Designated Fiber Stress 

for Wood Poles

Ronald W. Wolfe

Robert O. Kluge

United States

Department of

Agriculture

Forest Service

Forest

Products

Laboratory

General

Technical

Report

FPL–GTR–158
Abstract 

Wood poles have been used to support utility distribution 

lines for well over 100 years. Over that time, specifications 

for a “wood utility pole” have evolved from the closest 

available tree stem more than 15 ft in length to straight, 

durable timbers of lengths ranging up 125 ft and base diameters of as much as 27 in. The continued success of wood 

poles in this application is due in part to the development of 

consensus standards. These standards define the phrase 

“minimum acceptable” to the satisfaction of both users and 

producers. They also encourage more competitive pricing by 

relaxing species as well as quality limitations, opening the 

market to a broader range of available timber resources. The 

American National Standards Institute (ANSI) standard 

ANSI O5.1 is an internationally recognized standard that has 

served as a guide for selecting the quality and size of wood 

utility poles for more than 70 years. From its inception, this 

standard has addressed issues of relative load capacity as 

well as physical quality to allow for species substitutions. In 

2002, the relative strength evaluations previously published 

as a designated fiber stress took on added meaning when 

they were defined to represent the mean of the distribution of 

pole groundline strength values for various species. The 

change in meaning was accompanied by a more rigorous 

evaluation recognizing a change in strength with height and 

notation that pole strength distributions have a coefficient of 

variation of 20%. This paper reviews the history and philosophy of the ANSI designated fiber stress to help the 

reader more fully understand and appreciate the significance 

of changes adopted by the American Standards Committee 

O5 (ASC O5) in 2002. 

Keywords: pole, ANSI standard, fiber stress, strength, 

history 

June 2005 

Wolfe, Ronald W.; Kluge, Robert O. 2005. Designated fiber stress for wood 

poles. Gen. Tech. Rep. FPL-GTR-158. Madison, WI: U.S. Department of 

Agriculture, Forest Service, Forest Products Laboratory. 39 p. 

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

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

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University of Wisconsin. 

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Abbreviations 

ANSI American National Standards Institute 

ASA American Standards Association 

ASC O5 Accredited Standards Committee O5 

ASTM American Society for Testing and Materials 

COV coefficient of variation 

CSA Canadian Standards Association 

DFS designated fiber stress 

EPRI Electric Power Research Institute 

LRFD load-and-resistance-factor design 

MOR modulus of rupture 

NESC National Electrical Safety Code 

REA Rural Electrification Administration 

SI conversion factors 

English unit 

Conversion 

factor SI unit 

inch (in.) 25.4 millimeters (mm) 

square inch (in2) 6.45 square centimeter (cm2) 

foot (ft) 0.3048 meter (m) 

pound per square inch 

(lb/in2

) 

6.895 kilopascal (kPa) 

temperature (°F) [TF − 32]/1.8 temperature (°C) 

Contents 

Page 

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

Objectives ............................................................................ 2 

Background.......................................................................... 2 

Pole Strength........................................................................ 5 

 First National Standard Specification for Wood Poles ... 6 

 Subsequent Standards for Wood Poles ........................... 6 

 Addition of H-Class ...................................................... 10 

Wood Pole Test Data ......................................................... 10 

 ASTM Wood Pole Test Program.................................. 10 

 EPRI Wood Pole Test Program .................................... 11 

Nominal Strength ............................................................... 11 

 Initial Estimates............................................................. 11 

 Test Data and In-Service Performance ......................... 13 

 ANSI O5-02.................................................................. 19 

Conclusions........................................................................ 20 

Observations ...................................................................... 21 

Literature Cited .................................................................. 22 

Appendix A—Effect of 2002 Revision on Pole-Class 

Strength Distribution.......................................................... 24 

Appendix B—Wood Pole Test Data Sources .................... 26 

Appendix C—Data Used to Derive DFS for Long 

Poles................................................................................... 27 

Appendix D—Class Oversize Adjustment......................... 39 
Designated Fiber Stress for Wood Poles 

Ronald W. Wolfe,1 Research Engineer

Forest Products Laboratory, Madison, Wisconsin

Robert O. Kluge,2 Team Leader, Design Standards

American Transmission Company, Madison, Wisconsin 

Introduction 

Round timbers have been used structurally for centuries. 

Over the past 140 years, round timbers have been used 

primarily for utility structures. As in so many other areas of 

timber use, increased demand and competition from other 

materials have spurred development and efficiency in the 

structural use of poles. There is little doubt that Alexander 

Graham Bell and Thomas Alva Edison played key roles in 

initiating demand, but it is the influence of consensus standards (preservative treatment, strength tests, and material 

specification) that has enabled wood to retain a dominant 

role in support structures for telecommunications and electric power utilities. 

Prior to 1930, there was little demand for standards that 

rigorously addressed issues of the “design” capacity of wood 

poles. The National Bureau of Standards (1927) began development of the National Electrical Safety Code (NESC) 

for this purpose in 1913. Early concerns focused on durability and cost. American chestnut and western redcedar were 

preferred for their durability. The second edition of the 

NESC (1916) recognized southern pine as a viable pole 

species after the development of creosote treatment. In that 

edition, southern pine was listed along with western redcedar 

and chestnut as having an ultimate bending strength of 

5,000 lb/in2

. By 1927, northern white cedar was added, with 

an ultimate strength of 3,600 lb/in2

, and the bending strength 

of “dense” southern pine was rated at 6,500 lb/in2

. By this 

time, accumulating knowledge and experience related to the 

strength of wood poles indicated that somewhat higher 

values might be allowed (Wilson 1923). A Sectional Committee on Wood Poles was formed by the American Standards Association (ASA) to evaluate available data, and the 

NESC made provisions for use of new values as approved 

by the ASA. Decisions by this committee are discussed in 

detail by Colley (1932). 

____________________________________________ 

1

Former chairman of ASC O5. 

2

Former chairman of ASC O5 fiber stress subcommittee. 

In June 1931, standard specifications for wood poles were 

approved by the ASA (Jones 1931). These standards incorporated a pole classification system originally developed by 

the American Bell Telephone Company in a standard specification aimed at maintaining uniform reliability for pole 

structures, despite a range of species. The standards quantified pole capacity using existing cantilever test data to assess 

groundline stress at failure for each recognized pole species. 

This value was then used to determine pole groundline dimensions required for each pole length and species. While 

changes have been made to the designated fiber stress (DFS) 

values referenced by the ASA standard, this basic approach 

to classifying poles remained unchanged for over 70 years. 

As utility structure loads and importance increased, design 

became more of an issue. Most utility companies began 

using larger poles by the early 1960s, but they attracted little 

attention in the American National Standards Institute 

(ANSI) standard until 1972 when “H-class” pole sizes were 

introduced. Shortly thereafter, greater attention focused on 

the fact that the largest poles tested to that point were 55 ft in 

length with a butt diameter of 19 in., while the standard 

implied a knowledge of capacities for poles up to 125 ft long 

with a 30-in. butt diameter. Research conducted over the 

next 20 years pointed to the need for change to the traditional pole classification algorithm. As larger poles were 

shown to have a lower groundline stress at failure, the need 

to incorporate a pole “size” adjustment in the algorithm 

became apparent. This issue was added as an annex to the 

ANSI standard in 1986, but no changes were made to the 

pole size classification tables. 

In the late 1990s, discussions about revising the NESC 

recommendations for wood pole design focused on 

discrepancies between the results tabulated in the ANSI 

annex and DFS reported in the standard. The Accredited 

Standards Committee O5 (ASC O5), which was responsible 

for maintaining the ANSI O5.1 standard, adopted a more 

rigorous definition of the ANSI DFS, which enabled them to 

indirectly recognize a size effect without changing the existing pole size classification tables. Redefining DFS to be the 
2 

actual groundline strength rather than the groundline stress at 

which a pole would fail and recognizing a change in strength 

with height basically results in reductions in the load capacity of poles where the failures are likely to be well above 

groundline. While the recommended strength change with 

height is easy to apply, it does not recognize variations 

between species, and that detracts from the premise of uniform reliability across species.3

 

The DFS values are all derived to represent “dry” strength, 

yet tabulated dimensions are explicitly declared to be 

“green” dimensions (sec 5.2 ANSI O5.1–1986). The standard suggests a 2% reduction in circumference for drying, 

which corresponds to a 6% reduction in section modulus or 

dry moment capacity. For poles that remain above fiber 

saturation in the groundline region, the “dry” strength gives 

a estimated moment capacity that is 86% (1/1.16) of the 

DFS. No adjustment is suggested for this situation. 

Slight inconsistencies in the development of size classification are to be expected in a material specification. This has 

little impact on design when the design standards are independent of the material specification. In 2002, however, the 

ANSI O5.1 DFS values were considered to represent a basis 

for probability-based design of utility structures. The standard includes a statement about the DFS values that implies 

some degree of certainty that these truly represent average 

groundline strength values. It also states that poles have a 

bending strength coefficient of variation (COV) of 20% 

independent of species, species grouping, or size. 

Changes made to standard definitions to justify changes to 

the standard could have some impact on the assumed reliability of existing as well as future utility structures. It is 

therefore essential that some documentation be provided to 

explain the basis for the past as well as the current interpretation of ANSI fiber stress values. 

Objectives 

This paper reviews the history of the ANSI O5.1 standard 

specification for wood poles from the time that poles were 

first used as utility structures up to 2002, when the ANSI 

standard made its initial transition from being just a material 

specification to being a pole design standard. First, we trace 

the growth of the wood pole industry in response to rapidly 

changing technology in the telecommunications and electric 

 

3

 The pole specification was originally formulated by the 

American Standards Association (ASA) and was referred to 

as ASA O5.1. In 1972, the ASA was renamed the American 

National Standards Institute (ANSI). The designation for the 

pole standard was changed to ANSI O5.1 and the standard 

was managed by the Accredited Standards Committee O5 

(ASC O5). 

power utility industries. We next describe the evolution of 

the standard material specification, with emphasis on the 

evaluation and use of DFS values. Finally, we provide a 

summary and rationale for strength adjustments used to 

convert values derived from standard test procedures to 

values considered representative of utility poles. 

The objective of this review is to provide documentation for 

future generations of the decisions made by consensus committees that have significantly influenced the evolution of 

the ANSI O5.1 standard. The change made in 2002 was the 

most significant of the standard’s 70-year history. This paper 

is intended to highlight the significance of that change by 

offering a “before and after” glimpse. 

Background 

For over 140 years, wood poles have provided a low-cost 

solution to the problem of supporting utility wires for telecommunications and electric power. Wood poles presented a 

readily available resource when the first transcontinental 

telegraph line was constructed in 1861. By 1900, wood poles 

were being used to support electric power as well as telephone lines. Over time, utility companies gained experience 

at selecting poles that best met their needs for durability and 

load capacity. 

The growing demand for durable wood poles spawned the 

need for nationally recognized standards. Bell Laboratories 

initiated efforts to develop test standards and material specifications for wood poles. In 1913, the National Bureau of 

Standards initiated the NESC to support design consistency 

for utility supply and communication installations throughout the United States. This was the first standard to publish 

pole strength values. In 1915, the American Society for 

Testing and Materials (ASTM) published the first standard 

material specification for round timber (ASTM D 25). This 

standard was intended for timbers to be used as piling but, 

except for the larger allowable knots and more lenient requirements for spiral grain afforded poles, it was identical to 

the standard specification for wood poles proposed by Bell 

Laboratories that was adopted as a national standard 

15 years later. 

In 1924, the American Bell Telephone Company and the 

U.S. Independent Telephone Association’s American Standards Association (ASA) formed a Sectional Committee on 

Wood Poles. This committee used average pole strengths 

from the Bell Laboratories database to establish an interchangeable species–pole-size classification system. The Bell 

Laboratories strength data were obtained using a cantilever 

bending test that eventually was adopted by the ASTM and 

designated ASTM D 1036–49. This test provided estimates 

of relative groundline strength for each species. The ASA 

committee adopted a 10-class system comprising a series of 

tip load capacities increasing in increments of 25%. 

Groundline moments corresponding to these loads, divided 
3 

by the relative strengths determined from existing data, 

provided the groundline section property requirements for 

each pole species–class–length combination. 

In 1931, the ASA adopted and published the first standard 

specification for wood poles. This standard contained minimum pole quality specifications as well as minimum required dimensions for 10 pole classes. The first seven of 

these classes were considered structural; classes 8, 9, and 10 

were nonstructural. At that time, only four species were 

accepted for utility poles: western redcedar, Douglas-fir, 

southern pine, and northern white cedar. As new species and 

new data for accepted species were introduced, the committee refined its analysis of assigned DFS values and size class 

tables. 

President Roosevelt’s introduction of the Rural Electrification Administration in 1935 provided a definite impetus to 

the market for wood utility poles. By the end of World War 

II, annual production had grown to 8.1 million poles. To 

meet the growing demand, the ASA standard accepted a 

number of minor species in 1948. As there were no pole data 

for these species, the ASA committee assigned DFS values 

on the basis of published small–clear bending strengths 

adjusted for a relationship between full-size poles and small–

clear (ASTM D 143, ASTM 2000) test values determined for 

lodgepole pine. Questions about the selection of these DFS 

values prompted ASTM to sponsor a wood pole research 

project, beginning in 1955, to assess the bending moment 

capacity of distribution poles. The final report of that program (Wood and others 1960) became known as the ASTM 

wood pole report, and it formed the basis for revision of the 

ANSI DFS values in 1963. 

Large poles used for transmission structures became more of 

an issue with the introduction of six larger H-classes in 

1972. These structures were used to support high voltage 

power transmission lines. The greater potential costs and 

safety hazard resulting from failure of these structures 

prompted design engineers to apply more rigorous design 

standards than had traditionally been used for distribution 

poles. The addition of these larger pole classes sparked 

discussion about the efficacy of the existing DFS values that 

had been supported by tests of smaller poles, most of which 

failed close to the groundline. One result of this discussion 

was the addition of a strength–height function introduced as 

Annex A to the ANSI O5.1 standard in 1979. 

In the 1980s, a second major research project, sponsored by 

the Electric Power Research Institute (EPRI), was conducted 

to assess the load capacity of large poles. An initial effort 

was made as part of this research (Wolf 1979) to summarize 

published full-size pole test data in the form of a computerized database. Results of the ASTM- and EPRI-sponsored 

studies were compiled along with historic data. Statistical 

analyses of these data formed the basis for the probabilitybased-design guidelines that were added as Annex C to the 

ANSI O5.1 standard in 1986. 

Wood maintained its dominance as a utility pole material 

throughout the 20th century but gradually lost market share. 

As power utilities built higher voltage lines with large conductors and longer span, more steel came into use. Spincasting of steel-reinforced high strength concrete poles also 

had a major impact on markets for wood transmission structures. Poles fabricated from these materials are perceived by 

structural engineers as being more precisely engineered and 

therefore more reliable than wood poles. 

The shift to steel and concrete in the more highly engineered 

transmission structures was accompanied by recognition 

among engineering professionals of the benefits of promoting the use of a load–resistance factor design (LRFD). This 

design methodology, when tied to a statistical or reliability 

basis, generally shows greater benefit for materials with 

lower variability when used in non-redundant systems. A 

move by the NESC to adopt an LRFD design basis increased 

industry pressure for better, more reliable strength data. 

Efforts initiated by competing interests who supported the 

NESC led to a reevaluation of the ANSI O5.1 standard. 

Options debated by the ASC O5 included the following: 

1.Retain the philosophy of classifying poles according to 

their load capacity under standard test conditions and recognize size effect and strength disparities highlighted by 

EPRI research (Phillips and others 1985, Bodig and others 

1986a,b). 

2. Adopt a more rigorous analysis that attributes observed 

disparities to taper differences and height effects, and classify poles according to assumed groundline strength. The 

proposed analysis supported a groundline strength equal to 

the historic DFS for all species, but required designers to 

recognize a change in strength with height. 

Bodig and others (1986a,b) attempted to evaluate the 

strength of poles by assigning an ANSI class to each test 

pole and evaluating load capacity on the basis of ANSI 

minimum dimensions. This procedure accounted for effects 

of “over sizing” in that each test pole within a given class 

was assumed to have the dimensions dictated by minimum 

butt and tip circumferences and a linear taper. Pole load 

capacity, characterized as the groundline stress at failure, 

was assumed to provide a conservative estimate of strength 

at any location, as assumed dimensions were generally found 

to be conservative. This approach showed the ANSI DFS 

values to provide an appropriate estimate of mean strength 

for distribution poles and for Douglas-fir and southern pine 

transmission poles. It also indicated a slight size effect, 

varying by species, which made the DFS values unconservative when applied to western redcedar transmission poles. 
2 

In 2002, the premise of a size effect in poles prompted debate over possible changes to the DFS values. ASC O5 

accepted the explanation that large taper in western redcedar 

transmission poles caused them to fail in weaker wood well 

above groundline. Projecting failure location strength to the 

groundline, then adjusting for drying in service, conditioning, and a class-oversize effect justified leaving the ANSI 

DFS value unchanged. Adopting a change in strength with 

height adjustment to explain the apparent lower groundline 

capacity for larger poles when tested in the standard cantilever configuration required acknowledging a reduction in 

strength with height in the standard. It was felt that this 

reduction in strength with height would have little effect on 

the strength of wood transmission poles as the centroid of 

transverse loads in service is generally more than 2 ft from 

the tip, placing maximum stress closer to the groundline. 

Gravity loads, however, have the opposite effect: they tend 

to move the location of maximum stress higher on the pole, 

closer to the expected location of failure in the standard test. 

Table 1 compares values derived for poles over 50 ft in 

length using the traditional groundline stress-at-failure 

analysis and the approach adopted by the ANSI committee in 

2002. The data for this comparison are primarily from the 

EPRI-sponsored research conducted in the 1980s (Goodman 

and others 1981, Phillips and others 1983, Bodig and others 

1986a,b). All poles 50 ft and longer from these studies are 

included in this analysis. Table 1 basically compares the pole 

load capacity denoted as the groundline stress at failure 

under a standard test load to an estimate of groundline 

strength extrapolated from the failure location. The modulus 

of rupture at the failure location (MORfl) was determined 

from actual pole dimensions at the point of failure. This 

value was extrapolated to a groundline strength (MORgl) 

using a strength–height model presented in Annex A of the 

1992 standard and included in the main standard in 2002. 

Table 1 highlights the dilemma faced by ASC O5. Values 

shown in this table were all derived for poles ≥50 ft in length 

tested by the ASTM standard D 1036 cantilever test procedure. Groundline stress at failure is determined as the maximum groundline moment divided by the measured 

groundline section modulus. It represents a relative measure 

of pole capacity regardless of failure location. For long 

slender poles, the point of maximum stress (moment/section 

modulus) often occurs above the groundline. Assuming that 

wood pole strength decreases with height, it is possible that 

the point where stress exceeds strength will actually occur 

above the point of maximum stress. These effects make the 

longer pole appear weaker in terms of groundline stress at 

failure. 

Rather than change the tabulated DFS values and pole dimensions in response to test results, the ASC O5 chose a 

new definition of DFS that required no change. The new 

definition identifies DFS as the groundline strength of an 

ANSI-classified, dry, conditioned pole. This new definition 

was deemed to have little effect on distribution size poles as 

they generally failed close to the groundline, making the 

previous definition of groundline stress at failure synonymous with groundline strength. For poles in lengths 50 ft and 

longer, the groundline strength was derived considering 

change in strength from the failure location using a model 

originally proposed by Bohannan (1971) to model strength 

reduction with height. Other changes included a maximum 

increase of 10% for drying in service, compared with 16% 

referenced in 1963, and an adjustment for class oversize 

effect (see Appendix D). 

The change in definition of DFS solved one problem but 

created others. Drying and conditioning adjustments to the 

DFS on their own are not new; class oversize was not previously used to adjust the DFS values, but it was considered to 

present inherent conservatism in the sizing of poles. These 

factors took on new meaning, however, when the definition 

of DFS changed. The ANSI standard declares tabulated pole 

dimensions to be measured in the “green” condition. For a 

material specification, the 6% change from green to dry 

section modulus has little impact as these are treated as 

relative strength values. The dry/green section modulus ratio 

varies slightly between species, having little impact on relative pole dimensions when the largest pole in a class has a 

25% greater section modulus than the smallest. In 2002, 

however, the change in definition of DFS from a groundline 

stress at failure of the pole to groundline strength gave these 

adjustments new meaning. The fact that the dry strength, 

derived as 1.16 times green strength, was used to derive 

dimensions to be measured in the “green” condition means 

that the “green” section modulus is 16% below that required 

to carry the designated class load. This corresponds to a 63% 

probability of failure (see Appendix A). The class oversize 

adjustment added for poles 50 ft and longer in the 2002 

standard increased that probability of failure at the ANSIdesignated load. This means DFS applied to longer poles 

infers a lower reliability for transmission poles than for 

distribution poles 

The ANSI standard suggests that dry circumference will be 

2% less than that measured in the green condition. This 

corresponds to a 6% lower section modulus. To carry its 

class load using dry strength, a pole of average strength 

sized in the green condition will therefore have a load capacity 6% below the implied capacity. Assuming a 14% strength 

coefficient of variation (Wood and Markwardt 1965), this 

translates to an implied capacity that is 0.42 standard deviations above the average strength, corresponding to a 66% 

probability of failure. As the moment capacity decreases 

faster than the moment with pole height, this probability of 

failure is likely to increase slightly toward mid-height. 

4 
5 

Pole Strength 

Early evaluations of pole strength focused more on the 

comparison of wood species than on the selection of wood 

over other materials. It was not until the mid- to late 1960s, 

when the demand for power required larger, more reliable 

structures to support high voltage lines over long distances 

and through major metropolitan areas, that structural engineers became concerned about rigorous design standards 

for wood poles. 

One of the earliest public references to pole strength was 

the NESC. In 1915, the second edition of the NESC listed 

three species: creosote treated southern pine, American 

chestnut (no longer available), and western redcedar. All 

three species were denoted as having the same bending 

strength (5,000 lb/in2

). As species options increased, pole 

users realized that not all species have the same strength 

and that they could not simply replace a pole of one species 

with the same size pole of another species. So, between 

1920 and 1930, awareness of the need to classify poles by 

bending load capacity increased. In 1920, the NESC (3rd 

edition) adopted the ASTM definition of “Dense” (6 rings/ 

inch and 33% or 50% summerwood in annual rings more 

than 2 in. from pith) and assigned dense southern pine a 

bending strength of 6,500 lb/in2

. The 4th edition of the 

NESC, published in 1927, included northern white cedar 

with an ultimate bending strength of 3,600 lb/in2

. 

The first minimum quality specifications for wood utility 

poles were published in 1931. They included speciesdependent minimum dimensions as well as physical appearance properties and were presented as a consensus 

standard agreed to by a group comprising wood pole producers, users, and general interest members. The specifications were published as a nationally recognized consensus 

standard under ASA guidelines. 

In the first half of the 20th century, wood pole DFS values 

were changed a number of times in response to the introduction of new data. In the second half of that century, the 

introduction of new data resulted in a change to the definition of DFS rather than the values. When the ASA O5 

standard committee was initiated, scientists from the Forest 

Products Laboratory of the USDA Forest Service advocated the use of standard, small–clear test values as a more 

consistent basis for assessment of pole load capacity. They 

cited inconsistencies in the full-size pole test data that 

compromised the credibility of these data as a design basis. 

The values ultimately selected represented a compromise 

heavily weighted in support of poorly controlled full-size 

pole test data. In 1955, the ASTM initiated a study to provide new, more credible test data (small–clear as well as 

full-size pole strength tests) that influenced the last changes 

to DFS values implemented in the ASA–O5–1963 standard 

in the 20th century. 

Adoption of the larger transmission-size poles in 1972 

sparked some question about the applicability of the existing DFS values to larger poles. Tests of larger poles, conducted in the 1980s, suggested a significant reduction in 

groundline moment capacity of larger poles when tested 

following the standard full-size pole test (ASTM D 1036). 

This initiated debate in both Canada (Bhuyan and Chetwynd 1994) and the United States over how this should be 

acknowledged in the load-capacity-based pole size classification tables (ANSI O5.1 and CSA–O15 (CSA 1990)). 

Table 1—2002 ANSI derivation for DFS for poles ≥50 ft long from EPRI databasea

Fiber stress property (lb/in2

) 

Species n 

Groundline stress 

at failure MORfl MORgl DFS 

Douglas-fir 172 6,730 6,480 7,020 7,640 

Southern pine 122 7,580 6,825 7,800 8,530 

Western redcedar 105 4,200 4,310 5,030 6,040 

a

n is number of full-size poles tested in each length category. Groundline stress is based on measured 

 dimensions. MORfl is modulus of rupture at failure location; MORgl is estimated groundline strength 

 based on MORfl; DFS is groundline strength adjusted for conditioning, drying in service, and class oversize. 

 For southern pine and Douglas-fir, a 10% increase for drying in service is countered by a 10% reduction 

 due to high temperature pretreatment conditioning. 
6 

First National Standard Specification 

for Wood Poles 

By the time the first standard specification for wood poles 

was published in 1931, test data provided by the American 

Bell Telephone Company, American Telephone and Telegraph, Pacific Telephone and Telegraph Company, and the 

Engineering Experiment Station at State College Station in 

Raleigh, North Carolina, supported strength values of 

5,600 lb/in2

 for western redcedar and 7,400 lb/in2 for 

treated southern pine (Table 2). The values shown in 

Table 2 were derived using a cantilever bending test that 

had been developed by Bell Laboratories. This test was 

adopted as an ASTM standard in 1949. Descriptions of the 

individual poles tested to support this analysis showed they 

represented a range of conditions with respect to pole size, 

degree of conditioning, and quality. Colly (1931) points 

out that the sectional committee responsible for wood pole 

values chose to set strength values at half the standard 

deviation or 8% below the determined averages (Table 2, 

column 3) to address concerns over technical problems 

apparent in the data. The committee called this adjustment 

(γ = 0.92) a “variability factor.” This resulted in the values 

of 7,400 lb/in2

 for creosote treated southern pine, 

5,600 lb/in2

 for western redcedar, and 6,000 lb/in2 for 

chestnut. The value of 3,600 lb/in2 for northern white cedar 

was not adjusted. The committee decided that because this 

value had been used for at least 20 years with acceptable 

performance, it did not warrant a reduction to 3,300 lb/in2

. 

The 1931 standard included a total of 10 pole classes. Pole 

classes 1 through 7 were designated as poles having tip 

load capacities defined by the function 

 2 P = 5400 − 950C + 50C (1) 

where C is class (1 through 7) and P transverse tip load 

capacity of class C pole. 

The groundline moment resulting from this load divided by 

the DFS for the species gives the groundline section 

modulus required for each pole size class. This relationship 

has remained unchanged for load requirements for pole 

classes 1 through 7. Classes 8, 9, and 10 were defined 

simply by tip circumference and were intended to cover 

lightly loaded lines. 

Subsequent Standards for Wood 

Poles 

Few events influenced growth in the utility pole industry 

more than the creation of the Rural Electrification Administration (REA). Formed in 1936 to expand the availability 

of electric utilities, the REA so accelerated the demand for 

poles that by 1947, yearly production was 8.1 million poles 

(Wilson and Drow 1953). Soon after the ASA O5 committee completed the 1931 standard, there was a push from 

REA to develop improved wood pole specifications and to 

expand tables to recognize the use of many western softwoods. 

Because of discrepancies in species values resulting from 

inadequate control of moisture content, scientists from the 

Forest Products Laboratory in Madison, Wisconsin, proposed derivations based on an available database of 

bending strength values derived using standard green 

small–clear bending tests (ASTM D 143). The industry, 

however, preferred full-size pole data. The committee 

developed a compromise solution that involved derivation 

of a relationship between full-size poles and small–clear 

wood that would allow the use of small–clear data in instances where full-size pole data were missing or inadequate for deriving DFS. This approach was used by the 

War Committee on Specifications and Dimensions for 

Wood Poles formed in 1945 to prepare war standard specifications for species not previously used as poles. 

Table 2—Strength values available to 1930 ASA committeea

 Full-size poles Small clear green wood 

Species n 

Average 

MOR 

(lb/in2

) 

Relative 

strength

(lb/in2

) 

Approx. relative 

strength 

(%) 

Standard ultimate 

fiber stress 

(lb/in2

) 

Relative strength of 

standard value 

(%) 

Northern white cedar 56 3,621 41.2 45 3,600 45 

Western redcedar 151 6,065 69 67 5,600 70 

Chestnut 98 6,480 73.8 75 6,000 75 

Southern pine (creosoted) 121 8,026 91.4 — 7,400 92.5 

Southern pine (untreated) 55 8,784 100 100 8,000 100 

a Colley (1932). 
7 

ASA O5.1–1948 

After World War II, the ASA standards were revised to 

produce the new ASA O5.1–1948 standard, which attempted 

to resolve the discrepancies between values derived from 

standard tests of small–clear samples and those derived from 

full-scale pole tests (Wilson and Drow 1953). The committee dealt with this problem by first plotting the small–clear 

values (ordinate) versus pole-test values (abscissa) for average modulus of rupture (MOR) (Fig. 1). Drawing one line at 

45° through the origin indicated that for most species, the 

pole values exceeded the small–clear derivations. Lodgepole 

pine, which was considered to have given satisfactory performance for years when designed at 6,600 lb/in2

, showed 

the greatest difference between small–clear and full-scale 

pole test values for the species shown in Figure 1, except for 

western hemlock. Data from the Canadian Forest Products 

Laboratory at the time showed higher values for western 

hemlock, which would have pushed the value above the 

designated line. 

The ASA committee decided that a second line through the 

origin and the lodgepole pine coordinates should be used to 

define the fiber stress values for all species. This line defined 

values 20% greater for poles than for small–clear wood. A 

horizontal projection to this line from the (small–clear, pole) 

strength coordinates for each species defined the average 

value to be used in defining standard ultimate strength in the 

ANSI standard. Ratings assigned to new species were taken 

as less than or equal to 1.2 times the measured small–clear 

average MOR. 

To minimize the number of dimension tables required, stress 

categories were developed by dividing the range of projected 

pole strength values into nine segments. The general philosophy was to provide 10% increments, but there was some 

deviation based on historic precedence. With this grouping, 

the “design” values ranged from 6% to 20% above the MOR 

determined from small–clear tests. A similar procedure to 

that described in the previous text was employed using 

western redcedar as the basis. This procedure also gave 

values for poles that were greater than values from the average small–clear MOR values, but to a lesser degree. Table 3 

compares values derived using lodgepole pine and western 

redcedar bases to what was ultimately published in the 1948 

ASA O5.1 standards. 

Note that in 1948, the committee decided to retain the values 

for northern white cedar, western redcedar, and southern 

pine that were published in the 1931 standard. Species introduced by the war committee in 1945 were given the higher 

values derived using the lodgepole pine basis. 

ASA O5.1–1963 

The last changes to wood pole DFS values (Table 4) occurred in response to new information obtained from the 

ASTM wood pole research program (Wood and others 

1960). The 1963 ASA O5.1 standard added four species 

(Alaska cedar, Engelmann spruce, white spruce, and Sitka 

spruce), removed one species (pond pine), and changed 

fiber stress values for four species (Douglas-fir, southern 

pine, northern white cedar, and western redcedar). 

Figure 1—Comparison of tests of small–clear specimens and full-scale tests used as 

basis for DFS values on right (lb/in2

). 
8 

Table 3—Designated fiber stress values adopted in 1948 on basis of small–clear testsa

ASA fiber stress (lb/in2

) 

Species 

Clear wood MORa

(lb/in2

) 

Lodgepole pine 

basis 

Western 

redcedar basis ASA 1948 basis 

Northern white cedar 4,200 5,200 4,800 3,600 

Englemann spruce 4,200 5,200 4,800 — 

Western redcedar 5,100 6,000 5,600 5,600 

Ponderosa pine 5,000 6,000 5,200 6,000 

Jack pine 6,000 6,600 6,000 6,600 

Lodgepole pine 5,500 6,600 6,000 6,600 

Red pine — 6,600 6,000 6,600 

Western fir 5,800 6,600 6,000 6,600 

Western hemlock 6,100 7,400 6,600 7,400 

Douglas-fir 7,600 8,400 7,400 7,400 

Southern pine 7,800 8,400 7,400 7,400 

Western larch 7,500 9,400 8,400 8,400 

a Wood Handbook (USDA 1940). 

Table 4—ASA O5.1 1963 derivation of ASA fiber stress values based on standard small–clear specimen tests 

Adjusted strength (lb/in2

) 

FPL–39/ASTM Drying b,d,f

Species 

ASA 

O5-48 n

Pole 

MOR 

(lb/in2

) 

Clear wood a

(lb/in2) 

Green clear 

wood b,c,d

Small 

clear 

woode

 Pole MOR 

ASA O5.1–63 

(lb/in2

) 

Northern white cedar 3,600 — — 4,250 4,560 4,950 — 4,000 

Western redcedar 5,600 46 5,370 5,120 5,500 5,960 5,790 6,000 

Western larchg

 8,400 54 8,500 8,180 7,900 8,577 8,250 8,400 

Lodgepole pine 6,600 50 4,810 5,490 5,900 6,400 5,190 6,600 

Douglas-fir 7,400 45 8,180 7,590 7,340 7,960 7,940 8,000 

Yellow pineg

 7,400 106 8,990d 7,970g 7,275 7,900 8,240 8,000 

Loblolly — 9 8,990 7,340 6,700 7,272 — — 

Shortleaf — 40 8,450e

 7,300 6,600 7,220 — — 

Longleaf — 41 8,920 8,670 7,920 8,580 — — 

Slash — 11 9,600 8,570 7,820 8,480 — — 

Western fir 6,600 — — — — — 6,600 

Red — — — 5,950 6,390 6,930 — — 

Grandg

 — — — 6,060 6,510 7,060 — — 

Noble — — — 5,790 6,220 6,750 — — 

Silverg

 — — — 5,890 6,300 6,860 — — 

White — — — 5,700 6,120 6,640 — 

Western hemlock 7,400 — — 6,140 6,590 7,150 — 7,400 

Alaska cedar — — — 6,450 6,930 7,510 — 7,400 

Sitka spruce — — — 5,660 6,080 6,590 — 6,600 

White spruceg

 — — — 5,580 5,990 6,500 — 6,600 

Engelmann spruce — — — 4,540 4,880 5,290 — 5,600 

a

Values from Wood Handbook (USDA 1955). b

Reduced for variability effects. 

c

Increased 5% for form effect. 

d

Reduced for conditioning: 0.9 for boultonizing (Douglas-fir, larch) and 0.85 for steaming at <245°F (southern pine). e

Increased 8% for form effect. 

f

Increased 16% for drying in service. 

g

Clear wood strength values for these species changed enough in 1966 to have a significant impact on the 

 ASA derivation, but since these are minor species the changes were never recognized by the ASA standard. 
9 

The 1963 standard also made an explicit declaration that the 

specification was intended only for poles loaded as cantilever members subject to transverse load resulting in maximum moment at the groundline. 

More than 30 years of experience in observing the performance of wood poles and relating that performance to pole 

specifications weighed heavily in the selection of the DFS 

values used to classify poles using a standard definition of 

cantilever load capacity. Factors considered in the selection 

of DFS values included traditional adjustments adopted in 

the 1930s—variability, form factor, conditioning effects, and 

moisture content. The effects of high temperature conditioning on wood had been studied rather extensively. The other 

adjustments, however, which dated to discussions prior to 

1940, were not well supported by theory or data but had 

been referenced in selection of values given in previous 

standards. These factors are discussed in more detail in the 

section titled Nominal Strength. For those species for which 

there were no or limited full-scale pole test data, the committee referenced published values for small–clear bending 

strength (Wood and Markwardt 1965, USDA 1955). 

Table 4 compares the values given in the 1948 (column 2) 

and 1963 (column 9) versions of the ASA O5 standard. The 

intermediate columns show the scenarios considered by the 

committee that appeared to have the greatest influence on the 

assignment of DFS values. These include values derived 

from the ASTM wood pole report (Wood and others 1960) 

(Table 4, columns 3 and 4) and published clear wood values 

(column 5). Columns 6 through 8 show these values adjusted 

to take into account pretreatment conditioning, drying in 

service, and between-study variability. There was no 

standard set of parameters used to select a DFS value 

(column 9). This was, in the words of Wood and Markwardt 

(1965), “a system of rationalizations that afforded the technical basis supplemented by engineering judgment.” 

Most of the ANSI DFS values appear to be based on clear 

wood strength with no adjustment for grade effect (Table 4, 

last column). Values such as those for western hemlock, 

western larch, lodgepole pine, and the western fir group, 

unchanged from the 1948 standard, appear to be more 

closely aligned to adjusted small–clear strength values than 

to full-scale pole test results. Northern white cedar was 

assigned a value close to the published clear wood value 

with no adjustment. For Alaska yellow cedar, western redcedar, and spruce, selected values were close to the clear 

wood strength adjusted for form (Newlin and Trayer 1924), 

variability (0.93), and in-service drying. 

Changes to full-scale pole data and published small–clear 

strength values introduced subsequent to the 1963 standard 

had no effect on the ASA O5.1 values. In 1963, the Forest 

Products Laboratory played a dominant role in setting allowable stresses. The small–clear strength values referenced by 

ASA O5.1 were those published in the Wood Handbook

(USDA 1955). Industry concern over the dominant role 

played by the Forest Products Laboratory led the ASTM 

Committee on Wood to introduce “Standard methods for 

establishing clear wood strength values” (ASTM D 2555) in 

1966. In that standard, clear wood values for several species 

were changed from the values reported in the Wood Handbook. These changes, however, were never acknowledged by 

ANSI in the form of a revision to the 1963 DFS values. This 

supports the conclusion that these values were never tied to 

any algorithm that references published wood strength values. 

Table 5 shows how the ASTM changes to clear wood 

strength values (D 2555–66) would have affected the ASA 

fiber stresses. Despite the fact that no significant changes 

were reported for the strength values for yellow pine poles, 

there was a change to the way “species group” clear wood

Table 5—Effect of changes in clear wood strengtha

Adjusted clear wood 

strengthb

 (lb/in2) 

Species 

1963 ASA–O5.1 

 fiber stress 

(lb/in2

) 

D 2555–66 clear 

wood strengthb

(lb/in2

) Green 

Dryc

1.16 (Green) 

Southern pine 8,000 7,590 6,861 7,923 

Western larch 8,400 7,652 6,920 8,020 

Grand fir 6,600 5,839 5,860 6,800 

Noble fir 6,600 6,169 6,200 7,190 

Silver fir 6,600 6,410 6,440 7,470 

Western hemlock 7,400 6,637 6,670 7,730 

White spruce 6,600 4,995 5,020 5,820 

Englemann spruce 5,600 4,750 4,770 5,530 

a

ASA O5.1 DFS values not modified to recognize changes in clear wood strength 

 presented in ASTM D 2555 standard. 

b

Reduced for variability effects (0.93), increased 8% for form effect, adjusted −0.9 for 

 kiln drying for pine and larch (air-dry for other species, 1.0). 

c

Increased 16% for drying in service.
14 

stresses are derived. The ASTM D 2555 standard requires 

that these values be assessed by weighting according to 

standing timber volume (7,590 lb/in2) rather than using an 

average of the subspecies mean values. For the other species 

listed, average green clear wood values were changed by 

the ASTM D7 committee to reflect changes in the 

supporting data. 

In 1963, the ASA O5 committee reduced the number of pole 

classes by one. Classes 8, 9, and 10 were not rated as utility 

poles. Rather they were used in construction of pole buildings. The suggestion to drop all three classes was countered 

by the agreement to drop only class 8, leaving the gap between utility and construction poles. 

Addition of H-Class 

In 1972, along with the change from the “ASA” designation 

to “ANSI” came the addition of an appendix that provided 

information on the assumptions made in developing the 

tabulated pole size classes. Load capacities assumed for each 

pole class were moved from the main body of the standard 

into the appendix, and these values were expanded to include the new H-class poles. Unlike the standard classification for distribution poles that is inversely related to pole 

size/capacity, the new size designations (H1 through H6) 

correlate directly to pole size and capacity. To simplify 

discussion related to the ANSI size classes, an index (i) is 

used, ranging from 0 to 12 and correlating to pole size from 

class 7 (smallest of utility poles) to H6 (largest transmission 

pole). The class loads for these poles fit a quadratic function 

(Eq. (2)): 

 Class load = 1200 + 250i +50i

2 (2) 

Tip circumference values for ANSI class poles follow a 

linear function of this class index. 

 Tip circumference = 15 + 2i (3) 

The groundline circumferences are derived to provide the 

section moduli required to resist the groundline moments 

imposed by the class loads applied transversely at a location 

2 ft from the top of the pole (ASTM D 1036–1999). The 

estimated groundline circumference is then extrapolated to a 

circumference 6 ft from the butt to give the ANSI tabulated 

values. 

To estimate a 6-ft-from-the-butt circumference from the 

calculated groundline value, ANSI incorporates speciesdependent circumference tapers, also given in the ANSI 

appendix. The origin of these taper values is uncertain, and 

they appear to have changed slightly over time. In 1974, 

Bohannan and others evaluated survey data collected for 

southern pine, Douglas-fir, and western redcedar pole producers. Their analysis showed mean circumference tapers 

for these three pole species (0.34, 0.31, and 0.46 in/ft, 

respectively) to be 20% to 50% greater than the values noted 

in the ANSI appendix and up to 50% lower than values 

derived from tabulated class tip and 6-ft-from-butt 

circumferences. 

In addition to information on how size classes were determined, the ANSI annex provides some cautionary statements. Of particular importance is the note regarding the 

assumption that maximum stress occurs at groundline. This 

is theoretically correct only if the circumference at groundline is less than or equal to 1-1/2 times the circumference at 

the point of load. If this is not the case, maximum stress is 

likely to occur some distance above the groundline. Prior to 

2002, the DFS was derived to represent the average stress at 

failure rather than groundline MOR and could be used with 

ANSI minimum dimensions to conservatively estimate pole 

capacity regardless of the location of maximum stress.

Wood Pole Test Data 

In 1949, ASTM published the first standard test procedure 

for wood poles (ASTM D 1036). This standard actually 

includes two methods for testing poles as a structural element intended to be loaded as a cantilever having its maximum moment at a location close to the butt end. One method 

involves actually loading the pole as a cantilever beam, with 

one end rigidly anchored at and below “groundline” and load 

applied 2 ft from the other end. The second method, called 

the “machine,” incorporates a universal testing machine. The 

pole is placed on simple supports, and a vertical load is 

applied at the “groundline” location. Test results differ 

slightly as the machine method includes stress resulting from 

the dead weight of the pole. Bending moment resulting from 

dead load is maximum at a location close to mid-span. When 

added to the moment resulting from the groundline concentrated load, this results in a slight shift in the point of maximum stress away from the groundline. The difference is 

generally inconsequential for short poles, but it becomes 

more significant as pole length increases. 

ASTM Wood Pole Test Program 

The ASTM wood pole research program initiated in 1955 

strengthened the basis for characterizing the strength of 

domestic pole species. The study involved tests of full-size 

poles of five domestic species (southern pine, Douglas-fir, 

western redcedar, lodgepole pine, and western larch), along 

with matched small–clear specimens. Tests focused primarily on 30-ft poles but also included 25-ft and 55-ft poles. 

Yellow pine poles were tested using both the cantilever and 

machine tests described in ASTM D 1036. The other species 

were evaluated using only the machine test method. A total 

of 630 full-sized poles and 14,000 matching small–clear 

specimens were tested. 

Results of this study (Wood and others 1960) provided a 

basis for the 1963 revision of the ASA O5.1–1948 standard 

DFS values. Changes included increases to designated 

10 
15 

strength values for Douglas-fir, yellow pine, and western 

redcedar. In addition, the data strengthened the committee’s 

confidence in standard small–clear tests as a basis for evaluating pole strength, leading to an increase in the DFS for 

northern white cedar. Wood and Markwardt (1965) provide 

a synopsis of the committee decisions that led to these 

changes. Table 4 compares results published in the ASTM 

pole report, Wood and Marwardt’s work (FPL–RP–39), and 

the ASA O5 standards. 

Key points gleaned from the ASTM study that influenced 

subsequent decisions related to pole strength include the 

following: 

• Average groundline strength of poles is closely 

correlated to strength of small–clear samples taken 

from the butt section. 

• Change in strength with height is correlated to change 

in specific gravity. 

• Strength of young fast-grown poles having fewer than 

6 rings per inch in the outer 2 in. is significantly lower 

than that of poles having 6 or more rings in the outer 2 in. 

• Poles in the 55-ft length category have lower groundline 

MOR than do 25-ft poles. 

EPRI Wood Pole Test Program 

The second major contribution to the advancement of wood 

pole engineering came in the 1980s when the Electric Power 

Research Institute (EPRI) sponsored a research program to 

promote development of probability-based design guidelines 

for wood utility structures. This effort attracted the attention 

of the wood industry as a major step in the development of 

probability-based load-resistance-factor design (LRFD) for 

engineered structures and for promoting the image of wood 

poles as engineered components. This research addressed 

many issues that were questioned after the adoption of the 

H-class sizes in 1972. The testing phase focused primarily 

on the larger poles, including 122 southern pine poles ranging from 50 to 70 ft long, 172 Douglas-fir poles ranging 

from 50 to 95 ft long, and 100 western redcedar poles ranging from 50 to 70 ft long. The analysis phase included a 

review and compilation of the state of the knowledge on 

wood poles. The published results have been widely referenced by academia as well as private industry and form the 

basis for design and maintenance of wood pole transmission 

line systems. Results are published by the EPRI under the 

designations EPRI EL–2040 project 1352–1 and 

EPRI EL–4109 project 1352–2. 

As part of the EPRI-sponsored research from 1980 to 1985, 

a database was compiled from existing wood pole test data. 

This database comprised results of full-scale pole tests conducted in accordance with the ASTM D 1036 standard test 

procedure. The references for these studies are listed in 

Appendix B. These data were made available to the ANSI 

O5.1 committee for use in supporting current and future 

utility pole standards. 

In 1994, Bhuyan and Chetwynd (1994) used a large portion 

of the data generated by the EPRI project along with privately sponsored testing of poles to provide a basis for 

evaluating the classification stress values published by the 

Canadian pole standard (CSA O15) as well as ANSI. They 

reviewed values for western redcedar, yellow cedar, red 

pine, jack pine, lodgepole pine, southern pine, and Douglasfir. Their analysis of this data showed that test data for western redcedar, plantation-grown red pine, and some southern 

pine poles produced in the last 10 years and classified by the 

CSA O15 standard have load capacities that are one class 

lower than that implied by the standard. ANSI O5.1 classification stresses are greater than those of the Canadian Standards Association (CSA) for these three species. 

The two most significant conclusions drawn from this work 

are (1) as pole size increases groundline moment capacity 

decreases and (2) western redcedar pole test data do not 

support the classification stress values, especially for poles 

greater than 50 ft in length. 

Nominal Strength 

Throughout the recorded archives that constitute the history 

of the ASA/ANSI O5 material specification, there are numerous discussions related to the meaning of the strength 

values selected as a basis for classifying wood poles. Design 

values for wood poles have traditionally been established by 

the NESC, as was discussed in the Introduction. For the past 

60 years, however, the NESC has referenced the ASA/ANSI 

standard DFS values as a basis for wood pole design values. 

It is interesting to note that until the 2002 revision of ANSI 

O5.1, the DFS values were treated solely as a basis for relative size classification within the standard: the standard 

provided no assessment of these values in terms of the probability density function for bending strength. This changed 

in 2002 when the committee decided to redefine DFS as a 

groundline fiber strength that represents the mean value for a 

strength distribution having a coefficient of variation (COV) 

of 20%. The following text provides the background discussion that led the committee to make this change. 

Initial Estimates 

The natural variability of a material such as wood adds to the 

complexity of balancing design safety and economic feasibility. In the early 1930s, the ASA O5 committee adopted an 

analysis procedure that was assumed to give a 3% probability of failure for poles in a distribution system where each 

pole is assumed to carry its ASA O5 “class” load. This 

analysis relied on four basic assumptions (Colley 1931): 

11 
14 

1. Pole strengths are normally distributed. 

2.The average pole in any size class has an actual groundline 

section modulus that is 12.5% greater than that calculated 

on the basis of the minimum class circumference. 

3.Creosoted southern pine poles have a mean strength of 

8,026 lb/in2 and standard deviation of 1,348 lb/in2

; 

western redcedar has an average strength of 6,065 lb/in2

and standard deviation of 1,151 lb/in2. These data 

provided the basis for assigning a 17% COV for bending 

strength for all pole species.

4. Poles are used in a redundant system in which three in-line 

poles share load, each carrying load in proportion 

to its relative stiffness. 

The first two assumptions were reasonably conservative. The 

load increment between consecutive size classes was initially 

set at 25%. This meant that groundline section moduli for 

poles of a given length had a 25% increment between 

classes. If the average pole within a class is assumed to have 

a section modulus that falls midway between its class minimum and that of the next larger class, it will have an approximately 12% higher groundline bending moment 

capacity than that of a pole having the minimum class dimension. This means that the groundline stress on the average pole is 1/1.12 or 0.892 times4

 the DFS. 

Colley (1931) points out that the sectional committee on 

wood poles chose to set strength values slightly below the 

determined averages to address concerns over technical 

problems apparent in the data. There was concern among 

committee members that moisture content variations in test 

poles had a significant impact on species strength values that 

were being used to establish size classification tables. Rather 

than reject the data, the committee chose to apply a confidence adjustment for bending strength, publishing values 

calculated to be one-half the standard deviation below the 

measured mean. Data available at the time showed a range 

of within-sample strength COVs of 17% to 23%. Rather 

than derive this for each species, the committee chose a 

“variability factor” (γ = (1 − 0.5Ω)) based on the COV (Ω) 

of 17%. As a result, nominal resistance values were set as 

92% of the mean test strength. 

In 1963, the ASA committee adopted the same confidence 

adjustment on mean strength. At that time, the ASTM data 

showed COV values ranging from 10.8% to 16%. The 

committee assumed a COV of 14% and recalculated this 

variability factor to be 0.93. In 1963, this adjustment was 

 

4

 Today, the pole load increment varies from 27% for the 

smallest pole class to 14% for the largest, so the groundline 

stress on the average pole will vary from 0.88 to 0.93 times 

that estimated on the basis of minimum dimension. 

warranted on the basis of small within-class sample size 

rather than variations in moisture content below fiber saturation at time of test. 

In an attempt to assess the true meaning of the DFS value, 

Colley (1931) evaluated the product of “variability” and 

“oversize” adjustments in terms of standard normal deviations. For poles of classes 4 through 7, the product of class 

size adjustment (0.89) and variability (0.92) is 0.82. For a 

normal distribution having a COV of 17%, this factor times 

the mean gives a value equivalent to 1.055 standard deviations below the mean which equates to the 14th percentile of 

a normal distribution. 

The committee then attempted to show that this value actually represented the 3rd percentile of in-service pole strength 

by proposing that a utility distribution line represents a loadsharing system. Theoretically, if the load on any one pole is 

distributed in a manner to cause the loaded pole and two 

adjacent poles to deflect together, each pole might be assumed to carry a share of the load proportional to its relative 

stiffness. Given a perfect correlation between stiffness and 

strength, one might also assume that each pole is then loaded 

to the same proportion of its load capacity. This means that 

failure will not occur until all three poles have been loaded 

to capacity. So, the capacity of the system is limited by the 

average strength of three poles in series rather than the 

weakest pole. The distribution of the average values of three 

poles will have a variance equal to one-third the variance of 

the individual values. Equating the ANSI DFS (mean pole 

capacity, adjusted by the variability and class-oversize adjustment factors) to a fractile of a normal distribution using 

Equation (4) provides a basis for estimating the probability 

that a pole in a load-sharing system will fail at a load that 

represents the average single pole capacity. 

( ) ⎟

⎟

⎠

⎞

⎜

⎜

⎝

⎛ = − Ω

3

γ 0.892 X X 1 k (4)

where 

Ω is coefficient of variation (assumed to be 17%), 

γ variability of adjustment (assumed to be 0.92), and 

k normal distribution variable. 

Solving for k gives the value 1.827. Assuming that the adjusted strength representing the average stress on three adjacent poles is 1.821 standard deviations below the mean 

strength of three adjacent poles suggests that the DFS 

represents the 0.03 fractile of the distribution of strengths of 

three-pole load-sharing assemblies. 

This analysis involves three flawed assumptions. First, there 

is no rigid load distributing element in a distribution line that 

will redistribute load laterally to adjacent poles in such a 

manner that all three poles will deflect equally. If such a 

mechanism did exist, the correlation between stiffness and 

12 
15 

strength would not be perfect, so the reduction in variability 

would not be as large as 1/31/2. Second, the basis for the 

“variability factor” (0.92) was a lack of confidence in the 

data. The adjusted value was the committee’s estimate of 

mean performance for green poles that met their minimum 

specifications, and γ should not have been included in the 

expression. Finally, load sharing is an engineering judgment, 

not a material property, and has little meaning in a material 

specification. 

It does not appear that the NESC, which provided design 

recommendations on the basis of the ASA pole strength 

values, treated these values as though they were lower fractiles. It was standard practice at the time for timber engineering standards to recommend design values as 25% of the 

mean strength. The 5th edition of the NESC (1941) adjusted 

ultimate strength by a factor of 0.25 for grade B and 0.375 

for grade C. This would have been overly conservative if 

applied to a 0.03 fractile of the pole strength distribution. By 

the early 1960s, design standards for wood began referencing a lower 5% exclusion for strength divided by 2.1 to 

account for duration of load effects. When applied to a 

population having a COV of 30%, this results in 0.25 times 

the mean. So, by the 1960s, the NESC approach was recognized as compatible with recognized timber engineering 

standards. 

Test Data and In-Service Performance 

A major issue for most standards development has been the 

relationship between laboratory evaluation under controlled 

conditions and in-service performance. 

Early debate over development of the ANSI O5.1 standard 

focused on the efficacy of standard full-size pole tests compared with standard small–clear tests as a basis for setting 

pole size class requirements. The objective was to provide a 

classification system in which any pole of a designated 

“class” had the same tip load capacity, regardless of species. 

The small–clear test assessed the fiber strength that would 

control in the event of a groundline failure. 

The full-size pole test provided a more direct assessment of 

pole capacity independent of failure location. The full-size 

pole tests attracted greater support as being more representative of the end product. 

As engineered design gained importance for utility structures, the relationship between standard full-size pole tests 

and in-service pole loading began to be debated. The pole 

size classification system adopted by ASC O5, based on 

groundline stress at failure, was generally considered to 

provide an appropriately conservative estimate of load capacity for poles loaded as cantilever beams. When standard 

load conditions are applied to longer poles, tree shape and 

physiology have a significant influence on where and how 

the poles fail, leading to questions about a negative bias for 

longer poles that are likely to fail further away from groundline and to exhibit a lower groundline stress at failure. 

While standard tests provide a means of assessing relative 

value between species and pole sizes, they rarely address 

all design issues. In the case of wood poles, DFS values used 

to set pole size class include some adjustment to account 

for effects of high temperature conditioning and drying 

in service. 

Conditioning 

Preservative treatments are what make wood a viable option 

for use as poles. Their effect on strength is therefore of vital 

concern to pole producers. Many studies conducted to assess 

treatment effects on strength have concluded that any 

strength reduction is due to high temperatures involved with 

pretreatment conditioning. This conditioning sterilizes the 

pole and removes excess water, making the pole more conducive to treatment. 

The committee responsible for the 1963 version of the standard recognized potential strength reductions of 15% for 

steaming and 10% for boultonizing. The steaming reduction 

was supported by research studies conducted by Betts and 

Newlin (1915), Wilson and others (1930), Buckman and 

Rees (1938), Stamm (1956), and Thompson (1969). 

The ASTM pole research program (Wood and others 1960) 

provided data that suggested that previous studies on steaming effects were non-conservative. The average strength of 

treated western larch and southern pine poles was less than 

85% that of untreated poles. The difference in western larch 

was attributed to the lower specific gravity of the treated 

poles compared with that of untreated poles. In the case of 

southern pine, the committee attributed the large reduction to 

the fact that the poles were conditioned following the guidelines of the AWPA standards C1–55 and C4–55, which 

permitted the poles to be subjected to 259°F for 8 to 13 h. 

During the conditioning processes, temperatures reached as 

high as 267°F on four of the charges. These older standards 

were considered to be too destructive and not representative 

of state of the art in 1963. Erickson and Dohr (1959) compared bending strengths of longleaf pine poles at a range of 

conditioning times and temperatures and confirmed MacLean’s findings: on average, poles that were steam conditioned at 245°F for 15 h showed a 16% loss in bending 

strength. The data of Erickson and Dohr also showed a 

reduction in variability, which led to the observation that a 

13% loss in bending strength occurred at the 5th percentile. 

Drying Effects 

Wood drying in service has a mix of effects on load capacity. Wood strength and modulus of elasticity are known to 

increase with drying below the fiber saturation point, while 

section property is reduced as a result of shrinkage (Wolfe 

2000). Drying checks that open as poles dry may expose 

poorly treated wood to trapped dirt and moisture, promoting 

13 
14 

local decay pockets. Excessive drying and weathering in arid 

environments may cause poles to lose strength over time. 

For most heavy timber engineered applications, these effects 

are considered to be offsetting. 

Eggleston (1952) supported the premise that the green 

groundline strength basis was appropriate for the ASA pole 

size classification. He evaluated butt-soaked dry poles and 

concluded that except in soils so dry that pole butt moisture 

content would not exceed the fiber saturation point, the 

strength of the tops of southern pine poles in service is 

probably very close to the (saturated) groundline strength. 

Rhatigan and Morrell (2002) reported an average of 19% to 

40% moisture content at groundline for creosoted Douglasfir poles that had been in service for 35 years, suggesting 

that the assumption of dry in-service conditions is 

nonconservative. 

Kulp (1957) reported finding decay in poorly treated wood 

above groundline. He found that 1.7% of 1,885 southern 

pine poles placed in test installations were removed after an 

average 15 years of service; 5 failed within 13 years of 

service as a result of decay. These poles had been treated 

with coal tar creosote. Kulp reported that 1.7% of the remaining poles were deteriorated as a result of decay or termite attack after 6 to 18 years of service. In semi-arid 

regions of Montana and the Dakotas, the tops of 6.1% of 

butt-treated lodgepole pine poles were partly decayed after 

5 to 8 years of service, and the tops of 28% of butt-treated 

western redcedar poles had light to moderate shell-rot after 

9 to 15 years of service. Light top decay around spur marks 

was present in 21% of untreated northern white cedar. Kulp 

did not report moisture content, but the incidence of decay is 

a reasonable indication that moisture content exceeded 20% 

in these poles. 

Research in support of increasing pole strength for drying in 

service was reported by Wilson and others (1930) and by the 

ASTM pole study. These studies showed a 10% strength 

increase for shortleaf pine poles and 14% strength increase 

for lodgepole pine when dried to 20% moisture content. 

Further support for recognizing the dry strength of poles was 

gained from a survey conducted by the REA that showed 

that moisture content 4 ft above ground was lower than 20% 

in a high percentage of poles (Wood and Markwardt 1965). 

The REA data included more than 1,400 poles in a region 

north and east of 40 ̊N,100 ̊W (Kansas/Nebraska line, west 

of Missouri, north of Cincinnati, Ohio) and 700 poles to the 

south and east of that point. More than 1,200 poles were 

surveyed between mid-Kansas and the California/Nevada 

border; 33 were surveyed west of that point. The survey 

showed that moisture content at 4 ft above ground exceeded 

20% in 13% of poles surveyed in the southeast region, 4% in 

the moderate northeast region, and fewer than 1% of poles 

west of the Colorado/Kansas border. 

This information was available to the ASA committee in 

1963 when it decided to consider drying in-service to account for a 10% to 16% increase in bending strength values 

determined in the green condition. Discussions leading to the 

acceptance of these increases included concern for the length 

of time from installation to the point a pole actually reaches 

20% moisture content and the fact that for most poles, the 

highest stress occurs in a region where moisture remains at 

fiber saturation. In most cases, these increases offset reductions caused by pretreatment conditioning, resulting in values that appeared to be representative of the test data. It was 

only for species such as western redcedar that did not require 

high temperature conditioning that the drying effect noticeably increased strength above that determined in green pole 

tests. As the ANSI standard was considered a material specification and not a design standard at the time, no reference 

was made to these assumptions or to any required adjustments for poles to be used in wet environments. The DFS 

values were used by ANSI only to set relative dimensions in 

the pole size classification tables. They were not considered 

by ANSI O5.1 to be a design basis. 

In 1986, ASC O5 made an explicit declaration that the tabulated pole circumference values were to be measured in the 

“green” condition. This was done to address questions about 

undersized dry poles. The fact that the tabulated “green” 

circumferences are derived using estimates of “dry” groundline strength is of little consequence when the only concern 

is size classification to ensure species interchangeability. If 

the standard is to promote the DFS values as a basis for 

design, however, it should be remembered that the ANSI 

class loads are not a good measure of the load capacity of the 

classified “green” pole. 

As discussed earlier in this paper, a major problem with 

using dry strength to set a green dimension is that the failure 

probability for the target pole is greater than 50% at the 

designated class load. A green pole sized using the dry 

strength will have a strength that is over one standard deviation (or 16%) below the strength required to carry the designated class load. When this pole dries, it will have a section 

modulus 6% below that required of a dry pole. 

The offsetting effects of shrinkage and fiber strength increase with drying, combined with the results of the EPRIfunded research on poles taken out of service, suggest that 

taking an increase in strength for drying in service is nonconservative. The EPRI research included a limited set of 

74 western redcedar, 49 southern pine, and 11 Douglas-fir 

poles taken out of service after 30 years that had average 

strengths 20% to 30% less than that of new counterparts. 

The effect of drying in service most likely varies with exposure conditions as well as species. An across the board adjustment was not warranted, and ASC O5 provided neither 

recommendations nor warnings about adjusting load capacity for drying. 
15 

Pole Shape 

The relationship between in-service capacity and standard 

test capacity is influenced by physical aspects of the pole 

and test specimen. These may be placed in three categories: 

1. Form, relating to round as opposed to prismatic sections. 

2.Taper, relating to change in section property over pole 

length. 

3. Size, relating to weak-link theories, which suggest that 

average extreme fiber strength will decrease as size increases. 

Pole form—Form effect refers to differences in the strength 

assessment of round as opposed to rectangular sections when 

using the standard engineering model for stress in a bending 

member. This model, which estimates stress as an inverse 

function of section modulus, assumes that stress is uniform 

across the width and varies linearly with distance from the 

neutral axis of any section. Newlin and Trayer (1924) concluded that this model gives roughly an 18% higher value 

for stress at failure for round than for square wood sections. 

Basically, the researchers found that square and round sections having the same cross-sectional areas had the same 

bending moment capacity. Because the square section has an 

18% greater section modulus, its calculated stress at failure 

is 18% less than that of the round section. Therefore, when 

estimating the strength of a round section on the basis of 

published small–clear square section strength, some form 

effect should be recognized. 

Pole taper—Wood pole taper has played a major but largely 

unrecognized role in the history of structural performance of 

ANSI poles. Table 6 summarizes measured and implied 

wood pole tapers that have had some influence on this standard and on the history of pole performance. 

The ANSI pole size classification relies in part on measured 

pole taper values, but the size classification tables hold 

conservative implications in this regard for the pole user. In 

deriving the 6-ft-from-the-butt (6’FB) circumferences in the 

size classification tables, the values listed in Annex B of the 

ANSI O5.1 standard (see Table 6) are used to transfer required circumference dimension from groundline to the 

6’FB location. In selecting an ANSI pole, however, a pole 

designer must assume that it will have the designated minimum tip and 6’FB circumferences given for its class. The 

corresponding taper is the difference in these dimensions 

divided by the length between them. As the tip circumferences vary in increments of 2 in. and the 6’FB circumferences for any length change in increments exceeding 2 in., 

the implied tapers increase with pole size. Table 6 provides a 

comparison of 55-ft and 80-ft poles; the low values in these 

cases are for class 7 poles and the high values for class H6 

poles. 

Table 6 also provides measured tapers reported by Bohannan 

and others (1974) and Wood and others (1960). The circumference values reported by Bohannan were determined from 

measurements of poles ranging in length from 55 to 80 ft. 

These data showed no consistent relationship between taper 

and pole length. In all cases, however, the taper appeared to 

increase with groundline circumference as implied by the 

ANSI tables. Wood’s data, which was determined for 

25- to 55-ft-long poles, showed a definite increase in taper 

with pole length for southern pine and western redcedar, but 

the evaluation did not provide a comparison across size 

classes. 

Comparing ANSI Annex B values to those reported by 

Bohannan suggests possible undersizing for longer poles. 

While the ANSI butt taper values appear to be about average 

for the 25- to 55-ft poles reported by Wood, they fall below 

the median values reported in Bohannan’s survey for each 

species (0.021 vs 0.028 for southern pine, 0.017 vs 0.027 for 

Douglas-fir, 0.032 vs 0.039 for western redcedar). The 

implication is that on average, for poles >50 ft in length 

where the groundline is above 6 ft from the butt, the 6’FB 

dimension required to give the necessary groundline section 

modulus will be underestimated using the ANSI Annex B 

value. Projecting back to groundline using the implied tapers 

in the ANSI tables results in an undersize groundline dimension for the classified pole. 

Table 6—Pole taper values 

Taper (in./in.)a

 Pole Southern pine Douglas-fir Western redcedar 

Reference length 

 (ft) Low Med. High Low Med. High Low Med. High 

Bohannan 55–80 0.014 0.028 0.045 0.011 0.027 0.050 0.018 0.039 0.077 

Wood 25–55 0.015 — 0.029 0.012 — 0.017 0.024 — 0.042 

ANSI Annex B — 0.021 0.021 0.021 0.017 0.017 0.017 0.032 0.032 0.032 

ANSI tables 55 0.033 — 0.043 0.033 — 0.043 0.047 — 0.055 

ANSI tables 80 0.030 — 0.040 0.030 — 0.040 0.038 — 0.049 

a

Calculated as change in circumference per change in height. 
16 

Taper has a definite effect on how stresses are distributed in 

a pole. Stress is not distributed linearly in a tapered cantilever beam. The bending moment (M) for a cantilever increases at a constant rate from the point of load application 

to the reaction. However, when divided by the section 

modulus (S), which increases as a cubic function of distance 

from the load point, the quotient (M/S) or stress is an inverse 

quadratic function that increases at a decreasing rate. If the 

taper is steep enough or the cantilever long enough, the 

stress stops increasing with distance from the load point and 

may decrease. This means that the point of maximum stress 

in a tapered cantilever does not necessarily coincide with the 

point of maximum moment. 

For a linearly tapered, round cantilever beam of uniform 

homogeneous material, the quotient M/S will reach its maximum value at a point where the pole circumference is 

1.5 times the circumference at the point of load. Under these 

conditions, the distance from the point of load to the point of 

maximum stress (Xσ) can be estimated as 

τ σ 5 lp X = 0. C (5) 

where 

Clp is circumference at load point and 

τ circumference taper. 

For most distribution poles, Equation (5) suggests that the 

maximum stress location will be at groundline. The estimated maximum stress location for a class 1 55-ft pole, 

however, would be 6 ft above groundline for southern pine 

and 18 ft above groundline for the average western redcedar 

using tapers taken from Wood and others (1960) or median 

values reported by Bohannan and others (1974). 

Of course, this purely geometric derivation provides only a 

rough approximation of how stresses vary in an actual wood 

pole. There are other variables to consider when predicting 

where a pole will fail. If maximum stress occurs at 6 ft above 

ground in dry wood and the groundline is saturated, it is 

likely that the groundline stress will exceed the saturated 

strength before maximum stress will exceed dry strength. 

Physiological effects such as knots, juvenile wood, strength 

change with height, and diving grain exposed in the debarking–peeling process also present points of weakness that 

might initiate a failure at a location other than the point of 

maximum stress. 

Material variability—Because wood is a product of nature 

rather than a manufacturing process, its strength is much 

more variable than that of most commonly used engineering 

materials. Wood strength varies not only between trees but 

also within a tree. Between-tree variability in strength is 

attributed to growing conditions (e.g., nutrients, rainfall, 

length of growing season, competition) and within-tree 

variability to factors such as the transition from juvenile to 

mature wood, the occurrence of branches and reaction wood, 

and growth rate, which are partly related to species and 

partly influenced by the environment. 

Change in strength with height—Change in strength with 

height and size effect are two overlapping yet independent 

issues used to characterize the perceived reduction in 

strength with pole size. When ANSI O5.1 adopted the Hclasses in 1972, the specification treated these classes the 

same as the distribution pole classes in that the poles were 

sized as single pole cantilever beams. Unlike distribution 

poles, however, many of these larger poles were used in 

multi-pole structures (such as “H-frames”), which required 

cross-bracing that imposed critical stresses at close to midheight above groundline. The concern for critical stress 

occurring at or above mid-height prompted Bohannan (1971) 

to evaluate change in pole strength with height and to propose a strength–height adjustment that defined the lower 

boundary of strength change with height on the basis of the 

limited data available at the time. This model took the form 

⎟

⎠

⎞ ⎜

⎝

⎛ = − L

X Fx Fgl 1 0.5 (6) 

In developing this model, Bohannan reviewed the limited 

available data on change in pole strength with height. The 

ASTM pole data provided data on failure strength as a function of location. Other sources provided information on 

incremental tests of pole sections from the butt to the tip and 

on specific gravity changes with height (Okkonen and others 

1972). After collecting this information, Bohannan derived 

his model to characterize the lower boundary of strength as a 

function of location with respect to pole length. The model, 

based on measured dimensions of test poles, was intended to 

provide a conservative estimate of the strength of the pole at 

height x above the groundline. Annex A, which was introduced in 1976 and added in 1979, incorporated Bohannan’s 

model as nonmandatory information. It stated that the basic 

premise of the model is that it would be used only when 

actual dimensions are known and acknowledged that a reduction in fiber stress with height is not required when analyzing a pole based on minimum size and assumed linear 

taper values given in ANSI tables. 

Soon after appendix A was added, this adjustment became a 

major issue of concern among ASC members as it held 

implications beyond the concern for multi-pole structures. 

Members realized that strength reduction also had an impact 

on smaller poles. One member showed that a 60-ft class-4 

southern pine or Douglas-fir pole would have a critical 

section 26 ft above the groundline and require 3.5-in. greater 

circumference at that point. 

Loss of strength with height continued to be an issue of 

debate into the early 1980s despite the fact that it was 

published in an annex and was not an official part of the 

ANSI main standard. 
17 

As part of the EPRI-sponsored pole research conducted in 

the 1980s, undamaged sections of test poles were retested to 

assess change in strength with height. The retest sample 

included 58 southern pine, 87 Douglas-fir, and 32 western 

redcedar poles. Strength change with height based on measured dimensions was compared to strength change based on 

ANSI minimum dimensions. Results confirmed that the 

conservatism inherent in the use of ANSI minimum dimensions and the assumption of a linear taper countered any 

strength loss with height for Douglas-fir and western redcedar. Bodig and Goodman (1986) and Phillips and others 

(1985) confirmed the efficacy of Bohannan’s model for 

southern pine when based on measured dimensions. 

When the data are reevaluated using minimum class dimensions, however, the relationship becomes 

⎟

⎠

⎞ ⎜

⎝

⎛ = − L

X Fx 9400 1 0.176 (7) 

This means the strength reduction is roughly 9% by midheight when based on minimum class dimensions compared 

to 25% with actual dimensions. 

Size effect—Weibull (1984) developed his extreme value 

distribution on the basis of a weak-link theory that predicts 

reduction in mean strength as size increases. Wood (1956, 

1958) first noted a size effect in poles while working on the 

ASTM pole study and reported the effect for poles less than 

50 ft in length. Erickson and Dohr (1958) showed a decrease 

in tree strength with age: trees more than 200 years old have 

lower strength than younger trees. The effect of size on load 

capacity has been noted for most engineered applications of 

wood (Bohannan 1966). 

Figure 2 shows relative strength values for various species 

and pole size classes. The relative value in this case is expressed as a fraction of the 1963 DFS for southern pine 

(8,000 lb/in2

). Figure 2 suggests that for the ASTM tests, 

relative strength varies with pole size. However, it must be 

noted that the class 1-55 and 9-25 plots represented only 

5 or 6 tests in each species, whereas the class 6-30 samples 

ranged from 15 to 43 tests. This observation did not attract 

much debate in 1963 when the ASA standard dealt only with 

distribution poles and design capacity was a relatively minor 

issue. It did, however, attract attention in 1980 when test 

values from the EPRI-sponsored study were low by comparison with ASTM values. 

The change in strength with pole size is attributed in part to 

geometry and tree physiology. As noted previously, maximum stress in a tapered round cantilever of uniform properties will occur at the location where the circumference is 

1.5 times that of the point of load. For poles of a given 

groundline dimension, the longer the pole the greater the 

chances of failing at a location significantly above groundline. The greater the distance between groundline and the 

failure location, the greater the assumed difference between 

groundline strength and groundline stress at failure. Because 

the poles in the ASTM study were evaluated on the basis of 

groundline stress at failure and not actual groundline 

strength, there is some question as to the source of the 

apparent size effect seen in that data. However, the majority 

of these poles would be expected to fail close to the 

groundline. 

At the time the EPRI study results were published, the 

evaluation of groundline stress at failure was still considered 

to be the easiest way to deal with the evaluation of pole 

capacity. A tendency for a decrease in groundline stress at 

failure with an increase in pole size presented in the EPRI 

Figure 2—Relative strength (fraction of 8,000 lb/in2

) of five species in three size classes tested in 

ASTM wood pole research. LL is longleaf pine, SL shortleaf pine, Lblly loblolly pine, WL western larch, 

DF Douglas-fir, WRC western redcedar, and LPP lodgepole pine. 
22 

study raised some concern among committee members about 

acknowledging these results in the material specification. At 

that time, the ANSI O5.1 standard attributed no design 

significance to DFS values. These values were used only as 

a basis for establishing the pole circumference 6 ft from the 

butt for the ANSI pole size classification. The premise of 

species interchangeability inherent in this classification 

system was not violated if size effect was treated as a design 

consideration, rather than a material variable. The committee 

decided to acknowledge the EPRI results in Annex C as 

additional information for those interested in promoting the 

application of probability-based design to utility structures. 

In this way, the burden of using or ignoring these effects was 

left on the designer with relatively minor impact on the 

ANSI standard. Without code recognition by the NESC, 

probability-based design was not used for design. Nevertheless, interest in the Annex C values continued and this information provided the basis for calibration studies of 

competing design methods. 

Annex C 

Results of the EPRI-sponsored research to promote the use 

of probability-based design for wood transmission structures 

were summarized in the form of a condensed introductory 

guideline and published as Annex C of the ANSI O5.1 standard specification. This guideline was written to be compatible with the established groundline capacity classification 

system presented in the ANSI O5.1 standard. To address the 

size effect made apparent by this analysis, the technical 

advisory committee to that project recommended separating 

the data into two distinct length categories (<50 ft, ≥50 ft) 

with a discrete reduction in strength at the 50-ft length rather 

than adopting a continuous function to account for size 

effect. 

In the evaluation of these strength values, all data were 

adjusted to the minimum groundline dimensions corresponding to the ANSI class of each pole rather than using the 

measured dimension. The groundline circumferences used to 

determine groundline bending strength were the ANSI-class 

minimum values. These are the values an engineer assumes 

for an “ANSI pole” in calculating pole moment capacity. 

The values presented in Annex C are therefore appropriate 

only for use with ANSI-class pole minimum dimensions. 

In addition to differentiating by length and adjusting for the 

ANSI-class minimum dimensions, the larger pole values 

were adjusted to remove any pole size bias. As the EPRI 

study indicated, the effect of groundline circumference on 

pole strength and test poles was greater for western redcedar 

(largest pole H4-70) and Douglas-fir (largest pole H2-75) 

than for southern pine (largest pole 1-75), leading to the 

conclusion that the southern pine data had an inherent size 

bias. As a result, ASC O5 felt that all test values should be 

adjusted to the equivalent class 2-65’ circumference value. 

This was done using a linear regression between groundline 

stress at failure adjusted for ANSI minimum dimension 

(AMORgl) and the ANSI 6-ft-from-butt circumference (c) 

 AMORgl = a + bc (8) 

Regression parameters (a, b) are provided in Table 7 along 

with the ANSI minimum circumference at groundline for the 

class 2 65-ft poles (46.5 for Douglas-fir and southern pine, 

51.5 for western redcedar). Groundline stress values derived 

for poles greater than or equal to 50 ft in length (Bodig and 

others 1986a,b) are shown in the last column. The difference 

between mean distribution pole values (test poles ranging 

from 25 to 50 ft) (ANSI O5.1; App. C, Table C1) and the 

values for an ANSI class 2 65-ft pole (ANSI O5.1; App. C, 

Table C2) represents the effect of pole size on groundline 

stress at failure. 

Despite the effort to keep design information separate from 

the material specification, the ASC O5 was forced to address 

disparities highlighted by the Annex C guideline. An initiative by the Institute of Electrical and Electronic Engineers 

(IEEE) to revise utility structure design procedures of the 

NESC raised questions about the efficacy of the ANSI DFS 

values for design of transmission structures. The primary 

concerns included the effect of pole size and the bending 

strength values found for western redcedar poles. The IEEE 

requested a recommendation from the ASC O5 on values to 

use for design to retain a uniform reliability across species 

and class sizes. This prompted the ASC O5 to consider 

options for recognizing disparities between the ANSI fiber 

stress values adopted in 1963 and those determined on the 

basis of larger pole tests summarized in Annex C. An initial 

recommendation to recognize the effect of pole size was to 

use a stepped allowable fiber stress. With this option, large 

poles would have a single, reduced DFS that would be less 

than the published value for small poles. This would result in 

a sudden jump in the required circumference when that 

transitional pole size was reached. It seemed awkward and 

unrealistic that a slightly larger pole would suddenly have 

less strength, and this too was rejected. 

Table 7—Regression parameters to account for size 

effect on poles (Bodig and others 1986) 

Species 

Sample 

size a b 

ANSI 2-65 

circumference 

(in.) 

AMORgl

(lb/in2

) 

Douglas-fir 248 12,332 −96.1 46.5 8,890 

Southern pine 263 12,589 −89.4 46.5 10,190 

Western 

redcedar 

487 7,952 −53.5 51.5 6,310 

18 
21 

Wolfe and others (2001) proposed the use of a continuous 

function that would maintain the traditional approach to 

pole capacity evaluation, but it was rejected as too complicated and too costly for pole producers. This approach retained a relationship between the groundline moment 

capacity (Mgl) and groundline circumference (Cgl) similar to 

that used previously (Eq. (9)) but with parameters that were 

derived empirically from a statistical evaluation of the 5% 

lower exclusion bending strength versus groundline circumference (Eq. (10)). 

3

gl gl M = σkC (9)

3

gl gl

+ = B M AkγC (10) 

For poles used only in distribution structures, the value of 

“B” was set to zero giving an equation identical to 

Equation (9) except that the A parameter represented a 

L5%EL value for bending strength. A calibration factor γ is 

also included to adjust for the lower values found for western redcedar. 

For the three species used for transmission structures (southern pine, Douglas-fir, and western redcedar), the B parameter had a negative value. This means that the groundline 

moment capacity did not increase as rapidly with an increase 

in groundline circumference as it did using Equation (9), 

thus accounting for the observed size effect. 

This approach retained the simplicity of the ANSI approach 

to pole classification as well as species interchangeability, 

but it required too many changes to the established ANSI 

values. The perception that this would cause significant 

changes in class sizes prompted many producers to conclude 

that transition costs would be prohibitive. Targeting a lower 

5% exclusion value for bending rather than a mean value 

meant that the class loads would have to be reduced. Despite 

the fact that this could be done with little impact on class 

circumference values and would be compensated by warranting higher resistance factors in design, it gave the appearance of reducing the capacity of wood poles. ANSI DFS 

values had traditionally represented a value close to the 

mean strength at groundline for distribution poles: 5% lower 

exclusion values are roughly 30% less than the traditional 

DFS values. Finally, this approach acknowledges a larger 

gap between the ANSI DFS and measured mean strength of 

western redcedar than exists for the other species. Members 

of ASC felt this was not adequately substantiated. They 

explained the strength gap as the result of deficiencies in the 

test and analysis procedures when evaluating a species with 

large taper in the lower half, which caused failures to occur 

higher in the pole giving a groundline moment at failure that 

was not representative of groundline moment capacity. 

ANSI O5–02 

Rather than change the method of deriving DFS values and 

the tabulated circumferences, ASC O5 opted to change the 

definition of DFS from a relative measure of pole groundline 

moment capacity to an estimate of groundline strength (see 

Appendix A). Prior to 2002, the DFS was evaluated as the 

groundline stress at failure for a dry, treated pole loaded as a 

cantilever beam with a transverse load applied 2 ft from the 

top. The sole ANSI O5.1 use of this value was to represent a 

relative measure of pole capacity as a basis for pole size 

classification. A number of attempts were made by committee members (Colley 1932, Wilson and others 1953) to 

attach a level of statistical significance to the derivation, but 

their assessments were all based on assumptions that were 

never supported by data or any rigorous analysis of field 

performance. The ASA and later the ANSI standards were 

strictly material specifications that presented the DFS values 

as only a relative measure of pole strength. 

In 2002, the DFS was redefined as the fiber strength of dry, 

treated wood. The ANSI standard added statements of significance that these values represent the means of distribution of groundline fiber strength and that these distributions 

are known to have a COV of 20%. The standard states that 

these evaluations were made with sufficient precision to be 

used for probability-based design. The new definition introduced an acceptance of design liability by ASC O5 as the 

DFS value is now defined to represent wood strength at a 

specific location as opposed to relative pole capacity used 

simply to set pole size classes. 

Using data provided by the EPRI project (Phillips and others 

1985, Bodig and others 1986) (Appendix C) the committee 

derived groundline strengths (MORgl) for poles in lengths of 

50 ft and longer on the basis of measured modulus of rupture 

(strength) at the failure location (MORfl) and the strength–

height relationship proposed by Bohannnan (1971). Prior to 

2002, DFS values included adjustments for variability in test 

data attributable to handling and test procedures (0.93), 

high-temperature conditioning, and drying in service. This 

approach was retained for poles shorter than 50 ft based on 

the assumption that they normally fail at the groundline. For 

the 2002 revision of the ANSI O5.1 standard, however, the 

ASC O5 supported dropping the adjustment for variability 

and accepting a maximum 10% increase for drying in service, originally recommended by Wood and Markwardt 

(1965) for longer poles. 

A new adjustment, also adopted for long poles, was a classoversize adjustment that varies from 7% to 13%, depending 

on the pole class size. Due to its variation with class size, 

this adjustment requires an iterative procedure to match the 

product of adjusted DFS and average class section modulus 

to a designated class load moment (discussed in Appendix 

C). This adjustment decreases the probability that poles 

within a size class will have a groundline bending moment 

19 
22 

capacity that exceeds that imposed by the designated class 

load (see Appendix A). The resulting values are shown in 

the column labeled “DFS” in Table 1. Comparing these DFS 

values to the DFS values that would have been derived using 

the previous definition shows a 14% advantage for Douglasfir and southern pine and a 44% advantage for western redcedar. 

In 2002, groundline strength of poles in lengths 50 ft and 

longer were extrapolated from the failure location using the 

strength–height model previously presented in Annex A of 

the ANSI O5.1 standard. As change in strength with height 

is now an inherent part of the evaluation of groundline 

strength, it is required that this model be used when assessing the strength profile for ANSI poles. Adoption of this 

model to estimate groundline strength, however, detracts 

from species interchangeability. Groundline circumference 

values are still calculated on the basis of section modulus 

required to resist the class-load-induced groundline moment. 

Using the projected groundline fiber strength rather than the 

groundline stress at failure means that the groundline dimension does not account for the occurrence of a failure above 

groundline. The larger the taper of a pole, the greater the 

distance from groundline to the point of maximum stress and 

the greater the difference between groundline stress at failure 

and groundline strength. Unless the groundline circumferences are adjusted to account for this, poles of the same class 

but different taper will have different load capacity 

While the decision to change the DFS value from a relative 

measure of pole capacity to a measure of bending strength 

applicable only at groundline constituted a significant 

change to the standard, on its own this decision had a minor 

impact on solving the problem of size effect that was apparent in the test data. The change that had the greatest impact 

on maintaining the traditional DFS values was the adoption 

of a classification oversize adjustment. 

The analysis presented in Appendix C shows that the projected MOR at groundline, when adjusted for only conditioning and drying in service (5,530, 6,950, and 7,720 lb/in2

for western redcedar, Douglas-fir, and southern pine, respectively) is still significantly less than the ANSI DFS values 

for these species (6,000, 8,000, and 8,000 lb/in2

, respectively). Adjusting DFS values so that the average, rather than 

minimum-sized, pole in a class will be sized to carry the 

ANSI-designated class load gives DFS values of 6,040, 

7,640, and 8,530 lb/in2

, respectively. 

The effects of these changes on ANSI values are discussed 

in Appendix A. Small poles are still sized using DFS values 

that were derived assuming a 16% increase for drying in 

service but with no increase for class oversize. Long poles 

(>50 ft) are assumed to have a dry strength 10% greater than 

the measured green value. But they are also given a 7% to 

13% increase for class oversize. Despite the fact that poles 

are sized using values adjusted for drying in service, ANSI 

notes that the tabulated circumferences are intended to be 

measured in the green condition and that a reduction in 

section modulus of up to 6% due to shrinkage should be 

expected. This means that a minimum-sized pole of average 

strength and more than 50 ft in length could be overstressed 

by as much as 24% (1.1 × 1.13), whereas the averagestrength minimum-sized distribution pole would be overstressed by 16% if required to carry its class load in the 

green condition. If loaded in the dry condition, an average 

strength minimum-sized long pole would be overstressed by 

13% to 20% at the class load (6% due to shrinkage and 7% 

to 13% for the class oversize adjustment), whereas the 

smaller, minimum-sized distribution pole would be only 6% 

under capacity at the designated class load. 

The fact that the probability of failure of the average ANSI 

pole loaded to its class load in a standard cantilever test went 

from 50% or more (depending on strength variability assumed) prior to 2002 to 70% or more after 2002 (see Appendix A) is really not a critical issue as long as the pole 

user realizes that class loads are used simply to classify 

poles. There should be some concern, however, that the new 

definition of DFS for long poles has been inflated by 7% to 

13% above dry strength and a dry pole is likely to have a 

section modulus 6% below that of a green pole. This suggests that the DFS for long poles should be reduced by 20% 

when calculating the groundline strength for design purposes. 

The most obvious drawback to this change is the confusion it 

creates over the meaning of the ANSI fiber stress value. The 

engineer will still use the minimum dimensions permitted for 

the pole class selected to assess capacity. The pole selection 

must be made by first assessing the load profile. Using the 

ANSI dry DFS and the change in strength and circumference 

along pole length, the engineer must determine the required 

minimum section property profile for portions of the pole 

expected to dry in service and then decrease the DFS by 

10% (16% for short poles) to check section property required in any wet regions. Once a required profile has been 

established, it can be compared to ANSI-tabulated dimensions to select a class size. The difference in average failure 

probability at the designated class load for long and short 

poles will most likely be of little importance provided the 

design engineer references the tables simply as a catalog of 

ANSI size codes. 

Conclusions 

• The practice of classifying green poles by groundline 

circumference determined on the basis of dry strength is 

inappropriate for design applications. The ANSI standard 

should derive class dimensions on the basis of green 

strength and suggest an adjustment for drying above 

groundline. 

20 
21 

• The practice of sizing long poles so that the average size 

pole is defined by the minimum required section modulus 

and of sizing short poles so that the smallest permissible 

size pole is defined by the minimum required section 

modulus leads to inconsistency in pole reliability. ANSI 

should apply the same bias to all pole classes. 

• For species having ANSI DFS values derived from 

published small–clear tests, values should be modified 

to reflect the changes implemented in ASTM D 2555. 

• It is more conservative for taper values used to project 

6-ft-from-butt dimensions to err on the high side. Comparison of taper values reported in Annex B of the ANSI 

O5.1 standard appear low by comparison to those determined by Bohannan. This could result in undersizing for 

poles more than 40 ft in length. 

• The relationship between strength and pole size is not 

adequately characterized by the strength-reduction-withheight function adopted by ANSI. The relationship between pole size and strength as evaluated over the past 50 

years suggests that it would be technically more correct to 

reduce the allowable fiber stress with increasing pole size 

(volume) rather adopting an untested model to project 

groundline strength from strength at the failure location. 

• Pole design should focus on differences between stress 

and strength profiles, setting design as a fraction of the 

load at which the maximum difference is zero. The ANSI 

O5.1 standard currently instructs users to consider change 

in strength with height only when the point of maximum 

stress occurs above groundline. The DFS derivation 

adopted in 2002, however, extrapolates strength from the 

point where the difference between stress and strength are 

the greatest (failure location); not the point of maximum 

stress. Suggesting a design approach to prevent stress from 

exceeding strength would add credibility to the design 

method presented and would make wood pole strengths 

comparable to that of other materials. 

• The wood pole industry should continue to work on refining its knowledge of pole physiology and shape to provide 

more reliable wood pole structures and a more credible 

standard. Also, if the producers can be encouraged to revise the class circumference tables, then interspecies exchangeability can be restored. 

Observations 

The ANSI O5.1–02 standard appears to magnify the disparities between the main body of the standard and Annex C. 

Annex C presents what has been viewed as technically correct and compatible with established timber engineering 

methods compromised by what is politically and economically acceptable to users and producers. By keeping design 

information in an annex, separate from the material specification, the committee was able to maintain the standard size 

classifications established 40 years ago, despite the realization that they did not represent uniform reliability across 

species or across classes within a given species. 

The methods outlined in Annex C brought ANSI up-to-date 

with utility-structures engineering, but they suggest a need to 

recognize a size effect when designing wood pole transmission structures. The committee responsible for the ANSI O5 

standards (ASC O5) has attributed the perceived problem of 

a size effect to change in strength with height. The committee must decide if it wants ANSI O5.1 to remain a material 

specification or to assume the role of a wood pole design 

standard. The standard needs a stronger technical basis to 

bring it into alignment with Annex C. 

Before ANSI O5.1 moves too far into the realm of a design 

standard, it would be beneficial to develop technical support 

for assumptions made and to more carefully consider their 

impact on the market for utility poles. The fact that poles are 

generally classified in the green condition means that 

groundline section modulii derived using the new “dry” 

definition of DFS values for large poles may result in undersized poles. A size class adjustment is appropriate only when 

converting the DFS, used to set size class dimensions, to a 

design value. It is not appropriate to adjust test data for size 

class effect to derive a size class dimension as it leaves the 

false impression that all poles having a strength greater than 

or equal to the DFS in a given size class have the capacity to 

carry the designated class load. 

If loads on a large pole are located to induce maximum stress 

at groundline, where moisture could be at or above fiber 

saturation point, a pole whose required groundline dimension is derived on the basis of dry strength will be undersize. 

Thus, in addition to including a reduction in strength with 

height, the standard should include a moisture adjustment for 

critical sections being within 4 ft of the groundline. 

The fact that this new derivation requires recognition of a 

change in strength with height means that this becomes 

mandatory for all poles. The committee is basically assuming that small poles fail close to the ground where change in 

strength with height is minimal. The strength change profile 

(Eq. (7)) is undoubtedly conservative in the lower half of 

trees; even for small poles it indicates that the critical stress 

location is above groundline. The result is that the strength 

change with height evaluation is required for any poles 

designed using the new ANSI standard. If the committee 

wishes to design poles with this degree of sophistication, the 

strength profile must be derived to present a more accurate 

picture, not only of fiber strength change but of section 

property change. 
22 

Literature Cited 

ANSI. 1997. Specifications and dimensions for wood poles. 

New York, NY: American National Standards Institute. 

ASTM. 1966. Standard methods for establishing clear wood 

strength values. Standard D 2555–66. West Conshohocken, 

PA: American Society for Testing and Materials. 

ASTM. 1999. Standard test methods of static tests of wood 

poles. Standard D 1036–99. West Conshohocken, PA: 

American Society for Testing and Materials. 

ASTM. 2000. Standard test methods for small clear specimens of timber. Standard D 143–94. West Conshohocken, 

PA: American Society for Testing and Materials.

Betts, H.S.; Newlin, J.A. 1915. Strength tests of structural 

timber treated by commercial wood preserving processes. 

Bull. 286. Washington, D.C.: U.S. Department of Agriculture. 

Bhuyan, G.; Chetwynd, D. 1994. Strength and stiffness of 

new wood utility poles. Rep. 119D 1010. Montreal, Canada: 

Canadian Electricity Association. 

Bodig, J.; Goodman, J.R. 1986. Western redcedar data and 

size effect. Wood pole properties. Review and recommendations for design resistance data. EPRI EL–4109, Vol. 3. Palo 

Alto, CA: Electric Power Research Institute. 

Bodig, J.; Goodman, J.R.; Brooks, R.T. 1986a. Western 

redcedar data and size effect. Wood pole properties. 

Review and recommendations for design resistance data. 

EPRI EL–4109, Vol. 3. Palo Alto, CA: Electric Power 

Research Institute. 

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

1986b. Douglas-fir data—Wood pole properties. Review 

and recommendations for design resistance data. EPRI 

EL–4109, Vol. 2. Palo Alto, CA: Electric Power Research 

Institute. 

Bohannan, B. 1966. Effect of size on bending strength of 

wood members. Res. Pap. FPL–RP–56. Madison, WI: U.S. 

Department of Agriculture, Forest Service, Forest Products 

Laboratory. 

Bohannan, B. 1971. Height–strength relationship for round 

timbers. Working document for ANSI O5.1 committee use. 

Bohannan, B.; Habermann, H.; Lengel, J.E. 1974. Taper 

of wood poles. Gen. Tech. Rep. FPL–GTR–2. Madison, WI: 

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

Products Laboratory. 

Buckman, S.J.; Rees, L.W. 1938. Effect of steaming on 

strength of southern yellow pine. In: Proceedings, American 

Wood Preservers’ Association. 34: 264–296. 

CSA 1990. Wood utility poles and reinforcing stubs. 

CAN/CSA O15–90. Etobicoke, Ontario: Canadian Standards 

Association. 

Colley, R.H. 1932. Ultimate fiber stresses for wood poles. 

Bell Telephone System, Tech. Pub. Monograph B615. New 

York, NY: Bell Telephone Laboratories. 

Eggleston, R.C. 1952. Pole strength tests. Forest Products 

Journal. April, pp. 3–24. 

Erickson, E.C.O.; Dohr, A.W. 1959. Effect of steam conditioning on the strength of treated longleaf pine poles. Interim 

rep. Madison, WI: ASTM Wood Pole Program P&E. 

Goodman, J.R.; Vanderbilt, M.D.; Criswell, M.E.; 

Bodig, J. 1981. Probability based design of wood transmission structures—strength and stiffness of wood poles. EPRI 

EL–2040, Vol. 1. Palo Alto, CA: Electric Power Research 

Institute. 

Jones, R.L. 1931. New standard specifications for wood 

poles. Bell System Technical Journal, September, 

pp. 514–524. 

Kulp J.W. 1957. Second report on tests of service life of 

poles in REA-financed electric systems. REA Bull. 169–32. 

Washington, D.C.: U.S. Department of Agriculture, Rural 

Electrification Administration. 

MacLean, J.D. 1951. Rate of disintegration of wood under 

different heating conditions. In: Proceedings, American 

Wood Preservers’ Association. 47: 155–168. 

MacLean, J.D. 1953. Effective of steaming on the strength 

of wood. In: Proceedings, American Wood Preservers Association. 49: 88–112. 

National Bureau of Standards. 1927. National Electrical 

Safety code, initiated 1913. 

Newlin. J.A.; Trayer, G.W. 1924. Form factors of beams 

subjected to transverse loading. Rep. 181. National Advisory 

Committee for Aeronautics. Reprinted October 1941 as Rep. 

1310. Madison, WI: U.S. Department of Agriculture, Forest 

Service, Forest Products Laboratory. 

Okkonen, A.; Wahlgren, G.; Maeglin, R. 1972. Western 

wood density survey. Forest Products Journal. 22(7): 37–41.

Phillips, G.E.; Bodig, J.; Goodman, J.R. 1985. Background and southern pine data. Wood pole properties. Review and recommendations for design resistance data. EPRI 

EL–4109, Res. Proj. 1352–2, Vol. 1. Palo Alto, CA: Electric 

Power Research Institute. 

Rhatigan, R.G.; Morrell, J.J. 2002. Seasonal variations in 

moisture content of in-service poles in the Willamette Valley. In Proceedings, International conference on utility line 

structures. Ft. Collins, CO: Engineering Data Management 

Inc. and Colorado State University. 
21 

Stamm, A.J. 1956. Thermal degradation of wood and 

cellulose. Industrial Engineering Chemistry. 48: 413–417. 

Thompson, W.S. 1969. Factors affecting variation in compressive strength of southern pine piling. In: Proceedings. 

Washington, DC: American Wood Preservers’ Association. 

USDA. 1940. Wood handbook. Washington, D.C.: U.S. 

Department of Agriculture, Forest Service, Forest Products 

Laboratory. 326 p. 

USDA. 1955. Wood handbook. Washington, D.C.: U.S. 

Department of Agriculture, Forest Service, Forest Products 

Laboratory. 528 p. 

Wilkinson, T.L. 1968. Strength evaluation of round timber 

piles. Res. Pap. FPL–RP–101. Madison, WI: U.S. Department of Agriculture, Forest Service, Forest Products 

Laboratory.

Wilson, T.R.C. 1923. Results of some strength tests on 

wooden poles. Madison, WI: U.S. Department of Agriculture, Forest Service, Forest Products Laboratory. 

Wilson, T.R.C.; Drow, J.T. 1953. Fiber stresses for wood 

poles. No. D1619. Madison, WI: U.S. Department of Agriculture, Forest Service, Forest Products Laboratory. 

Wilson, T.R.C.; Carlson, T.A.; Luxford, R.F. 1930. The 

effect of partial seasoning on the strength of wood. In Proceedings. Granbury, TX: American Wood Preservers’ Association.

Wolf, J. 1979. Bending strength of roundwood timber poles: 

a summary and evaluation of literature. Unpublished. Madison, WI: U.S. Department of Agriculture, Forest Service, 

Forest Products Laboratory. 

Wolfe, R.W. 2000. Design stress derivation for ANSI poles. 

In: Proceedings, International conference on utility line 

structures; 2000 March. Ft. Collins, CO: Colorado State 

University and Engineering Data Management Inc. 

Wolfe, R.W.; Bodig, J.; Lebow, P.K. 2001. Derivation of 

nominal strength for wood utility poles. Gen. Tech. Rep. 

FPL–GTR–128. Madison, WI: U.S. Department of Agriculture, Forest Service, Forest Products Laboratory. 

Wood, L.W. 1956. Testing poles for the American Society 

for Testing and Materials. Interim rep. P&E 134. Madison, 

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

Forest Products Laboratory. 

Wood, L.W. 1958. Strength tests of wood poles. Interim 

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

Service, Forest Products Laboratory. 

Wood, L.W.; Markwardt, L.J. 1965. Derivation of fiber 

stresses from strength values of wood poles. Res. Pap. 

FPL–RP–39. Madison, WI: U.S. Department of Agriculture, 

Forest Service, Forest Products Laboratory. 

Wood, L.W.; Erickson, E.C.O.; Dohr, A.W. 1960.

Strength and related properties of wood poles. ASTM final 

rep. Conshohocken, PA: American Society for Testing and 

Materials. 

23 
28 

Appendix A—Effect of 2002 

Revision on Pole-Class 

Strength Distribution 

In 2002, six basic changes were made to the ANSI standard 

that influence the definition of the designated fiber stress 

(DFS) for poles in lengths greater than or equal to 50 ft in 

length (referred to as “long” poles in this appendix). These 

changes were as follows: 

1. Redefining DFS to be groundline strength rather than 

groundline stress at failure 

2. Adopting a change in strength with height function as 

part of the evaluation of pole properties 

3. Dropping recognition of the influence of test procedure 

variability from derivation of strength of long poles 

4. Incorporation of a class oversize adjustment for long 

poles 

5. A reduction in the long pole strength adjustment for 

drying in service 

6. Declaration of a 20% coefficient of variation for pole 

bending strength 

The change in definition of DFS allowed ANSI to acknowledge the observed “size effect” (Bodig and Goodman 1986) 

discussed in Annex C of the O5.1 standard as being a result 

of a change in strength with height rather than a groundline 

dimension. The advantage of this change is that it provides 

a basis for retaining established DFS and therefore the 

ANSI class groundline dimensions for all pole classes. 

To justify maintaining the same DFS for long poles, ANSI 

adopted as mandatory the change-in-strength-with-height 

function previously presented in the “non mandatory” Annex A. The inverse of this function was used to estimate 

groundline strength for long test poles that failed above 

groundline. As the inverse of this function was used to 

evaluate groundline strength, the function becomes a mandatory adjustment for anyone concerned about the strength 

of poles in locations other than the groundline. 

The 2002 derivation of DFS for long poles did not include 

the adjustment for variability inherent in the established 

DFS values. The “variability” adjustment, originally introduced in 1931 and re-approved in 1963, was proposed to 

counter the effect of testing poles that had unknown moisture gradients at the time of testing. This was basically a 

confidence adjustment that reduced the measured mean by a 

half standard deviation. In 1963, ASA O5 elected to use a 

14% coefficient of variation for bending strength of poles 

on the basis of results published in the ASTM wood pole 

report (Wood and others 1960). As there were no changes 

to the derivation of DFS for poles under 50 ft in length in 

2002, the variability adjustment (0.93) is still considered to 

be an inherent part of those values. This adjustment was not 

used, however, in the 2002 derivation of DFS for long poles 

(≥50 ft). 

In 2002, ANSI adopted an adjustment for class oversize 

effect for long poles. This adjustment was previously acknowledged as a design consideration (Annex C of the 

ANSI O5.1 standard), but it was not included in the previous derivation of DFS values, which are used to set the pole 

class dimensions. This adjustment basically realigns the 

designated class load with the median-size rather than the 

minimum-size pole in a given size class. Multiplying the 

minimum allowable section modulus for a pole class–length 

combination by the DFS thus gives the moment capacity of 

the median “long” pole and the minimum “short” pole. This 

suggests that for a given pole class, an average strength 

median sized “long” pole and an average strength minimum 

sized “short” pole have the groundline moment capacity 

implied by their designated class load. 

Other changes adopted in 2002 that affect the conclusions 

about the ability of a pole to resist the implied groundline 

moment capacity include the adjustment for drying in service and the assumed strength variability. In 1963, two 

drying-in-service adjustments were considered, 1.1 and 

1.16. The latter value was preferred for derivation of DFS 

for the primary pole species. In 2002, an adjustment of 1.1 

was adopted for long poles. The implication of this change 

is that short poles that are more likely to be critically 

stressed at groundline have a larger increase for drying inservice than are poles more likely to fail above ground. 

As for the variability of pole bending strength, the 2002 

version of ANSI O5.1 declares that wood poles have a 

strength coefficient of variation of 20%. This value is 

higher than previously referenced by ANSI and higher than 

that obtained for data presented in the ASTM study (Wood 

and others 1960) or the EPRI-sponsored study (Bodig and 

others 1986). 

The following analysis provides a comparison of pole 

strength as defined by ANSI standards before and after the 

2002 revision. All assumptions are based on values either 

provided in ANSI O5.1 standards or taken from discussions 

that support those values. This analysis includes only those 

parameters that were modified in 2002. It assumes that 

removing adjustments for variability, drying in service, and 

class oversize effects from ANSI DFS values truly represents mean groundline strength of green poles adjusted for 

the effects of pretreatment conditioning. 

This analysis also focuses only on the strength at the 

groundline. There is little support for the premise that 

24 
27 

strength change with height is the same for all species and 

pole classes, and we know that tree geometry varies with 

species as well as growing conditions. For this comparison 

of effects of the 2002 change, it is not necessary to address 

that issue except to acknowledge that the adjustment for 

change in strength with height function adopted by ANSI 

adds a degree of conservatism to the analysis of pole capacity. 

Table A1 lists normalized distribution parameters for pole 

strength implied by the ANSI derivation of DFS. Three 

variables listed (drying in service, variability, and class 

oversize adjustment) are those that vary between “long” and 

“short” poles in the 2002 standard and between long poles 

before and after 2002. The “normalized mean” value is the 

ratio of the mean strength in the green treated condition to 

DFS. The normalized strength distributions thus have a 

mean value (λ) with the same coefficient of variation (COV 

= Ω) as the actual pole strengths. For example, on average, 

a long, class 2, “green” pole will have a mean strength that 

is 82% of the DFS value and a standard deviation (λΩ)of 

0.16. Mean strength is less than DFS because the DFS 

includes increases for drying and class oversize. Short poles 

have a mean strength of 93% of the DFS because of the 7% 

reduction for variability in combination with the 16% increase for drying (1/[(1.16)(0.93)]). 

A distribution of pole moment capacities can be generated 

using a method known as random-products generation. This 

is done by randomly choosing a normalized strength from a 

normal distribution having a mean value λ and a coefficient of 

variation Ω and multiplying by a randomly chosen normalized section modulus (random class value divided by the 

minimum required for the class). The section modulus is 

selected assuming a uniform distribution with a range from 

1.0 to the maximum for the pole class (see Table A2). Repeating this process 6,000 times for each of the four 

scenarios provides a basis for evaluating the effect of the 

2002 changes on implied pole capacity. 

As the revisions for long poles incorporated in 2002 shift 

average capacity from the minimum-size to the median-size 

pole in the group, it is not surprising that fewer poles within 

a pole class have the groundline moment capacity implied 

by the designated class load. The comparison of load capacity distributions shows that at the 5th percentile, the load 

capacities for the H6 and the class 2 long poles were 57% 

and 55% of the implied capacity, respectively. For the short 

poles, the 5th percentile value was 62% of implied capacity; 

for class 2 poles evaluated prior to 2002 ANSI parameters, 

poles at the 5th percentile had 73% of the implied capacity. 

The mean strength of the H6 and class 2 long poles was 

91% of the implied capacity. As stated earlier, this is due to 

the use of a dry strength being used to size a green pole. 

Values for short poles and the pre-2002 analyses showed 

the mean strength to be 102% of implied capacity. 

Overall, only 35% the distribution of long poles was capable of carrying the implied groundline moment. The 2002 

analysis of short poles and the pre-2002 analysis resulted in 

more than 54% of poles in a given class having capacity 

that exceeded that implied by the designated class load. 

This analysis suggests that the ANSI 2002 change created a 

gap in the inherent capacity of short and long poles. This is 

of minor consequence for ANSI O5.1. Pole sizes corresponding to any ANSI class are the same as they were prior 

to 2002. For the utility engineer, however, it is noteworthy 

that only slightly more than a third of the long poles selected on the basis of groundline moment divided by the 

ANSI DFS have the groundline moment capacity to resist 

that DFS in the green condition. 

For those concerned about the future use of round poles as 

engineered components, there is a definite advantage to 

cooperating with those committees responsible for design of 

utility structures (ANSI C2/NESC and ASCE committee on 

wood) to assure that they understand the basis for the difference between the ANSI DFS as applied to “long” and 

“short” poles. 

Table A1—Normalized distribution parameters for 

ANSI poles before and after 2002 

Distribution parameter 

Variablea

 Long poles b 

 H6 2 

Short poles c 

 All 

Pre-2002 

All 

Drying in 

service 

 1.10 1.10 1.16 1.16 

Variability 1.0 1.0 0.93 0.93 

Class oversize 1.06 1.11 1.0 1.0 

Strength COV 0.2 0.2 0.2 0.14 

Normalized 

mean (λ) 

 0.858 0.819 0.927 0.927 

Normalized SD 0.172 0.164 0.216 0.151 

a

 SD is standard deviation.

b Poles ≥50 ft. c

 Poles <50 ft. 

Table A2—Maximum value for uniform distribution of 

normalized section moduli 

Class 2 1 H1 H2 H3 H4 H5 H6 

Max 1.22 1.20 1.19 1.17 1.16 1.15 1.14 1.13 

25 
28 

Appendix B—Wood Pole Test 

Data Sources 

Alexander, J.B. 1934. Strength of lodgepole pine telephone 

poles. Circular 42. Vancouver, B.C.: Western Forest Products Laboratory. 

Alpin, E.N. 1967. Strength of red pine poles. Rep. No. OP–

X–10. Ottawa, Ontario: Eastern Forest Products Laboratory. 

Anon. 1947. Strength of telephone poles of eastern cedar, 

red pine, and jack pine. Circ. 31. Ottawa, Ontario: Eastern 

Forest Products Laboratory. 

Betts, N.W.; Hein, A.L. 1914. Tests of Rocky Mountain 

woods for telephone poles. Bull. 67. Washington, D.C.: 

U.S. Department of Agriculture, Forest Service. 

Bodig, J.; Goodman, J.R. 1986. Wood pole properties. 

Vol. 3, Western redcedar data and size effect. EPRI EL–

4109, Res. Project. 1352–2. Palo Alto, CA: Electric Power 

Research Institute. 

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

1986. Wood pole properties. Vol. 2, Douglas-fir data, EPRI 

EL–4109, Res. Proj. 1352–2. Palo Alto, CA: Electric Power 

Research Institute. 

East Kentucky Power Cooperative Inc. 1980. Eleven new 

poles. Tested at Colorado State University, Fort Collins. 

Fox, S.P. 1977. Strength of ACA and PCP treated white 

spruce, western hemlock, and amabilis fir. Unpublished 

report. Vancouver, B.C.: Western Forest Products 

Laboratory. 

Goodman, J.R.; Vanderbilt, M.D.; Criswell, M.E.; 

Bodig, J. 1981 Probability-based design of wood transmission structures—strength and stiffness of wood poles. EPRI 

EL–2040, Vol. 1. Palo Alto, CA: Electric Power Research 

Institute. 

Idaho Power Co. 1981 Forty-nine western redcedar poles 

removed from service and tested at Colorado State University, Fort Collins, Colorado. 

Jessome, A.P. 1955. An investigation of the bending 

strength of 25-foot red pine poles by the machine test 

method. Progress Rep. No. 9, Proj. No. 0–74–1. Ottawa, 

Ontario: Department of Northern Affairs and National 

Resources, Forest Products Laboratory. 

Jessome, A.P. 1960. An investigation of the bending 

strength of 35-foot untreated lodgepole pine poles by the 

machine test method. Department of Northern Affairs and 

National Resources Progress. Progress Rep. No. 5, Proj. No. 

0–74–1, Ottawa, Ontario: Forest Products Laboratory. 

Kennedy, D.E.; Wakefield, W.E. 1948. The strength of 

jack pine poles infected with pocket rot. Circ. 65. Ottawa, 

Ontario: Canadian Dept. Mines and Resources, Domestic 

Forest Service, Forest Products Laboratory. 

McElhannay, T.A.; Perry, R.S. 1928. Tests of relative 

strength of green-cut and fire-killed western cedar pole 

timber. Circ. 22. Forest Service, Department of Interior. 

Canada. 

McGowan, W.M. 1960. Strength of Douglas-fir telephone 

poles. Tech. Note 15. Vancouver, B.C.: Western Forest 

Products Laboratory. 

McGowan, W.M. 1962. Strength of western hemlock 

power and communication poles. Tech. Note 27. Vancouver, B.C.: Western Forest Products Laboratory. 

McGowan, W.M.; Smith, W.J. 1965. Strength and related 

properties of western redcedar poles. Dept. Pub. 1108, 

Vancouver, B.C.: Western Forest Products Laboratory. 

Platt, J.C. 1955. Breaking strength of red pine poles. 

Rep. 55–587. Ontario, Canada: Hydro-Electric Power Commission. 

Platt, J.C. 1958. Strength tests of jack pine poles. Rep. 

58–149. Ontario, Canada: Hydro-Electric Power Commission of Ontario. 

Platt, J.C. 1958. The strength tests of wood poles: western 

redcedar and red pine. Rep. 58–206. Ontario, Canada: 

Hydro-Electric Power Commission. 

Phillips, G.E.; Goodman, J.R.; Bodig, J. 1986. Wood 

pole properties. Vol. 1, Background and southern pine. 

EPRI EL–4109, Res. Proj. 1352–2. Palo Alto, CA: Electric 

Power Research Institute. 

Public Service Company of Colorado. 1980. Twenty-five 

new mechanically pretreated poles. Tested at Colorado State 

University, Fort Collins. 

Thompson, W.S. 1969. Effect of steaming and kiln drying 

on the properties of southern pine poles. Part i. Mechanical 

properties. Forest Products Journal. 19(1): 21–28. 

Wilson, T.R.C. 1917. Results of some strength tests on 

wooden poles. Madison, WI: U.S. Dept. of Agriculture, 

Forest Service, Forest Products Laboratory. 

Wood, L.W.; Erickson, E.C.O.; Dohr, A.W. 1960. 

Strength and related properties of wood poles. West 

Conshohocken, PA: American Society for Testing and 

Materials. 

26 
27 

Appendix C—Data Used to 

Derive DFS for Long Poles 

In 2001, Martin Rollins of H.M. Rollins Company proposed 

an alternative approach to accounting for the observed 

“size” effect in wood poles on behalf of the North American 

Wood Pole Coalition. His proposal was unique in that it 

recognized a size effect but did not require any modification 

to the ANSI classification tables. He presented this proposal 

to the ANSI O5.1 committee in the form of a committee 

report titled Evaluation of Fiber Stress Values for Wood 

Utility Poles in ANSI Standard O5.1. Although it never 

received critical review outside the fiber stress subcommittee, the report played a significant role in the committee’s 

decision to adopt proposed changes to the ANSI standard. 

Essentially, Rollins attributed “size” effect to a “height” 

effect. His premise was that the groundline strength does 

not change with increased circumference, but the load 

capacity as characterized by groundline stress at failure goes 

down because longer poles fail further from the ground in 

weaker wood. His solution was to change the definition of 

“designated fiber stress” (DFS) from groundline stress at 

failure to groundline strength. Rather than use a groundline 

strength derived from published small–clear strength test 

data, he decided to estimate groundline strength by adjusting the modulus of rupture (MOR) of full-size poles tested 

in the green condition. The primary adjustment accounted 

for the change in strength with height. This Rollins borrowed from annex A of the existing ANSI standard. Other 

adjustments included a 10% increase for drying in service, a 

pretreatment conditioning adjustment (0.9 for kiln drying 

and boultonizing, 1.0 for air drying), and a classification 

oversize adjustment that varied from 1.07 for a class H4–

70-ft pole to 1.117 for a class 3–60-ft pole (Appendix D). 

The height adjustment was originally derived as a conservative estimate of strength change with height when the actual 

dimensions are known. This adjustment varies linearly from 

1.0 at groundline to 0.75 at midheight. The inverse varies 

from zero for a failure occurring at groundline to 1.33 if the 

failure occurs at midheight. 

The class oversize adjustment is intended to recognize the 

fact that the majority of poles will have a groundline circumference larger than the tabulated minimum value for 

their size class. Any increase to groundline stress results in 

a proportional decrease in the required goundline section 

modulus. The ASC O5 committee voted to apply an increase to the projected groundline strength so that an average strength pole will be required to have the average, 

rather than the minimum, section modulus for its pole class 

to carry its class load. Determination of this increase involves an iterative process. The average groundline strength 

must first be determined on the basis of all pole tests. For 

each pole class–length combination, the standard-test moment divided by the average groundline strength gives the 

initial estimate of groundline section modulus required. 

Poles are initially classified by these minimum section 

modulii. 

The class load is calculated as 

2 P 1200 250i 50i i = + + (C1)

where i is an index corresponding to the size class (0 for 

class 7, 12 for class H6), and Pi is the ANSI-designated load 

for class index. 

The class oversize adjustment is defined as 

( ) /(2 ) Acos = Pi+1 + Pi Pi

This simplifies to 

Pi A 1 50(3 i)/ cos = + + (C2) 

Application of this adjustment requires an iterative evaluation of minimum class size. Test poles are first classified 

using the estimated groundline strength (i.e., 8,000 lb/in2

 for 

southern pine and Douglas-fir) to match the pole to an 

ANSI load class (Pi) (Eq. (C1)) and required minimum 

groundline dimension. The pole’s projected groundline 

strength is then adjusted by Acos (Eq. (C2)). The average of 

adjusted values is used to recalculate minimum required 

section modulii for each pole class, and the test poles must 

be reclassified to recognize possible pole class change. This 

is repeated to convergence, resulting in a groundline 

strength that gives an average within-class section modulus 

the average capacity to carry the designated class load. 

Table C1 shows the values derived at three stages in the 

derivation of the DFS for long poles. These values were 

derived using test data provided by EPRI-funded research 

on transmission-sized poles (Bodig and Goodman 1986, 

Bodig and others 1986b, Phillips and others 1985). MORgl

is an estimate of the average groundline strength as projected from the measured stress at failure. MORgl’ is MORgl

adjusted for drying in service and conditioning. DFS is the 

designated fiber stress which, when multiplied by the minimum section modulus for a designated pole class-length, 

gives the average pole moment capacity for that class 

length. 

Table C2 summarizes the resulting minimum groundline 

circumferences and oversize adjustment factors corresponding to the new DFS values. The oversize adjustment is ratio 

of average to minimum section modulus for each pole class. 

Test-pole projected groundline strengths were multiplied by 

these factors in the derivation of the average species DFS 

values. Values used for each pole are provided in column 8 

of Tables C3, C4, and C5. 

The “variability” adjustment originally incorporated in the 

first ASA O5 standard and continued in 1963 to counter 

effects of variability introduced by sampling, handling, and 
28 

test methods was not used in this derivation. This was 

discussed in Rollins’ report but not included in his final 

tally of strength values. If we assume a coefficient of variation of 14%, the adjustment of a half standard deviation 

would be 0.93 and the values reported in Table C1 would be 

changedto 5,630 lb/in2

 for western redcedar, 7,130 lb/in2 for 

Douglas-fir, and 7,960 lb/in2

 for southern pine. MeasuredCOV values ranged from 15% to 19% for the test poles. 

This would give variability adjustments ranging from 0.90 

to 0.92. 

Table C1—Derivation of DFS for long poles 

Species 

MORgl

(lb/in2

) 

MORgl′

a

(lb/in2

) 

DFSb

 

(lb/in2

) 

West redcedar 5,030 5,530 6,040 

Douglas-fir 7,020 6,950 7,640 

Southern pine 7,800 7,720 8,530 

a

Adjusted for class oversize. 

b

Adjusted for conditioning and drying-in-service. 

Table C2—Minimum class groundline circumference and corresponding class oversize 

adjustment used to derive DFS for long poles

 

Groundline circumference (in.) for different species 

 and various pole lengths (ft) 

 Western redcedar Douglas-fir Southern pine 

Class i

Class load 

(lb) 50 55 60 70 50 65 75 50 65 75 

Oversize 

adjust 

7 0 1,200 33 34 35 37 31 34 36 30 33 34 1.13 

6 1 1,500 36 37 38 40 33 36 38 32 35 37 1.13 

5 2 1,900 39 40 41 44 36 39 41 35 38 40 1.13 

4 3 2,400 42 43 45 47 39 43 45 37 41 43 1.13 

3 4 3,000 45 47 48 51 42 46 48 40 44 47 1.12 

2 5 3,700 48 50 52 54 45 49 52 43 47 50 1.11 

1 6 4,500 51 53 55 58 48 52 55 46 51 53 1.10 

H1 7 5,400 55 57 58 62 51 56 59 49 54 57 1.09 

H2 8 6,400 58 60 62 65 54 59 62 52 57 60 1.09 

H3 9 7,500 61 63 65 69 57 62 65 55 60 63 1.08 

H4 10 8,700 64 66 68 72 59 65 69 57 63 66 1.07 

H5 11 10,000 67 70 72 76 62 68 72 60 66 70 1.07 

H6 12 11,400 70 73 75 79 65 71 75 63 69 73 1.07 
Table C3—Groundline data for southern pine

Groundline Groundline

Failure stress at circumLength MORfl location failure MORgl ference Oversize AMORgl

ID (ft) (lb/in2) (in.) (lb/in2) (lb/in2) (in.) adjust (lb/in2)

 P010 50 9,371 12 9,476 9,481 46.0 1.100 10,429

 P014 50 6,955 42 7,113 7,250 44.8 1.100 7,975

 S004 50 9,632 60 9,673 10,227 43.1 1.108 11,332

 S005 50 5,665 96 6,830 6,246 42.0 1.108 6,921

 S006 50 6,871 168 8,603 8,207 41.5 1.117 9,165

 S007 50 7,717 90 7,967 8,454 43.1 1.108 9,368

 S008 50 7,748 51 8,734 8,151 43.2 1.108 9,032

 S012 50 6,935 114 7,277 7,796 42.0 1.108 8,639

 S034 50 6,448 0 6,448 6,448 45.1 1.100 7,093

 S039 50 5,509 159 6,149 6,512 41.5 1.117 7,272

 S040 50 6,699 168 7,114 8,002 43.6 1.108 8,867

 S041 50 6,171 177 7,175 7,449 43.7 1.108 8,254

 S042 50 7,190 15 7,195 7,296 44.5 1.100 8,026

 S055 50 9,169 30 9,447 9,444 41.5 1.117 10,545

 S062 50 4,506 225 5,621 5,762 44.0 1.108 6,385

 S074 50 6,080 36 6,018 6,300 42.8 1.108 6,981

 S081 50 8,475 0 8,475 8,475 43.6 1.108 9,391

 S082 50 8,953 213 9,993 11,281 41.3 1.117 12,598

 S098 50 5,196 180 5,729 6,294 42.4 1.108 6,974

 S105 50 7,340 93 7,446 8,067 44.0 1.108 8,939

 S108 50 7,386 48 7,669 7,746 44.6 1.100 8,521

 S109 50 9,905 0 9,905 9,905 41.9 1.108 10,976

 S111 50 5,564 96 5,506 6,135 43.4 1.108 6,798

 S112 50 5,792 120 6,451 6,554 41.6 1.117 7,319

 S113 50 5,915 54 6,059 6,242 42.1 1.108 6,916

 S114 50 7,711 96 8,985 8,502 43.0 1.108 9,421

 S116 50 9,047 9 9,024 9,127 43.0 1.108 10,113

 S120 50 5,666 240 7,062 7,383 42.8 1.108 8,181

 S001 65 9,539 84 10,005 10,169 49.1 1.100 11,186

 S009 65 6,074 87 5,966 6,490 51.4 1.100 7,139

 S013 65 6,735 264 7,071 8,363 48.7 1.108 9,267

 S016 65 8,523 306 10,191 11,007 50.4 1.100 12,108

 S022 65 7,662 12 7,735 7,730 48.0 1.108 8,566

 S025 65 7,252 186 8,568 8,405 49.1 1.100 9,245

 S031 65 6,001 384 6,667 8,001 51.2 1.100 8,801

 S035 65 6,402 88 6,566 6,846 49.0 1.100 7,531

 S037 65 5,316 440 7,233 7,088 50.3 1.100 7,797

 S049 65 4,649 387 5,991 6,199 52.0 1.093 6,773

 S053 65 6,733 380 8,619 8,977 50.3 1.100 9,875

 S063 65 5,098 324 6,818 6,699 52.7 1.093 7,319

 S064 65 4,095 258 5,058 5,057 50.6 1.100 5,563

 S066 65 5,601 150 5,793 6,298 50.5 1.100 6,927

 S073 65 7,105 6 7,066 7,137 53.5 1,093 7,797

 S100 65 6,462 312 7,774 8,393 49.0 1.100 9,232

 S101 65 6,306 297 6,884 8,075 50.8 1.100 8,882

 S102 65 4,918 360 6,632 6,557 49.0 1.100 7,213

 S107 65 7,980 0 7,980 7,980 51.4 1.100 8,778

 S110 65 3,783 330 4,345 5,000 48.6 1.108 5,540

29
Table C3—Groundline data for southern pine (continued)

Groundline Groundline

Failure stress at circumLength MORfl location failure MORgl ference Oversize AMORgl

ID (ft) (lb/in2) (in.) (lb/in2) (lb/in2) (in.) adjust (lb/in2)

 S115 65 7,031 264 8,404 8,731 51.0 1.100 9,604

 S002 65 6,321 234 7,714 7,639 47.1 1.108 8,465

 S003 65 6,522 210 7,876 7,717 49.0 1.100 8,489

 S010 65 10,534 0 10,534 10,534 47.4 1.108 11,673

 S011 65 5,784 276 6,899 7,262 47.6 1.108 8,047

 S014 65 6,548 21 6,556 6,651 48.1 1.108 7,370

 S015 65 3,380 396 4,233 4,507 46.3 1.108 4,994

 S030 65 6,714 57 7,097 7,009 46.1 1.108 7,766

 S032 65 5,926 306 6,578 7,653 49.2 1.100 8,418

 S033 65 4,047 441 6,267 5,396 48.0 1.108 5,979

 S046 65 8,925 39 9,321 9,189 47.6 1.108 10,183

 S047 65 6,508 252 8,045 7,994 48.1 1.108 8,858

 S054 65 8,467 21 8,783 8,600 46.8 1.108 9,530

 S059 65 8,785 0 8,785 8,785 46.9 1.108 9,735

 S069 65 4,492 369 5,981 5,989 48.8 1.108 6,637

 S075 65 7,865 240 8,557 9,556 47.0 1.108 10,590

 S078 65 7,366 0 7,366 7,366 48.6 1.108 8,162

 S079 65 3,966 456 6,539 5,288 47.4 1.108 5,860

 S080 65 7,280 180 7,854 8,394 46.6 1.108 9,302 

 S083 65 6,747 342 7,722 8,996 48.2 1.108 9,969

 S088 65 4,612 360 5,730 6,149 46.4 1.108 6,814

 S092 65 7,755 30 7,751 7,930 48.8 1.108 8,788

 S094 65 5,337 375 6,683 7,116 46.0 1.108 7,885

 S095 65 5,157 396 7,030 6,876 48.4 1.108 7,619

 S106 65 8,013 54 8,410 8,345 45.4 1.117 9,319

 S117 65 6,498 102 7,288 7,027 46.1 1.108 7,786

 S118 65 7,955 198 8,675 9,315 45.8 1.108 10,322

 S119 65 6,151 300 7,137 7,898 48.6 1.108 8,752

 S018 65 6,483 264 8,108 8,050 49.2 1.100 8,855

 S019 65 5,398 396 8,731 7,197 45.5 1.117 8,037

 S021 65 8,391 154 9,499 9,466 43.9 1.117 10,570

 S023 65 10,338 0 10,338 10,338 42.8 1.117 11,544

 S027 65 8,371 78 8,412 8,882 43.9 1.117 9,918

 S028 65 7,189 45 7,618 7,436 46.4 1.108 8,240

 S036 65 4,960 210 3,588 5,869 48.5 1.108 6,503

 S038 65 7,776 168 8,980 8,876 42.4 1.125 9,985

 S045 65 5,777 189 7,062 6,713 43.2 1.117 7,496

 S051 65 8,012 42 8,224 8,268 43.2 1.117 9,233

 S052 65 7,353 228 8,280 8,839 43.3 1.117 9,870

 S058 65 6,812 312 7,913 8,848 45.6 1.117 9,880

 S060 65 4,319 309 4,965 5,594 44.9 1.117 6,246

 S070 65 5,812 261 6,514 7,197 44.0 1.117 8,037

 S071 65 7,427 210 7,709 8,788 43.3 1.117 9,813

 S077 65 6,296 99 6,376 6,792 42.8 1.117 7,584

 S085 65 8,208 39 8,173 8,451 46.0 1.108 9,365

 S091 65 5,022 360 6,478 6,696 43.7 1.117 7,477

 S097 65 7,894 72 7,555 8,337 45.1 1.117 9,309

 S099 65 7,705 171 7,994 8,817 43.4 1.117 9,845

30
Table C3—Groundline data for southern pine (continued)

Groundline Groundline

Failure stress at circumLength MORfl location failure MORgl ference Oversize AMORgl

ID (ft) (lb/in2) (in.) (lb/in2) (lb/in2) (in.) adjust (lb/in2)

 S103 65 5,207 312 7,097 6,763 42.8 1.117 7,552

 S017 75 6,616 132 6,989 7,222 52.7 1.100 7,945

 S020 75 4,947 444 6,523 6,596 50.8 1.108 7,309

 S024 75 5,447 236 6,285 6,409 48.5 1.108 7,102

 S026 75 7,166 258 8,712 8,573 47.0 1.117 9,573

 S029 75 8,023 36 8,172 8,211 49.1 1.108 9,099

 S043 75 6,238 504 8,698 8,317 49.3 1.108 9,217

 S044 75 8,419 189 9,450 9,570 49.7 1.108 10,604

 S048 75 7,989 204 9,410 9,180 48.8 1.108 10,173

 S050 75 7,390 75 7,903 7,760 47.8 1.117 8,666

 S056 75 5,475 405 7,231 7,300 49.3 1.108 8,089

 S057 75 6,768 105 7,016 7,252 55.7 1.093 7,924 

 S072 75 9,302 0 9,302 9,302 49.4 1.108 10,308

 S086 75 7,872 225 8,711 9,187 46.8 1.117 10,259

 S087 75 9,740 42 10,139 10,007 50.9 1.108 11,089

 S089 75 5,629 396 8,678 7,505 49.3 1.108 8,317

 S093 75 11,204 0 11,204 11,204 49.9 1.108 12,415

 S104 75 6,624 282 7,246 8,072 55.1 1.093 8,819

Average 6,826 7,579 7,798 1.10 8,618

Adjusted for drying in service (1.1) and conditioning (0.90) 0.99

Adjusted value 7,720 8,532

31
Table C4—Groundline data for Douglas-fir 

Groundline Groundline

Failure stress at circumLength MORfl location failure MORgl ference Oversize AMORgl

ID (ft) (lb/in2) (in.) (lb/in2) (lb/in2) (in.) adjust (lb/in2)

 D010 50 7,561 12 7,715 7,650 46.9 1.10 8,415

 D013 50 7,197 72 7,492 7,737 45.6 1.10 8,510

 D031 50 7,338 24 7,482 7,513 47.3 1.10 8,264

 D039 50 7,483 0 7,483 7,483 47.3 1.10 8,231

 D057 50 6,656 0 6,656 6,656 45.1 1.10 7,322

 D058 50 5,785 42 5,765 6,030 48.7 1.09 6,589

 D091 50 7,094 96 7,156 7,822 46.2 1.10 8,604

 D092 50 6,992 24 7,046 7,158 46.2 1.10 7,874

 D093 50 5,363 90 6,315 5,875 45.7 1.10 6,463

 D094 50 4,769 150 5,411 5,580 47.2 1.10 6,138

 D095 50 6,874 33 7,016 7,101 45.4 1.10 7,811

 D097 50 5,773 24 5,738 5,910 47.5 1.10 6,501

 D100 50 6,917 18 7,067 7,040 47.2 1.10 7,744

 D101 50 5,302 0 5,302 5,302 49.7 1.09 5,793

 D108 50 5,866 75 6,640 6,326 45.0 1.11 7,010

 D110 50 7,792 9 7,875 7,861 47.2 1.10 8,647

 D118 50 5,461 48 5,512 5,727 46.6 1.10 6,300 

 D120 50 6,370 66 6,750 6,805 47.0 1.10 7,486

 D128 50 5,741 36 5,828 5,949 45.8 1.10 6,543

 D137 50 5,727 0 5,727 5,727 46.2 1.10 6,300

 D148 50 7,613 54 7,783 8,033 45.4 1.10 8,837

 D152 50 7,801 0 7,801 7,801 45.6 1.10 8,581

 D156 50 8,594 24 8,846 8,799 46.3 1.10 9,678

 D164 50 7,446 30 7,625 7,669 45.2 1.10 8,436

 D168 50 8,768 0 8,768 8,768 44.2 1.11 9,716

 D009 50 6,692 18 6,698 6,811 47.6 1.10 7,492

 D012 50 5,326 6 5,380 5,357 50.2 1.09 5,853 

 D034 50 7,043 36 7,340 7,298 48.8 1.09 7,973

 D038 50 5,829 151 5,853 6,828 48.5 1.09 7,460

 D096 50 5,541 18 5,636 5,639 46.8 1.10 6,203

 D099 50 9,137 48 9,423 9,583 48.2 1.09 10,470

 D103 50 4,708 174 5,295 5,663 50.3 1.09 6,187

 D105 50 6,175 138 6,667 7,128 48.5 1.09 7,788

 D124 50 6,042 168 6,713 7,217 48.6 1.09 7,885

 D125 50 7,076 0 7,076 7,076 49.3 1.09 7,731

 D011 50 6,881 0 6,881 6,881 50.8 1.09 7,472

 D037 50 6,540 69 7,351 7,009 52.1 1.09 7,611

 D056 50 5,375 0 5,375 5,375 50.5 1.09 5,873

 D142 50 5,769 54 6,091 6,088 52.2 1.09 6,611

 D008 50 5,514 0 5,514 5,514 55.1 1.08 5,955

F043 50 5,792 85 6,630 6,312 53.0 1.09 6,854

F061 50 5,520 150 5,740 6,459 55.3 1.08 6,975

 F044 51 6,498 84 7,288 7,061 54.2 1.08 7,626

 F067 51 2,942 228 3,957 3,754 53.9 1.08 4,055

 D001 65 8,768 0 8,768 8,768 49.8 1.10 9,645

32
Table C4—Groundline data for Douglas-fir (continued)

Groundline Groundline

Failure stress at circumLength MORfl location failure MORgl ference Oversize AMORgl

ID (ft) (lb/in2) (in.) (lb/in2) (lb/in2) (in.) adjust (lb/in2)

D004 65 5,559 261 6,051 6,884 51.7 1.10 7,572

D015 65 6,473 60 6,656 6,773 51.5 1.10 7,450

D027 65 6,216 48 5,979 6,444 50.6 1.10 7,089

D028 65 7,679 12 7,646 7,748 50.2 1.10 8,522

D030 65 5,420 276 6,229 6,805 49.4 1.11 7,541

D040 65 5,682 30 6,001 5,811 49.3 1.11 6,439

D045 65 4,822 360 5,493 6,429 53.2 1.09 7,025

D051 65 6,828 90 7,122 7,313 50.8 1.10 8,045

D054 65 7,320 300 7,863 9,400 49.9 1.10 10,340

D059 65 6,559 54 6,617 6,831 50.4 1.10 7,514

D072 65 6,322 294 6,694 8,072 54.0 1.09 8,820

D073 65 6,570 270 6,904 8,203 50.6 1.10 9,024

D088 65 5,740 255 6,093 7,069 50.8 1.10 7,776

D098 65 6,123 78 6,618 6,497 49.4 1.11 7,199

D130 65 7,793 240 7,598 9,469 48.2 1.11 10,493

D141 65 4,734 276 5,008 5,944 50.2 1.10 6,538

D144 65 6,003 90 6,405 6,430 49.8 1.10 7,073

D150 65 7,045 204 7,795 8,293 50.6 1.10 9,122

D153 65 7,984 30 8,084 8,165 50.5 1.10 8,981

D163 65 7,416 156 8,078 8,380 49.3 1.11 9,286 

D021 65 5,750 84 6,316 6,130 46.6 1.11 6,792

D090 65 6,253 108 6,227 6,794 46.4 1.11 7,529

D127 65 6,943 0 6,943 6,943 46.0 1.12 7,753

D135 65 8,967 120 9,505 9,838 44.5 1.12 10,985

D146 65 6,094 54 5,977 6,347 46.4 1.11 7,033

D149 65 9,189 60 9,660 9,614 46.0 1.12 10,736

D155 65 7,350 252 9,114 9,028 45.1 1.12 10,081

D159 65 6,601 72 6,789 6,971 45.5 1.12 7,784

D166 65 7,199 192 8,508 8,386 44.9 1.12 9,365

D003 65 6,800 84 7,023 7,249 43.3 1.12 8,095

D005 65 7,858 180 7,825 9,061 46.2 1.12 10,118

D007 65 6,516 120 6,707 7,149 43.8 1.12 7,983

D019 65 5,747 78 5,859 6,098 45.5 1.12 6,809

D026 65 7,795 204 7,595 9,175 45.7 1.12 10,246

D033 65 5,964 138 6,373 6,640 44.4 1.12 7,414

D036 65 7,491 64 7,391 7,862 46.3 1.12 8,779

D042 65 8,354 4 8,359 8,379 45.7 1.12 9,356

D047 65 6,805 39 7,169 7,007 44.2 1.12 7,824

D055 65 7,424 30 7,476 7,592 44.6 1.12 8,478

D064 65 6,216 36 6,326 6,386 44.2 1.12 7,131

D066 65 6,223 102 6,639 6,729 46.4 1.11 7,457

D070 65 6,298 18 6,338 6,383 44.3 1.12 7,127

D075 65 5,789 0 5,789 5,789 45.0 1.12 6,464 

D083 65 6,540 78 6,356 6,939 43.7 1.12 7,749

D085 65 6,847 162 7,170 7,776 43.4 1.12 8,683

33
Table C4—Groundline data for Douglas-fir (continued)

Groundline Groundline

Failure stress at circumLength MORfl location failure MORgl ference Oversize AMORgl

ID (ft) (lb/in2) (in.) (lb/in2) (lb/in2) (in.) adjust (lb/in2)

D089 65 6,556 12 6,583 6,615 43.8 1.12 7,386

D111 65 8,428 36 8,336 8,658 44.4 1.12 9,668

D117 65 8,092 0 8,092 8,092 44.4 1.12 9,036

D122 65 7,309 30 7,577 7,474 45.8 1.12 8,346

D129 65 7,203 6 7,270 7,235 43.2 1.13 8,139

D133 65 7,695 120 7,528 8,442 44.4 1.12 9,427

D134 65 6,654 162 7,274 7,557 43.2 1.13 8,501

D002 65 7,444 0 7,444 7,444 52.4 1.10 8,188

D063 65 6,018 168 6,272 6,869 52.3 1.10 7,556

D065 65 5,357 75 5,391 5,671 52.2 1.10 6,238

D071 65 6,415 240 6,319 7,795 53.9 1.09 8,516

D107 65 6,438 216 6,920 7,658 52.6 1.09 8,367

D114 65 5,883 54 6,065 6,127 52.1 1.10 6,740

D121 65 5,935 294 6,890 7,578 52.8 1.09 8,280

D126 65 6,943 120 6,886 7,617 53.0 1.09 8,322

D154 65 8,093 6 8,105 8,129 51.0 1.10 8,942

D165 65 6,147 216 7,261 7,312 52.3 1.10 8,043

D006 65 7,696 0 7,696 7,696 57.8 1.09 8,357

D014 65 4,574 108 4,419 4,970 54.5 1.09 5,430

D016 65 5,259 36 5,337 5,402 56.2 1.09 5,867

D017 65 7,213 54 7,428 7,512 56.9 1.09 8,158

D041 65 6,976 63 7,026 7,316 56.6 1.09 7,945

D043 65 7,082 18 7,260 7,177 55.2 1.09 7,842

D049 65 5,056 384 5,412 6,741 58.9 1.08 7,281

D050 65 8,011 31 8,050 8,198 55.3 1.09 8,958

D052 65 5,747 300 6,538 7,380 62.0 1.07 7,931

D061 65 5,348 0 5,348 5,348 61.6 1.08 5,776

D062 65 3,538 450 3,789 4,717 66.4 1.07 5,048

D067 65 4,081 0 4,081 4,081 59.5 1.08 4,407

D068 65 6,633 12 6,698 6,692 55.4 1.09 7,312

D074 65 6,800 0 6,800 6,800 55.4 1.09 7,430

D076 65 5,665 156 6,336 6,401 55.7 1.09 6,952

D077 65 5,507 96 5,865 5,927 54.0 1.09 6,475

D078 65 5,197 240 6,269 6,315 55.6 1.09 6,899

D081 65 5,716 204 6,644 6,728 55.0 1.09 7,351

D104 65 6,938 48 6,918 7,193 55.9 1.09 7,811

D109 65 5,182 216 5,800 6,164 55.7 1.09 6,694

D112 65 8,550 3 8,567 8,569 55.2 1.09 9,362

D116 65 6,837 6 6,799 6,867 56.2 1.09 7,458

D131 65 4,217 408 5,794 5,623 56.8 1.09 6,106

D138 65 7,092 0 7,092 7,092 56.3 1.09 7,701

D139 65 7,223 18 7,278 7,320 58.2 1.09 7,949

D140 65 5,394 240 5,787 6,554 57.4 1.09 7,117

D147 65 6,962 0 6,962 6,962 56.4 1.09 7,560

D157 65 6,129 216 6,918 7,290 56.8 1.09 7,917

34
Table C4—Groundline data for Douglas-fir (continued)

Groundline Groundline

Failure stress at circumLength MORfl location failure MORgl ference Oversize AMORgl

ID (ft) (lb/in2) (in.) (lb/in2) (lb/in2) (in.) adjust (lb/in2)

D160 65 5,286 9 5,265 5,321 57.1 1.09 5,779

D167 65 6,545 24 6,543 6,663 57.0 1.09 7,236

D018 65 5,325 78 5,199 5,650 60.2 1.08 6,102

D044 65 7,160 12 7,252 7,224 59.3 1.08 7,802

D115 65 6,886 81 7,029 7,323 61.0 1.08 7,909

D020 75 5,708 258 6,218 6,829 54.4 1.10 7,512

D023 75 5,265 516 4,969 7,020 61.3 1.09 7,623

D024 75 5,871 42 5,783 6,032 52.6 1.10 6,635

D025 75 5,303 84 5,206 5,602 56.8 1.09 6,121

D029 75 5,977 132 6,298 6,525 51.8 1.11 7,230

D032 75 7,016 12 6,966 7,070 53.4 1.10 7,777

D035 75 5,392 192 5,977 6,142 52.9 1.10 6,756

D046 75 6,306 318 6,869 7,905 53.4 1.10 8,696

D048 75 7,756 57 7,506 8,048 52.8 1.10 8,853

D060 75 6,177 37 6,182 6,326 53.6 1.10 6,958

 D069 75 4,721 408 5,294 6,295 53.2 1.10 6,924

 D079 75 6,344 48 6,381 6,544 52.3 1.10 7,198

 D080 75 4,804 192 4,820 5,472 52.7 1.10 6,020

 D082 75 8,075 0 8,075 8,075 52.4 1.10 8,883

 D084 75 4,127 336 4,284 5,249 56.5 1.09 5,735

 D086 75 6,938 210 6,975 8,008 52.2 1.10 8,809

 D087 75 6,068 165 6,633 6,780 52.6 1.10 7,458

 D102 75 7,131 78 7,401 7,503 53.3 1.10 8,254

 D106 75 7,472 234 8,433 8,779 53.5 1.10 9,657

 D113 75 8,203 66 8,365 8,562 53.0 1.10 9,419

 D119 75 7,905 108 8,549 8,488 52.8 1.10 9,337

 D123 75 6,566 0 6,566 6,566 51.6 1.11 7,276

 D132 75 6,025 66 6,594 6,289 54.1 1.10 6,918

 D136 75 8,202 0 8,202 8,202 51.8 1.11 9,089

 D143 75 7,549 90 7,640 8,007 52.3 1.10 8,808

 D145 75 5,690 222 6,772 6,626 52.1 1.10 7,288

 D151 75 6,784 180 6,889 7,661 52.3 1.10 8,427

 D158 75 6,494 150 6,487 7,179 52.8 1.10 7,897

 D161 75 6,137 192 6,053 6,991 52.3 1.10 7,690

 D162 75 7,235 102 7,104 7,737 53.0 1.10 8,511

 D022 75 6,000 216 6,252 6,956 56.9 1.09 7,600

 D053 75 3,912 432 5,025 5,216 55.3 1.10 5,738

Excluding PSC pole tests 6,480 6,730 7,020 1.04 7,720

Adjusted for conditioning and drying in service 0.99 0.99

Adjusted value 6,950 7,643

35
Table C5. Groundline data for western redcedar 

Groundline Groundline

Failure stress at circumLength MORfl location failure MORgl ference Oversize AMORgl

ID (ft) (lb/in2) (in.) (lb/in2) (lb/in2) (in.) adjust (lb/in2)

 C91 55 3,410 294 2,980 4,547 54.1 1.09 4,968

 C92 55 4,140 270 4,310 5,425 51.1 1.10 5,967

 C93 55 5,060 72 4,960 5,401 53.4 1.10 5,941

 C94 55 5,160 228 4,830 6,450 53.7 1.09 7,047

 C95 55 2,830 174 2,640 3,340 54.2 1.09 3,649

 W001 50 4,049 150 4,122 4,738 51.0 1.10 5,211

 W002 50 3,295 180 3,081 3,991 52.1 1.09 4,361

 W003 50 4,363 54 4,409 4,604 49.4 1.10 5,064

 W005 50 4,256 174 4,321 5,119 50.2 1.10 5,631

 W009 50 3,939 48 3,692 4,131 54.6 1.09 4,514

 W035 50 4,253 150 4,365 4,976 55.7 1.09 5,404

 W036 50 4,231 132 4,149 4,852 53.3 1.09 5,301

 W037 50 4,229 276 3,850 5,639 58.9 1.08 6,090

 W040 50 4,205 120 4,672 4,758 55.4 1.09 5,167

 W046 50 3,935 138 3,213 4,542 59.5 1.08 4,906

 W064 50 3,079 246 3,688 4,043 52.0 1.09 4,417

 W065 50 3,985 42 4,009 4,154 52.2 1.09 4,539

 W066 50 5,521 102 5,536 6,127 53.3 1.09 6,694

 W067 50 3,038 78 2,780 3,286 56.5 1.09 3,569

 W069 50 3,585 60 3,699 3,806 52.0 1.09 4,159

 W082 50 4,673 202 4,909 5,807 50.9 1.10 6,388

 W083 50 4,257 198 4,247 5,268 50.2 1.10 5,794

 W084 50 5,445 0 5,445 5,445 50.5 1.10 5,990

 W094 50 3,873 246 3,312 5,085 53.8 1.09 5,556

 W096 50 6,187 178 6,185 7,473 50.2 1.10 8,220

 W004 60 4,099 108 4,003 4,487 50.6 1.11 4,972

 W008 60 6,126 148 6,076 6,948 49.6 1.11 7,699

 W011 60 5,205 210 4,910 6,258 51.6 1.11 6,935

 W013 60 4,793 204 5,122 5,730 51.4 1.11 6,349

 W015 60 4,627 192 4,517 5,468 59.4 1.09 5,938

 W016 60 3,576 192 3,372 4,226 58.6 1.09 4,589

 W017 60 4,246 228 4,308 5,195 50.4 1.11 5,757

 W018 60 4,867 160 4,771 5,581 55.3 1.09 6,097

 W019 60 2,538 166 2,179 2,926 69.6 1.07 3,131

 W020 60 4,403 240 3,819 5,451 59.0 1.09 5,920

 W021 60 4,613 96 4,384 4,997 60.7 1.09 5,427

 W022 60 4,384 192 3,876 5,181 61.2 1.09 5,626

 W023 60 4,401 198 3,901 5,231 55.8 1.09 5,715

 W024 60 3,719 174 3,363 4,322 67.9 1.07 4,624

 W025 60 2,189 180 2,004 2,558 66.7 1.07 2,749

 W026 60 5,210 196 4,889 6,178 52.4 1.10 6,796

 W027 60 4,718 168 4,106 5,452 58.9 1.09 5,920

 W028 60 5,216 198 5,001 6,200 58.9 1.09 6,732

 W031 60 3,689 270 3,256 4,707 62.5 1.08 5,084

 W034 60 4,942 108 4,784 5,410 62.6 1.08 5,843

36
Table C5. Groundline data for western redcedar (continued)

Groundline Groundline

Failure stress at circumLength MORfl location failure MORgl ference Oversize AMORgl

ID (ft) (lb/in2) (in.) (lb/in2) (lb/in2) (in.) adjust (lb/in2)

 W038 60 3,845 210 3,858 4,623 62.2 1.08 4,993

 W039 60 3,833 150 3,705 4,357 54.1 1.10 4,792

 W041 60 4,542 300 3,384 5,979 60.0 1.09 6,493

 W042 60 4,360 180 4,207 5,095 64.7 1.08 5,502

 W043 60 4,433 24 4,244 4,520 63.6 1.08 4,882

 W044 60 4,746 180 4,402 5,546 48.8 1.11 6,145

 W045 60 4,617 264 4,413 5,856 51.4 1.11 6,489

 W047 60 3,987 162 3,640 4,582 50.0 1.11 5,077

 W048 60 5,089 210 4,965 6,119 63.2 1.08 6,608

 W049 60 4,841 330 4,337 6,455 63.5 1.08 6,971

 W052 60 4,170 150 4,261 4,740 49.6 1.11 5,252

 W053 60 3,336 198 3,192 3,965 56.9 1.09 4,332

 W054 60 4,298 102 4,344 4,681 47.3 1.12 5,227

 W055 60 3,217 0 3,217 3,217 65.8 1.07 3,457

 W056 60 3,395 192 3,260 4,012 63.4 1.08 4,333

 W057 60 2,455 156 2,563 2,806 60.6 1.09 3,047

 W058 60 4,525 264 4,947 5,739 55.0 1.10 6,313

 W063 60 3,778 0 3,778 3,778 60.1 1.09 4,103

 W068 60 3,023 210 2,844 3,635 58.9 1.09 3,947

 W070 60 3,742 34 3,759 3,846 61.3 1.09 4,176

 W071 60 4,000 108 4,190 4,379 48.5 1.12 4,890

 W072 60 5,075 180 5,225 5,930 48.4 1.12 6,622

 W073 60 4,209 66 4,044 4,444 48.5 1.12 4,962

 W074 60 2,020 270 2,012 2,578 66.2 1.07 2,770

 W075 60 4,366 192 3,822 5,160 58.2 1.09 5,638

 W076 60 4,189 372 5,358 5,585 62.8 1.08 6,032

 W077 60 4,849 186 4,455 5,698 59.4 1.09 6,188

 W078 60 4,333 226 4,779 5,289 55.2 1.10 5,818

 W079 60 5,343 184 4,907 6,265 63.4 1.08 6,766

 W080 60 5,119 12 5,165 5,169 49.1 1.11 5,727

 W081 60 4,061 330 4,003 5,415 61.8 1.08 5,848

 W087 60 4,040 228 4,067 4,943 62.4 1.08 5,339

 W088 60 4,858 198 5,163 5,774 62.2 1.08 6,236

 W089 60 4,778 142 4,459 5,390 63.4 1.08 5,821

 W092 60 4,628 228 4,677 5,662 50.4 1.11 6,275

 W093 60 4,393 6 4,413 4,414 50.0 1.11 4,891

 W097 60 4,620 46 4,228 4,795 55.8 1.09 5,239

 W098 60 5,219 214 5,205 6,297 50.8 1.11 6,977

 W099 60 5,202 208 4,986 6,240 48.2 1.12 6,968

 W100 60 5,122 0 5,122 5,122 60.4 1.09 5,562

 W006 70 5,196 0 5,196 5,196 59.2 1.09 5,677

 W007 70 4,191 282 3,569 5,191 64.8 1.09 5,637

 W010 70 3,898 300 3,310 4,903 64.7 1.09 5,324

 W012 70 4,213 264 3,992 5,140 63.5 1.09 5,582

 W014 70 4,171 192 3,919 4,801 69.0 1.07 5,159

37
Table C5. Groundline data for western redcedar (continued)

Groundline Groundline

Failure stress at circumLength MORfl location failure MORgl ference Oversize AMORgl

ID (ft) (lb/in2) (in.) (lb/in2) (lb/in2) (in.) adjust (lb/in2)

W029 70 3,740 72 3,517 3,933 70.7 1.07 4,227

 W030 70 3,702 216 4,335 4,343 60.0 1.09 4,745

 W032 70 5,819 264 5,692 7,099 61.2 1.09 7,757

 W033 70 4,991 372 4,420 6,655 61.2 1.09 7,271

 W050 70 4,407 174 4,445 5,001 70.7 1.07 5,375

 W051 70 4,667 90 4,362 4,973 59.2 1.09 5,433

 W059 70 4,685 162 4,090 5,268 60.0 1.09 5,756

 W060 70 3,983 150 4,074 4,438 59.3 1.09 4,849

 W061 70 4,387 228 4,640 5,196 56.9 1.10 5,716

 W062 70 3,880 234 3,819 4,618 57.4 1.10 5,080

 W085 70 4,524 88 4,506 4,812 56.5 1.10 5,293

 W086 70 4,206 358 4,623 5,565 56.3 1.10 6,122

 W090 70 5,040 180 4,800 5,747 56.2 1.10 6,321

 W091 70 5,117 246 5,405 6,150 55.7 1.10 6,766

 W095 70 4,030 300 4,345 5,069 56.8 1.10 5,576

Average 4,310 4,200 5,026 1.19 5,494

COV 19% 19%

Adjusted for conditioning and drying in service 1.1 1.1

Adjusted initial value 5,528 1.22 6,043

38
31 

Appendix D—Class Oversize 

Adjustment 

The use of discrete class sizes to classify utility poles introduces inaccuracy in the estimate of groundline strength. The 

ANSI O5.1 standard was developed as a specification for 

poles on the premise that the tabulated minimum circumferences for each size class provide the section modulus required to carry the designated class load and poles falling 

between consecutive class minimums are oversize for their 

respective class load. Annex C values were derived to correct for this by evaluating pole strengths as a weighted average capacity by first assigning an ANSI size class to each 

test pole, then dividing the failure moments at groundline by 

the minimum section modulus for the respective pole class 

rather than using the measured section property: the larger 

the pole within a given class size, the larger the groundline 

modulus of rupture (MOR). When applied to the minimum 

class dimension, these values give the average pole capacity. 

 In 2002, it was proposed to change the way values are 

derived in the specification. For poles in lengths of 50 ft and 

longer, test poles were first classified. The groundline 

strength of each test pole was then increased by the mean-tominimum section modulus ratio (A) for its ANSI class. 

The required section modulus is directly proportional to the 

class load, and the class load is defined by the function 

 

2 P 1200 250i 50i i = + + (1) 

The mean-to-minimum-section modulus ratio (A) is equivalent to the ratio of loads as given by Equation (2). These 

“oversize” adjustment values are given in Table D1. 

 

i

i

i P

P

A +0.5 = (2) 

The average of the size-adjusted strength values was then 

used to recalculate the minimum section modulus required 

for each class. In some cases, this required reclassification as 

some poles moved to the next larger class. The process was 

repeated until no further change was required. The use of 

this process, along with a 10% increase for drying in service, 

resulted in average adjusted fiber strength values for poles 

50 ft and longer that were close to those used for poles less 

than 50 ft in length. As a result, the tabulated minimum 

circumference values remained unchanged, but they must 

now be recognized as providing reduced capability to carry 

the designated class load when compared to the minimum 

circumference values given for shorter poles. 

ANSI tables represent a catalog of pole sizes. Class loads 

merely provide a basis for the size classification. The fact 

that values are derived differently for pole lengths either side 

of 50 ft is something to be documented but has little impact 

on design as long as users realize that class load is not synonymous with capacity. 

Table D1—DFS adjustments for ANSI structural pole classes 

DFS adjustments for ANSI structural pole classes 

 0 1 2 3 4 5 6 7 8 9 10 11 12 

Factor 7 6 5 4 3 2 1 H1 H2 H3 H4 H5 H6 

Ai 1.12 1.13 1.13 1.12 1.12 1.11 1.10 1.09 1.09 1.08 1.07 1.07 1.07 

Bi 0.96 0.96 0.96 0.96 0.96 0.97 0.97 0.97 0.97 0.97 0.98 0.98 0.98 

39