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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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    3

    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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    6

    ( ) )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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    7

    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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    8

    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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    9

    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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    10

    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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    11

    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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    12

    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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    13

    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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    14

    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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    15

    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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    16

    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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    17

    [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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    18

    ( ) 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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    19

    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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