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Bolted Wood-Steel and Wood-Steel-Wood Connections:
Verification of a New Design Approach
M. Mohammad
Research Associate, Department of Civil Engineering, Royal Military College of
Canada, P.O. Box 17000, STN Forces, Kingston, Ontario K7K 7B4
J.H.P. Quenneville
Associate Professor, Department of Civil Engineering, Royal Military College of
Canada, P.O. Box 17000, STN Forces, Kingston, Ontario K7K 7B4
(Word Count = 6504)
ABSTRACT: This paper covers the verifications tests carried out at the Royal
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2
Military College of Canada (RMC) on wood-steel-wood and wood-steel bolted
connections. Thirty groups of specimens were tested. Specimens configurations were
selected in such a way to include fundamental brittle and ductile failure modes cases.
Comparisons between experimental results and predictions from proposed equations
developed from steel-wood-steel bolted connections are given. Proposed design
equations were found to provide better predictions of the ultimate loads than current
CSA standard O86.1 design procedures especially for bearing. However, row shear-
out predictions seem to over-estimate the strength. An adjustment using the reduced
(effective) thickness concept is therefore proposed. Experimental observations on
specimens that failed in row shear-out indicated that shear failure occurred over a
reduced thickness. Stress analysis confirms findings on the reduced thickness. The
research program is described in this paper along with the results and the proposed
design equations for wood-steel-wood and wood-steel bolted connections loaded
parallel-to-grain.
Key words: wood-steel-wood, wood-steel, bolt, connection, strength, failure, design,
thickness.
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BACKGROUND
In timber structures, different types of connectors and connections
configurations are used. Bolted timber connections however, are one of the most
popular types used in North America. Steel plates are used to connect timber
structural members to secure a proper transfer of forces from one structural member
to another. Two steel side plates are often used to connect members and are often
referred to as steel-wood-steel connections (SWS). However, due to some structural
and architectural requirements, a single steel plate may be used. Steel plates could
be inserted within the timber member or could be installed between two timber
members. Those kind of connections are known as wood-steel-wood connections
(WSW). Occasionally, a steel plate is used to transfer the load from one single
member to another. This is usually used in light timber structures and are referred to
as wood-steel connections (WS).
The problem of predicting the strength of multi-bolted connections is a well
known one. For the last fifty years, the strength of connections that failed in a ductile
fashion has been understood and predicted well by the European engineering
community. As a result of this, some European wood design codes emphasize the
importance of using many small diameter fasteners instead of few large diameter
ones so as to obtain the ductile failure modes. In North America however, the
engineering community has been slow to adopt the European Yield Model (EYM) and
the normal practice in bolted connection design is to use fewer large diameter bolt to
save on fabrication costs.
The current design equations in the Canadian design code (CSA 1994), are
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4
based on work by Johansen (1949) and further modified by Larsen (1973) and were
first introduced into the Canadian design code in the 1989 edition (CSA 1989). In this
design approach, failure is assumed to be governed by bearing (crushing of wood)
and/or bending of the bolts (Mode I and Mode III according to the European Yield
Model (EYM)). This assumption results in a ductile failure mode for connections. This
is not always the case even when minimum requirements for spacing, end distances
and edge distances are met. Consequently, the failure modes that show brittle failure
(which are typical in connections with multiple fasteners) had to be addressed by
modifying the EYM.
However, when using large fasteners, brittle failure modes such as splitting,
row shear-out, tearing and a combination of tearing and shear-out (known as a group
tear-out) are the norm. This has been confirmed by test results from various sources
(Yasumura et al. 1987, Mass et al. 1988, Mohammad et al. 1997, Quenneville and
Mohammad 2000), and especially for multiple fastener connections and connections
with low slenderness ratio (l/d) fasteners. These modes of failure can not be
predicted by the EYM, resulting in discrepancies between design strength values and
actual experimental ones.
The current design model in the Canadian code for bolted timber connections
(CSA 1994) assumes that connections will fail in a ductile manner. To account for
situations where the connections show a brittle behaviour (generally connections with
multiple bolts), modifications factors were introduced. Test results conducted on
double shear steel-wood-steel bolted connections using 12.7 mm or 19.1 mm bolts
have shown that the current Canadian design approach leads to conservative design
strength values. Connections resistances as calculated from the O86.1-94 design
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5
N)sMIN(e,f'Jtnn2p bvrwsruRS =
code were found to be as low as one third of the experimental values (Quenneville
and Mohammad, 2000). This leads to the connections being over-designed.
It is recognized that there is a need for a fundamental design method for bolts.
The most desirable approach would be similar to the one used for other construction
materials (i.e. steel), where a two-step process is utilized. The first step would be to
check yield failure in the bolt, and is calculated by multiplying the capacity of one bolt
in the connection times the number of bolts. The second step consists of checking
the failure around the bolt, and is calculated by determining bearing and the
combined tension and shear capacity of wood. This step depends on the joint
configuration, spacing, end distances, etc.
Proposed Design Equations for Steel-Wood-Steel Connections
In an attempt towards developing a more rational approach to determine
connections design strength, a set of equations has been developed by Quenneville
to predict the ultimate strength of connections based on the actual failure modes and
mechanisms observed during tests (Quenneville and Mohammad, 2000). Design
equations were derived so that specified strength values for materials as listed in
O86.1-94 could be used. Failure modes covered in these design equations were row
shear (RS), group tear-out (GT) and bearing (B). The connection strength (p u) would
be the minimum of puRS, puGT and puB. Design equations proposed by Quenneville
(1998) are given below.
Row shear-out:
[1]
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( ) )2))f(d-1)(s-(nt(N)MIN(e,sf'2tnp tgrrwbvwsuGT ++=
Nntdf0.8p rwwuB1 =
++
=dt
51
f
f
)f(ff
61
Nnn2
df0.8p ww
y
sw
srswuB2
= 0.25,
t
)sMIN(e,N0.0851.085MAXff'
w
bvO86.1v
w
y
sw
srswuB3
f
f
)f(f
f
3
2Nnndf0.8p
2
+=
Group tear-out:
[2]
Bearing:
[3]
[4]
[5]
where,
Validation tests on SWS bolted connections were carried out to compare
predictions from proposed equations with those of the O86.1-94 values (Quenneville
and Mohammad, 2000). A reasonably good agreement was found especially for row
shear-out and group tear-out. However, proposed design equations for WSW and
WS bolted connections were not validated and there was a pressing need to carry
out some extra connection tests to ensure that changes in bolted connection design
equations were well justified and that they have been validated for practical
situations.
So, the main objective of this research project was to verify if the proposed
design equations are valid for WSW and WS connections with single or multiple
bolts. This will be accomplished by comparing predictions from the proposed design
equations with experimental results. These tests are also required to complement the
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results database already available at the RMC.
MATERIALS AND TEST PROCEDURES
Specimens
Thirty groups of 10 replicates each were used in this study. Details of the
groups are given in Table 1. Specimens were made of either glulam or lumber
members. Glulam specimens consisted of either a single member (130mm wide) or
two members (80mm wide each) with a steel plate in the middle (WSW connections).
The reason for choosing the two sizes of glulam members was to compare the
response of the connections, using a single wood member with a slot in the middle or
when using two separate members. Glulam specimens were either Spruce-Pine (S-P)
grade 20f-EX or Douglas fir 20f-EX and were either 80mm x 190mm or 130mm x
190mm. A slot of 10 mm wide was made in the center of the 130mm x 190mm glulam
members to accommodate a 9.5mm steel plate using a chain saw rigged to a cutting
table (referred to in Table 1 as Insert). Other groups were fabricated with two
members of 80mm x 190mm sandwiching a 9.5mm steel plate. One group was
fabricated with a single 80mm glulam member and a steel side plate (group 12).
Lumber specimens were made of either one, two or three members nailed together
and were bolted to a steel plate (WS connections). Specimens made with lumber
were S-P-F, No. 2 and better 38mm x 140mm. These groups were fabricated using
two or three lumber members nailed together using 90mm nails. Two types of bolts
were used in this study, 12.7mm and 19.1mm. No particular effort to have matched
group was made for these specimens. However, glulam billets were purchased in
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different batches. Upon delivery, billets were cut and pieces were distributed to
alternate groups. Glulam and wood members were stored in a conditioning chamber
to attain a 12% equilibrium moisture content (EMC). Detailed connections
configurations are listed in Table 1. Although 10 physical specimens were tested
per group, 20 connections were tested (2/specimen).
Test Set-up and Procedures
A typical test set-up is shown in Figure 1. All bolts were finger tight to allow a
self-alignment. Specimens were loaded parallel to grain and were fabricated with
identical connection configurations at each end. A universal loading machine (MTS)
was used to apply the load. A monotonic tension load was applied through the central
steel plate (WSW) or through the side steel plate (WS). Four linear variable
displacement transducers (LVDTs) were used to record the slip of the wood side
member(s) with reference to the steel plate (two at each end). A data logging system
was used to record the machine load and slip from the four LVDTs. An initial pre-load
of about 1.0 kN was applied to the specimens. The test was displacement driven at a
rate of 0.9mm/min. (0.035in/min.) in accordance with ASTM standard D07.05.02
(ASTM 1988). Tests were stopped upon failure, when the load dropped with no
recovery.
RESULTS AND DISCUSSION
Test results are given in Table 2. Figure 2 shows typical load-slip relationships
for all specimens in group 5 (ductile) and group 9 (brittle). The ultimate strength
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values for each group was determined and the lower 5th
percentile value was
calculated using a two-parameter Weibull distribution based on a 75% confidence
level (ASTM 1994). Test results were adjusted for a normal duration of loading by
dividing by a factor of 1.25 (CWC 1995). Calculated values for connection are
presented in Table 2 for comparison. O86.1-94 values represent the lateral strength
resistances as determined from Clause 10.4.4 (CSA 1994). Predictions using the
equations proposed by Quenneville and modes of failure observed for each group
are given as well.
It should be noted that although 10 physical specimens were tested per group,
20 connections were tested (2/specimen). The 10 ultimate values are the lower
resistance of each 10 pair of connections.
Comparison Between Selected Groups
Connections configurations (i.e. loaded end distance, thickness, number of
shear planes ..etc) have a significant effect on their ultimate strength.
Generally increasing the end distance from 5d to 10d has increased the
ultimate strength considerably. Specimens in group 5 with an end distance of 5d have
a lower 5th
percentile value compared to those in group 6 with an end distance of 10d
by a factor of 0.66. The 5th
percentile value for specimens in groups 7 and 8 with
equal end distances (5d) are found to be the same.
When comparing the ultimate strength of specimens in groups 5 and 10 with
exactly the same configurations, except that group 10 was fabricated with two
members sandwiching a steel plate, it is evident that their 5th
percentile strength
values are similar (25.7 kN compared to 28.9 kN). The same could be found when
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comparing specimens in groups 7 and 8. This indicates that for this type of
connections, using two members sandwiching a steel plate or using a single one with
a steel plate in the middle does not have any significant influence on the 5th
percentile strength of the connections. This is quite interesting, since the cumulative
thickness of wood in group 8 is less than that of group 7 (120mm compared to
180mm).
Observations on Failure Modes
Generally, two dominant types of failures were observed in all specimens.
These were row shear-out and bearing. Splitting was observed but was not as
significant. Almost all WSW connections specimens fabricated with 12.7mm bolts
failed due to wood bearing. Yielding of the bolts was observed as well. Connections
fabricated with a single bolt and an end distance of 5 times the bolt diameter (5d) and
with a single 130mm wood member, exhibited considerable crushing prior to failure.
However, the final failure was mainly in row shear. Group 6 with a single bolt and an
end distance of 10 times the bolt diameter failed mostly in bearing. Splitting failure
which were observed after significant deformation, were followed by localized row
shear-out failure. Group 7, with two members and two bolts in a row, failed in row
shear-out, however, few specimens failed in splitting. Group 8, with two bolts in a row
and a spacing and end distance of 5d, failed mostly in row shear-out. Hardly any
signs of bearing were observed. The failure scenario for these few specimens was as
follows: splitting developed first, resulting in a sudden drop in the load, followed by
failure in row shear. Failure in WS connections specimens fabricated with a single
glulam member (group 12) was characterized by a row shear-out failure in one row
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followed by failure in the second. Bearing deformation was obvious in few specimens.
Groups of specimens in WS connections fabricated with 2 or 3 lumber members
failed ultimately in row shear-out, however, they exhibited considerable amount of
bearing deformation prior to failure. Localized drops in the load corresponded with
row shear-out failure in the individual members, with the member adjacent to the
steel plate being the first to fail, followed by the second or the middle (in the case of 3
members connections).
Comparison Between O86 Predictions and Experimental Values
Predictions using the current design code (CSA 1994) were found to be
conservative compared to the validation tests results (the lower 5th
percentile).
Excluding the groups for which the O86.1-94 predictions are zero, the ratio between
O86.1-94 values and the experimental results was found to be between 0.53 to 0.89
(with the exception of group 23 and 24) with an average of 0.73. It can be seen in
Figure 3 that the O86.1-94 predictions when plotted against the 5th percentile
experimental, lie below the 45o
line, thus considered to be conservative.
One reason for such discrepancies between O86 values and those of the
experimental ones could be attributed to the axial tensioning force that develops in
the bolt once the plastic hinge is developed. This axial force reinforces the
connection and results in the connection sustaining higher loads than anticipated. In
fact the axial force may even alter the mode of failure completely in some cases (i.e.
where row shear-out strength is not much higher than bearing strength of wood). The
influence of the axial force was obvious especially in group 6, where almost all tested
specimens exhibited that effect. This was observed in the load-slip curve as a
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discontinuity in the envelope around 55 kN, followed by an increase in the capacity of
the connection due to the development of a plastic hinge and the axial tensioning
force in the bolt. This effect is more pronounced in SWS connections due to the
anchorage provided by the steel side plates as compared to wood side members. It is
not surprising that O86 predictions for bearing based on the EYM under-estimate the
failure load, since the EYM does not take into consideration the axial tensioning force
that develops in the bolt. That explains the higher values for the experimental tests
compared to O86.1-94 predictions. Other reasons could be attributed to group and
loaded end distance modification factors (JG
and JL) used in the calculations of the
O86.1-94 values. These factors are very restrictive resulting further in under-
estimating the capacity of bolted connections.
VALIDATION OF PROPOSED EQUATIONS (QUENNEVILLE, 1998) FOR WSW
AND WS CONNECTIONS
In order to verify if the proposed design equations provide some reasonable
accuracy for WSW and WS connections, strength values calculated using the
proposed equations are compared with experimental values (5th
%). Figure 3
suggests a reasonable correlation between the experimental and the minimum
predicted values which are based on failure due to bearing (B), indicating that the
bearing equation (which are based on the EYM) are adequate for predicting bearing
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failures. However, as can be seen in Figure 4, where row shear-out failure governs,
row shear-out predictions (RS) were found to be considerably higher than the 5th
percentile as determined from tests, unlike RS predictions for SWS connections
(Quenneville and Mohammad 2000). The following discussion describes a theory
proposed by Jorissen (1998) that could explain such a discrepancy between
predictions using the proposed equations for row shear-out and those obtained from
tests for WSW and WS. It is based on the stress distribution underneath the bolts
and the observed failure patterns during the tests.
Analysis
In the new design equations for row shear-out predictions of bolted WSW
connections proposed by Quenneville (1998), the row shear-out failure was assumed
to occur over the full thickness of the wood member (tw was taken as equal to the full
thickness of the wood side member). This assumption is valid for connections with
rigid type of fasteners, where the embedment stress is usually assumed to be
uniformly distributed over the full timber thickness. This was validated with visual
inspection of failed specimens, where the shear failure plane occurred across the full
timber thickness (see Figure 5-a)). Good agreements were found between
predictions from Quenneville proposed equation for row shear strength and those
from validation tests for SWS connections (Quenneville 1998).
However, row shear-out predictions for WSW and WS connections calculated
over the full thickness of the wood side members were found to be high, compared to
the validations tests (the lower 5th
percentile). Further inspection of failed WSW and
WS connections revealed that row shear-out failure did not occur over the full
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thickness of the wood side member of the WSW connection, but over a reduced
thickness which was smaller than the wood member full thickness (tw), as can be
seen in Figure 5-b). Considering that similar type of fasteners (12.7 mm or 19.1 mm
bolts) were used in both types of connections and with more or less similar
thicknesses of wood members, the influence of material properties was
eliminated. This difference in the row shear failure pattern could be attributed to the
assumed embedment stress distribution along the fastener length. A different
embedment stress distribution clearly takes place in connections with a middle steel
plate, compared to those with steel side members.
Embedment stress distribution for rigid dowel type fasteners is assumed to be
uniform along the fastener length, Figure 6-a), and c). Figure 6-b) and d) shows the
assumed embedment stress distribution along the fastener for connections with non-
rigid dowel type fasteners and with steel side plates or middle steel plate. Unlike
connections with rigid type of fasteners, a uniform embedment stress distribution is
assumed only over a specified length y. Now, for connections with a middle steel
plate, this assumption results in crack propagation near the shear planes over a
length ye, which is assumed to be slightly bigger than y. This means that the shear
force (F) is acting over a reduced thickness which is less than the full thickness (tw) of
the wood side member (Jorissen 1998). Jorrisen refers to this specified thickness as
the effective thickness. The result is a lower than anticipated shear strength for the
connection with a middle steel plate. In an attempt to verify this theory, the thickness
across which row shear-out failure plane took place was measured for all groups that
exhibited row shear-out failure and the mean value for each group is presented in
Table 3, column 3 (except for those groups fabricated with 2 or 3 lumber members,
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where the thickness of a single member was considered as being the experimental
one). These values were found to be smaller than the full thickness of connections
wood members (column (2)). The mean ratio between the measured thickness and
full thickness of the wood member was found to be 0.65, Table 3.
In SWS connections with non-rigid fasteners, though the embedment stress
distribution is also assumed to be uniform over a length y, smaller than the member
full thickness, different stress distribution is associated with this type of connections,
see Figure 6-b). This distribution does not seem to influence the propagation of the
cracks near the shear planes. The development of infinite number of plastic hinges in
the fastener (unlike the case for WSW connections) leads to nearly uniform stress
distribution underneath the fastener. Failure usually occurs across the full thickness
for the type of fasteners used in the validation study of bolted SWS connections. This
may not necessarily be the case for connections with higher slenderness ratio, where
row shear failure is not dominant normally. This may lead to the conclusion that the
effective thickness theory could be applicable only for connections with a middle steel
plate. Adjustment to the wood side members thickness (tw) may be necessary to
account for that phenomenon in order to achieve better predictions for row shear-out
in WSW .
The following discussion describes how to determine the effective thickness
(ye) across which the row shear-out failure takes place for WSW and WS
connections, based on the stress distribution along the fasteners, for rigid and non-
rigid dowel type fasteners.
Effective Thickness
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ww ftdF =
e
s0.3C by =
Johansens Yield Model for double shear symmetrical connections with a
single internal steel plate and assuming a rigid dowel type fastener (see Figure 6-a)
and 6-c)) for failure Mode I, is given by the following equation:
[6]
For Mode II (Figure 6-d), the failure mode is given by the following equation:
[7]
Using Eq. [7], the length y can be determined as follows:
[8]
Additional stress analysis given by Jorissen (1998) indicated that for
connections with more slender dowel type fasteners, cracks propagate near the
shear planes over a thickness ye, which is assumed to be slightly bigger than y (see
Figure 6-d). For connections with rigid dowel type fasteners, it can be assumed that
y = ye = tw. According to Jorrisen, the value ye was determined using linear
interpolation where y < ye < tw and was given by the following equation:
[9]
where,
Cy is a constant (0 < Cy < 1.0), which was calculated based on the following
equation:
wfydF =
wfdF
y =
yt
ytC1y
w
wye
+=
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[10]
For connections with a single fastener, Cy was taken as equal to 1.0. Since
ye = tw for rigid dowel type of fasteners, Eq. [10] does not influence the calculation for
rigid dowel fasteners. Eq. [10] was derived empirically using test results (Jorissen
1998).
Measured values of the effective thickness (ye) across which the row shear-
out failure took place were found to be comparable with those calculated using Eq.
[9], except for connection made with lumber (WS), in which, due to the presence of
discontinuity, the row shear failure occurred at the interface between the two wood
laminates. From Table 3, the mean ratio of the measured effective thickness and that
of the member thickness was found to be 0.65. However, excluding the WS group
measurements brings the mean value up to 0.8, which corresponds well with the
computed effective thickness calculation (0.85).
It should be noted that the effective thickness approach should only be applied
to determine the strength of the row shear-out failure mode. For bearing failure
modes I, II and III, the member thickness must be used in design calculations. Group
tear-out calculations should also be based on the thickness of the wood members
since for group tear-out, failure must occur over the entire thickness. Partial tension
failure is unlikely to occur. Test observations for group tear-out failures confirm that
requirement.
To determine the appropriate row shear-out equation, predictions using
equation [1] were plotted against the 5th
% as determined from tests for the original
group. A linear regression analysis was carried out to determine the best fit line. A
modification factor of 0.8 was found to be appropriate for row shear-out predictions in
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( ) N)sMIN(e,f'Jt0.8n2np bvrwsruRS =
WSW and WS connections.
In order to account for the reduced thickness and to simplify the proposed RS
equation, a reduction factor could be introduced in the proposed row shear-out
equation [1] for WSW and WS connections. This leads to the following equation:
[11]
Using the calculated effective thickness (ye ) instead of the full thickness (tw ) in
the row shear-out equation (Eq. [11]), leads to lower predictions for row shear-out. In
Table 4, group tear-out and bearing predictions calculated based on the full thickness
are shown together with predictions based on the calculated effective thickness
(Column (6), Table 4). Row shear-out predictions based on the effective thickness,
provide a better agreement with the 5th
percentile from tests, where the row shear-out
failure controlled.
In Figure 7, the 5th % values determined experimentally are plotted against
model predictions including the modified row shear-out predictions. A better
agreement was found between test results and row shear-out predictions (see Table
4). This should not be understood as being only an empirically derived or a simple
curve fit exercise. It is based on the stress analysis described earlier and on the
laboratory observations.
Based on the above discussion, it is evident that introducing a factor of 0.8 in
the row shear-out proposed design equations for WSW and WS connections leads to
better predictions. Using the full thickness of the wood side members in WSW and
WS connections tend to overestimate the row shear strength.
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CONCLUSIONS
Based on the validation tests of the proposed design equations for WSW and
WS bolted connections, it can be concluded that:
1. Current Canadian design code (O86.1-94) leads to over-designed WSW and
WS bolted glulam connections, especially with multiple bolts, where it under-
estimates the failure loads.
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2. Proposed design equations for WSW bolted connections (Quenneville and Moham
predictions of the ultimate loads than current design procedure.
3. Improved predictions for row shear-out can be achieved if the effective thickness princ
factor of 0.8 was found to be suitable for row shear-out strength predictions of WSW an
ACKNOWLEDGMENT
Funding from the Academic Research Program, from the Military Engineering Resea
Royal Military College of Canada and from the Canadian Wood Council (CWC) is greatly app
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to express their gratitude to Mr. Lee and Ocdt Carriere who conducted the experimental tests.
REFERENCES
American Society for Testing and Materials (ASTM). 1988. Standard test methods for
mechanical fasteners in wood. Standard D1761-77, ASTM, Philadelphia, PA.
American Society for Testing and Materials (ASTM). 1994. Standard specification for
computing the reference resistance of wood-based materials and structural connections for
design. D5457-93, ASTM, Philadelphia, PA.
Canadian Standards Association. 1989. Engineering design in wood ( limit states
design). Standard O86.1-M89. Canadian Standard Association, Rexdale, ON.
Canadian Standards Association. 1994. Engineering design in wood ( limit states
design). Standard O86.1-94. Canadian Standard Association, Rexdale, ON.
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Canadian Wood Council. 1995. CSA Commentary, Wood Design Manual, Canadian
Wood Council, Ottawa, ON.
Johansen, K.W. 1949. Theory of timber connectors. IABSE Journal, No. 9. pp.249-
262.
Jorissen, A. 1998. Double shear timber connections with dowel type fasteners. Ph.D.
Thesis, Technical University of Delft, Delft, The Netherlands.
Larsen, H.. 1973. The yield load of bolted and nailed joints. Structural Research
Laboratory, Technical University of Denmark, IUFRO Division 5, p.14.
Mass, D.I., Salinas, J.J. and Turnbull J.E. 1988. Lateral strength and stiffness of
single and multiple bolts in glued-laminated timber loaded parallel to grain. Engineering
Centre, Research Branch, Agriculture Canada, Report No. C-029, Ottawa. ON.
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Mohammad, M., Smith, I. and Quenneville, J.H.P. 1997. Bolted timber connections:
investigations on failure mechanism. Proceedings of IUFRO S5.02 Timber Engineering
Denmark.
Quenneville, J.H.P. 1998. Predicting the failure modes and strength of brittle bolted
connections. Proceeding of the 5th World Conference on Timber Engineering (WCTE), Mo
144.
Quenneville, J.H.P. and Mohammad, M. 2000. On the failure modes and strength of
steel-wood-steel bolted timber connections loaded parallel-to-grain. Canadian Journal of Civ
Yasumura M., Murota T. and Nakai H. 1987. Ultimate properties of bolted joints in
glued-laminated timber. Report to the Working Commission W18-Timber Structures. Dublin
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d = bolt diameter, mm
e = end distance, mm
F = applied load resisted by one side of the wood member, N
ftg = specified strength in tension parallel to grain at the gross section, MPa
fv = specified strength in shear parallel to grain, MPa
fw = embedment strength of wood member, MPa
= 63G (1-0.01d), for parallel to grain loading
GT = group tear-out strength, N
Jr = factor for number of rows
= 1.0 for 1 row, or for 1 bolt per row
= 0.8 for 2 rows, (2 or more bolts in a row)
= 0.6 for 3 rows, (2 or more bolts in a row)
N = number of bolts in a row
nr = number of rows
ns = number of shear planes
RS = row shear-out strength, N
sb = bolt spacing in the row, mm
sr = row spacing, mm
tw = thickness of the wood side member, mm
y = thickness along which the embedment stress is assumed to be uniform, mm.
ye = effective thickness, mm.
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Table 1. Summary of specimens configurations.
GroupWood
1)
TypeSpecimensize
d nr N Type e sb sr MeanCOV 5th
%
. (mm) . . (mm) . (kN) (%) (kN)
1 S-P G2)
2@80x152 19.1 1 1 WSW 134 (7d) N/A N/A 57 8 35
2 S-P G 2@80x152 12.7 1 1 WSW 89 (7d) N/A N/A 46 16 24
3 S-P G 2@80x190 19.1 1 1 WSW 191 (10d) N/A N/A 68 15 37
4 S-P G 2@80x190 12.7 1 1 WSW 127 (10d) N/A N/A 56 13 30
5 S-P G 130x190 19.1 1 1 Insert 95 (5d) N/A N/A 42 8 26
6 S-P G 130x190 19.1 1 1 Insert 191 (10d) N/A N/A 65 8 39
7 S-P G 2@80x190 19.1 1 2 WSW 95 (5d) 95 (5d) N/A 103 9 61
8 S-P G 130x190 19.1 1 2 Insert 95 (5d) 95 (5d) N/A 95 6 61
9 S-P G 2@80x190 19.1 2 2 WSW 95 (5d) 95 (5d) 95 (5d) 181 9 107
10 S-P G 2@80x190 19.1 1 1 WSW 95 (5d) N/A N/A 48 8 29
11 S-P G 2@80x190 19.1 2 1 WSW 95 (5d) N/A 95 (5d) 106 6 69
12 S-P G 1@80x190 19.1 2 1 WS 95 (5d) N/A 95 (5d) 52 12 28
13 SPF L3)
2@38x140 12.7 1 1 WS 63 (5d) N/A N/A 14 14 8
14 SPF L 2@38x140 12.7 1 2 WS 64 (5d) 64 (5d) N/A 25 14 13
15 SPF L 2@38x140 19.1 1 1 WS 95 (5d) N/A N/A 29 12 15
16 SPF L 2@38x140 19.1 1 2 WS 95 (5d) 95 (5d) N/A 48 22 22
17 S-P G 130x190 19.1 2 2 Insert 95 (5d) 95 (5d) 95 (5d) 175 8 106
18 D-fir G 130x190 19.1 2 2 Insert 95 (5d) 95 (5d) 95 (5d) 181 13 96
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19 D-fir G 2@80x190 19.1 2 2 WSW 95 (5d) 95 (5d) 95 (5d) 225 13 120
20 S-P G 2@80x190 19.1 2 2 WSW 191 (10d) 191 (10d) 95 (5d) 243 11 141
21 S-P G 130x190 19.1 2 1 Insert 95 (5d) N/A 95 (5d) 88 11 49
22 S-P G 130x190 19.1 2 2 Insert 191 (10d) 191 (10d) 95 (5d) 227 6 147
23 SPF L 2@38x140 12.7 1 2 WS 127 (10d) 127 (10d) N/A 24 16 12
24 SPF L 2@38x140 19.1 1 2 WS 191 (10d) 191 (10d) N/A 47 23 20
25 SPF L 3@38x140 12.7 1 2 WS 64 (5d) 64 (5d) N/A 16 18 8
26 SPF L 3@38x140 19.1 1 2 WS 96 (5d) 96 (5d) N/A 26 19 11
27 S-P G 130x190 12.7 2 2 Insert 127 (10d) 127 (10d) 64 (5d) 115 12 64
28 D-fir G 130x190 12.7 2 2 Insert 127 (10d) 127 (10d) 64 (5d) 140 6 89
29 S-P G 2@80x190 12.7 2 2 WSW 127 (10d) 127 (10d) 64 (5d) 142 11 80
30 D-fir G 2@80x190 12.7 2 2 WSW 127 (10d) 127 (10d) 64 (5d) 143 5 931)
Both S-P and D-fir glulam were from 20f-EX grade and SPF lumber was No. 2 and better Structurallight Framing.
2)Glulam
3)Lumber
Table 2. Validation tests results and predictions using O86.1-94 and proposed equations (Que
Proposed Design Equations
3)
Group
(1)
5th
%1)
test
(2)
O86.1 94
(3)
Ratio(4)/(2)
(4)puRS PuGT PuB1 puB2 puB3 puB4 puB5
pu Min.Observedfailuremode
.. (kN) .. .. (kN) ..
1 35 23 0.65 71 71 55 30 39 -- -- 30 B/RS
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2 24 13 0.53 49 49 39 17 18 -- -- 17 B/RS
3 37 30 0.82 94 94 55 30 39 -- -- 30 B
4 30 17 0.56 68 68 39 17 18 -- -- 17 B
5 26 21 0.80 38 38 41 28 39 -- -- 28 B/RS
6 39 28 0.71 65 65 41 28 39 -- -- 28 B/RS
7 61 02)
0 94 94 110 61 78 -- -- 61 RS8 61 0 0 65 65 82 56 78 -- -- 56 RS
9 107 0 0 150 244 219 122 156 -- -- 122 RS/S
10 29 0 0 52 52 55 30 39 -- -- 30 RS
11 69 0 0 105 203 110 61 78 -- -- 61 RS
12 28 0 0 52 101 55 31 57 49 39 31 RS
13 8 0 0 10 10 18 10 17 16 9 9 B
14 13 0 0 18 18 36 16 34 32 18 16 RS
15 15 0 0 14 14 25 15 28 24 19 14 RS/B
16 22 0 0 25 25 50 29 57 48 38 25 RS
17 106 0 0 105 179 164 111 156 -- -- 105 RS
18 96 0 0 119 188 183 119 165 -- -- 119 RS19 120 0 0 173 259 244 131 165 -- -- 131 RS
20 141 125 0.89 232 296 219 122 156 -- -- 122 B
21 49 0 0 76 151 82 56 78 -- -- 56 RS
22 147 115 0.78 140 201 164 111 156 -- -- 111 RS/B
23 12 16 1.33 31 31 36 16 34 32 18 16 RS/B
24 20 30 1.5 38 38 50 29 57 48 38 29 RS/B
25 8 0 0 29 29 54 20 34 36 18 18 RS/B
26 11 0 0 41 41 75 34 57 52 38 34 RS/B
27 64 57 0.89 124 171 118 59 72 -- -- 59 B
28 89 60 0.67 141 200 131 64 75 -- -- 64 B
29 80 57 0.72 186 234 157 67 72 -- -- 67 B
30 93 60 0.65 212 274 175 73 75 -- -- 73 B
1)
Adjusted for the normal duration of loading by dividing over a factor of 1.25.
2)
No values are given for groups with an end distance below the 7d limit (O86.1-94).
3)
From Quenneville (1998).
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Table 3. Comparison between measured and calculated
effective thicknesses.
Group
Woodmemberthickness
MeasuredeffectiveThickness
Ratio(3) / (2)
EffectivethicknessEq. [9]
Ratio(5) / (2)
(1) (2) (3) (4) (5) (6)
.. (mm) .. (mm)
1 80 --1)
-- 68 0.85
2 80 -- -- 80 1.00
3 80 -- -- 68 0.85
4 80 -- -- 80 1.00
5 60 48 0.8 58 0.97
6 60 57 0.9 58 0.97
7 80 49 0.6 55 0.69
8 60 42 0.8 52 0.87
9 80 58 0.7 55 0.69
10 80 62 0.8 68 0.85
11 80 75 0.9 68 0.85
12 80 41 0.5 76 0.95
13 76 38 0.5 67 0.88
14 76 38 0.5 54 0.71
15 76 38 0.5 76 1.00
16 76 38 0.5 71 0.93
17 60 54 0.9 52 0.87
18 60 54 0.9 48 0.80
19 80 56 0.7 50 0.63
20 80 -- -- 55 0.69
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21 60 54 0.9 58 0.97
22 60 -- -- 52 0.87
23 76 38 0.5 73 0.96
24 76 38 0.5 57 0.75
25 114 38 0.33 80 0.70
26 114 38 0.33 60 0.5327 60 -- -- 60 1.00
28 60 -- -- 60 1.00
29 80 -- -- 72 0.90
30 80 -- -- 66 0.83
Mean 0.65 0.85
STD 0.19 0.131)
information not available.
Table 4. Test results and predictions using O86.1-94 andproposed design equations.
Group5th %test
O86.1-94 PuGT PuB puRS2)
Mod.
... (kN) ...
1 35 23 71 30 56
2 24 13 49 17 39
3 37 30 94 30 75 3)
4 30 17 68 17 543)
5 26 21 38 28 30
6 39 28 65 28 52
7 61 01)
94 61 75
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8 61 0 65 56 52
9 107 0 244 122 120
10 29 0 52 30 42
11 69 0 203 61 84
12 28 0 101 31 42
13 8 0 10 9 814 13 0 18 16 14
15 15 0 14 14 11
16 22 0 25 25 20
17 106 0 179 105 84
18 96 0 188 119 95
19 120 0 259 131 138
20 141 125 296 122 1863)
21 49 0 151 56 61
22 147 115 201 111 112
23 12 16 31 16 253)
24 20 30 38 29 30 3)
25 8 0 29 18 233)
26 11 0 41 34 33
27 64 57 171 59 993)
28 89 60 200 64 1133)
29 80 57 234 67 1493)
30 93 60 274 73 1703)
1)O86.1 has no provision for an end distance less than 7d.
2)RS values multiplied by a factor of 0.8.
3)Bearing failure governs.
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LIST OF FIGURE CAPTIONS
Fig. 1. Typical specimen in testing apparatus.
Fig. 2. Typical load-slip envelopes that exhibit ductile and brittle behaviour.
Fig. 3. Comparison between test results and O86.1-94 or proposed equations prediction
Fig. 4. Comparison between test results and row shear-out predictions using proposed d
Fig. 5. Comparison between row shear-out failure in bolted connections: a) SWS; b) WS
Fig. 6. Comparison between embedment stress distribution in SWS and WSW or WS co
Fig. 7. Comparison between test results and O86.1-94 or modified row shear-out predict
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