Timber Rivet 

Connections in 

U.S. Domestic Species 

Marshall Begel 

Ronald W. Wolfe 

Douglas Stahl

United States

Department of 

Agriculture 

Forest Service 

Forest 

Products 

Laboratory 

Research 

Paper 

FPL−RP−619 
Abstract 

This paper discusses results of tests conducted to expand the 

data base on the performance of timber rivet connections in 

U.S. domestic species to verify existing and proposed design 

procedures. Eight-, 80-, and 200-rivet connections were 

tested. The 8-rivet joint tests illustrate the effects of material 

type, rivet length, and load direction on the behavior of rivet 

connections when gross wood failures are avoided. The 

80-rivet connections, loaded perpendicular to grain, show 

the effect of rivet spacing as well as test support conditions 

on failure mode and strength. The 200-rivet joints, with load 

applied parallel to grain, provide data for previously reported 

failure modes, including the rivet yield/wood crush modes 

modeled by European yield model (EYM) and gross wood 

failures. The 200-rivet tests also show the effect of rivet 

spacing on joint strength and failure mode. 

The tests were conducted in two phases. In Phase I, small 

specimens were made from Southern Pine and ponderosa 

pine solid sawn lumber, HemñFir and Southern Pine glulam, 

and parallel strand lumber. The small specimens failed by a 

combination of rivet yielding and wood crushing as described by EYM, rather than by a gross failure of the wood 

block. Joint yield load was determined by fitting test data to 

a curve and finding the load at 5% offset displacement. 

Basic trends can be seen in the data, but yield load is a poor 

predictor of ultimate load. 

April 2004 

Begel, Marshall; Wolfe, Ronald W.; Stahl, Douglas C. 2004. Timber rivet 

connections in US domestic species. Res. Pap. FPL-RP-619. Madison, WI: 

U.S. Department of Agriculture, Forest Service, Forest Products 

Laboratory. 15 p. 

A limited number of free copies of this publication are available to the 

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In Phase II, large specimens were made from Southern Pine 

glulam. Larger rivet spacing, wider specimens, and longer 

rivets generally increased joint capacity; end spacing had no 

significant effect. Short continuously supported beams had a 

higher capacity load than did longer simply supported 

beams. The results show that when wood failure modes 

occur, thicker glulam has higher capacity, contradicting the 

size effect of the design code. 

Keywords: glulam rivet, timber rivet, connection, 

design model

Contents 

 Page 

Introduction........................................................................... 1

Objectives and Scope............................................................ 1

Literature Review.............................................................. 2

Definition of Terms........................................................... 2

Procedures............................................................................. 3

Phase I. Small Connection Tests....................................... 3

Phase II. Large Connection Tests ..................................... 4

Results................................................................................... 7

Phase I. Small Connection Tests....................................... 7

Phase II. Large Connection Tests ..................................... 8

Analysis of Results ............................................................. 10

Phase I. Small Connection Tests..................................... 10

Phase II. Large Connection Tests ................................... 12

Conclusions......................................................................... 12

Literature Cited ................................................................... 13

AppendixóResults of Large Connection Tests.................. 13
Timber Rivet Connections in 

U.S. Domestic Species 

Marshall Begel, General Engineer 

Ronald W. Wolfe, Research General Engineer 

Forest Products Laboratory, Madison, Wisconsin 

Douglas C. Stahl, Associate Professor

Architectural Engineering and Building Construction 

Milwaukee School of Engineering 

Milwaukee, Wisconsin 

Introduction 

The timber rivet connection has become common in Canadian glued-laminated (glulam) and sawn timber construction 

as a cost-effective alternative to bolted connections. Timber 

rivets are hardened steel nails from 40 to 90 mm long and 

3.2 by 6.4 mm in ovalñrectangular cross section. They are 

driven with their major cross-sectional dimension parallel to 

the wood grain. Timber rivets are installed through predrilled steel side plates (typically 6.4 mm thick) in rectangular arrays with minimum spacing of 15 to 25 mm. The rivetpattern density (rivets/unit area) has a profound effect on the 

failure mode. Close spacing promotes a brash failure characterized by wood tear-out within the area defined by the rivet 

array. Wider spacing tends to cause a combined fastener 

yield and wood crushing mechanism to dominate, resulting 

in greater ductility. The latter failure mode is common to 

other dowel-type fasteners and is described by the European 

yield model (Aune and PattonñMallory 1986). 

Although timber rivets have been used in Canada for more 

than 30 years (Madsen 2000), they were first adopted in the 

United States through the 1997 edition of the National Design Specification for Wood Construction (NDS). Provisions 

in the NDS (AF&PA 2001) are based on, but less comprehensive than, the Canadian O86.1 code (CSA 2001). The 

NDS, for example, limits the use of rivets to only Southern 

Pine and Douglas Fir glulam, and it does not include the 

general connection design procedure for non-standard connection geometries that is included in the Canadian code as 

an appendix, which effectively limits applications for designers in the United States. Additional shortcomings of the 

Canadian and U.S. code design procedure are as follows: 

1.The analysis of stresses around the fastener group, which 

leads to prediction of connection strength when wood 

failure controls, is well reasoned but too complex for a 

closed form solution; designers must refer to tables of 

constants. 

2.The procedure includes a volume effect on shear strength, 

so predicted connection strength decreases as timber 

thickness increases. While the concept of a size effect is 

accepted in several areas of timber design, there are insufficient test data in the literature to support its inclusion in 

this analysis. 

3.The procedure uses the capacity of a single rivet as an 

input, but it does not provide guidance on how to estimate 

this capacity for wood species not currently included in 

the code. 

Objectives and Scope 

This paper discusses the results of tests conducted to expand 

the data base on performance of timber rivet connections in 

U.S. domestic species. The primary objective was to provide 

a range of data for verifying existing and proposed alternative design procedures. Current design procedures consider 

failure mode as well as material strength in predictions of 

joint capacity. The tests were therefore required to exhibit 

failure modes dominated by wood fracture as well those 

dominated by rivet yield and localized wood crushing. 

Test joints included a range of rivet sizes (40- to 90-mm 

lengths) and material specific gravity (0.4 to 0.55) and 

strength in U.S. domestic species. Joint tests included 

8-, 80-, and 200-rivet connections. The 8-rivet joint tests 

illustrate the effects of material type, rivet length, and load 

direction on the behavior of rivet connections when gross 

wood failures are avoided. The 80-rivet connections, loaded 

perpendicular to grain, show the effect of rivet spacing 

and test support conditions on failure mode and strength. 

The 200-rivet joints, with load applied parallel to grain, 

provide data to compare the effects of wood tear-out failure 

modes on rivet yield. 
2 

Literature Review 

The first test data were reported by McGowan (1966) and 

the development of the timber rivet was chronicled by 

Madsen (2000), but without doubt the most influential work 

on timber rivets is that of Foci and Long worth (1975). Foci 

and Long worth presented rational analyses of the conditions 

leading to a rivet yield failure mode and to a wood failure 

mode; they included limited testing to verify and calibrate 

the analyses. These authors presented a number of equations 

to predict the strength of timber rivet joints based on their 

geometry and material properties. The correlation of predictions and test results was remarkably good. The authors 

concluded that the wood tension failure mode should rarely, 

if ever, control and then only for unusually thin wood members. They also presented analogous developments for perpendicular-to-grain loadings. Their limited experimental 

work served to show that the analysis gives good predictions. The work by Foci and Long worth has proven to be 

quite robustóthere have been no published alternatives to 

their basic analysis of concentrically loaded connections, and 

their procedure has been adopted by the Canadian and 

U.S. codes. 

The first study of timber rivet behavior in solid timber presented promising results with an important caveat. In testing 

timber rivet connections parallel and perpendicular to grain, 

Karacabeyli and Fraser (1990) reported that the strength, 

loadñdisplacement response, and failure modes of these 

connections in Douglas Fir solid timber and glulam were 

essentially the same. They noted, however, that the absence 

of major defects such as checks and splits from the solid 

timber specimens limited the value of their results as an 

indication of how rivet connections would behave in real 

solid timber. 

Karacabeyli and others (1998) described a large testing 

program to evaluate the behavior of rivet connections in 

solid timber (Douglas FirñLarch, SpruceñPineñFir, and 

HemñFir). Their solid timber specimens, like those in the 

study by Karacabeyli and Fraser 1990), were defect-free in 

accordance with testing norms (ASTM 1988). They noted 

that when rivet yielding is the failure mode, connections in 

solid timber are 80% to 90% as strong as connections in the 

same species of glulam. For connections with wood failure 

modes, the authors noted that these failure modes are similar 

for solid timber and glulam. They reasoned that the work of 

Foci and Long worth (1975) should therefore apply to 

solid timber. 

As previously noted, the equations developed by Foci and 

Long worth (1975) use wood tensile and shear strength to 

determine strength of the riveted connection in wood failure 

modes. Shear strength values for solid timber are typically 

40% to 50% of shear strength values for glulam of the same 

species group, as a result of cracks and checks in solid timber. Therefore, Karacabeyli and others (1998) concluded that 

riveted connections with solid timber should have a 50% 

reduction in design strength from connections with glulam of 

the same species group. The adjustments they suggested 

were adopted by the Canadian code. 

Buchanan and Lai (1994) tested timber rivet connections in 

Monterey pine (Pinus radiata) and investigated application 

of the European yield model (EYM) for rivet connections. 

The EYM describes failure as a combination of fastener 

yield and localized wood crushing (Aune and Pattonñ 

Mallory 1986), and it has been accepted for use with all 

other dowel fasteners (AF&PA 1999, 2001). Buchanan and 

Lai (1994) showed that EYM predictions were quite good, 

but their work did not lead to changes in code. 

Definition of Terms 

The following terms describe the rivet connection configuration (see also Fig. 1): 

Figure 1óConnection geometry variables: (from left) parallel-to-grain connection, perpendicularto-grain connection, and view through thickness of connection. 
3

Row line of rivets parallel to load direction 

Column line of rivets perpendicular to load direction 

sp spacing of rivets measured parallel to grain 

sq spacing of rivets measured perpendicular to grain 

L rivet length (standard lengths used in this study 

were 40, 65, and 90 mm) 

p rivet penetration into wood (p = L ñ plate thickness 

ñ 3.2 mm) 

a distance from end of wood member to closest 

column of rivets for parallel-to-grain loading 

eq distance from loaded edge of wood member to 

closest column of rivets for perpendicular-to-grain 

loading 

ep distance from rivet to non-loaded edge for parallelto-grain loading 

es distance from edge of steel plate to closest rivet 

Procedures 

Testing for this research was conducted in two phases. 

Phase I was intended to illustrate the effects of material type, 

rivet length, and load direction on the behavior of rivet 

connections when gross wood failures are avoided. Phase II 

was planned to provide a full range of previously reported 

failure modesóincluding gross wood failures and rivet 

yield/wood crush modes modeled by EYM. 

Phase I consisted of 130 tests of small connections. The 

specimen geometry and test procedures were consistent with 

those used by previous researchers (Buchanan and Lai 1994, 

Karacabeyli and others 1998), so these data could be fairly 

compared with data in the literature. 

Phase II included 45 tests of large connections in 22 configurations. These tests were intended to illustrate the effects 

of connection configuration, load direction, and support 

conditions on failure mode and capacity. The large connection test specimens were similar to those used by previous 

researchers (Buchanan and Lai 1994, Foci and Long worth 

1975, Karacabeyli and others 1998, McGowan 1966), but 

our specimens had riveted plates on both faces of the glulam, 

making a symmetric connection that better simulates real 

applications. 

Wood for all test specimens was preconditioned at 18°C and 

65% relative humidity to achieve a target equilibrium moisture content of 12% prior to fabrication of the connection. 

After fabrication, the specimens were stored under the same 

conditions for at least 1 week until testing. Test joints were 

fabricated with plates applied to opposing surfaces of the 

wood members. All side plates, as called for by NDS specifications, were A36 steel with predrilled 7.14-mm holes. 

Rivets were driven either by hand or by a hand-held pneumatic hammer. 

Phase I. Small Connection Tests 

Small specimens (Fig. 2) were designed to fail by a combination of rivet yielding and wood crushing, as described by 

EYM, rather than by gross failure of the wood block. All 

specimens had 6.4-mm steel plates on opposite sides, with 

four rivets installed through predrilled holes in each plate. 

Rivet spacing was 25.4 mm parallel and perpendicular to 

grain. The specimens were supported at the bottom of the 

Figure 2óSmall specimens for (a) perpendicular-to-grain and (b) parallel-to-grain loadings. 
4 

steel plates and loaded by pushing down on the top. We 

assumed that if gross wood failure is avoided there is no 

difference between compressive and tensile loading; either 

one is really shear or lateral loading on the fastener groups. 

All specimens were cut from solid or glulam timber with 

cross-sectional dimensions of 140 mm wide (along glue 

lines) and 184 mm deep (perpendicular to glue planes). 

Specimens with 90-mm rivets were fabricated with the 

140-mm and 184-mm dimensions switched so that the tips of 

the long rivets would not overlap at the center. Parallel-tograin specimens were 152 mm long; perpendicular-to-grain 

specimens were 203 mm long. Specimens were made with 

solid sawn Southern Pine, solid sawn Ponderosa Pine, 

glulam Southern Pine, glulam HemñFir, and parallel strand 

lumber (PSL), a structural composite lumber product available in sizes large enough for timber rivet use. Table 1 

shows the wood material, rivet length, and loading direction 

for the specimens. Ten replicates of each specimen 

were used. 

All wood material was conditioned and cut at the Forest 

Products Laboratory and delivered to the Milwaukee School 

of Engineering for connection fabrication, final conditioning, and testing. The glulam was fabricated commercially. 

All tests were conducted on an open-loop hydraulic test 

machine with manual control over the loading rate. Load 

was measured with a 90-kN capacity electronic load cell; 

deformation of the connection was measured with a pair of 

symmetrically mounted 12.7-mm linear variable differential 

transformers (LVDTs) (Fig. 3). The LVDTs were mounted 

by clamping the LVDT body into a fitting screwed to the 

wood specimen and clamping the core extension rod into a 

plexiglass fitting that was hot-glued to the steel plate. 

The two displacement readings were averaged to produce a 

single measure of connection deformation. The rate of connection deformation was monitored during the test, and the 

loading rate was adjusted manually as necessary to keep the 

deformation rate close to the 2-mm/min target. For all tests, 

deviation from the target deformation rate was insignificant. 

Readings from the load cell and LVDTs were recorded at 1-s 

intervals. After testing, specimens were split apart with a 

chisel to observe the condition of the rivets. 

Phase II. Large Connection Tests 

The large rivet test connections were fabricated using Southern Pine glulam. Twenty-two joint configurations were 

tested; 14 configurations were tested with load applied parallel to grain and 8 with load applied perpendicular to grain. 

Joint configuration variables, summarized in Tables 2 and 3, 

included rivet length (L), spacing parallel (sp) and perpendicular (sq) to grain, end distance for parallel-to-grain (a) and 

perpendicular-to-grain (eq) loading, and glulam thickness. 

For parallel-to-grain tests, most glulam specimens were 

2.0 m long by 419 mm wide. The first four specimens tested 

(9ñ1, 9ñ2, 10ñ2, 10ñ2) were necked down from a width of 

381 mm at the grip to a width of 279 mm at the rivet connection in an effort to reduce edge distance (ep). Glulam 

Table 1óSpecimens for small connection tests 

Wood material 

Loading direction with respect 

to wood grain 

Rivet 

length (L) 

(mm) 

Parallel 40 

 65 

 90 

Perpendicular 40 

 65 

Solid-sawn Southern Pine 

 90 

Solid-sawn ponderosa pine Parallel 65 

Perpendicular 65 

Glulam HemñFir Parallel 65 

Perpendicular 65 

Glulam Southern Pine Parallel 65 

Perpendicular 65 

Parallel strand lumber Parallel 65 

Perpendicular 65 Figure 3óSetup for small connection tests. 
5

Table 2óConfigurations for large parallel-to-grain specimens (mm) 

Rivet pattern Glulam Steel side plates Joint 

IDa 

sp sq L a Thickness Width Width Length Thickness 

P3 25.4 12.7 40 50 79 419 229 508 6.4 

P4 150 79 419 229 610 6.4 

P5 31.8 25.4 40 50 130 419 254 584 6.4 

P6 150 130 419 254 686 6.4 

P7 25.4 12.7 40 50 130 419 229 508 6.4 

P8 150 130 419 229 610 6.4 

P9 31.8 25.4 65 50 130 279 254 584 9.5 

P10 150 130 279 254 686 9.5 

P11 25.4 12.7 65 50 130 419 229 508 6.4 

P12 150 130 419 229 610 6.4 

P13 31.8 25.4 90 50 216 419 254 584 9.5 

P14 150 216 419 254 686 9.5 

P15 25.4 12.7 90 50 216 419 229 508 6.4 

P16 150 216 419 229 610 6.4 

a

 Study plan included two joint configurations (joints P1 and P2) in 79-mm-thick glulam using 

 40-mm rivets spaced 32 mm parallel and 25 mm perpendicular to the grain. Although these 

 tests were not conducted, the original numbering was retained to maintain compatibility 

 between the final report, the raw data, and the study plan. 

Table 3óConfigurations for large perpendicular-to-grain specimens 

Rivet patterna

 (mm) Glulam Steel side plates (mm) 

Joint 

ID sp sq p eq

Thickness 

(mm) 

Length 

(m) Width Length 

Thickness 

Q1 25.4 25.4 65 102 130 2.3 180 276 6.4 

Q2 73 130 0.8 180 324 6.4 

Q3 25.4 12.7 65 73 130 2.3 143 324 6.4 

Q4 127 130 0.8 143 324 6.4 

Q5 25.4 25.4 90 73 216 2.3 180 381 6.4 

Q6 73 216 0.8 180 381 6.4 

Q7 25.4 12.7 90 102 216 2.3 143 324 6.4 

Q8 127 216 0.8 143 324 6.4 

a

See text for definition of terms. 
6 

specimens for the perpendicular-to-grain tests were fabricated with a width of 381 mm and length of 3 m. Each of 

these 3-m lengths was cut to provide two test samplesóone 

2.3 m long and the other 0.8 m longóto provide a basis for 

evaluating the effect of continuous support as opposed to 

simple span bending support. 

Rivet patterns used for the test joints are summarized in 

Tables 2 and 3. For most test connections, the steel side 

plates were 6.4 mm thick; 9.5-mm-thick plates were used for 

parallel-to-grain test configurations predicted to develop 

ultimate loads higher than 500 kN. Rivet holes were drilled 

through the steel plates in rectangular 10×10 arrays for 

parallel-to-grain tests and 4×10 arrays for perpendicular-tograin tests. For parallel-to-grain specimens, row spacing was 

12.7 or 25.4 mm and column spacing was 25.4 or 32 mm; 

for perpendicular-to-grain specimens, row spacing was 

25.4 mm and column spacing was 12.7 or 25.4 mm. Steel 

plates were sized to accommodate rivet patterns with a minimum edge distance of 12.7 mm. 

Application of load parallel to grain required a grip to transmit large forces from the test machine to one end of the 

glulam specimen (Fig. 4). The grip consisted of two 9.5-mm 

steel plates 0.4 m wide by 1.1 m long, each containing a 

3×4 matrix of 20.6-mm-diameter holes and a single row of 

three 23.8-mm-diameter holes. The plates were attached to 

the wood using 19-mm-diameter bolts with 102-mmdiameter shear plates inserted into seats drilled in the surface 

of the wood. The number of bolts varied with the expected 

capacity of the rivet connection. The grip and rivet plates 

were bolted to steel load blocks that were in turn attached to 

the test machine heads. The load blocks were 229 wide by 

102 mm deep. 

Two sets of blocks were used: one set was the thickness of 

the thickest specimen (216 mm) and the other was the thickness of the intermediate specimen (130 mm). The thinner 

block was also used for the thinnest (79-mm) glulam specimens. The grip and rivet plates were held against the load 

blocks by six 22.2-mm bolts. A 54-mm hole was drilled in 

each load block to accommodate a high-strength threaded 

rod, which connected the block to the load head of the tension machine. The rod passed through the hole and through a 

spherical washer, which could swivel to reduce any moment 

resulting from test eccentricities. Displacement was measured between the steel rivet plate and the glulam at the centerline of the rivet pattern (Fig. 5). The LVDT was attached 

to the steel plate at the geometric center of the rivet pattern 

and aligned with the load direction. Reference bars, oriented 

perpendicular to the LVDT, were held above the surface of 

the plate and fastened to the glulam beam at the exposed 

edges just beyond the plate edges. 

Figure 4óSetup for tension parallel-to-grain test.

Figure 5óMeasurement of 

displacement parallel to grain. 
7

For the tests conducted perpendicular to grain, two boundary 

conditions were compared (Fig. 6). In one case, the connections were tested using a continuous bearing support on the 

bottom edge of a 0.76-m-long specimen; in the other case, 

the connection was placed at the center of a 2.1-m beam 

span. In both cases, load was applied to edges of the plates 

extending above the top edge of the timber. 

The load head was grooved to fit over the plate edge and 

provide lateral restraint while applying a vertical shear load. 

Displacement was measured as the movement of the load 

head relative to the top of the test specimen, using one 

LVDT on each side of the load head (Fig. 7). 

For both parallel- and perpendicular-to-grain loadings, force 

was applied using a screw-type universal testing machine. 

Load and displacement channels were recorded three times 

per second. The test was stopped when load dropped off by 

10% of the maximum value. For most loadings, loads were 

applied at a displacement rate of 2 mm/min; for parallel-tograin loading beyond a load of 89 kN, the load rate was 

slowed to 1 mm/min. 

As noted previously, placing rivet plates on both faces of a 

wood specimen eliminates a potentially troublesome source 

of eccentricity that does not exist in most real applications. 

To make the test connections behave even more like real 

connections, out-of-plane movement of the steel plates was 

restricted by bolting the plates to the load heads (for parallelto-grain loading) or by grooves on the loading head (for 

perpendicular-to-grain loading). This simulates the restraining effect, in real applications, provided by the wood to 

which the other end of the steel plate is attached. 

Results 

For each specimen, load and displacement measurements 

were recorded at specific intervals until a time well past the 

point of joint failure. These measurements can be plotted as 

data points to form a loadñdeflection (Pñδ) curve (see 

Figs. 8 and 9). The first 5% to 10% of the graph (measured 

by load value) is not considered an accurate characterization 

of joint behavior for two reasons: (1) variances in joint 

construction require a small amount of load to distribute 

forces among the fasteners, and (2) measuring equipment is 

not considered accurate in its low ranges. The slope of the 

curve generally decreases steadily until one of two failure 

mechanisms takes effect, and the shape of the curve will 

indicate the general type of failure experienced by the joint. 

If the curve levels off and load decreases in value smoothly 

with increased deflection, ductile yielding of the rivets and 

bearing failure of the wood has taken place in the joint. If the 

load value drops in one or more steps as deflection increases, 

a brash and sometimes sudden wood failure has occurred. 

Phase I. Small Connection Tests 

All specimens failed by a combination of rivet bending and 

localized wood crushing. Representative sets of loadñ

displacement plots are shown in Figures 8 and 9. 

Figures 8 and 9 show parallel- and perpendicular-to-grain 

loadings, respectively, for solid Southern Pine with 65-mm 

 Figure 6óSpecimens for perpendicular-to-grain tests. 

Figure 7óTest of rivets loaded 

perpendicular to grain. 
8 

rivets. Two observations are noteworthy. First, the peak 

loads are similar. Second, the differences in curve for parallel loading (simple curve from initial slope to zero final 

slope, Fig. 8) compared with a perpendicular loading (complex curve with intermediate straight region between about 

1.5- and 4.5-mm displacement, Fig. 9) are typical of all 

materials and rivet lengths tested. 

Results for all small connection specimens are listed in 

Table 4. Data include the observed load when the test deflection reached 4.8 mm, which can be compared to results from 

previous studies by Buchannan and Karacabeyli, and the 

maximum load achieved during the test. The dominant yield 

mode was determined by inspection of the rivets after testing. Figures 10 and 11 illustrate typical yield modes. 

The only unusual occurrence with regard to the glulam was 

that 6 of the 10 Southern Pine specimens for parallel-tograin loading split severely while the rivets were being 

installed. This set of 10 specimens was discarded. This was 

the last batch of specimens to be tested, and it was the only 

set of specimens to exhibit this kind of behavior. We have 

no definitive explanation, but we suspect that the conditioning chamber was not properly monitored and the specimens 

were inadvertently over-dried prior to fabrication. 

The only yield modes observed were modes IIIm (single 

curvature) and IV (double curvature) (nomenclature from 

AF&PA 1999, 2001), both of which include a straight portion at the rivet head where rotation of the rivet is restrained 

by the fixity of the tapered head in the undersized hole in the 

steel plate. All specimens with yield mode data in Table 4 

were split to inspect the rivets, and at least 25 of the 32 rivets 

had the designated yield pattern. 

Phase II. Large Connection Tests 

Peak loads parallel to grain ranged from 228 kN for a tightly 

spaced array of 40-mm rivets to 1,010 kN for wider spaced 

90-mm rivets. Peak loads perpendicular to grain ranged from 

175 kN for tightly spaced 65-mm rivets to 375 kN for wider 

spaced 90-mm rivets. Test results for large connections 

loaded parallel- and perpendicular-to-grain are summarized 

in Tables 5 and 6, respectively; more detailed information 

about each replicate is presented in the Appendix. The format of the data in Tables 5 and 6 matches that of Table 4 

with the exception of failure mode description. For the large 

connection tests (Tables 5 and 6), the codes used to classify 

failure modes are R for rivet bending and localized wood 

crushing, and W for gross wood failure. The effect of failure 

mode on the loadñdisplacement plots is shown in Figure 12. 

While some failures were easily classified using this simple 

code (Fig. 13, for example), many of the observed failures 

exhibited a combination of failure modes. For perpendicularto-grain tests, splits along the grain appeared in conjunction 

with rivet bending and wood bearing deformation. These 

failures are considered ìwoodî failures rather than ìrivetî 

failures. For parallel-to-grain tests, it was difficult to tell 

which component failed first in a number of hybrid failures. 

Wood failure was observed more commonly in the specimens that had closely spaced rivet patterns than in those with 

widely spaced patterns. Perpendicular-to-grain specimens 

with closely spaced rivets tended to have a single large crack 

extending along the grain from the top rivet column (column 

farthest from loaded edge), whereas specimens with wider 

spacing were more likely to have cracks along nearly every 

column (Fig. 14). 

As noted previously, four parallel-to-grain specimens were 

necked down to minimize the horizontal edge distance. This 

created stress concentrations that appear to have caused one 

Figure 8óTypical results for small specimens loaded 

parallel to grain. Ten replicates of solid-sawn Southern 

Pine, 65-mm rivets. Fitted curve and 5% offset line 

used for analysis. 

Figure 9óTypical results for small specimens loaded 

perpendicular to grain. Ten replicates of solid-sawn 

Southern Pine, 65-mm rivets. Fitted curve and 5% 

offset line used for analysis. 
9

Table 4óSmall connection test resultsa

Woodb Loading direction 

L

(mm) 

Load at 

5% 

offsetc

(kN/rivet) 

Initial 

sloped

(kN/mm/ 

rivet) 

Max. load 

up to 

4.8 mm 

(kN/rivet) 

Maximum 

load 

(kN/rivet) 

Dominant 

yield mode 

SSP Parallel 40 0.97 5.67 3.37 (4) 3.52 (4) IIIm

 65 1.81 6.99 3.75 (13) 3.97 (13) IIIm 

 90 2.02 5.61 4.75 (10) 5.17 (10) IV 

 Perpendicular 40 0.69 2.61 2.50 (8) 2.93 (7) IIIm

 65 1.05 4.12 3.88 (11) 4.38 (10) IV 

 90 0.76 4.58 4.25 (10) 5.25 (9) IV 

SPP Parallel 65 1.68 10.4 3.62 (7) 3.74 (9) IIIm

 Perpendicular 65 1.02 5.17 3.50 (10) 4.64 (8) IV 

GHF Parallel 65 2.07 6.75 5.09 (5) 5.90 (9) ód 

 Perpendicular 65 1.47 3.15 4.04 (14) 4.93 (12) ód

GSP Perpendicular 65 0.63 5.02 4.13 (8) 4.94 (7) ód

PSL Parallel 65 2.10 7.01 3.83 (4) 3.92 (5) ód

 Perpendicular 65 2.19 2.45 5.14 (4) 5.55 (5) ód

a

Average for 10 replicates. Values in parentheses are coefficients of variation (%). b

SSP is solid sawn Southern Pine; SPP, solid sawn ponderosa pine; GHF, glulam HemñFir; GSP, 

 glulam Southern Pine; and PSL, parallel strand lumber. c

Initial slope and load at 5% offset were evaluated directly for average of 10 replicates; 

 variability was not measured. 

d

Rivet yield modes not assessed for these specimens. 

Figure 10óYield mode IV is characterized by 

reverse curvature of rivet between tip and head 

(solid-sawn Southern Pine, 90-mm rivets).

Figure 11óYield mode IIIm is characterized by 

single curvature of rivet below head (solid-sawn 

Southern Pine, 65-mm rivets).
10 

specimen to break in the glulam at the cross-section transition rather than at the joint. Evaluation of the test loadñ

deflection curve showed that the joint had reached its maximum load before the break, and thus the curve properly 

reflected joint behavior to maximum load. 

Analysis of Results 

Phase I. Small Connection Tests 

Figures 8 and 9 show curves of the form 

e y = a tanh(bx) + cx + dx (1) 

where a, b, c, and d are constants and e is the integral of the 

natural logarithm. The curve represents the average of all 

tests in a particular set. The curves were fitted by the 

DavidonñFletcherñPowell (FletcherñPowell 1980ñ81) 

(quasi-Newton) algorithm. The initial slope of the curve is 

the derivative of Equation (1) at x = 0, ab + c. Figures 8 

and 9 also show straight lines with this slope, offset by a 

displacement equal to 5% of the rivet ìdiameter,î for which 

we use the average of the rivet cross-sectional dimensions 

to derive an offset of 

0.24mm

2

6.4mm 3.2mm 0.05 = + (2) 

Table 5óLarge connection parallel-to-grain test resultsa

Joint ID 

Load at 5% 

offset 

(kN/rivet) 

Initial slope 

(kN/mm/rivet) 

Max. load up 

to 4.8 mm 

(kN/rivet) 

Maximum load 

(kN/rivet) 

Failure 

modeb

P3 0.61 4.23 1.30 1.36 W 

P4 0.89 3.30 1.50 1.57 W 

P5 0.88 7.67 2.90 2.94 R 

P6 1.02 6.87 2.52 2.53 R 

P7 0.78 3.78 2.08 2.09 W 

P8 0.80 4.57 2.34 2.34 W 

P9 1.08 12.5 3.14 3.14 R 

P10 1.11 7.86 3.16 3.16 R/W 

P11 1.15 4.56 2.06 2.14 W 

P12 1.23 5.62 2.09 2.09 W 

P13 1.17 7.02 4.51 4.79 R 

P14 1.14 9.02 4.48 4.71 R 

P15 0.95 5.31 2.70 2.79 W 

P16 1.33 5.10 3.66 3.66 W 

a

Averages for 2 or 3 replicates with 200 rivets/connection. 

b

R indicates rivet bending and localized wood crushing; W indicates gross wood failure. 

Table 6óLarge connection perpendicular-to-grain test resultsa

Joint 

ID 

Load at 

5% offset 

(kN/rivet) 

Initial slope 

(kN/mm/rivet) 

Max. load up 

to 4.8 mm 

(kN/rivet) 

Maximum load 

(kN/rivet) 

Failure 

mode 

Q1 0.82 1.40 2.94 3.24 W 

Q2 0.81 1.10 2.67 2.91 W 

Q3 1.33 1.13 1.85 1.88 W 

Q4 0.92 1.08 2.10 2.22 W 

Q5 0.90 1.57 2.93 3.59 W 

Q6 1.06 1.19 3.49 4.69 W 

Q7 1.56 1.09 2.33 2.78 W 

Q8 1.25 0.90 2.23 2.80 W 

a

Averages for 2 replicates with 80 rivets/connection. 
11

The ìload at 5% offsetî displacement in Table 4 is the load 

at the intersection of the average fitted curve with the offset 

straight line as described, consistent with the definition of 

nominal yield values used in NDS (AF&PA 2001). The 

displacement at this intersection is also reported. 

Basic trends can be seen in the data, but none is completely 

consistent across all measures of capacity (Table 4). At the 

5% offset measure, longer rivets produced higher load capacities, except 90-mm rivets perpendicular to grain. At the 

same measure, parallel-to-grain capacity was higher than 

perpendicular-to-grain capacity for all wood types except 

PSL. At the 4.8-mm measure, longer rivets produced higher 

load capacities and parallel-to-grain capacity was higher than 

perpendicular-to-grain capacity for all wood types except 

solid-sawn Southern Pine with 65-mm rivets and PSL. At 

the peak measure (maximum load), longer rivets produced 

higher load capacities, but there is no trend for parallel-tograin capacity relative to perpendicular-to-grain capacity. 

The load and displacement at 5% offset were determined as 

single values for each suite of 10 replicates, so variability is 

not reported. The code definition of nominal yield value as 

the load at 5% dowel diameter offset seems unfortunate. 

The correlation coefficient (r2

) for 5% offset load as a predictor of peak load is only 0.16. The load at a displacement 

of 4.8 mm, the average of rivet cross-sectional dimensions, 

is the measure of ultimate load used in the literature 

(Buchanan and Lai 1994, Karacabeyli and others 1998). 

The correlation coefficient for load at 4.8 mm as a predictor 

of peak load is 0.84. This suggests that if the goal of design 

guidelines is to achieve a desired reliability at ultimate load 

levels or to achieve an appropriate factor of safety at service 

load levels, then defining connector capacity at higher load 

levels makes sense. 

Figure 12óTypical loadñdisplacement plots for large 

parallel-to-grain specimens. See Appendix for 

description of test specimens. 

Figure 13óTypical failure mode for large parallelto-grain specimens. 

Figure 14óTypical failure modes for large perpendicular-to-grain tests. 
12 

Phase II. Large Connection Tests 

Statistical analysis of parallel-to-grain test results showed the 

effect of the design variables on connection performance. 

The peak load of the specimens with 32- by 25-mm rivet 

spacing was 64% higher than that of the specimens with 

25.4- by 12.7-mm rivet spacing. The peak load of the 90-mm 

rivets was 52% higher than that of the 65-mm rivets and 

95% higher than that of the 40-mm rivets. End distance had 

an insignificant effect on any dependent variable, though this 

finding would most certainly change if even smaller end 

distances were tested. Note that the end distance required by 

NDS (AF&PA 2001) for these connections is 102 mm. 

The effect of glulam thickness can be seen by comparing 

joints P7 and P8 with P3 and P4, respectively. The maximum load of the 130-mm-thick glulam (P7 and P8) was 51% 

higher than that of the 79-mm-thick glulam (P3 and P4) 

(Table 5), consistent with common sense but contrary to 

predictions of the code analysis. As with the small specimens, the code use of the load at 5% offset is unfortunate 

because this load is a poor predictor of peak load (r2

 = 0.36). 

A much more sensible definition of connection capacity 

would be the maximum load achieved before deflection 

reached 4.8 mm. 

Comparable analysis of perpendicular-to-grain results 

showed that the independent variables of vertical spacing, 

rivet length, and support type had a significant effect on 

peak load. The peak load for the 25.4-mm spacing was 57% 

higher than that for the 12.7-mm spacing (Table 6). Similarly, the peak load for the 90-mm rivets was 33% higher 

than that for the 65-mm rivets. The short continuously supported beams had a peak load that was 15% higher than that 

of the longer simply supported beams. 

In some parallel-to-grain tests that were considered wood 

failures, inspection of the rivets showed significant plastic 

bending. The presence of bent rivets means that before the 

test was stopped, the rivets and the surrounding wood substrate were stressed beyond their elastic limits. Because tests 

were continued beyond peak load, one cannot be sure as to 

whether inelastic bending occurred before or after gross 

wood failure. Of course, at low load levels there is no reason 

to expect a difference in rivet behavior in a specimen that 

will eventually experience gross wood failure and one that 

will fail by an EYM mode. 

As load levels increase, the rivets bend and the wood surrounding the rivets compresses. If the wood does not rupture, the rivets and surrounding wood will go into the inelastic range. If the wood still does not rupture, a classic EYM 

failure occurs (for example, Fig. 12, specimen 14-2). Such 

was the case for specimens with failure mode R (rivet bending and localized wood crushing) described in the Appendix.

If wood rupture occurs before final EYM failure, either 

catastrophic sudden failure could occur (Fig. 12, specimen 

16-2) in or a crack may redistribute forces so that loading 

could continue. Subsequent failure could occur by either 

EYM (Fig. 12, specimen 15-2) or additional wood rupture 

(Fig. 12, specimen 12-2). 

For the parallel-to-grain rivets, per rivet values for the 

200-rivet connections ranged from 80% of the values for the 

8-rivet connections to 92% of the values for the 80-rivet 

connections. Similar results were found for perpendicularto-grain per rivet values. 

Conclusions 

The tests described in this report are consistent with existing 

literature and therefore can be used in combination with the 

literature to provide verification for analytical models. This 

report includes the first published results for rivets in Southern Pine timber, thus providing important support to the 

NDS recognition of rivet connections in this species. The 

results suggest that the current code provisions are reasonable, although an alternative analysis could improve certain 

aspects of the code. 

The results of testing large specimens loaded parallel to 

grain show the important behavioral differences between 

connections limited by EYM-type failure modes and those 

limited by wood failure modes. When EYM-type failures 

occur, the connection maintains positive stiffness to large 

displacements. When wood failure modes occur, the connection loses stiffness suddenly and can fail catastrophically, 

and at a lower per-rivet load than does a similar EYM-failed 

joint. The displacements associated with EYM-type failure 

are far beyond the usual range of serviceability, but they 

indicate that the connections may be suitable for use in 

seismic design if the designer can ensure that an EYM-type 

failure mode will control. Different strength reduction factors (for load resistance factor design, LRFD) or safety 

factors (for allowable stress design, ASD) might be justified 

for different modes. A primary requirement of any design 

procedure should be to predict whether a given connection 

will fail by EYM-type mode or wood mode. 

Wood failure occurs when the energy built up in the connection exceeds the sum of the tensile and shear capacities of 

the wood fiber on the boundary of the rivet group before it 

exceeds the accumulated rivetñwood bearing strength of 

individual fasteners. Rivet bending and wood crushing lead 

to a more uniform stress redistribution and higher peak loads 

at larger displacements than do block shear failures. 

The effects of rivet length and spacing are as one might 

expect: longer rivets provide increased capacity, and rivet 

spacing influences whether the connection will exhibit 

EYM-type failure or wood failure. Rivet spacing of approximately 32 mm along the grain and 25 mm across the 

grain seems to be adequate for rivets up to 90 mm in length. 

A more detailed study would be required to determine exactly what minimum spacing is required to ensure EYM-type 

failure for each rivet length. 
13

Our results show that the NDS requirements for end distance 

are conservative, as connections with end distances of 50 

and 150 mm performed similarly. Again, a more detailed 

study with end distances between 25 and 100 mm might be 

able to show true minimum end-distances to ensure EYMtype failures. The results show that when wood failure 

modes occur, thicker glulams have higher capacity. This 

finding contradicts the size effect in the code analysis. 

The wood engineering community has devoted a significant 

amount of brain power and printed space over the years to 

establishing a standard definition for the capacity of a wood 

connection. The current NDS bases connection capacity on 

the load at yield, which is defined by the 5% offset method 

(values in NDS are calibrated down to approximate load at 

proportional limit, but that is somewhat beside the point). 

Test results show that load at yield is a poor predictor of 

ultimate load. While that may be a satisfactory state of 

affairs when design strictly follows an allowable stress 

approach, it is not reasonable today. Basic design for seismic 

events considers the ability of a structure to maintain load 

under significant displacements, so information about behavior beyond the 5% offset is critical. In addition, the LRFD 

philosophy provides a framework within which a designer 

can assess the suitability of a structure with regard to ultimate behavior where knowledge about behavior beyond the 

5% offset is critical. 

Literature Cited 

AF&PA. 1999. General dowel equations for calculating 

lateral connection values. Tech. Rep. 12. Washington, DC: 

American Forest and Paper Association. 

AF&PA. 2001. National design specification for wood 

construction. Standard ANSI/AF&PA NDSñ2001. 

Washington, DC: American Forest & Paper Association. 

ASTM. 1988. Standard test methods for mechanical fasteners in wood. Standard ASTM D1761ñ88. Philadelphia, PA 

American Society for Testing and Materials. 

Aune, P.; PattonñMallory, M. 1986. Lateral load-bearing 

capacity of nailed joints based on the yield theory: Theoretical development. Res. Pap. FPLñRPñ469. Madison, WI: 

U.S. Department of Agriculture Forest Products Laboratory. 

20 p. 

Buchanan, A.H.; Lai, J.C. 1994. Glulam rivets in radiata 

pine. Canadian Journal of Civil Engineering. 21: 340ñ350. 

CSA. 2001. Engineering design in wood (limit states design). Standard CSAñO86ñ01. Toronto, Ontario: Canadian 

Standards Association. 

FletcherñPowell, R. 1980ñ81. Practical methods of optimization. Vol. 1.Unconstrained optimization. New York, NY: 

John Wiley & Sons. 

Foci, R.O.; Long worth, J. 1975. Analysis and design of 

grip lam nailed connections. ASCE Journal of Structural 

Division, 101(STñ12): 2537ñ2555. 

Karacabeyli, E.; Fraser, H. 1990. Short-term strength of 

glulam rivet connections made with spruce and Douglas-fir 

glulam and Douglas-fir solid timber. Canadian Journal of 

Civil Engineering. 17: 166ñ172. 

Karacabeyli, E.; Fraser, H.; Deacon, W. 1998. Lateral and 

withdrawal load resistance of glulam rivet connections made 

with sawn timber. Canadian Journal of Civil Engineering. 

25: 128ñ138. 

Madsen, B. 2000. Behaviour of timber connections. North 

Vancouver, BC, Canada: Timber Engineering Ltd. 

McGowan, W.M. 1966. A nailed plate connector for gluedlaminated timbers. Journal of Materials. 1(3): 509ñ535. 

AppendixóResults of Large 

Connection Tests 

Definitions of failure modes used in the tables: 

BS (block shear)óShear-block pull-out exhibiting shear 

failures in at least three planes and end-grain tensile failure. 

In most cases, the shear block was confined to an area beneath the plate, with edges perpendicular to the original 

surface defined by rivet holes. 

DSP (double shear plane)óTwo splits passing completely 

through the timber thickness. Tensile failures were not 

apparent in all instances. In some cases, the shear planes 

extended to the shear-plate seat in the timber grip. In other 

cases, the splits ran the length of the timber but did not pass 

all the way through the grip end. 

NB (nail bearing)óBearing deformation of the wood followed by surface splitting along interior nail rows. In some 

instances, surface splits were apparent along interior rivet 

rows 5 to 7. 

SSP (single shear plane)óSingle split in the timber extending through the thickness and along the length, often to the 

timber grip. This failure mode was not accompanied by 

tensile rupture. In most cases, the split occurred along an 

edge row of rivets. One end of the plate-displacement reference bar was fastened to the wood outside this split so that 

the loadñdisplacement plot indicated an increase in the rate 

of plate displacement relative to the wood, despite the fact 

that no significant bearing deformation occurred at the rivetñ

wood interface. In some instances, the split could be attributed to shake or the existence of juvenile wood in the edge 

lamination, which was subject to the greatest shear stress. 
14 

TP (tension perpendicular to grain)óLongitudinal split 

passing through the nail pattern rather than along an edge. 

The split ran the length of the timber, wide at the connection 

end and tapering to an incomplete failure at the grip. One 

explanation is that load eccentricity, possibly resulting from 

surface splits in one edge lamination, caused the plate to 

rotate slightly. This slight rotation placed a combined tension 

perpendicular to grain and shear stress on the weak plane, 

causing a fracture to propagate toward the grip. 

Results of 30 parallel-to-grain large connection tests (two plates with 100 rivets per plate) 

Joint 

ID 

sp

(mm) 

sq

(mm) 

L

(mm) 

a

(mm)

Glulam 

thickness

(mm) 

Load at 

4.8-mm 

displacement 

(kN) 

Peak 

load 

(kN) 

Displacement at peak 

load (mm) 

Failure

modea

P3ñ1 25.4 12.7 40 50 79.4 310 348 5.91 DSP 

P3ñ2 50 79.4 310 241 1.41 DSP 

P3ñ3 50 79.4 310 228 2.40 SSP 

P4ñ1 150 79.4 310 311 1.88 SSP 

P4ñ2 150 79.4 310 274 1.05 SSP/TP 

P4ñ3 150 79.4 313 356 6.59 BS 

P5ñ1 32 25.4 40 50 130 547 558 7.05 NB 

P5ñ2 50 130 614 616 11.26 NB 

P6ñ1 150 130 490 492 6.15 NB 

P6ñ2 150 130 518 519 6.34 NB 

P7ñ1 25.4 12.7 40 50 130 518 383 3.52 BS, 2 

P7ñ2 50 130 451 453 5.91 BS, 2 

P8ñ1 150 130 451 401 4.55 BS, 1 

P8ñ2 150 130 534 537 7.38 TP/SSP 

P9ñ1 32 25.4 65 50 130 534 627 3.10 NB 

P9ñ2 50 130 534 628 3.00 NB 

P10ñ1 150 130 534 615 4.23 SSP 

P10ñ2 150 130 534 647 3.91 NB 

P11ñ1 25.4 12.7 65 50 130 379 408 5.35 BS 

P11ñ2 50 130 379 446 2.22 BS 

P12ñ1 150 130 379 406 2.06 SSP 

P12ñ2 150 130 379 429 2.18 TP/SSP 

P13ñ1 32 25.4 90 50 216 905 931 8.83 NB 

P13ñ2 50 216 897 987 10.71 NB 

P14ñ1 150 216 897 874 4.05 NB 

P14ñ2 150 216 918 1,010 10.05 NB 

P15ñ1 25.4 12.7 90 50 216 918 574 3.26 BS, 1 

P15ñ2 50 216 504 542 7.38 BS, 1 

P16ñ1 150 216 504 664 2.03 BS, 1 

P16ñ2 150 216 504 800 3.81 BS, 1 

a

For block shear (BS) failure, numeral indicates number of planes affected. 
15

Results of 16 perpendicular-to-grain large connection tests (two plates with 40 rivets per plate) 

Joint ID 

sp

(mm) 

sq 

(mm) 

L 

(mm) 

a 

(mm) 

Glulam 

thickness 

(mm) 

Load at 

4.8-mm 

displacement 

(kN) 

Peak 

load 

(kN) 

Displacement at 

peak load 

(mm) 

Failure 

mode 

Q1ñ1 25.4 25.4 65 102 130 228 255 7.49 W 

Q1ñ2 102 130 242 264 8.59 W 

Q2ñ1 73 130 255 289 9.89 W 

Q2ñ2 73 130 171 176 6.67 W 

Q3ñ1 25.4 12.7 65 73 130 171 124 2.91 W 

Q3ñ2 73 130 160 175 11.33 W 

Q4ñ1 127 130 176 180 5.66 W 

Q4ñ2 127 130 237 286 12.22 W 

Q5ñ1 25.4 25.4 90 73 216 232 288 8.88 W 

Q5ñ2 73 216 286 373 14.07 W 

Q6ñ1 73 216 271 377 15.48 W 

Q6ñ2 73 216 181 210 7.33 W 

Q7ñ1 25.4 12.7 90 102 216 192 235 10.88 W 

Q7ñ2 102 216 168 193 7.98 W 

Q8ñ1 127 216 188 255 13.37 W 

Q8ñ2 127 216 228 255 7.49 W