Dowel-Nut 

Connection in 

Douglas-fir Peeler 

Cores 

Ronald W. Wolfe 

John R. King 

Agron Gjinolli 

United States

Department of 

Agriculture 

Forest Service 

Forest 

Products 

Laboratory 

Research 

Paper 

FPL−RP−586
Abstract 

As part of an effort to encourage more efficient use of 

small-diameter timber, the Forest Products Laboratory 

cooperated with Geiger Engineers in a study of the structural 

properties of Douglas-fir peeler cores and the efficacy of a 

“dowel-nut” connection detail for application in the design of 

a space frame roof system. A 44.5-mm- (1.75-in.-) diameter 

dowel-nut connector was found to be economically 

feasible at a design capacity of 44.5 kN (1.0 × 104

 lbf) for a 

127-mm- (5-in.-) diameter Douglas-fir peeler core. Variables 

that affect joint strength and failure mode are location of the 

dowel nut, wood moisture content, presence of reaction 

wood, and grain angle orientation with respect to force vectors. The results of this study provide a basis for deriving 

design properties for peeler core structural application in a 

space frame roof system and the foundation for establishing 

a database to support small-diameter timber design and 

construction standards. 

Keywords: Dowel nut, small-diameter timber, Douglas-fir, 

design properties 

July 2000 

Wolfe, Ronald W.; King, John R.; Gjinolli, Agron. 2000. Dowel-nut 

connection in Douglas-fir peeler cores. Res. Pap. FPL–RP–586. 

Madison, WI: U.S. Department of Agriculture, Forest Service, 

Forest Products Laboratory. 16 p. 

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

public from the Forest Products Laboratory, One Gifford Pinchot Drive, 

Madison, WI 53705–2398. Laboratory publications are sent to hundreds 

of libraries in the United States and elsewhere. 

The Forest Products Laboratory is maintained in cooperation with the 

University of Wisconsin. 

The United States Department of Agriculture (USDA) prohibits discrimination in all its programs and activities on the basis of race, color, national 

origin, sex, religion, age, disability, political beliefs, sexual orientation, or 

marital or familial status. (Not all prohibited bases apply to all programs.) 

Persons with disabilities who require alternative means for communication 

of program information (Braille, large print, audiotape, etc.) should contact 

the USDA’s TARGET Center at (202) 720–2600 (voice and TDD). To file a 

complaint of discrimination, write USDA, Director, Office of Civil Rights, 

Room 326-W, Whitten Building, 1400 Independence Avenue, SW, Washington, DC 20250–9410, or call (202) 720–5964 (voice and TDD). USDA 

is an equal opportunity provider and employer. 

Contents 

Page

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

Objective and Scope ..............................................................1

Material..................................................................................1

Methods .................................................................................2

Dowel-Nut Connection ......................................................2

Pilot Tests ..........................................................................2

Joint Tests ..........................................................................3

Compression Tests.............................................................3

Results....................................................................................3

Pilot Test Failures ..............................................................3

Test Joint Failure Modes....................................................3

Compression Tests.............................................................4

Influencing Factors ............................................................4

Analysis .................................................................................5

Joint Strength .....................................................................6

Conventional Design Values for Strength..........................6

Joint Stiffness.....................................................................6

Joint Design Recommendations.........................................7

Compression Tests.............................................................7

Conclusions............................................................................8

Recommendations for Future Study......................................8

Literature Cited ......................................................................9

Appendix A—Pilot Tests of Dowel-Nut Connector ............10

Appendix B—Test Results...................................................11

Appendix C—Design of Dowel-Nut Connectors.................16
Dowel-Nut Connection in 

Douglas-fir Peeler Cores 

Ronald W. Wolfe, Research Engineer

Forest Products Laboratory, Madison, Wisconsin 

John R. King, Engineer 

Geiger Engineers, Bellingham, Washington 

Agron Gjinolli, Visiting Scientist 

Forest Products Laboratory, Madison, Wisconsin 

Introduction 

Structural application of small-diameter timber is currently 

limited by conventional construction standards. The round 

tapered shape is not compatible with conventional framing 

and cladding methods. Round timber connections are generally custom made and their design properties are based on 

limited test data, making these connections expensive and 

their design capacity conservative. 

The Forest Products Laboratory (FPL) has initiated a research program to assess uses for small-diameter timber in an 

effort to encourage more efficient use of this material. Research on structural applications includes the characterization 

of material and connection properties, processing, design of 

demonstration structures, and economic feasibility. The goal 

of this work is to encourage innovative uses in structures by 

providing not only the information needed to design structures but also some basis for evaluating economic feasibility. 

Geiger Engineers cooperated with the FPL in a study to 

assess the efficacy of using Douglas-fir peeler cores, a byproduct of wood veneer production, in structural applications. Geiger’s interest included verification of design assumptions made in reference to guidelines noted in the 

National Design Specification for Wood Construction 

(ANSI/AF&PA 1997) and finding an economically feasible 

connection. The FPL’s interest was in strengthening the 

technical basis necessary to support development of smalldiameter timber design and construction standards. 

Objective and Scope 

The objectives of this study were as follows: 

• To provide a basis for deriving design stresses for 

127-mm- (5-in.-) diameter Douglas-fir peeler cores to be 

used as axially loaded elements in a space frame truss roof 

assembly 

• To add to the limited knowledge on techniques for 

making more efficient use of small-diameter round timber 

This paper describes test procedures and results for a very 

specific application. Its scope is limited to the consideration 

of only one joint configuration designed to resist only axial 

tension and the axial compressive strength and stiffness of 

the material. The target application assumes that tensile 

stresses carried by the timber will be limited by the connection. In compression, it is assumed that the critical buckling 

load will be the limiting factor that makes stiffness the critical issue for design. 

Material 

The round timbers used in this study were 127-mm- (5-in.-) 

diameter Douglas-fir peeler cores. A peeler core is the cylindrical piece left after veneer is peeled from a log. The growth 

rate of the timbers used to produce these cores ranged from 

4 to 14 rings/in (1.6 to 5.5 rings/cm) (hereafter referred to as 

rpi). The pith was rarely centered, so it was not unusual for 

the wood on the outer surface to be within 8 to 10 annual 

rings of the pith. Joint tests were conducted using 1.8-m- 

(6-ft-) long specimens, and the material compression tests 

were conducted with members cut to a length 4 times the 

average diameter (or approximately 0.5 m (1.6 ft) in length). 

The peeler cores had a slight taper, varying in diameter by as 

much as 13 mm (0.5 in.) over the 1.8-m (6-ft) length of the 

joint test samples. Knots were relatively small because these 

cores were graded to meet the No. 2 lumber grade knot requirements. The cores were machined and kiln dried at FPL 

prior to testing. 

The test connections were fabricated from relatively high 

yield (0.39 to 0.41 GPa (56 to 60 × 103

 lb/in2)) steel. Each 

joint consisted of a “dowel nut” cut from solid, round, 

31.8- to 44.5-mm- (1.25- to 1.75-in.-) diameter hot roll steel 

and a 19-mm- (0.75-in.-) diameter threaded rod. 
2 

Methods 

Tests included the tensile capacity of the dowel-nut connection and the compressive strength and stiffness of the peeler 

core sections, as well as the evaluation of nondestructive test 

parameters useful as quality indicators. Pilot tests were conducted to gain some knowledge of joint capacity and influence of fabrication variables. 

Moisture content and specific gravity (ASTM 1993a) samples were taken after the joints were tested, but prior to the 

compression tests. Cross sections were examined to quantify 

growth rate, extent of pith eccentricity, proportion of juvenile 

wood, and occurrence of reaction wood. Pith eccentricity was 

noted as an index (0 to 1) calculated as the ratio of the distance between the pith and the geometric center of the cross 

section and the radius. Knot size and location were also 

mapped for each compression sample. 

Dowel-Nut Connection 

The dowel-nut connection evaluated (Fig. 1) is similar to that 

commonly employed in “knock-down” furniture. It requires 

that two holes be drilled in the peeler core: one along the 

diameter of the peeler core and the second along the 

centroidal axis. The first hole is large enough to accommodate the dowel nut and is centered 254 mm (10 in.) from the 

end. The second hole is 25.4 mm (1 in.) in diameter—large 

enough to accommodate the threaded rod, which is screwed 

into a threaded hole in the dowel nut. 

Because of the difficulty of holding a round section in the 

test machine, the test connections were used to mount the 

samples in the test machine. The joints were tested in axial 

tension. The threaded rods protruding from each end of the 

test sample were mounted to opposing heads of the test 

machine and in series with a 222-kN (50 × 103

 lbf) load cell. 

Consequently, two joints were tested at one time, one on each 

end of the peeler core. The two dowel-nut holes on a single 

test sample were oriented at 90° to each other. In each case, 

only one joint failed, but the load–displacement characteristics were recorded for both joints. 

Pilot Tests 

Preliminary analysis of the dowel-nut connection in Douglasfir suggested that these joints were capable of attaining a 

design load close to 44.5 kN (1.0 × 104

 lbf). A total of 16 

pilot test joints were fabricated and tested. The initial pilot 

test was conducted on green wood. After this test, the failed 

ends were cut off, and the joints were re-machined and retested in the dry condition. Moisture change from the time of 

fabrication to test had a negative effect on joint behavior. 

On the basis of these tests, we decided that all test cores 

should be kiln dried prior to joint fabrication because the 

intended application for these joints is a protected environment where the wood will equilibrate at a moisture content 

of 15% or less. 

The predominance of compression-parallel-to-grain failures 

in the pilot tests motivated us to increase the diameter of the 

steel dowel from the 31.8-mm (1.25-in.) size used in the pilot 

tests to 44.5 mm (1.75 in.). We also decided to place a 

12.7-mm- (0.5-in.-) wide by 0.7-mm- (0.03-in.-) thick 

stainless steel strap with a threaded “T-bar” tightening device 

(Fig. 2) just above the dowel hole and a truss plate in the end 

grain oriented perpendicular to the dowel-nut hole to reinforce the resistance of the core to tension-perpendicular-tograin forces imposed by the round dowel. The clamp was 

tightened with a torque of roughly 67.8 N⋅m (50 lbf⋅ft). Pilot 

tests of the dowel-nut connection are described in detail in 

Appendix A. 

Figure 1—Dowel-nut connection detail. 

Clamp 

Dowel 

Nut 

Figure 2—Steel-strap pipe clamp with 

threaded T-bar tightening device. 
3 

Joint Tests 

A total of 32 test joints were fabricated to assess the load 

capacity of the dowel-nut connections. Load–deformation 

data were collected for each joint. A linear variable differential transducer (LVDT) with a 6.4-mm (0.25-in.) displacement range was attached to each threaded rod, with its core 

bearing on the end of the test timber. Load was applied 

maintaining a constant displacement rate of 0.254 mm/min 

(0.01 in./min) to attain maximum strain in 5 to 10 min. 

As the test machine heads were moved apart, the displacements measured by the LVDTs were recorded and plotted 

against the tensile force measured by the load cell. The test 

continued until the load either dropped off or displacement 

exceeded the range of the LVDT. In all cases, the LVDT 

displacement range exceeded the linear load–displacement 

limit for the test joints so that characteristic curves were 

recorded to maximum serviceability limits. 

After the test, the joints were inspected and disks were cut 

for moisture content and specific gravity determination. The 

joints were visually inspected to note the failure mode, any 

influence caused by knots or other defects, pith index, and 

growth rate (Fig. 3). 

Compression Tests 

Testing wood in the green condition has the advantage of 

saving the time and expense of kiln drying prior to testing. 

Since no machining was required for the test and we were 

fairly confident of the failure mode, tests on green wood 

(green tests) were deemed appropriate. The results of green 

tests can be adjusted using dry/green ratios for strength and 

modulus of elasticity (MOE). These ratios are published in 

ASTM D2555 (ASTM 1998). As the ASTM values are 

based on tests of small clear specimens, we deemed it 

necessary to check their application to peeler cores by testing 

both green and dry peeler cores. 

A total of 38 compression parallel-to-grain tests were conducted following procedures given in ASTM D198 (ASTM 

1998) to assess member strength and stiffness. Following this 

standard, the tests were conducted on a length less than 

17 times the radius of gyration. In the case of a round timber, 

the radius of gyration is 0.25 times the diameter, giving a 

length limit of 4.25 times the diameter. The peeler core 

specimens were cut to a length of 4 times their average diameter (or 0.50 m (20 in.) in length). Specimen ends were 

carefully cut perpendicular to their axis. The specimens were 

then placed in a test machine and loaded in compression 

parallel to grain. Load was applied using a constant displacement rate of 1 mm/min (0.04 in/min) (strain rate = 

0.2%/min). Loading continued to a displacement beyond the 

point of maximum load. Data recorded included the load on 

the specimen as a function of load head movement. Load 

head movement was assumed to provide a measure of compressive strain over the length of the specimen up to proportional limit. The location of compression failure was noted 

on each test specimen. 

Results 

This study provided a basis for evaluating the effects of 

material properties, processing, and joint configuration on 

the load capacity of dowel-nut connections. Moisture content 

of the wood at the time of fabrication and characteristics of 

the timber in the area of the connection appeared to have a 

significant influence on failure mode and capacity of these 

connections. 

Pilot Test Failures 

The first round of pilot tests failed in combined compression 

parallel to grain and tension perpendicular to grain. There 

was a tendency for the dowel bearing force to push both 

halves of the peeler core apart, rotating about a confinement 

ring placed just above the dowel-nut hole. As a result, we 

decided to retest the joints with the confinement band further 

away from the dowel-nut hole and a truss plate pressed into 

the end of the timber to help resist tension perpendicular 

to grain. 

By the time the second group of pilot test joints had been cut 

and re-machined, their moisture content had dropped from 

25%–30% to 9%–20%. These drier specimens no longer 

exhibited excessive compression parallel to grain. Instead, 

they failed in shear and tension parallel to grain. 

Test Joint Failure Modes 

The load–displacement curves for the test joints exhibited a 

characteristic initial stiffening, followed by a linear portion 

that extended beyond 80% of maximum load. There was little 

Figure 3—Comparison of growth rate to pith index and 

amount of reaction wood. Log on left has centered pith 

and no reaction wood, but a growth rate of only 4 rpi. 

Log on right has high pith index and reaction wood, 

but a more desirable growth rate of 8 rpi. 

.7r 

Reaction wood 
4 

sign of the compression-parallel-to-grain deformation prevalent in the pilot tests conducted at higher moisture content. 

Only the last specimen tested (joint 16) showed any sign of 

bearing deformation parallel to grain. 

The joint test results (Appendix B) exhibited three predominant failure modes (Fig. 4). The most common failure mode 

was combined shear and tension parallel to grain. The second 

most common failure mode was one in which no sign of 

tensile failure was evident; shear in two planes caused a 

section to pop out of the end. In a few instances, the failure 

initiated as tension perpendicular to grain, appearing as a 

single split in a plane parallel to and through the dowel hole. 

In many cases, tension perpendicular-to-grain and shear 

apparently acted together in splitting the timber from the 

dowel hole to the end. In several instances, the failure appeared as a complete tension-parallel-to-grain failure on both 

sides of the dowel hole with no accompanying shear. One 

joint with this kind of failure (joint 10) failed at a relatively 

low load (106.75 kN (2.4 × 104

 lbf/in2)) as a result of low 

quality wood around the dowel nut. On one side of the 

dowel, the wood was 80% juvenile and reaction wood; on the 

other side, the grain of 50% of the wood was oriented 30° to 

45° to the load direction. Finally, one joint (joint 1) failed as 

a result of knots in the area of the dowel hole. Joint 1 failed 

at 57.8 kN (1.3 × 104

 lbf/in2) but it was not included in the 

analysis of results because no knots should be permitted 

within three diameters of the dowel hole. 

Eccentric loading was noted for two test joints, joints 5 and 

12. This resulted from a slight misalignment of the longitudinal hole drilled for the threaded rod. The misalignment was 

noticeable; the dowel nut protruded approximately 6.3 mm 

(0.25 in.) from one surface of the peeler core. It is difficult to 

assess the true effect of this eccentric load on the basis of two 

tests, especially when the resulting joint strength values fall 

at opposite extremes of the joint strength distribution. 

Compression Tests 

Failures for the compression tests appeared as compression 

wrinkles in the surface. These wrinkles were most apparent 

around the knots. For the sample with the lowest compressive 

MOE (Ec) (sample C-5-7), the failure involved shear displacement initiating at a drying check in an area of juvenile 

wood close to one surface. In several cases where there was a 

high pith index, the swivel bearing plate of the active loadhead tilted to accommodate the lower Ec of the juvenile 

wood zone. 

The compression test load–displacement curves have several 

distinct characteristics. They all begin with initial stiffening 

where a non-uniform end surface levels out. As the stresses 

become more uniformly distributed (>2.54-mm (>0.1-in.) 

displacement for 508-mm (20-in.) gauge length and 44.5 kN 

(1.0 × 104

 lbf/in2)), the curves become fairly linear to 50% of 

the maximum load. Beyond 60% of the maximum load, the 

curves exhibit an elastoplastic region, which then generally 

turns plastic beyond 3.81 mm (0.15 in.). Slopes that are 

measured for determination of MOE are therefore measured 

primarily between deformations of 2.54 and 3.81 mm (0.1 

and 0.15 in., respectively) or strains of 0.5% and 0.75%. 

Influencing Factors 

Factors other than moisture that appeared to have a strong 

influence on joint performance included wood growth characteristics, processing variables, and connection details. 

Wood growth characteristics such as reaction wood, pithassociated wood, and angled grain can be controlled by 

grading. Processing variables such as drying and machining 

can be controlled to minimize fabrication and service problems. However, the connection details must be evaluated by 

testing to assure that a solution for one problem does not 

create other problems. 

Reaction wood, pith-associated wood, and angled grain can 

all be detected by visual inspection. Reaction wood appears 

as an area of unsymmetric growth rings (see Fig. 3), readily 

apparent on the cross section of the timber. Pith-associated 

wood is present in all peeler cores. The fact that pithassociated wood is often located off-center increases the 

probability that it will occur in the tension zone to the side 

of the dowel nut unless a conscious effort is made to drill the 

dowel-nut hole through this wood. 

Angled grain occurs as a result of knots or a bend in the tree 

stem that is bent down when the veneer is cut. An annual ring 

pattern on the surface of the peeler core consisting of concentric elliptical rings (Fig. 5) is an indicator of grain that is not 

parallel to the surface. This pattern occurs where several 

annual rings pass through the exposed surface over a short 

Figure 4—Shear, tension-perpendicular-to-grain, and 

tension-parallel-to-grain failures in test joints 

(left to right). 
5 

length. It indicates that the tree had a slight bend or protuberance that defined the annual ring orientation at that location. 

The more circular the concentric rings, the steeper the localized grain just below the surface at that location. 

Reaction and pith-associated wood should be avoided in 

areas subject to high tensile stress. This wood is weaker and 

less stiff than other wood and also exhibits greater longitudinal change with change in moisture content than does mature 

wood. When stresses are applied at an angle to the grain, 

strength is influenced by the proportion of wood loaded in 

tension perpendicular to grain. Wood loaded at this orientation is assumed to be have one-third the shear strength parallel to grain. 

Problems with pith-associated and reaction wood can be 

minimized by proper detailing of the joint. For test joint 5, 

80% of the area to one side of the dowel hole exhibited a 

brash tensile failure classified as reaction wood and/or pithassociated wood failure. On the other side of the dowel hole, 

wood fibers in 50% of the failure area were oriented at an 

angle at least 30° to the direction of principle stress. Angled 

grain was also a problem in joint 10. Joints 5 and 10 failed at 

approximately 106.75 kN (2.4 × 104

 lbf/in2). Joints 3, 6, and 

12 also exhibited pith-associated and reaction wood, but their 

failure loads were 164.6 kN (3.7 × 104

 lbf/in2), 124.5 kN 

(2.8 × 104

 lbf/in2), and 160 kN (3.6 × 104 lbf/in2), respectively. The difference was that the dowel holes of the 

stronger joints had been drilled through the pith and reaction 

wood areas and showed little sign of angled grain. 

Six of the 16 joint test samples (joints 1, 3, 5, 6, 10, and 

12) had some amount of reaction wood. Joints 5 and 10 

contained significant amounts of both reaction wood and 

angled grain in the tensile area on either side of the 

dowel hole. 

Processing variables include drying and machining. Dimensional change with drying and drying checks can create 

problems that reduce joint load capacity. Holes drilled in the 

wood will become smaller as the wood dries. If the steel 

dowel nut is inserted in a tight-fitting hole when the wood is 

green, the shrinkage will result in tension-perpendicular-to 

grain stresses in the wood in the area of the dowel, initiating 

a drying check in a plane parallel to that of the dowel hole. 

Drying checks that were oriented at an angle <45° to the 

dowel hole reduced the effective shear area or tensionperpendicular-to-grain area, increasing the probability of an 

early failure. When wood is dried prior to drilling, such 

stresses are not a problem. When the dowel hole is oriented 

to a drying check at an angle >45°, the check is less likely 

 to contribute to a weak plane for shear- or tensionperpendicular-to-grain stress1

The joint configuration parameters that apparently exert the 

greatest influence on performance are dowel-nut diameter 

and confinement. Naturally, the larger the diameter, the 

greater the end grain surface area subject to compression and 

the smaller the area subject to tension. These two areas 

should be balanced slightly in favor of tension to encourage 

the more ductile compression failure. Confinement had a 

significant effect on both failure mode and load capacity. The 

adjustable steel confinement band placed 25.4 mm (1 in.) 

above the dowel-nut hole used in conjunction with a truss 

plate located in the end grain and oriented perpendicular to 

the dowel-nut hole served to reinforce the specimens against 

tension-perpendicular-to-grain failures. The lower incidence 

of bearing deformation and Poisson’s effect in pushing fibers 

apart made the bands seem less highly stressed in the samples 

from the final test compared with those from the pilot tests. 

However, in one test (test 3), the steel band actually failed in 

tension, and in another test, the confinement stress that built 

up during the test left compression-perpendicular-to-grain 

bearing deformations in the surface of the timber (joint 16). 

In these cases, failures were predominately due to tension 

perpendicular to grain, implying that joint capacity would 

have been increased by using a second band or by pressing 

the metal plate connector into the end grain as was done for 

the pilot tests. 

Analysis 

The peeler cores evaluated in this study were relatively fast 

grown, ranging from 4 to 14 rpi. The average specific gravity 

was low: 0.4 compared with 0.45 generally reported for 

coastal Douglas-fir. The low specific gravity was most likely 

due to the high proportion of juvenile wood. Because juvenile wood normally predominates in the first 10 to 20 years 

of growth (≤8 rpi), the entire section could be assumed to be 

juvenile wood for a 127-mm- (5-in.-) diameter peeler core. 

Figure 5—Concentric elliptical annual 

ring patterns indicate angled grain 

below wood surface. 

Concentric 

ellipse pattern 
6 

Cores with a higher growth rate naturally have a lower proportion of juvenile wood and therefore higher specific gravity, as indicated in Figure 6. This does not mean that slower 

growth in general will yield higher density. Density in mature 

wood is more a function of the percentage of latewood, 

which could easily be greater in fast-grown material than in 

slow-grown material. 

Joint Strength 

Appendix C provides an assessment of the expected maximum strength and design stress for bearing parallel to grain, 

tension, and shear. These assessments are based on the dowel 

bearing capacity of coastal Douglas-fir (ANSI/AF&PA 

1997), published values for clear wood (ASTM 1998, 

D2555), and tensile capacity of sawn timber. 

The average strength for 16 tests was 129 kN (2.9 × 104

 lbf). 

Because two joints were used for each test, the actual sample 

was 32 joints. For the 16 joints that did not fail, strength was 

assumed to be greater than or equal to that of the joints that 

did fail, making the true average strength for the 32 joints 

greater than or equal to that determined for the 16 tests. 

A Student’s t-statistic was used to estimate a 5th percentile 

value for joint strength assuming 31 degrees of freedom, the 

measured mean strength, and the sample standard deviation. 

The resulting value was 93 kN (2.1× 104

 lbf). Assuming a 

predominately tension-type failure and an adjustment of 

2.1 to account for duration of load and factor of safety, 

the design capacity for a normal load duration would be 

44 kN (1.0 × 104

 lbf). 

Conventional Design 

Values for Strength 

The average strength of round timbers (ASTM 1998, Bodig 

and others 1986, Wolfe 1989) is generally noted to be comparable to or slightly higher than that of dimension lumber, 

with roughly two-thirds the variability. The reduction in 

variability is attributed to the fact that the wood fibers of a 

round timber are parallel to the surface and thus parallel to 

the principal stresses. The grain in a tree flows around knots 

and remains continuous. In lumber, on the other hand, the 

wood fibers are often sawn at a slight angle to the surface. 

The grain around knots in a piece a lumber is often discontinuous and cut at steep angles to the surface, which causes 

the knots to have a more deleterious effect on strength. Fiber 

discontinuity is often most obvious along the corners of 

rectangular lumber sections. This occurrence of diving grain 

reduces the effectiveness of the rectangular section. Design 

stresses derived for rectangular members should therefore be 

conservative when applied to a round section of comparable 

material quality. 

Following the National Design Specification for Wood

(NDS) (ANSI/AF&PA 1997) for wood and the strength ratio 

assumed for Construction-grade nominal 4- by 4-in. (standard 89- by 89-mm) lumber, the normal load–duration design 

capacity of the dowel-nut connection would be 44 kN 

(1.0 × 104

 lbf) for a dowel bearing failure mode, 44 kN 

(1.0 × 104

 lbf) for tensile failure, and 31 kN (0.7 × 104 lbf) 

for shear failure. However, the ASTM D2899 standard for 

derivation of shear design stresses for round timbers 

includes a strength ratio of 0.75 (Phillips and others 1985) in 

adjusting from small clear to full-sized round timber strength. 

Using this value and a coefficient of variation (COV) of 

20% for round wood gives a design value of 47 kN 

(1.05 × 104

 lbf) (App. C). 

Design values derived following the ASTM recommendations appear to be in line with the value derived from the test 

data. The design capacity of the dowel-nut connector would 

be 44.5 kN (1.0 × 104

 lbf) for normal load–duration shear, 

tension parallel to grain, and bearing. Tension perpendicular 

to grain, which also contributes to joint failures, could be 

restrained using a second clamp or a metal plate connector 

(“truss plate”) pressed into the end grain. 

The dowel bearing design for bolts in double shear, outlined 

in the NDS for wood published by the American Forest & 

Paper Association (AF&PA), gives predictions of maximum 

load that are close to the predicted value of 137 kN 

(3.1 × 104

 lbf) as opposed to the measured value of 

131 kN (3.0 × 104

 lbf). 

Joint Stiffness 

The slope of the load–displacement curves for the test joints 

ranged from 96 to 204 kN/mm (55 to 117 × 104

 lbf/in), with 

an average value of 150 kN/mm (856 × 103

 lbf/in) and 18% 

COV. The linear range of the curve extended to at least 80% 

of the maximum load. Failures ranged from ductile parallelto-grain compression to brittle tension failures. Brash or 

brittle failures involved tension-perpendicular-to-grain 

stresses or reaction wood. In many cases, the tensionperpendicular-to-grain stresses occurred at places where the 

Figure 6—Increase in specific gravity with growth rate 

resulting from decrease in proportion of juvenile wood 

with age; rpi designates rings/in. 1 in. = 25.4 mm. 
7 

grain was oriented at an angle to the axis as a result of the 

influence of a knot or other defect in the tree. 

Joint Design Recommendations 

The greater ductility of dowel bearing failure in compression 

parallel to grain makes this mode preferable to tension. The 

pilot test joints (App. A), which were made with green lumber, all exhibited this more ductile failure, while only one of 

the 16 joints tested in the dry condition did. Apparently, 

there would be some advantage to reducing the size of the 

dowel nut. This would increase the effective tensile area, 

reducing tensile stress at design load, and have minimal 

effect on the shear area. Reducing the bearing area to the 

point controlled by compression parallel to grain and maintaining confinement to inhibit tension-perpendicular-to-grain 

splitting would reduce the chance of a brash failure in the 

event of an unpredicted overload. 

Compression Tests 

Compressive strength and MOE were measured under both 

green and dry conditions. All test results were adjusted to 

12% by interpolation of the ASTM D2555 dry/green 

(12%/30%) ratios (1.99 for stress and 1.27 for Ec). These 

results were compared using a t-test for mean difference, and 

the dry/green ratios were checked using least squares linear 

regression. 

The ultimate compressive stress (UCS) for specimens tested 

at moisture content exceeding 20% averaged over 21 MPa 

(3.1 × 103

 lb/in2), with 11% COV. Despite the apparent 

influence of unsymmetric pith and knots, neither pith index 

nor knot size appeared to have a significant correlation to 

compressive strength. The average UCS from green tests 

adjusted to 12% moisture content using the ASTM dry/green 

ratio was predicted to be 33 MPa (4.8 × 103 lb/in2

). 

Results of dry tests showed an average UCS of 31 MPa 

(4.5 × 103

 lb/in2). Because these samples were dried below 

12% moisture content, their UCS values should have averaged higher than that predicted using the “green” test results 

and the ASTM dry/green ratio. Results of a t-test to compare 

adjusted values for these two samples suggest that the mean 

values were significantly different at the 0.05 level of significance. The difference may have been due to the proportion of 

juvenile wood or to the fact that drying checks masked the 

reduction in cross section, giving a high estimate of cross 

sectional area. 

Given that the ASTM D2555 dry/green ratio gave an unconservative estimate of member strength at 12% moisture content, an adjustment factor was derived from the data. A least 

squares linear regression analysis of compressive strength as 

a function of moisture content (Fig. 7) indicated that the 

dry/green ratio is 1.78 rather than 1.91. Using this adjustment 

resulted in an average UCS at 12% moisture content of 

29.6 MPa (4.3 × 103

 lb/in2) with a 12% COV. 

The MOE (Ec) for the green samples averaged 3.86 GPa 

(560 × 103 lb/in2

) with a 23% COV. Adjusting these values to 

12% moisture content using the ASTM dry/green ratio (1.27) 

gave an estimated dry MOE of 4.48 GPa (650 × 103

 lb/in2). 

The dry tests showed an average Ec of 5.16 GPa (750 × 

103

 lb/in2) with a 20% COV. Assuming the ASTM dry/green 

ratio could be extended beyond the 12% to 30% range, we 

adjusted these values upward to 12% moisture content. This 

gave an average value of 4.97 GPa (720 × 103

 lb/in2) and 

18% COV. Although the 4.48-GPa and 4.97-GPa values 

(650 and 720 × 103

 lb/in2, respectively) were not significantly different at the 0.05 level of significance, we nevertheless used a least squares regression of Ec against moisture 

content to check the ASTM ratio. The regression suggested 

that the dry/green ratio should be 1.62. Combining the samples using this ratio gave an average Ec of 5 GPa (730 × 

103

 lb/in2) with a 22% COV and 12% moisture content. 

Figure 8 compares UCS and Ec for the two samples adjusted 

to 12% moisture content. In this case, the adjustments were 

1.78 for UCS and 1.62 for MOE going from 30% to 12% 

moisture content. The ASTM D2899 standard procedure for 

the derivation of compression-parallel-to-grain design stress 

gives a value of 18 MPa (2.6 × 103

 lb/in2) at 12% moisture 

content. Removing the adjustment for duration of load (1.52) 

gives a value of 27.5 MPa (4 × 103

 lb/in2). This is close to 

the estimate made using the derived dry/green ratios. The 

ASTM D2899 compression-parallel-to-grain design value 

gives a short column design capacity of 231 kN 

(5.2 × 103

 lbf) for a 127-mm- (5-in.-) diameter round 

Douglas-fir timber at 12% moisture content. 

The longitudinal MOE is normally assumed to be the same 

for bending, tension, and compression. Thus, the Ec values 

measured for the full-size compression tests should provide 

an estimate of the E required for column stability calculations 

following the NDS procedures. However, measured values 

were lower than expected. Using the ASTM D2555 dry/green 

adjustment for Douglas-fir MOE to derive an average value 

Figure 7—Relationship between UCS and moisture 

content for two samples suggests a dry/green ratio 

of 1.78. 
8 

of 4.4 GPa (650 × 103

 lb/in2) limits the design load to 39 kN 

(8.8 × 103

 lbf) at 12% moisture content for a 2.43-m (8-ft) 

column. The additional 14 dry tests support increasing this 

average value to 5.03 GPa (730 × 103

 lb/in2), which would 

increase the design load to 41.4 kN (9.3 × 103

 lbf). 

Conclusions 

A 44.5-mm- (1.75-in.-) diameter dowel-nut connector 

appears to be economically feasible at a design capacity of 

44.5 kN (1.0 × 104

 lbf) for a 127-mm- (5-in.-) diameter 

Douglas-fir peeler core. To achieve a reliable connection, 

however, the fabricator should be aware of variables that 

affect joint strength and failure mode. We conclude the 

following: 

• This design stress is based on the assumption that the 

dowel nut is centered seven diameters (311 mm 

(12.25 in.)) from the end of the peeler core. The derivation considers bearing capacity, tension, and/or shear 

parallel to grain (ASTM D2899) and tension perpendicular to grain. 

• The hole for the dowel-nut connector should be drilled 

after drying the peeler core to prevent problems caused 

by shrinkage of the hole and to assure that any drying 

checks are not aligned with the primary shear plane of the 

joint. 

• Reaction wood should be aligned in the plane of the 

dowel as opposed to being in the areas stressed in tension 

on either side of the dowel. 

• Angled grain should be avoided in the region of the connection. Such grain can be detected by a pattern of concentric ellipses formed by annual rings on a tangential 

surface of the peeler core. A pattern with a major/minor 

axis ratio of 3 or less should be avoided (Fig. 9). 

• In this study, the 5th percentile axial compressive strength 

of full-sized sections was close to the value predicted following ASTM D2899 design procedures. Adjusted to 

12% moisture content, the 127-mm- (5-in.-) diameter 

peeler core sections had a short column design capacity 

of 186 kN (4.2 × 104

 lbf) and a 2.4-m (8-ft) column capacity of 41.4 kN (0.93 × 104

 lbf). 

Recommendations 

for Future Study 

1. Tests should be conducted to assess the value of machining prior to drying as a means of minimizing the effects of 

drying checks. A saw kerf cut prior to drying may help 

control where splitting occurs. Holes drilled parallel to 

the log axis along the centroid may reduce splitting by 

opening an inner surface to evaporation. 

2. Modification of joint parameters should be evaluated for a 

possible increase in joint efficiency. Reductions in dowel 

diameter, combined with the same or increased end distance, will increase the effective areas for tension parallel 

and perpendicular to grain as well as shear parallel to 

grain. By increasing resistance to tension and shear, the 

failure mode may be changed to a more ductile bearing 

deformation with little change in the average capacity of 

the connection. 

Figure 8—UCS adjusted to 12% moisture content 

using a dry/green (12%/30% moisture content) ratio 

of 1.78 for UCS and 1.62 for Ec. 

Figure 9—Joint detail: timber characteristics. Dowel 

nut should pass through the pith (A). If the surface 

annual ring pattern has a L/W ratio (B) less than 3, 

the timber should be rejected for structural applications. The dowel nut should be oriented at least 45¤

 from any severe drying checks (C). Any knots greater 

than 1.2 mm (0.05 in.) in diameter should be at least 

3 times the dowel diameter from the dowel hole (D). 

A 

B 

C 

D 

L W 

>45o 3D 

>1.2 mm 

D 
9 

Literature Cited 

ANSI/AF&PA. 1997. National design specification for 

wood construction., Washington, DC: American National 

Standards Institute/American Forest and Paper Association. 

ASTM. 1993a. Standard test methods for specific gravity of 

wood and wood-base materials. ASTM D2395–93. West 

Conshohocken, PA: American Society for Testing and 

Materials. 

ASTM. 1993b. Standard practice for establishing structural 

grades and related allowable properties for visually graded 

lumber. ASSTM D245–93. West Conshohocken, PA: 

American Society for Testing and Materials. 

ASTM. 1995. Standard test methods for mechanical fasteners in wood. ASTM D1761–88. West Conshohocken, PA: 

American Society for Testing and Materials. 

ASTM. 1998. West Conshohocken, PA: American Society 

for Testing and Materials. 

ASTM D198. Standard methods of static tests of lumber 

in structural sizes. 

ASTM D2899. Standard practice for establishing design 

stresses for round timber piles 

ASTM D2555. Standard test methods for establishing 

clear wood strength values. 

Bodig, J.; Goodman, J.R.; Phillips, G.E.; Fagan, G.B. 

1986. Wood pole properties, Vol. 2. Palo Alto, CA: Electric 

Power Research Institute. 

Phillips, G.E.; Bodig, J.; Goodman, J.R. 1985. Wood pole 

properties, Vol. 1. Palo Alto, CA: Electric Power Research 

Institute. 

Wolfe, R. 1989. Allowable stresses for the upside-down 

timber industry. In: Proceedings, International conference on 

wood poles and piles; 1989 October 25–27; Fort Collins, 

CO. Ft. Collins, CO: Colorado State University, Engineering 

Data Management, Inc: D1–D11. 
10 

Appendix A—Pilot Tests of 

Dowel-Nut Connector 

Eight joints were tested to assess their sensitivity to variations in joint configuration. Variables in these tests included 

dowel-nut diameter, type and location of the confinement 

device, specimen moisture content, and specimen length. 

Four test peeler cores were used. The cores were initially 

1.8 m (6 ft) long. After the first series of tests was completed, 

a 0.3-m (1-ft) length was cut from each end of the core, so 

that cores for the second series were 1.2 m (4 ft) long. 

In all cases, the steel dowel nut was centered 0.25 m (10 in.) 

from the end of the specimen. For the first two tests, the steel 

dowel was 31.8 mm (1.25 in.) in diameter. When we observed ductile compression parallel-to-grain failures, we tried 

increasing the size of the dowel to 44.5 mm (1.75 in.) to 

determine whether joint load capacity could be increased 

without changing the mode of failure to tension parallel 

to grain. 

The results of the pilot tests are summarized in Table 1. 

The joint test sequence was as follows: 

1. The first joint test (1A) included a 31.8-mm (1.25-in.) 

dowel-nut connector, a shipping band confinement strap 

fastened with a metal crimp, and a peeler core with a relatively high moisture content (probably 25%–30% at time 

of test). Because of the loose fit, the joint exhibited an 

initial failure at a relatively low load (62.3 kN (14 × 

103

 lbf)). Then, the confinement strap caused the load to 

climb to 70.3 kN (15.8 × 103

 lbf). A second failure 

caused the load to drop to 48.9 kN (11 × 103

 lbf). 

We then reinforced the failed end with a spiral bolt helix 

clamp and reloaded to 62.3 kN (14 × 103

 lbf) before the 

final failure. The failure mode in this case was a combination of tension perpendicular to grain, which caused an 

initial crack, and reduction in size of the effective shear 

plane, followed by shear failure. 

2. The second test (1B) was similar to the first except that a 

spiral bolt helix clamp was used on both ends of the 

specimen. In this case, the maximum load rose to 89 kN 

(20 × 103

 lbf), and the failure mode was a combination of 

compression parallel to grain and tension perpendicular to 

grain. Both ends of the specimen appeared to exhibit approximately the same level of compression. We thought 

that improving confinement by increasing the dowel bearing area would increase joint capacity. 

3. For the third test (8B), the diameter of the dowel nut was 

increased to 44.5 mm (1.75 in.) and a steel strap pipe 

clamp with a “T-bar” bolt type adjustment was tested. 

The pipe clamp was placed just above the dowel-nut hole. 

Because of the taper in the peeler core, we were not able 

to put the same clamp on the larger end, so we used the 

spiral bolt helix clamp. Both ends of the specimen were 

compressed equally. At a load of 93 kN (20.9 × 103

 lbf), 

the smaller end split in tension perpendicular to grain at 

the part beyond the clamp. We reclamped this end and reloaded to 97.14 kN (21.84 × 103

 lbf). This was a ductile 

failure in which the load dropped to 66.7 kN 

(15 × 103

 lbf). 

4. For the fourth test (10B), we placed two strap clamps on 

each end of the specimen. In this case, the failure was 

confined to compression parallel to grain, reaching a 

maximum load of 110.7 kN (24.9 × 103

 lbf) and holding 

that load. We stopped the test at a total deformation in 

excess of 25 mm (1 in.). The remaining tests were conducted on the same peeler cores as those used for the previous tests except that the failed material was removed 

from the specimen (0.3 m (1 ft) from each end) and the 

ends were re-machined. 

5. Results of joint test 1A2 were significantly different than 

those of test 1A1. In this case, the maximum load of 

150.3 kN (33.8 × 103

 lbf) resulted in a combined shear 

and tension parallel-to-grain failure. Each specimen end 

had only one pipe clamp. The upper end, away from the 

active load head, was reinforced with an embedded metal 

plate connector (MPC) (truss plate) to aid in resisting tension perpendicular to grain. In the case of test 1A2, the 

pipe clamps were located 25 mm (1 in.) away from the 

dowel-nut hole, toward the end of the peeler core. Unlike 

the joints in test 8B, joint 1A2 exhibited no tendency to 

split as a result of tension perpendicular to grain. Rather, 

the failure was a combination of shear and tension parallel to grain on the specimen end without the MPC. The 

material was quite a bit drier the second time around, and 

drying checks that opened up played a role in the failure. 

6. Results of the 1B2 test were similar to those of 1A2 

except that a second clamp was added to the end without 

the MPC reinforcement. Maximum load was 143.2 kN 

Table 1—Results of pilot tests of dowel-nut connector 

 

Moisture 

content (%) 

Specific 

gravity 

ID 

Length 

(m (ft)) 

End 

A 

End 

 B 

End 

A 

End 

B 

Max load 

(kN (×103 lbf)) 

1A1 1.8 (6) — — 0.43 0.47 70.3 (15.8) 

1A2 1.2 (4) 9 11 — — 150.3 (33.8) 

1B1 1.8 (6) — — 0.41 0.39 89.0 (20.0) 

1B2 1.2 (4) 9 11 — — 143.2 (32.2) 

8B1 1.8 (6) — — 0.40 0.41 97.1 (21.8) 

8B2 1.2 (4) 20 18 — — 118.7 (26.7) 

10B1 1.8 (6) — — 0.45 0.47 110.7 (24.9) 

10B2 1.2 (4) 18 19 — — 138.0 (31.0) 
11 

(32.2 × 103

 lbf). The failure was the same as that observed for test 1A1—combined shear and tension parallel 

to grain at the dowel-nut hole. This failure occurred on 

the specimen end with the two pipe clamps. 

7. The retest of peeler core 8B was also similar to that of 

peeler core 1A except that the MPC end was placed on 

the bottom, toward the active load head. In this case the 

MPC end failed at a load of 118.7kN (26.7 × 103

 lbf). 

8. For the 10B retest, strap clamps were located 25 mm 

(1 in.) outside the dowel connector holes, MPC was on 

the active load-head end, and the clamp was on the opposite end. Failure at 138 kN (31 × 103

 lbf) occurred on the 

upper end of the specimen. The mechanism of failure was 

shear and tension parallel to grain. 

The shear and tension-parallel-to-grain failure mode is less 

desirable than the compression-parallel-to-grain mode because it is more brash, causing a sudden drop in joint capacity. This seems to be influenced primarily by the change in 

moisture content. As the compression parallel-to-grain 

strength of wood almost doubles from green to dry, it is 

possible that increasing the dowel-nut diameter caused the 

failure mode to shift to tension parallel to grain in the dry 

material. 

Appendix B—Test Results

Tables 2 and 3 give the results of tension tests of the 

dowel-nut connection in 127-mm- (5-in.-) diameter Douglasfir peeler cores and compression tests of full-size sections, 

respectively. The joint tests (Table 2) are described in the 

main body of this report. Tables 3 and 4 compare tests conducted in green and dry conditions, respectively. The “slope” 

value listed in these tables refers to the slope of the load–

displacement curve determined by linear regression over the 

displacement range of 20% to 40% of the maximum load. 

The elastic limit value was determined as the point where the 

linear regression deviated from the measured load by more 

than 1%. In most cases, the load–displacement curves for the 

two end-joints were almost parallel, to the point where one 

began to fail. Figure 10 shows the extreme example where 

the slopes ranged from 96 to 189 kN/mm (550 to 1,080 

× 103

 lbf/in). Note that beyond 5 kN/mm (29 × 103 lbf/in), 

both joints exhibited accelerating deformation. 

Compressive load–displacement slopes and maximum loads 

are shown in Tables 3 and 4. Results reported in Table 3 are 

from samples tested at varying moisture content levels, all 

greater than 20%. Results reported in Table 4 are from samples tested at moisture content under 12%. All results were 

adjusted to 12% moisture content using linear interpolation 

dry/green ratio derived for moisture content changes from 

30% to 12% (1.78 for ultimate compressive stress and 

−1.25 for modulus of elasticity (ASTM D2555)). 

Table 2—Results of dowel-nut connection tests in Douglas-fir peeler cores 

ID 

Growth ratea

 

(r/cm (rpi)) 

Pith 

index SG 

MC 

(%) 

Maximum load 

(kN (×103

 lbf)) 

Slope 

 (kN/mm (×103 lbf/in)) 

Elastic limit 

(kN (×103 lbf)) Commentb

J-5-1 2 (2.0) 0.50 0.44 10 101 (22.6) 158.5 (905) 89 (20) Joint knot affects parallel tension 

J-5-2 3 (7.5) 0.14 0.41 10 159 (35.7) 159.7 (912) 142 (32) Shear and tension perp to grain 

J-5-3 5 (12.3) 0.57 0.45 10 165 (37.1) 155.3 (887) 151 (34) Shear and tension perp to grain 

J-5-4 2 (5.2) 0.60 0.37 8 126 (28.2) 120.5 (688) 111 (25) Shear 

J-5-5 3 (7.0) 0.14 0.41 10 106 (23.9) 173.3 (990) 89 (20) Tension perp to grain 

J-5-6 3 (7.5) 0.67 0.43 11 123 (27.6) 146.7 (838) 120 (27) Shear and tension perp to grain 

J-5-7 3 (7.3) 0.54 0.40 10 117 (26.4) 134.3 (767) 116 (26) Tension parallel to grain +shear 

J-5-8 4 (11.3) 0.55 0.49 12 164 (36.9) 186.7 (1,066) 133 (30) Shear 

J-5-9 2 (6.3) 0.29 0.45 10 131 (29.5) 204.8 (1,170) 129 (29) Shear 

J-5-10 4 (10.0) 0.20 0.45 10 108 (24.4) 130.6 (746) 107 (24) Tension parallel to grain 

J-5-11 2 (4.5) 0.09 0.44 7 102 (23.0) 187.9 (1,073) 102 (23) Tension parallel to grain + shear 

J-5-12 4 (9.0) 0 0.49 11 161 (36.2) 131.0 (550) 129 (29) Tension parallel to grain 

J-5-13 3 (8.0) 0.28 0.4 10 139 (31.3) 113.6 (649) 125 (28) Tension parallel to grain + shear 

J-5-14 2 (5.3) 0.31 0.43 10 102 (23.0) 164.2 (938) 102 (23) Shear + tension perp to grain 

J-5-15 2 (5.7) 0.21 0.43 10 137 (30.8) 150.2 (858) 120 (27) Brash tensile failure parallel to 

grain; splintery latewood 

J-5-16 3 (8) 0.85 0.43 10 127 (28.6) 114.9 (656) 120 (27) Compression parallel to grain; 

ductile tension perp to grain 

Average 0.4 7.7 131 (29.1) 151.6 (856) 120 (26.5) 

COV 7% 53% 16% 17% 15% 18% 8% 14% 

5th percentile 0.4 0.8 95 (20.5) 114.0 (596) 103 (20.3) 

a

rpi designates rings per inch; r/cm, rings per centimer. 

b

Parallel designates parallel to grain; perp, perpendicular to grain. 
12 

Table 3a—Summary of compression parallel-to-grain test results for green Douglas-fir peeler coresa—SI units

ID 

Growth 

rate 

(r/cm) 

Pith 

index SG 

MC 

(%) 

A

(cm2

) 

Slope 

(kN/mm) 

Pmax 

(kN) 

UCS 

(MPa) 

MOE 

(GPa) 

UCS @ 

12% MC 

(MPa) 

MOE @ 

12% MC 

(GPa) 

C-5-1 3 0.45 0.44 25 3.04 120.8 278 22 4.8 37 5.7 

C-5-2 3 0.22 0.39 24 3.43 112.3 309 22 4.0 35 4.7 

C-5-3 2 0.39 0.39 23 3.22 102.8 292 22 3.9 34 4.5 

C-5-4 2 0.50 0.42 23 2.98 91.9 262 21 3.7 33 4.3 

C-5-5 2 0.20 0.41 23 3.16 121.9 322 24 4.7 38 5.4 

C-5-6 3 0.50 0.45 22 3.30 92.8 326 24 3.4 36 3.9 

C-5-7 3 0.90 0.42 21 3.08 56.6 216 17 2.2 25 2.5 

C-5-8 3 0.95 0.38 22 3.22 81.8 278 21 3.1 31 3.5 

C-5-9 4 0.40 0.40 23 3.29 100.5 261 19 3.7 30 4.3 

C-5-10 3 0.25 0.45 25 3.36 99.8 269 19 3.6 32 4.2 

C-5-11 3 0.27 0.35 23 3.15 107.7 239 18 4.2 28 4.8 

C-5-12 4 0.40 0.46 22 2.90 45.5 228 19 1.9 29 2.2 

C-5-13 4 0.23 0.38 23 3.16 85.8 268 20 3.3 32 3.8 

C-5-14 3 0.31 0.43 22 3.19 124.9 318 24 4.8 36 5.4 

C-5-15 6 0.41 0.42 20 3.43 94.2 314 22 3.4 31 3.7 

C-5-16 2 0.33 0.39 21 3.30 108.4 282 20 4.0 30 4.5 

C-5-17 3 0.56 0.42 24 3.29 87.6 305 22 3.2 36 3.8 

C-5-18 3 0.21 0.37 23 3.16 107.9 292 22 4.1 35 4.8 

C-5-19 3 0.31 0.46 24 3.04 62.9 250 20 2.6 32 3.0 

C-5-20 4 0.86 0.50 23 3.30 108.9 328 24 4.0 37 4.6 

C-5-21 2 0.20 0.45 22 3.12 123.1 342 26 5.8 40 6.6 

C-5-22 2 0.20 0.37 20 3.07 98.8 318 25 4.7 35 5.2 

C-5-23 4 0.80 0.41 20 3.02 102.1 293 23 5.0 33 5.5 

C-5-24 3 0.25 0.39 21 3.10 89.1 271 21 4.2 31 4.7 

Average 0.42 23 3.18 97.0 286 22 3.9 33 4.4 

COV 9% 6% 4% 21% 12% 11% 23% 11% 23% 

5th percentile 0.40 20 2.9 62.0 229 17.6 2.3 27 2.7 

a

Pmax, maximum load; UCS, ultimate compressive stress; MOE, modulus of elasticity. 
13 

Table 3b—Summary of compression parallel-to-grain test results for green Douglas-fir peeler coresa—inch–pound units 

ID 

Growth 

rate 

(rpi) 

Pith 

index SG 

MC 

(%) 

A

(in2

) 

Slope 

(×103 lbf/in) 

Pmax 

(×103 lbf) 

UCS 

(×103 lb/in2) 

MOE 

(×103 lb/in2) 

UCS @ 

12% MC 

(×103

 lb/in2) 

MOE @ 

12% MC 

(×103

 lb/in2) 

C-5-1 8 0.45 0.44 25 19.6 690 62.4 3.18 700 5.24 835 

C-5-2 8 0.22 0.39 24 22.1 641 69.5 3.15 580 5.04 683 

C-5-3 5 0.39 0.39 23 20.8 587 65.7 3.17 570 4.93 664 

C-5-4 6 0.50 0.42 23 19.2 525 59.0 3.07 540 4.78 629 

C-5-5 4 0.20 0.41 23 20.4 696 72.4 3.54 680 5.51 792 

C-5-6 7 0.50 0.45 22 21.3 530 73.3 3.45 500 5.19 575 

C-5-7 8 0.90 0.42 21 19.9 323 48.5 2.43 320 3.54 363 

C-5-8 7 0.95 0.38 22 20.8 467 62.5 3.01 450 4.53 518 

C-5-9 9 0.40 0.40 24 21.2 574 58.6 2.76 540 4.36 633 

C-5-10 8 0.25 0.45 24 21.7 570 60.4 2.78 520 4.54 618 

C-5-11 7 0.27 0.35 22 20.3 615 53.7 2.65 610 4.06 706 

C-5-12 9 0.40 0.46 22 18.7 260 51.2 2.74 280 4.13 322 

C-5-13 9 0.23 0.38 23 20.4 490 60.3 2.95 480 4.60 560 

C-5-14 8 0.31 0.43 22 20.6 713 71.6 3.48 690 5.26 795 

C-5-15 14 0.41 0.42 20 22.1 538 70.5 3.20 490 4.45 547 

C-5-16 5 0.33 0.39 21 21.3 619 63.3 2.97 580 4.37 661 

C-5-17 8 0.58 0.42 24 21.2 500 68.6 3.23 470 5.16 553 

C-5-18 7 0.21 0.37 23 20.4 616 65.6 3.21 600 5.00 700 

C-5-19 8 0.31 0.46 24 19.6 359 56.1 2.87 370 4.55 434 

C-5-20 11 0.86 0.50 23 21.3 622 73.7 3.46 580 5.45 679 

C-5-21 5 0.86 0.45 22 20.1 703 76.9 3.82 840 5.66 960 

C-5-22 5 0.20 0.37 20 19.8 564 71.5 3.62 680 5.13 765 

C-5-23 9 0.80 0.41 20 19.5 583 65.8 3.38 720 4.79 808 

C-5-24 8 0.25 0.39 21 20.0 509 60.9 3.04 610 4.47 695 

Average 0.41 22 20.5 554 64 3.13 558 4.78 646 

COV 9% 6% 4% 21% 12% 11% 23% 11% 23% 

5th percentile 0.35 20 18.99 355.31 51.3 2.56 336 3.90 391 

a

Pmax, maximum load; UCS, ultimate compressive stress; MOE, modulus of elasticity. 
14 

Table 4a—Dry compression parallel-to-grain tests of samples cut from joint test specimensa—SI units 

ID 

Growth 

rate 

(r/cm) 

Pith 

index SG 

 MC 

(%) 

cxb

(cm2) 

Slope 

(kN/mm) 

Pmax 

(kN) 

UCS 

(MPa) 

MOE 

(GPa) 

UCSc

(MPa) 

MOEc

(GPa) 

J-5-3 5 0.57 0.45 10 132.81 135.8 412 31 5.2 30 5.0 

J-5-4 2 0.60 0.37 8 132.29 134.6 385 29 5.2 26 4.9 

J-5-5 3 0.14 0.41 10 116.23 85.3 319 27 3.7 26 3.6 

J-5-6 5 0.57 0.45 10 127.65 134.1 410 32 5.3 31 5.2 

J-5-7 3 0.54 0.40 10 123.65 98.9 331 27 4.1 26 4.0 

J-5-8 3 0.14 0.41 10 140.16 147.3 437 31 5.3 30 5.2 

J-5-9 3 0.67 0.43 0 132.29 159.3 445 34 6.1 32 6.0 

J-5-10 3 0.54 0.40 0 126.16 119.5 354 28 4.8 27 4.7 

J-5-11 4 0.55 0.49 0 125.13 172.7 455 36 7.0 32 6.5 

J-5-12 2 0.29 0.45 10 126.16 86.2 355 28 3.5 27 3.4 

J-5-13 4 0.20 0.45 10 130.23 122.4 351 27 4.8 26 4.6 

J-5-14 2 0.09 0.44 7 130.23 171.6 495 38 6.7 34 6.2 

J-5-15 4 0 0.49 11 126.61 131.5 388 31 5.3 30 5.2 

J-5-16 3 0.28 0.40 10 129.71 135.0 393 30 5.3 29 5.1 

Average 0.40 7.3 127.80 130.3 394.5 31 5 29 5.0 

COV 8% 58% 4% 21% 13% 11% 20% 9% 18% 

5th percentile 0.37 −0.14 118.13 81.6 304.47 24.73 3.4 24.30 3.4 

a

Load assumed to be uniformly distributed to cross section. Strain gauge length in all cases was 508 mm (20 in.). 

b

Cross section. 

c

Data for UCS and MOE were adjusted to 12% MC. 
15 

Figure 10—Load–displacement curves for end joints. 

Table 4b—Dry compression parallel to grain tests of samples cut from joint test specimensa—inch–pound units 

ID 

Growth 

rate 

 (rpi) 

Pith 

index SG 

MC 

(%) 

cxb

(in2) 

Slope 

(×103 lbf/in) 

Pmax 

(×103 lbf) 

UCS 

(×103 lb/in2) 

MOE 

(×103 lb/in2) 

UCSc

(×103 lb/in2) 

MOEc

(×103 lb/in2) 

J-5-3 12 0.57 0.45 10 20.59 776 92.7 4.50 750 4.30 732 

J-5-4 5 0.6 0.37 8 20.51 769 86.6 4.22 750 3.84 708 

J-5-5 8 0.14 0.41 10 18.02 487 71.8 3.98 540 3.80 526 

J-5-6 12 0.57 0.45 10 19.79 766 92.1 4.65 770 4.44 752 

J-5-7 7 0.54 0.4 10 19.17 565 74.3 3.88 590 3.70 573 

J-5-8 7 0.14 0.41 10 21.73 841 98.2 4.52 770 4.31 753 

J-5-9 8 0.67 0.43 10 20.51 910 100.2 4.88 890 4.66 863 

J-5-10 7 0.54 0.40 10 19.56 682 79.7 4.07 700 3.89 678 

J-5-11 11 0.55 0.49 7 19.4 986 102.4 5.28 1,020 4.68 946 

J-5-12 6 0.29 0.45 10 19.56 492 79.8 4.08 500 3.90 489 

J-5-13 10 0.2 0.45 10 20.19 699 78.9 3.91 690 3.73 673 

J-5-14 4 0.09 0.44 7 20.19 980 111.2 5.51 970 4.31 904 

J-5-15 9 0 0.49 11 19.63 751 87.3 4.45 760 4.35 755 

J-5-16 8 0.28 0.40 10 20.11 771 88.4 4.40 770 4.20 745 

Average 0.40 9.6 19.80 744 88.7 4.47 748 4.21 721 

COV 8% 13% 4% 21% 13% 11% 20% 9% 18% 

5th percentile 0.40 7.41 18.32 465.9 68.4 3.59 — 3.52 487.80 

a

Load assumed to be uniformly distributed to cross section. Strain gauge length in all cases was 508 mm (20 in.). 

b

Cross section. 

c

Data for UCS and MOE were adjusted to 12% MC. 
16 

Appendix C—Design of Dowel-Nut Connectors 

This appendix provides a derivation of estimates of the average strength of dowel-nut connectors in Douglas-fir peeler cores and 

a recommended nominal design value based on published strength values (ASTM D2555 (ASTM 1998)) and conventional design derivation procedures (AF&PA National Design Specification (ANSI/AF&PA 1997) and ASTM D245 and ASTM D1761 

(ASTM 1993b and ASTM 1995, respectively). 

Assumed joint parameters 

Nominal diameter of dowel nut Dd = 1.75 in. 

Nominal diameter of peeler core Dc = 5 in. 

Bearing length in timber (core diam–threaded rod hole diam) 1 = Dc − 1 in. 

Bearing area, projected Ab = DdDc 

Tensile area At = π(0.5 Dc)

2 − Ab 

Dowel center–end distance NDc = 10 in. 

Shear area Ashr = (NDc − 0.5 Dc)Dc 

Bearing parallel-to-grain failure 

Following the NDS guideline for dowel (bolted) connections loaded parallel to grain at a reduced end distance (<7D), the expected average strength of the connection would be as follows: 

Dowel bearing strength (NDS Table 8A) Fes = 5,500 lb/in2

Expected ultimate bearing strength of test joint Fes Dd1

d

c

7D

ND

= 3.143 × 104

 lb/in2 

Nominal design value (NDS) Z = Dd1(Fes/4) Z = 9.625 × 103 lb/in2

Design for wind Zw = Z 1.15 lb Zw = 1/107 × 104

 lb/in2 

Tension parallel-to-grain failure 

NDS design stress for Douglas-fir North Construction grade 4 by 4 Ft = 950 lb/in2 

Design load capacity for 10-year dry use Ft At = 1.034 × 104 lb/in2 

Expected average ultimate strength 

 ASTM D245, assuming COV of 20% 1 1.645(0.2)

t 2.1 t

−

F A

 = 3.23 × 104

 lb/in2 

Shear parallel-to-grain failure 

This type of failure appears as a block pushed between two shear planes. 

Shear area—two shear planes Av = 2Ashr 

NDS design shear stress for Douglas-fir Fv = 95 lb/in2

 North Construction grade 4 by 4 

Design load capacity based on sawn timber Fv Av = 7.125 × 103

 lb/in2 

Green timber design value based on ASTM D2899 Fvg = 947 ⋅ 5.47

(1−1.645[0.13])

(The factor 1/5.47 consists of a 0.75 adjustment for SR 

 and 1/4.1 for DOL and FS) Fvg = 136.103 lb/in2 

Shear strength, average at 12% MC (ASTM D2555) Fvd = Fvg1.48 

 Fd = 201.433 lb/in2 

Load capacity Fvd Av = 1.511 × 104 lb/in2 
17