compression
parallel to the grain. The principal exceptions to this
rule are beams that have large knots on or near the
tension-face, beams that have bad diagonal or cross-
grain, and beams that contain deep checks along the
neutral plane. The elastic limit of the beam is closely
related to the strength of the wood in compression
parallel to the grain, while the modulus of rupture is
most dependent upon the quality of the w:ood that is
subjected to tensile stresses."

Probably the most troublesome part of detailing con-
nections in timber is to make the necessary provision-
that the cross-bearing strength be not exceeded. In-
deed, it will usually be found upon examination of
typical structures that this point has not been consid-
ered. I have found that a great many designers con-
sider that the crushing of the fibres of the timber in
side-bearing is not a serious matter. The idea is preva-
lent that, after an initial crushing, no further deforma-
tion will take place, and that the structure will still be
in a working condition. The fallacy of this idea will be
realized by anyone who has seen an actual test made on
the crushing strength of timber across the fibres. In de-
signing timber framed structures, this weakness of tim-
ber must be carefully considered, otherwise, the conse-
quent deformation may unduly stress other parts. Fig.
14 shows that the average elastic limit is about 570 lb.
per square inch, therefore the working stress may be
taken at 285 lb. per square inch. Here, again, these

TIMBER FRAMING 33

■

values are for green material. Seasoned material
should show a marked increase of strength, and for
air-seasoned Douglas fir, protected from moisture, the
working stress may well be increased from 285 lb. per
square inch to 350 lb. per square inch.

Reference to Fig. 14 shows that the average elastic
limit of Douglas fir for end compression or pressure
against the ends of the fibi*es is 3612 lb. per square inch.
The working stress as recommended by Table I for build-
ings is 1800 lb. per square inch, which gives a safety
factor of two on the basis of the elastic limit. I prefer
in general to limit the pressure to 1600 lb. per square
inch. When providing for the stresses of compression,
tension, and shear, the nature of the detail should always
receive consideration in deciding the exact amount of
working stress. Under some conditions 1800 lb. per
square inch and even slightly more will provide a suffi-
cient factor of safety. For example, the basic* working
compressive unit-pressure of a lag screw in timber may
be well taken at 1800 lb. per square inch, since there is
a close and uniform fit of the lag screw in the bored hole.
On the other hand, the actual pressure of the toe of the
batter-post of a truss against the shoe-plate may be
almost as great as the ultimate strength of timber in
end compression, depending altogether upon how accu-
rately the carpenter shapes the batter-post to fit the
shoe. It is evident, then, that in details involving diffi-
cult cuts, a relatively low unit pressure should be used,
and in straight cross-cuts a corresponding higher work-
ing stress may be justly employed.

This intentional variation in unit stress applies with
greater force to the consideration of details involving
the stress of tension. Timber has a high tensile resist-
ance, and where it is certain that no other stress exists
than simple, uniformly distributed tension, the working
stress of 1800 lb. per square inch or even higher is not
excessive. However, as will be shown later, secondary

♦Basic, as distinguished from average unit stress on di-
ametral section of lag-screw, to be discussed in Chapter IV.

TIMBER FRAMIIJG

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TIMBER FkAMING

g
P

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30VnAV JO iN30il3<i

36 TIMBER FRAMING

stresses of indeterminate but nevertheless large amount
may exist, as in the ease of tension chords in trusses.
For this reason I prefer 1500 lb. per square inch as a
general limit for Douglas fir in tension.

The ever-present season-crack in timber as an incip-
ient cause of failure by shear along the fibres is sufficient
justification for limiting the working stress in shear to
150 lb. per square inch, and it is wise to decrease this
stress to 100 lb. per square inch, unless it is necessary to
exercise the strictest economy in the design of the par-
ticular detail.

Time Element as Affecting the Strength of Timber

No mention has yet been made of the effect of time
on timber; in other words, the effect on the ultimate
strength of long-continued loads. This is a quality
characteristic of timber alone, as compared to the other
materials ordinarily used in construction, and may well
be referred to as fatigue, although the term 'fatigue,' as
usually understood in a technical sense when applied to
metals refers to a different thing, namely, the effect on
the strength of the material of a great number of repeti-
tions of loadings, all within the ultimate strength;

J. B. Johnson in his 'Materials of Construction' notes
that ''Timber is entirely different from other forms of
building material in this, that it constantly yields under
heavy loads, and will finally fail under little more than
half of the load required to break it on a short time test,
such as is ordinarily given in a testing machine. R. H.
Thurston reported a few time tests on small wooden
beams, 1 in. square and 4 ft. long, in the Transactions of
the American Association for the Advancement of Sci-
ence for 1881. He found that 60% of the breaking load
would break the beams if left on some nine months* * *
the author has made about 75 tests in crushing endwise.
* * * Longleaf yellow-pine sticks, 40 in. long and 2 in.
square, were cut from a single plank, and these had sea-
soned three years in the dry. Each stick was dressed to
about 1.5 in. square, and then cut into specimens 3 in.
long. The alternate specimens were tested in compres-

TIMBER FRAMING 37

sion endwise in a testing machine, as is ordinarily done,
and the strength was found to be exceptionally uniform.
The intervening specimens were then loaded in succes-
sion, with various percentages of the average ultimate
strength of the two adjoining specimens, and these loads
left upon them until failure occurred." Mr. Johnson
concludes *'But little more than one-half the short-time
ultimate load will cause a column to fail if left upon them
permanently. Or, the ultimate strength of columns
under permianent loads is only about one-half the ulti-
mate strength of these same columns as determined by
actual tests in a testing machine."

A paper, reporting the behavior of timber under long-
continued loads was read before the American Society
for Testing Materials in June 1909 by H. D. Tiemann.
The tests were on twenty beams 2 by 2 by 40 in. of long-
leaf yellow-pine on a 36-in. span, and were made at the
Yale Forest School. An abstract of this paper appears
in Engineering News for August ^6, 1909, Vol. 62, No.
9, pages 216-217.

Mr. Tiemann concluded that **dry longleaf pine beams
may be safely loaded permanently to within at least
75% of their immediate elastic limit, provided no in-
crease in dampness occurs, and deflection will ultimately
cease (practically) under this load. No perceptible de-
flection will occur because of the time-eflfect and loads up
to within 20% of the immediate elastic limit. (* Im-
mediate' signifies 'caused by an immediate load or live
load as by an ordinary machine-test.') Loads greater
than the immediate elastic limit are dangerous, and will
generally result in rupture if continued long enough,"
The length of time necessary to cause such failure is
shown to be as small as a year or less. It is evident that
the 'critical' load, in either beams or columns, does not
correspond to the elastic limit.

The time element is the cause of the recommendation
of the American Railway Engineering Association, as
given in Table I. ' ' To compute the deflection of a beam
under long-continued loading instead of that when the

38 TIMBER FRAMING

load is first applied, only 50% of the corresponding
modulus of elasticity given in the table is to be em-
ployed."

TIMBER FRAMING 39

CHAPTER IV
Washers and Pins

Compression on Surfaces Inclined to the Direction of
Fibres. Resistance of Timber to Pressure from Cylin-
drical Metal Pins. The comparative weakness of soft
timbers like Douglas fir to compression across the fibres
has been mentioned, likewise the prevailing use of the
standard cast-iron and malleable-iron washers with bolts
and rods. This practice is altc^ether too common, and
results mainly from an unwarranted and ignorant con-
fidence on the part of the designer and constructor in
the word 'standard.' Standard details in steel con-
struction are usually the result of careful and intelli-
gent study, and long experience, but even standard de-
tails in steel will not be suitable for every case. In tim-
ber construction the term 'standard' means even less,
and many details to which this term is applied are unfit
for even the average case.

If the use of the standard 0. G. cast-iron washer and
the standard malleable-iron washer is to be condemned,
what may be said of the employment of small circular
WTOUght-steel washers with timber of low cross-bearing
strength? These latter washers were originally de-
signed for use with hard woods, yet they are constantly
employed in ordinary construction, with the result that
the washers are usually found to be drawn far into the
wood, and the resistance to further tension in the bolt
is nil.

It is not intended to make the assertion that all bolts
require large washers to fulfill their function properly.
Relatively short bolts acting in shear, as in splice joints,
may be designed with small washers with safety. All
bolts and rods, however, which act in tension should be

TIMBER FRAMING

F^O. 16. DIMENSIONS OP SPECIAL CAST-IRON WASHEBS.

provided with washers of ample area, and even those
joints in which the bolts aet principally in shear will
be greatly strengthened and their effective life increased
by the use of washers of generous size. Except for tem-
porary work, the employment of latter washers than the
standard will always be economical, when all the factors
are considered. Especially is this true where the timber
work is exposed to the weather, as the tightening of the
nuts on the bolts when washers of insufficient size have
been used will crush the timber, exposing it to the
weather with consequent decay.

TIMBER FRAMING 41

F. L. Bixby made a series of tests on the efficiency of
the standard 0. G. cast-iron washer in the course of
thesis work at the University of California in 1904.
These tests have been reported by me.* Mr. Bixby found
that the standard washer would develop from 30% to
45% only of the strength of the corresponding rod.
Special 0. G. washers were then designed with the fol-
lowing diameters :

f-in. bolt, 4.38-in. diameter washer
f-in. bolt, 3.44-iii. diameter washer
Hn. bolt, 2.96-in. diameter washer.

Tests were made on these washers as on the standard
washers. The results, rated on one-half the stress re-
quired to strip the thread of the rod in tension, showed
average efficiencies from 86% to 91% as against the low
efficiencies noted above for the standard washers. A
typical load-compression curve of Mr. Bixby 's tests is
shown in Fig. 17. The timber was Douglas fir in all
cases.

The tests show conclusively that for all washers with
tension rods pulling across the fibres of Douglas fir, the
area of the washer should be such that the unit bearing-

stress will be the same fraction of the elastic limit of the
timber for cross-bearing as the unit stress in the rod is
of the elastic limit of the rod in tension. By reference
to Fig. 14 of the preceding chapter, the average elastic
limit of green Douglas fir for compression across the
fibres is seen to be 570 lb. per sq. in.f Consequently, if
the elastic limit of steel rods is assumed to be 32,000 lb.
per sq. in., the corresponding unit working-stress for the
washers, when used with green timber, should be 285 lb.
per sq. in. This low stress may seem to be extravagant,
inasmuch as many timber trusses of Douglas fir are
giving service with much higher washer-pressure. It

^Engineering News, Vol. 71, No. 13.

tMr. Bixby does not note whether the timber of his tests was
green or air seasoned. The elastic limit as found by his tests
for bearing across the fibres was from 410 to 677 pounds per
square inch.

TIMBER FRAMING

may be definitely stated, however, that in the event
of a considerable overload, the washers would crush
the timber of the truss chords to such an extent that
excessive deflection with a probable consequent failure
of the truss would occur, while the rods would not
be dangerously overstressed. With such a truss, where
the rods are designed for 16,000 lb. per sq. in., and
the washers for a bearing pressure of 400 lb. per
sq. in., the critical strength of the truss would not be
lessened were the sizes of rods to be decreased until their
unit stress became 22,400 lb. per sq. in. In other words,
looking at the matter simply from the commercial stand-
point, the engineer may increase the unit stress in the
rods of a truss, and design the truss with smaller rods
and larger washers than are ordinarily used, and still
feel confident that'the truss is as safe as one having con-
servative unit-stresses in the rods and high bearing-
pressures under the washers.

In the framing of the Panama-Pacific International
Exposition buildings, special ribbed cast-iron washers
were designed and used. Their dimensions are shown
in Pig. 16. These washers were designed for a bearing
pressure on the timber of 350 lb. per sq. in. and a cor-
responding unit stress in the steel of 16,000 lb. per sq. in.

Compression in Inches

Fig. 17. typical washeb cubve.

TIMBER FRAMING 43

In this case, the rather high unit-stress for cross-bearing
on the timber was justified by the fact that all the work
Was of a temporary nature, and first cost, consistent with
safety, was the governing factor in the design.

To check the strength of the washers as designed, some
tests were made on the various sizes, at the University
of California. The first washer tested was a |-in. washer.
It was placed in the testing machine fitted with a short
length of bolt and nut, the nut bearing against the head
of the testing machine, and the washer bearing across
the fibres of a short block of Douglas fir. The bolt fitted
into a hole bored into the wood. Although a total load of
23,000 lb. was sustained without fracturing the washer,
the latter sank into the timber i in. under the pressure,
and the test had to be discontinued without having
reached the capacity of the washer in cross breaking-
strength. The other washers were broken by making
them bear against the ends of the fibres of the block of
Douglas fir, the head of the testing machine resting di-
rectly upon the washer. The ends of the wooden block
were not exactly parallel, so that the machine head bore
eccentrically on the washer, producing bending stresses
in the cast iron. As this condition often exists in actual
construction, the bolts being not exactly normal to the
timber, the results of the tests may be taken as typifying
the strength of such washers under the most adverse
conditions of construction.

Table II gives the results of the tests. The specimens
tested were from two different foundries ; in the table of
results the two sets are designated as A and B
washers.

The dimensions of the washers tested were all as shown
in Pig. 16, except that in the case of f-in. washers, h
If-in. instead of 1 in. and t was i^-in. instead of J.
Similarly, for the J-in. washers, t was i^-in. instead of
i-in. Due to the rather poor showing of the B foundry
washers, the dimensions of the f-in. and J-in. washers
were increased as noted in the preceding paragraph and
to the dimensions shown in Fig. 16. The effect of shrink-

44 TIMBER FRAMING

Tabi^ II

FAILURE TESTS OF EXPOSITION WASHERS

Washers from Foundry A
Size Ultimate Load-Lb. Remarks

Failure at edge of head.
Failure through head.
Failure through ribs: flaws.
Failure through ribs: flaws.
Failure through head.
Failure at edge of head.
Failure through head.
Washers from Foundry B

Failure through ribs near rim.
Failure through ribs near rim.
Failure through centre of hole.
Failure through head.
Failure through head.
Failure at edge of head.
Failure at edge of head.
Failure through head.
Flaws.
Small flaws.
Small flaw.
Small flaw.

age and flaws is much greater in the smaller sizes of
washers than in the larger sizes, consequently there
should always be more metal than might be called for
by strict observance of theoretical dimensions. Further,
in driving bolts and tightening nuts, the washers are
subject to considerable hammering, and for these reasons
any thickness of metal less than J in. is not advisable.

In Jacoby's 'Structural Details,' there is published
the result of a series of tests on cast-iron and malleable-
iron washers by H. M. Spandau, which shows that ribbed
cast-iron washers of equal strength with the 0. 6. type
may be designed, and at a saving in metal of from 30 to
50%. There are also given details of ribbed cast-iron
washers used as standard by the Atchison, Topeka &
Santa Fe and the Union Pacific railroads. Further tests
of cast-iron washers were made by L. R. Rodenhiser in
the course of thesis work at Cornell University.*

*ComeU Civil Engineer, Vol. 23, No. 2. The tests herein

20,000

28,000

16,000

17,000

33,000

23,500

18,000

20,500

7,500

9,500

16,000

14,000

25,600

17,000

23,000

26,000

23,000

TIMBER FRAMING 45

Jacoby in his * Structural Details,' page 246, allows a
25% increase in the allowable unit bearing-pressure,
when the bearing does not cover the full width of the
member. While there may exist a theoretical reason for
such increased resistance, I do not believe that actually
such a condition will exist, and I do not recommend any
such increase in bearing-pressure. On small jobs, it will
usually be found more economical to use small square
steel plate-washers instead of special cast-iron washers.
The proper sizes of plates needed to give the desired
unit bearing-pressures may be easily computed, and the
information given to the contractor by means of typical
sketches and tables. The thickness of such plates should
not be less than one-half the nominal diameter of the
threaded portion of the rod or bolt. For example, if a
unit bearing-pressure of 350 lb. per sq. in. be used, a f-
in. bolt in tension should be provided with a | by 33 by
3}-in. plate.

Compression on Surfaces Inclined to the Direction of
Fibres. The preceding discussion relates only to wash-
ers bearing normally on and across the fibres of the tim-
ber. Where the direction of pressure is inclined to the
direction of the fibres, the area for bearing need not be
so large, and it will usually be found convenient and
economical to use washers of special design, either of
cast iron or of plate steel.

The resistance to compression offered by the fibres of
the timber when the pressure is exerted at an inclination
with the direction of the fibres, namely, neither normal
nor parallel to the fibres, is a subject on which there is
some difference of opinion. H. S. Jacoby, in his 'Struc-
tural Details' develops the following formula,

n = p sin^ O -\- q cos^
in which
n = the allowable unit stress on a surface which

recorded were apparently made for washers bearing against
the ends of the fibres of the wood. For this reason, the tests
are not of practical benefit in establishing the proper size of
washer for bearing across the fibres in timber like Douglas fir.

TIMBER FRAMING

Dot and Dash Un*s rapratenr \/aco6ys K3rfi>uJa:-/7^p^f^^-'-^oo^«:
O erred Linei reorasen^ Fbrmtj/a:- /?■ ^^C/^-^}^^

INCLIMO) TO

makes an angle with the direction of fibres
p ^= the allowable unit stress against the ends of the

the fibres
q ^= the allowable unit stress on the sides of the fibres
Pig. 19 shows this equation plotted, using the values p
= 1800 lb. per sq. in., and q = 285 lb. per sq. in. These

^ TIMBER FRAMING 47

values are approximately one-half the^ values for the
elastic limits as shown on Fig. 14.

Malverd A. Howe published* the results of some tests
made to determine the allowable bearing pressure on in-
clined surfaces for various timbers. These results are
shown in the diagrams of Fig. 18, which are taken from
the article just mentioned. Mr. Howe recommends the
formula

= Q+{p-q) (w)

which equation corresponds closely to the values as de-
termined by the tests. Fig. 20 shows this formula plot-
ted, using the values p = 1800 and q ==-- 285, as before.

The allowable unit stresses for bearing on inclined
surfaces of timber, as shown by Fig. 19 and 20, differ ma-
terially. The effect on the design of a truss joint from
using the two sets of allowable pressures will be shown
in a subsequent article. Mr. Jacoby's curVe is the more
economical in material,, but the results of Mr. Howe's
tests are not to be ignored. The data available at pres-
ent should be supplemented by further tests.

Details of beveled cast-iron washers are shown in Fig.
21. Attention is called to the thickness of the base of
these washers. It is intended that the washers should be
set into the timber to the depth of the base plate. For
inclined bolts or rods, the use of these washers will give
a neat and efficient detail, and one comparatively easy
to construct.

Resistance of Timber to Pressure from a Cylindrical
Metal Pin. The bearing of a round metal pin in a
closely fitting hole in timber, as occurs in the case of a
spike, screw, or bolt with a driving fit, is a special case
of bearing on inclined planes. The subject is discussed
in Jacoby's * Structural Details,' Chapter II, Article 23,
where the following- statements are made :

**In framing, it is sometimes necessary to use metal
pins or bolts as beams in transferring stresses from one
timber to another. This involves the determination of

^Engineering News, Vol. 68, No. 5, and Vol. 68, No. 10.

TIMBER FRAMING

ft
O

IS
5

/o' ^o* JO' '^' >5cr w

-500

/o' eo' JO* -»• ^ 6cr 70' ao'

l/<7/ues of -O- in Degrees,
rormula :- r?" p s/n^ -^acos*^.

/SCO ^
4OO0^

Fig. 19. curve foe douglas fib based on jacoby's formula.

n = normal intensity on inclined planes.

p == normal Intensity on ends of fibres.

q = normal intensity across fibres.

e Wangle made by plane with direction of fibres.

^/00(A

\6CfO

/o^ 2cr JO* ^4cf ^50" 6(y 7o" «>•

. l/a/ues of -&- in Degrees.
Formula ba^ed on indentation of 0.0o

Formu/a> /?« ^^Cp-^j^j^, lA/here

Fig. 20. curve for douglas fir based on howe's formula.

n = normal intensity on inclined planes.

p = normal intensity on ends of fibres.

q = normal intensity across fibres.

e = angle made by plane with, direction of fibres.

TIMBER FRAMING 49

the pressure of the fibres of the wood upon the cylin-
drical surface of the pin. When the resultant of the
pressure is perpendicular to the fibres of the wood, the
magnitude of the resultant is the same as if the bearing
surface were the diametral section of the pin. But when
the direction of the resultant is parallel to the fibres of
the wood, the c^se is entirely different because the re-
sistance of the fibres to lateral compression is much less
than to longitudinal compression.''

The following formulas are deduced by Jacoby :

Let
P = the total safe load on the pin
h = height of the timber bearing against the pin
d = the diameter of the pin
p = safe unit stress for compression parallel to the

fibres, or for bearing on the ends of the fibres
" q = safe unit stress for compression perpendicular to

the fibres, or across the fibres
u = the unit pressure normal to the surface of the pin
= the angle which u makes with the direction of the

fibres (complement of angle used above)
0'=the special value of for which the transverse

component of 1^ = ^

**For wood having a ratio of 0.25 between the safe
unit bearing on side and on the end of the fibres, re-
spectively, 0' = 15°, and P = 0.62 hdp. An experi-
mental determination for long-leaf yellow pine, but in
which the timber was tested to its ultimate strength,
gave an average coefiScient of 0.63 for 5 tests. The speci-
mens were prevented from splitting by means of clamps.
The plane of division between the fibres crushed sidewise
was marked in every case, and gave an average value
for 0' of 15^°. When the resultant of the pressure of
the wood on a round pin is perpendicular to the fibres,
the magnitude of the safe bearing value is to be taken as
hdq, that is, the pressure is the same as if the pin were
square or rectangular in cross-section."

With the preceding theory, and in particular with the
statement as to the bearing across the grain, I am not in

50 TIMBER FRAMING

accord. The tests on spiked, screwed, and bolted joints
do not show the difference in strength between end and
cross-b6aring that the theory would indicate exists, as
•will be seen from the record of the tests to be given later.*

The following theory is presented to cover the case
now under discussion. Fig. 22 shows the case of a cylin-
drical metal pin bearing against the ends of the fibres
and across the fibres, respectively. Let the nomenclature
be as before, with the additional terms, as follows :

n = safe unit stress for compression on a surface in-

, clined to the direction of the fibres.

5' = the component of n, parallel to the direction of

fibres.
5'' = the component of n, perpendicular to direction of
fibres.

If the assumption be made that the various differen-
tial inclined surfaces will resist pressure simultaneously,
in radial lines, equivalent in amount to the allowable unit
stress, n, and if the law of variation of pressures on in-
clined surfaces be known, the. capacity of the timber or
the safe load on the pin may at once be determined.
For example, let it be assumed that the law governing
the allowable pressures .on inclined surfaces is accord-
ing to Jacoby's formula, and that

n==p sin^ O -\-q cos^

Then, referring to Fig. 22.

s = n sinQ= (p sin^ © + ^ cos^ 0) sin

The pressure s^ acts upon the differential area hr cos

e de

Therefore

Integrating and substituting the limits,

For a pin of diameter of 1 in. and length of 1 in., P
will represent the average unit stress on the diametral
section of the pin = p'

♦See also 'Tests of Duplex Hangers,' Chapter XII.

TIMBER FRAMING

or p' (unit stress in lb. per sq. in.) = § p + i g.

Similarly, for the case of Pig. 22, where the direction
of P is perpendicular to the direction of the fibres, it
may be shown that

p" ^ ^ p + § 9i where p" = average unit stress on
diametral section of the pin.

Using the values of 1800 lb. per sq. in. and 285 lb. per
sq. in, for p and q, respectively, in the above formula,

p' ^= 1295 lb. per sq. in., and

p"^ 790 lb. per sq. in.

If, on the other hand, the variation of the pressure
on inclined surfaces be taken from Howe's formula,
and the same numerical values for p and q used as be-
fore, it may be shown that

TIMBER FRAMING

Bearing ago/nsr Bearing across fibres.

£hc/3 of Fibres.

Bearins across flOttES-^' F>N.

B£Aftf/VS A&1INST £NOS ^ flBirea-^'^^-

» FiBRea- iYa Pin

BE4t»tN^ -^n^iNaT £Naa or Fianci-lV Pin.

Fll!. 23, DETAILS OF TEST^PIBCES CBED IK PIN EXPEBIMENTS.

TIMBER FRAMING

B^A/flN^ ACf^Si f/BVCS~ //f'ff/'/.

p \\MW^MMM i

Oeformafion in lnei>»:

BEA/ftf^ ASAIN9T Ends or Fibres— ^ Pin.

\ FW^

«;

•orm

/»

^ BEAR/NS AG/ktNST Sf^OS or riBRES-/i^' PiN.

D«formoTion in Inches.

54 TIMBER FRAMING

p' = 1120 lb. per sq. in., and that
p''= 675 lb. per sq. in.

In an effort to throw light on this question, I tested
a number of small blocks of Douglas fir, made with half
holes, and each fitted with a short piece of bolt having a
tight fit in the holes. Two sizes of bolts were used, one
J-in. diameter, and the other IJ-in. diameter. The tests
were made both for bearing against the ends of the
fibres and across the fibres. The details of the test pieces
are shown in Fig. 23, while the stress deformation curves
are shown in Fig. 24.

The results are not determinate. While the stress de-
formation curves for the two sizes of bolts bearing across
the fibres show the same value (approximately) for the
elastic limit, there is quite a variation in the curves for
end bearing, although the ultimate strengths are nearly
the same. The blocks with the IJ-in. pins were much
better specimens, both in respect to quality of timber
and grade of workmanship than those with the |-in. pins,
as the end cuts of the latter were not true. However,,
all uneven blocks were shimmed, and it is not probable
that the difference in strength was due altogether to
either quality of timber or workmanship.

It is probable that the diameter of pin affects the re-
sults, and that the formulae developed would hold more
closely for pins of a large diameter, since in this case the
effect of the alternate rings of spring and summer wood
would not be so marked.

The elastic limit for bearing across the grain is seen
to be approximately 1500 lb. per sq. in., while for bear-
ing against the ends of the fibres, the ultimate strength
is approximately 4000 lb. per sq. in. in the case of the
f -in. pin, and 6000 lb. per sq. in. in the case of the IJ-in.
pin, or an average of 5000 lb. per sq. in. The tests were
not suflScient in number, nor were the blocks made care-
fully enough to use the results as working data. They
do indicate, however, that the pressure of a circular
metal pin in a timber block is a function of the relative

TIMBER FRAMING

S/ip of Joint in Inches

ZSOO
ZOOO
ISOO

/ooo

5O0

-%
X

re^stances of the timber for compression against the
ends of the fibres and across the fibres.

The validity of the assumptions on which the formulae
for average unit bearing-pressures on pins are based
may be questioned. While it is not contended that the
pressure distribution on a circular pin is accurately
known, the formulae have a rational basis, and give re-
sults that are in approximate accord with such tests as
have been described, and others which will be discussed
in a subsequent chapter. The factors entering into any
theoretical solution of the question are many and com-
plex, and absolute working values may be assured only
through further tests.

Joints Framed With Shear Fins. Fig. 25 shows a
detail of a joint in which circular pins of metal or hard-
wood are used to trajismit stress. The detail is one
which was used extensively in the framing of Exposi-

56 TIMBER FRAMING

tion buildings, for splicing tension members of trusses,
fastening bolsters on columns to receive the ends of
beams, girders, knee-braces or trusses. The resistance
of such a joint involves several elements, namely,
strength of the pins in shear, resistance of the pins to
distortion, bearing resistance of the timber, both against
the ends of the fibres and across the grain, and the
strength of tlie bolts in bending, shear, and tension.

The method of framing the connection, when used as
a tension splice is as follows. Two splice-pads are spiked
on the sides of the main timber. The bolt holes are
bored, the bolts are driven and the nuts tightened. The
holes for the pins are then bored, and the pins driven.
A close fit for all pins is thus assured with a minimum
amount of labor. The joint has certain obvious ad-
vantages over some of the splice- joints which are in
general use in timber framing. It does not depend in its
action upon difficult and expensive cuts and daps of the
timber, and is much cheaper than a joint composed of
steel fish-plates with attached lugs.

A number of tests of such joints were made in 1913,
and described in detail in Engineering .News, Vol. 71,
No. 12, and Vol. 72, No. 9. The materials used were oak,
common gas pipe, full weight steel pipe, l^-in. extra
heavy pipe, solid steel, Hawaiian Ohia, Australian hick-
ory, and Australian ironbark. All pins were 2 in. diam.
The ultimate strength of the joints expressed in pounds
per linear inch of pin ranged from 1950 lb. for the gas
pipe pins to 2800 lb. for the Ohia pins. The slip of the
joints at these loads was about | in. The average load
at the apparent yield point was approximately 1800 lb.
per linear inch of pin, with a slip of 0.03 in. From the
results of the tests, it may be stated that joints of this
kind may be framed using 2-in. pins of extra heavy
steel pipe, solid steel pins, pins of Hawaiian Ohia and of
Australian iron bark, with working loads of 800 lb. per
linear inch of pin. Sufficient bolts must be provided
to take a total stress of one-half the load on the joint.

Oak and gas-pipe pins are practically worthless. The

TIMBER FRAMING 57

disadvantage of this joint is the eflfect of shrinkage,
which, if such occurs, may allow a slip. For this reason,
this type of joint is best suited to seasoned timbers.
Metal pins are to be preferred to hardwood pins, unless
the latter are thoroughly seasoned. Furthermore, the
construction is not suited to very thick joints, since the
cross-shrinkage of the wood will allow the timbers (o
spread and loosen the pins.

58 TIMBER FRAMING

CHAPTER V
Spiked, Screwed, and Bolted Joints

The strength of spiked joints in light timber framing,
and of bolted joints in heavy timber framing is, perhaps,
the most vital subject in timber design, yet it is a field
in which but few tests have been recorded. This state-
ment is particularly worthy of remark when applied to
the resistance of bolts to lateral forces, producing bend-
ing and shear in the bolts and compression on the timber.
Numerous tests have been made on the resistance of
spikes and screws to withdrawal from timber, but these
will not be discussed in the present article. Working
values for such cases will be given in the specifications
of the concluding article. For detailed information on
the results of the tests that have been made, the reader
is referred to the texts of Jacoby, Howe, and others.

Lateral Resistance of Spikes and Nails. The experi-
ments on the lateral resistance of spikes and nails are in
three sets, as far as can be determined.

The first series of tests was made by F. B. "Walker
and- C. H. Cross in 1897 at the University of Minnesota.
The results are published in the Journal of the Associa-
tion of Engineering Societies, Vol. 19, December 1897.

The second series of tests was made by H. D. Darrow
and D. W. Buchanan at Purdue University in 1898-
1899, and the results are published in the Proceedings
of the Indiana Engineering Society of 1900.

An abstract and review of both these sets of tests is
given in Jacoby 's 'Structural Details.'

The third series of tests was made in 1907 by C. K.
Morgan and F. Marish, at the Iowa State College. These
tests are described by M. I. Evinger in Bulletin No. 2,
Vol. IV, of the Engineering Experiment Station of the
Iowa State College, entitled * Holding Power of Nails in
Single Shear.'

TIMBER FRAMING

Tables III, IV, and V show the results of these tests.

•

Table III — Walker and Cross

STRENGTH

OF WIRE NAILS AT THE ELASTIC LIMIT

OF JOINT FOR

WHITE AND

NORWAY PINE

Size of

Strength of

Size of

Strength, of

Nail

Nail, Lb.

Nail

Nail, Lb.

6D ....

SOD

226

8D . . . .

40D

275

lOD . . . .

112

50D

342

16D . . . .

112

60D

362

20D . . . .

218

SOD

500

In the experiments of Walker and Cross, the timber
used was white pine, Norway pine, and oak, with average
compressive strengths of 4840, 5820, and 6600 lb. per
sq. in., respectively. The timber cleats in the tests were
surfaced. This surfacing must have influenced the
strength of the joints at the lower loads, and possibly at
loads up to the elastic limit, as there is considerable
friction in a tightly spiked joint with rough timbers.

Table IV — Darrow and BtLchanan

Kind of

Nail Size
Common 2iD.
32).
4D.
6D.
8D.
lOD.
16D.
20D.
40D.
60D.
Finish 4D.
6D.
SD.
lOD.
12D.
Fence SD.
lOD.
Fine 3D.

((

ULTIMAT

E FAILURE

OF JOINT FOB

[ne and

OAK

-Yellow Pine-^

, Oak ^

Wire

Cut

Wire

Cut

• • •

130

• • •

160

• • •

180

184

289

198

213

211

344

240

317

314

429

361

427

454

573

724

932

762

822

855

1079

891

1066

930

1112

1350

1631

1450

1360

1745

1874

2000

1860

1770

• • • •

106

163

186

262

216

209

299

273

264

282

359

405

537

451

583

• • • •

498

• • •

637

• • • •

690

686

709

780

855

912

1030

1092

121

• • •

164

• • • •

60 TIMBER FRAMING

The timber of the tests of Darrow and Buchanan
was well seasoned yellow pine and oak, and the average
compressive strengths are given as 7000 and 10,200 lb.
per sq. in. for the two sticks of yellow pine used and
5300 lb. per sq. in. for the oak.

The strength of the timber used in the tests of Morgan
and Marish, white pine, yellow pine, spruce, fir, and oak,
is not given, but the statement is made that the lumber
was obtained from the stock piles in a local lumber yard,
and that it was not thoroughly seasoned at the time the
experiments were started.

At first inspection, .the results of the different tests
appear to differ by so great an amount that any de-
ductions are worthless. However, when all the factors
that enter into the tests are considered, the variation in
the strength of the different joints can be largely ex-
plained. The ultimate strengths of the blocks of timber
in the three sets of tests varied considerably. The
method of measuring the slip of the joints was rough in
all cases. In at least two of the series of tests, the slip
of the joint was measured on one side of the joint alone.
This method can never give the true values of the de-
formations, and erratic results are to be expected. My
experience in testing timber joints has been that if meas-
urements are made carefully, the platted load-deforma-
tion curves will be remarkably smooth. Also, in order
to obtain a curve that will express fairly accurately the
relation of load to deformation, the slips of the joints
must be measured closer than ^ inch.

For purposes of comparison of the results of three sets
of tests Table VI has been prepared, which gives the re-
sistance of the various sizes of nails at the elastic limit
of the joints. To bring all experiments to the same
basis, the values of "Walker and Cross, which are for
white and Norway pine, have been increased by the
factor 1.45. This factor represents the ratio of the
average or effective unit bearing pressures of yellow
pine to white and Norway pine at the elastic limit
for the case of a round metal pin bearing against the

TIMBER FRAMING

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62 TIMBER FRAMING

ends of the fibres in a close fitting hole in accordance
with the formula developed in the preceding chapter.
Similarly, the values of the tests of Darrow and Bu-
chanan, which are for the ultimate strength of the
joints, have been reduced by the factor 0.41, which
represents the i;atio of the strength of the joints at the
elastic limit to that at the ultimate strength multiplied
by the ratio of the average unit bearing pressures at the
elastic limit for the timbers of the two tests.

The values in the fourth column were selected as the
average values for a slip of joint of ^ in., and were
taken from the curves in the bulletin of the tests. In all
three sets of experiments, it was found that the elastic
limit of the joints corresponded to a slip of the joints of
approximately y^ inch.

A. W. Muenster, in the Journal of the Association of
Engineering Societies, discussing the tests of Walker
and Cross, proposed expressing the shearing resistance
of a wire nail at the elastic limit of the joint' by the
formula Cd^, C being a coefficient dependent on the tim-
ber, and d being the diameter of the nail. For white
and Norway pine, C would be approximately 5500.
Further he proposed to take the working values at about
60% of the strength at the elastic limit, or a coefficient
of 3300.

Table VI

LATERAL STRENGTH OF WIRE NAILS AT ELASTIC LIMIT OF JOINT IN

YELLOW PINE

Nail / Strength in Pounds x

Walker Darrow and Morgan

Size of and Cross Buchanan and Marish

8D 128 148 120

lOD 163 296 175

12D ... 250

16D 163 350 260

20D 316 . 382 375

30Z> 328 ... 540

40D 400 595 480

50D 495 ... €30

mD 525 764 775

SOD 725

• •

TIMBER FRAMING 63

W. K. Hatt, in presenting the results of the experi-
ments of Darrow and Buchanan to the Indiana Engi-
neering Society, recommended that the safe working
values of a wire nail in yellow pine be expressed by the
formula :

S = 9.5 D, where D is the * penny' weight of the nail.

Table VII gives the working values for the strength of
wire nails, (1) by Muenster's rule, but using a co-
efficient, C = 4000, and (2) by the rules 8 = 8D. The
value of C =4000 is obtained by applying a safety factor
of two to the value (7 = 5500, and multiplying the re-
sult by 1.45, as explained previously, to bring the co-
efficient to the basis of yellow pine.

Mr. Hatt states that the rule which he proposes should
not apply to nails heavier than 20Z).

From consideration of all the tests, I recommend
using for the case of nails in Douglas fir the formula
S = 4000 d^, or for nails up to and including 20D nails
the values of the third column of the preceding table,
which represent 8 = 8Z>.

Since the preceding discussion of nailed joints was
written, there has been published a partial report of
some tests on the resistance to lateral forces of several

•Table VII

SAFE WOBKING VALUES FOB THE LATEBAL STBENGTH OF WIBE NAILS

IN YELLOW PINE

Ck)mparison of Results by Methods of Muenster and Hatt, with

Modified Coefficients

•

Size of Nail 8 = 4000 d= 8 = SD

6D 53 lb. 48 lb.

SD 62 " 64 "

lOD 88 " 80 "

12D 88 " 96 "

16D : 110 " 128 "

20Z> 165 " 160 "

302) 194 " 240 "

40D 226 " 320 "

501) 268 " 400 "

60D 322 " 480 "

SOD 364 " 640 "

(34 TIMBER FRAMING

sizes of common wire nails, made at the Forest Service
Laboratory, Madison, Wisconsin.*

The timber used was thoroughly air-dry (average
moisture 13.8%) long-leaf yellow pine. A deflectometer
was used to measure the slip of the joints. For this
reason, I feel that this set of tests should be given espe-
cial attention, and the complete report, when published,
carefully studied. In framing the test-joints, A-i^^- holes
were bored in the cleats for each nail, but the blocks were
not bored. No nails were placed in checks, knots, or other
defects. Table VIII gives the results of the tests. In the
fifth column, I have inserted the value of C in the for-
mula 8 = Cd^. The average value of C is seen to be
about 7000. Thus the working values for the nails of
Table VII are but slightly over one-half the elastic limit
as found by these tests.

Mr. Wilson makes the following comments on the tests :

1. The elastic limit is well defined, and at a compara-
tively small deformation.

2. No definite relation between size of nails and de-
formation at the elastic limit is apparent.

3. It seems probable that the load per nail is inde-
pendent of the number of nails in the joint.

4. In efficiency per unit-weight, the smaller nails in
general seem to have the advantage.

Until the complete set of tests has been made and pub-
lished, it is unwise to. make definite comments. The slip
at the elastic limit is small, or, expressed in another
manner, the elastc limit of the nail occurs at a low de-
formation, in fact, at a deformation of only about one-
half that found by the other investigators quoted. It is
possible that the same point on the stress-deformation
curve has not been taken by the various investigators.

This discussion of the strength of nailed joints is not
complete without mention of the tests of J. C. Stevens,
consulting engineer, of Portland, Oregon, noted in the

♦Tests Made to Determine Lateral Resistance of Wire Nails,'
by Thomas R. C. Wilson, Engineer in Forest Products, Forest
Service, Engineering Record, Vol, 75, No. 8, February 24, 1917.

TIMBER FRAMING 65

Engineering Record of January 13, 1917, Vol. 75, No. 2.
The tests are only briefly described. Ordinary 2 by 12-
in. yellow-pine planks were nailed to fir sills and then
sheared off. From the results of these tests a working
load of 210 lb. for a 16D nail, and 250 lb. for a 20D nail
were used in the heel-plates of wooden flumes. In the
light of other tests these working values are high, but
within the elastic limit.

The additional matter just given leads me to reafiirm
my recommendation that the working strength of wire
nails in lateral shear, when used with Douglas fir, be
taken in accordance with the formula 8 = 4000 D^.

Common Wood-Screws. The strength of ordinary
wood-screws in single shear was investigated as thesis
work in Cornell University by Andrew Kolberk and
Milton Bimbaum and discussed by them in the Cornell
Civil Engineer, Vol. 22, No. 2, November 1913. '*The
screws were the ordinary cut, flat-head screws, made by
the American Screw Co. of Providence, Rhode Island.''
The timber employed in the joints was cypress, yellow

Table VIII

Lateral Resistance of Nails in Aib-Dby Long-Leaf Pine

Test value of
Load in lb. per nail Slip in inches C at elastic
Size At At At At limit in

of elastic ultimate elastic ultimate formula

nail limit strength limit strength 8 = Cd^

*30D 355 779 0.021 0.54 7340

•40D 394 845 0.026 0.58 6950

♦50D 450 1261 0.019 0.69 6710

*60D 422 1144 0.018 0.70 5240

t30D 333 783 0.034 0.88 6880

t40D 389 1125 0.028 1.08 6860

t50D 544 1615 0.036 1.18 8120

t60D 589 1644 0.039 1.32 7300

♦Timbers with grain parallel, load parallel to grain.

tTimbers with grain at right angles, load parallel to piece
receiving points of nails.

In the first series each value is based on four tests of 3-nail
and four tests of 6-nail joints. In the second series each value
is based on two sets of 3-nail and two tests of 6-nail joints.
The average value of C at the elastic limit for all tests is 6920.

TIMBER FRAMING

pine, and red oak. The average strength in end bearing
of the three timbers was found to be 4980, 7580, and 8440
lb. per sq. in., respectively. The thickness of the timber
cleats varied with the length of screw used, but the test
piece was always arranged to make the screws act in
single shear. It was found by experiment that screws
could not be driven closer than 2^ in. from the edge per-
pendicular to the direction of fibres without danger of
splitting the wood. It was also often found impossible to
drive the screws without previously boring holes, and in
such cases the size of hole was made equal to the diam-
eter of the screw at the root of thread. In oak, and in
the case of the large screws in yellow pine, separate holes
had to be bored for the shank and for the threaded por-
tion of the screw. A hole was also bored for the head of
the screw, thus bringing it flush with the surface of the
wood. Every part of the screw was thus brought into
action. The procedure in testing was to measure the
force at each gV-iii- slip up to a maximum slip of -^ in.
As this slip is more than would be allowed in practice,
it was not thought necessary to carry the tests to the
ultimate capacity of the joints.

B^G. 26. FORCES ACTING ON SCREW IN JOINT.

a = thickness of side piece.

6 = depth of penetration Into centre piece.

Fig. 26 shows in a general way the forces acting on a
screw in a joint. The screw in such a joint was but
slightly deformed in cypress, the wood being so soft as
to crush readily without bending the screw. In yellow
pine and oak, however, the screws were bent in the
characteristic reverse curve typical of nails and screws.

TIMBER FRAMING 67

Curves in which the load per screw in pounds was
platted against the slip in inches showed the following
results : In cypress, while a joint having a thinner side-
piece might be stronger than one having a thicker side-
piece during the first increments of slip, it did not con-
tinue to be stronger during the last few increments of
slip. In yellow pine and oak, however, the joint with
the thinner side-piece, once being the stronger, continued
so during the entire test.

Investigation relative to the determination of the
proper proportion of the length of screw to the thickness
of side-piece, showed that in general, the joint with the
thin side-piece was the strongest. A f-in. side-piece and
2i-in. screw gives a ratio of 0.3 between side-piece and

Table IX

LATEBAL STBENGTH OF SCBEWS IN YELLOW PINE AND OAK COBBE-
SPONDINQ TO A SLIP OF JOINT OF V«2 INCH

YeUow Pine

Length, of

Gauge of

Thickness of

/ ]

Load in Lb. ^

Screw, In.

Screw

Side Piece

V«-in.

slip

rin. slip

758
924

1002
1094

477
557

657
750

437

585

586

696

590
770
780

810
936
960

453
563
526

580
702
631

1
i

414
520
508

530
622
610

606
632

797
802

TIMBER FRAMING

Length of Gauge of

Screw, in. Screw

2 18

2 16

Thickness of
Side Piece

-Load in Lb.-

A-in. slip ^-in.Blip

472 643

458 625

382

5S1

410

601

322

514

405

573

346

432

460

537

451

501

320

365

394

504

417

509

317

360

394

473

404

488

Red Oak

707

1144

933

1308

543

754

710

932

580

842

750

1154

442

634

557

809

■

627

874

700

912

513

681

600

767

456

639

529

632

514

601

603

802

383
523

416
590

TIMBER FRAMING 69

screw; a l-in. side-piece and 2i-in. screw has a ratio of
0.4. For a 2-in. screw, the J-in. side-piece gave the
strongest joint, especially with the screws of the smaller
gauges. This is a proportion of 0.375. For the l^-in.
screw, the strongest joint was that with a side-piece f -in.
thick, or a proportion of 0.4. Thus it would seem that a
side piece of about 0.4 of the length of the screw will
give the strongest joint. Or, conversely^ to obtain the
strongest joint, the screws should have a length of ap-
proximately 2i times the thickness of side-pieces. In
the joints with yellow pine and oak timber, it was found
that the strength of the joint varied as the square root of
the penetration of the screws into the centre timber,
while for the cypress joints the proportion varied as the
cube root of the penetration into the centre timber.

Table IX gives the strength of the screws at a slip
of joint of ^ and -^ in. for the joints framed with yel-
low pine and oak. This table has been condensed from
the results published in the Cornell Civil Engineer, the
values for slips above ^ in. being omitted for the reason
that this slip has been found to represent the elastic
limit in the case of nailed joints, and the table may thus
be compared with those of the nailed joints.

**In the case of yellow pine it was found that the
strength of the joint varied with its weight, or specific
gravity. The heavier joints invariably gave the larger
results. In order to reduce all the values to a common
standard, the weight of a joint of average specific grav-
ity was computed for each size of side-piece. The re-
sults were compared with each other and corresponding
differences in strength and weight noted. From the
averages of these values the mean difference in weight
of joint was found. Joints were reduced in this manner
to the strength corresponding to the weight of a joint of
average specific gravity. A difference of 0.1 lb. in the
weight of a joint was found to make a difference of 10
lb. at the ^-in. slip and 30 lb. at the ^-in. slip.

Table X gives the average lateral resistance per
screw for the yellow pine and oak joints at the assumed

TIMBER FRAMING

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TIMBER FRAMING

elastic limit or for a slip of joint of ^ in. The table
also shows the relation of these loads to values obtained
by multiplying the square of the diameter of the screw
by an arbitrary factor of 8750, and also shows the re-
lation of the loads to the values obtained by multi-
plying the gauge of the screw by an arbitrary factor
of 45. With the exception of the No. 18 gauge
screws, the table shows that the arbitrary formula, 8 ==

//z^e'-O'S^P/ore

jzrt-^ TrPL B

7VP5£- C.

mtt^ i:l|

nj jlr 'k e ; 1
ll 1 < 1

iiiiiiin

i..,i..r

Fig. 27. types of lag-scbewed joints.

Type A is an 8 by 10-in. or 12-in. timber, 1 ft. 5i iQ. long.

Type B is composed of two li to 2 in. by 8 in. pieces, 1 ft. 1 in.
long, and an 8 by 8-in. timber, 1 ft. 1 in. long.

Type O is composed of two If to 2 in. by 8-in. pieces, 1 ft. long,
and an 8 by 12-in. timber, 1 ft. 2 in. long.

S//p in /nches:

Fig. 28. load curves fob lag-scbewed joints with steel
plates. cubves show avebage fob one lag sciffiw.

72 TIMBER FRAMING

8750 d^, where S = the resistance of the screw at a slip
of joint of ^ in., and d = the diameter of the screw,
holds fairly well, and may be adopted for determining
working loads. The difference between the actual re-
sistances as shown for yellow pine and oak is small, and
in conformity with the relative properties of the two
timbers.

For determining the diameters of the screws, the
standard rule

d = 0.0578 + 0.01316 G
was used, where d = the diameter of the screw, and =
the gauge of the screw.

The working loads for Douglas fir may be taken as
shown in Table XI, whose values have been computed
from the formula

8 = 4375 d^

Table XI

SAFE LATERAL RESISTANCE OF COMMON WOOD-SCBEWS WITH

DOUGLAS FIB

Gauge of Safe Lateral Gauge of Safe Lateral

Screw Resistance Screw Resistance

12 205 lb. 20 450 "

14 256 " 22 529 "

16 315 " 24 615 "

18 380 "

Lateral Resistance of Lag Screws. The action of a
lag screw when subjected to lateral shear and flexure
is similar to that of a nail, but more like that of a com-
mon wood-screw. Since the diameter of a lag screw in
proportion to its length is considerably greater than
that of a nail, the bending is less marked, and the re-
sistance is dependent almost wholly upon the bearing
strength of the timber.

The only data on the resistance of lag screws is given
in Kidder's 'Architects and Builders Pocket Book,' and
in Thayer's 'Structural Design.' In the latter volume,
the lag screw is treated as if it were a bolt, when com-
puting its resistance. Kidder shows a detail of the end
joint of a scissors truss, in which a thin steel plate is
lagged to the truss chord. When used in this manner

TIMBER FRAMING 73

with Douglas fir the value of a J by 4i-in. lag screw with
a steel plate of ^ in. minimum thickness is given at
2100 lb., similarly, the value of a f by 5 in. lag screw
is given at 2800 lb. There is no statement made as to
the basis on which these values are selected. Thayer's
figures for the same conditions are 1125 lb. and 1575 lb.,
respectively.

In an article, already mentioned,* I published the
results of some tests on lag-screwed joints made at the
University of California in 1913. In these tests, f by
4i in. lag screws were used to fasten 2 in. planking to
12 in. blocks of timber, with the screws bearing against
the fibres of the planking, and across the fibres of the
blocks. The results are given in Table XII. The test
joints were made by lagging two 2 by 8-in. planks to
the sides of a 12 by 12-in. timber block.

The failure in joints a, c, e, and / was due to the screws
splitting the side of the main timber. In joints b and d,
the attached piece was sheared by the screws.

No detail measurement of slip was made, but it was
noted that the first large movement appreciable to the
eye occurred at a load of approximately 2400 lb., and
was between ^ and i inch.

More recently, I made a series of tests on lag-screwed
and bolted joints at the Panama-Pacific International

Table XII

1913 TESTS OF LAG-SCBEWED JOINTS

First Slip Ultimate Load

Total Load per Per

Load, Screw, Total Screw,

Joint Lb. Lb. Lb. Lb.

o(6 screws) 11,000 1,833 23,000 3,833

6(6 screws) 16,000 2,667 30,000 5,000

c(8 screws) 21,000 2,625 30,000 3,750

d(6 screws) 16,000 2,667 31,000 5,167

e(8 screws) 20,000 2,500 30,000 3,750

/(8 screws) 18,000 2,250 29,000 3,625

Average value of one screw 2,423 4,187

*Enffvneerinff News^ Vol. 71, No. 13.

TIMBER FRAMING

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TIMBER FRAMING 75

Exposition, through the courtesy of the Tinius Olsen
Testing Machine Co. of Philadelphia. These tests have
been described in a recent article.* In these tests, careful
measurements of the slips of the joints were made, corre-
sponding to constant increments of load, in addition to
making autographic records of the load-deformation
curves through the medium of the autograph attachment
on the machine.

The lag-screwed joints consisted of four joints in
which a J-in. steel plate was fastened to a timber block,
and ten joints in which wooden plates varying from IJ
to 2 in. thick were lagged to 8 by 8-in. and 8 by 12-in.
blocks. The detail makeup of these joints is shown in
Fig. 27, a, &, and c. Pig. 28 shows the curves for the
joints with steel plates, while Pig. 29 shows the typical
curves for the all-timber joints. The results of the tests
are given in Table XIII.

Reference to Pig. 28 indicates that the working values
for lag screws given by Kidder correspond to slips of
joint of 0.08 in. for the % by 4i-in. screws and 0.12 in.
for the J by 5-in. screws.

With respect to the safe working values to be adopted,
there may be some difference of opinion. In the case of
the joints with steel plates, the first break in the load-
slip curve occurs at a total load on the joint of about
6000 lb. With the all-timber joints, there is also a slight
break in the curves at about 8000 lb., but a marked break
at a total load on the joint of approximately 23,000 lb.
for the l-in. screws and 18,000 lb. for the f-in. screws.
It was found for the steel-plate joints that when the load
was removed at about 8000 lb. the joint did not fully re-
cover its slip, although the remaining set was probably
due to the initial adjustment of the joint.

In conformity with the tests on nailed joints, it is
recommended that the working values for lag screws be
taken at one-half the loads producing a slip of ^ in.
The safe lateral resistances would then be as follows :

*Engineering News, July 20, 1916.

TIMBER FRAMING

Lb. per Screw

Metal plate tagged to timber, } by 4Hn- 1^ screw 1030

Metal plate lagged to timber, | by 6-in. lag screw 1200

Timber planking lagged to timber, j by *)-ln. lag screw . , . 900
Timber planking l^ged to timber, i by E-ln. lag screw 1050

There appeared to be do reduction in stiffness for those
joints in which the lag screws bore across the fibres of the
timber in the centre block during the first portion of
the tests.

In the Proceedings of the American Railway and
Maintenance of Way Association, Vol. 10, 1909, there is
published a method for determining the safe lateral re-

o\^\ aUo\ d60\ ciBo\ /.bo \ /la ? I /Uo I ,\6o \

Flo. 29.

S///0 in inches-

D CURVES FOB LAO-SCBBWS

-where

sistance of a track spike. The discussion is also given in
Jaeoby'a 'Structural Details.' Referring to Pig. 30.

■ P=

4Z,-)- 6e

P^safe lateral resistance of the spike.
p = masimum safe unit bearing stress of the timber
on the ^ike.

6 =^ diameter or side of the spike.
h and e, as shown in the figure.

TIMBER FRAMING ^^

Applying this formula to the joints of the tests hav-
ing a steel plate lagged to the timber, assuming the
dianfeter of the lag screw to be constant throughout its
length and equal to its nominal diameter and using a
limiting bearing-stress of 1300 lb. per sq. in. on the
timber, there is found for the case of the |-in. lag screws
with a i-in. plate, a safe lateral resistance of 770 lb. per
lag screw. Similarly for a J by 5-in. lag screw, the safe
lateral resistance is found to be 1040 lb. The maximum
flexural stress in the lag screw is 10,770 lb. per sq. in. for
the f-in. lag screw and 13,400 lb. per sq. in. for the f-in.
lag screw. These values do not differ greatly from those
of the tests.

In Table XIV, the recommended values of Kidder and
Thayer, the values as derived from the formula just
given and those recommended by this article are given.

Lateral Resistance of Bolts. The values for the
strength of bolted joints as computed by the methods and
formulae given by the various text and hand books differ
widely. Probably in no other instance in structural
engineering is there a greater discrepancy between the
results of the methods of design of tlie various authori-
ties than in the case of the design of a bolted joint.
Again, although this connection is one used constantly in
ordinary construction, both temporary and permanent,
apparently no tests have ever been made on the actual
strength of bolted timber- joints from which working
values might be selected with the assurance of reasonable
accuracy. In fact, the only tests which are on record, as

Table XIV

OOMPABISON OF SAFE LATEBAL RESISTANCE OF LAO SCREWS WHEN
USED TO FASTEN A METAL PLATE TO TIMBER AS GIVEN BY VA-
RIOUS AUTHORITIES.

f by 4i-in. i by 5-in.

Lag Screw, Lag Screw,

Authority Lb. Lb.

Kidder 2100 2800

Thayer 1125 1575

Theoretical formula 770 1040

Tests in present chapter 1030 1200

TIMBER FRAMING

far as I have been able to determine, are those of E. E.
Adams, first published in the California Journal of
Technology of the University of California in 1904, and
later discussed by me.*

The present discussion will review the common meth-
ods of designing bolted joints, and compare the results
with those of the tests on 24 bolted joints which I made
in 1915 in conjunction with the tests on lag-screwed
joints previously described. For a detailed description
of these tests, the reader is referred to Engineering News
for July 20, 1916.

Current Methods of Design. Fig. 31 represents a
splice in the tension chord of a truss. The thickness of

Fig. 30. diagram of stress on tback-spike.

the chord is 2L, while the thickness of either splice is L.
The total stress in the chord is P, of which it is assumed
that each splice-pad takes half.

Kidder in his * Architects and Builders Pocket Book'
states with regard to the design of bolted timber-joints,
''When the pieces joined together ar^ not more tJian
two inches thick, so that they can be tightly drawn
together, thereby producing a good deal of resistance
from friction, the bolts may be considered as rivets, and
proportioned for shearing and bearing only, the bending

*Engineering News, Vol. 71, No. 13. Since writing this
chapter, I am informed that Harold A. Thomas of Rose Poly-
technic, Terre Haute, Indiana, has made a series of tests on
bolted timber joints which are to be published in the technical
press.

TIMBER FRAMING 79

moment being neglected. When the pieces of wood are
more than two inches thick, the bolts should be propor-
tioned for shearing, bearing, and flexure." Where bend-
ing is considered, Kidder recommends that the bending
moment, M, be taken as 1/12 P times the distance be-
tween the centres of the splice pads. In Fig. 32, M would

nzE

1^ — * *-

Fig. 31. splice in tension choed of truss.

then be equal to 1/12 P times 3L or \ PL, H. S. Jacoby
in his * Structural Details' shows the bending moment to
be equal to J PL when the pressure against the bolt is
considered to be uniform over its entire length. M. A.
Howe in *A Treatise on Wood Trusses' uses the same
formula as Jacoby, ov M = ^ PL, W. J. Douglas in
Merriman's * American Civil Engineers Handbook' dis-
regards the bending in bolts, where the pieces joined are
less than three inches thick, otherwise the bolt is con-
sidered as a restrained beam, and computed as recom-
mended by Kidder. The working stress used for bend-
ing in the bolts in the preceding formulae is 22,500 lb.
per sq. in. Horace R. Thayer* in 'Structural Design,'
Vol. 1, recognizes the necessity of a method of computa-
tion for bolt sizes more consistent with practice. He
recommends that the "pressure on the bolt be considered
as concentrated on the portions of the bolt immediately
adjacent to the contact faces of the timbers; the pressure
to be considered uniform over the length on which it
exists. This length a is to be such, that the resultant
bending in the bolt will just equal the flexural or shear-
ing strength of the bolt.

Thus referring to Fig. 32, which shows the assumed
distribution of load on the bolt, where

fif = maximum allowable flexural unit stress in bolt.

♦See also Engineering Neios, Vol. 71, No. 17, p. 923.

TIMBER FRAMING

^ = allowable unit bearing stress against ends of
fibres of timbers.

d = diameter of bolt.

Assuming that the shear on the bolt need not be con-
sidered, and that the bending on the bolt alone need be

considered, a =

, 0' = d f ^

^8 \^
'62B J

For Douglas fir with iron bolts, Mr. Thayer uses S
= 12,000 lb. per sq. in., and B = 1250 lb. per sq. in.

tsy<[

Fig. 32. diagbam of stbess on bolt.

For steel bolts, S is increased to 16,000 lb. per square
inch.

In Table XV is shown the working strength of
such joints, as determined by the methods of Jacoby,
Howe, Kidder, Merriman, and Thayer, as well as from
curves taken from actual tests and from a formula de-
rived by myself. In the case of all these authorities, the
lowest and highest values are shown ; namely, the values
for joints with 2 in. and 2J-in. splice-pads and those
with 6-in. splice-pads. This table emphasizes the wide
variation in the prescribed practice of designing bolts.

There is this radical difference between Thayer's
method and. the others quoted, namely, using Thayer's
analysis the value of L in the joint may be in excess of
a without decreasing the strength of the joint. In other
words, the joints with a 12-in. centre timber and 6-in.
side-timbers will have the same strength as the joints
with side-timbers of a width equal to the distance (k.

TIMBER FRAMINQ

TIMBER. FRAMING

Description of Teste of Bolted Joints. As noted
p^eviousl3^, the 1915 series of tests embraced twenty-
four bolted joints, the bolts varying in size from f to 1
in. diam. Fig. 33a and b shows the details of all the
joints tested. The tests were compressive tests. The
washers were standard cut-steel pressed washers through-
out. Before commencing the tests, all nuts were loosened
so that friction would not aflEect the results. This was to
approximate the condition of a joint after shrinkage has
taken place. Careful measurements were made of the
slip of the joints, in addition to recording the load-de-
formation curves by means of the autographic attach-
ment of the testing machine.

Table XVI gives a summary of the results of the tests.
A careful study of the curves representing the relation
of load to slip has led to the conclusion that for the test
joints having the same number and diameter of bolts,
except for those joints in which the bolts bear across the

Trpc A

TypE

Fig. 33. details of joints tested in 1915. sizes of timbers

and bolts as listed in table xiv.

fibres of the centre timber, the load-slip or load-deforma-
tion curve is practically a constant. In other words, for
the limits of the tests, it appears that the strength of a
bolted joint depends upon the number and size of the
bolts, and is nearly independent of the thickness of the
timbers forming the joint.' Fig. 34 shows the curves
representing the relation between slips and loads for
each set of joints having bolts of the same diameter.
These curves are drawn as an average of the curves of
the individual tests.

For those joints where the bolts have bearing across
the fibres of the centre timber, the load-slip curves ap-
pear to coincide with those of the end-bearing joints up
to a total load of approximately 30,000 lb., or up to

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TIMBER FRAMING

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TIMBER FRAMING 85

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TIMBER FRAMING

FiQ. 3

3 AFTER TESTING.

bolt is as shown in Pig. 36, namely, tringular in shape,
concentrated near the contact surfaces and of such ar-
rangement as to produce the condition of a restrained
beam with some point of contraflexure at each contact
surface.

In order to conform to these conditions b must equal
2a. Then B' = J B.

(7l>

,^A

i^g^

'nn

.S-^L.

Fl<3. 36. ASSOUED DISTBmUTION OF I«AD OH BOLT.

TIMBER FRAMING 87

The distribution of bending moment will be as shown,
there being two equal maximum moments of amount,

itf = 2/27 PL where Z = a + & and must be less or
equal to t.

Under the assumption that beyond a certain minimum
value of t, or width of side piece, the strength of the
joint is independent of the length of bolt required, the
length I may be found, such that the moment resulting
from the load on this length of bolt will just equal the
flexural strength of the bolt. This method is similar to
that of Thayer, the diflference being in the assumed dis-
tribution of load, the resulting moment and the com-
puted strength of bolt.

The maximum moment will occur at a point of zero
shear, or at a point distant i b = a either side of the
contact face.

Let the moment be computed at point X

^■^> CM'-sO-j^'

Bid

^ 4 ^ Bid y 7 ^ L dRj2. ^^^^

j2^ 7rd^S.27
32Bd

With the safe stresses for flexure of bolts and bear-
ing for the timber, the strength of the bolts is easily
determined ; and for all thicknesses of side pieces in ex-
cess of the limiting value I, the strength will be constant.
For thicknesses of side pieces under I, the safe strength
of the joint would be a question of bearing of the bolts
against the timber, if the pressure distribution be as-
sumed to remain constant, resulting in a decrease in the
strength of the joint. The results of the tests do not
bear out this assumption. It is believed that in joints
where the thickness of side timbers is less than the
limiting value I, the pressure-distribution diagram, while
holding to the general triangular shape, changes in its

88 TIMBER FRAMING

relative dimensions a and b, within the limits, where
a = and a = it. Further, it is held that the ratio of
a to t is always such, that the resulting bending moment
on the bolt bears the same relation to the capacity of the
bolt in bending as the maximum intensity of pressure on
the timber bears to the resistance of the timber in
bearing:

Following out these assumptions, the strength of the
test joints with 2-in., 2J-in., and 6-in. side timbers have
been calculated, using 16,000 lb. per sq. in. as the flexural
stress in the bolts, and 1300 lb. per. sq. in. as the limiting
unit-bearing pressure on the timber. These computa-
tions were made by the use of the diagram shown in Fig.
37. The results are entered in the bottom line of Table
XV, and are seen to coincide approximately with the
values taken from the test curves.

The preceding discussion of the theory of bolted joints
tacitly assumes the case of comparatively thick timbers
with bolts of small diameters. As the ratio of the di-
ameter of bolt to the thickness of splice-pad is increased,
the pressure-distribution on the bolt will change from
the triangular shape to a trapezoidal shape, and finally,
for the case of short thick bolts of great stiffness, the
pressure-distribution will become uniform along the
length of the bolt. In other words, the limit in this
direction will be the case where the strength of the joint
is determined by crushing of the timber. Obviously, the
joints tested do not fall within this class. The strength
of bolted joints where the pressure distribution is trape-
zoidal may be found by diagrams constructed along the
lines, of Fig. 37.

From the results of the studies on bolted joints, the
safe working loads for bolts in double shear, having all
end bearing, may be taken as shown in Table XV. For
bolts having side bearing on the timber the safe loads
may be taken as two-thirds the values of the table.
While the tests did not consider bolts in single shear
alone, working loads for this condition may be taken at
one-half the values of the table.

TIMBER FRAHING

II It,- ill

Ij j j 8 i 8 i i i U J

u»

, ^ s

..itV

V. .V:'

jt j-y I

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y i. ■■■ " » !

. j^ J

'% A. :: ::: ,_

/ , f. » • 5 !.

f J ,<-- ; •

4/,^ S +

t^ ' ::: :.

It is hardly necessary to state that further tests are
required to confirm the theory presented, but in the
absence of sueh tests, the method advocated is believed
to be reasonable, and to give results that are safe.

90 TIMBER FRAMING

CHAPTER VI
End Joints of Trasses

The timber truss, with its details of joints, forms
perhaps, the most important and interesting subject, of
timber design. A roof truss of the Howe type is the
simplest form of truss, as far as the calculation of
stresses is concerned. Consequently we find in practice
that the main sections of such roof trusses are usually
of the proper size, but that the details, particularly the
end joints, are often quite deficient in strength to de-
velop the calculated stresses. Thus the truss that the
designer imagined confidently had a large safety factor,
may actually be not far from failure. The live load
used in the design is in many instances the saving
factor of the truss.

It is not the intention in what follows to treat of the
solution of' stress diagrams in roof trusses. In such
mathematical discussion of primary or secondary
stresses as is given, th£ reader's knowledge of the de-
rivation and the proper application of the different
formulae used, is assumed. In other words, it is taken
for granted that he has a working knowledge of struc-
tural mechanics. It is proposed to discuss the different
details of a typical roof truss from the standpoint of
both theoretical and practical efficiency.

The number of types of end joints that may be used
might almost be said to be legion, and no attempt will
be made to list or describe them all. No one type can
be specified that will be applicable to all trusses. The
individual case must govern the selection. For ex-
ample, a roof truss that rests upon a masonry wall will
usually require an altogether different type of end
joint from a truss which frames into a post. The con-
sideration of clearance may decide whether a shoe is

TIMBER FRAMING 91

necessary, or whether the batter post of the truss may
simply dap into the lower chord and be bolted thereto.
Certain simple end-details can be used only with
trusses of small chord-stresses. Again, the considera-
tion of wind action in a truss which is a unit of a portal
frame may involve the necessity of an end bolster, with
carefully detailed connection to post, truss-chord and
knee-brace.

There are, however, a few types of end joints which
have marked superiority, both from directness of action
of the stresses, and from simplicity and ease of fabri-
cation and framing. The cardinal principle may be
set down as a basis for design of all end joints, that the
complete thrust of the batter post should be taken by
one line of action alone. To illustrate, a joint should
never be framed so that the thrust is taken partly by
lugs dapped into the lower chord, and partly by an in-
clined bolt. Tests* have demonstrated conclusively
that the two systems will not act together, and that
either the lugs or the bolt will take the whole stress up
to the commencement of failure, before the other sys-
tem will come into action.

In the following paragraphs, eight types of end joints
are detailed and cost estimates of six are given. In
order that a comparison of these different details may
be made upon the same basis, a typical truss known as
the English roof truss has been chosen. The span is
70 ft., the distance between trusses is 24 ft., and the
load has been assumed at 40 lb. per sq. ft. of horizontal
projection. Of this total load, 13.5 lb. per sq. ft. has
been considered as acting at the lower-chord panel-
points, being made up of one-half the weight of the
truss plus the weight of the ceiling. The skeleton
diagram of the truss, together with the stress diagram

*Engineering Record, Vol. 42, November 17, 1900. This
article by F. E. Kidder shows full size end joints after failure,
and discusses the tests. The article is summarized and the
illustrations reproduced in *Jacoby*s Structural Details,* pages
274-276. Jacoby also discusses other tests of end joints.

92 TIMBER FRAMING

is shown in Fig. 38. The stresses in the various mem-
bers are indicated on the left half of the truss, while the
required sizes are marked on the right half.

The following working stresses have been used in
designing the detail^.

TiMBEB Lb. per

&q. in.

Tension on net section 1,500

Compression, end-bearing . ! 1,600

Compression, cross-bearing 285

Shear, longitudinal 150

Steel

Tension 16,000

Shear 10,000

Bearing 20,000

Lb. each

f-in lag screw in steel plate 1,030

J-in. lag screw in steel plate 1,200

All computations incident to the design are given, so
that the method may be followed by the reader. For
the sake of simplicity, it is assumed that the truss in
all cases rests upon a masonry wall, although no details
of support are shown. It is also assumed that the centre
Une of support passes through the intersection of the
centre lines of the chords.

Types A and B. Where the clearances, relative in-
clination of upper chord to lower chord, and the magni-
tude of the stresses will permit, the simplest and cheap-
est details for end joints are those represented by
Types A and B, shown in Fig. 39 and 40.-

In these two details, the thrust of the upper chord is
taken completely by bearing and shear on the lower
chord. As shown by the detail computations, the re-
quired length of the uncut portion of the chord for
shear is 41 in. The inclined bolts in both details take
no calculable stress. In Type A, the length of the cut
d is determined from the curves of Fig. 19 and 20, of
Chapter IV, which give the allowable pressures on
surfaces inclined to the direction of the fibres. On ex-
amining these curves, it is evident that if the pressures

TIMBER FRAMING

shown by the curve of Fig. 20 be used, this type of detail
must be abandoned, unless the sections of chords are
materially increased, as the required cuts in the chord
will be too deep. If, however, the safe bearing pressure
be taken in accordance with the curve of Fig. 19, the
required depth of toe-cut is 5 in. measured on the nor-
mal face of the strut. This leaves an uncut portion of
the chord with an effective area of 4J in. X 10 in. =

'^^^

^^^

:??^ 'S*

5^^S; «5:

O § ^

Sca/e of Stress Dkkftwns

^-£0.6 '472 -5€l2 -675

Sjoan 70''0'

<^1

FlO. 38. SKEXETON AND STBESS DIAGBAM:S OF TRUSS.

94 TIMBER FRAMING

42.5 sq. in. The average tensile stress on this section

49 000

is then —425" =1200 lb. per sq. in. With the centre

lines of truss members and end support intersecting
in a point, as assumed in finding the stresses in the truss
members, there will be excessive secondary bending due
to the couple of the horizontal component of the thrust
of the upper chord and the resultant tension in the
lower chord, the latter acting at one-half the uncut depth
of the lower chord. The moment of such couple in Fig.
39 is 49,000 lb. X 4 in. = 196,000 in. lb. To overcome
this moment, which would stress the chord to failure,
the line of action of the reaction of the truss, or the
centre line of support, must be placed at such a distance
to the left of the intersection of the centre lines of
the chords, that the couple formed by the reaction
of the truss and the resultant of the vertical thrust
of the upper chord on the lower chord will equal
the moment of the first couple. It has been assumed
that the thrust of the upper chord is uniformly dis-
tributed over its toe. Consequently the vertical com-
ponent of this thrust will also be uniformly distributed
over the same area* In this case the vertical com-
ponent of the thrust happens to coincide in position
with the line of the truss reaction. Therefore the line

of action of the truss reaction must be placed ^g *^^^

♦It is obvious that with a normal cut on the upper chord, as
in this detail, the whole thrust of the upper chord must be
taken on the toe of the post, and that the inclined cut can
take no pressure. The assumption of a uniformly distributed
pressure over the toe of the upper chord requires that there
exist another component of the stress in the chord which is
normal to its centre line, otherwise the joint cannot be in
equilibrium. This component is small, and will be resisted
either by friction of the toe on the cut in the lower chord, or
by tension in the bolts. In effect, such assumed distribution
of forces assumes that the direction of stress in the upper
chord is inclined slightly to the centre line of the chord, and
is coincident with a line drawn between the centres of the
normal cuts at the ends.

TIMBER FRAMING

= 7 in. to the left of the centre of the toe, or 7 in. to
the left of the intersection of the truss members.

In Type B, the cuts of the upper chord are so modi-
fied that neither of the two surfaces are normal to the
centre line of the chord. Consequently, each surface
will take a component of the total stress. The angle
between the two bearing surfaces may of course vary;
in this case it has been made 90°. Computations made
to determine the depth of cut in the lower chord, which
are shown in detail in Fig. 40, indicate that this depth
must be 4| in. This distance determines the length of
the other surface, which is found to be more, than suffi-
cient for its component of pressure. There then re-
mains to be determined the position of the centre line
of support, so that no secondary bending will exist in
the lower chord. For this condition, the moment of all

t^7

V 1 ^ ' €

E^G. 39. END JOINT, TiPE A.

Depth of toe: Computations

9 = 60°, therefore n = 1400 lb. per sq. In., from Fig. 19,
Chapter IV.

Required area in bearing = — = 40 sq. in.

1400

Required depth = — - = 5 in.

Length of chord projection for shear:

JMOO =41in. = 3ft.5 1n.
150 X 8

96 TIMBER PRAMINO

forces acting on the lower chord must equal zero. Each
normal component of stress on the two bearing sur-
faces of the upper chord is resolved into its horizontal
and vertical components. The product of the resultant
of these horizontal forces by the distance between such
resultant and the centre line of the uncut portion of
the lower chord is 49,000 lb. X 3| in. = 183,500 in. lb.
The resultant of the two vertical components of the
normal pressures is found to lie at a distance of 3J in,
to the left of the reaction. The new centre line of sup-
port must therefore be placed at a distance to the left
of this point such that the product of the vertical re-

Fio

40. END JOINT, TYPE B.

COMPTTTATIONS

Depth Of toe:

e = 74°, therefore n = 1680 lb. per
Chapter IV.

BQ. In.,

from

Fig.

Required area in

-"««=S=

= 32.8 sq

in.

Required depth =

= JM =4.1 1„.

Pressure on Inclined bed:

e = lfi°, therefore n = 400 lb. per sq. in.
Required area in bearings ' =32.1 aq. in.
Actual area = 16 in. x 8 in. = 120 sq. in.

TIMBER FRAMING

action by this distance will equal the moment of 183,500
in. lb. found above. This length is 6i in., so the new
centre line of support must lie 3 in. to the left of the
intersection of the centre lines of the upper and lower
chords. The long projection of the lower chord beyond
the point of intersection of its centre line with that of
the upper chord makes it extremely improbable that the
Types A and B end joints could be used in an actual
case. Undoubtedly some form of shoe would be found
necessary.

Types C and D. Types C and D, shown in Fig. 41
and 42, are comparable to the common end-detail of a
steel truss, in which the two chords deliver their
stresses into a common gusset plate. This is particu-
larly true for Type D. In Type C, each stress is com-
pletely taken in bearing by the steel tables riveted to

Chord.

A//// beann^ edge of rabies.

\^Qr^Q' Chord

Fig. 41. end joint, type c.

Computations

56 500
Bearing area of tables for upper chord = — *— — =35.3 sq. in.

1600

Qg OA

Assuming two tables each side of chord, depth of table = — '- —
= 1.103 in. = li in.

Rivets required in each table =-: — = three f-in. rivets.

4 X 4420

Thickness of side plates for bearing against rivets = ^ in.

Shearing area required for each table = - — *■ — -^94.2 sq. in.

4 X 150

94 2
Distance required between tables = -^^-^^ = 11.75 in. = 11} in.

98 TIMBER FRAMING

11.75

Treating side plates as columns, r== 0.0903. ■^/*"=/^ /^q/.«
= 130.

Moment of rotation on each table = -^^4^ X i (1.125 + 0.3125)

= 10,130 pound-inches.

10 130
Stress in bolts = ^ — *-j^ = 1450 lb. Use two i-in. bolts for

Z X 3.5

each table.

40 AAA

Bearing area of tables for lower chord = '^^ = 30.6 sq. in.

1600

Assuming two tables each side of chord, depth of tables = — —

= 0.955. Use 1 in.

49 000
Rivets required in each table = = three f-in. rivets.

4 X 4420

Shearing area required for each table = -l?i522_ = 81.6 sq. in.

4 X 150

Distance required between tables =—— =10.2 in. Use 10-in.

bolts as for upper chord.
Area between last table in upper chord and end of upper chord

greater than minimum area of 94.2 sq. in.
Net section of lower chord = (8-2) (8-1) =42 sq. in.

Unit stress in lower chord = iM2? == 1170 lb. per sq. in.

Bill of Matebial fob One End-Connection po«nda

Two A-in plates with area of 9.2 sq. ft., at 12.72 lb 117.4

Four flats, 1^ by 3 by 8 in., at 7.65 lb 30.6

Four flats, 1 by 3 by 8 in., at 6.80 lb 27.2

Twenty-four rivet heads at 0.136 lb 3.3

Thirteen bolts, i by 9 in., at 0.65 lb 8.5

Total weight of steel 187.0

dosT OF One End-Connection
187 lb. steel at %0M $7.48

the side plates. As will be seen by the detailed compu-
tations, the thickness of the tables is a factor of the
capacity of the timber in end bearing. A sufficient dis-
tance between tables must be given to provide the
necessary shearing area ; bolts must be provided to hold
the tables in their notches; the side plates must be
thick enough, acting as columns between the lines of
bolts, to take the largest stress; and last, each table

TIMBER FRAMING 99

must have enough rivets to hold its individual portion
of the total stress.

In principle, this type of end joint is perfect. There
is no eccentricity of the principal stresses, and conse-
quently no secondary stresses. As a type of end detail
from the standpoint of field work, the joint is not so
good. It will be found practically impossible to cut
and fit the notches for the tables with sufficient ac-
curacy so that each table will take equal and uniform
bearing. Moreover, any deficiency in bearing will be
extremely difficult, if not impossible, to remedy by

Lag

Oiom

Dotted c/rc/es s/joty Lag Screws in far plate.

Fig. 42. end joint, type d.

Computations

Number of lag ecrewe in upper chord ^^ °°' — ^47.
1200

Number of lag screws in lower chord = .

1200

Thichnesa of plate ^ A In.

Bill of Matebial Pounds

Eighty-eight I by 5-ln. lag screws at 0.95 lb 83,6

Two A-ii- plates with area of 15.4 sq. ft, at 12.75 lb 198.5

Total weight of steel 279.1

Cost of One End-Connection
279.1 lb. steel at (0.01 $11.16

100

TIMBER FRAMING

shimming or similar methods, on account of inaccessi-
bility after the side plates are bolted to the timber. The
joint is not susceptible to field examination for defects,
and this is perhaps its worst feature.

Fig. 42 (Type D) shows a modification of Type C, in
which the steel tables are abandoned, and the chord
and batter-post stresses are transmitted to the gusset
plates by means of lag screws acting in shear. Except
for the consideration of economy, this joint has all the
advantages of Type C, and none of its disadvantages.
All stresses are concentric, and with good inspection
during construction, one may rely on a close fit of the
lag screws in the timber. The holes in the steel plates
for the lag screws are drilled to a diameter of ^^ in.

Chord,

'Mardt^ocd 3foc^
/ Bent Pfate-r^CT-t'O'

-^-^ — -^

i' 2"^ Pipe
Pin.

Bo/ster '4-%'^ Boftz

Fig. 43. end joint, type e.

Computations

Area required for bearing between upper and lower chord =

?M2i = 99 gq, in.
285

Depth of lugs = ^^^525 = 1.915 in. Use 2-in. lugs.

2 X 1600 X 8

Thickness of lugs for bending:

Bending moment one lug = 24,500 lb. X 1^ = 36,750 pound-
inches.
Thickness of lug assumed to be 1 in.

no 'TRA

Required section modulus =: ■ '^^ =1.465.

25,000

Required thickness of lug = 1.05 in. Use 1-in plate as

assumed.

TIMBER FRAMING IQl

LengtlL required for shear between lugs= ' — =20.4

2 X 150 X 8

in. Use 1 ft. 8i in.

Depth of toe = ^^^^ ^ ^ = 3.84 in. Use 4 in.

49,000

Bearing stress of 1600 lb. per sq. in. is used, as timber
fibres are confined and therefore capable of taking full
end compression.
Stress in bolster = horizontal component of stress in two f-in.

bolts, = 4830 lb. X 2 X 0.5 = 4830 lb.

Number of shear pins required = — ^^^=0.75. Use one

800 X 8

2-in. pin.

Bill of Matebial fob One End-Gonnegtiom

One plate, 1 by 8 by 12 in., at 27.20 lb 27.20

One plate, 1 by 8 in. by 4 ft., at 27.20 lb 109.00

One bolt, f by 25 in., at 3.52 lb 3.52

One bolt, f by 33 in., at 4.54 lb 4.54

Four bolts, f by 15 in., at 1.54 lb 6.16

Two washerii», f by 3f by 3f in., at 1.40 lb 2.80

EHght washers, ^ by 3f by 3f in., at 1.02 lb 8.16

Total weight of steel 161.38

One 2-in. extra heavy steel pipe-pin.
One bolster 6 by 8 in. by 6 ft, 24 ft. B.M.
One hardwood block.

Cost of One End-Connection

Steel, 161.38 lb., at $0.04 $6.46

Pipe pins, 1, at $0.25 0.25

Bolster, 24 ft. B.M., at |0.04 0.96

Hardwood block, 1, at |0.50 0.50

Total cost end connection $8.17

greater than the diameter of the shank of the lag screw ;
the holes in the timber are to be bored in accordance
with the specifications to be given in the concluding
article of this series. The lag screws are to be screwed
and not driven into place. Lag screws are better fitted
for this type of joint than are bolts, as it would be
practically impossible to bore a hole from one plate as
a template and strike the corresponding hole in the
opposite plate exactly. Consequently, were bolts to be
specified, it would undoubtedly be found after the

102 TIMBER FRAMING

fabrication of the joint that the bolts were sprung or
bent into place, and their value in shear would be
questionable.

This type of end detail is well suited to trusses of an
A-shape, resting upon posts. The side plates in such
cases may be extended to engage the top of the post,
and thus to give considerable stiflEness to the building-
frame.

Type E. In Fig. 43 is shown a detail of end joint
using a steel shoe made of two plates. This detail is
similar to that shown in Jacoby's 'Structural Details,
p. 262, Fig. 63b. The points to be noted in a design of
this type are (1) the depth of lugs cut into the chord
to give the required bearing area, (2) the thickness of
the projecting lug to resist bending, (3) the distance
between lugs in order that there may be sufficient shear-
ing area to take the increment of stress, (4) the depth
of the vertical end-cut of the upper chord so that the
necessary area be provided for bearing against the
fibres of the timber, (5) the length of horizontal cut on
the upper chord for distributing the vertical component
of its thrust across the fibres of the lower chord, and
finally (6) the size of the inclined bolts for holding the
joint together.

Attention should be called to the fact that in this
detail, the centre line of the upper chord intersects the
base of the shoe practically at the toe of the shoe. In
consequence, there will be a tendency for the shoe to
rotate in a counter-clockwise direction, bringing an
uneven distribution of bearing over the base of the
shoe, with a possible crushing of the fibres of the wood
at the top of the lower chord under the toe of the shoe.
This tendency to uneven bearing pressure will be coun-
teracted by the action of the inclined bolts, which, if
always tight, will come into play with the application
of the load to the truss, and will cause a readjustment
of the joint stresses and a consequent approximately
uniform distribution of the vertical bearing pressure.

From the standpoint of field work, this type of shoe

TIMBER FRAMING 103

is an excellent one, though extravagant of steel. It is
simple in its action and comparatively easy to frame
into the timber. Care must be taken, of course, to see
that both lugs have an even bearing against the chord.
As has been noted already, this is a hard thing to se-
cure, and is a fault of all shoes having more than one
bearing surface. However, the shoe has only two lugs
to fit, as against eight in Type C. Moreover, the in-
spection for uniformity of bearing is easy to make, and
if necessary shimming can be done readily and effec-
tively.

This shoe can only be used for stresses requiring not
more than two lugs, hence its field of application is
limited. Another defect is that the forge work is diffi-
cult with the thickness of plate used. Especially is this
true of the bending of the end of the inner plate to form
the inner lug. Incidentally, this detail forms a good
example of the consideration of actual unit working
stresses as compared to purely theoretical values, as
mentioned in the first article of this series. "With a 2-in.
depth of lug, the bearing pressure against the ends of the
fibres is assumed to be 1600 lb. per sq. in. On account
of the fillet formed in bending the plate, the actual bear-
ing area will be decreased and the actual unit working
stress will probably be found to be around 1800 lb. per
square inch.

Type F. Type F, illustrated in Fig. 44, is a modifi-
cation of Type E, in which steel tables riveted to the
shoe plate are substituted for the lugs of Type E. Its
advantages are (1) the main plate may be reduced to
the minimum thickness as required by consideration of
'shear and tension alone, (2) any number of tables may
be used, and (3) the forge work is less than in the pre-
vious type. A point to be considered in a shoe of this
type is that no table should be placed under the end
of the batter post. The notch for the table will invari-
ably be made deeper than the table itself, so that the
vertical' bearing of the steel table on the timber chord
cannot be counted upon to distribute load.

104

TIMBER FRAMING

•I
•I

M
II

-A

.i'loo'

3' W Countersunk FfVet5,each taif^e

A 5

6'x8'

/3^\>^ \^^^' /^^^i!'^:

JL «- iSp

e- %" 3o/n

/-z" P/pe
P/o

Bo/ster
i^c7shers A//// bearing edge of tabJes

E*IG. 44. END JOINT, TYPE F.

Computations

Depth of toe as in type C, 4 in.

Area required for bearing between upper and lower chord =

?M?5 = 99 sq. in.
285

A 10-in. depth will therefore be required for the upper chord,

giving an area of 8 by 13 in. = 104 sq. in.

Depth of tables (assuming three used) = ^^'^^^ =1.275

3 X 8 X 1600

in. Use lA by 3 in.
Assuming three rivets in each table, stress in each rivet =

iM25 = 5450 lb. Use three ^-in. rivets in each table.

«7

Thickness of plate for bearing against rivets = | in.

Thickness of plate for shear = — ^^'^^^ = 0.614 in.

10,000 X 8

49,000

16,000 X (8 in. -2.80 in.)

Thickness of plate for tension

0.59 in.
Make plate f in. thick.

. - . X. « .^,1 1.3125 in. + 0.625 in. ^^ 49,000

Moment of rotation of tables = — X —

3 ^3

= 15,800 pound-inches.

Stress in bolts = 15:522. = 4520 lb.

Add stress due to pin in bolster = i X i X 800 lb. X 8 in.=

800 lb.

Total stress in two bolts =: 5320 lb. Use two f-in. bolts.

TIMBER FRAMING 105

Using two i-in. diagonal bolts, the horizontal component in the
bolster will be as in Type C, requiring one pin.

40 AAA

Distance required between tables for shear = — ' — =

3 X 8 X 150

13.6 in. Use 13f in.

Bill of Material fob One End-Connection

Pounds

One plate, f by 8 in. by 5 ft. llf in., at 17 lb 101.50

Three plates, 1^^ by 3 by 8 in., at 8.95 lb 26.85

One bolt, | by 23 in., at 3.28 lb 3.28

One bolt, f by 30 in., at 4.13 lb 4.13

Eight bolts, f by 16 in., at 1.62 lb 12.96

Ten washers, ^ by 3f by 3f in., at 1.02 lb 10.20

Two washers, f by 3f by 3f in., at 1.40 2.80

Nine Hn. rivet heads, at 0.24 lb 2.16

Total weight of steel 163.88

One 2-in. extra heavy steel pipe-pin.

Ft. B.M.

One bolster, 6 by 8 in. by 7 ft 28.00

One 2 by 8 in. by 40-ft. extra length upper chord 26.50

54.50
(Total B.M. = 53; use one-half only as labor will be prac-
tically the same.)

Cost of One End Connection

Steel, 163.88 lb. at, $0.04 $6.55

Pipe-pin, one, at $0.25 0.25

Lumber, 54.50 ft. B.M., at $0.04 2;18

$8.98

The size of diagonal bolts in both Types E and F are
not susceptible of computation, but are determined by
judgment and experience. In the present instance, two
l-in. bolts have been used for the diagonals, and two
f-in. bolts for holding the lugs or tables in their
notches. The sizes of the vertical bolts are found as
shown in the detailed computations.

Type O. Fig. 45 illustrates a detail of end joint in
which a cast-iron shoe is used to transmit the thrust of
the batter post to the lower chord. The details of such
a shoe may be arranged in several ways, but the form
shown represents a rather common type. As there are

106 TIMBER FRAMING

no diagonal bolts, the vertical pressure of the shoe on
the lower chord is not uniform, and hence a toe has
been provided, extending beyond the end of the batter
post. The depth of the lugs are determined by limi-
tations of end bearing on the timber, and their spacing
by considerations of shearing of the timber. As the
area of the base of the shoe is large, the first lug may
be placed underneath the batter post. The thickness
of the lugs is found by treating them as projecting
cantilevers taking shear and bending, using the stresses
shown in the detail computations. The number, size,
and arrangement of ribs is largely a question of judg-
ment, remembering that cast iron is rather brittle, and
that the shoe must consequently be well stiffened to
resist tension and bending. The thickness of the metal is
determined by the requirements of tensile and flexural
stresses and by general considerations of a minimum
thickness for castings to resist the unknown stresses of
shrinkage and the probability of unseen blowholes.

Type H. Type H, illustrated in Fig. 46, is a cheap
and, with well-seasoned timber, effective type of end
joint, where the circumstances of clearance will permit
of its being used. The principles involved in its design
are simple; the pins take the whole thrust, the bolts
being assumed to resist a tension equal to one-half the
total stress on the lower chord. The bolster must be
investigated for shear on the uncut portion, and ten-
sion on the net section back of the surface of applica-
tion of the upper chord. It must be emphasized that
the whole effectiveness of the joint depends on the
question of whether any shrinkage of the timber, sub-
sequent to the fabrication of the joint, is to be appre-
hended. If unseasoned timber is likely to be used in
the framing, the detail should not be used, as the cross-
shrinkage of the bolster and chord will allow the pins
to become loose, and an undue strain will come upon
the bolts with a consequent slip of the joint. This detail
of end joint may be further modified by omitting the
shear pins, and notching the bolster into the lower chord,

TIMBER FRAMING

/-?' Pipe Pin- ^■^~ Bo/rs-8-^-'3^'^3^- «

Top yietv of Casting

Fig. 45. end joint.

Computation 8
Depth of toe:

e^aO", (1^1400 lb. per aq. in., from Fig. 19, Chapter IV.

Required area In bearing = ^^■°'^° = 35 aq. in.
1400

Required depth o( vertical cut = — = 4.4 sq. in.
Required area of horizontal cut:

6 — 30°. n = 650 lb. per sq. in., i=_?M^=43.2 aq. in.
650

Length of horizontal cut =- — ^"^'^ = 7.1 in.

Maximum unit bearing-pressure of shoe on lower chord:

Aa designed, toe of shoe la 9 In. beyond point where the
vertical component of the thrust of upper chord inter-
sects base of shoe. Call this distance u^9 In. Let
length of shoe be I^^S ft. 9} In. =33} in. Let p =

maximum unit stress desired. Then p^2 1 2 — —I

X-^, where p^V^rtlcal reectlon^g^^, j^_
/., Width of chord

Therefore p=:250 lb. per aq. In.

Depth of luga as in Type E, 2 In.

Thickness of lug; as in Type E, bending moment on lug

= 24,500 lb. X li In. = 33,700 pound- inches.

3.700 In, lb. __„ ._ , ^

—mm 8-*5=*x

8 in. X f .

Required section modulua =

108 TIMBER FRAMING

Required thickness of lug =rt = 2.62 in., make 2f in.
Distance required between lugs as in Type E = 20^ in.
Moment of sotation of lugs = 24,500 lb. X i (2 in. + J in.)
= 33,800 in. lb.

Stress in bolt back of lug = , 33.800 ^ __ ^^^^^ -^^

(21 in. + f in.)
Use one li in. bolt.

Bill of Material fob One End Connection

One casting, weight 118 lb 118 lb. cast iron

Lb. steel

Two U by 16i in. bolts, at 6.71 lb 13.4

Four t by 16 in. bolts, at 1.63 lb 6.5

Two washers, i by 6f by 6| in., at 6.2 lb 12.4

Four washers, A hy 3i by 3ft in., at 1.01 lb .- 4.0

36.3
One 2 in. extra-heavy pipe-pin.

One bolster, 6 by 8 in. by 6 ft 24 ft. B.M.

Cost of One End Connection

118 lb. cast iron, at $0.0325 $3.84

36.3 lb. steel, at $0.04 1.45

One 2-in. pipe-pin, at $0.25 0.25

One bolster, 24 ft. B.M. lumber, at $0.04 0.96

Total cost of one end connection $6.50

in which case a small shrinkage will not cause the joint
to slip. This modified Type H is shown in the detailed
roof truss of Fig. 71.

General Sununary of End Joints. For the details of
end joints using lag screws, the question may aris^ as
to whether or not all of the lag screws may be counted
upon as acting together. I believe that this condition
will be realized approximately; certainly to the same
extent that the lugs of the other types of shoes will act
together. As has been noted before, the holes for the
lag screws should be either punched or drilled, prefer-
ably the latter, to a diameter not greater than ^ in.
larger than the nominal diameter of the lag screw. It
is to be emphasized that all lag screws are to be
screwed, and not driven into place, in holes of the proper
diameter. First, a hole should be bored with a length
and diameter equal to the length and diameter of the
unthreaded shank of the lag screw, continuing with a

TIMBBH FRAMING

FIO. 46, END JOINT, TYPE H.

Computation 8

Required area of bolster for shears; — ^--- — ^327 aq. in.

Required length of uncut portion of bolster for shear = — -
= 40.5 in. *

Required number of 2-ln. plna = ^^•'"^^ — = 7.7. Use 8 pins.

8 X 800 lb.
Required net area of bolster lor tension:

Maximum stress on cut portion of bolster ^ 4 X 6400 lb. ^

2B,600 lb.
Net section of bolster as detailed ^ 3 b]' 8 in. ^ 24 sq. in.

Unit stress In tension = JM5i =1070 lb. per sq. In.

24
Total stress In bolts = i x 49.000 lb. = 24,500 lb.
As detailed, have eleven |-ln. bolts, 35,500 lb.

Bill or Matemal fob One End Connection Lb.

Five i by Sll-in. bolts, at 2.17 lb 10.9

Six i by 13i-in. bolts, at 1.4 lb. 8.4

Twenty-two washers, ^ by 3g by 39 In., at 1.01 lb 22.2

One dowel, 2 by 4 In., at 3.6 lb 3.6

Total weight of steel 45.1

Eight 2-ln. extra-heavy steel pipe-pins.

One bolster, 8 by 12 In. by 6 ft 48.0 ft. B.M.

Cost or One End Connection

Steel. 4B.1 lb., at *0.04 »1.80

Bight plpe-plns, at I0.2E 2.00

One bolster, 48 tt B.M. lumber, at 10.04 1.92

Total cost of one end connection $6.72

110 TIMBER FRAMING

second hole of a length and diameter of the threaded
portion of the shank at the base of thread. Careful and
insistent inspection is necessary to secure the condition
of lag screws screwed into place, as the carpenter will
almost invariably drive the screws into place, if not
watched.

In all the shoes with riveted lugs, special care must
be exercised to see that good riveting is secured. This
statement may seem so self-evident as to be foolish to
mention. It must be remembered, however, that the
steel and iron work on a timber-framed structure is
usually let to a small iron-shop, if not to a blacksmith,
and careless work is to be anticipated. Attention must
also be paid to the milling of the bearing faces of the
lugs. It is a curious fact that the average iron-worker
regards any piece of steel or iron as being so superior
to timber in strength, that he does not consider de-
fective forge-work or riveting as of much importance.
In his opinion, a failure of a steel shoe on a timber truss
would be an impossibility. Again, in fabricating the
truss, the steel shoes themselves should always be used
as templates in cutting the notches for the tables, and
boring the holes. The use of a well-made shoe, neatly
finished, will result in more careful work on the part of
the carpenter when framing the truss, than if the shoe
is roughly made. To secure the best results, the steel
shoes, as well as all the iron work, should be detailed
and marked carefully and plainly. All the work should
be carefully inspected before it is allowed to leave the
shop. This inspection should preferably be done before
painting the iron, if painting is to be done.

In the details shown in Fig. 43 and 44, the hole for the
diagonal bolts in the upper chord should be specified as
i in. larger than the diameter of bolts. This is to allow
the upper chord or batter post to slip easily into the toe
of the shoe when the load is brought to bear upon the
truss, and to thus take care of a possible untrue fit of the
batter post into the shoe. With this arrangement, no
bending of the diagonal bolt will result, when the batter

TIMBER FRAMING HI

post wedges into the toe of the shoe. I have examined
many roof trusses before erection, where shoes with tables
or lugs have been used,, and have found in a number of
instances that the toe of the batter post did not touch
the shoe. Obviously in such a case, when the truss was
erected and the load applied, this bolt must have been
badly overstrained, both in tension and bending, if it
had a driving fit in the upper chord.

The preceding investigations show that the cast-iron
shoe has the advantage in economy over the other
metal shoes investigated for the case under discussion,
while the end joint using 2-in. shear pins is the cheap-
est of all. It should not be assumed that this same rela-
tion as regards economy, holds for all trusses. In gen-
eral, it has been my experience that joints of Types F
and G, and especially Type F, are the most suitable and
reliable for trusses with end stresses of some magni-
tude. Other joints of different types may be used for
special cases, but it is believed that the types here
shown will cover all the cases the engineer is likely to
meet, where single-stick chords are used. The unit costs
used may be questioned. It is not contended that the
relative costs of joints as given represent the actual costs
of each detail. Many factors that would influence the
price cannot here be considered, and hence the net costs
of the different types of joints are only roughly approxi-
mate.

112 TIMBER FRAMING

CHAPTER VII

Intermediate Joints of Trusses

An intermediate joint in a truss differs from the end
joint only in the smaller stresses to be considered, and in
the existence of a length of adjacent chord sufficient to
take the component of the diagonal stress in the web by
longitudinal shear in the timber. The discussion of the
end joints of Types A aud B will therefore apply in
principle to all other joints of the typical truss shown,
if the tension of the rod at the panel point be substi-
tuted for the end reaction. Details of intermediate
joints in roof trusses furnish a good indication of the
extent of the designer's knowledge of structural me-
chanics. Frequently, and especially where the truss is
counterbraced in the central bays, the opposing diag-
onals are merely butted against one another with no
provision for transmitting the component of the diagonal
stress to the chord.

Fig. 47, 48, and 49 illustrate three methods of detail-
ing the joints at panel points No. II and IV of the
English roof truss of Pig. 38 of Chapter VI. Of these
details, the two shown in Fig. 47 and 48 are the most
common. The details of Fig. 49 is seldom used ; never-
theless it is the most consistent and logical in principle,
and the simplest of construction of the three types
shown. This statement can best be brought out by a
detailed discussion of the three types.

Type A, Fig. 47. Since neither of the two bearing
surfaces of the indent is normal to the longitudinal axis
of the strut, both surfaces exert a pressure on the chord,
which can be determined by resolving the stress in the
member, 11,500 lb., into two components, perpendicular,
respectively, to the planes of the two bearing surfaces.
The angle which each bearing plane makes with the di-

TIMBER FRAMING 113

rection of fibres of the timber determines at once the
allowable unit bearing-pressure, as discussed in Chapter
IV. This unit pressure requires a certain minimum
bearing-area, and the necessary depth of indent is there-
by determined. Each of the two components must be
investigated in this manner.

It is evident that the shape of the chord indent and
the length of the two bearing surfaces can be found only
by a *cut and try' method. For example, in the present
instance, the web stress, 11,500 lb., is resolved into the
two components 8200 lb. and 6000 lb. The 8200 lb.
component has an inclination of 30° with the direction
of fibres, or the bearing plane makes an angle of 60°
with the fibres. By reference to Fig. 18, Chapter IV, the
safe unit pressure for this angle is seen to be 1400 lb.

per sq. in. The required bearing area is then j^ =

5.86 sq. in., which necessitates a depth of indent of -y-

= 1 in. Similarly the limiting inclination of the 6000-
Ib. component is 12^^°, which is the angle which the
bearing surface makes with the direction of fibers. The
safe unit pressure for this angle is 410 lb. per sq. in.
The required area for bearing for this surface is then

-rrr = 14.6 sq. in., which corresponds to a length of cut

14 6

of -g^ = 2.5 in. In a similar manner, the depth of in-
dent for panel point No. VI is found to be IJ in., cor-
responding to = 76°, n = nOO lb. per sq. in., giving
required area of 6.4 square inches.

The required angles for the cuts must be noted care-
fully on the plans in order to secure the conditions in
the field that are assumed in the design of the joint. As
each diagonal of the truss may have a different slope
for its end cuts, the most careful and accurate workman-
ship will be necessary to secure the desired results.
This type of joint therefore violates the principle that
all carpenter work should be made as simple as possible.

T^ype B, Pig. 48. In this detail, the end cuts of the
struts are normal cuts. The length of the chord in-

114

TIMBER FRAMING

dents can be determined at once, since the total stress
of the strut must act on the normal face of the strut
alone. The conditions in this detail are exactly as dis-
cussed for the case of the Type A, end joint, in the pre-

7;^% 7Jt;;:*%' dasher "^
Fig. 47. intermediate joint, type a.

ceding chapter. For panel point No. II, the angle O is
30°, 71 = 670 lb. per sq. in., the required area in normal

cut is g^Q = 17.2 sq. in., and the required length of cut
is ^ = 2.87 in. For panel point No. VI, = 60°,

n =1400 lb. per sq. in., the required area of cut is "JTao"

8 2
= 8.2 sq. in., and the required length of cut is -y =

1.37 in., or If in. The force necessary to hold the strut
in equilibrium, and which must be developed by friction
along the normal cuts of the strut will now be found.
With the assumption that the thrust in the strut acts
uniformly over the area of the normal cuts, the moment
developed is the thrust in pounds multiplied by the
eccentricity in inches, which latter is the distance be-
tween the centres of the normal bearing areas at the ends
of the strut. The moment is therefore 11,500 lb. X 3f
in. = 38,800 in.-lb. The length of the strut is approxi-
mately 153 in., therefore the force to be developed in
38800

friction is

153

= 254 lb. This f rictional force will act

parallel to the normal cut of the strut. Assuming that
the coefficient of friction of wood on wood is 0.20, the

TOIBBR FRAMING 115

effective resistance may be counted upon as amounting
to 0.20 X 11,500 lb. =2300 lb. In addition to this trie-
tional force, such joints shonld be always well toe-nailed.
Two 16D nails will give a resistance of 256 pounds.*

Both of the details of Kg, 47 and 48 involve a con-
siderable depth of cut into the chord. In the upper
chord, theoretically, the indent is of no consequence,
dnce the chord is in compression, and a tight joint is
assumed ; actually, however, there is a considerable loss
in efficiency. In the lower, or tension chord, the depth

Fig. *

of cut is important, so that any detail reducing the depth
of indent la to be favored.

There is also some eccentricity in the action of the
various forces around the panel points in both details.
For example, in Fig. 48, panel point No. VI, the hori-
zontal component of the thrust of the strut acts at the
centre of the toe, while the resultant tension in the lower
chord acts at the centre of the unuut depth of chord. The
moment of this couple is therefore (49,000 lb. - 39,000 lb.
= 10,000 lb.) X ^ in- = 33,375 in.-lb. The vertical com-
ponent of the stress in the strut is 6000 lb., and also may
be taken as concentrated at the centre of the toe. This
force forms a couple with the tension in the vertical rod
(10,000 lb.) less the concentration at the panel point
(4000 lb.), or a resultant tension of 6000 lb. The amount
of the couple is therefore 6000 lb. times the horizontal

•See Chapter V.

116 TIMBER FRAHINa

distance between the centre of the rod and the centre of
the normal cut on the strut, or 6000 lb. X 3f in. = 22,500
in.-lb. These two moments are in opposite direction of
rotation, therefore the resultant moment is 33,375 in-lb.
-22,500 in. -lb. = 10,875 in.-lb. The net section of the
chord, taking out the hole for the rod and the dap in the
chord, is 6 in. wide by 6| in. deep = 39,7 sq. in. The net
section modulus is i X 6 in. X (H in.)' = 43.8. The
resultant tension in the chord is therefore -gjjY +
^^=12,350 lb. + 248 lb. = 12,598 lb. per sq. in.
While the secondary stress is negli^pble in thk case, it
does not follow that it can always be ignored, and any
truss designed for high unit working-stresses should

Fra. 49.

have its joints investigated for secondary stresses. It
should also be borne in mind that any variation in the
relation of the web members meeting in a panel point,
resulting from careless detailing or framing may in-
crease these secondary stresses to a considerable amount.
In roof trusses employing details of intermediate
joints of Types A and B, it will often be found that,
with the condition of the centre lines of all members
meeting at a common point satisfied, the toes of the struts
will either bear against the rod, or the hole for the rod
will cut away part of the strut. Sometimes this condi-
tion cannot be avoided if the stmt is to be dapped into
the chord. If it so happens that the rod has not a driv-

TIMBER FRAMING 117

ing fit in the chord, which condition will usually exist,
especially with an upset rod and a deep chord, the toe
of the strut will have bearing against the chord for only
a part of its width. The result of this condition will be
that the actual bearing area may not be over one-half of
what was assumed in design, and the unit bearing stress
may consequently be double the allowable.

Type C, Fig. 49. The disadvantages of details of in-
termediate joints of Types A and B, as shown in the
preceding paragraphs are lacking in the detail of Type
C, illustrated in Fig. 49. In this joint, the strut has a
full bearing on the butt block, and the butt block, in
turn, utilizes the total width of the chord for bearing.
Also, the detail takes advantage of the full bearing pres-
sure in end compression of the butt block on the chord,
resulting in a minimum depth of cut into the chord.
Nearly all the cuts are normal, and the others are simple.
All the cuts cau be easily and accurately laid out and
made by the carpenter. The length of the butt block
can be adjusted to fit all conditions of possible inter-
ference with other connections. Its minimum length is
determined by longitudinal shear. The bolt through the
end of the butt block holds the block securely in its
socket. Whether there is any actual tension in the bolt
depends upon the length of the butt block. This can be
determined at once by inspection. If the line of the
thrust of the strut falls within the base of the block,
there can be no tension in the joint. However, it is well
to provide at least a f-in. bolt to bind the joint together
thoroughly. I have used this joint in many trusses of
all types, and have found it to be an extremely satis-
factory detail in all cases.

In Fig. 50, alternate details of intermediate joints are
illustrated. These may be used for panel points No.
II and VI, using the unit bearing pressures on incUned
planes in accordance with the curve recommended by
Howe, and shown in Chapter IV, Fig. 20. The lower
values for bearing-pressures result in deeper chord in-
dents than for the details of Fig. 47 and 48, and also

TIMBER FRAMING

necessitate increasing the strut from a 4 by 6-in. to a
$ by 6-in. timber, in the ease of Type B, in order to pro-
vide sufficient bearing area against the upper chord.

;^'-:5^'*-^ i^asher

FlO. &0, IKTEBMEDIATB

The detail calculations need not be repeated, as they
are similar to those made for Fig. 47 and 48.

Applying the lower bearing pressures to the butt
block detail, or Type C, it will be found that the 4 by
6-in. strut must be replaced by a 6 by 6-in. strut, in order
to provide sufficient bearing area for the inclined cut of
the butt block. The alternative would be to use a hard-
wood timber, such as oak, for the butt block.

This type of truss, with its small inclination of web
struts to upper chord, will usually require attention in
order that the joints may have sufficient bearing in ac-
cordance with the allowable unit working-stresses
adopted.

TIMBER FRAMING 119

CHAPTER VIII
Tension and Compression Splices

In Fig. 51 to 56 are presented details of six types of
tension splices, the details shown being for the centre
panel of the lower chord of the English roof truss of
Fig. 38, Chapter VI. As in the case of the end- joint de-
tails for the same truss, the calculations for the design
are fuUy shown in the figures. In all cases, the splice is
designed for only the computed stress in the chord. This
fact will influence any deductions that may be made re-
garding the comparative economy of the different types.
It will be seen that in som<e cases it would be impossible
to increase the capacity of the splice without weakening
the main member beyond the allowable limit. Most
specifications provide that all splices be made of suffi-
cient strength to develop the main section of the member
spliced, regardless of the possible smaller computed
stress existing in the member.

The details here shown are not presented as covering
the whole field of tension splices. They do show, how-
ever, some of the most efficient forms. Various modi-
fications are possible, as for example, the substitution of
square or rectangular keys of hardwood or metal for the
shear pins for Fig. 56, the omission of the wooden splice
plates of Fig. 53, and the tabling of the main member
itself. For a description of the various forms of splicing
timbers, including the lapped and scarfed splice, the
reader is referred to the texts ^of Jacoby, Howe, Thayer,
Kidder, and others.

The detailing of tension splices is a problem to be de-
cided for the individual case, in conformity with the
circumstances of importance of the connection, cost of
materials, quality of workmanship to be expected, possi-
bility of occasional inspection after completion, and the

120 TIMBER FRAMING

particular requirements of the splice. It will usually
be found, however, as iu the case of other truss connec-
tions, that certain details stand out as superior to the
many that may be used, and that such type or types may
be employed successfully for almost all the eases that
will arise, with minor modifications.

For convenience of reference the details shown may
be listed as follows: Fig. 51, the Bolted Fish-plate
Type, Fig. 52, the Modified Bolted Fish-plate Type,
Fig. 53, the Tabled Fish-plate Type, Pig. 54, the Steel-
Tabled Fish-plate Type, Fig. 55, the Tenon-Bar Type,
and Fig. 56, the Shear-Pin Type. The advantages and
disadvantages of each type will be discussed briefly, in
order that an intelligent selection may be made for any
actual ease.

Bolted Fish-Plate Type. The size of the bolts in this
detail are computed in accojdance with the formula

M^-JP X(y + Y)» ■"1'^''^ i'=^the thickness of splice
pad, or fish-plate, and C'^the thickness of the main

^ID£ E LCI/ATI ON

«1

£ Hex Nuts, 2-3%' H^ashets, each bolt

TOP View.

Fig. 51. bolt^ fish-plate splice.

Computation a

Net area required = ^^'"°'' '^ - =2G aq. In.
IBOO

Since the end detail reaulred an 8 by S-!n. chord, the splice
pads or fish-plateB will be made 8 In. deep. Plates 2 by 3 In.
are not eufficlent, so use 3 by 8-ln. ABSumlng the bolts to be

TIMBER FRAMING 121

, If in. diameter, the net area of fish-plates will be 6 in. X [8 in.
- (2 X If in.)] =27.0 sq. in., and the unit tension will be

39,000 lb. _. ^44g j^ . j^ .pj^jg calculation assumes the

27
bolts spaced in pairs, and not staggered. However, with the
large diameter of bolts used, the net section of chord should
be figured as if the bolts occurred in pairs.
Six bolts are arbitrarily selected for each side of joint.

Bending moment on one bolt= =^^^X [(if X 3 in.) - (i X

8 in.)] =11,380 in.-lb.

11380
Required section modulus = =: 0.474 in.

24000

(J«= Mil =4.82 in. and d = 1.67 in. Use If-in. bolts.
0.098

The unit bearing pressure on the diametral section of the

bolts = — ^^^ = 619 lb. per sq. in., which is less than

1.75 X 6 X 6

one-half the allowable.

Distance Required Between Bolts Inches

per bolt

Total shearing area required = ^ = 260 sq. in., or 3.61

Area required for transverse tension = 39,000 lb. X 0.1 __ ^^^

150 X 6 X 6
Adding diameter of bolt 1.75

Required spacing of bolts 6.08

Use 6-in.

Bill of Material foe One Splice Pounds

Twelve If by 18Hn. bolts at 12.6 lb 151.5

Twenty-four If-in. nuts at 3.2 lb 76.8

Twenty-four 3f-in. circular washers at 0.4 lb 10.0

Total weight of steel 238.3

Two 3 by 8-in. pieces, 4 ft. 6. in. long = 18.0 ft. B.M.

Cost of One Splice as Detailed

Steel, 238.3 lb. at $0.04 $9.55

Timber, 18 ft. B.M. at $0.04 0.72

$10.27

Bill of Material for One Splice, Using If -in. lateral pins.

Pounds
Twelve If by 15Hn. lateral pins, at 11.37 lb. (including

nuts) 137.0

122 TIMBER FRAMING

Pounds.
Twenty-four 3Hn. washers at 0.4 lb 10.0

Total weight of steel 147.0

Timber as before.

Cost of One Splice

Steel, 147.0 lb. at $0.04 $5.88

Timber as before 0.72

$6.60

timber. As explained in Chapter V, this formula is that
used by Jacoby and Howe, and is based on the assump-
tion of uniform bearing of the timber along the length
of the bolt. The use of this formula results in an ex-
cessive diameter of bolts being required, not only add-
ing to the cost of the splice, but decreasing the capacity
of the main timber for tension.

Besides the bending in the bolts, the net section of
main timber and fish-plates must be investigated for suf-
ficient area to resist the computed stress; the bearing
pressure of the timber against the bolts must not exceed
the allowable unit working-stress, the distance between
bolts, and also the distance between any bolt and the end
of the timber, must be sufficient for longitudinal shear
on the timber, and also for transverse tension. In pro-
portioning the splice for the latter stress, it may be as-
sumed that the transverse tension tending to split the
timber along the centre line of bolts is equal to one-tenth
of the longitudinal stress in the chord. The working
stress for transverse tension is taken at 150 lb. per sq. in.

For determining the necessary spacing of bolts, the
net distance as required by longitudinal shear in the
timber is found, and to this distance is added the net
length required for resfsting transverse tension. To
the sum of these two is added the diameter of the bolt.
This combined distance is the minimum that should be
used. The distance of the last bolt from the end of the
timber should theoretically be one-half that of the com-
puted bolt spacing.. On account of the tendency of
timber to check at the ends, those bolts in the details act-

TIMBER FRAMING 123

ing in shear, as in the splice now under discussion, have
been placed a distance of six inches from the end of the
timbers.

In the figure, the bolts are shown as If in. diam., and
full-size nuts are indicated. If they can be obtained at
a reasonable cost, standard lateral bridge-pins will be
cheaper, and an alternate cost estimate has been pre-
pared on this basis. The actual diameter of the lateral
pin is l{i in. instead- of IJ in. The reduction in size is
not important, since it is still within the computed
necessary diameter.

The disadvantages of this detail are the large sizes of
bolts, with a corresponding loss in efficiency of the total
splice, as has been noted. As was discussed in Chapter
V, the general theory on which the sizes of bolts has
been computed is believed to be incorrect, and such a
joint would seldom be used in an actual case.*

*This statement needs some explanation. Where, as is the
case under discussion, the pressure distribution on the bolt is
assumed to be uniform, and the diameter of the bolt is then
made of such a dimension that the bolt will have a resistance
to bending sufficient to withstand the bending moment re-
sulting from such uniform pressure distribution, the design
cannot be said to be inconsistent, and it is believed that the
action will be as assumed. Further, it may be said that no
bolt of lesser diameter will give as high a total resistance
per bolt of the joint. It is obvious that if the bearing of the
timber on the bolt is uniform along the length of the bolt,
and if the bolt is lai'ge enough to resist the resultant bending,
the capacity of the joint is limited by the safe unit bearing-
pressure of timber on a cylindrical metal-pin. If at the same
time, the flexural strength of the bolt is attained, the infer-
ence might be drawn that the design was the most economical
that could be made. Such an inference would be correct, were
the price of metal the same for all sizes of bolts, or for stock
bolts or lateral pins. The criticism that I make of the design
under discussion is that bolts of a smaller diameter are not
given credit for the resistance that they can develop. A joint
framed with the bolts nearly two inches in diameter has the
appearance of a monstrosity when actually viewed in the
field, and always excites the ridicule of the carpenter. It is
granted that the carpenter's opinion has no bearing on the
case, if the design is correct. However, the carpenter in this

124 - TIMBER FRAMING

Modified Fish-Plate Type. The principles of design
of this detail are the same as in the previous type, except
that the sizes of bolts are proportioned from the values
given in the previous article of this series. One-inch
bolts have been chosen; there is no necessity for using
this size as against either a smaller or larger diameter.
In general, the fewer bolts there are to place, the less
will be the cost of labor, and the more certain will be the
combined action. Against these considerations must be
weighed the amount of metal in the bolts, and the avail-
ability of the chosen size. Stock bolts are of course,
cheaper than special sizes.

As the working values used here were taken from the
results of tests in which the action of washers did not
play a part, the splice is detailed with standard malleable
washers, which will allow the joint to be drawn together
fairly tightly without crushing the timbers. An esti-
mate of an alternate detail, in which the bolts have been
spaced at the minimum distance allowable, and in which
standard pressed-st^el washers are used has also been
prepared.

The modified fish-plate splice is easily framed, and
for many joints is the most economical, when all factors
are considered. All bolts are to have a driving fit in the
timber. This is a condition that can easily be obtained
with good inspection. The simplest method of assuring
a driving fit with bolts is to examine the size of the bit
which the carpenter uses, and to see that all holes are
bored from one side only. In case a bolt has not a
driving fit, it should be withdrawn, and another bolt of
the next larger size be used. For this reason, it is well
to detail such joints with a slightly larger spacing of
bolts than is actually required.

Tabled Pish-Plate Type. The tabled fish-plate joint
for the case under consideration is simple and effective.

instance knows what is undoubtedly true, that the designer
did not realize that bolts of a smaller diameter are capable of
developing much more resistance to lateral shear than is stated
in the textbooks.

TIMBER FRAMING 125

The stress of the chord and fish-plates is taken in ten-
sion, shear, and end-compression of the timber, with
comparatively small secondary tension in the bolts. The
bolts thus act in their most efficient manner, not being

mm^mmmmmfmr^Mmwi'

dpqcpq^QepQapQjl ; (m®®®&^^

ee-/' Boffs,2'5rand.Malieable iVashers,each bo/r
SIDE ELEVATION.

•/ a'

a-dT^a" Sp/'/ce Pa ds- 6-8

L-LJ_U ','''■ ,■ irt^ i ■' ' ■■ ' '■■' '■■ ' ' ■■' '■■' '■■ ' ^L-_L

3 e'^e' ^= =:

, ^

3900 O"" \

■ ■ ■ •

■ J ^ ' ^^"

TOP I//EIV.
Fig. 52. modified fish-plate splice.

Computations

Bolts of 1 in. diameter will be used. The strength, of one

bolt in double sheaf with a thickness of fish-plate of 3 in. is,

from Chapter V, 2664 lb.

39000
Number of bolts required = =14.6. Use fourteen 1-in.

2664
bolts.

Distance required between bolts (total shearing area re-
quired, as before, 260 sq. in.) : Inches

260

Spacing of bolts for shear = — = 1.55

14 X 6 X 2

39000 V 1

Spacing required for transverse tension = — p — - — = 0.31

150 X 14 X 6

Adding diameter of bolts 1.00

Required spacing of bolts 2.86

Bolts will be spaced 2 in. staggered.

Required area of chord and plates for tension need not be
investigated.

Bill of Material fob One Splice ^^ .

Pounds

Twenty-eight 1 by 16i-in. bolts at 4.81 lb 135.00

Fifty-six 1-in. standard malleable washers at 0.75 lb 42.00

Total weight of steel 177.00

Two 3 by 8-in. pieces 6 ft. 8 in. long = 28.00 ft. B.M. timber.

126 TIMBER FRAMING

Cost of One Splice

Steel, 177 lb. at $0.04 $7.08

Timber, 28 ft. B.M. at $0.04 1.12

$8.20

For a rigid comparison with the previous type, the spacing

of the bolts would be decreased to 3 in., and circular steel

pressed washers used. The bill of material and cost would

then be as follows:

BILL OF MATERIAL p^^^^^

Twenty-eight 1 by 15^ in. at 4.60 lb 128.6

Fifty-six 1-in. circular washers at 0.16 lb 9.1

Two 3 by 8-in. pieces 5 ft. 3 in. long = 21.0 ft. B.M. 137.7

Cost of One Splice

137.7 lb. steel at $0.04 $5.55

21 ft. B.M. timber at $0.04 0.84

$6.39

subjected to lateral forces. All cuts of the timber are
square, and where the amount of the stress to be trans-
ferred across the joint in the chord can be taken by not
more than two tables on either side of the chord joint, the
detail may be regarded as reasonably certain in its ac-
tion. Washers of generous size must be provided, in
order that the joint may be well pulled together at the
time of framing and the bolts be able to hold the tables
in place when the stress comes into the splice. The cal-
culations for Fig. 53 show a moment of 39,000 in.-lb. to
be counteracted by the tension of the bolts in each table,
acting about the vertical cut in the chord, or the bear-
ing end of each table. It is obvious that if the bolts
should fail to hold the fish-plates in place, this moment
would have to be taken by the plates acting as beams
in flexure. The net section modulus of the plate at the
plane of the cut for the table is ^ X 8 in. X (24 in.) ^ =

8.35. The flexural stress would therefore be gg^ — '■ =
4680 lb. per sq. in., and the maximum stress on the fish-
plate would be 4680 lb. + 2^x8 =^^^^ ^^- P^^ ^- ^^-^

TIMBER FRAMING

127

^-^'x<9'' Spiice Pads- e-O

TiT?

»■ =--f

<JSf

•IT

n
*• It.

>tr

I6U'

|l
I*-

V6V

■^

-ii __ i [_J

« iL?

" 3900O

16'// 3li

SIDE . £L€\//I HON

rr^^" h^%

t J^»

e-0

/?- %" Bo/ts, 34-^ "3%'* Sfi' M7sher9.

TOP \/IEW,

Fig. 53. tabled fish-plate splice.

Computations
Depth of cut for table and chord: Area required for cut =

?5552 = 1.52 in. Make 1^ in.

1600 X 2 X 8

39000
Length of table for shear: Area required = =

16.25 in. Make 16i in. 8 X 2 X 150

Size of bolts required: The resultant stress in the fish-plate
acts at the centre line of the uncut portion, while the resultant
of the pressure of the table on the chord acts at half the depth
of the cut. The total thickness of fish-plate should be 4 in.,
since a 3-in. piece of timber would not give suflacient area for
tension. There is thus a couple acting on the fish-plate equal
to one-half the stress in the chord multiplied by one-half the
thickness of the fish-plate, or 19500 lb. X 2 in. = 39,000 in.-lb.
This moment must be resisted by tension in the bolts acting
about the bearing face of the tables. The bolts should be
placed at the centre of the tables. Their lever arm is there-
fore 8 in., and their stress ^^52? = 4875 lb. Two f-in. bolts

will be provided. In addition, for binding the joint together,
eight f-in. bolts will be placed as shown in the detail.

Bill of Material fob One Splice Pounds

Twelve f by 14Mn. bolts at 1.50 lb 18.00

Twenty-four washers ^ by 3§ by 3f in. at 1.02 24.50

Total weight of steel 42.50

128 TIMBER FRAMING

Two 4 by 8-in. pieces 6 ft. long = 32.00 ft. B.M. timber.

Cost of One Splice

Steel, 42.50 lb. at $0.04 J1.70

Timber, 32 ft. B.M. at $0.04 1.28

$2.98

a value far beyond the allowable safe stress. The joint is
therefore dependent to a large degree on the tightness
with which the timbers are held in place by the bolts,
and excessive shrinkage in the timber would allow the
fish-plates to be overstrained. In such a joint, if thor-
oughly seasoned timber is not certain to be employed, the
fish-plates should be given a generous section, and addi-
tional bolts over those required by the computations
should be provided. Spiking in the form of toe-nailing
will also assist in holding the fish-plates in place. The
bolts resisting the tension due to the incipient bending
should be placed at the centre of the tables, in order that
the fibres of the fish-plates will receive equal bearing
under the washers.

Steel-Tabled Fish-Plate Type. The calculations
necessary in the design of this type of splice are similar
to those of the Type C end joint. The net area of steel
in the plates, the bearing area of tables, number of rivets
in the tables, their spacing for longitudinal shear on the
timber, the number of bolts to hold the tables in position,
and the net area of timber must all be sufficient to hold
their respective stresses.

The splice is an effective one, and is fairly economical
where good work in the fabrication of the metal can be
obtained. The detail works well for joints carrying
heavy stresses. The objections that may be offered
to the splice, outside those of cost of materials, are the
number of tables that may be required, necessitating
careful fitting into the timber, in order that snug and
uniform bearing may be assured between steel and
timber. For this reason the detail may be listed in the
class that especially requires good and careful inspection
on the part of the engineer.

Tenon-Bar Type. The bar and tenon splice is one of

TIMBER FRAMING

129

rJ* S" T 9'

^''^'

hi [J I ' I

IBS

6-^''^T-0''6"nfbles,bearing edges mi/fed
/lU Rivets ^^ / /^// 25b//;y ^^

*5/Z7r ELE\/ATION

= — n

•I — =r

— Xssooo^^

TOP ]/l5yV,

Fig. 54. steel-tabled fish-plate splice.

Computations

9QAAA

Bearing area required for tables = =24.4 sq. in.

1600

24 4
Total combined depth of tables = ^ = 1.53 in. Make \\ in.

2X8

Will use tables f by 8 in., requiring eight tables in all.

39000
Each table transmits = 9750 lb., and requires three

J-in. rivets, as determined by bearing on a i-in. plate.

Net section of \ by 8-in. steel plate = i X [8 in. - (3 X 5 in.) ]

= 1.34 sq. in.

39000
Net section of one plate required = = 1.22 sq. in.

2 X 16000
Size of bolts required to resist moment on tables: Moment

= 9750 lb. X \ in. = 4857 in.-lb. • Tension in bolts = ^^ =
1400 lb. Will use two f-in. bolts. ^*

Bill of Materials foe One Splice Pounds

Two i by 8 by 4 ft. 4 in. plates at 29.4 ft 58.8

Eight i by 3 by 8-in. tables at 5.1 lb 40.7

Twenty-four f-in. rivet-heads at 0.14 3.4

Twelve f by 9Hn. bolts at 1.14 13.7

Total weight of steel 116.6

Cost of One Splice
Steel, 116.6 lb. at |0.04 $4.67

130 TIMBER FRAMING

the older types of timber splices, and was formerly
used to considerable extent in bridge work, but is not
often seen at the present time. It is distinguished
from all other splices by its simplicity and directness.
There is but one bearing surface, consequently the area
taken out by the bar is a large part of the gross area
of the chord. As the bar is rectangular in shape, the full
end-bearing value of the timber can be taken advantage
of, and there is no cross-tension on the timber tending
to split the chord. The detail computations consider
the size of bar for bearing against the ends of the fibres,
and for bending in the bar, the required distance be-
tween the bar and the end of the timber for shear, the
net section of chord, and the area of tension bolts, using,
of course, the area at the root of threads. It should be
observed that the length of the bar is determined by the
long diameter of the hexagonal nut of the bolts, so that
sufficient distance may be obtained for tightening the
nuts. As the bar is a short beam in bending, the high
unit flexural-stress of 24,000 lb. per sq. in. is permissible.
For holding the splice firmly together, two 2 by 8-in.
pads have been provided, bolted through the chord. This
will be necessary wherever a single stick is to be spliced.
In the case of a built-up chord, such as is usual with a
railroad or highway bridge of long span, the entire chord
is never spliced at one point, the splices in the individual
timbers of the chord being staggered. Packing blocks
are provided between the sticks, through-bolts being
used to bind the whole together thoroughly.

SheaT-Pin Type. In this detail, the tension is trans-
mitted by shear in the pipe or hardwood pins, and a
consequent secondary tension in the bolts. The work-
ing values are in accordance with the results of tests, as
has already been described in Chapter IV. With
thoroughly seasoned timber, the detail is a reliable one.
It should be remembered that this detail should not be
employed with very green timber, as the ability of the
pins to transmit shear is a function not alone of the end-
bearing value of the timber, but also of its strf^ngth in

TIMBER FRAMINO

131

/-^^^J /'4'4'

2'2''*a'' ^/ice Pads 4 -%" Bolts.

e-Z'^'^J-y Rods, 2 Hex, Nuts.each.
SIDE EUl/ATION

2 Bar3-2'4"^2f^''''/''-/'^

3' -4'
TOP VIEW

Fig. 55. tenon-bar splice.

Computations

Size of rod: area required = ^^'^^^ ^^' =1.22 sq. in.

16,000

A l^-in. rod has an area at root of thread of 1.295 sq. in. Use
this size.

The long diameter of a IHn. hexagonal nut is 2f in., hence
the distance from the side of the timber to the centre line of
the bolt must be slighty more than If in. Will make this dis-
tance 1^ in.

Size of bar required: The bearing of the timber against the
bar will be assumed to be uniform per unit area of bearing.
Hence the bending moment on the bar will be 19,500 lb. X
[1^ in. + (i X 8 in.)] =19,500 lb. X 3.4375 in. = 67,000 in.-lb.

The required bearing area is — '- ' = 24.4 sq. in.

1,600

The required width of bar is therefore ^^^^ = 3.07 in. Use
3 in. ^

Use a fiber stress of 24,000 lb. per sq. in., since the case is
that Qf a short beam restrained at the ends to some extent.

The necessary section modulus is -^~- = 2.79 in.

^=2.79= ?-A^
6 6

24000
Therefore h = 2.37 in.

Will use bar 2| by 3 in.

132 TIMBER FRAMING

The shearing length required, or the distance between the

edge of bar and the end of the timber, is — ?5522 — = 16.23

150 X 2 X 8

in. To this distance will be added one-half the width of the

bar, making the distance from the centre of bar to the centre

of splice, say, 1 ft. 5i in.

Bill of Material fob One Splice Pounds

Two steel bars, 2| by 3 in. by 1 ft. 5| in. at 35.6 lb 71.2

Two li-in. rods, 3 ft. 4i in. long at 20.2 lb 40.4

Four hexagonal nuts at 2.0 lb 8.0

Four f by 13i in. at 1.4 lb 5.6

Eight f-in. standard malleable washers at 0.23 lb 1.8

Total weight of steel 127.0

Two 2 by 8-in. pieces 5 ft. 6 in. long = 15 ft. B.M. timber.

Cost of One Splice

Steel, 127 lb at $0.04 $5.08

Timber, 15 ft. B.M. at $0.04 0.60

$5.68

cross bearing. In the action of the splice, there is a
couple on the pin, tending to spring it out of its hole.
Any shrinkage in the timber will allow some slip of the
joint, because of the action described.

General Summary of Tension Splices. As was stated
in the case of the estimates of costs of the different types
of end joints, the figures representing the costs can be
regarded as only approximate. The actual amount of
labor required for each type of detail, whether end
joint or tension splice, is difficult to estimate accurately.
In the costs given herein, all timber in place has been
figured at the same rate, and the same statement applies
to the steel, whether such steel is forged, riveted, or is in
the form of bolts. This assumption would be justified
only in the case of a very large job. On a small job com-
prising only a few roof trusses, this method of estimating
costs would probably be seriously in error. Again, prices
of material fluctuate, not alone in relation to time, but
also with the situation of the job. It will be recognized,
therefore, that not alone must the prices of the different
types of tension splices be taken as only approximate.

TIMBER FRAMING 133

but that their relative costs must be regarded as only
comparatively accurate.
While the relative cost of any one type of tension splice

5IOC CUl/ATION
IS- Z' Shear Pins.

v^pm^nm

B-^' Bolts,

TOP yicw.

flo. 66. sheab-pin splice.

Computations
Using fish-plates of 3 by S-ln. timbers, tbe net section ot
plates will be 4 by g in. = 32 sq. In. The unit tensile stress
la tbe plates will then be ^^^ =1216 lb, per sq. la.

Quiring i^i^=4.03 or lour i-iu. bolts.

The number of 2-ln. pins required will be ■ """"- =6.1.
Use six pins. * ^ ^"'*

The tension to t>e taken in the bolts wtti be 19,500 lb., re-
19500 _
4830 "

For developing the bolts, plate washers, | by 41 by 41 in.
will be used.

Bul of Material fob Omb Splice Pounds

Bight i by 16i-ln. bolU, at 2.32 lb 18.56

Sixteen washers, g by 4i by 4i in., at 1.92 lb 30.70

Total weight ot steel 49.26

Twelve 2-ln. pipe-pins or hardwood pins.
Two 3 by 8-ln. pieces 4 fL long = 16 ft. B.M. timber.
Cost op One Splice

Steel, 49.3 lb. at 10.04 $1.97

Twelve pipe-pins at (0.10 1.20

Timber, 16 It. B.M. at (0.04 0.64

13.81

134 TIMBER FRAMING

is a vital factor to be considered, other considerations
than cost alone will generally decide the detail to be
used. For example, where the roof truss is to be ex-
posed, wooden splice-pads may be objectionable from the
standpoint of appearance. In such a case, steel plates
are a necessity, unless a laminated chord be used, and
the problem will then resolve itself into a question of
using a bolted-steel fish-plate splice, or a tabled-steel
fish-plate.

In the case of joints in which the stresses to be re-
sisted are comparatively small, the modified bolted
fish-plate splice will be found satisfactory. In the case
of the truss illustrated here, the number of bolts re-
quired is too large from a practical standpoint, and the
tabled fish-plate detail is, perhaps, the most satisfactory
of the types shown. Where the stresses are still larger,
the tabled-steel fish-plate will be found to offer an eco-
nomical solution. Of the various types of splices illus-
trated, the bar-tenon type is the only one that is prac-
tically free from the effects of shrinkage of the timber.
Next may be classed the bolted fish-plate, followed by the
steel-tabled fish-plate, the wooden-tabled fish-plate, and
lastly, the shear-pin splice.

Shrinkage is a factor that is almost impossible to
avoid. For this reason, in all timber design, it is well to
use conservative stresses. In the case of a roof truss,
there usually exists a considerable safety factor in the
live load assumed in the design. Where the timber
work is protected from the weather, and at the same
time is accessible for inspection, the joints should be
carefully watched, and the bolts tightened as the timber
shrinks. If the structure is a building that is heated, the
full shrinkage may occur in a few weeks. Green timber
may take one or two years to season completely. It
should be emphasized, therefore, that thoroughly sea-
soned lumber only should be used for construction which
will not be accessible for inspection and maintenance.

In addition to the requirements of computed stresses
in tension splices, the unknown stresses of erection must

TIMBER FRAMING 135

be provided for. A splice joint should always, therefore,
have some general stiffness in addition to its capacity to
resist the known stresses. An illustration of this state-
ment is seen in the detail of the bar-tenon type, in which
two 2 by 8 in. splice-pads are used, although not re-
quired theoretically. The general statement may be
made that, where it is possible to secure chords of full
length so that splicing may be eliminated, it is better
to use the long sticks, even at a considerable increase in
the unit cost of the timber.

Compression Splices

Compression splices may be divided into two classes,
those which take compression only, and those which may
be called upon at some time to take either flexure or
tension, or a combination of each.

As in the case of tension splices,* it is not the purpose to
discuss here the many types of joints which are em-
ployed in timber construction. Each type has its ad-
vantages and disadvantages. The reader who is inter-
ested in the subject will find a very complete description
and discussion of timber joints in Chapter II of Jacoby 's
'Structural Details.' Three of the most common types
of compression splices are shown in Fig. 57. These may
be termed the butt joint, the half lap, and the oblique
scarf, respectively. The figure illustrates the funda-
mental difference between the butt type and all others,
namely, that the former has only one surface of contact,
and the others two. In accordance with the principle
already mentioned, that all timber joints should be made
as simple of fabrication as possible, the butt joint is
superior to the others, whose efficiency is largely de-
pendent upon the experience and the care of the car-
penter in framing. It is self evident that one bearing
surface is always better than two.

In the butt joint shown, the thickness of the splice-pads
and the number and size of the bolts may be varied to ac-
commodate the conditions existent in the member,
whether the splice-pads are required only to hold the

136

TIMBER FRAMING

main timbers firmly in position, or must transmit tension
or compression across the joint. In cases where. the main
member must resist considerable flexure, it may be
necessary to use metal or hardwood shear pins in addi-
tion to the bolts.

The oblique scarf has more flexural strength than the
half -lap; otherwise I consider it much inferior to the
half-lap. There is much less timber in straight end-
bearing in the oblique scarf than in the half -lap, and the

V\An

tJM

^:.-.-

^-i:]

Fig. 57. types of compression splices.

oblique cut, if unresisted .by the normal cuts would tend
to separate the two timbers.

In erecting a structure^ in which the member to be
spliced is vertical in position, the bolts through one end
of the splice pads should be placed in the field, so that
the joint may come to a full bearing before the bolt holes
are bored.

Pig. 58 and 59 show two details of the upper-chord
joint of a small timber highway bridge, in which the
batter-post frames into the upper chord. Here again

TIMBER FRAMING

137

the comparison may be made between a somewhat com-
plicated detail, as illustrated by Fig. 58, and a joint in
which but one straight cut is required, as in Fig. 59.
The latter detail fulfills all the functions of the former,
and there seems to be no need for the double cut of the
first detail. If any further fastening between the chord
and batter post is required than is given by the detail
of Fig. 59, it is best obtained by splice pads across the
joint, either of steel or timber.

General

All of the joints of the present and previous chapters
are of the class that may be termed the open type, as

Fig. 58. complicated upper-chord detail.

opposed to the closed, or housed class, the latter embrac-
ing the simple housed, the cogged, the halved, the dove-
tailed and other joints. The housed joints, with the ex-
ception of the halved joints, are seldom seen in building
construction. The housed joints are used in mine tim-

Fig. 59. simple upper-chord detail.

138 TIMBER FRAMING

bering, both above and underground, and also in crib-
bing. Aside from the fact that they are more difficult to
construct, they offer more chance for decay than the
simple open joints, and their efficiency is but a small
proportion of the total strength of the timber. For ex-
ample, the main timber is badly cut by an entrant tenon,
compression across the fibres is introduced, and the effect
of cross shrinkage of the timber is a maximum. The
use of such joints requires large timbers working at a
low unit-stress, or, in other words, at a low efficiency.

TIMBER FRAMING 139

CHAPTER IX

Main Members of Trusses

Compression Chords and Struts. The importance
which is attached to the details of a truss may be in-
ferred from the fact that their design has been dis-
cussed before that of the main members. A case where
the details of a truss are amply sufficient for the stress-
es that may come upon them, and the main members
are insufficient, seldom if ever occurs.

The present treatment of compression members may
be divided into two sections; first, a discussion of
solid timber struts subject either to concentric com-
pression alone, or to a combination of concentric com-
pression and cross-bending, and second, a discussion
of built-up timber struts, straight or curved, and tak-
ing either simple compression or compression combined
with bending. The bending in either case may be due
to an eccentricity of the primal stress or may be pro-
duced by transverse loading.

The calculations for the design of intermediate truss-
joints, as given in the previous article, indicated that
the struts of a timber truss are usually dependent on
the required area for end-bearing, rather than upon
the allowable working-stresses from the standpoint of
column action. This condition will prevail in the aver-
age roof -truss encountered in practice, where the length
of the struts is short.

If, in a building truss, the roof joists rest directly
upon the upper chord and are rigidly fastened thereto,
the latter may be considered to be supported by the
joists, and column action will not enter into the deter-
mination of its section. When, on the other hand, the
roof joists frame parallel to the truss, or when the
joists do not rest directly upon the upper chord, the

140 TIMBER FRAMING

chord must be considered as a column, and its allow-
able unit stress determined by the application of a
column formula.

Formulas giving the safe working-stresses for timber
columns are to be found in almost every text-book on
structural engineering, and in most specifications for
steel structures. The two formulas most commonly
used are those of the American Railway Engineering
Association, and the U. S. Department of Agriculture,
Forestry Division. The first mentioned is expressed as
follows :

p = working unit stress, in pounds per square inch.

C = safe fibre stress in end compression, in pounds
per square inch.

L = length of column in inches.

d = least diameter or dimension of column in inches.

The U. S. Department of Agriculture formula is more
complicated. It is

_ r r/_700jfl5c^\l

^ ~ ^ t \ 700 -f 15c 4- c /J

In this expression, p and C represent the same quanti-
ties as in the formula above. In addition,

c = — , where

L = length of column in inches, and

d = least diameter or dimension of column in inches.

Two other formulas may be quoted, that of Milo S.
Ketchum for mill buildings, and* that of the Seattle
Building Ordinance, 1914. Using the same nomenclature
as before, these two formulas are

p = C (l - ji^ ^ ) , (Ketchum)

P = C (i-^ot) (Seattle)

The value of C recommended by the American Rail-
way Engineering Association for railroad bridges and
trestles is 1200 lb. per sq. in. for Douglas fir. For high-

TIMBER FRAMING 141

/aoo teoo

160C J600

1^00 MOO

/200 /200

iOOO 1000

600 eoo

600 600

400 400

200 ZOO

10 10 ^ 40 SO eo

Fifl. 60.

way bridges and trestles, C may be increased to 1500 lb.
per sq. in., and for buildings a value of 1800 lb. per sq.
in. is allowed. Ketelium's recommended value of C for
Douglas flr is 1200 lb. per sq. in., while that of the
Seattle Ordinance is 1600 lb. per square inch.

For purposes of comparison, these various formulas
have been platted in Fig. 60, for the ease of use in a
building structure, protected from the weather. With
the exception of Ketehum's formula, the values as given
by the various formulas are not far different between the
limits of 15 -^ and 35 -^ . Sixteen hundred pounds per
square inch is the value recommended in this article for

142 TIMBER FRAMING

use with Douglas fir m building construction with a
grade of lumber of No. 1 Common. The selection of any
one of these formulas as against the others for use in
designing will usually not materially affect the section

of the strut or column. The value of -^ should not
exceed 60 for any column or strut.

When computing the necessary size of the upper

chord of a truss, the area of sections taken out by rods,

bolts, washers, etc., must not be forgotten, and the
proper allowance must be made for these losses of ef-
fective area. The fit of the web compression-members
or the butt-blocks, if used, may be assumed to be such
that the joint is 100% efficient, if an allowance for their
cuts would mean a serious increase in cost. Otherwise,
it will be better to make some allowance for poor fitting.
The relation of actual load on the truss to probable load,
as well as other considerations make this point one to be
decided by the engineer for the particular case in hand.

Composite or Laminated Compression Members. In

beginning this discussion, the general statement may be
made that composite or laminated columns and struts,
that is, columns built up of dimension stock, and spiked
or bolted together, should be avoided whenever possible.
In designing roof trusses for armories, skating-rinks, etc.,
it is often necessary to use trusses with arched chords.
In such instances, laminated chords acting in compres-
sion and in tension may be a necessity. The requirement
of bending the individual members of the chords to a
curve of a comparatively short radius decrees that these
members be either built of boards or of comparatively
thin planking. The peculiar complications of such a
truss will be discussed a little later.

For the general case of built-up columns, it may be
argued that the average quality of timber in such a
composite column is higher than that of a solid stick.
On the other hand, and by far counter-balancing this
small advantage, is the practical impossibility of making
the built-up column act as a single stick. Composite

TIMBER FRAMING 143

columns may be separated into two classes, the first class
comprising those constructed of a number of boards or
planks, laid face to face, and bolted or nailed together,
and the second class, consisting of those columns built of
several laminations, with their edges tied together by
cover plates. These two classes are illustrated by Fig.
61, a and h. Tests* have shown conclusively that the
first class, when bolted together at the ends and the mid-
dle, will act as individual sticks. Expressed in another
way, it may be said that the strength of a composite

V\X'%WVOXV

Fig. 61, a and &. types of laminated columns.

column, without cover-plates, and bolted together, is the
sum of the strengths of the individual boards or planks,
acting as separate columns with a length equal to that
of the whole column. When, in place of, or in addition
to, the bolting, such laminated sticks are spiked together
thoroughly, the total strength of the column is in excess
of the sum of the individual sticks.

In an endeavor to throw some light on this subject, I
made a few tests in 1915 on some small composite col-
umns. The results were published in Engineering News,
Vol. 75, No. 7, February 17, 1916: Five built-up col-
umns, constructed in three different ways, with a 3 by
4-in. section, and 23 in. long, were tested in compression
to failure, and for comparison, two solid timbers, of the
same cross-section and length, were also tested. The de-
tails of the test columns are shown in Fig. 62, while the
load-deformation curves are shown in Fig. 63. The lum-
ber was No. 1 common Douglas fir, with ends true and
square, and surfaced. The individual boards were sur-

*See 'The Elasticity and Resistance of the Materials of
Engineering/ by Wm. H. Burr, 1905 edition, pp. 539-541; The
Materials of Engineering/ by J. B. Johnson, pp. 682-683; also
'Structural Details/ by H. S. Jacoby, pp. 210 and 217.

144

TIMBER FRAMING

faced on one side. Fig. 64 shows the columns after
failure.

The column of Type b was found to be as strong ad
the single stick. At the ultimate load, however, the in-
dividual pieces separated somewhat. The failure in this

[:^

mm\

ll'l; !

'1 i 'f

)

w p;

i!l

« «

; I

> J, _, ..

\ \\\'

'4 •

" ■

' "I'l

_ ,.

' 1 •
• •

• ♦

...

.,:■■ 4

• •

1 1

1 ■ I'

e •

• •
• •

• •

• •
• •

...

• •

*: ♦

l*>

...L

• •

■■

' ' 1
1 f

«M

^1 B«

■

• •

•• r

• •

-. ..

. \

■ m

■

,•:!!',■■•

Y < *

• •

I'ir:: ■

•ir^

■ -

liii ^^

! ' 1

■ III

04.

lli,

! '^

Fig. 62.

i>. c

DETAILS OF TEST COLUMNS.

column was a combination of crushing, resulting from
straight compression, and of tension, due to the bending
of the individual sticks. The columns of Type c were
far deficient in strength as compared to the other types.
The ratio of length to least diameter of the individual
boards of the c columns was 46, while the corresponding

quantity for the a columns was 31. The ^ for the

solid sticks was 7f . The ultimate strengths of these
three types of columns as computed by the formula of
the U. S. Department of Agriculture, assuming the ulti-
mate strength of the timber in end-compression to have

TIMBER FRAMING

145

been 4500 lb. per sq. in., and assuming the sticks to have
acted as individual columns, would be as follows:

Table XVII

Type a Type 6 Type c

— 31 31 46

Ultimate strength (computed).. 29,600 lb. 29,6001b. 21,4001b.

Ultimate strength, (actual) 49,0001b. 50,0001b. 38,0001b.

Efficiency 98% 100% 76%

Average of values of line (2)

and strength computed as solid

sticks 39,900 lb 35,800 lb.

While the small number of the tests, and the diminu-
tive size of the specimens does not warrant forming too

•

1 '

soc

CO-

V*-

i~N.

\t/'

•/

j \

1V«^

-5r

y^"'

k'^

'\l

re'>

^/l /I

90 '

1 \ 1

i\i

}

f%/V

1 \

-J \

OO-

■el

! Vi !\

20 C

lJ 'vi Iv

r— 1

t — ■

toe

^i _

^ M

h ■ ■ 1
J

ate

O.IO

0« format ion m /nch^s

a/0

Fig. 63. load-defobmation curves of laminated columns.

definite conclusions to be extended to large-size columns,
it is evident that the columns of Type & or the 'cover-
plate' type are much superior in strength to the plain
laminated type. This is what might be expected from a
theoretical standpoint. An inspection of the values of
the ultimate strengths as given in Table XVII indicates
that the actual strengths of the columns of Types a and
c are not far from the mean of the strengths computed
first as a solid stick and then .as a summation of indi-
vidual sticks. In this connection, it is well to note that
the spiking of these composite columns was exceedingly
thorough, and such effective spiking could not be ex-
pected in actual construction. Even though equivalent

146 TIMBER FRAM:nQ -

spiking were to be specified and shown on the drawings
of actual columns, U would be a Herculean task for the
inspector to secure this result in the field. When, in
addition to the practical impossibility of obtaining
sufficient nailing, it is remembered that in the case of
a laminated chord of a truss, and also in the ease of
many composite columns, the individual boards or

planks splice at various points throughout the length
of the chord or column, it will be realized that another
element of weakness is introduced, namely, the failure
of the carpenter invariably to secure perfect butt-joints
in the splices. Any such imperfect splice will, of course,
put an additional stress upon the spikes, which must
then transmit the load of the spliced timber to the ad-
joining boards.

From a consideration of the above factors, and until
further tests prove otherwise, I- recommend that the
strength of a composite, coluimi of the type of Fig. 61a _
be taken at 80% of the mean of the strengths computed
(1) as a solid stick, and (2) as a summation of the
strengths of the individual sticks considered as indi-
vidual columns. For columns of the type as illustrated

TIMBER FRAMING 147

in Fig. 616, or the ^cover-plate' type, I recommend that
the strength be taken as 80% of that of a solid stick of
equal cross-section and length.

Curved Laminated Tnuis-Chords in Gom.pression and
Tension. The preceding discussion has considered
only straight columns or struts. The much more com-
plicated case of a curved laminated truss-chord must
now be treated. The subject is one that is generally
avoided in the few text-books on timber-framing, or at
best is dismissed with brief mention. However, as has
been stated in a former paragraph, the case of a lam-
inated truss-chord, acting in compression or in tension,
is one that occurs frequently, and it therefore becomes
of vital importance to establish, if possible, average safe
working-stresses for such chords. In addition to the
average unit-stress on the section, there is introduced
the complication of secondary stresses. Due to the fact
that in the solution for the stresses of such a truss, the
chords are assumed to be straight between panel-points,
when actually they are curved, there is produced a bend-
ing in the chords,- equal to the total main stress in the
chord multiplied by the maximum eccentricity of the
centre line of the chord measured from the straight line
connecting the two adjacent panel-points, except as this
bending moment may be modified by conditions of con-
tinuity and fixedness of the chord. Further, as has been
mentioned previously, there exists a considerable initial
stress in each lamination of the chord, resulting from
springing the boards to the required curve during con-
struction. The amount of the bending due to the as-
sumption, in the stress analysis, of a chord of straight
segments may be computed ; also the initial stress in each
board due to framing to a curve may be found. The
difficulty arises in determining the actual efficiency of
such a composite beam in resisting the bending due to
the eccentricity of the main stress,* and in deciding for

♦Reference is made to 'Graphical Analysis of Roof Trusses,'
by Charles E. Greene, 1905 edition, p. 82, and the following
statement is quoted from the paragraph entitled 'Curved

148 TIMBER FRAMING

what length of time the modulus of elasticity of the
timber remains constant. With regard to the first con-
sideration, the efficiency of the chord-section to with-
stand bending depends on the number and position of
the splices of the boards, and the ability of the nails to
resist, without slip, the longitudinal shear between the
laminations. Remembering that the nails are also
called upon to resist the shear due to column action, and,
to a large extent, that due to the initial bending of the
boards, and further, as shown in a previous article, that
nailed joints slip at a comparatively small load, it will
be realized that the laminated chord should not be
credited with a high efficiency.f

The complications and the uncertainties of the prob-
lem can best be appreciated by a practical example. For
this purpose, Fig. 65 shows a skeleton diagram of one of
the three-hinged roof arches of the main group of build-
ings of the Panama-Pacific International Exposition. In
this figure the sizes of the various members of the arch
are indicated. Fig. 66 gives a detail of a portion of the
lower chord of the arch. The maximum compressive

Beams:' "If the planks are bent to the curve and laid upon
one another, this combination is not nearly so effective as the
former (scarfed boards side by side, the plane of the boards be-
ing parallel to the plane of the loading — H. D. D.), but it can
be more cheaply made. The lack of efficiency arises from the
unsatisfactory resistance offered to shear between the layers
by the bolts or spikes. The strength to resist bending moment
will be intermediate between that of a solid timber and that
of the several planks of which it is composed, with a deduction
of one for a probable joint. If the curved member has a
direct force acting upon it and a moment arising from its
curvature, the treatment will follow the same lines; but the
joints, if there are any, will be more detrimental in case there
is tension at any section. Such curved pieces are sometimes
used in open timber trusses for effect, but their efficiency is
low on account of the large moment due to curvature."

tin framing a curved laminated truss chord, such as the
one under discussion, the required curve is marked out on the
floor of the fabricating platform, and blocks are then nailed to
the floor along the curve. These blocks hold the boards in
position. One after the other, the boards are then bent to the

TIMBER FRAMING 149

stress is 42,400 lb., while the maximum tension is 45,400
lb. These stresaeB are due to the foUowiug loads:

Lb. per ad. ft.

(1) Dead load plue live load 35

(2) Dead load plus wind load

The wind on the side walls, or the vertical portion of

'U^M^

*U^^

■^^

■a-IZ-r^a'surlecetl nif

^
W
^

the truss, was taken at 20 lb- per sq. ft., and the wind
on the roof in accordance with Duchemin's formula,
with P^30 lb. per sq. ft. Per the condition of deac.

curve, eacb eucceeaive board being nailed to the preceding one.
Both chords are usually tramed on the floor in their correct
relative position, and the web members, struts and rods, are
then placed in the truss. The truss is often fabricated com-
pletely before the blocks are released. The chords are thus
to some extent maintained In their correct shape by tbe action
of the web members against the butt-blocks, and the butt-blocks
against the first boards. The nails and the bolts binding tbe
board or boards between the butt-blocks are in shear. The
statement above, of which thla note Is an explanation, is
believed, tlien, to be a reasonable one.

150 TIMBER FRAMING

load alone, the lower chord is in compreseion, with a
stress of approximately 24,200 pounds.

Referring to Fig. 66, the eccentricity of the centre
line of the chord from a straight line connecting the

I Z 3 4

Lot^er Chord- I2'I''I2'' Board%
Surfaced to %' Finished Thickness.

FlO. 66. DETAIL OF PART OP LOWER CHORD.

adjacent panel points is 1 in. The bending moment due
to this eccentricity, disregarding the effect of continuity
of the chord, is therefore 42,400 in.-lb. As the chord is
continuous, and held rigidly at the panel-points by the
long butt-blocks, the apparent bending moment may be
reduced by the factor |, making the effective bending
moment 31,800 in.-lb. If the chord were a solid stick,
the section modulus would be i X 12 X (10.5) = = 220.5,
and the maximum fiber stress would correspondingly be
-ggng = 144 lb. per sq. in. The actual efficiency will be
taken in accordance with the recommendations of Mr.
Greene, namely, as the average between the section
modulus of the solid stick of equivalent cross-section and

TIMBER FRAMING 151

the sum of the section moduli of the separate boards,
minus one. In other words, the efficiency of the chord

will be taken at 220^ =0.54, or say one-half. The

actual maximum fibre-stress due to the bending will
therefore be twice 144 or 288 lb. per square inch.

For finding the initial stress due to the springing of
the boards to the curve of the chord, the formula

will be used. This formula is one of the forms of ex-
pressing the bending moment in any beam according to
the Common Theory of Flexure.
Af = bending moment in inch-pounds.

E = modulus of elasticity.

/ = moment of inertia in inches.

R =. radius of curvature in inches.

We may also write the equation,

M = K ^ 6d% where

K = maximum unit fiber-stress in any board,
h = width of any one board, and
d = depth of any one board, both in inches.
Equating the two expressions, we have

whence

Khd^ —

EI
R ~

K —

= E

1
12

dE
R

bd^

1
R

The radius of curvature may be assumed, for prac-
tical purposes, to be constant for all boards of the chord,
and its value will be. taken at 485 in. Then,

K=XXKX-^ X 1,500,000 = 1350 lb. per sq. in.
The average gross unit compression in the chord is

424Q0

12 X 10 5 ^^ 337 lb. per sq. in. To find the maximum unit

compressive stress in the chord, the three values found
above must be added. Thus, adding 288 + 1350 + 337,
the maximum compressive stress is seen to be 1975 lb.
per square inch.

152 TIMBER FRAMING

The unsupported length of the boards may be taken
at 33 in., or slightly more than the distance between the
ends of the butt-blocks. Due to the continuous chord,
and the long butt-blocks bolted through the chord, which
produce to a great degree the effect of 'fixedness,' the
length of the column may be reduced to i"X 33 = 16^
in. The ratio of length to least width is therefore 16^

X y =19. The safe unit stress is then 1170 lb. per

sq. in., using the U. S. Department of Agriculture for-
mula, with C = 1600. If, on the other hand, the allow-
able unit-stress be determined on the basis of column
action of the chord as a whole, the unit-stress will be,
using the same formula, 1600 lb. per sq. in. Using the
recommendations set forth previously, the allowable unit
fibre stress would be 80% of the half sum of 1170 + 1600,^
or 1100 lb. per sq. in. The computed maximum stress
is therefore considerably in excess of the allowable.
It must be stated, however, that the lumber in these
laminated chords was clear and straight grained, it being
so specified and furnished. Consequently, its ultimate
strength was considerably in excess of the average grade
to which the column formula applies. Further, the com-
puted fibre stresses are for the condition of dead load and
wind. With dead load alone acting on the truss, the
maximum fibre stress would be 1650 lb. per sq. in. Con-
sidering the fact that this truss was for a temporary
building, the unit stresses, while high, were considered
as safe. The strength of the long butt blocks, dapped into
the chords, and bolted thereto is an important factor in
stiffening the laminated chord in compression, and unless
such construction exists, the effective length of column
should be taken as the panel length of the truss.

In a similar manner, the maximum tensile stress mav
be found. The stress due to springing the boards to
position is the same as before ; the stress due to second-
ary bending is greater in the proportion of the principal

stresses, or 4^Jqq X 288 = 308 lb. per sq. in. The av-

TIMBER FRAMING 153

erage unit stress in tension on the whole chord is ~J26~
= 360 lb. per sq. in. An allowance must be made in
this case for splicing of the boards ; it will be assumed
that the chord has an efficiency of 75%. The average

stress of 360 lb. per sq. in. must therefore be increased

4
by the factor -j , resulting in an actual unit stress of

480 lb. per sq. in. Finally, combining all the unit
stresses we have a total unit stress of 1350 + 308 + 480
= 2138 lb. per sq. in. For the grade of timber used, this
is not an excessive unit stress.

The change in the modulus of elasticity of the timber
has been mentioned. It is a' fact, established by tests,
that if a load be left on a timber beam for some length
of time, the modulus of elasticity of the timber will
drop to approximately one-half its value for temporary
loads. This phenomenon is generally expressed by the
recommendation that, in computing the deflection of
timber beams, the modulus of elasticity for a 'dead' or
constant load be taken at one-half the value used for
4ive' or temporary loads. It is believed reasonable,
therefore, to state that while the stresses due to spring-
ing the boards of the truss-chord illustrated above are
actual stresses at the time of framing, a change in the
properties of the timber eventually takes place, resulting
in a decrease in the modulus of elasticity, and conse-
quently, a diminution in the stress due to shaping the
boards to the curve. Just how long a time is required
for this change to take place it is difficult to say, the
time being dependent to some extent on the original
moisture-content, the amount of bending introduced in
the chords, and the protection from the weather in the
structure of which it is a part.* It is believed the initial

♦For "the purpose of obtaining some definite measurement of
the amount of this initial stress remaining in such laminated
chords, I conducted some tests on the boards of the chords of
one of the Trusses 'A' of the Panama-Pacific International
Exposition (one of the same trusses just discussed) through
the kindness of C. H. Munson, assistant to the Director of

154 TIMBER FRAMING

stresses are reduced within a few months nearly one-
half.

The complicated conditions existing in a curved, lam-
inated truss-chord, will now be appreciated; also the
force of the statement that such a section should be
avoided whenever possible. The calculations and the
reasoning of the above discussion may seem to be both
doubtful in accuracy and cumbersome. I am frank to
admit that the result reached in the illustration chosen
rests upon a number of assumptions whose validity can-
not, perhaps, be definitely proved. However, the main
tenets are true: the initial stress due to springing the
boards to a curve does exist, and approximately to the
amount computed, when the boards are first bent ; after-
ward, this stress undoubtedly decreases; also there does

Works. On September 29, 1916, tliree laminations were re-
moved from a chord which was built approximately two and
one-half years previous. The length of the chord, and the
middle ordinate of the arc of the approximate circle to which
the boards sprang back on being released were measured, and
from these measurements the radii of the circles to which the
boards returned have been computed. Of the three boards
measured, the radii of their respective circles were 95, 81, and
96 ft., or an average of 91 ft. The same boards were again
measured on October 14, 1916, and the respective radii were
found to have increased to 128 ft., 129.5 ft., and 132 ft, or an
average of 130 ft. As the fibre stress due to the curvature
is in direct proportion to the radius of curvature, it may be

40
stated that the measurements indicated that of the

(130-40)
initial stress remained in the boards up to the last date men-
tioned, or, in other words, that approximately 45% of the
initial stress still remained in the chords. This calculation
is on the assumption that the modulus of elasticity of the
timber had not changed. The boards were again inspected
after about three weeks, when they had nearly straightened
out. The results of these experiments were not such as to
justify any definite conclusions. The boards did not form a
true circular arc after being removed from the truss, so that
accurate measurements were impossible. They were stored
inside a warehouse, and lay on their edges. They were, there-
fore, free to take their natural shape, except as the friction of
the floor held them to the curved shape.

TIMBER FRAMING 155

exist a secondary stress of bending due to the curve of
the chord. The actual amount of the reduction of the
initial bending stresses is somewhat uncertain, but the
assumptions made herein are believed to be fair.

Curved laminated chords are more efficient in tension
than in compression, and a truss with the compression
chord of solid members, even if broken and spliced at
every panel-point, is in many cases to be preferred over
one in which both compression and tension chords are
curved laminated sections. This statement is made ad-
visedly. I have seen instances of laminated curved com-
pression-chords in a badly buckled condition.

Timber Tension-Members. The tension chord of a
truss, when framed in timber, needs no further discus-
sion. It has been shown above, and also in the treat-
ment of end-details, that secondary stresses very often
add considerably to the primary stresses. Timber will
seldom fail in straight tension ; the details will give first.
For this reason, it might seem reasonable to use a much
higher unit stress in tension than has been recommended
in these articles. However, because of the uncertainties
of the actual amount of secondary stresses, and the varia-
tion in the structure of the material, it is recommended
that from 1500 to 1800 lb. per sq. in. be taken as the
extreme limit for tension in the case of live and dead
loads for permanent structures.

Tension-Rods. The selection of the proper size of
tension rods is not merely the problem of dividing the
maximum stress by the allowable unit stress. Certain
other factors enter into the problem from the practical
standpoint, and these will be discussed briefly.

To the computed stress in a tension rod of a truss, as
found from the stress analysis, it is well to add an initial
tension. In fabricating a truss, camber is usually intro-
duced, and largely by means of springing the chords,
cutting the web compression members to fit, and holding
the truss in this strained position by tension in the rods.
The amount of this initial tension that should be added
may be taken at from 1500 lb. for the smaller roof

156 TIMBER FRAMING ^

trusses to 3000 lb. for the larger trusses, the values
given being for each rod of the truss.

Either plain or upset rods may be used, the former
being cheaper for the shorter rods, and the latter eco-
nomical for the longer rods. The dividing line for any
case can be determined easily from local prices. In
using plain rods, it must not be forgotten to take the
area at the base of the threads as the net section in de-
termining the size of rod to be used. If upset rods are
specified in designing the truss, great care must be ex-
ercised to see that no welded rods are furnished. It is
the custom in some small shops, when rods with upset
ends are specified, to weld * upsets' to the body of the
rods. This practice results largely from the fact that
such shops have no upsetting machines. Welds in plain
rods are, of course, a possibility, and thQ inspector must
needs watch for them, especially in long rods, but their
occurrence is not so probable as it is in the case of upset

■

rods.

Another factor to be considered in the design of ten-
sion steel is the quality of steel to be expected. This
consideration will aJlect the working stress to be used.
On the Pacific Coast at least, re-rolled steel is used al-
most exclusively for the stock sizes of rods. The princi-
pal objection to re-rolled steel is its variable composition,
as shown by its fibrous, laminated fracture. Again, al-
though medium steel may be specified, wrought iron may
be furnished. The quality of the material in the tension
rods of a truss can only be determined by tests, and
these are not always convenient or possible to make. It
is of interest to the designer df such trusses, therefore,
to be familiar with the limitations of the market, and his
design may be modified accordingly. For the purpose
of indicating the nature of the metal commonly fur-
nished under specifications calling for medium steel cor-
responding to standard specifications, there is given be-
low the results of some tests on various sizes of rods
taken from material submitted by contractors under
specifications calling for medium steel to correspond to

TIMBER FRAMING 157

Table XVIII

RESULTS OF TESTS ON ROUND STEEL TRUSS RODS

Nominal dimensions, inches 1^ li If

Actual dimensions, in 1.498 1.238 1.129

Actual area, sq. in 1.7624 1.203 1.001

Yield point, actual load, lb 53,070 37,780 33,150

Yield point, lb. per sq. in 30,112 31,404 33,117

Ultimate strength, actual load, lb.. 88,130 61,620 51,960

Ultimate strength, lb. per sq. in 50,005 51,221 51,908

Elongation in 8 inches, in 2.00 2.30 2.35

Elongation, per cent 25.00 28.75 29.25

Dimensions, reduced section 1.240 0.960 0.824

Area, reduced section 1.207 0.7238 0.5332

Reduction of area, per cent 31.5 39.83 46.73

Character of fracture: li-in. round steel truss rods, charac-
teristic, badly laminated, small distinct bars In mass; lHn-»
part cup, characteristic, badly laminated; l^-in., characteristic,
slightly laminated.

the standard specifications adopted by the Association
of American Steel Manufacturers.

Comparing these results with the specifications of the
Association of American Steel Manufacturers, the ulti-
mate strength is below the limit for structural steel
(60,000 lb. per sq. in.), and somewhat below the limit
set by the specifications of the American Society for
Testing Materials (55,000-65,000 lb. per sq. in.) ; the
elastic limit is satisfactory ; the percentage of elongation
is satisfactory, the requirement being a minimum per-

o . 1 -xi- £ 1400000 1400000

centage on an 8-in. length of uit. strength = "sioor
= 27.5% for material not over J in. in thickness, with
an allowable deduction of 1% for each |-in. increase in
thickness, except that the minimum elongation shall not
be less than 18%. For the l^-in. rod, the minimum
elongation required is, according to the rule above,
24.5% ; and for the l^-in. rod, the minimum is 21.5%.

The testing company reporting the above tests classi-
fied the material as wrought iron. In my opinion, this
classification was erroneous; the fracture had somewhat
of the fibrous texture characteristic of wrought iron, as
opposed to the silky or granular fracture of steel, but

158

TIMBER FRAMING

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TIMBER FRAMING 159

was actually what may be described as a mild 'mongrel'
steel. The laminated composition ia characteristic of the
material, and shows to better, or worse, advantage in the
case of plates. Fig. 67 shows a typical specimen.

Table XIX gives the results of some other truss rods
and also of bolts. These results are introduced for the

F DEFECTIVE Bli>BOLLED STEEL.

purpose of giving a general idea of the fairly uniform
characteristics of the steel of this class. It will be
noticed that the elastic limit is quite high.

For use in rods, or plate connections taking tension
alone, it is believed that the material may be used with
confidence, employing a stress of 16,000 lb, per sq. in.
for dead and live load. For important work, such as
where the full live load is a certainty, this steel should
not be used, and rigid adherence to the standard specifi-
cations should be required.

160

TIMBER FRAMING

CHAPTER X

Bracing-Trusses — Details of Howe-Type Roof Truss —
Lattice Trusses — Truss Connections to Posts

Bracing-trusses in building construction may serve
one or all of three purposes, first, that of stiffening
the top or compression chords of the main roof-
trusses, second, providing general stiffness to the build-
ing against wind, and third, supporting the roof joists

Building kVaU

Fig. 68. general plan of bbacing-tbusses.

directly, and transferring this load to the main roof-
trusses. In the latter event, the bracing-trusses are
generally referred to as purlin trusses. It might
seem at first thought that the most economical ar-
rangement of framing would be secured by using pur-
lin trusses, and utilizing them as bracing-trusses. This
is not necessarily true, however. While, by this scheme,
the upper chords of the main trusses are relieved of
cross-bending from the roof joists, the additional ma-

TIMBER FRAMING

161

terial necessary in the bracing-trusses to enable them to
carry the roof is usually considerable. In addition, more
steel is required in the main roof -trusses, because the
shear in these trusses is a constant from the supporting
columns to the point of attachment of the purlin trusses,

Ooubte R6cf J049H - Sp/tte /%^«

vs*.v — =- '

Fig. 69. method of trussing boof joists.

l'2'*a'nof; /- 3'- ^'^^rticai

r B

Fig. 70. detail of bracing-teuss.

and for requirements of general stiffness, the rods of
each main truss between the purlin trusses cannot be
altogether omitted, even though the shear due to roof
covering and joists is zero. Because of the varying fac-
tors of span and spacing of the main roof -trusses, direc-
tion of slope of the roof, and possible limitations of clear-
ances and ceilings, no hard and fast rule as to the most
economical arrangement of framing can be stated. I
have found, however, that in actual cost of construction,
there is little difference between a roof framed with the
joists resting directly upon the chords of the main
trusses, and one in which purlin trusses are employed.
In the case of a brax;ing-truss which carries no roof

162 TIMBER FRAMING

load, the principal points to be observed are that the
chords have a section capable of taking compression, and
that the bracing-truss has a good and rigid attachment
to the chords of the main trusses. Theoretically, the
lower or tension chord of a roof -truss needs no stiffen-
ing. Practically, however, it is well to support it later-
ally, not alone to keep it from warping out of shape, but
also for the purpose of adding general stiffness to the
building frame.

The actual stress which maj come upon a bracing-
truss is usually indeterminate. In many cases, however,
a definite scheme of wind bracing may be provided, in
which the bracing trusses play an important part.
For example, the roof may be stiffened to act as a hori-
zontal beam against wind pressure, transferring the
wind loads to the end or side-walls, or to columns and
walls. In such cases, diagonal rods are usually intro-
duced in the plane of the roof-joists, the upper chords
of the roof-trusses and the bracing-trusses acting as the
chords of the horizontal wind-trusses. This is an ef-
fective way in which to stiffen a building against wind,
provided that the connections are carefully studied, and
made strong enough properly to fulfill their respective
functions and provided that the walls are well braced.
Pig. 68 illustrates the general scheme.

For buildings of small height and truss spans, suffi-
cient stiffness may be obtained by trussing the roof
joists, similar to the detail shown in Pig. 69. A detail
of a bracing-truss which is easily framed and is efficient
is shown in Pig. 70.

The rjequirements of bracing in timber-framed build-
ings are no different from those of steel-framed build-
iilgs. The general conditions of provision for wind pres-
sure, arrangement of main trusses and bracing-trusses
are the same in either type of building, with due allow-
ance for the nature of the roof to be supported. In this
connection, the excellent texts of Ketchum and Tyrell
on the subject may be profitably studied by the reader
interested in the construction of mill buildings, and

TIMBER FRAMING 163

other buildings having large open spaces, necessitating
long columns and roof trusses. As was remarked in the
introductory chapter, steel-framed buildings of these
types are generally designed by a competent engineer.
The same building, however, if framed in timber, is often
planned by an architect unacquainted with the funda-
mental principles of structural engineering. That the
consideration of roof bracing is vital will be appreciated
by reading the account of two recent failures of timber
buildings, one in Salt Lake and the other in Atlanta,
Georgia*. The last named failure resulted in loss of life.
It is significant that in both cases definite information
as to the plans of the buildings, as well as accounts of
the failure, were hard to obtain, there being an evident
desire on the part of both the architect and the municipal
authorities to hush up the matter.

Details of Howe-Type Roof Truss

In Chapters IV, V, and VI the design of the details
of timber trusses has been discussed and illustrated
by typical joints of open-panel trusses. In this chap-
ter, it is desired to show a complete truss, designed-
in accordance with the principles set forth in the pre-
ceding chapters. Another type of roof -truss, the lattice-
truss, as distinguished from the open-panel type, will
also be described, and illustrated with a typical case.
The design of the supporting columns is so closely
interwoven with the subject of roof-trusses, as to re-
quire simultaneous treatment. In general, the illus-
trations of roof-trusses given in the text-books con-
sider trusses supported on masonry walls. While this
is a common case, the engineer or architect is con-
fronted frequently with the problem of a timber-framed
building, that is, a building with timber roof-trusses
and timber columns, forming a structural frame sup-
porting the walls and roof, which may be of wooden
sheathing or of corrugated iron. Indeed, this case
has been the most common in my experience. Here,

* Engineering News, Vol. 75, No. 25 and Vol. 76, No. 2.

164 TIMBER FRAMING

many of the details which might be used in connec-
tion with masonry walls have had to be either dis-
carded or modified. The timber columns do not merely
support the dead weight of the roof -trusses ; they be-
come a part, with the trusses, of a definite structural
frame, technically termed a * transverse bent,' which
resists the lateral forces of the wind and stiffens the
building. This subject was mentioned above, in speak-
ing of the function of bracing-trusses. The designer of
such a * transverse bent' must consider carefully wind
forces, and the means of providing for them. In this
chapter it is not the intention to treat of wind forces in
any detail, but the connections of trusses to columns will
be discussed.

In Chapter VI, Fig. 38, is shown a diagrammatic ele-
vation and stress-diagram of a 70-ft. span Howe-type
timber roof-truss. Fig. 71 of the present chapter gives
the truss completely detailed. The spacing of the
trusses is assumed at 24 ft., and the loading at 38 lb.
per sq. ft. of horizontal projection of roof surface. For
simplicity, all loads are assumed to act at the upper
chord of the truss. This assumption is somewhat in
error, as the dead weight of truss and bracing trusses
should be taken as distributed between the upper and
lower chords. The resultant error is, however, small
and can be neglected in this case.

In the design, it has been further assumed that the
roof -trusses are for a building of the * mill-building'
type, that is, having a definite structural frame of
timber trusses, columns, wall-girts, etc., and that some
stiffness against lateral forces is desired, although no
definite length of columns has been established, nor
have any wind stresses been computed.

This detail presents what, in my opinion, is the most
economical and efficient truss for such buildings. It is
designed with conservative unit-stresses in all its de-
tails; it is simple of construction, and direct in its
action. The rods are slightly larger than is required
by the stresses indicated on the stress diagram in order

TIMBER FRAMINO 16^

to allow for initial tension when fabricating the tnisi^.
Washers of ample size are provided so that the rods
can work to their full capacity. The butt-block type
of intermediate joints has been used; hence the diag-
onal struts have full bearing at their ends. Attention
is directed to the detail of the end- joint. This detail,
where circumstances will permit of its use, is believed
to be the most eflScient and at the same time, the cheap-
est, that can be found. (If the lower chord could ex-
tend beyond the post, the end details, Types A and B
of Chapter VI could be used). The roof -joists rest di-
rectly upon the upper chord. While this introduces the
secondary stresses of bending into the chord, the joists at
the same time support the chord laterally in its weak-
est dimension, acting as a column, and so permit rather
high unit fibre-stresses. To give sufficient lateral sup-
port, it is required that the roof -joists be well spiked
to the chord, and that they also be well spiked at their
laps. In the end-panels, bolsters have been provided
for the lower chord. These bolsters, being well bolted
to the chord, not only take care of any secondary
stresses due to the action of the end-detail, but also
provide additional bolting space for the attachment of
the lower chord to the column.

The truss rests concentrically on the posts, hence
there is no bending in the post due to eccentric loads.
To accomplish this result, it is necessary that the knee-
braces be cut and framed into post and bolster after the
truss has been erected and all of the dead load of the
roof is in place. Otherwise, the slight deflection of the
truss, when the roof load is placed, will cause the knee-
braces to transfer a horizontal thrust to the post, with
consequent bending in the post. Possibly this may
seem an unnecessary refinement in timber-framing;
however, the cost of cutting and fitting the knee-braces
after the truss is erected and the roof loads are in place
is not excessive, and I believe that the slight additional
expense is justified.

Note the bolting of knee-braces to post and truss

Timber framing

TIMBER FRAMING 167

chord with cast-iron bevelled washers. The use of such
washers, set into the knee-brace, insures that the brace
is capable of withstanding some tension. If computa-
tions for wind show that the knee-braces must take
considerable tension, metal side-plates, bolted or lag-
ged to the knee-brace and to the post or truss-chord
and bolster may be substituted for the bolts shown.
Without going into the theory of wind stresses in trans-
verse framed-bents at this time, it may be said that the
exact distribution of the wind moment in the column
in this case is somewhat indeterminate, not only from
the fact that the condition of the base of the post is un-
known, that is, whether * hinged' or * fixed,' but also
from the conditions here present of three intersections
of the truss with the post, namely, the connection of the
knee-brace to the post, and the connections of the up-
per and lower chords to. the post. However; the state-
ment may be made that the maximum bending-moment
in the column occurs either at the foot of the knee-brace
or at the intersection of the lower chord with the post.
For this reason, the detail here shown, with the post
spliced over the truss, the splice-pads with a total ca-
pacity in bending equal to that of the post, and extend-
ing well below the foot of the knee-brace, and bolted
thoroughly to the post and truss, provides a condition
of maximum stiffness consistent with simplicity of
fabrication and ease of erection. If the knee-brace is
omitted, as is common in many instances, the computa-
tion of the bending moment due to wind, and the forces
acting on the bolted joints, is a simple matter. For the
purpose of presenting the conditions of wind bending
in such a case, diagrammatic representations of these
moments have been prepared and are shown in Fig.
72, a, h, and c, which illustrate the influence of the end
conditions of the columns. Fig. 72a represents the
bending for a column with *free' or 'hinged' ends; Fig.
12h shows the bending for a column with 'fixed' ends;
while F'ig. 72c illustrates the bending for an inter-

168

TIMBER FRAMING

mediate condition, namely, for a post half-way between
hinged or fixed at the base.

Fig. 71 also shows the detail of the bracing-trusses
between the main roof -trusses. Such a truss, as was ex-
plained in the preceding article, may play an important
part in resisting the wind on the building; its chords
may, in such a case, take stresses, either of compres-
sion or tension, that can be calculated with a reasonable
amount of accuracy. Its connection with the main
truss must then be carefully studied, particularly the
splices, in order that it may fulfil its part properly in
the general scheme of bracing the building. On the
other hand, the bracing trusses may take no stresses
that can be computed. Nevertheless, no roof-truss
should be constructed without due attention to the
need of bracing trusses. The detail shown here is
simple and cheap, and at th« same time is effective.

"T":a

/ I

~~^— :r

I A

'a/e/mum
Moment 'M,

/!
/ I

Mar/mum
Moment* Mj

M, > My >M^

Fig. 72. bending-moment diagram fob wind-stbesses in

COLUMNS.

Note the * T ' section of the chords, enabling them to
take compression, as well as to tie the main trusses
together.

The splice of the lower chord merits mention possibly.
Besides the steel-tabled fish-plate splice, which has a
capacity of the computed stress in the chord in this
panel, two wooden stiffening splice-pads are bolted
through the chord to give the tru^s additional stiffness
during erection. Some criticism may perhaps be made
of placing a tension-splice at the point of greatest

TIMBER FRAMING Ifid

chord-stress. While such a position for the splice is
not desirable and should in general be avoided, the
lengths of available timbers will sometimes require
placing the splice at the centre of the truss. If due
attention is given to the detail, and conservative unit-
stresses are used, there need be no apprehension of the
strength of the truss.

It should be noted that Fig. 71 does not represent a
working detail, and is not intended as such. To make
this detail a 'working' drawing, for shop and field,
sundry additional information should be given, such as
a shop drawing of the steel splices, spacing of all bolts,
etc.

Lattice-Trusses

The lattice-truss, diagrammatically illustrated in
Fig. 73, is often employed in roof construction for
moderate spans. This type of truss, as distinguished

Fig. 73. outline of half-elevation of lattice-truss.

from the open-panel type, was probably first used for
bridges. Burr and Falk, in their 'Design and Con-
struction of Metallic Bridges,' Edition 1905, note '*A
later type of timber bridge which was most extensively
used in this country was invented by Ithiel Towne in
January, 1820, which was known as the Towne lattice-
bridge. This timber bridge was among those used for
railroad structures. It was composed of a close timber-
lattice, heavy planking being used for the lattice mem-

170 TIMBER FRAMING

bers, and they were all joined by wooden pins at their
intersections. This type of timber structure was com-
paratively common not longer than twenty-five years
ago, and probably some structures of its kind are still
in use. The close lattice work with its many pinned
intersections made a safe and strong framework, and it
enjoyed deserved popularity. It was the forerunner in
timber of the modern all-riveted iron and steel lattice-
truss. It is worthy of statement, in connection with the
Towne lattice, that its inventor claimed that his trusses
could be made of wrought or cast-iron as well as timber.
In many cases timber arches were combined with
them. ' '

In the case of a railway bridge of the latticed type, the
chords of these trusses are firmly supported laterally, the
top chords by the upper lateral-bracing, and the lower
chords by the floor system, and also by the lower lateral-
bracing. In using the lattice-truss as a roof-truss,
especial care must be taken to see that the unit com-
pression-stress in the upper chord does not exceed the
safe unit-stress for the chord treated as a long column.
Due to the necessity of making the chords of a lattice-
truss deep for bolting, and the use of thin material
for the web members, the truss a& a whole is rather thin
and deep, as compared with a truss of the Howe type,
for example. It will therefore have a tendency to twist,
and must be braced accordingly.

When the roof-surface slopes from the centre of
the truss to the ends, the upper chords may be given
the slope of the roof, or the truss may be constructed
with horizontal chords, and the roof-surface furred,
either by means of a low studded-wall or short post
and roof girder. The truss with a sloping upper-chord
is somewhat more difficult to construct, as the diagonals
have different lengths, and the intersections of diag-
onals and upper chord are not uniform. For this
reason, I prefer in general to build such trusses with
parallel chords, and then to construct a studded-wall
on the upper chord. In such an event, the trusses must

TIMBER FRAMING 171

be tied together by bracing trusses or struts, with pos-
sibly the additional precaution of stiffening the upper
chord laterally by means of a 2-in. plank nailed to the
upper edges, forming a * T ' section.

It is hardly necessary to say that the lattice-truss
is an indeterminate structure, and that the exact stress-
es cannot be found by the ordinary methods for solving
the stresses in a roof -truss. It is customary to consider
the truss as a combination of a number of Warren
trusses, each taking its proportion of the total load.
The chord-stresses are, of course, the sum of the chord
stresses in the individual Warren trusses. Or, the
lattice-truss may be computed as a plate girder in
determining the approximate chord-stresses. For find-
ing the stresses in the diagonal web-members, the shear
at any section may be divided by the number of web-
systems, and the quotient resolved into the line of the
diagonal. For finding the required number of bolts or
nails fajstening the webs to the chords, the stress in the
diagonal must be resolved into the two components
parallel to and perpendicular to the chords, respect-
ively. The component perpendicular to the chords, or
the shear in the section, acts through the bolts across
the grain or fibre of the chord-timbers, and hence may
be the feature determining the size of the bolts.*

To illustrate the design of a typical latticed roof-
truss, there is here given the computations for a truss
of this type, shown in Fig. 74. The span is 40 ft., the
spacing 24 ft., and the total loading, including the
weight of the truss, 35 lb. per sq. ft. of projected hori-
zontal roof-surface. The figure gives a half elevation
of the truss. For finding the maximum stresses in the
diagonal web-members, the stress diagram shown in
Fig. 756 has been constructed.

All intersections of web-members are spiked, and

♦This is strictly true of the end-diagonals only; the size of
the bolts in the intermediate diagonals are determined from
consideration of pin action, as explained in the detail com-
putations.

TIMBER FRAMING

TIMBER FRAMING 173

the intersections of webs and chords are spiked in addi-
tion to the bolting. Fig. 74 shows the attachment of
the lateral bracing which consists of a bracing truss at
the centre of the span, and a strut between trusses at
the quarter-points of the span. For a truss of this 'span
and loading, it is not necessary to use fillers between webs
and chords, but for trusses of greater span or loading
such fillers may be required.

The lattice-truss is not an efficient truss from the
standpoint of material, but where timber is compara-
tively cheap, and steel in the form of rods, plates, etc.,
is either expensive or difficult to procure, this type of
truss may be the most economical to use. The lumber
required is all dimension stock, and may be had in any
lumber-yard, and the bolts are stock bolts. The skill
required in framing is small, and the ordinary house-
carpenter may do a satisfactory job. The weak feature
of the latticed roof -truss is usually found to be the ten-
sion-splice of the lower chord. This splice should be
carefully designed, and an ample number of bolts be
provided to take the chord stress at that point. For
determining the chord stress at any point, a simple
method is to construct a bending-moment diagram
similar to that shown for this truss, in Fig. 15a.

Computations for Lattice-Truss. Span, 40 ft.

Spacing of trusses, 22 ft. Depth of trusses : The depth

of a truss with horizontal chords, namely, the vertical

distance between the centre lines of the chords, should

be between J to ^ of the span, i being an economical

ratio. In a lattice-truss with both chords horizontal,

the depth should be a simple proportion of the span,

in order to secure good intersections of diagonals with

the chords. The depth in this case will be taken at

5 ft., which is a ratio of depth to span of h

Loading of horizontal projection of roof, 35 lb. per sq. ft.

Total load on truss = 22 by 35 by 40 ftr= 30,800 lb.

Gross reaction = i total load = 15,400 lb.

Maximum bending moment = 4 X 30,800 X 40 = 154,000 Ib.-ft.

Maximum chord stress = 1^M25 = 30,800 lb.

30800
Required net area of tension chord = ^ . = 20.4 sq. in.

174

TIMBER FRAMING

The chords will be composed of two 2 by 10-in. tim-
bers, giving a gross area of 40 sq. in., or twice the net
area required. This may appear excessive, but 2 in. is
the minimum thickness that should be used, and the

Fig. 75. diagram of bexding-moment and maximum stbesses

in members of lattice-tbuss.

width of 10 in. will give ample bolting and spiking
space, which is a vital requirement. The strength of
the truss depends upon the bolting and spiking, and
for bolting to be effective, the bolts must not be placed
too close together, nor too close to the ends or the edges
of the timbers.

Stresses in Diagonal Members. (See Fig. 7db.) This
figure represents one of the four web-systems, each
forming a Warren truss. The panel loading for one

TIMBER FRAMING

175

such system is 2^J X 22 X 35 lb. = 1925 lb. The maxi-
mum diagonal stress is 5400 lb. The same result is
reached by working from the end reaction. Thus,
dividing the total reaction by the number of web-sys-
tems, and resolving such shear into the line of the
diagonal, there results,

i^ X 1.407 = 5400 lb.

Bolting and Spiking of Diagonals. If spiking alone
were to be counted upon for fastening the web-mem-

Moment on 3ott

Forces in l/erticol Plane
(a)

Moment on Boft

forces in homontal Pione

Resuifant Moment
^ 6400*'

(c)

Pig. 76. moment-diagrams foe diagonals of lattice-tbuss.

hers to the chords, each web-member would have to
have suflScient spikes fastening it to the chord to trans-
mit the stress in such diagonal. A 30D spike is 4^ in.
long ; this size of spike is about the largest that should
be used. The safe resistance of a 30D spike to lateral
shear is 194 lb. ; eleven spikes is about the maximum
number that may be used without danger of splitting

176

TIMBER FRAMING

the timber. The maximum resistance of spiking is
therefore 11 X 194 lb. = 2134 lb. In order that both
chords may act together, they should be bolted through,
in addition to the spiking. For the web-members with
a stress of 2700 lb. (see Fig. 75 6), two f-in. bolts and
nine SOD spikes will be used. This will take care of all
panel-points except the first four from the ends.

mtr

Fig. 77. connection of truss and post.

For the first four panels from the end, the bolts will
act as pins, with forces as shown in Fig. 76, a, &, and c.
In Fig. 76a and h, the web-stresses are resolved into
their components in horizontal and vertical planes, the
reactions found, and the moments computed. Fig. 57c
shows the resultant moment to be 5400 in.-lb. Four
J-in. bolts will give a resisting moment of 4200 in.-lb. at
a flexural stress of 16,000 lb. per sq. in. Using four

f-in. bolts, then, there is required in addition — -^^ —

X 5400 = 1200 lb. to be taken by spikes directly into

TIMBER FRAMING

177

the chords from one diagonal. The detail shows six
SOD spikes, which are good for 1164 pounds.

Where the end diagonals intersect the chords and
posts, it is essential to secure a strong connection. To
accomplish this purpose, two of the end diagonals have
been made 2 by 10 in., and 1-in. bolts are used through
the post splice-pads. The actual bending on these bolts
is diflBcult, if not impossible to determine, depending
somewhat upon the efiBciency of the filler between the

Pkids.

E^G. 78. TRUSS AND POST CONNECTION.

web-member and the chord. The bolting shown will
be ample, however, to take care of the stresses indicated
by Fig. 75 &. Theoretically, the centre web-members
take no stress for uniform loading. For this reason,
they have been reduced in size to 2 by 6 inches.

Lower-Chord Splice. For determining the chord
stress at the point of splice in the lower chord, the
bending-moment diagram of Fig. 75a has been con-
structed. On a base of one-half the span of the truss, a
rectangle is drawn with a height proportional to the
maximum bending-moment at the centre. The base and
the left side of the rectangle are then divided into the
same number of equal parts, in this case, eight. From

178

TIMBER FRAMING

the upper right-hand comer of the rectangle, radiating
lines are drawn to the division lines of the left side.
The intersections of these radiating lines with the ver-
ticals erected on the base line at the points of division
determine the parabola representing the bending-mo-
ment. From this bending-moment curve, the moment
at the point of the splice is seen to be 124,500 Ib.-ft.,

and the chord stress is therefore — g — = 25,000 lb.
The detail shows fourteen J-in. bolts, which from the

Fig. 79. poorly designed tbuss and post connection.

values given in Chapter V, have a resistance of 26,600
pounds.

Upper-Chord Splice. As the upper chord is in com-
pression, a true butt-joint will be assumed, and the
splice-pads designed merely for holding the joint to-
gether, and supplying some tensile resistance for un-
known erection stresses. The post has been assumed
as an 8 by 12-in. timber. As detailed, the truss rests

TIMBER FRAMING 179

directly upon the post, and for stiffness against lateral
forces, two 4 by 12-in. splice-pads are provided, well-
bolted to the post and to the truss.

The preceding discussion illustrates the method of
design of a lattice-truss. While, as was noted pre-
viously, the stresses are indeterminate, the approximate
stresses can be found, and a reasonably rational de-
sign made. In some instances, particularly where there
is a ceiling-load to support, it may be advisable, and
even necessary, to make the chords of four pieces, in-
stead of two. Kidder's handbook gives the sizes for
lattice-trusses of various spans and spacing, and
recommends in all cases chords constructed of four
timbers. This practice, in my opinion, is not advisable,
since the deeper the chords, the more space there is
available for bolting and spiking. It might be men-
tioned that the use of four planks to each chord results
in what are termed * cumulative stresses.' In other
words, the two chord-timbers next to the web-members
must not only take their own proportion of the total
chord-stress, but must also transmit the part of the
total stress borne by the two outside chord-timbers.
This construction, therefore, results in an overstraining
of portions of the inrter chord-timbers.*

Truss Connection to Post

The method of connection of truss to post, illustrated
in Fig. 71 and 74, furnishes what in my opinion, is the
most efficient detail that can be devised. If there are

♦In this connection, note the following statement from Kid-
der's 'Architects and Builders Pocket Book,* edition 1905, page
898, "The bottom chord should also be bolted every two feet
between the joints, as this member is in tension. The top
chord, being in compression, will be tied sufficiently by the
bolts at the joints, and by a short bolt on each side of the butt
joint." This statement is misleading; in a lattice-truss with
each chord built of four sticks, the upper chord needs through
bolting to the same extent as the bottom chord. For splicing
the tension member, special bolting and spiking is required.
In general, the bolts between the joints will have to be spaced
closer than 2 ft. centre to centre.

TIMBER FRAMING

interior posts, that is, if there are several trusses across
the width of the building, this detail can only be used
for the outside, or wall posts, and the connection to the
interior post or posts, will have to be modified. Pig.
77 and 78 illustrate two methods that may be used.
Fig. 77 being for the Howe truss of Fig. 71. For this
case, the end-detail of Fig. 71 must be abandoned, on
account of lack of room, and a steel shoe substituted.

TIMBER FRAMING 181

As between the two details of connections to post
shown in Fig. 77 and 78, the particular circumstances
of the building to be framed must determine the type of
connection. The* wind shear to be transferred across
the post may require special treatment with a special
detail. The post in the detail of Fig. 78 is considerably
weakened by * gaining' the bolsters into the post, and
this reduced section is at the critical point for resist-
ance to bending stresses from wind.

Fig. 79 shows a detail of connection of truss to a wall
post which is not good, but which I have seen used to a
considerable extent. In this detail, there is eccentricity
of loading, and a consequent bending in the post, not-
withstanding the. fact that the centre lines of the post,
lower-chord and batter-post intersect in a common
point. Were the end of the chord to bear snugly
against the post, and were the iron tie-strap to provide
suflBcient tensile connection between truss and post, the
joint would not produce bending in the post. In actual
construction, however, the truss will invariably be cut
slightly short to facilitate erection. Even with an
initial snug fit, shrinkage of the post will soon destroy
this tight fit. Similarly, the post will shrink away
from the yoke, and the value of the tie-strap be largely
lost. The detail thus gives a false impression of stiff-
ness. It is true that the joint may be tightened after
shrinkage has taken place by shimming and wedging,
but the chance of this extra work being done after the
building is completed is small, and ^ny connection
which minimizes the effect of shrinkage is to be pre-
ferred.

It is sometimes instructive to learn how not to do
things. Fig. 80 is a detail of a truss and post connec-
tion used in one of the concession buildings at the
Panama-Pacific International Exposition. It is repro-
duced here to illustrate how it is possible to use a great
quantity of material without obtaining great strength,
and especially without gaining an appreciable amount
of stiffness. Expensive construction does not neces-

182 TIMBER FRAMING

sarily mean strength. The principal defect of the con
struction shown is its lack of stiffness. The only ties
between truss-chords and posts are the small steel tie-
straps or yokes, fastened with lag-^screws. As these
yokes bear across the fibre, of the posts, their maximum
resistance limited by this stress is 36 X 300 = 10,800 lb.
As a matter of fact, the pressure across the face of the
post would never be uniform, as the strap is not stiff
enough to so transmit the pull. The pressure would
all be concentrated near the sides of the post, and
would crush the corners of the post, should any amount
of pull come upon the strap. As shown above, shrink-
age of the timber would soon destroy the efiBciency of
these yokes.

Both end-details are objectionable. The double cut
on the end connection of the left truss, with a small
shearing area is ineflBcient. Double cuts similar to this
introduce cumulative stresses, as the total horizontal
component of the thrust of the batter-post must event-
ually come upon the shearing area between the inner
or lower end-cut of the batter-post and the end of the
bolster. The cast-iron shoe of the truss on the right is
poorly designed. Here again the two different depths
of lugs introduces cumulative stresses in shear. The
thickness of 1 in. for the first lug with a depth of 2 in.
is altogether too small. If the full stress ever came
upon this lug, it would fail through flexure. For a unit-
bearing pressure of 1600 lb. per sq. in. this lug would
be stressed in flexure to 31,200 lb. per sq. in., acting as a
cantilever. No bolts are provided to hold the inner lug
in its cut in the timber.

While, because of the large live load figured on the
truss, and the safety factor, it was in no danger of failure,
the designs, of which this is an example, are not only un-
economical, but the owner of such a building is not get-
ting security in proportion to money expended. A
stronger and stiffer structure could have been secured at
a less expense. When a competent engineer checks such a
design, and points out, for example, the weakness of the

TIMBER FRAMING 183

tie-straps or the end-details, he is sometimes accused of
attempting to add material unnecessarily, and to prove
the claim, the sizes of the different members are pointed
out, as sufficient evidence of the safety of the structure.
Sometimes, one of the most difficult things to make an
owner realize, is that heavy members of trusses and
posts do not necessarily indicate a strong construction.

184 TIMBER FRAMING

CHAPTER XI

Theory of Column-Action — Tests of Timber Columns

Considered from the standpoint of safety of eon- .
struetion alone, the design of a solid timber post to
support a concentric vertical load is merely a question
of selecting the column formula to be used, and provid-
ing the required area as determined by this formula.
In this respect, a column is no different from a strut of
a timber truss. Colimin formulas were discussed to
some extent in Chapter X, and working values may be
selected from those formulas.

Theory of Column Action

The phenomenon of column action is best established
by the Bankin or Gordon formula, and without attempt-
ing to go extensively into the mathematics of this * theo-
retical' formula, it will be of interest to examine briefly
the history and derivation of the component parts of the
expression. A full discussion of column action may be
found in any standard work on structural mechanics, for
example, Merriman's * Mechanics of Materials,' Church's
* Mechanics of Engineering,' Burr's *The Elasticity and
Resistance of the Materials of Engineering,' and others.
The following discussion is presented for the purpose of
emphasizing the importance of the effect of eccentric
loading, by calling attention to what many engineers
are inclined to forget, namely, that bending is an im-
portant part of long-column action even with concentric
loading. I cannot do better than quote from the text of
Burr mentioned above, as follows: ''There is a class of
members in structures which is subjected to compressive
stress, and yet whose members do not fail entirely by
compression. The axes of these pieces coincide, as nearly
as possible, with the line of action of the resultant of the
external forces, yet their lengths are so great compared
with their lateral dimensions, that they deflect laterally,

TIMBER FRAMING 185

and failure finally takes place by combined compression
and bending. Such pieces are called 'long columns/ and
the application to them, of the common theory of flexure,
has been made in Article 24.'' And from Article 24, ''A
4ong column' is a piece of material whose length is a
number of times its breadth or width, and which is sub-
jected to a compressive force exerted in the direction of
its length. Such a piece of material will not be strained
or compressed directly back into itself, but will yield
laterally as a whole, thus causing flexure. If the length
of a long column is many times the width or breadth, the
failure in consequence of flexure will take place while the
pure compression is very small." Mr. Burr then de-
velops Euler's formula for long columns, which is

P = j^3 , where E is the modulus of elasticity and

I is the moment of inertia.

* * It is to be observed that P is wholly independent of the
deflection, that is, it remains the same, whatever the de-
flection, after the column begins to bend. Consequently,
if the elasticity of the material were perfect, the weight
P would hold the column in any position in which it
might be placed, after bending begins." The above
formula is the basis of *Hodgkinson's formula,' for the
resistance of long columns.

'*Two different formulas were first established for
use in estimating the resistance of long columns; they
are known as * Gordon's formula' and 'Hodgkinson's
formula. ' . Neither Gordon nor Hodgkinson, however,
gave the original demonstration of either formula. The
form known as Gordon's was originally demonstrated
and established by Thomas Tredgold — while that known
as ' Hodgkinson 's formula' was first given by Euler. In
1840, however, Eaton Hodgkinson, F.R.S., published the
results of some most valuable experiments made by him-
self on cast and wrought-iron columns — and from these
experiments he determined empirical coefficients ap-
plicable to Euler 's formula, on which account it has
since been called Hodgkinson 's formula. Mr. Lewis

186 TIMBER FRAMING

Gordon deduced from the same experiments some em-
pirical coefficients for Tredgold's formula, since which
time it has been known as Gordon's formula.''

With this brief history of the origin of these famous
column formulas, we may go at once to the derivation of
Gordon's formula for long columns. ** Since flexure
takes place if a long column is subjected to a thrust in
the direction of its length, the greatest intensity of the
stress in a normal section of the column may be consid-
ered as composed of two parts. In fact the condition of
stress in any normal section of a long column is that of
a uniformly varying system composed of a uniform stress
and a stress-couple. In order to determine these two
parts, let S represent the area of the normal cross sec-
tion ; /, its moment of inertia about an axis normal to the
plane in which flexure takes place; r, its radius of gy-
ration in reference to the same axis; P, the magnitude
of the imposed thrust ; f, the greatest intensity of stress
allowable in the column; and D, the deflection corre-
sponding to /. Let p' be that part of / caused by the
direct effect of P, and p" that part due to flexure alone.
Then, if h is the greatest normal distance of any ele-
ment of the column from the axis about which the mo-
ment of inertia is taken, by the common theory of flexure,

c^PD = '~jr-y therefore p = — j —
Also,

p' = -^, therefore p' -f p'' = /= -^ ( ^ + ^ — )

fs
Hence, ^"^1+ ^'^^^

This equation may be considered one form of Gordon's
formula.

Burr then shows that D = a —, in which expression a

is a constant. Making this substitution, and expressing 1
in terms of 8 and r, there results the formula,

1 -j- a —
r2

TIMBER FRAMING 187

The preceding treatment illustrates clearly that col-
umn action, for long columns, or, as has been stated, for
timber columns whose length is greater than twenty times
the least cross dimension, consists of a uniform compres-
sion plus a cross-bending. Since there is a flexural stress
on such columns for concentric loads, it is obvious that
the addition of bending moment due to eccentric loading
decreases the strength of the column, and that the pro-
portion of flexural stress to the total stress for eccentric-
ally loaded columns is greater the longer the column in
proportion to its least width. Further, it may be said
that eccentrically loaded columns are in the realm in
which the fewest tests for strength have been made.
It follows, then, almost as an axiom, that eccentrically
loaded columns should be avoided wherever possible, and
that where they must be used, careful consideration of
the maximum combined unit-stress must be made in
order that such maximum unit-stress shall not exceed the
safe unit-stress for the column.

Uneven ends on columns, or ends not exactly normal to
I the axis of the stick will produce eccentricity. In fabri-
cating steel columns, care is always taken to mill the ends
of the column to a true and even bearing, and the bearing
or base plates are usually planed to an even surface and
a uniform 'thickness. On the other hand, the timber
' column even though it may carry heavy loads, as in ware-
house or heavy mill-construction, is at best trued by a
carpenter's square.

The ultimate strength of timber columns is not a mat-
ter of definite knowledge, because of the lack of suffi-
cient tests on full-sized columns. This is especially true
of long columns; for example, columns whose length is
from 40 to 60 times the least width. It is interesting to
note that the formulas of the American Railway Engi-
neering Association give a value of zero for the safe

working stress for a column whose -^ is 60. On the other

hand Ketchum would allow a working stress of 480 lb.
per sq. in. for this column, and the formula of the U. S.

188 tiMBGR FRAMtN(!^

Department of Agrieulture, Forestry Division, gives a
value of approximately 500 lb. per sq. in. for the same
column, when G is taken at 1600 lb. per sq. in. This
wide variation in recommended working values is un-
fortunate, and might cause the layman to believe that
the engineer's formulas were worthless. The practical
meaning of this variation is that the actual strength of
columns with a large g- is uncertain. Columns with a
greater -J than 20 will generally fail by lateral buckling,
a fact which has been definitely proved by tests on full

Teste of Timber

The published data on the strength of full-sized tim-
ber columns is meagre; practically all of the recorded
tests were made by the United States Government at the
Watertown arsenal. Undoubtedly some other tests on
small columns have been made in technical schools and
colleges, but the results of these are not generally known.

In Fig. 81 the ultimate strengths of the timber col-

TIMBER FRAMING 189

umns tested at the Watertown, arsenal are plotted. These
values were taken from the digest of the tests made by
J. B. Johnson and W. H. Burr. As these tests were
published in 1882, I have added the results of some sub-
sequent tests made at the same laboratory, also a few
tests made by W. L. Huber on small Douglas fir columns,
but with a wide variation in the ratio of length to least
width. These sticks were 1.7 in. square, and the tests
were a part of the regular course in the testing labora-
tory of the University of California. The Watertown
arsenal tests of 1882 were on yellow pine and white
pine. The size of these columns varied from 5.3 by 5.3
in. by 27 ft. 6 in. to 8.25 by 8.25 in. by 15 ft. For the
case of the white-pine tests, I have arbitrarily increased
the recorded values by the ratio 1.52 in order to give
more data on the variation of strength with the change

in ^. This procedure is in error to some extent; it

would be correct only if the ratio of the compressive
strengths of the two timbers was the same as the ratio
of the respective moduli of elasticity in bending, and if
the moduli of elasticity bore a constant relation to the
respective compressive strength throughout the range of
the tests. It will be seen, by referring to Pig. 14 of
a preceding chapter that the ratio of the compressive

strengths of long-leaf yellow pine to white pine is jr^ =
1.475, while the ratio of the respective moduli of elastic-
ity is ]^]^3()0QQ = 1.425 ; the average is 1.45. The Water-
town arsenal tests on short columns of the same timbers,
that is, columns which failed by compression alone with
no lateral deflection, showed the average ultimate
strength of yellow pine to be 4442 lb. per sq. in., while
the same quantity for the white pine was 2414 lb. per
sq. in. These figures give a ratio of 1.84. The ratio
used (1.52) is the average of the ultimate strengths of

the columns with an "t^^ 22 and over. The difference
between the two figures shows the infiuence of bending.

190 TIMBER FRAMING

Table XVIII gives the regults of some tests on Douglas
fir columns published in * Tests of Metals/ 1896. These
results are also incorporated in Pig. 81.

Table XVIII

TESTS ON DOUGLAS FIB COLUMNS*

Least

Ultimate

width.

strength,

Modulus of

in.

Length

lb. per sq. in.

elasticity

8.12

26 ft.

0.125 in.

38.4

2600

1,651,000

10.12

29.6

3700

1,875,000

10.04

29.9

2700

1,785,000

8.18

29.3

3371

1,704,000

10.10

8.00 in.

19.7

3500

1,875,000

10.06

8.00 "

19.9

3700

1,639,000

8.21

3.75 "

19.5

3600

1,756,000

10.12

6.00 "

14.8

3900

1,854,000

10.08

6.00 "

14.9

3400

1,393,000

8.08

11.8 "

14.9

4249

1,791,000

10.12

3.90 "

9.9 \

4312

1,743,000

9.98

4.05 "

10.0

4138

1,792,000

10.07

8.13 "

8.0

4100

1,963,000

7.92

7.98 "

9.9

2600

1,904,000

10.07

1.94 "

5.1

4626

1,675,000

8.13

4.00 "

5.0

3988

• •••••••

♦Tests of metals, Watertown Arsenal.

Various Formulas for Ultimate Strength. In Fig. 81
are shown some of the various formulas for ultimate
strength of yellow-pine timber columns. W. H. Burr,
from the results of the Watertown arsenal tests advocates
the following straight line formula

p = 5800 - 70 "^ ^ p being the ultimate strength in

pounds per square inch, this formula to be used only

between the limits 20 ^ and 60 -^ .

On the basis of the same tests, J. B. Johnson proposed
the parabolic formula

p = 4500 -1.0 -^, this formula to be used between

the limits -j- == 1 and -^ == 50. At the latter limit the
parabola is tangent to the curve of Euler's formula

AirEI

p = -jj- when E = 1,620,000 lb. per sq. in. This

TIMBER FRAMING 191

formula is for partially seasoned yellow-pine columns.
For dry long-leaf yellow-pine columns, he proposed the

■formula p = 6000 - 1.5 ^ .

The U. S. Department of Agriculture formula, is also
shown, with C = 4500 lb. per sq. in.

W. H. Burr in his text already quoted, states that
some 1200 tests on full sized specimens of square and
rectangular yellow-pine columns were made by C. Shaler
Smith for the Ordnance Department of the Confederate
Government, and that the results indicated that the' fol-
lowing formulas represented the ultimate strengths of
the columns.

1. For green, half-seasoned sticks answering to the
description, *good merchantable lumber'

5400
"^250 (J2

2. For selected sticks, reasonably straight and air-
seasoned under cover for two years and over

8200

300 d2

3. For average sticks cut from lumber which had been
in open-air service for four years and over

5000
P= 1+ J_^

250 d2

The tables for strength of timber columns as given
in Trau twine's * Handbook' are based on the Shaler
Smith formulas. These formulas are of the Rankin or
Gordon form. It is of interest to note that the curves
of the Shaler-Smith formulas do not fit any of the tests
of the Watertown arsenal, as may be seen by reference to
Fig. 81, where the last formula has been plotted.

In Fig. 81, I have plotted a curve of the Rankin-
Gordon type which seems best to fit the results of the
tests there shown, and find as noted in the figure, that
the coeflScient a has a value of about 1750 instead of 250,
as found by Mr. Smith. As no numerical results of Mr.

192 TIMBER FRAMING

Smith's tests are to be found, no comment can be made
with regard to the difference between his proposed
formulas and those of later engineers.

The elastic limit is high in proportion to the ultimate
strength in a timber column. The average ratio as
measured by the stress-deformation curves of Mr.
Ruber's tests is about 84%, while on some similar tests
on 3^ by 3^ in. redwood columns, I found the propor-
tion about 90%.

On the basis of the proposed column formula,

5000
1750 d2

there is given in table XIX, the ultimate strength of
D.ouglas fir timber columns, for the case of partially
seasoned timber of the No. 1 common grade.

Table XIX

ULTIMATE AND WORKING STRENGTHS OF DOUGLAS FIR COLUMNS

5000

Formula: p^ 1 I^

^ "^ 1750 d^

^' Strength in lb. per sq. in.

d Ultimate Working

10 4740 1355

12 4630 1325

14 ., 4510 1290

16 4350 1245

18 4210 1205

20 4060 1161

22 3910 1120

24 3760 1075

26 3600 1030

28 3450 986

30 3310 946

32 3150 902

34 3020 864

36 2880 823

38 2740 785

40 2620 750

42 2490 712

44 2370 678

46 2260 646

TIMBER FRAMING I93

_ Strength in lb. per sq. in.

d Ultimate Working

48 2160 618

50 2060 589

52 1960 560

54 1870 535

56 1790 512

58 1715 490

60 1635 421

Working Strength of Timber Columns. Taking into
consideration the adverse effect on the strength of timber
of knots or oblique grain, the possibility of uneven end-
bearing, eccentricity due to imperfect beam or girder
connections, the eflEect of long-continued loads of large
magnitude, and the relatively few tests on full-size sticks
of a large ratio of length to least width, a safety factor
of 3^ on the basis of ultimatie strength as given by the
tests quoted above would seem to be the lowest that
should be used, and this only for buildings. The factor
should be increased to five for unprotected structures,
such as bridges or other outdoor construction.

lable XIX also gives the safe unit stresses for build-
ings based on the modified Rankin-Gordon formula with
a = 1750^ and a safety factor of 3 J.

194 TIMBER FRAMING

CHAPTER XII

Column Splices and Oirder Connections — Floor Oirders
and Joists — Joist Hangers — Mill Construction

Column Connections. Other conditions than the al-
lowable stress under column action often, determine the
size of a post in a timber-framed building, for example,
the required cross-sectional area at the ends of the post to
provide bearing area for the beams, girders, or trusses
resting on the post, or the requirement of a general
minimum size of column to give the proper stiffness to
the building.

Except in the case of columns supporting floors or
roof-bays of uniform size, the ideal condition of con-
centric loading will seldom be realized. Unless care is
taken in the detailing of connections, wall columns will
usually be loaded eccentrically, producing bending in the
posts, the amount varying not only with the numerical
value of the load and its eccentricity, but also with the
nature of the connections. The case of truss connections
to posts was discussed in*the preceding chapter, where it
was pointed out that many details involve considerable
resultant bending.

Fig. 82, a, h, and c illustrate details sometimes seen
in building designs. The defects in these three details
are self-evident. In a, it is almost certain that the girders
have not sufficient bearing area to prevent crushing of the
fibres. If the upper post is working at an efficient stress,
the fibres at the top and bottom of the girders must be
stressed above their elastic limit. This condition will
produce settlement of the upper floors, which, added to
the shrinkage of the timbers, will crack plaster walls, or
produce uneven floors.

In Fig. 82& sufficient area for bearing is given to
the girders by the bolster, but both the top and the bot-

TIMBER FRAMING

195

torn of the bolster are probably over-stressed in cross-
bearing. The shrinkage in this case will be that of the
bolster only.

Fig. 82c shows the most defective details. Here the
settlement because of shrinkage is the greatest.

With the use of a hardwood bolster, the crushing of
the fibres of the bolster may be reduced, and possibly
eliminated, although.it must be remembered that even
oak has an elastic limit across the grain of only approxi-

^ -^

r~iTr~i

^^^

(C)

^VJ

CO)

(b)

^£7

Fig. 82. examples of defective details.

mately 920 lb. per sq. in., for green timber, or about
50% greater than Douglas fir.*

To overcome the disadvantages of wooden bolsters,
metal post-caps of cast-iron, wrought-iron or steel are
commonly employed. Standard post-caps, usually of
pressed steel, can be bought in the open market. Typical
details of post-cap framing, are shown in Pig. 83, the
illustration being taken from 'Structural Timber,' En-
gineering Bulletin No. 2, published by the National

♦These values are from the table of unit stresses adopted by
the American Railway Engineering Association, as given in
Table I, Chapter III.

196

Timber framing

CO
00

1 !

—

^ — -

Po~5

■

TIMBER FRAMING 197

Lumber Manufacturers Association. Some of the more
common of these post-caps are the Duplex, Goetz, Van
Dorn, and on the Pacific Coast, Falls caps, and others.
The prices of these vary considerably, and on a large
job, it may be possible to build up structural post-caps
that will meet the requirements and at the same time be
cheaper. Four-way post-caps are open to the objection
of resulting in unequal shrinkage, where wooden girders
are used, since the joists supported by the girders will
drop an amount equal to the shrinkage of the girder,
while the joist or beam resting on the post-cap will not
drop. This will occur even with the use of joist-hangers,
except that when hangers of the Duplex type are used,
the shrinkage will be only half of that of the type of
hangers which fasten over the top of the girders, since
the Duplex joist-hanger is secured to the girder by
means of a circular nipple inserted into the girder at
slightly above the centre of the depth of the girder.

Where the absence of ceiling will permit, the details
of joints shown in Fig. 84, a, b, c, and d, will be found,
on analysis, to be free from the defects of the connec-
tions shown in Fig. 82. The bolster-blocks are either
dapped into the lower post, or bolted and keyed. In
either case, they have end bearing, while their section
may be large enough to provide ample bearing for the
girders. Where the size of posts decreases with the suc-
ceeding stories, the trimming of the end of the lower
post to the section of the upper post will ordinarily pro-
vide sufiicient area for the bearing of the bolster blocks.
The size of the bolts in Fig. 84, a and b, may be de-
termined by taking moments about the centre of the
lower bolster bearing. Thus, in Fig. 84a, neglecting
the eflEect of the lower bolts, because of their short dis-
tance above the end of the bolster, the tension in the
upper bolts may be found from the equation

T =

where P is the reaction of the girder, a is the horizontal
distance from the centre of the upper end of the bolster

198 TIMBER FRAMING

to the centre of the gain in the post, and k is the dis-
tance from the upper bolts to the lower end of the block.
The working stress for the circular keys or pins may
be taken from the tests mentioned in the preceding
chapters. Pipe pins, 2 in. external diameter, and of
extra-heavy section may be considered good for 800 lb.
per lin. in. of pin. Oak, as has been shown, is prac-
tically worthless, and the same is true of gas-pipe. The

bolts should be designed to take a tension equal to the
reaction of the girder.

It will be noted that in these details, the normal spac-
ing of the joists has been modified at the post, to pro-
vide a joist at either side of the post. This is an inex-
pensive way to secure considerable stiffness in the build-
ing. The two joists are to be either spiked or bolted to
the post, as the requirements for stiffness may warrant.
In Pig. 84d the girders are shown tied together across

TIMBER FRAMING 199

the column by means of two wooden splice pads. The
two sections of posts may be similarly tied together by
the use of splice-pads, with fillers under them, of the
thickness of the girder-splices. A one-inch thickness of
girder-pad will usually give sufficient tie, if it is long
enough to give the required spiking or bolting area.

Connection of Joists to Girders. The cheapest and
most satisfactory manner of supporting floor joists is to
rest them upon the girders. There is no device that is as
satisfactory as putting the support directly under the
load, without resorting to bending or shearing of metal.
In buildings with wooden or corrugated-steel walls, the
extra height of building and consequent expense result-
ing from this form of construction will be justified. On
the other hand, in the case of a building with masonry
walls, and several stories high, resting the floor joists on
top of the girders in place of attaching them to the gird-
ers by means of metal hangers may add six or seven feet
to the height of wall. From the standpoint of cost of
construction alone, the cost of the extra walls may be
considerably more than the cost of the necessary joist-
hangers.

The danger of unequal shrinkage resulting from the
use of hangers has been mentioned. On the other hand,
when all the joists rest upon the girders directly, while
the floors will settle uniformly through shrinkage, the
floors will not remain level, since the wall-bays will drop
at their inner ends the amount of the girder shrinkage
plus the joist shrinkage, and the outer or wall ends will
settle the amount of the joist shrinkage alone.

"Where it is found necessary or advisable to employ
joist-hangers, special hangers may be designed, or some
of the standard makes on the market may be used. The
standard makes may be divided into two classes, those of
the duplex type which, as has been mentioned, are se-
cured to the girders by meaps of an inserted nipple, and
those which fasten by arms or straps which fit over the
tops of the girders. The two types are illustrated by
Fig. 85 and 86.

200 TIMBER FRAMING

Joist-hangers should not be . used indiscriminately,
that is, without investigation as to their fitness for the
particular case, and their ability to withstand the par-
ticular loads. Engineering News of November 20, 1902,

Fig. 86. dttpu:x joist-hanoeb.

Vol. 48, page 420, describes the collapse of a building in
Minneapolis through failure of joist-hangers, although
these were of special design, and not standard hangers.
The late F. E. Kidder discusses the general question of

TIMBER FRAMING 201

joist-hangers, in connection with this failure, in the sub-
sequent issue of January 15 and February 5, 1903, Vol.
49, Engineering News, Kidder notes several tests as fol-
lows: (1) a standard hanger of the second clas^ men-
tioned above made of f by 24-in. wrought iron, which
failed at 13,750 lb., or a unit stress in tension on the
iron of 7333 lb. per sq. in. ; (2) a Van Dom hanger (type
2), where the arms began to straighten out at 13,300 lb.,
and failed at 18,750 lb. ; (3) a double stirrup of f by 2^-
in. wrought iron carrying two 8 by 12-in. timbers over
one 12 by 14 in. failed at a load of 28,825 lb. on each
side, or at a tensile stress of 15,273 lb. per sq. in. ; (4) a
duplex No. 35 hanger with a nipple 2J in. in diameter
and 3i in. long, broke under a load of 39,950 lb. The
bearing under the lower half of the nipple was 1977 lb.
per sq. in., yet the compression on the wood and the
effect on the girder was slight. The hanger failed by
breaking of the iron directly under one of the nipples.
Another duplex hanger of the same size failed at- 38,000
pounds. These hangers are shown in Fig. 87.

Kidder points out that the first point of weakness in a
joist-hanger of the stirrup type is the bending of the top
strap, and the crushing of the fibres on the joist side of
the top of the girder ; the second point of weakness is the
bending of the bottom of the stirrup supporting the
joist, or the tendency to shear. The tests quoted above
show that the metal of a joist-hanger does not fail by
direct tension. ♦

Referring to the first test noted by Kidder, the equiva-
lent load on the 6 by 12-in. beam would be 26,000 lb. A
6 by 12-in. beam on a 10-ft. span is good for 17,100 lb.
at a maximum fibre stress of 1800 lb. per sq. in.; the
safety factor was therefore approximately 1^. The
double stirrup of f by 2i in. failed at 28,825 lb. on each
side, or at an equivalent load on the beam of 57,650 lb.
An 8 by 12-in. beam on a 10-ft. span will carry 22,800
lb. ; the safety factor was therefore 2 J.

In a catalogue of a standard joist-hanger there is pub-
lished the result of some tests made for the company by

202 TIMBER FRAMING

TIMBER FRAMING 203

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204 TIMBER FRAMING

a firm of testing engineers. The letter from the engi-
neer is reproduced in the catalogue, and I quote the fol-
lowing significant statement: **The joist-hangers being
tested on this occasion were taken from their regular
stock, and of the following sizes, 2 by 12-in. and 4 by
12-in. They were mounted on heavy headers luith thin
pieces of plate to prevent the arms from crushing the
rough pine used and in order that the strength of the
hanger could be tested, and not the lumber.''*

In the same catalogue is published a similar letter of
later date, embodying the results of further tests. These
results are given in Table XX, in which I have also
noted the equivalent load that the corresponding beam
would stand with a 10-f t. span at a maximum fibre stress
of 1800 lb. per sq. in., and the safety factor of the beam
at the initial load (load at which the arm of the hangers
began to rise). Certain extracts of the letter are also
quoted as follows :

• *' All of the above hangers were mounted on eucalyptus
timbers,! adhering as nearly as possible to the usual form
of construction, and were spiked to the headers, after
which the joists were dropped into place.

"Two readings were taken : The first one (initial load)
at the moment the arm of the hanger began to rise, and
the final (maximum load) when the arms straightened,
and the timbers crushed so that further recording was

*The italics are my own ; this point is not emphasized in the
catalogue.

tThe strength of eucalyptus timber, grown in California, was
investigated by B. L. Soul6 and Thomas Williamson in 1904,
as thesis work at the University of California in a series of
95 tests. The elastic limit of the timber in crushing, at right
angles to the direction of the fibres, was found in 19 tests to be
as follows: Maximum 1679 lb. per sq. in., minimum, 964 lb.
per sq. in., and average 1368 lb. per sq. in. The corresponding
moisture content was 49.6%, 36.4%, and 48.4%. The results of
the tests on the hSngers are not, therefore, directly applicable
to Douglas fir, which, when green has an elastic limit across
the fibres of 570 lb. per sq. in., as against 1368 lb. per sq. in. of
eucalyptus. The strength of a joist-hanger is just as much a
question of the capacity of the timber as of the hanger.

TIMBER FRAMING 205

useless. Crushing of the timbers occurred when testing
che 6 by 12-in. and 10 by 14-in. sizes."

The values in the table representing the unit bearing-
pressures on the bottom of the joists are not necessarily
the correct values; in fact, it is certain that they are
incorrect. The values there given are computed on the
basis of an even distribution of loading over the seat of
the hanger. This condition probably existed in no case,
but the values for the thin joist are more nearly correct
than for the thicker joists, as the ratio of thickness of
metal to thickness of joist (span of seat of hanger acting
as a beam) is much greater than in the case of the large
joists. In the latter instances, the pressure would all be
concentrated at the sides of the joist, and the unit pres-
sure may have been three or four times that given in the
table.

One other point is of interest to mention, in connection
with Kidder's report of the test of the duplex hanger.
He states, as was noted, that the unit bearing-pressure
of the 2J by SJ-in. circular nipple was 1977 lb. per sq. in.
at the time of failure, with but small effect in compres-
sion on the wood. H. S. Jacoby, in his * Structural De-
tails' notes with respect to the duplex or Goetz hanger,
*4t may be assumed, according to the results of tests,
that the safe load is limited only by the safe bearing-
value of the cylindrical bearing surfaces on the sides
of the fibres of the supporting beam. As shown, the
effective bearing area equals the horizontal projection
of the cylindrical surface when the direction of pressure
is perpendicular to the fibres.'' If the average unit
pressure of a cylindrical metal pin be taken as the limit-
ing pressure perpendicular to the grain, the nipple of
the hanger (for longleaf pine) would have crushed the
girder at a load of 10,500 lb. for green timber, and pos-
sibly 15,000 lb. for dry timber. This is additional evi-
dence that such a consideration of cylindrical bearing is
in error, as was already discussed in Chapter IV. On
the basis of the theory there proposed, Kidder's com-

206 TIMBER FRAMlNa

puted unit bearing-pressure would represent approxi-
mately the elastic limit of the timber.

Calculations

Capacity of two circular nipples, 2i by 3i in., according to
usual method = 2 X 2| in. X 3i in. X 520 lb. per sq. in. X
1.50. (Increase of 50% for probable condition of seasoned
timber; value of 520 lb. per sq. in. is from Table I, Chapter III,
for longleaf pine, green condition of timber, elastic limit in
compression across the fibre) == 15,650 pounds.

Capacity of same nipples by method of Chapter IV = 2 X 2 J
in. X 3i in. X [(S X 520) + (i X 3500)] X 1.50 = 45,600 lb.
(The figure 3500 is the elastic limit for green longleaf pine,
as taken from Forest Service Bulletin 88.) The actual load
on the hanger as tested was 39,950 pounds.

Mill Construction. The preceding discussion and
the illustrations of details have not considered the ques-
tion of fire risk, and some of the details are open to criti-
cism from this standpoint. This subject is one that is
treated to considerable extent in Kidder's 'Pocket Book/
and in other books. Engineering Bulletin No. 2, of the
National Lumber Manufacturer's Association, entitled
'Structural Timber, Heavy Timber Mill Construction
Building,' dealing with this construction, has been
issued recently. Valuable information, including many
tables for strength of timber structural members, is also
given in the 'Structural Timber Hand Book on Pacific
Coast Woods,' publisbed by the West Coast Lumber-
men's Association, with headquarters at Seattle.*

From the standpoint of fire-protection in buildings
with a timber-framed interior construction, and brick
walls, it is advisable to have all sections of beams, gird-
ers, and posts of as large section as practicable, even at
the cost of economy in framing. Beams framing into

♦Both of these publications are of great value to those en-
gaged in timber construction; the former for its presentation
of the requirements of mill construction in accordance with
the standards of the National Board of Fire Underwriters, and
the latter for its many tables of strength of beams and
columns.

TIMBER FRAMING 207

walls should be self-releasing in case of fire, so that if
the timber beams burn and fall, they will not pull the
wall with them. Similarly, many standard post-caps are
designed with the idea that the girders will pull out of
their seats, if they burn and fall, without pulling down
the post with them. It may be of interest to quote the
definition of mill construction from the bulletin of the
National Lumber Manufacturers' Association. This type
of construction is divided there into three classes as
follows :

*'l. Floors of heavy plank, laid flat upon large girders
which are spaced 8 to 11 ft. on centres. These girders
are supported by wood posts or columns spaced from 16
to 25 ft. apart. This type is often referred to as * Stand-
ard Mill Construction. '

*'2. Floors of heavy plank laid on edge and supported
by girders which are spaced from 12 to 18 ft. on centres.
These girders are supported by wood posts or columns
spaced 16 ft. or over apart, depending upon the design
of the structure. This type is called *Mill Construction
with Laminated Floors.'

**3. Floors of heavy plank laid flat upon large beams
which are spaced from 4 to 10 ft. on centres, and sup-
ported by girders spaced as. far apart as the loading
will allow. These girders are carried by wood posts or
columns located as far apart as consistent with the gen-
eral design of the building. A spacing of from 20 'to
25 ft. is not uncommon for columns in this class of fram-
ing where the load is not excessive. This type is more
generally known as * Semi-Mill Construction'."

Also, from the Building Code Kecommended by the
National Board of Fire Underwriters:

** Wooden girders or floor timbers shall be suitable for
the load carried, but in no case less than 6 in., either
dimension, and shall rest on iron plates on wall ledges,
and where entering walls, shall be self -releasing. "Walls
may be corbelled out to support floor timbers where
necessary. The corbelling shall not exceed 2 inches.

*'So far as possible, girders or floor timbers shall be

208 TIMBER FRAMING

single sticks. Width of floor bays shall be between 6 and
11 feet.

*'The practice in mill-construction of supporting the
ends of beams on girders by means of metal stirrups or
bracket hangers is objectionable. Experience has shown
that such metal supports are likely to lose their strength
and collapse when attacked by fire.*

** Floors shall not be of less than 3 in. (2J in. dressed)
flooring laid crossways or diagonally."

♦No difference in insurance rate, however, will be made for
this factor alone.

TIMBER FRAMING 209

CHAPTER XIII

Foundations

The three cardinal principles of foundation design
are met when (1) the safe bearing-pressure on the soil
is not exceeded, (2) all footings exert the same pressure
per unit area on the soil, and (3) the individual footings
are each strong enough to withstand the loads coming
upon them. The fulfilment of the preceding conditions
involves the careful calculation of all loads coming upon
the several piers or wall-footings, and the proportioning
of the details so that such piers and footings may be
strong enough in all their parts to distribute the in-
dividual loads with safety.

To an engineer, these principles are so self-evident
that it seems redundant even to mention them. In
structures of importance, such as bridges and steel-
framed buildings, careful attention is paid to all these
considerations. In timber-framed buildings, the second
principle, that of providing equal bearing on all foot-
ings, is commonly neglected. The footings of a timber-
framed building, unlike those of a steel-framed build-
ing, are usually the last details to be designed. The
common practice is to compute approximately the maxi-
mum load on any one pier, design this pier accordingly,
and either make all the others the same size, or else to
establish their dimensions arbitrarily.

In a similar manner, the sills of walls are often made
of the same size throughout the building, although the
different walls will probably carry widely varying
loads. Much of the cracking of plaster walls in dwell-
ing-houses is on account of the unequal and often iur
sufficient bearing of the foundation on the soil.

Foundations may be divided into two kinds for the
purpose of this discussion, permanent and temporary.

210 TIMBER FRAMING

It is not my intention to discuss the design of permanent
footings. At the present time these are usually built of
concrete, either plain or reinforced. Brick is also some-
times used. Discussion of the design of concrete foot-
ings and piers may be found in the numerous texts on
concrete and reinforced concrete.

When a timber post rests on a concrete footing, it may
be necessary to use a steel base-plate under the post.
This will serve two purposes, first, to distribute the load
of the post over the concrete, in order that the safe unit
compressive stress be not exceeded, and second, that
there may be an impervious surface between the con-
crete and the ends of the fibres of the timber. With re-
gard to the first consideration, it must be remembered
that timber can safely withstand a unit pressure of 1600
to 1800 lb. per sq. in., in end bearing, while concrete
should not be stressed in compression over 350 to 450 lb.
per sq. in. Standard steel base-plates, of several different
makes, may be purchased, or a plain plate may be used.
In either case, the plate should be well painted. .Further,
the bottom of the post should be treated with a good
brand of wood-preservative. In no case should the end
of the post be allowed to rest directly upon the concrete,
as moisture will attack the post, and cause decay of the
timber. The standard base-plates are fabricated with
lugs fitting closely around the sides of the post. If a
plain base-plate is used, it will be advisable to provide
a dowel, embedded in the concrete base, and extending
an equal distance into both the concrete and the post.
The dowel may be a short piece of round steel rod, say
IJ in. diam. by 6 in. long, or else a piece of heavy or
extra-heavy steel pipe. In general no dowel should be
used with a diameter of less than one inch.

Timber Foundations. Foundations made of timber
are seldom used now except for temporary structures.
Not many years ago, it was a common practice in Cali-
fornia to use timber footings for dwelling-houses. For
this purpose, redwood or cedar was employed, since

TIMBER FRAMING 211

both these varieties of timber resist decay to a consider-
able extent, even when embedded in the earth. Two
kinds of redwood are found in California, the Coast
redwood, or Sequoia sempervirens, and the Sierra red-
wood, or Sequoia gigantea, the latter being used prin-
cipally in the San Joaquin valley. To my knowledge,
the former is generally considered the better timber of
the two for use in foundations, although I am by no
means sure that such opinion is based upon anything
but prejudice. Good sound cedar is practically as good
as redwood, although I prefer redwood myself. Here,
again, the preference may be based on prejudice, as I do
not know of any tests establishing the length of time
either redwood or cedar will resist decay, when buried
in the earth. Indeed, there are so many factors, such
as quality of timber, character of soil, amount of mois-
ture, etc., affecting the life of a timber in contact with
earth, that no single series of tests would establish a
definite result. Fence-posts made of redwood or cedar
have withstood the ravages of decay for many years. On
the other hand, I have seen some redwood posts decayed
after a few years' service.

Ordinary timber, such as Douglas fir, if in contact
with the soil, or alternately wet and dry, will rot in
a short time. It may be said that one to five years'
service is all that can be expected from such timber,
provided that it is untreated. For this reason, . such
timber is usually treated with somie wood-preservative
when placed in foundations. Some of these so-called
wood-preservatives are, however, almost useless. Fur-
thermore, even when using a good preservative, care
should be taken to see that the timber is thoroughly
dry, and that the preservative is well worked into the
fibres of the timber, otherwise it will not be effective.
Painting the timber lightly is a needless expense, since
such treatment is of little value in adding to the life of
wood exposed to underground conditions.

In the following discussion, I wish to consider briefly
typical details of timber footings. Fig. 88, a, 6, and c,

212

TIMBER FRAMING

show some types of timber footings that I have seen
used in buildings. It is hardly necessary to point out
the defects in these details. It is obvious that in Fig.
88a, the two planks, m, add no strength to the footing,
and serve only to tie the sticks of the lower planking
together. The plaak n must distribute the whole load of
the post to the lower layer of planks. With a soil-pres-
sure of any appreciable amount, this plank must deflect
to such an extent that the bearing of the soil is taken
almost entirely by the middle plank of the lower layer.

I-
r

-J
±

</77

^-n

-m

rai

a>)

—

~ L

-1

—

1 "
-1

F^_j — 1

— 1

(c)

Fig. 88. types of defective footings.

thus increasing the intensity of soil-pressure over that
computed (assuming that computations were made).

In Fig. 88&, the distributing-cap is so short that again
the middle plank of the lower layer is a<;ting as the
footing. The footing of Fig. 88c is somewhat stiffer, on
account of the three layers of planking. Further layers
of planking will, naturally, strengthen the footing, and
in this way a detail can be constructed sufficiently stiff
to distribute the load on the post uniformly over the
area of the foundation, but this kind of foundation is
neither efficient nor economical.

It may appear a waste of time and space to discuss
such a detail, yet, as stated previously, I have seen it

TIMBER FRAMING

213

used extensively. Indeed, in checking the designs of
the various foreign, State, and eoneession buildings, sub-
mitted to the Division of Works of the Panama-Pacific

Exposition, I found such details quite common. The

fact that these structures were not designed in any one
locality would seem to indicate that this type of footing
is used extendveiy for temporary structures.

The best spread timber-footing for a small foad is
illustrated in Fig. 89. This detail is efficient and eco-

214 TIMBER FRAMING

nomical, and subject to rational analysis. Its design
involves only the consideration of bearing pressure on
soil and timber, and bending and longitudinal shear in
timber. There is no tendency for the distributing-cap
to be split by the punching effect of the post.

Fig. 90 shows a similar detail for a larger footing.
The outer stringers are added to tie the bearing planks
together. This figure also illustrates a typical detail for
post and girder connection. The corbel shown is not
for the purpose of reducing the unit bearing-pressure
across the under side of the girders, but rather to allow
for the possibility of the girders being cut too short to
meet over the centre of the post. The 2 by 14-in. splice-
pads not only tie the girders together, and so add gen-
eral stiffness to the floor, but they also furnish a certain
amount of end-restraint or continuity in bending to the
girders, in case the actual centre lines of bearing of the
girders are unsymmetrical with regard to the centre
line of the post. The 2 by 6-in. braces may or may not
be necessary, depending upon the height of the floor
above the ground. Such braces are an effective means
of stiffening a floor against vibration from machinery.
Where such bracing is necessary, the post should be
braced in all directions, and it will usually be sufficient
to brace only every other floor bay. For the bracing in a
plane normal to the plane of the girders, the joists im-
mediately over the post may be spaced so as to allow the
batter-braces to be spiked to the joists.

In Fig. 91 there is outlined a typical timber-footing
for the case of a column extending through the floor.
In this detail the girders are supported by short posts
alongside and fastened to the main post. A modi-
fication of this detail is shown in Fig. 92, where the
short posts are eliminated, and the main post is cut to
receive the girders. Because of the expense of cutting
the post and the weakening of the post resulting there-
from, the detail of Fig. 91 is to be preferred. The detail
computations for the design of the timber footing of Fig.
92 is given, using the typical building of Fig. 93, al-

TIMBER FRAMING

f- Iff Stid Brid^nq

FlQ. 91. TIMBEB B-OOTINO FOB

though such a building, unless built for very temporary
purposes, would have concrete piers.

Pile Foundations. Where the conditions of soil are
such that piles are necessary, the details shown in Fig.
90 and 91 may be modified by resting the bottom of the
post on the top of the pile. lu such eases, however, a
bolster, either of timber or of iron, should be placed be-
tween the ends of the post and the pile, in order to pre-
vent moisture from attacking the post.

For piles of ordinary length, it will generally he found
economical to arrange the spacing of floor-hays so that
the full load-eapacity of the pile may be utilized. The
capacity should be determined either by test- piles or by

216

TIMBER FRAMING

comparison \s^ith piles used under similar soil-conditions,
supplemented by borings to determine the nature of the

"H/55^Sw5"

KltttA

i\a»ji^H

tr*i4'i

5»J

■^

F=l

^1 I I I I I I I I I^

I I I I I

■ III

I I

tP ^3^&

<»^/<' ^

Pig. 93. typical timbeb-fbamed building.

Computations
Assuming tar and gravel roof-covering weigMng 8 lb. per
cu. ft. and Douglas fir weighing 3.5 lb. per ft. B.M.

Dead load of roof per sq. f t ^ 17 lb.

Dead load of floors per sq. ft = 15 lb.

Live load on roof per sq. ft = 28 lb.

Live load on floors per sq. ft = 85 lb.

TIMBER FRAMING 217

Total dead load on footing:

Roof = 3,800 lb.

Floors = 10,100 lb.

Post and footing = 1,100 lb.

Total = 15,000 lb.

Total live load on footing, assuming that 60% reaches the
footing:

Roof = 3,760 lb.

Floors = 34,300 lb.

Total = 38,060 lb.

Total live load plus dead load = 53,060 lb.

Assuming allowable pressure on soil of 2 tons per sq. ft., area

required = 5?55£ = 13.25 sq. ft.

4000

Footing therefore will be 3 ft. 8 in. square.

Bearing area required under post = = 186 sq. in. Use

285

14 by 14-in. short post.
Sill will have overhang of iillM = 15 in.

Load on overhang = 4000 X 3.66 X 1.25 = 18,300 lb.
Bending moment = 18,300 X 7.5 = 137,500 Ib.-in.
Requires, for bending, an 8 by 10-in. timber laid flat.

Maximum shear == 18,300 lb.

18300 V 3
Area required for longitudinal stress = ^-^— = 183 sq.in.

2 X 150
Use a 14 by 14-in. sill.
Overhang of planking = 15 in. Load on overhang for 12-in.
width = 4000 X 1.25 = 5000 lb.

Bending moment It = 5000 X 7.5 = 37,500 Ib.-in. Requires a
4 by 12-in. plank for bending.

Maximum shear = 5000 lb.

Area required for longitudinal shear = — — = 50 sq. in.,

2 X 150

therefore a 4 by 12-in. plank is all right.
Bearing required for floor-beam.

Load = (14X16)(15-[0.8X85]) ^ 93^^ ^^ j^ ^^ ^^^

required = ^^^^ = 3.26 in.
285 X 10

Therefore taper bottom of 10 by 10-in. column as shown.

218 TIMBER PltAMINO

underlying soil. Borings should be made in any event
in order that full knowledge may be obtained of existing
conditions. This statement applies not only to investi-

FiQ. 94. 95.

'ANCHOSAOES.

gations to determine the capacity of piles and elevation
of ground-water, but also to the study of foundation
conditions in the case where spread footings are to be
used. Generally a pile of an^verage length of 40 ft.,
and a butt of from 12 to 14-in, diameter driven properly

TIMBER FRAMING 219

to refusal may be expected to carry twenty tons with-
out settlement.

Where conditions are such as to justify temporary
construction, the piles may be cut off just above the
ground-level, and posts used to carry the floor-girders,
or the point of cut-off of the piles may be raised so that
no posts are required. The question of relative cost will
be the main factor in determining which method of the
two will be used, and this must be computed for each

FiS. 97. TYPE OF COLUMN'

individual ease. Other things being equal, the first
method is to be preferred, as the top of the piles, if
they are cut off at any appreciable distance above the
ground, are likely to be a considerable distance from
their computed positions. This condition will disar-
range the floor-system. For a permanent structure
the timber piles must be cut off below the permanent
ground-water level in order to prevent decay.

In this connection, it should be noted that 'one-pile'
footings should only be used to support posts canning
no other floor than the first. For posts extending
through the first story of the building and supporting
some of the other floors, with the position of the posts
determined by wall or girder-ties, provisions must be

220 TIMBER FRAMING

made for the foundation piles being at least six inches
from their theoretical position. The use of a 'two-pile*
footing with a wide cap will be necessary, even where
the load coming upon the footing could be safely sup-
ported by one pile alone.

Anchorage for Columns. When uplift must be con-
sidered, which may occur in high, narrow structures,
piles are important in providing anchorage. If the con-
dition of the soil does not necessitate piles, the uplift
must be handled either by burying the timber footing in
the ground, or by using a large concrete footing to fur-
nish the required weight. The latter condition is quite
common. A timber-framed building is comparatively
light. With a high narrow building, or a high building
with only two or three posts in the direction of the width
of the building, it may be desirable, or even necessary,
to anchor the wall-columns against uplift, or else to
secure the ends of the columns rigidly to the founda-
tion. In this way the columns may be considered as
practically * fixed' in the computation of stresses, and
the computed stresses in the columns resulting from
wind reduced accordingly.

For anchoring columns securely to the foundations,
various details may be employed, ranging from the
simple expedient of using two or four thin straps, bolted
or lag-screwed to the post (see Pig. 94) , to the somewhat
elaborate detail shown in Fig. 95, 96, and 97, which is
made from plates, angle-bars, and anchor-rods.

The problem of anchoring a timber post to a concrete
foundation is much simpler than where the footing is
constructed of timber, either of the spread-foundation
type, or a grillage resting on piles. A timber founda-
tion is subject to shrinkage, and in such cases, where
anchor-rods are used, the detail of connection should be
arranged so that the nuts on the anchor-rods may^ be
tightened after the shrinkage has taken place. Anchor-
ages of timber columns to timber grillages are unsatis-
factory at the best, since it is practically impossible to

TIMBER FRAMING 221

keep the connections tight. With concrete piers the case
is different. Anchorages of the types shown in Fig. 94,
95, and 96, with the anchor-rods or plates embedded in
the concrete can be expected to remain tight irrespective
of conditions.

In Fig. 97, it will be noted that there is a bed-plate
underneath the column and stiff ener-angles. Attention
is called to the fact that if the bolts through the column
are loose, it will be impossible to draw these bolts tight
by screwing the nuts on the anchor-rods, as the pull of
the anchor-rods is not against the column, but directly
against the foundation timbers.

Perhaps the most satisfactory anchorage that can be
devised for a timber column, when the stresses in the
anchor ties are of considerable magnitude, and when,
especially, a rigid anchorage is desired, is the detail
shown in Fig. 98, which is an adaptation of the tenon-
bar splice. In designing such an anchorage, the width
of the bar is determined either by the required area for
bearing against the ends of the fibres of the timber in
the post, or by the minimum width necessary for the
size of anchor rod used. The height or thickness of the
bar is found from considerations of bending, the total
uplift in the post or pull in the anchor-rods being con-
sidered uniformly distributed along the bearing-length
of the anchor-bar. As such bar is a short beam in bend-
ing, it is allowable to use a high unit fibre-stress^ in
flexure. The usual unit b ending-stress of 16,000 lb. per
sq. in. may be increased to 24,000 pounds.

In order to. keep the bending-moment in the anchor-
bar at a minimum, it is advisable to use hexagonal nuts
on the anchor-rods. This will allow the rods to be placed
nearer the post than if the ordinary square nuts are
employed.

This detail can be relied upon at all times. There is
no initial slip in the anchorage, when the uplift comes
upon the post, since the nuts on the anchor-rods can be
drawn up tightly at the time of framing, and can al-
ways be maintained in this condition. The detail also

222

TIMBER FRAMING

'Z ^^Sftse/ Bar

■i /Anchor /9od$.
tfeAo^onat NutB
at^ Upper £nc^.

COtiCf9tTt Pi£R

Mfl
s

II
II

CoNcnere Ptcif,

Fig. 98. appboved form or foundation.

has the advantage over the others shown above, in that
there is only one bearing-surface of metal upon timber.

TIMBER FRAMING 223

CHAPTER XIV
Miscellaneous Structures

In the preceding chapters I have endeavored to pre-
sent the underlying principles of structural mechanics
as applied to timber framing, to show that timber con-
struction is worthy of the same study as construction
in steel or concrete, and to point out details of fram-
ing that best fulfil the requirements of the particular
structure ^ under consideration. This presentation has
been in somewhat logical order, based upon the various
structural features of a timber-framed building. Thus
the text has considered in sequence: grading-rules,
working-stresses, washers and pins, strength of nailed,
screwed, and bolted timber- joints, the design of typical
truss- joints, design of truss-members, bracing, columns,
joist and girder connections, and foundations. While
the discussion has been limited almost wholly to build-
ing design, it does not follow that the principles, meth-
ods of design, and details are applicable only to timber-
framed buildings. The building has been chosen
rather as an example, for the reason that the design and
construction of a large timber building of the mill-
building type includes practically all of the problems
that will arise in timber-framing.

When other structures, such as bridges, trestles,
.towers, flumes, etc., are considered from a structural
standpoint, the loads and their applications, and also
the allowable working-stresses, may vary, but the same
principles of design will hold. Many of the details
of connections that are used in building construction
may be employed in these other types of structures. As
was stated in the opening chapter, timber railroad-
structures, such as a combination timber and steel
bridge, or a timber trestle, are well standardized. The

224

TIMBER FRAMING

same is true of the larger highway bridges. To a con-
siderably less degree, Hume-desigo follows certain
standards.

In the present article, there is presented a somewhat
superficial treatmeut of a few timber stmetures that
may be classed as ' miacellaneoua structures.' In the
order of their diseus^on, these are (1) flumes, (2)
head-frames, and (3) water-towers.

Flume-Design. To illustrate a typical problem in
dume-design a timber fiume will be assumed 9 ft. wide
on the inside, and with a maximum depth of water of
5 ft. 3 in., as shown in Fig. 99. Further, it will be

assumed that the timber is not Douglas fir, but Cali-
fomian mountain pine. The following unit working-
stresses will govern the design :

Tension and bending, 800 lb. per sq. in.

Bearing across fibres, 200 lb. per sq. in.

Bearing against the ends of the fibres, 700 lb. per sq.

Longitudinal shear, 100 lb. per sq. in.
The figure shows the typical section as framed, and
the sizes of the different members. The details of eon-

TIMBER FRAMING 225

nections are found as follows. The cross-ties will be
placed 3 in. above the flow-line and the bents will be
spaced 2 ft. 8 in. centre to centre.

Detail Calculations

(1) Vertical post.

Total pressure per lin. ft. = P = i X 62.5 X (5.25)^ = 860 lb.
This pressure acts at § of the depth of the water.
The reactions at A and B are then

Pressure at A = P, = ^4^ X 860 lb. X 2.67 = 735 lb.

q 75

Pressure at B = P^ = ttk X 860 lb. X 2.67 = 1565 lb.

O.oU

In determining the moment on the post, such moment ma^
be found by considering the total pressure P as uni-
formly distributed along the length of post. This method
is not exact, but is approximately correct.

M = lX 2300 lb. X 5.50 ft. = 1580 Ib.-ft.

The resisting moment of a 6 by 6-in. timber is 2400 Ib.-ft.,
while the maximum resisting moment of a 4 by 6-in. tim-
ber is 1600 Ib.-ft., both computed at 800 lb. per sq. in.,
maximum unit fibre-stress in bending.

The post will be gained into the cross-tie at the top and into

the sill at the bottom. The required bearing-areas at the

top and bottom, are then

735
Area at A = -rzr^ = 3.68 sq. in.

A X •« 1565 _ „_ ,

Area at B = -~-r- =: 7.87 sq. in.

If the posts, cross-ties, and sills are 6 in. wide, the gain at
the top must be f in., and the gain in the sill If inches.

(2) Stringers.

The trestle bents will be assumed to be spaced 8 ft. centre

to centre.
The load on one stringer will then be
Water, 3.25 X 8 X 5.25 X 62.5 = 8530 lb.
Plume, assume weight, 400

8930 lb., say 9000 lb.

Moment on stringer = J X 9000 X 8 = 9000 Ib.-ft.
Resisting moment of 6 by 10 in. timber = 6650 Ib.-ft.
Resisting moment of 6 by 12 in. timber = 9600 Ib.-ft.

The above calculations indicate that the vertical posts,
cross-ties, and sills could be 4 by 6 in., the posts to be
gained 1 in. into the cross-ties, and 2 in. into the sills.
The cross-ties, as far as requirements of strength are

226

TIMBER FRAMING

concerned, could be made 4 by 4 in. This design, with
the detail computations, has been taken from an actual
case, a flume intended for a hydro-electric development.
The sizes actually used were : cross-ties, 6 by 6 in. ; posts,
6 by 8 in. ; sills, 6 by 8 in. ; and stringers, 6 by 12 in.
See Fig. 99, which also shows the* substructure. The
engineers in this case used their calculations more as a
general guide as to minimum requirements than to de-
termine the actual sizes, and judgment and experience
influenced the selection of practically every section. In
this case, also, the decay of the timber through alternate
wetting and drying was considered in employing sec-
tions larger than required by computations for actual
strength. Standard practice has determined certain
minimum sizes of timber for use in such cases : for larger
flumes, requiring larger sections from a purely theo-
retical standpoint, the margin of safety to provide
against decay could be reduced and the actual sections
used would be nearer the sections determined from con-
siderations of strength alone.

As an illustration of a larger flume, Fig. 100 shows one

gi''ef-/Oj I — ,

3 ^A»3

~^= — L. /- -^'MS'-Z^O'/f)^

J-. i6?' ^a/fA: ' i''^ Borfmm

^JifM 'sae*. yyy.^ : *tf«i(. x</yy^ invvxt ^yyA-y .\v>nv -^^^f^ >xwv j«««^ '«cwp i^

Fig. 100. cross-section of flume at exposition.

flume used in the filtration-plant of the Panama-Pacific
International Exposition. This flume is 7 ft. 9 in. high,
and 12 ft. 6 in. wide. The side-posts are two 3 by 8 in.
timbers, bolted and spiked to a 4 by 6 in. sill and a 4

TIMBER FRAMING 227

by 6 in. cap, or yoke. Between the bottom and top yoke
is a 3 by 8 in. filler, which stiffens the side-posts. In
addition to the bolting and spiking of the posts to the
top and bottom yokes, short wooden blocks were bolted
and spiked to each yoke to receive the thrust of the posts.
The side-posts were spaced 4 ft. centre to centre. Every
third post was braced to the sills by two 2 by 6 in. braces.
The flume was carried by five 6 by 12 in. stringers, hav-
ing a span of 12 ft, the stringers being supported by 8 by
12 in. sills, resting upon 3 by 12 in. plank. All material
was Douglas fir. As the plant was of a temporary nature,
the sections were determined from considerations of
strength and deflection alone. . In flumes of this size, de-
flection of the stringers, and of the side-posts must be
taken into consideration, as leakage may result if there
is appreciable deflection in the side-posts or the sills
under load.

Two types of joints for the flume-lining were used;
one, using i by 4 in. battens with asphaltum in the joints,
and the second an untreated spline-joint. The superin-
tendent of the plant favored the results obtained from
the spline-joint. I inspected the plant carefully several
times, and as far as I was able to determine by observa-
tion, neither joint showed any material advantage over
the other. Both were almost free from leakage.

In Engineering News, Vol. 76, No. 23, December 21,
1916, there occurs an article entitled * Rectangular
Wooden Flumes' by J. C. Stevens. Mr. Stevens treats
of many practical details of timber-flume construction,
but makes what I consider a radical statement to the
effect that it is a waste of lumber to place battens on the
inside of a flume. His recommendation is to place the
battens on the outside of the sides, and on the under-
side of the floor, cutting them between the posts and
sills. This is not the common practice on the Pacific
Coast, where there are many large hydraulic enterprises
using timber flumes. In triangular logging flumes,* the

♦See Bulletin No. 87, U. S. Department of Agriculture,
'Flumes and Fluming,' by Eugene S. Bruce.

228 TIMBER FRAMING

battens are usually placed on the outside of the flume,
and are often made continuous, the side-supports being
cut accordingly. However, the two cases are not at all
comparable.

Head-Frame for a Mine. Without going into an ex-
tended discussion of the design of head-frames, it is de-
sired to give one example of the application to a timber
head-frame of some of the principles and details advo-
cated in preceding articles. For this purpose, there is
shown in Fig. 101, a typical head-frame, taken from the
Engineering and Mining Journal of October 11, 1913.
The cut is from an article describing different types of
timber head-frames. Sufficient data are not given by
either the article or illustration to enable any of the
stresses to be computed, and consequently no stress-
analysis has been made. Indeed considerations of stiff-
ness of the frame under working-conditions will require
more bracing than might be computed from conditions
of actual load and wind-forces. In a structure of this
nature it is highly desirable not only that it shall be
rigid under the racking received in the hoisting and
dumping of ore, but also that the details are such that
the joints may be kept tight.

Referring to the figure, the first criticism to be made
is of the use of diagonal steel rods and horizontal struts
of timber. Not only is this the most expensive system of
framing under the usual conditions, where steel is ex-
pensive, and timber is comparatively cheap, but it re-
sults in bad details at the joints, making the whole struc-
ture difficult to tighten. With the rods horizontal, and
the struts in an inclined position, the amount of steel
would be the minimum, and the timber, while greater in
quantity, would not add appreciably to the cost of the
structure, and would be more than offset by the saving
in other features. Consider one of the panel-points of
the frame, where two rods intersect with the vertical
post. Not only is the post cut to receive the horizontal
strut, but in addition, two inclined holes of considerable
size must be bored through the post at the same point.

TIMBER FRAMING 229

greatly weakening the post. It is safe to say, without
knowing the size of all the rods, that 25% of the post is
cut away. While this proportion may not seem to be
large in itself, in this case it represents an area of some
60 sq. in. of timber that is useless. Expressed in other
terms, an area equivalent to a 6 by 10-in. stick is not
available. With such a detail, large sticks are a neces-
sity.

Near the top of the head-frame, the inclination of the
rods is considerable. It is evident that to pull the posts
snugly against the horizontal struts, there must be ex-
erted on the nut of the rod a force approximately twice
that which is actually pulling the rear and front posts
together. In other words, a large proportion of the stress
in the rod, when tightening is in progress, is exerted in a
vain attempt to lift the whole frame from its foundations.
On the other hand, if the rods were horizontal, every
pound of stress placed upon them when^ tightening the
nuts would be exerted in pulling the posts tightly against
the inclined struts. This statement is, needless to say,
based on the assumption that the washers are large
enough to prevent their being crushed into the timber of
the posts.

A similar criticism may be made of the system of
bracing below the loading-floor. Here "would seem to be
an excellent opportunity to use a truss of the Howe type,
which would serve to carry the weight of the loaded floor
to the supports, and could also be utilized to brace the
whole frame against lateral forces. For this purpose, the
bays of the truss should be counter-braced, and special
attention should be paid to the end-connections of the
truss, to see that they are capable of transferring both
tension and compression to the posts.

With the design shown in Fig. 101, it is difficult to see
how tension could ever exist in any but the main posts of
the head-frame. It will be seen that the two intermediate
posts are anchored to the concrete foundations. Ap-
parently, however, there is no definite connection be-

230 TIMBER FRAMING

tween these intermediate or minor posts and the main
frame.

The final point of criticism in the design is the type of

post-ancborage. Since the details are shown to a small
scale, it is difficult to be sure just what such details are.
It would appear that the anchorage consists of steel
straps or anehor-rods flattened at one end to straps,

TIMBER FRAMING

232 TIMBER FRAMING

buried in the concrete, and bolted through the posts. If
the actual tension that may come into the posts is small,
this detail may be entirely satisfactory ; in fact, it is not
my intention to imply that such a type of detail may not
be strong enough in any case. There can be no chance of
failure, provided that length of anchorage, size of strap-
bolts, and number and size of bolts are sufficient for the
stresses they will be called upon to carry. But, as wag
pointed out in Chapter XIII, such a detail, once in
place can never be tightened. It would seem therefore,
that a detail which could be adjusted at any time by
tightening the nuts of the anchor-rods would be ad-
visable, particularly as such a type of detail would not
necessarily be any more expensive.

In order to emphasize the criticisms of Fig. 101 that
have been made, there is shown in Pig. 102 a revised ele-
vation of the head-frame, with some of the most important
connections detailed. The two main posts have been re-
duced in size, since with the type of connections used, a
much larger proportion of their section is available for
use. No sizes have been noted on any of the rods or
timbers; without knowing the loads there has been no
attempt to compute stresses. The tie-rods have been
placed in a horizontal position, and the diagonal mem-
bers made compression timbers. Where two such di-
agonal struts intersect on the posts at a common point,
butt-blocks have been used. For the support for the
loading-floor a truss has been introduced, bolted at its
ends to the main posts by means of splice-pads. The first
vertical post to the left of the main inclined post of the
frame, has been moved in to the left from its former posi-
tion. This is to enable a more positive tie to be made at
its upper end to the cross-truss, and also to allow another
system of cross-bracing to be introduced. Clearance for
cars would determine the position of this post.

Suitable washers of generous area should be provided
for all bolts and rods, so that crushing of the fibres of
the timber would not occur. This provision will add
much to the life of the structure. It would also be ad-

TIMBER FRAMING

O'Dioirt • /■*^-//,f/i.

I OF TAHK-TOWEB.

234

TIMBER FRAMING

visable to treat thoroughly the contact faces of all tim-
bers, and the contact faces of all metal and timber with
a good wood-preservative. The life of the head-frame
will be lengthened by painting the inside of all bolt and
rod-holes in the timber, the ends of struts, and the cut-
faces of the timbers into which the ends of the struts
and the butt-blocks fit. It might even be advisable to
use castings at all panel-points, similar to the construc-

FlG. 104. FLOOR-PLAN FOB TANK.

tion used in standard timber railroad-bridges, even
though the stresses might not, and probably will not,
require the use of such castings. Metal base-plates
under all posts are a necessity to prevent moisture from
creeping up the timber.

Water-Tower. Another type of miscellaneous struc-
tures will be discussed in this chapter, namely, a water-
tower. An example from actual practice will again be
chosen, and an alternative design shown. In this case
the first design (see Fig. 103), was submitted by a firm
manufax3turing and selling wooden tanks and pipes.
Fig. 105 shows the tower as re-designed and built.

The principal points of difference between the design
of Fig. 103 and 105 are (1) the omission of the column
splices, in Fig. 105 (2) the change in details of the inter-
sections of the diagonal struts with the posts, and (3)
the revision of the post-anchorages.

Taking up these three points in the order mentioned.

TIMBER FRAMING 235

not only is a saving in cost made in the revised design
by the omission of the column-splices, but a much
stronger tower is secured. Attention is called to the
splice proposed by the tank-manufacturers. The type of
splice may be classed as the oblique scarfed joint with
fish-plates. In a previous chapter objection has been
made to this type of splice for timbers carrying heavy
compression, because of the two surfaces of contact, re-
quiring an accurate fit to be made. The fish-plates as
required in the design were two 1^ by 10 in. by 6 ft. oak
plates, fastened with eight J by 14J-in. bolts. Why oak
was specified is not clear. The strength of oak in end-
bearing and in cross-bending is practically the same as
Douglas fir, of which latter timber the posts are com-
posed. The shearing strength of oak is not quite 25%
greater than that of Douglas fir. The fear of splitting
the splices if they were made of Douglas fir may have
influenced the designer to substitute oak. If, however,
splices were necessary, which in this case they were not,
since the length of the posts is not abnormal, and long
timbers were easy to obtain, a better detail would have
been secured by using a straight normal cut for the post,
and thicker splice-plates of Douglas fir.

The detail of the intersection of the inclined struts
with the posts is, perhaps, the worst feature of the de-
sign of Fig. 103. Note that the horizontal 3 by 10-in.
girts are set into the posts approximately one-fourth of
the depth of the posts. A note appears on the draw-
ings, * Bracing for central post-girts only.' In this post
over 50% of the cross-sectional area of the central post
is in cross-bearing on the timber. A rough calculation
indicates that the unit compression on the central post
is 320 lb. per sq. in. for dead load with no wind. While
this unit compression is not excessive, it is evidently
more than the designer anticipated, since 3 by 10-in. oak
plates are specified between the tops of the posts and the
bottom of the stringers. The inclined braces bear di-
rectly*across the fibres of the 3 by 10-in. girts. This is
a poor detail, both on account of unit bearing-pressures.

236 TIMBER FRAMING

and of timber-shrinkage. The girts would shrink and
leave a somewhat loose fit between post, girt, and braces.

The third point is the anchorage for the posts to the
foundation. Each of the four comer-posts is tied into
the concrete foundations by one f by 2i-in. tie-strap,
fastened with two bolts, presumably } in. diam. The
strength of such anchorage, measured by the safe re-
sistance of the bolts, is not over 3000 lb., measured by the
tie-strap in tension, 10,000 lb., and by the weight of the
concrete, 5472 pounds.

Turning to the revised design, as shown in Fig. 105,
it will be seen that the post-splices have been omitted and
the connection of inclined struts modified by introducing
butt-blocks set into the posts ; the latter change brings all
post-timbers in end-bearing, and frees the joints from the
effect of shrinkage of the timber, since the shrinkage of
Douglas fir parallel to the fibres is small, and negligible
for ordinary lengths of timber. The horizontal tie-rods
now extend through the three outside posts. It is there-
fore necessary that horizontal struts be used in addition
to the rods, in order that the increment of wind-shear may
be transferred across the posts, from one system of brac-
ing to the other. A more effective manner of handling
this problem, although more expensive, would have been
to have placed a tumbuckle on either side of the mid-
dle exterior posts. It is evident, that, with wind on the
tower, the stress in any horizontal tie-rod will not be
the same on both sides of the post in the revised de-
sign, and the difference of shear must be carried from one
comer-post to the opposite one and back to the mid-
dle post by compression in the two 2 by 8-in. girts. The
central post is tied to the exterior posts by two 2 by
8-in. girts, in two directions, the girts being bolted to
the posts. These girts are therefore able to develop
both compression and tension. Finally, the size of the
column-footings has been increased, steel base-plates in-
troduced, and the anchorages strengthened, each post
having two anchor-straps lag-screwed into the posts.

In the re-design of the tower, a wind-load of 30 lb.

TIMBER FRAMING

287

•

Structure adeim fhf» /ine
not chcnged

9.9'^arl /^^O^^

''?»*i,^^^

^pped r/ato posti

e-z-^6'

11 \ rootf/7^9,€'0'*e<r.

D£r/ifL Ce/vr/ML Post.

Fig. 105. bevised design of tank-toweb.

238

TIMBER FRAMING

per sq. ft. of exposed surface was used. It should be
stated that the tower was to be completely enclosed, and
that it stood on the top of a hill, near the ocean shore,
and exposed to the full force of the wind on the entire
surface of structure enclosing the tower. The rather
heavy anchorages are therefore justifiable.

The following data regarding the design of head-
frames and ore-bins, contributed by Robert S. Lewis,
professor of mining at the University of Utah, will be of
value to those interested in mining structures.

Head-Frames. The small head-frames used in pros-
pecting and development work are seldom designed by
an engineer. Their construction and planning is gen-
erally left to a carpenter or contraxjtor, and the excel-
lece of the design depends upon the previous experience
of the carpenter or contractor. In a mining district
there is often a striking similarity in design of the dif-
ferent head-frames, either because of a common builder,
or because the design of the first frame to be erected was
copied by the builders of the other frames. The average
life of a wooden head-frame may be taken at 10 years,
hence, if the mine is likely to have a longer life, it is
desirable that the frame be made of steel.

There are two general types of head- frames: the A-
frame type and the four-post type, both shown in Fig.

"A" Frame . four Post Frame .

Pig. 106. types of head-frames.

106. The former is a simple type of frame. All stresses
are determinate. The objection that the sheave is
mounted by bolting the bearings to the front frame, thus
bringing a pull on the bolts, can be met by special meth-

TIMBER FRAMING

239

ods of mounting the sheave. However, to prevent the
skip or cage from striking the front bracing, the front
posts must be set back from the shaft. Consequently the
sheave must be large enough to bring the hoisting-rope
to the centre of the shaft. This may require the use of a
very large sheave, entailing great inertia and wear on
the rope on account of slippage when starting and stop-
ping. This disadvantage is not serious unless a heavy
load is handled at high speed.

AC-2AB
AB- Safe working had
for rope, including
tending stresses.

Not over
6 degrees.

Drum.

45 degrees or over.

Pull.

Position of BacksfvyJ
Influence Diagram.

Fig. 107. hoisting-diagbam and influence diagram.

In the four-post type the sheaves are mounted on hori-
zontal timbers placed on top of the structure. This
frame lends itself to rigid construction in either wood or
steel. The joint at B is indeterminate, consequently the
stresses cannot be computed exactly. The frame strad-
dles the shaft, and there is no particular difficulty en-
countered in mounting the sheave, its diameter being
computed after the size of the rope and the allowable
bending stresses have been determined. However, the
diameter of the sheave governs the width of the frame, as
the horizontal timbers at the back of the frame must clear
the sheave. When the loads to be handled are large, an
extra post may be included between the other two, thus
making a six-post frame. When a three or four compart-
ment shaft requires a head-frame, a bent may be in-
cluded between each compartment.

The height of the head-frame should be the sum of the
following three quantities : the height of the landing-floor
above the collar of shaft, which is determined by local
conditions, the height of the skip or cage, and an allow-

240 TIMBER FRAMING

ance for overwinding. This allowance should be i to f
of a revolution of the drum for direct-acting hoists.

The next step is to determine the position of the hoist-
ing-drum. This should be placed so that the fleet angle
of the rope, leading from a position on the extreme edge
of the drum, is not greater than 6° to prevent a tendency
of the rope to climb the grooves in the drum. Then the
angle between the horizontal and the rope should not be
less than 45° and should not be more than 55°. This is
to avoid 'lashing' of the rope, and the necessity for em-
ployment of extra sheaves to support the rope.

The distance between panel-points is assumed, as there
is no definite practice in this regard. It is generally be-
tween 12 and 20 feet.

When the position of the hoisting-drum is decided, the
next point is to find the position of the back-stay of the
frame. The following methods have been used :

1. Placing it parallel with the resultant of pull and
loads.

2. Placing it parallel with the hoisting-ropes.

3. Placing it just outside the resultant of pull and load.

4. Empirical position, making the pull twice that of
the load.

5. Placing it at 30° with the vertical.

Method No. 4 is satisfactory, and gives a safe position.
It is not far from the position given by method No. 5.

The front-width of the frame at the bottom is calcu-
lated to prevent overturning by wind-pressure. The
wind is assumed to blow around the first timbers and
strike the rear-posts also, unless the entire structure is
housed, when the total exposed area can be computed.
The wind-pressure should be taken at 30 lb. per sq. ft.
The front top-width must be such as to carry shaft-
guides, and to give sufficient clearance between the cage
or skip and the frame. The guides are from 4 by 6 up
to 8 by 10 inches.

A method of computing the stresses is given in
Ketchum's 'Design of Mine Structures.' For the four-
post frame, a simpler approximate method is as follows :

TIMBER FRAMING 241

In Fig. 106, consider that the structure BCD is rigid
and takes most of the wind and live loads, and that
ABDE serves merely to support the sheaves and transmit
these loads to BCD, For the side-elevation of BCD, the
resultant of the rope-pulls is resolved into horizontal and
vertical components. AH the horizontal forces are as-
sumed to be applied at B and act on BCD, The reactions
of the vertical components at A and B are computed and
the structure BCD is assumed to carry those at B, while
the front-posts carry the reactions at A. It may be
assumed that the wind-loads at A, F, G, and E are car-
ried by BDC, or they may be considered as producing
stresses in the cross-bracing in ABDE. This method will
give a structure which, if anything, errs on the side of
safety. The cross-bracing in ABDE would be propor-
tional to the size of the posts. Since part of the area of
some of the sections is cut away for joints, the size of
these members should be increased to provide the desired
strength.

Ore-Bins. Ore-bins serve, in general, to regulate the
movement of ore between the mine and railroad, or be-
tween mine and mill, so that a temporary cessation at
one end does not stop operations at the other end of the
system. In most cases, ore-bins are designed to hold
from two to three days' supply of ore, but local condi-
tions may call for modification of these figures. In a
country where Sunday transportation of ore is forbidden
by law, the capacity of an ore-bin must be such that it
affords ample storage-space from Saturday to Monday,
which means that the bin should hold practically a three
days' supply of ore.

Timber, steel, and reinforced concrete are used for
making ore-bins. The largest are usually built of steel.
Medium and small-sized bins are generally built of tim-
ber, because of its availability, ease of working, and
cheapness. At some mines, timber bins of large capacity
have been built for the reason that the mine was begin-
ning production, and it was impossible to obtain steel
within a reasonable time.

242

TIMBER FRAMING

Pig. 108-111 show side-elevations of the most common
forms of bins. The triangular shape, shown by the dot-
ted lines in Fig. 108 is used sometimes, but is not eco-
nomical, on account of the large amount of timber re-

Fig. 108, 109, 110, and 111. types of ore-bins.

quired. Flat-bottom bins, such as are shown in Fig. 109,
may have discharge-gates on one side, both sides, or in
the bottom. The bin shown in Fig. 110 is self -emptying.
Fig. Ill shows a bin that requires only one railroad
track because of its central discharge-gates.

Flat-bottom bins cannot be emptied without shoveling.
However, this shape gives the maximum capacity for a
given floor-space and height, the bottom is protected
from wear, and the cost of the bin-bottom is from one-
third to one-half that for an inclined bottom. A flat-
bottom bin is also cheaper for a given storage-capacity.
In order to be self -emptying, the bottom of a bin must
slope at an angle of 45° or more. The bottom should be
protected from wear by steel plates.

In figuring the capacity, the weight of ore is usually
assumed at 100 lb. per cu. ft. Where the ore is dumped
into the bin from several fixed points, the ore will stand
in a series of cones and the full capacity of the bin can-

TIMBER FRAMING

243

not be obtained. An allowance for this cooditioD should
be made in computing the capacity of the bin.

In ease the bin is to be part of a mill, its length should
conform to the floor-plan of the mill. Since increasing
the height of a bin increases its cost considerably, a long
narrow bin ia the most economical. The bin may be di-
vided by partitions into pockets or compartments. The

Fia. 112. ISFtUESfE DIAGBA

number and aze of these pockets depend mainly on the
number of different classes of ore that must be kept sep-
arate. In a long bin, partitions add rigidity to the struc-
ture. Pockets are sometimes made 20 ft. long. As a
rule, 20 ft. is also taken as the limiting width and height
for a wooden bin. If it is made larger, the timbers must
be of such a size that steel construction would be more
economical.

When a bin has to support additional loads, such as
crushers, or an engine and train of ore, the timbers must
be designed for the extra load. The sudden stopping of a
train results in severe stresses.

Foundations for bins must be high enough to permit
the passage of wagons or cars under the loading-chutes.
If the chutes are arranged to discharge at points one-

244 TIMBER FRAMING

third of the width of the car from the side next to the
bin and at quarter-points along the length of the car, no
shoveling will be needed to load the car to capacity.

For small or medium-sized bins, the vertical posts
along the sides and ends of the bins are unsupported
throughout their length. For large bins the load on
these beams becomes so great that they must be rein-
forced by the addition of horizontal ties. A single beam
may be used, in which case an I-beam or two channels,
placed back to back, may be more economical than a large
wooden beam. Tie-rods, from two to five vertical postd
apart, run from the front to the back of the bin. To
protect the tie-rods from falling ore, a beam with its up-
per part beveled and shod with a steel plate is some-
times placed just over the rod. Since the horizontal
beams and the tie-rods are generally designed to carry
all the loads, the vertical posts need be only large enough
to carry their load from one horizontal beam to the other.
With such construction, the ends of the bin may be
a source of weakness. Longitudinal tie-rods in a bin are
not desirable, so there is an unsupported span of the full
width of the bin at the ends of the horizontal beams.
The beams must be designed for this condition.

The stresses may be computed as shown in Fig. 112.
The weight of ore being known, the weight of ACD for
1 ft. in width can be computed. This weight, W, acts
as the centre of gravity of the triangle, 0. The total

pressure against CD, or P = ^ wh^ i h- sin e ^'^^^^

= angles of repose of the ore.

W = weight of ore per cubic foot.

h = total height, or CD.

The load P is assumed to be applied at a point i h
above the bottom. R is the resultant of W and P. Its
normal component, R, is the total normal pressure against
the bottom AD. To construct the graphic diagram, R
(in pounds) = i AD (in feet) times OD (in pounds),
from which GD can be found and laid off on a suitable
scale. Then the area FEDG = total normal load on bot-
tom-beam ED. One-half of the horizontal component

TIMBER FRAMING 245

of the normal pressure against the bottom is assumed to
be applied at E and causes bending in the rear-post.
One-half of the component of the bottom-pressure parallel
to ED is assumed to be applied at E to cause compres-
sion in beam ED. The spacing of the bottom-beams is
now assumed and their size calculated. Supports may be
used along the bottom-beams to keep them within a
reasonable size. The front-pressure may be computed
by either of the methods discussed above. Since a flat-
bottom bin has only the weight of the ore to be carried
by the bottom, the pressure against the sides is calculated
by the formula for finding P. The total pressures found
for the side and inclined bottom of a bin may be assumed
to act uniformly over the length of these beams. This
simplifies the design and introduces no serious error.
The walls of the bin are assumed to be smooth, so that the
angle of friction is taken as zero. ^

246 TIMBER FRAMING

CHAPTER XV

Wind-Pressure and Wind-Stresses
Working Drawings

The subject of wind-pressure and wind-stresses is an
unsatisfactory one to discuss. There is a wide varia-
tion in opinion as to the wind-pressure that should be
adopted for different types of structures, and for differ-
ent heights of the same structure. Again, the authori-
ties differ as to the unit-stresses that should be allowed
in designing for wind. Finally, several methods are in
use for finding the stresses resulting from wind. The
latter statement applies more particularly to the steel-
framed office-building than to the mill-building type of
structure.

In order to bring out these points more clearly, I
quote from Milo S. Ketchum's * Structural Engineers'
Handbook.' This authority specifies in regard to mill-
buildings, as follows :

**Wind Loads. The normal wind-pressures on trusses
shall be computed by Duchemin's formula, with P =
30 lb. per sq. ft., except for buildings in exposed loca-
tions, where P = 40 lb. per sq. ft. shall be used.

*'The sides and ends of buildings shall be computed
for a normal wind-load of 20 lb. per sq. ft. of exposed
surface for buildings 30 ft. and less to the eaves ; 30 lb.
per sq. ft. of exposed surface for buildings 60 ft. to the
eaves, and in proportion for intermediate heights."

Also, after defining the unit working-stresses for dead
and live loads,

'*When combined direct and flexural stress due to
wind is considered, 50% may be added to the allowable
tensile and compressive stresses.''

In the case of steel highway-bridges, Mr. Ketchum
specifies as follows:

TIMBER FRAMING 247

Wind Loads. The top lateral bracing in deck-bridges
and the bottom lateral bracing in through-bridges, shall
be designed to resist a lateral wind-load of 300 lb. for
each foot of span ; 150 lb. of this to be treated as a mov
ing load.

*'The bottom lateral bracing in deck-bridges, and the
top lateral bracing in through-bridges, shall be designed
to resist a lateral wind-force of 150 lb. for each foot of
span. In bridges with sway-bracing, one-half of the
wind-load may be assumed to pass to the lower chord
through the sway-bracing. For spans exceeding 300 ft.,
add in each of the above cases 10 lb. additional for each
additional 30 feet.

**In trestle-towers, the bracing and columns shall be
designed to resist the following lateral forces, in addi-
tion to the stresses due to dead and live loads: The
trusses loaded or unloaded, the lateral pressures specified
above ; and a lateral pressure of 100 lb. for each vertical
linear foot of trestle-bent.''

For direct wind-stresses, not combined with flexural
wind-stresses, the above specifications allow an increase
of 25% in the unit working-stresses; when direct and
flexural wind-stresses are combined with dead and live
load stresses, the unit working-stresses may be increased
50%. These specifications, while for steel structures,
should also apply to timber structures, except possibly as
regards the increase in unit working-stresses.

In the case of buildings of the mill type, a number of
experiments have been made on small models, some of
which would indicate that the ordinary assumptions as
to the action of wind on buildings of this type do not
hold. Albert Smith of Purdue University has found
that in some instances there is tension in certain truss-
members which by the commonly accepted method of de-
sign would take compression, and vice versa. In other
words, he finds a suction on certain portions of the roof
in such a building, instead of a pressure, or instead of
neither suction nor pressure, as would be shown by the
ordinary analysis.

248 TIMBER FRAMING

Wind-pressure on a building produces bending in the
columns, just how much bending is a disputed question.*
There is no doubt that the specifications of Mr. Ketchum,
if followed consistently, will result in a building of safe
design. The question to be decided by the engineer is
whether or not, such specifications, when applied to
timber buildings, are too severe.

R. Fleming, of the American Bridge Company, has
made a study of all available discussions in technical
literature on this subject, and has published several
articles on the subject in the Engineering News. Re-
cently, in an endeavor to standardize the various con-
flicting specifications, he has proposed a set of Specifica-
tions for Structural Steel Work.f

On the subject of wind-pressure, Mr. Fleming pro-
poses the following specifications:

** Wind-Pressure. AH steel buildings shall be designed
to carry wind-pressure to the ground by steel-framework.

**Buildings of Class No. 1 (mill buildings) not over
25 ft. to the eave-line shall be designed to resist a hori-
zontal wind-pressure of 15 lb. per sq. ft. on the sides of
the building, and the corresponding normal component
on the roof according to the Duchemin formula for
wind-pressure on inclined surfaces.

** Where buildings are more than 25 ft. to the eave-
line, the horizontal pressure shall be taken at 20 lb. per
sq. ft., and the corresponding normal component on the
roof.

* * Only the excess of the wind-stresses obtained by this
paragraph over the wind-stresses according to Clause 13
(stresses due to dead and live load) need to be con-
sidered. In arriving at this excess the wind included in
the total uniform loads designated in Clause 13 shall be
assumed at 10 lb. per square foot.

♦The difficult point to determine is the exact distribution of
the reactions at the foot of the columns resulting from wind,
and the amount of 'fixedness' existing in the column, the latter
being dependent on the presence or absence of sufficient
anchorage.

^Engineering Record, Vol. 74, No. 24, Dec. 9, 1916.

TIMBER FRAMING 249

*'For combined stresses due to wind and other loads,
the above mentioned stresses (working stresses for dead
and live loads) may be increased 50%, provided the sec-
tion thus obtained is not less than that required if wind
forces be neglected. ' '

It is my experience in checking over many designs,
and observing the sizes of members and connections of
buildings which have stood for several years, that the
greater number of buildings of the mill-building type
which are supposed to be designed for a wind-pressure
of 20 to 30 lb. per sq. ft. of exposed surface, would not
stand over half this pressure if consistently figured ac-
cording to the commonly accepted methods of design.
The greatest weakness is found in the knee-brace con-
nections to trusses and columns* and in the columns
themselves. In the case of timber- framed buildings,
this comment applies to the actual section of column as a
whole ; in the case of steel-framed buildings the incon-
sistence in design may lie in the relative strength of the
column section as a whole and the details. For example,
an analysis of a steel column built up of four angles
laced together, will often show that, while the moment
of inertia of the column-section as a whole is sufficient
to take the bending due to a 20 or 30 lb. wind, the lac-
ing-bars are far deficient to withstand the compression
due to wind shear, as they are usually constructed of
'flats,' with a large ratio of length to radius of gyration.

In the case of a timber-framed building, of moderate
dimensions, a rigid adherence to the standard specifica-
tions, such as Ketchum's quoted above, will often give
results that are out of reason, when compared to build-
ings that have long given service, and whose strength
no one would seriously question. This can best be
brought out by a typical example. Consider a timber-
framed building of the mill-construction type, with
trusses 16 ft. centre to centre, 30-ft. span, and with a
height to the eaves of 15 ft., height from floor to foot
of knee-brace 11 ft., distance from floor to bottom of
trusts 15 ft., and an over-all height of 24 ft. At 20 lb,

250 TIMBER FRAMING

per sq. ft., the total wind-pressure on one bay is 16 X 24
X 20 = 7680 lb. Assuming the wind-reactions to be
equally divided between the windward and leeward
columns (the usual assumption in design), the reaction
on one post is 3840 lb., and the moment at the foot of the
knee-brace is 3840 X H X 12 = 507,000 Ib.-in. Using
a maximum fibre-stress of 1800 lb. per sq. in., the re-
quired section-modulus of the column is j^g^^ = 282 in.

(It is assumed that the dead load and the direct wind-
load will not stress the post over 1000 lb. per sq. in. in
addition to the 1800 lb. due to the wind). This section
modulus corresponds to a 10 by 14-in. post. Yet a build-
ing of this size and type \^ith posts 10 by 14 in., 16 ft.
centre to centre carrying a corrugated-iron roof and
walls, would be considered a monstrosity, and rightly so.
It is true that a smaller post might be used, if the col-
umns are fixed at the base. Assuming that the columns
are rigidly fixed at their bases, the wind-pressure produc-
ing bending in the posts would then be 19^ X 16 X 20 =
6240 lb. The reaction producing bending would be 3120
lb., and the column-bending, 3120 X 5^ X 12 = 206,000
Ib.-in. The required section-modulus of the timber is

..gQQ = 114 in., which is furnished by an 8 by 10-in.

timber. Even this size of post is too heavy. In addition,
it will be found to be diflScult to design an anchorage that
will develop the required fixing-moment, and which can
be depended upon to remain tight under all conditions.
In the first instance taken, the building with columns

hinged at the base, the stress in the knee-brace is -j- X

3640 lb. X 1.41 = 19,000 lb. In the second case, columns

fixed at the base, the knee-brace stress is j- X 3120 lb.

X 1-41 = 10,400 lb. Both these stresses will require a
well designed connection of knee-brace to both column
and post. A bolt or two, and a few spikes will not suffice.
It must also be remembered that the connection must be

TIMBER FRAMING 251

designed so that it will be able to withstand both ten-
sion and compression.

For a building of the type and size just described, I
believe that a wind-pressure of 10 to 15 lb. per sq. ft. of
exposed surface is sufficient for the design of the trusses
and posts. The girts should be designed for a load of
15 to 20 lb. per sq. ft. of tributary area. The posts
should be tied into the foundations, since the dead load
coming on the posts is small. I have often found it
necessary to provide more concrete in the post-footings
than is required from considerations of unit soil-pres-
sure, in order to give rigidity to the building. In such
cases, -^ by 3-in., or f by 3-in. strap-iron, boltecf to the
posts, and anchored in the concrete footing will give a
certain * fixedness' to the posts. The necessity of such
anchorage can be determined easily: if the anchorage is
merely to prevent overturning of the building, the direct
wind-load in the column should not be allowed to exceed
about 80% of the computed dead load in the column. If
the weight of the concrete footing is utilized to give
fixedness to the column and thus reduce the wind-bend-
ing, it may be found that a considerable mass of concrete
is necessary.

In referring to overturning of the building, I have in
mind incipient overturning, or a lifting of the windward-
post off its base. It is almost inconceivable that the build-
ing could overturn as a whole. A *mind's-eye' picture
of the probable action of such a building under a terrific
wind will emphasize the enormous strain that would
come upon the knee-brace connections, and will bring
home the fact that such connections are the most im-
portant in the whole structure in resisting lateral forces.

In making the above recommendations for a reduced
wind-pressure to be figured on mill-construction build-
ings, I am considering localities not subject to cyclones or
tornadoes. It is my b^ief, that a carefully designed
timber-framed building, with connections intelligently
studied, will be perfectly safe ui\(ier all conditions that
may arise on the Pacific Coast, at least, provided that the

252 TIMBER FRAMING

girts are designed for 20 lb. wind-pressure per sq, ft.,
and the frame for 15 lb. Indeed, in some places, and for
some buildings, I would not hesitate to reduce the fore-
going pressures to 15 and 10 lb. respectively. A 15-lb.
wind-pressure would be produced by a gale of a velocity
of 60 miles per hour, which seldom occurs even for a few

minutes.

The corresponding unit-stresses to be employed in con-
nection with the wind-pressures advocated above should
not exceed 2000 lb. per sq. in. for combined dead and
wind load, including flexural and direct wind-load
stresses. Unless the building is of unusual length, the
end-waHs offer considerable resistance for transferring
the wind to the ground, the roof acting as a horizontal or
inclined truss delivering the wind-load to the end-walls.
For this reason, such walls should be well braced, with
diagonal bridging.

Working Drawings

Not only must the designer of timber-structures be
able to compute the necessary sizes of the members, and
the details of the connections; he must be able also to
present his design clearly to the builder. This state-
ment is not peculiar to timber-framing, yet it needs to
be emphasized, especially in this connection. The engi-
neer accustomed only to steel design, and even the engi-
neer versed in reinforced concrete design, is prone to
leave much to the detailer, knowing that standard prac-
tice will govern many details, and that he will check
over such details after the design is completed, before
fabrication of the structural steel or the steel reinforc-
ing-bars is begun. As has been stated already in these
pages, standard details, in timber design, do not exist.
Left without working-details, and given sizes of main
members, the carpenter will build a structure. Whether
such structure will be safe, depends largely upon the
carpenter's experience. A timber cut too short may be
spliced with comparative ease, even if not with full
safety, and by means of saw, hammer, and a few nails, a

TIMBER FHAMINO 253

makeshift connection tiiat may appear to be of Mufflcient
strength can always be accomplished.

In the preparation of drawings for a timber-framed
structure, two conditions present themselves, (1) when
the structure is to be built by contract, and (2) when it
is to be constructed by day labor or force-account. In
the first case, the engineer may require detail drawings
to be furnistied, just as for a steel-framed structure,
such details to be checked and approved by him before
any material is bought. For tlie steel and iron-work,
this method may be entirely satisfactory, provided that
the contract drawings and specifications show clearly
just what is wanted, since such detailing will in all prob-
ability be done by an experienced structural draftsman.
This is providing that the job is of sufficient magnitude
so that the steel and iron-work will be fabricated by a
shop of some size. For the timber-work, it will be neces*
sary for the designer practically to detail the job com-
pletely, as only in this manner can the desired connec-
tions be shown clearly. A case of an all-timber structure
where the designer can show a diagrammatic plan, eleva-
tion, and sections, giving sizes of members, and main
dimensions, and expect a draftsman to draw up satis-
factory details is practically an impossibility.

In the second case, where the structure is to be built
directly from the designer's plans, with no other details,
particular care should be taken to see that every impor-
tant member and connection is shown clearly. The steel
should be detailed accurately and fully, the number and
length of all rods, bolts, etc., listed, and all steel should
be designated in accordance with a (denr system of
marking. In this work, one day in the office is worth at
least two in the field. Then; is no Ix^ttcr check that can
be applied to drawings than to prepare an accurate list
of every piece of material in the structure. In fact, I
know of no better method to make one realize? the con-
venience, not to say necessity, of fully-d(!tailed drawings,
than to be compelled to make a complete detailed esti-
mate of cost. While in the case of small timbor-struc-

254 TIMBER FRAMING

tures, and for some larger ones, it is the custom of the
contractor or carpenter to order bolts and other small
steel material as he needs them in the course of construc-
tion, such a course will not be satisfactory on a large
structure. Even on a small job, it is an inefficient and
wasteful method.

The engineer will sometimes be called upon to furnish
plans and specifications for timber structures in isolated
localities, where all material needed for the job must be
purchased beforehand, and shipped to the site, and
where mistakes in ordering material or in showing de-
tails may cause serious delay and expense. For such a
condition, I believe that it pays well to mark every bolt
and rod, that is to say, all bolts of a certain length and
diameter are to be given a special mark, as a letter or
group of letters, or a combination of letters and figures,
in accordance with some definite system. For example,
all bolts in columns may be given the prefix C, as C-1,
C-2, etc. Not only should these marks appear in the
bolt-list after the particular bolt-size, but the marks
should be placed on the bolts on the drawing in the ele-
vation of the column. Further, the bolts should be
shipped in bundles of one size and length, bound together,
and tagged. This is, perhaps, going outside the domain
of strict design, and into the field of detailing and con-
struction, yet it should be a part of the designer's task, in
the case under consideration, to detail the work, and to
draw his specifications for the contractor furnishing the
iron-work so that the field work will be a minimum. This
suggestion as to marking applies to all steel of whatever
shape. Rods should be tagged, and structural shapes
plainly marked by painting, with the corresponding
marks at the proper places on the drawings. It is highly
desirable to mark the cutting lengths of the important
timbers on the drawings ; it is much simpler, and better,
for the designer, who, at the time, has the structure well
in mind, to note the lengths of timbers, than for the car-
penter to compute the lengths. Objection might be made
to this statement, on the ground that it puts the -re-

TIMBER FRAMING 255

sponsibility for accuracy on the engineer, rather than on
the carpenter. For a contract drawing, such a conten-
tion may hold ; for the detail drawing as required under
our present assumption, the argument is unsound. Thor-
. ough checking is essential, and such checking should al-
ways be given, even at the expense of having the owner
annoyed by an apparent needless delay in the comple-
tion of the drawings. After the structure is well under
way, and the work is progressing smoothly and rapidly,
the owner will forget any small delay in getting out the
plans and specifications; on the other hand, he will sel-
dom forget a mistake.

In the preparation of drawings for timber-framed
structures, there should be a general plan, framing-
plans, elevations, cross and longitudinal sections, and
details. The exact number of drawings, it is hardly
necessary to state, will depend altogether on the kind of
structure, and its simplicity or complexity. In general,
the plans as opposed to elevations, sections, and details,
should be to the scale of eight feet to the inch, or, as com-
monly called, i-in. scale. In some cases, it may be ad-
visable, for the sake of clearness, to use a larger scale, as
i in.; and certain small part^of the general plans may
need to be re-drawn to a J-in. scale, in addition to the
smaller scale. No matter how many parts of the build-
ing may be drawn to a large scale, as i in., a complete
plan to a ^-in. or i-in. scale is needed, in order that the
entire structure may be seen at a glance. The eleva-
tions can usually be shown to a ^-in. scale, and the gen-
eral cross and longitudinal sections to a ^-in. or ^-in.
The details should be at a scale not less than J-in.

In the case of a frame building of the mill-construc-
tion type, taking a typical example, of a building 100
ft. long, and one bay wide, trusses say 40-ft. span, corru-
gated-iron sides and roof, and floor of timber construc-
tion, about 3 ft. oflE the ground, the following plans will
show the work completely:

(1) One sheet, to a ^-in. scale, showing the four eleva-
tions, with all window and door openings, the doors and

256 TIMBER FRAMING

windows being lettered or numbered to correspond with
details of same.

(2) Foundation-plan, to a i-in. scale, showing size and
position of piers and wall-footings, with i-in. or ^-in.
details of the individual footings and piers.

(3) Floor-framing plan, to a J-in. scale, showing sizes
of joists and girders and posts, with all dimensions of
spacing of same, and centre lines of truss-posts, and
first-floor posts.

(4) Roof -framing plan, showing main trusses, with
their proper letters, bracing-trusses, bracing, roof-joists,
roof-covering.

(5) Cross-section for the building to a ^-in. scale, com-
pletely detailed as to roof-joists, trusses, columns, and
floor-construction.

(6) Miscellaneous- timber details to a |-in. scale, as
may be necessary.

(7) Details of all fabricated steel to a 1-in. scale.

In general, such scales as ^-in., and f-in., should be
avoided, although no hard and fast rule can be made.
An architect employs a J-in. scale to show details on a
contract drawing. It is often convenient, therefore, to
use the same scale when preparing structural drawings
for an architect; the architect's tracings may be super-
imposed on the structural drawings, and vice versa.
Mistakes of clearances may sometimes be found in this
manner. However, on the other hand, there is often an
advantage in re-drawing the architect's outlines within
which the engineer must confine his work ; errors of scale
are discovered in this manner. The converse is also
true; the architect may find mistakes in the engineer's
drawing when he lays it out on' the architectural sheets.

It is unwise to furnish a drawing that is badly out of
scale, even if it is fully and accurately dimensioned.
This statement holds for construction in any material,
but is especially true in timber framing, as the carpenter
is almost sure to scale some timbers. For this reason,
considerable erasing, and even re-drawing and re-tracing
will be well worth the effort and expense, if, by such

TIMBER FRAMING 257

extra work, a drawing badly out of scale may be made to
scale. Serious errors on the carpenter's or contractor's
part may thus be avoided.

Finally, a general and comprehensive note should be
placed on all structural drawings. This procedure may
not be in accordance with the theory held by many, that
written instructions are specifications, and as such,
should not appear on the drawings. If this view is held,
allow the specification writer to incorporate such instruc-
tions to the contractor in his specifications, but be
verbose to the extent of repeating the more important
points on the drawing in a general note. The specifica-
tions, bound separately from the plans, often become
separated from them. Notes on a drawing cannot be de-
tached from the details. Finally, it is a curious fact that
a note on the drawings carries about twice as much
weight with a carpenter as an obscure sentence in the
specifications.

258 TIMBER FRAMING

CHAPTER XVI
Speciflcations for Timber Framing

The following specifications are primarily for timber-
framed mill buildings, to be constructed of Douglas
fir. The unit stresses for timber, as given, are for par-
tially air-seasoned timber, as distinguished from thor-
oughly seasoned material or from green timber. Fur-
ther, the unit stresses are for the grade of timber known
as No. 1 Common.

When the conditions are different from those just out-
lined, as, for example, green timber, structures ex-
posed to the elements, or lumber containing No. 2 Com-
mon, lower unit stresses are to be used, and the propor-
tional decrease in stresses must depend upon the judg-
ment of the designer, in accordance with the particular
conditions.

For the case of bridges, either railway or highway,
the specifications of Milo S. Ketchum, as given in his
* Structural Engineers Handbook' shall be used.

Contract Plans

Unless specifically stated otherwise, the plans to be
furnished are to be what are known as 'Contract
Drawings.' That is, the drawings and specifications
are to show the structure in such detail that the exact
amount of all material may be determined without re-
sorting to computations for strength of any member
or detail of the structure, but subsequent shop and field
details will be required, the same to be checked in a
general way by the engineer for strength, but not for
accuracy of detail-dimensions.

To this end, there shall be furnished, in general, a
foundation-plan, framing-plan, sections, and elevations,
and typical details of all connections, sufficiently di-

TIMBER FRAMING 259

mensioned and noted, so that the detailer may under-
stand fully the requirements of the design. The speci-
fications shall state the kind and quality of all material
entering into the structure and shall give all other in-
formation and requirements that the fabricator and
erector of the structure will need in order to produce
a workmanlike job in conformity with the requirements

of the design.

In the case of building-plans, the scope of such plans
may be more specifically stated as follows. There shall
be furnished a general ground-plan, foundation-plan,
floor-framing plan or plans, depending upon the num-
ber of floors, roof-framing plan, typical sections, cross
or longitudinal, and details of all important connec-
tions.

Scale. The scale for framing-plans shall be i in. or
i in. to 1 ft. The same scale shall be used for elevations
and small sections. Larger sections in which it is desired
to show connections of members in addition to the gen-
eral arrangement of structural members, shall be on a
. scale of i in. or f in. to 1 ft., preferably the former. De-
tails of steel and iron-work, as shoe-plates, washers, etc.,
shall be at a scale of not less than f in. and preferably to
a scale of IJ in. to 1 foot.

Detail Specifications — Structures of the Mill-Building

Type

Under this class will come mill-buildings, power-
houses, pump-houses, machine-shops, armories, skating-
rinks, amusement pavilions, exposition buildings, etc.

Roof Loads. For localities where a snow-load can-
not occur, the following minimum loads shall be used.

1. Dead Load. The dead load shall consist of the
weight of the roof-covering, rafters, purlins, roof-
bracing truss, and ceiling, where the latter occurs. The
weight of the roof -covering, rafters, and purlins shall
be taken as applied at the panel-points of the upper end
of the truss. The weight of the roof-truss for light
trusses may be considered as concentrated at the upper

260 TIMBER FRAMING

chord. For roof-trusses, in which the dead weight of
the roof -truss is over 15% of the total dead and live
load supported by the truss, and including the weight
of the truss itself, the weight of the truss shall be con-
sidered as applied equally at the upper and lower-chord
panel-points. The weight of the ceiling, where such
occurs, shall be considered as concentrated at the lower
chord.

2. Live Load. The live load on the roof shall be
taken at 20 lb. per sq. ft. of projected area for rafters
and purlins, and the same figure shall be used in com-
puting bending in the top chords of the trusses. The
roof-covering shall be designed for a load of not less
than 30 lb. per sq. ft. and computations shall be made
both for strength and stiffness. The live load on the
trusses shall be taken at not less than 15 lb. per sq. ft.
of tributary area.

3. Wind Load. The wind load on the roof of build-
ings not over 25 ft. to the eaves, shall be considered as
applied normal to the roof-surface, and the amount
of such normal wind-load shall be computed by Duche-
min's formula.

^ 2 Bin e ,
^ = ^ i4-8in'e > ^^^^^

p = normal pressure on roof in lb. per sq. ft.
P = 15 lb. per sq. ft.

= angle which the plane of the roof -surf ace makes
with the horizontal.

For buildings over 25 ft. in height to the eaves, P in
Duchemin's formula shall be taken at 20 lb. per sq. ft.

Walls. For buildings not over 25 ft. in height to the
eaves, the wall-covering and girts or studs shall be de-
signed for a horizontal wind-pressure of not less than
20 lb. per sq. ft, and the columns, when forming a
transverse bent with the roof-trusses, shall be designed
for a horizontal wind-pressure of 15 lb. per sq. ft. For
buildings over 25 ft. in height to the eaves, the above
pressures shall be increased 5 lb. per square foot.

TIMBER FRAMING 261

All roof-trusses and columns shall be designed for
the maximum of the two following conditions.

1. Dead load plus wind load.

2. Dead load plus live load.

Unit Working-Stresses

Douglas Fib Timber (Grade No. 1 Common)

Lb. per sq. in.

Tension with fibres 1,500

Compression, end-bearing 1,600

Compression across fibres 300

Bending, extreme fibre-stress 1,500

Modulus of elasticity:

(o) For dead load only 1,200,000

(ft) For live load only 1,600,000

Shearing with grain 150

Longitudinal shear in beam 175

Columns :

For columns under 15 diameters 1,200

For columns over 15 diameters P = 1600 | 1 - 777, tB

V 60 d/

where P = unit working-stress in lb. per sq. in for centric loads

L = unsupported length of column in inches

d = least width of column in inches

Steel : Lb. per sq. in.

Tension 16,000

Shear 10,000

Bearing 20,000

Cast Iron:

Bending, extreme fibre stress 4,000

Tension 3,500

Pressures on Inclined Surface of Timber. The safe
unit working-compression on timber on surfaces in-
clined to the fibres shall be taken in accordance with
the formula:

n = p sin^ Q + q cos^
Where w = allowable unit compression on inclined

surface
p = allowable unit compression on ends of

timber
q = allowable unit compression across fibres
= angle which surface makes with the di-
rection of the fibres.

262 TIMBER FRAMING

•

Pressure of Circular Iron Pin on Timber. ' The safe
average unit stress on the diametrical section of an
iron pin bearing on timber in a close fitting hole, shall
be taken as follows :

1. When the direction of loading is parallel to the
length of the fibres,

p' = h + iq

2. When the direction of loading is perpendicular to
the length of the fibres,

r = ip + h

In these formulas, /?' and p'^ are the safe average
unit-stresses on the diametrical sections of the pin,
parallel and perpendicular, respectively, to the direc-
tion of fibres.

Strength of Nails When Used with Douglas Fir.

1. Lateral strength of wire nails. The safe working-
resistance of wire nails or spikes to lateral shear for
static loads, bearing either against the ends of the
fibres of the timber, or across the fibres, shall be taken

as follows :

Safe lateral
Size of nail resistance, lb.

6D 48

8D 64

lOD 80

12D 96

16D ^ 128

20D 160

30D 240

40D . . , 320

50D 400

60D 480

SOD 640

2. Resistance of wire nails to withdrawal. The safe
working-resistance of wire nails or spikes to with-
drawal from timber, when the nail or spike is driven
perpendicular to the fibres of the timber, shall be
taken at 75 lb. per sq. in. of contact surface of wood
and nail. For nails driven parallel to the fibres, the

TIMBER FRAMING 263

safe loads shall be taken at 25 lb. per sq. in. of contact
surface of wood and nail.

Strength of Common Wire Screws When Used with
Douglas Fir: 1. Lateral resistance of screws. The
safe working-resistance of wood screws to lateral
shear, for static loads, bearing either against the ends
of the fibres, or across the fibres, shall be taken as fol-
lows:

Safe lateral
Gauge of screw resistance, lb.

12 205

14 256

16 .315

18 380

20 450

22 529

24 615

The length of the screw shall be approximately two
and three-quarters times the thickness of the side-piece.

2. Resistance to withdrawal. The safe working
resistance of wood screws to withdrawal from timbers,
when the screw is inserted perpendicular to the direc-
tion of fibres shall be taken as follows :"

Safe resistance to
withdrawal per linear
Gauge of screw inch of insertion, lb.

4 75

8 100

12 125

16 140

20 150

22 170

28 185

For screws inserted parallel to the fibres, the safe
working resistance to withdrawal shall be taken at
75% of the above values.

Strength of Lag-Screws When Used With Douglas
Fir: 1. Lateral resistance when used in fastening
planking to large timbers. The safe working-resistance
of lag-screws to lateral shear when used in fastening
planking to large timbers shall be as follows :

264 TIMBER FRAMING

i by 4i in. lag-screws 900 lb.

J by 5 in. lag-screws 1050 lb.

The thickness of such planking shall not exceed f of
the length of lag-screw.

2. . Lateral resistance when used in fastening steel
plates to timbers. The safe working-resistance of lag-
screws to lateral shear when used in fastening metal
plates to timbers, when such plates are not less than
i in. to i in. thick, shall be taken as follows :

i by 4 in 700 lb.

f by 4 in 860 "

1 by 4i in 1030 "

I by 5 in 1200 "

3. Resistance of lag-screws to withdrawal ; The safe
working resistance of lag-screws to withdrawal, when
inserted perpendicular to the fibres, shall be taken at
180 lb. per sq. in. of the surface obtained by multiply-
ing the nominal diameter of screw by the length of the
threaded portion of screw, excluding the tapering end.

Strength of Bolts

1. Bolts in Double Shear — ^All end-bearing on tim-
bers. The strength of bolted timber joints should be
computed by the methods explained in the text of
Chapter V.*

For a given diameter of bolt, the length I shall be
found, by use of the formula

32 5

when l = a-\-}> (See Fig. 36).

d = diameter of bolt in inches.

Sf = maximum allowable flexural unit stress in

the bolt = 16,000.
B = maximum allowable unit bearing-stress against
the ends of fibres of timber.

♦For a practical solution of the strength of bolted joints it
will be necessary to construct diagrams similar to that of Fig.
37 of Chapter V. Tables can then be prepared that will cover
the ordinary range of construction.

TIMBER FRAMING 265

^ = thickness of splice-pad = one-half thickness
, of main timber.
a, h, B^, P, P and P2 = as shown in Fig. 36.

The joint will be in one of the two following classes:

A. Thickness of splice-pad equal to or greater than
computed value of I.

B. Thickness of splice-pad less than I.

If the point falls under class A, the strength of the
joint for one bolt shall be found by means of the formula :

P = iBtd

where P = total safe load on joint for one bolt in double

shear and bending.
If the joint falls under class B, it shall be still further
classified as follows :

a. Pressure uniform along length of bolt.

b. Pressure distribution along length of bolt trape-
zoidal in shape, with a maximum intensity B at contact
faces of main timber and splice-pads. The lowest unit
pressure will be B% at the centre of main timber, and at
the outside-faces of the splice-pads. The value of B'
will vary between the limits B' = and B' = B,

e. Pressure distribution along length of bolt trian-
gular, but with varying values of a and 5, the limits
being a = and a = i,

2. Bolts in Double Shear — Centre or main timber
with bearing across the fibres, splice-pads or outside
timbers with end-bearing.

JJke values two-thirds those of Class l.f

Drift-Pins: The safe resistance of round drift-pins
to withdrawal, when such drift-pins have been driven
perpendicular to the fibres of the timber and in holes of
a diameter not greater than ^ of the diameter of the
drift, shall be taken at 180 lb. per sq. in. of contact-
surface of wood and metal.

When such drift-pins are driven parallel to the fibres,

tThe case of metal plates bolted to timbers is purposely
omitted, as I prefer to await the publication of tests which
have been made.

266 TIMBER FRAMING

and in holes of a diameter not to exceed | of the diam-
eter of the pins, the safe resistance to withdrawal shall
be taken at 90 lb. per sq. in. of contact-surface of wood
and metal.

The safe resistance of such drift-pins to pulling
through the timber in the direction of driving shall be
taken at not to exceed 60% of the above values.

Shear-Pins: The safe working-resistance of 2-in.
circular shear-pins of solid steel, extra heavy steel pipe,
Hawaiian ohia, or Australian iron bark shall be taken
at 800 lb. per linear inch of pin. Bolts shall be pro-
vided with a total capacity in tension equal to one-half
the total load on the joint. Pins shall be spaced not
closer than six inches centre to centre.

The shear-pin joint shall be used with seasoned tim-
ber only.

General Conditions of Framing

Special attention shall be paid to laying out column
centres, and the general arrangement of trusses, posts,
girders, and joists, in order that a stiff structural frame
may be secured. To this end, the roof-trusses shall be
well braced, both upper and lower chords, by means of
bracing-trusses, or their equivalent. The wind-stress
on the building shall be carried to the foundations
through the structural frame, and all parts thereof
shall be consistently designed to accomplish this pur-
pose.

Roof-joists shall be lapped over truss-chords not less
than 12 in., and spiked well to each other and to the
truss-chord, or if such roof-joists abut over the chord,
splice-pads not less than 2 ft. long shall be provided on
both sides of each joist.

Knee-braces to trusses, when used, shall be attached
rigidly to the truss and post, and shall meet the truss at
a panel-point only.

Interior floor-columns, when possible, shall be in line
with the wall-posts, and these cross-lines of posts shall
be well tied together by means of girders or joists. For

TIMBER FRAMING 267

this purpose, the joists shall, when possible, be so spaced
at the posts, that two joists shall tie the lines of columns
together. When the girders frame into the posts, the
girders shall be well tied to the posts by means of splice-
pads.

Girders framing into the sides of posts shall be sup-
ported, when possible, on side-bolsters, dapped into the
posts and bolted to them, and such bolsters shall have
all end-bearing.

In buildings of a height of 20 ft. or over to the eaves,
diagonal bracing-rods shall in general be provided in at
least every other one of the outside-bays, in the plane
of the upper or lower chords of the roof -trusses. Special
cases may occur where this requirement is not neces-
sary, and this will depend upon the judgment of the
engineer.

When timber posts rest upon concrete foundations, a
steel or iron base plate shall be provided between the
concrete and the bottom of post.

For all structures exposed to the weather, special at-
tention shall be paid to the detailing of joints, in order
that such finished joints shall shed rather than hold the
water that would tend to collect from rain. In the
case of such structures, all bearing-surfaces of timber
to timber, and timber to iron, and also all ends of tim-
ber, shall be treated with one good coat of wood-pre-
servative. When the importance of the structure will
permit, steel or iron bearing-surfaces shall be provided
at the ends of timber that is bearing against the side of
timber.

Roof Covering

Corrugated Steel: When used for permanent build-
ings, corrugated steel shall never be less than No. 24
gauge. When this weight of steel is used, the maximum
spacing of purlins shall not exceed 4i ft. The end laps
shall be not less than 6 in. and side laps not less than two
corrugations.

Timber Sheathing: Timber sheathing shall be dress-
ed and matched lumber, free from loose knots, and of a

268 TIMBER FRAMING

width not to exceed 6 in. Purlins or rafters shall be
spaced so that the deflection for a live load of 30 lb. per

sq. ft. shall not exceed j^ of the span of the sheathing,
using the formula:

^ = 128 W

where A = centre deflection in inches for uniform

loading.
w = load per square foot.
I = length of clear span in inches.
JS; = modulus of elasticity.
/ = moment of inertia.
For sheathing of a nominal thickness of one inch, and
covered with prepared roofing, the spacing of rafters for
permanent buildings, shall preferably not exceed two
feet.

Details of Roof Trusses

When the roof joists rest directly upon the upper
chords of trusses, the bending-stresses in the chords re-
sulting from such condition of loading shall be computed,
and the sum of the direct compression and compression
due to bending shall not exceed 1600 lb. per sq in. nor
shall the direct compression exceed the safe unit-stress
considering the chord as a column.

In computing bending in the chords, such chords may
be regarded as continuous beams, supported at the panel
joints.

The area of holes for bolts and rods through both com-
pression and tension members of timbers shall be de-
ducted, in order to obtain the net section to resist com-
pression, tension, and bending, the diameter of holes for
rods being assumed as i in. larger than the nominal di-
ameter of rod or upset end of rod.

In general, full deduction shall be made for notches
cut in truss chords for butt-blocks and web-members.
In special cases, where such provision will add consider- >.
ably to the cost, and where the designer is to have full
control of framing, such deduction need not be made in

TIMBER FRAMING 269

the case of the compression-chord, provided that such
notches occur only on the compression-side of the chord
regarded as a continuous beam for transverse bending,
and provided that the butt-block detail is employed.

All tension-rods shall be of steel, conforming to the
Manufacturers' Standard Specifications.

If upset ends are used, such upsetting shall be done by
machine. No welding will be allowed.

All rods on trusses shall be given an initial tension of
at least 1500 lb. and allowance for such tension shall be
made in the design.

All joints of end or batter posts of trusses with the
lower chord shall be provided with a proper detail, capa-
ble of developing the computed stresses in the truss mem-
bers. Such detail of end-joint shall provide definite lines
of action, and such joint shall be, as far as possible, a
simple joint, depending for its strength upon one type
of detail.

When inclined bolts are used to connect the main mem-
bers of an end-joint, such bolts shall not have a greater
slope than 60° with the centre line of lower chord.

In details of end-shoes employing lugs or tables set
into the lower chord, the spacing of such lugs or tables
shall be arranged so that no lug or table occurs directly
under the end of the upper chord or the batter-post.

The holes in the timbers for inclined bolts in details
employing end-shoe plates shall be J in. larger than the
nominal diameter of bolt.

No daps in chords for butt-blocks shall be less than
5 in. deep.

The minimum thickness of metal in shoe-plates shall
be f inch.

Steel Lugs and Tables : (Applies particularly to end-
shoe plates and tension-splice plates) .

The bearing faces of lugs or tables shall have a smooth
even surface. If rolled bars are used for tables, they
shall be milled on the bearing edges.

The bolts holding the lugs or tables in the notches in

270 TIMBER FRAMING

the timber shall be placed as near to the lugs or tables as
possible.

When rivets are countersunk on one side, in plates
less than f in. thickness, the values shall be taken at
7500 lb. per sq. in. for shear, and 15,000 lb. per sq. in.
for bearing.

All holes in metal over |-in. diam. shall be drilled, not
punched.

No steel lug or table shall have a thickness of less than
I inch.

Details of Columns

No column shall have a greater ratio of length to least
width than 60.

Columns may be considered as fixed at the ends, where
provision can be made for obtaining the condition of
fixedness assumed.

When bending resulting from wind occurs in columns,
the combined stress because of dead load, direct wind-
compression, and wind-bending shall not exceed the safe
unit-stress as given by the column formula, increased by
25%, considering the width of the column in the plane
of bending.

When the column is rigidly supported laterally at the
point of maximum combined stress, such maximum com-
bined unit-stress may equal but shall not exceed 2000 lb.
per sq. in., and the total combined unit-stress at the
centre of the section of column that is unsupported later-
ally, shall conform to the stress allowed by the column
formula, as outlined above.

Built-up columns shall be avoided whenever possible.
The strength of built-up columns, composed of two or
more sticks bolted together, either with or without pack-
ing-blocks, shall be considered as equal to the combined
strength of the single sticks, each considered as an inde-
pendent column.

When it is necessary to employ columns built of plank-
ing, such columns shall preferably be of the * cover-plate'
type, in which the edges of the interior planks are tied
together by cover-plates. The strength of such a built-up

TIMBER FRAMING • 271

column shall be considered as 80% of a solid stick of the
equivalent cross-sectional area. The strength of a built-
up column composed of planks laid face-to-faee and
spiked together thoroughly shall be considered as 80%
of the mean of the strengths computed (1), as a solid
stick, and (2), as a summation of the strengths of the
individual sticks considered as separate columns.

When columns are built of large timbers placed at a
considerable distance from each other, such timbers shall
be tied together by means of 2 by 12-in. lacing-plates, in-
clined at an angle of approximately 60° with the axis of
the column, and fastened to the column by means of lag-
screws, not less than J by 6 inches.

When such columns are built of two timbers, laced on
both sides, the effective moment of inertia of such built-
up columns shall be taken as 80% of the theoretical
moment of inertia of the column.

Curved laminated compression members shall be avoid-
ed when possible. Where it is necessary to employ such
members, the strength shall be computed in accordance
with the principles of Chapter IX, taking into account
average unit-compression, flexural stress resulting from
bending and from eccentricity of loading, and initial
flexural stress resulting from springing the boards to a
curved shape.

Bolts. All bolts shall be provided with cast-iron or
steel plate washers, of a size such that the unit-stress in
cross-bearing on the timber under the washer shall not
exceed the safe unit-stress for cross-bearing when the bolt
is stressed in tension to 16,000 lb. per square inch.

Bolts shall preferably be spaced not closer than 6 in.
centre to centre, and not less than 6 in. from the end of
any timber, nor less than 2^ in. from the sides of any
timber. This rule shall apply to bolts of sizes not to
exceed 1 in. diam. For larger sizes the minimum. dis-
tances given above should be increased accordingly.

All bolts except as otherwise specified, shall be driven
in holes of a driving fit.

Inclined bolts through timber shall preferably be pro-

272 * ' TIMBER FRAMING

vided with beveled cast-iron washers, instead of using
standard washers and cutting inclined daps in the
timber.

Lag-ScrewB. All lag-screws shall be screwed, not
driven into place.

All lag-screws fastening timber to timber, shall be pro-
vided with standard circular pressed-steel washers under
their heads.

Holes for lag-screws in steel plates shall be drilled to a
diameter of ^ in. larger than the nominal diameter of
the lag-screw. In placing lag-screws, a hole shall first be
bored of the same diameter and depth as the shank, and
the hole tljen continued with a diameter equal to the
diameter of the screw at the root of the thread.

Drift-Pins. Drift -pins shall preferably be round, with
or without heads, and shall be driven in holes of a diam-
eter of approximately 80% of the diameter of the pin,
and of a length somewhat larger than the length of the
drift-pin.

Tension-Splices. Tension-splices shall be of such a
type that the effects of cross-shrinkage of the timber will
be a minimum. Neither the tabled-steel fish-plate, nor
the shear-pin splice shall be used on timbers over 8 in.
thick, since the cross-shrinkage of the timber will allow
the splice-plates or pads to separate.

INDEX 273

INDEX

Page

Anchorage, for columns 220

Bolts, lateral resistance of 77

Bolts, strength of, in double shear 264

Bracing-trusses 160

Checks, in timber 13

Chipped-grain 13

Chords, compression 139

Column-action, theory of 184

Column-anchorages 220

Column-connections 194

Columns, details of 269

Composite compression members 142

Compression chords and struts 139

Compression on inclined surfaces 45

Compression-splices 135

Connection of joists to girders 199

Connection of truss to post 179

Connections, column .^. 194

Contract drawings 258

Contract plans 258

Corrugated steel for roof-covering 267

Curved laminated truss-chords 147

Dead load 259

Design of flumes 224

Design of head-frames 228

Design of water-tower 234

Details of columns 270

Details of Howe-type roof-truss 163

Details of roof-trusses 268

Dimension lumber 22

Drift pins 265

Drawings, contract 258

Drawings, working 258

End-joints of trusses 90

Fir bridge-stringers 23

Fir car-material 23

Fir timbers 22

Fish-plate type of splice 120

Fleming, R., discussion of wind-stresses by 248

Flume-design 224

Foundations 209

274 INDEX

Foundations, pile 215

Grading rules 11

Head-frames, design of 228

Head-frames, discussion of, by Robert S. Lewis 238

Howe roof-truss, details of 163

Intermediate joints of trusses 112

Joints, end of trusses .-. 90

Joints, intermediate of trusses 112

Joist-hangers 200

Ketchum, M. S., discussion of wind-pressure by 246

Knots 14

Leg-screws, resistance to withdrawal of 263

Lag-screws, lateral resistance of 72

Laminated compression members 142

Lattice-trusses 169

Lewis, Robert S., discussion of head-frames and ore-bins by 238

Live load 260

Load, dead 259

Load, live 260

Load, wind 260

Loads, roof 259

Lugs, steel 269

Mill construction 205

Miscellaneous structures 223

Nails, lateral resistance of 58

Nails, resistance to withdrawal of 262

No. 1 common lumber 24

No. 2 common lumber 24

Ore-bins, discussion of, by Robert S. Lewis 241

Pile foundations 215

. Pins 47

Pins, drift : 272

Pins, shear 266

Pitch-pockets 17

Pitch-shakes 13

Pitch-streak 13

Plans, contract 258

Roof-covering 267

Roof loads 259

Roof-trusses, details of 268

Roof-truss, Howe type, details of 163

Sap, in timber 17

Scales for drawings 259

Screws, lag, lateral resistance of 72

Screws, wood, lateral resistance of 66

Screws, wood, resistance to withdrawal of 263

Shear-pin joint 55

INDEX 275

Shear-pins 266

Sheathing, timber, for roof-covering 267

Specifications 258

Spikes, lateral resistance of 58

Splices, bolted fish-plate type 120

Splices, compression 119, 135

Splices, design of 78

Splices, tension 119

Splits 13

Steel tables and lugs 269

Stresses, unit working, for timber 261

Stresses, unit working, for steel 261

Stresses, unit working, for cast-iron 261

Standard sizes of lumber 19

Steel tables 269

Tenon-bar type of splice t 128

Tension-members made of timber 155

Tension-rods 155

Tests of timber columns 188

Timber columns, tests of 188

Timber columns, working strength of 193

Time element, effect of on strength of timber 36

Torn grain in timber 13

Trusses, bracing 160

Truss, connection of to post 179

Trusses, end-joints of 90

Truss, Howe-type, roof, details of 163

Trusses, lattice 169

Trusses, intermediate joints of 112

Unit stresses for timber 261

Unit stresses for steel 261

Unit stresses for cast-iron 261

Unit stresses, recommended by American Railway Engi-
neering Association 27

Walls, specifications for 260

Wane, in timber 13

Washers 39

Water-tower, design of 234

Western hemlock, description of qualities of 24

Wind load 260

Wind pressure and stresses 246

Wood-screws, lateral resistance of .^ 66

Working drawings 252

Working-stresses for cast-iron 261

Working-stresses for timber 261

Working-stresses for steel 261

Yard lumber 11

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