NBSIR 76-1154

Consideration of Fracture

Mechanics Analysis and Defect

Dimension Measurement

Assessment for the TransAlaska Oil Pipeline Girth Welds

Volume I

October 18, 1976

This is a Final Report

NOTE: This document has been prepared for the use

of the Materials Transportation Bureau,

Department of Transportation. Responsibility

for its further use rests with that agency.

Prepared for

Materials Transportation Bureau

U.S. Department of Transportation

Washington, D. C. 20590

NBSIR 76-1154

CONSIDERATION OF FRACTURE

MECHANICS ANALYSIS AND DEFECT

DIMENSION MEASUREMENT

ASSESSMENT FOR THE TRANSALASKA OIL PIPELINE GIRTH WELDS

VOLUME I

October 18, 1976

This is a Final Report

Prepared for

Materials Transportation Bureau

U. S. Department of Transportation

Washington, D. C. 20590

TABLE OF CONTENTS

Page

Foreword 1

List of Symbols 2

Executive Summary 6

1. Introduction 9

2. Problem Definition and Scope 12

3. NBS Data on Material Properties 14

4. Weld Defect Measurement Assessment 79

5. Remarks on Pipeline Design and Stress Analysis 113

6. Assessment of Material Property Data 118

7. Fracture Analyses 129

8. Weld Defect Summary 166

9. Conclusions 174

FOREWORD

The National Bureau of Standards (NBS) received a request in June,

1976, from the Department of Transportation, Office of Pipeline Safety

Operations, for technical assistance in evaluating anticipated requests

for waivers involving girth welds in the trans-Alaska oil pipeline. To

respond to that request, NBS formed a Trans-Alaska Pipeline Project (TAP P)

involving a number of NBS personnel. Mr. Harold Berger was named Leader

of the TAPP and Dr. John H. Smith was appointed as Deputy. Technical task

groups within the project were headed by Dr. Richard P. Reed (Fracture

Mechanics Analysis), Dr. M. Linzer (Nondestructive Evaluation), Dr. Robert

C. Placious (Assessment of Radiographic Measurements), and Dr. Joan

Rosenblatt (Statistics and Mathematical Analyses). Legal assistance

was provided by Robert G. Hayes and Gordon Fields. Liaison with the

NBS Office of Programs was through Dr. 3ruce W. Morrissey.

Major report sections were prepared by M. B. Kasen, J. F. LeBrecque,

H. M. Ledbetter, H. I. McHenry, D. T. Read, R. P. Reed, R. E. Schramm,

L. L. Sparks, and R. L. Tobler from the NBS Boulder Laboratory and by

J. T. Fong, D. A. Garrett, C. G. Interrante, J. A. Lechner, R. C. Placious,

J. Rosenblatt, and J. H. Smith of the Washington Laboratory. Critical

reviews and analyses were provided by R. deWit, E. Passaglia, and R. Thomson.

Many others contributed to this project and report. A partial listing includes the following: F. Biancaniello, G. Birnbaum, C. Brady, 8. H. Colvin,

N. M. Crockett, C. E. Dick, J. J. Filliben, M. Ganoczy, A. J. Goldman,

D. T. Goldman, D. E. Harne, G. E. Hicho, N. Hsu, L. W. Ketron, H. H. Ku,

J. W. Motz, J. Moulder, M. Reeve, R. G. Rehm, T. P. Royston, A. W. Ruff,

N. Sanchez, R. Schofer, D. Schwab, G. C. Serig, D. Wilson and W. Wilson.

In addition, several consultants contributed ideas, reviews, and in

some cases, reports to this project. Consultants included Professor J. A.

Begley, Ohio State University; E. L. Criscuolo and D. Polansky, Naval

Surface Weapons Center; Professor R. E. Green, Johns Hopkins University;

Dr. R. Halmshaw, Royal Armament Research and Development Establishment;

Professor G. R. Irwin, University of Maryland; Professor R. C. McMaster,

Ohio State University; Professor P. C. Paris, Washington University; and

Dean R. D. Stout, Lehigh University.

Fact finding visits were made to several organizations including The

Welding Institute, Southwest Research Institute, Rockwell International

Rocky Flats Laboratory, and the Alyeska Pipeline Service Company. The

open exchange of information and cooperation are gratefully acknowledged.

We are pleased to acknowledge the assistance of all these people and

the many others who contributed to this NBS Trans-Alaska Pipeline Project.

Harold Berger

John H. Smith

01
LIST OF SYMBOLS

For the purpose of this report, the following notation is used.

Some extra symbols, used only in certain sections, are defined there;

A

A

a

a

5

Aa

*

a

da

dN

B

b

BSI

BWI

C

CVN

c

CIT

COD

D

°m

D

t

DOT

E

Area under load-displacement curve of test specimen.

(Section 3E).

Area of separated metal = Bb (Section 31 and Appendix 0).

Depth of defect into pipe wall.

Crack depth in test specimen.

Allowable defect parameter of Draft British Standard.

Crack extension in test specimen.

Effective crack depth in Begley-Landes-Uilson model.

Fatigue crack growth rate, which equals C(AK)

n according

to the Paris equation.

Test specimen thickness.

Length of uncracked ligament in test specimen * W-a.

British Standards Institute.

British Welding Institute.

Empirical constant in the Paris fatigue-crack-growth

equation.

Charpy V-notch

1/2

Cranfield Institute of Technology.

Crack opening displacement = <5.

Radiographic film density.

Depth of defect measured by RI technique.

True depth of defect.

Department of Transportation.

Young's modulus.

n 0
F c Ligament closing force in the Begley model.

E

q

Total absorped energy (dial energy).

E m Energy to Max Load (initiation energy).

E

p Propagation energy = Eg-E

m

.

EPRI Electric Power Research Institute.

G, G

d

Film gradient.

HAZ Heat affected zone.

J J-integral (the energy per unit of new separational area

that is released at the tip of a perfect crack in an elastic solid during an infinitesimal, virtual increment

of forward crack extension).

J T Measure of fracture toughness derived using the J-integral

c concept.

K Stress-intensity factor.

K* Fracture toughness derived from the equivalent-energy concept.

Kj

c Plane-strain fracture toughness.

Kj

c Static fracture toughness derived from J-integral concept.

K

Jd Dynamic fracture toughness derived from J-integral concept.

aK Stress-intensity range in fatigue.

L Distance over which residual stresses act in the Begley

model

.

i Length of defect parallel to pipe wall.

LEFM Linear elastic fracture mechanics.

M Folias correction factor for pipe wall curvature.

m Plastic constraint factor in conversion between J

T and

5 or J and 6.

1C

c

M

back Special function occurring in the Irwin surface-flaw

equation.

N Number of fatigue cycles, i.e., service lifetimes.

NBS National Bureau of Standards.

n Empirical exponent in the Paris fatigue crack growth

equation.

OPSO Office of Pipeline Safety Operations.
F

GY

M

max

P . min

R

R

r

RI

S

SWRI

T, t

t

^GY

t

m

W

W

M

AX

A

6

6

Applied load.

Arrest load for fast fracture

Fast-fracture (cleavage) load

General -yield load

Maximum load

Maximum load in fatigue.

Minimum load in fatigue.

Special function occurring in the Irwin surface-flaw

equation.

Stress ratio = cr . /a . (Section 7A). min max

Through-wall reduction (Section 4B and Appendix L).

Pipe radius.

Rockwel 1 Internati onal

.

Span in bend test.

Southwest Research Institute.

Thickness of penetrameter and shim.

Thickness of pipe wall.

Time to general yield

Time to maximum load

Test specimen width.

Absorbed energy associated with maximum load

Change in thickness of the sample traversed by the

x-ray beam.

Special function occurring in the Srawley-Gross equation

for stress intensity factor of test specimen.

Displacement due to residual stress in the Begley model.

Crack opening displacement = COD.

Critical COD.

Applied tensile strain.

04
£

y

U

\

V

a

a

Ao

°c

cmax

ami n

a

w

Yield strain = o /E.

Linear absorption coefficient (x-radiation)

Fong safety factor

Poisson's ratio.

Applied tensile stress.

Flow strength.

Stress range * c? v - a. . 5 max mi n Ligament closing stress in the Begley model

Maximum a in fatigue.

Minimum a in fatigue.

Ultimate tensile strength.

Residual stress.

Yield stress (0.2% offset).

Dynamic yield stress.

Static yield stress.

05
EXECUTIVE SUMMARY

A number of buried girth welds in the 48 inch diameter trans-Alaska oil

pipeline contain radiographically detected defects somewhat larger than

permitted by the Department of Transportation (DOT) Regulations, 49 CFR

Part 195. The Alyeska Pipeline Service Company (Alyeska) has requested

waivers for these welds from DOT. The waiver request is based upon a fracture mechanics analysis. The information presented by Alyeska in

support of the waiver request contains proposed allowable defect size

curves for the weld defect types involved and descriptions and measurements of the weld defects in terms of both length along the weld

ci rcumference and depth in the through-wall direction.

In anticipation of this waiver submission, DOT requested assistance

from the National Bureau of Standards (NBS) in evaluating the fracture

mechanics analysis and the nondestructive evaluation (NDE) methods

used to detect and determine dimensions of sped fic weld defects. This

report is the response to that request.,

r

”

f Material properties needed for the fracture mechanics analysis

were determined from samples of field welds. NBS measured a number

of properties, for example, elastic, tensile, fracture toughness, and

fatigue, to confirm that our data agreed with more comprehensive test

program results reported in the open literature or reports from the (British)

Welding Institute. In addition, NBS determined the dynamic fracture

toughness to assess the effect of high rates of loading on the toughness

and allowable defect sizes. The effect of aggressive environments on

cracking of the pipe and welds was also assessed.

For ttois defect size and orientation case, there is no accepted theoretical

formulation or empirical data base to determine a specific allowable

defect size curve. Therefore, it was necessary to use safety factors

and other judgment considerations in modeling the allowable defect

size curves. This may lead to a more conservative design approach

than would otherwise prove necessary. Future work to provide experimental

verification of the allowable defect size curves would permit performancerelated regulations to be written. These regulations might provide the

same level of safety in a more cost-effective manner.

Taking into account the material properties, the stresses expected

in the pipeline and the several fracture mechanics models considered

in this report, NBS concludes that the Alyeska allowable defect size

curve for nonplanar defects is reasonable. For the planar defects

and arc burns, the Alyeska curves are reasonable except for small defect

lengths where the Alyeska curves allow unlimited defect depth; the other

fracture mechanics analyses do not allow unlimited defect depth.

06
The waiver request also reported weld defect sizes, both in

terms of length and depth, as determined from existing field radiographs

of the girth welds. The regulation now calls for defect length measurement from the radiographs. However, there can be differences in

reported defect lengths by different interpreters . Comparing results

from the several organizations who measured length in this case, there

were a number of major discrepancies.

Defect depth measurements are not normally made from radiographs

and, of course, were not originally intended from these field radiographs.

The depth measurement can be made in cases where good control of the

radiographic procedure, particularly in terms of x-ray energy and film

processing, can be maintained. With this control, the relationship

between film density and the thickness of metal can be predicted.

Taking into account the expected variations encountered in field

situations, it might have been anticipated that the field radiographs

would show large variations from the predicted contrast. This large,

observed variation in contrast makes any defect depth measurement which

depends on film density observations in the weld and defect area subject

to error.

Three methods were used to measure defect depths from the Alyeska

radiographs. Southwest Research Institute (SWRI) used both a densitometer

method and a visual comparison technique. Rockwell International (RI)

employed a visual comparison method involving only the weld and defect

area. In both SWRI methods, comparisons were made to the film contrast

for the shim-penetrameter image on the pipe wall.

Possible sources of error for the SWRI densitometer method include

uncertain values of the shim-penetrameter thickness and the inaccuracy of

measuring radiographic film densities in the areas of the weld (which does

not have a constant thickness) and the defect (which may not normally

have a predictable, uniform shape). The film density measurements used

to determine defect depth from the field radiographs were made with a densitometer with measuring aperture of 1 mm. This does not permit the

accurate measurement of defect sizes less than about 1.25 mm in width

or diameter, a limitation that could have been surmounted by the use of

a more sophisticated densitometer. For the depth measurements made by

the SWRI densitometer method, estimates of possible errors have been made.

For the depth measurements made by the SWRI. method of visual comparison

of contrast in the defect area and the penetrans ter-shim area, no assessment

of accuracy was possible.

07
The visual depth measurements made by RI have been analyzed. Based

on laboratory x-ray calibration data, the RI method tends to underestimate

the defect depth in increasing amounts for defects of greater depth.

However, these corrections are based on laboratory x-ray films. For

the Alyeska films, there are many field radiographs whose contrast falls

outside the contrast limits predicted by the film response and expected

radiographic procedure. This poses a dilemma since the reason for this

change in film contrast is not known. If the contrast variations are

due to extreme changes in x-ray voltage or changes in film processing,

an accurate depth determination by the RI method may not be possible.

However, if the contrast changes were caused by shim thickness variations

or densitometer reading errors, the accuracy of the RI method may not

be affected.

In support of the RI measurements there is information presented to

demonstrate that the RI depth measurements tend to be larger (more conserva

tive) than those made by the SWRI densitometer method. This is important

because these two sets of measurements were taken by independent approaches

and because the SWRI measurements should not be influenced by film contrast

variations. It is also demonstrated that RI measurements of several

natural defects in welds furnished to NBS by Alyeska, showed good

correlation with laboratory radiographic measurements. Since the RI method

is the only one used to measure the depth of all weld defects still under

consideration for waiver, more comparisons may be desirable in order to set

limits on this method.

The effective depth of an arc burn cannot be determined directly by

depth measurements from the radiographs. However, a reliable correlation

between arc burn diameter or width and the effective depth (which includes

the heat-affected zone) has been established. Therefore, the effective

depth of the arc burn can be determined from the diameter or width measured

radiographical ly

.

In summary, the fracture mechanics analyses provide a method to

evaluate the performance of the pipeline girth welds in terms of allowable

defect sizes. Defect lengths can be measured accurately, although some

discrepancies have been noted. Arc burn diameters or widths can be

measured and a method for converting that measurement to an effective

depth is described. For other defect types, the depth measurement becomes

more difficult and is subject to possible error as described in this report

08
1

.

INTRODUCTION

The trans- Alaska oil pipeline, which stretches 800 miles from Prudhoe Bay

to the terminal at Valdez, is being constructed by the Alyeska Pipeline Service

Company (Alyeska). Field construction of the pipeline began in 1975 following

several years of negotiation between the oil companies involved and the United

States Government. The stipulations for the agreement [G-l , exhibit D] 1 between

the oil companies and the Department of Interior reached in 1974 provide that

the design, material and construction, operation, maintenance and termination

practices employed in the pipeline system would be in accordance with safe and

proven engineering practice, and would meet or exceed several designated standards, including those set out in the Department of Transportation (DOT) Regulations, 49 CFR Part 195, "Transportation of Liquids by Pipeline'1 . 2 It was

further stipulated that all mainline girth welds would be inspected by radiography, a requirement beyond that normally called for in the DOT regulations for

1 iquid pipelines.

Soon after field welding began, some doubts were raised as reported by

Mr. E. L. Patton, Chairman and Chief Executive Officer of the Alyeska Pipeline

Service Company, about the quality of the radiographic film interpretation.

3

As

construction of the pipeline progressed, a number of pipeline girth welds were

buried or made relatively inaccessible. Reinterpretation of the radiographs of

some of the buried or inaccessible welds showed weld defect sizes beyond those

permitted by 49 CFR, Part 195.

On August 30, 1976, Alyeska requested that DOT waive the requirements of

49 CFR Part 195 for some of these girth welds. Documentation supporting Alyeska'

request for a waiver consisted mainly of a four-part study entitled: "Fracture

Mechanics Study of Buried Field Girth Welds. The study tracks the DOT

published suggested waiver application. 5 The Alyeska waiver request depends

upon fracture mechanics analysis. The analysis leads to a proposed acceptable

defect size upon which a decision to accept or reject a pipe weld defect can

be based. The analysis requires information which includes the defect type,

and dimensions, and the pipeline mechanical properties, operating stresses

and projected thermo-mechanical history.

1. See reference G-l listed in Appendix G (U.S. Government documents).

2. See reference G-2.

3. See reference G-3 for public statement by E. L. Patton dated July 20, 1976.

4. See reference H-l listed in Appendix H (Alyeska documents).

5. See Vol . 41, Federal Register, pp. 34375-77, (August 13, 1976) [G-4].
In anticipation of the Alyeska waiver request, the Office of Pipeline

Safety Operations (OPSO) of DOT approached the National Bureau of Standards

(NBS) in June 1976 to request technical assistance in evaluating the

fracture mechanics analysis and the nondestructive evaluation (NDE) methods

used to detect and determine dimensions of weld defects. The fracture

mechanics aspects of the NBS developed program were established shortly

after NBS was contacted by OPSO. It was agreed by the Materials Transportation

Bureau and NBS 5 that NBS would evaluate the fracture mechanics analysis

and independently. assess the material property data.

The NDE aspect of the NBS program has been more flexible because of

the changing nature of the anticipated Alyeska waiver request. As

initially presented by OPSO, the problem was one of evaluating an ultrasonic system which would be used to inspect welds from inside the pipe.

The ultrasonic system, under development as a result of a contract

between Alyeska and Holosonics, Inc., presented the inspection information

in an image format so that direct comparisons could be made with radiographic

inspection. The system was expected to be used to inspect welds for

which radiographic information was unavailable or in question.

The Holosonics ultrasonic system was demonstrated to the President's

Fact Finding Team on the Trans-Alaska Pipeline System Problem7 on July 13,

1976, i n Fairbanks, Alaska. Prob lems with the system were noted, however,

and no request for its use has been received i.v OPSO. The NBS role in

regard to the ultrasonic inspection included technical preparation for

the evaluation of the system and the preparation of a set of guidelines

by which any proposed alternate NDE system could be judged. Those

guidelines are included as part of this report.

After it had been established by OPSO that the existing field

radiographic inspection records would be used to provide the defect

dimensions necessary for the fracture mechanics analysis, NBS agreed to

review the methods used to measure these dimensions. An assessment of that

methodology is also included in this report.

The NBS fracture mechanics analysis developed alternative procedures to

the one used by Alyeska to determine allowable defect size. Bearing in mind

that fracture mechanics as applied to this problem is an inexact science, these

analyses will assist OPSO to determine if the allowable defect size curves

proposed by Alyeska are reasonable. There are several methods by which the

fracture mechanics analysis can be made, and therefore, this report includes

several such analyses.

6. See NBS Statement of Work dated September 8, 1976 [G-5]

.

7. See report by J. W. Barnum dated July 17, 1976 [G-6].

10
The assessment of the radiographic methodology determines limits,

where possible, on the accuracy of the defect size measurement. The

defect sizes, when coupled with the fracture mechanics analyses, pennit the

assessment of the likelihood that a given weld defect will result in a failure.

At least three precautionary observations should be made at the

outset of this report. One, there is no presently accepted procedure in

the Uriited States for fracture mechanics analysis of ductile materials. AOdinona 1 large scale testing and elastic-plastic analysis research is

required to produce a standardized procedure.

A second precautionary observation is that the radiographic measurement

of the size of a weld defect in the through-wall direction (hereinafter

referred to as the depth dimension) is not a widely accepted practice.

Since there is no accepted industry standard for this measurement, one

that is critical to the fracture mechanics analysis, the variables that

can influence the measurement are described and assessed so that decisions can

be made concerning the defect size.

Finally, it is worth noting that when all the technical information on

stress, defect size, and material properties is collected and subjected to

fracture mechanics analyses to arrive at an assessment of the Alyeska

waiver request, it becomes necessary and prudent to adopt several so-called

safety factors in engineering- type calculations to account for variables in

all phases of measurements ranging from laboratory data to field and

fabrication quality control. These safety factors discussed here are

independent of the analysis requirements published by DOT. A more detailed

discussion of the implications of safety factors in fracture analysis is

included in this report.

11
2. PROBLEM DEFINITION AND NBS SCOPE OF WORK

The Alyeska waiver request consists of several large volumes of supporting

documents. These documents contain information on the material properties of

the base metal, heat-affected zone, and weld material. Material property

data, including fracture toughness, fatigue-crack-growth rate, stress

corrosion threshold and tensile properties are used by Alyeska to perform

a fracture mechanics analysis and determine allowable defect sizes. A

knowledge of the stress spectrum the pipeline will experience over its

proposed 30-year lifetime is also required since this is necessary to

assure that stress corrosion and fatigue will not cause defects to grow to the

point that a through-wall crack occurs. The other major part of the waiver

request concerns the specific welds for which a waiver from the requirements

of 49 CFR part 195 is being made. The weld data submitted by Alyeska

included a description of the defect as to location and type, and dimensional

information concerning the defect length and the depth, or amount of steel

missing in the through-wall direction. The waiver documents are listed,

along with other Alyeska material in Appendix H of this report.

The two major aspects of the waiver justification are: (1) the

fracture mechanics analysis which allegedly provides a demarcation between

defect sizes that are acceptable (will not cause failure) and those that

are not, and (2) Alyeska's measurement of actual weld defect dimensions

which, after factors of safety are applied, are compared to the fracture mechanics

analysis.

The actual number of welds involved in the waiver request changes as the Alyeska Pipeline Service company continues to repair and/or replace

suspect welds in the fields. NBS has been informed that this remedial work

is continuing as long as weather conditions in Alaska permit. The types of

weld defects for which a waiver is requested include, incomplete penetration,

incomplete fusion, elongated slag inclusion, gas pocket, spherical porosity

and cluster porosity, wormhole porosity, hollow bead, crater cracks and arc

burns. These weld defect types are described generally in the Alyeska

Welding and Radiographic Support Manual included in Appendix H of this

report. The lengths of weld defects involved in the waiver request vary

appreciably between the various types of defects; these are generally a few

inches or less. However, a few defect lengths in the Sept. 10, 1976 submission

to DOT (see Section 8 ) are in excess of 7 inches. Weld defect depth

dimensions submitted by Alyeska (before NBS analysis of the accuracy of

these figures and before a new set of depth figures were obtained by

Rockwell International personnel for DOT) are qenerally small (less than

0.040 in.) but a few are in the range greater than 0.060 in. in depth. The

nominal wall thicknesses of the pipe are 0.462 and 0.562 in.

The NBS tasks were to evaluate the Alyeska fracture mechanics analysis

and methods of defect measurement. To accomplish that it was necessary to

conduct several studies, all of which are described in this report. In

Section 3, material property data generated by NBS are given. These tests

were necessary in order to verify the material property data used in the

fracture mechanics analysis.

12
The nondestructive evaluation aspects of the problem are discussed in

Section 4. Although NBS did not actually measure the sizes of the weld

defects, the methods used to obtain these dimensions are reviewed and

analyzed so that the accuracy of the measurements may be assessed. Where

possible this assessment is given. However, there remain several parts of

the measurement methods for which sufficient information for an assessment

was not available. These are defined.

In Section 5, the stress spectrum for the pipeline is examined; as

indicated above, the stress spectrum is needed to determine the growth of

defects during the pipeline's design life. Although the stress spectrum

has been established, some of the uncertainties in these calculations are

discussed.

The assessment and uncertainty estimates given in Section 6 put the data on

material properties in perspective and permit the fracture mechanics analyses to

be made using defined data, constants, and conditions.

The fracture mechanics analyses are presented in Section 7. Several

analyses are given, representing several different models presently under

consideration by the scientific community. For the range of defect sizes

under consideration in this case, there is no experimental data base nor

a three-dimensional elastic-plastic analysis. The several analyses given

are compared to those proposed by Alyeska for the various types of defects.

This Section also includes a discussion of the various safety factors involved

in the calculations. For each class of defects considered in the waiver

request, several theoretical curves are presented for proposed allowable

weld defect sizes.

A summary of the weld defect size information is given in Section 8. This

section presents defect sizes as aiven by Alyeska, SWRI , and RI. It also includes

RI arc burn measurements^ after NBS adjustment for depth. Important information

concerning the loc ation o f wel ds' and t he steal material is discussed in this

section.

The conclusions are presented briefly in Section 9. The appendices or, tain

many supporting documents, references, consultant rocrts and explanatory

technical material.

In a project of this magnitude and time scale, it is not surprising to find

that some aspects of the information are not as complete as one would expect to

realize in a longer duration program. Whenever this is true, it has been pointed

out and the implications of the missing information described. In spite of

this, it is our belief that this report meets our primary objective, namely to

provide DOT with the technical data and analyses to permit DOT to make their

decisions concerning the Alyeska waiver requests.

The units of measurement are consistent within each section of the report.

For example, material property da^ta are given primarily in metric units

whereas" defect dimensions are given in Englfsh~units. Multiple units are

given in a few parts of this report.

NBS recognizes and has attempted to respond to the urgent need for this

study as_pressed for by DOT. NBS has responded by forming a Trans-Alaska Pipeline

Project Task Force which includes employees throughout the Bureau. The work

described in this report was accomplished over the three month period beginning

on approximately July 1, 1976.

13
3 NBS DATA ON MATERIAL PROPERTIES

3A . WELDMENT SELECTION

3B . SPECIMEN GEOMETRIES AND LOCATIONS

3C . ELASTIC PROPERTIES

3D. TENSILE PROPERTIES

3E . FRACTURE TOUGHNESS OF HEAT- AFFECTED ZONE

3F . FATIGUE-CRACK GROWTH RATES

3G . SIMULATED-SERVICE FATIGUE PROPERTIES

3H . SIGNIFICANCE OF ARC BURNS

31. CHARPY V-NOTCH IMPACT SPECIMEN TESTING

3J. STRESS-CORROSION CRACKING TESTS

14
3A. WELDMENT SELECTION

The following criteria were used to select girth welds from the

trans-Alaska pipeline for study by NBS.

(1) Welds must be 1975 production- line type.

(2) Welds must be identified positively.

(3) All welding procedures that enter anticipated variance requests

must be represented by two complete girth welds.

Criteria 1 and 2 were met by obtaining the original radiographic

films of each weld. Dates on the films confirmed 1975 production. The

radiographic "fingerprint,” that is, locating radiographic landmarks such

as the distances of the longitudinal seam welds from the top button

position, confirmed the weld number. This permitted NBS to detect

immediately that one of the original radiographs furnished by Alyeska

(for NBS weld 2) was incorrect. This clerical error was promptly

corrected. Alignment-sheet numbers on the radiographs permit precise

location of the weld in the pipeline.

Welding procedures 102A-1 and 104A-1 were expected to enter into

waiver requests, according to Alyeska personnel (Mr. Joseph Willing,

Alyeska Pipeline Service Company). Two welds of each procedure were,

therefore, required. These procedures differ mainly in the type of

shielded-arc electrode. The 102A-1 procedure uses AWS E8010-G (Phoenix

Cel 80) electrodes, while the 104A-1 procedure uses AWS E7010-G (Lincoln

Shield Arc 65+) electrodes. Alyeska records were relied on for weldprocedure identifications. These were checked by chemical analyses of

the electrodes and the welds. Both wet-chemistry and emission-spectroscopy

results were obtained by Dr. R. W. Burke of the NBS Analytical Chemistry

Division. These analyses for nickel content are included in Table 1.

They show that Alyeska mislabeled the NBS #4 weldment as 102A instead of

104A. This error made it impossible to meet criterion 3 since time was

not available to obtain and analyze a second 102A weld.

Weldments received at NBS are listed in Table 1. Weldment 88070 of

procedure 302A-1 will not enter into waiver considerations. However, it

had been already cut from the pipeline and was included to provide

additional base metal for study. To facilitate specimen preparations,

approximately 20 cm (8 in) of base metal was retained on both sides of

each weld. All welds were quartered and labeled for shipping as shown

in Fig. 1. A photograph of an as-received weldment is shown in Fig. 2.

Welds 88070 and 88119 were selected by M. B. Kasen of NBS, Boulder,

during a visit to the pipeline July 15-24, 1976. Arrangements for

selecting and delivering the remaining weldments were made then. Quick

delivery was sought consistent with the minimum disruption of the Alyeska

remedial-weld program.

15
Table 1. Trans-Alaska-pipel ine girth welds received at NBS for study.

Weldment No.

NBS Alyeska

Weld

Procedure

Wt. Pet.

Nickel b

Alignment

Sheet

Date

Radiographed

Date Received

at NBS

1 88070 302A1

0.15 122 14 Oct. 1975 26 July 1976

2 88119 104A-1 0.67 123 14 Oct. 1975 26 July 1976

3 79907 102A-1 0.15 123 17 Aug. 1975 5 Aug. 1976

4

a 75061 104A 0.90 135 9 May 1975 5 Aug. 1976

5 88503 104A-1 0.79 123 17 Oct. 1975 19 Aug. 1976

a. Alyeska identified this weldment as a 102A type, but chemical analysis shows

it is actually a 104A type,

b. See Appendix F for details of chemical analysis.

Top Button at 12 O'CLOCK

a O'CLOCK

Fig. 1. Schematic of method for quartering and labeling the transAlaska pipeline girth weldments for shipping to NBS,

Boulder. Each quarter is to be stenciled with the weld

number and with markings 1-2, 2-3, etc.

18
1

Fig. 2. Quartered Alyeska weldment 8S503 received at MBS, Boulder.

II
3B. SPECIMEN GEOMETRIES AND LOCATIONS

Five cylindrical weldments of the Alaskan pipeline were supplied

to NBS^ Boulder by Alyeska for testing and evaluation. The weldments

are shown schematically in Fig. 3. All five cylinders were cut into

quarters approximately the same size.

Ten types of test specimens were prepared from the weldments:

1. Compact tensile (CT) for J-integral tests

2. Tensile (weld) (T)

3. Tensile (base metal) (T)

4. Surface-flaw fatigue (SF)

5. Arc-burn fatigue (AB)

6. Bend fatigue ( BF ) for crack-growth tests, weld metal

7. Bend threshold (BT) for stress corrosion tests, weld metal

8. Bend (B) for crack-growth tests, base-metal

9. Elastic (E)

10.

Charpy impact (Cl)

where the abbreviations in parentheses were used to mark the specimens.

Diagrams of these specimens are shown in Fig. 4.

Detailed specimen-location maps for all quarters are shown in Figs.

5-9. Shaded areas on each quarter map indicate heat-affected zones

from which no specimens were taken. All specimens are located relative

to the distance from top-dead-center (TDC) for the particular weld. Topdead-center indicates the point at which the welding begins and proceeds

in opposite directions around the pipe. The distance of each edge from

TDC is indicated in the upper corners of each map.

Each specimen is identified by a hyphenated code number. The first

digit is the NBS, Boulder number substituted for the Alyeska weld number;

NBS numbers 1 through 5 are Alyeska numbers 88070 (302A-1), 88119 (104A-1),

79907 (102A-1), 75061 (104A), and 88503 (104A-1), respectively. The number following the hyphen is the specimen number from the weldment; specimens

from each weldment are numbered consecutively in the order they were prepared. Letters following this number indicate the type of specimen. For

example, 2-26SF indicates the specimen was the twenty-sixth taken from NBS

number 2 (Alyeska 88119, 104A-1) weldment, and the specimen is a surfaceflaw type. A summary of the origins of all test specimens prepared at

NBS, Boulder is given in Table 2.

18
9 o'clock

quarter

4X

TDC

i

12 o'clock

quarter

" &

'

76 cm

Longitudinal Weld

3X' ——2/

4X

6 o'clock

quarter

3 o'clock

quarter

>r'- Girth Weld 88070.

Procedure 302A-J

(NBS 1)

9 o'clock

quarter

TDC

Fig. 3. Schematic diagrams of cylindrical weldments of trans-Alaska pipeline

containing girth welds. All welds are shown as shaded areas. Topdead-center, where the girth welds begin, is marked TDC. Specimen

locations within each quarter are shown in Figs. 5-9. See Fig. 2

for a photograph of one of the weldments.

13
9 o'clock

quarter

TOC

9 o'clock

quarter

9 o'clock

quarter

20
960

9

TENSILE (weld metal) (T)

All dimensions in cm.

BEND FATIGUE

(weld)

SURFACE-FLAW FATIGUE (SF)

21
All dimensions in cm.

Fig. 4. Diagrams of test specimens prepared from trans-Alaska pipeline.

Shaded areas indicate weld zone. All dimensions are in centimeters (1 cm = 0.39 in).

22
12

O’CLOCK

OUARTEH 30

CLOCK

QUARTER

334cm

hum

IOC

48cm

Irom

TOC

48cm

ham

IOC

147cm

hum

TOC

s III su

<T>Ii! §I

v\

535£

M1 1 UM <3.

ij £ ^ ' r vv 3

OJ

xs- u aj 3 T- W

•o

x

o

o

<u 3 • • z a • <j ~o i— -o O E (— 4-> r—

s_

o OJ •r—

a

CL i- 3 U 4-> . 2

4- •<—

X

c

*o

X <— OJ <-

X

i— on 4-> 0.4-1 4-> O) a) E on i_ 3 C r— •<- c on •<- O CD on *r“- • on » fsj O *-“N <u S- ID OJ <— -O +J o ai II X

=?te OJ <13 4-c 4-> +> O •T— E

CO

U T3 <13

U 4-

ea a> "O O) r—

o

ir nc -t-> Q. r— ^

4- r— 4-»

X on

<H •r- <JJ •*>c

O 1 on

E on -t—

!-» +-> ca

O O) sc

O <13 OJ J- a;c

CO <v L. OJ <13

M •<— C3 X fO s_

<0 on

cj on E a; +j at •a 3 'i- X e s- a; on o -LJ a <o T3 -a n- C OJ

<13 <— -<-> OJ

X

"O on

X OJ <13

U 4~> l— <13 CO 3 3

OJ CL) on

c - 3 i- ,— »i

•<- L- <a • <13 on OJ

<13

c

C OJ QJ 4L> -lc -a OJ <- -a S- C

on <u lsl *o <13 OJ

<y

x <13 3 C

O

>1

u 4-> 4J OJ on i »— +j •<- E 3 "O

<£ <o OJ on •<—

O <13

X s_

c

o •<— QJ 4- 1 OJ o a; on T3 © on 3 <— CL C 1 on on OJ Q.

on

o on 4-

E

O

a, sc

O 4- •<— 4J

<13 C_3 0)

O -o

E

E on on C on r— <U

L • u

O on < -LJ OJ <— OJ •<- c <13 <13 •M l a. +j •<- <_> s- < on <13 C 4-J -r-

<13 CM

U <13 <13 *0 3 O o O OJ X C

cr co

c -i 2: 4-> -rLO

cr

<_O ^J
12

O'CLOCK

QUARTER 3

O'CLOCK

QUARTER

333cm

Rom

TOC

47cm

from

TOC

47cm

from

TOC

140cm

from

TDC 5

V*

Fig.

6.

Quarter

maps

of

Alyeska

weldment

88119

(NBS

02),

weld

procedure

1Q4A-1.

See

caption

of

Fig.

5.
12

O'CLOCK

QUARTER

3

0

CLOCK

QUARTER

340cm

liom

TOC

41cm

tcom

TOC

41cm

liom

TOC

I4bcw

liom

IOC _ M' 1

1

‘ u-k.-

T

"*•**/..

<

5

1

eu2

5

©

**u

o

c_j ©

09

5

Uo

£

E

u

un

Quarter

maps

of

Alyeska

weldment

79907

(NBS

#3),

weld

procedure

102A-1.

See

caption

of

Fig.

5.
12

OCIOCK

(jllAHIHI 3U

CIOCK

UUAH1IH

J•

tin

hum I

DC

bhiu

hum

IOC

SJr.m

hum

IOC

"I 5a 5a i ! : I i I I I 50^ £<55 £o> cn

ao

if!

ga^'» =o 2 - - s n J < ^ a ^ ^ ^ : n

'v.w- 1 y i 'V'?

£2

uS

cc

cn Lu

26

Quarter

maps

of

Alyeska

weldment

75061

(MBS

44),

weld

procedure

104A.

See

caption

of

Fig.

5.
<T) C*5

cn

CD

27

Quarter

maps

of

Alyeska

weldment

88503

(NBS

#5),

weld

procedure

104A-1.

See

caption

of

Fig.

5.
Table 2. Specimen-weldment matrix showing origins of test specimens.

SPECIMEN IDENTIFICATIONS

WELD IDENTIFICATIONS

Compact

Tmsi/a

Tensila ime/dl

Tons/la ibasal

Surface- Flaw

fatigue

Arc Burn

Fatigue

Band

Fatigue

Band

Thrashold

Band

<basal

Elastic

1

Charpy

Impact NBS Alyasha Prvcadura

1 88070 302A-1 1-1T 1-2T 1-3T

1-4T

1-5T

1-6T 1-33T 1-34T 1-35T

1-26AB

1-27AB

1-28AB

1-29A8

1-7B 1-88 1-9B MOB

1-11B 1-128

1-13B

1-148

1-15B 1-16B

1-17B 1-188 1-19B 1-20B

1-21B 1-22B 1-368

1-378 1-38B

1-30E

1-31E 1-32E

j

2 88119 104A-1 2-1CT

2-5CT

2-9CT

2-13CT

2-17CT 2-21CT

2-2T 2-6T

2-10T

2-14T

2-18T

2-22T

2-25SF

2-26SF

2-27SF

2-28SF

2-29SF

2-30SF

2-38F

2-7BF

2-11BF 2-15BF

2-19BF

2-23BF

2-31BF 2-328F

2-33BF

2-34BF

2-35BF

2-36BF

2-4BT

2-8BT

2-12BT

2-16BT

2-20BT

2-24BT

A-1

through

A-50

3 79907 102A-1 3-1CT

3-5CT

3-6CT

3-10CT

3-11CT 3-15CT

3-16CT

3-20CT

3-21CT

3-25CT

3-26CT

3-30CT

3-2T

3-7T

3-12T

3-17T 3-22T

3-27T

3-3BF

3-8BF

3-13BF

3-18BF

3-23BF

3-28BF

3-4BT

3-9BT

3-14BT

3-19BT

3-24BT

3-29BT

B-1 through

B-29

C-1 through

C-26

0-1 through

0-35

4 75061 104A 4-3SF

4-4SF

4-5SF

4-6SF

4-9SF

4-30SF

4-1BF 4-7BF

4-11 BF 4-13BF

4-26BF

4-28BF

4-2BT

4-8BT

4-10 BT

4-12BT

4-27BT

4-29BT

4-148

4-15B

4-16B

4-17B 4-188

4-198

4-20B

4-21B 4-22B

4-238

4-248

4-25B

5 88503 104A-1 5-5CT

5-7CT

5-10 CT

5-12CT

5-17CT

5-19 CT

5 20CT

5-22CT

5-25 CT 5-27CT

5-6T

5-11T 5-16T 5-18T

5-21T 5-26T

5-1SF

5 2SF

5-3SF

5-14SF 5-15SF

5-24SF

5-48F

5-8BF

5-9BF

5-13BF

5 -23 BF 5-28BF

5-33 BT 5-34BT

5-35 BT 5-36BT

5-37BT

5-388T

5-29B

5-3QB

5-31B

5-32B

•

28
3C. ELASTIC PROPERTIES

Elastic constants of samples of the trans-Alaska pipeline were determined between -40°C (-40°F) and 60°C (140°F) using dynamic methods. Roomtemperature measurements were made with a 10-MHz pulse method,

1

and temperature effects on the elastic constants were determined with a pulsesuperposition method. 2 The elastic constants reported here are: Young's

modulus, the shear modulus, the bulk modulus (reciprocal compressibility),

and Poisson's ratio.

Room-temperature elastic constants of the pipeline base metal (specimen 1-30E) and the pipeline weld metal (specimen 1-32E), measured along the

weld direction, are shown in Table 3 together with the elastic constants of

unalloyed iron. 3 For practical purposes, the elastic constants of the base

metal and the weld metal are identical; and these materials are slightly

softer elastically than unalloyed iron.

Changes of the elastic constants with temperature are shown in Fig. 10.

As expected for an iron-base material in this temperature region, 3 all the

elastic stiffnesses decrease linearly with increasing temperature, and

Poisson's ratio increases linearly. The elastic constants of the pipeline base metal are given in Table 4 for selected temperatures.

The measured quantities =* longitudinal wave velocity and vt = transverse wave velocity are believed to have systematic errors between

0 and +1% and 0 and +2%, respectively. This estimate is based on previous studies in our laboratory using the same experimental conditions

and a specimen similar to the pipeline steel. The t 2 a imprecisions of

v£ and vt are ± 1.4% and ± 0.66%, respectively. The total uncertainty in

the mass density is estimated to be t 0.1%, and it was neglected. Thus,

from standard formulas, the total (systematic plus imprecision) uncertainty in E is +5%, -3% and in v is +3%, -1%. The point-to-point imprecision in the temperature dependence of the elastic constants is a few

parts in 104 .

While the results are not reported here, the anisotropy of the base

material was studied by measuring nine ultrasonic sound velocities in

three mutually perpendicular directions. A definite anisotropy was

detected; for example, the shear modulus varied about ten percent with

direction. This anisotropy is probably related to the mill rolling of the

steel. The elastic constants reported in the tables were determined by

averaging the anisotropic elastic constants over all directions.

1. E. R. Naimon, W. F. Weston, and H. M. Ledbetter, Cryogenics 14,

246-249 (1974).

2. W. F. Weston, E. R. Naimon, and H. M. Ledbetter, in: Properties of

Materials for Liquefied Natural Gas Tankage , ASTM STP 579 (Amer. Soc.

Test. Mater., Philadelphia, 1975).

3. H. M. Ledbetter and R. P. Reed, J. Phys. Chem. Ref. Data 2_, 531-618

(1974).

29
Table 3. Room-temperature elastic constants of the trans-Alaska

pipeline base metal compared to those of unalloyed iron.

Base Weld-Zone

Iron Material Material

(302A)

Young's modulus, lO

1 ^ N/m^ (10^ psi) 2.130(30.9) 2.050(29.8) 2.060(29.9)

Shear modulus, 10^ N/m^ (10

6

psi) 0.828(12.0) 0.794(11.5) 0.797(11.6)

Bulk modulus, 10^ N/m^ (10^ psi) 1.680(24.4) 1.640(23.6) 1.660(24.1)

Poisson's ratio, dimensionless 0.289 0.291 0.294

Table 4. Elastic constants of the trans-Alaska pipeline base metal at selected

temperatures; primary units are 10'* N/ms secondary units in parentheses are 106 psi, Poisson's ratio is dimensionless.

Temp.

°C(°F)

Young's

Modulus

Shear

Modulus

Bulk

Modulus

Poisson's

Ratio

-40 (-40) 2.09 (30.3) 0.810 (11.7) 1.65 (24.0) 0.290

-20 ( -4) 2.08 (30.1) 0.805 (11.7) 1.65 (23.9) 0.290

0 ( 32) 2.07 (30.0) 0.800 (11.6) 1.64 (23.8) 0.291

20 ( 68) 2.05 (29.8) 0.794 (11.5) 1.64 (23.8) 0.29T

40 (104) 2.04 (29.6) 0.789 (11.4) 1.63 (23.7) 0.292

60 (140) 2.03 (29.4) 0.784 (11.4) 1.63 (23.6) 0.292

30
ELASTIC

CONSTANTS

RELATIVE

TO

22°C=72°F

TEMPERATURE, °F

-40 -20 0 20 40 60 80 100 120 140 1 1 1 1 1 1

Fig. 10. Temperature variations of the elastic constants of the trans-Alaska

pipeline steel. Data points, which were determined at 5°C intervals,

are omitted for clarity. Room-temperature (22°C = 72°F) values of

the elastic constants are given in Table 3.

31

3D. TENSILE PROPERTIES

Several tensile properties - yield strength,

ultimate strength, elongation, and reduction in area -- were determined for the trans-Alaska pipeline base metal and welds (102A-1 and

104A-1) with test procedures used previously in this laboratory

J

Flat base-metal specimens oriented parallel to the pipe axis

were 1.1 cm (0.43 in) thick with a reduced section 2.5 cm (1.0 in)

long and 1.0 cm (0.43 in) wide. Weld-metal specimens from the two

welding procedures had their axes along the pipe girth; their reduced

section length was 3.2 cm (1.25 in) and their diameter was 0.4 cm

(0.14 in).

The important experimental details are as follows: Crosshead

speed on the test machine was constant at 0.1 cm/min (0.04 in/min).

Three test environments were used: (a) air at 60°C (140°F) achieved

by a surrounding electrical heater; (o) air at 20°C (68°F); and (c)

the ullage space above liquid nitrogen at -20°C (-4°F). Temperatures

were stable within ± 0.5°C (0.9°F) except near the finish of the tests

when mechanical deformation caused temperature increases up to 2°C

(4°F) in the necked region. Two data recordings were made simultaneously: load-cell output versus cl ip-on-extensomater output, and

load-cell output versus elapsed time. Elongation was determined by

gage-mark separation.

Inaccuracies in these measurements are estimated to be less than

+3 percent for load and about ± 5 percent for strains, according to the

equipment manufacturers. Relative "imprecisions are estimated to be about ± 5

percent. Results of this study _are

^ summarized ^ln Table 5 and in Figs. 11-13.

Differences in the yielding behavior of the base and weld metals

are shown in the stress-strain curves in Fig. 11. The strain-hardening

behavior of the base metal is typical of mild steels. 2 Small oscillations often occurred in the plastic regions of the weld-metal stressstrain curves. After reaching ultimate load, all specimens necked down

locally. Upon fracturing, the base metal usually showed a "delamination,

or split, parallel to the original pipe surface; this traveled in both

directions from the fracture surface as far as 0.6 cm (0.25 in). Two

weld specimens from procedure 102A-1 showed a void on the fracture surface.

Examples of fractured tensile specimens are shown in Figs. 14-15.

1. K. A. Warren and R. P. Reed, Tensile and impact properties of

selected materials from 20 to 300°K, NBS Monogr. 63, U.S. Gov't.

Print. Office (1963).

2. NBS, Cryogenics Div., data reported in Cryogenic Materials Data

Handbook, PB 171809-5 (1961).

32
Table 5. Tensile properties of trans-Alaska pipeline materials; primary units

are 10® N/m2; secondary units in parentheses are 103 psi; elongation

and reduction in area are expressed as percentages.

Specimen Temp.

°C(°F)

Yield Strength

(0.2% offset)

111 timate

Strength

Elongation at

Fracture, %a Reduction

in Area, %

1 -2T -20 ( -4)

Base Metal

6.1(88)

1x7W

6.1(89) 44 68

1-5T 20( -4) 5.2(76) 5.8(84) 37 70

1-34T -20 ( -4) 5.4(79) 6.1(89) 41 70

1 -IT 20( 68) 5.0(73) 5.7(83) 37 71

1-4T 20 ( 68) 5.1(74) 5.7(82) 35 69

1-35T 20 ( 68) 5.5(80) 5.9(85) 41 70

1-3T 60(140) 4.9(71) 5.0(72) 40 73

1 -6T 60(140) 5.0(73) 5.5(80) 34 69

1-33T 60(140) 5.1(74) 5.7(82) 37 75

3-7Tb

-20 ( -4)

Weld Metal, Procedure 102A-1

4.7(68) 5.7(83) 15 55

3-22T -20 ( -4) 4.6(66) 5.9(85) 19 69

3-2T 20 ( 68) 4.1(59) 5.2(76) 24 69

3-12T 20( 68) 4.1(60) 5.4(78) 23 68

3-1 7T 60(140) 3.9(57) 5.2(75 20 69

3-27Tb 60(140) 4.8(69) 5.0(72) 14 78

2-2T -20< -4)

Weld Metal, Procedure 104A-1

4.6(67) 6.0(87) 23 67

2-6T -20( -4) 4.4(64) 5.4(79) 21 71

5-6T -2Q( -4) 4.7 68) 5.7(83) 21 66

5-1 IT -20( -4) 5.0(72) 5.7(83) 24 68

2-1 0T 20( 68) 4.8(70) 5.2(76) 23 73

2-1 4T 20 ( 68) 3.8(55) 5.2(75) 22 69

5-21

T

20( 68) 4.7(68) 5.5(80) 20 67

5-26T 20( 68) 4.1(60) 5.4(78) 21 68

2-18T 60(140) 4.6(67) 5.3(77) 17 70

2-22T 60(140) ! 4.3(62) 5.1(74) 15 68

5-16T 60(140) 3.5(51) 4.7(68) 17 65

5-18T 60(140) 4.7(68) 5.5(79) 12 42

a. In 2.54 cm (1 in) gage length.

b. Void visible on fracture surface.

33
STRESS

(10

8

N/m

2

)

STRESS

(10

8

N/m

2)

STRAIN (%)

Typical stress-strain curves for the trans-Alaska pipeline

base metal and weld metals from procedures 104-A and 104A-1.

Gage lengths was 2.54 cm = 1 in.

1 1, u 4

Fig. 11.

STRESS

(

1

0

3

psi

)

STRESS

(10

3

psi)
PERCENT

STRENGTH

(10

8

N/m

2)

-20

TEMPERATURE (°F)

20 60 100 140

TEMPERATURE (°C)

Fig. 12. Temperature variation of yield strength (0.2* offset) and

ultimate strength of trans-Alaska pipeline base metal and

weld metals. Error bars indicate data spread. Symbols

have same meanings as in Fig. 13.

Fig. 13. Temperature variation of the elongation and reduction in

area of trans-Alaska pipeline base metal and weld metals

35

STRENGTH

(10

3

psi)
I—I—I—I—I

I—I—I

I—I

I—I—I—I

0 5 10 cm

Fig. 14. Examples of fractured weld-metal tensile specimens

that were used to determine yield strength, ultimate strength, elongation, and reduction in area.

0 5 10 15 cm

Fig. 15. Examples of fractured base-metal tensile specimens.

36

3E . FRACTURE TOUGHNESS OF HEAT-AFFECTED ZONE

Tests were conducted to determine the fracture toughness of the

heat-affected zones (HAZ) in trans-Alaska pipeline steel weldments.

J xc is reported as an index of fracture toughness. Ji is defined as

the rate of change of potential energy with respect to crack area, and

it is proportional to the energy required to initiate fracture in a flawed specimen subjected to monotonically increasing loads. Using

experimental procedures described previously^, J-resistance curves were

determined for specimens of the 104A-1 and 102A-1 weld processes at

20°C and -20°C. The tests at 20°C were conducted in room air, while

tests at -20°C were conducted with the specimens immersed in an alcohol

bath that was cooled with dry ice. Compact specimens, 1.01 cm thick,

were machined with starting notches located 0.38 cm from the weld centerline so that subsequent fatigue cracks would sample the HAZ. (See

Fig. 4.) All specimens were then subjected to sinusoidal fatigue load

cycling at 15 Hz. Using minimum and maximum fatigue loads of 0.6 and

6 kN, precracks were initiated and propagated at 20°C until the overall

crack-1 ength-to-specimen-width ratio, a/W, was about 0.6. The fatigue

loads were small, less than half the loads required in subsequent J-tests.

At least five specimens were tested for each weld-process/temperature

combination. The J-test procedure first suggested by Landes and Begley,

and successfully employed in NBS studies^ 0 f ferritic steels and their

welds, was used. This procedure involves loading each specimen to a predetermined displacement such that a range of subcritical crack extensions is obtained. The test machine is operated in ram-displacement

control. Upon reaching the predetermined displacement, the specimen is

unloaded and heat treated to mark the crack extension. The specimen is

then fractured into halves and crack extension is measured at three points

equidistant across the specimen thickness; the average crack extension is

defined as Aa. For each test the J value was calculated using J = 2A/Bb,

where A = area under load-displacement curve, B = specimen thickness, and

b = length of uncracked ligament. Test specimens are shown in Fig. 16,

and a macrophotograph of the fracture surfaces is shown in Fig. 17.

The J-versus-crack extension curves are shown in Fig. 18. As recommended by Landes and Begley, the J/2aflow ^ ine was constructed to account

for apparent crack extensions due to crack tip blunting that occurs prior

to actual material separation. The flow stress of the material is defined

as the average of the yield and ultimate stresses:

- MPa LlS'O ±

37
The critical J values are taken at the points of intersection of the

J/2jfiow line wlth the J'Aa curves. The critical J values, denoted

are listed in Table 6. The uncertainties in Jig values are estimated at ± 10% and are due to (1) the uncertainty in fitting curves to

the data and (2) the fact that the J/2cTflow ^ ne is an approximation

for crack-tip blunting.

The results show that Ji c for the 104A-1 weld process HAZ at 20°C

and -20°C is 80 kJm-2; jj c f0 r the 102A-1 weld process HAZ at 20°C is

100 kJm-2, and the temperature reduction to -20°C apparently increases

the toughness to 110 kJm-2.

Direct Ki c measurements using the ASTM E 399-74 method were not

possible since rapid, unstable crack extension in this material at 20

and -20°C did not occur. However, Kj c values can be estimated from Jj c data using the approximation:

K,

c

- (EJ

Ic ) 1/2

(2)

4

where E is Young's modulus. Ledbetter's elastic modulus measurements

were used to obtain the Ki c values listed in Table 6. These values serve

as estimates of Kj

c

, with uncertainties estimated at + 20 percent. In

view of data presented by Read and Reed^ for an aluminum alloy, it appears

that Kj

c estimates based on Eq. (2) may be conservative.

Although the critical crack-tip opening displacement (COD) at the

initiation of crack extension was not measured, it can be calculated

using the approximation:

(3)

Here, m, the plastic constraint factor, has a value between 1 and 2.1,

dependent on materials. For deeply notched 1 X_I bend specimens of transAlaska pipeline steel, Harrison^ indicated that m has a value near 1.6.

This value was used to obtain the critical COD estimates listed in Table 6.

The uncertainty in COD is ± 30%, due primarily to uncertainty in m, which

appears to be specimen-geometry dependent.

5

1. R. L. Tobler, R. P. Mikesell, R. L. Durcholz, and R. P. Reed, in

Properties of Materials for Liquefied Natural Gas Tankage , STP 579

(Amer. Soc. Test. Mater., Phi 1 a . , 1975), pp. 261 -287.

2. H. I. McHenry and R. P. Reed, submitted for publication.

3. D. T. Read and R. P. Reed, Intern. J. Fracture, forthcoming.

4. H. M. Ledbetter, Elastic properties section of this report.

5. J. D. Harrison, in Fracture Mechanics Study of Buried Field Girth

Welds, Part 3 (Alyeska Pipeline Service Company, August 1976).

38
Table 5. J-integral fracture-toughness results for trans-Alaska pipeline HAZ specimens.

Specimen a/W

Aa mm( in

)

J

kJ/mZ (1n*lb/in Z

)

J Ic kJ/m^ (in-lb/in^)

COD

3

mm ( i n

)

K,

b

(ks1-in) ,/Z

Procedure 104A-1

,

T = 20°C ( 68° F

)

2-5CT 0.619 0.0254(0.001) 49.5(283)

2-9CT 0.594 0.0508(0.002) 70.2(401)

2-1 3CT 0.536 0.2032(0.008) 120.7(690)

2-17CT 0.596 0.3048(0.012) 160.4(917) 80(457) 0.114(0.0045) 128(117)

2-21CT 0.577 1.1938(0.047) 285.2(1630)

5-2CCT 0.626 0.4064(0.016) 181.6(1038)

5-22CT 0.613 0.2286(0.009) 128.3(733)

S-25CT 0.626 0.5538(0.022) 200.9(1148)

Procedure 104A-1 , T = -20°C (-4°F)

5-5CT 0.597 0.1016(0.004) 89.6(512)

5-7CT 0.610 0.2540(0.010) 133.7(764)

5-1 OCT 0.616 0.5334(0.021 ) 189.2(1081

)

80(457) 0.107(0.0042) 128(117)

5-1 9CT 0-619 0.0889(0.0035) 66.5(380)

5-27CT 0.622 0.1270(0.005) 99.4(568)

Procedure 102A-1

,

T » 20°C (68°F)

3-5CT 0.608 0.4826(0.019) 182.3(1042)

3-oCT 0.625 0.6350(0.025) 264.2(1510)

3-1 OCT 0.615 0.1270(0.005) 115.5(660) 100(571) 0.152(0.0060) 143(131) . 3-15CT 0.512 0.0381(0.0015) 69.3(396)

3-16CT 0.595 0.1016(0.004) 103.3(593)

Procedure 102A-1 , T * -20°C ( -4°F)

3-1 CT 0.620 0.0508(0.002) 80.0(457)

3-HCT 0.609 0.1778(0.007) 142.1(312)

3-2CCT 0.610 0.1016(0.004) 113.2(647)

3-21CT 0.6C5 0.0381(0.0015) 49.9(285) 110(623) 0.143(0.0053) 151(133)

3-25CT 0.606 0.0254(0.001

)

63.2(361

)

3-25CT 0.610 0.0508(0.002) 39.6(512)

3-30CT 0.513 0.5342(0.023) 231 .9(1325)

a. Calculated from Eq. (3).

b. Calculated from Eq. (2).

3&A

Fig. 16. One-centimeter-thick compact specimens used for J-integral tests.

Fig. 17. Fracture surfaces showing: (a) fatigue pre-crack

zone, (b) the heat-tinted crack-extension zone

due to J test at 20°C, and (c) final fracture zone

obtained after cooling to low temperatures to avoid

subsequent deformation of (b).

33
CRACK EXTENSION. Aa (in)

O O.OI 0.02 0 03 0.Q4 0.05

CRACK EXTENSION. Aa (mm)

Fig. 18. J-integral results for trans-Alaska pipeline HAZ specimens.

J

INTEGRAL

(in

lb/in 2)
3F. FATIGUE-CRACK GROWTH RATES

Due to a variety of factors, including periodic shut-downs,

the trans-Alaska pipeline will experience stress fluctuations

during its projected thirty-year service lifetime. The property

governing subcritical flaw growth under cyclic loading conditions

is the fatigue-crack growth rate, da/dN, where a is the crack length

and N is the number of cycles. This report presents fatigue-crack

growth rates as a function of the stress-intensity-factor range,

AK, for the pipeline steel and its welds in air and in a corrosive

envi ronment.

The four-point bend specimens. Figs. 4 and 19, chosen for these

tests were approximately 1 . 1 cm wide (W), 1.2 cm thick (B), and 23

cm in span length (S). After etching each specimen, notches were

machined in the base metal, weld metal, or heat-affected zones

(HAZ). The notch orientation was such that the cracks were propagated in the thickness direction of the pipeline sections. All

base-metal specimens were from weld NBS #4. Weld-metal specimens

were from welds NBS #2 and NBS #4, and HAZ specimens were from welds

NBS #2 and #3. Differences due to welding procedure were not distinguishable from the scatter among replicate testsOne series of tests was conducted on the base metal, weld

metal, and HAZ in room air at 20°C ( 68 ° F ) and on the base metal in

air at 60± 1°C (14P±_2°F)- __The cycl ing frequency was 10 Hz.__ A second

test series was conducted on base-metal, weld-metal / and HAZ spec i men s in

a corrosive environment at. both specimen temperatures of and" 60°C~ _ The corrosive environment consisted of Groenveld and Elsea's^ electrolyte: a 0.004 wt. pet. H2SO4 solution, to which 5 drops per liter

of a 5g phosporous/40 ml carbon-disulfide "poison" was added. The

frequency was 0.1 Hz, except for one base-metal test at 0.01 Hz.

A constant-current power supply was used to produce a current density of about 0.1 mA/cm^ at the crack-growth region. This procedure charges the crack-tip material with hydrogen and creates a simulated stress-corrosion environment that tends to enhance fatiguecrack growth.

The specimens were first pre-cracked in room temperature air,

using sinusoidally varying loads. The minimum/maximum load ratio

(Pm in/ pmax) was constant at 0.05. Pmax ranged between 1.3 and 3.1

kN, but was constant for each test. The da/dN values were calculated

1. G. P. Groenveld and A. R. Elsea, in Hydrogen in Metals , I. M.

Bernstein and A. W. Thompson, Eds. (Amer. Soc. Metals, Metals

Park, Ohio, 1974).

41
by the central -difference method and crack lengths were monitored by compliance measurements.

2 Using methods described by McHenry and Reed , the functional

relationship between relative crack length (a/W) and specimen compliance, EBC (E = Young's modulus, B = thickness, C = deflection

per unit load), was determined using calibration specimens. Calibration correlations are shown in Fig. 20, and a base-metal calibration specimen is shown in Fig. 21. Using a computer program,

the crack-length/compliance data were fitted to a four-term powerseries equation. Crack lengths were calculated from compliance

measurements taken during da/dN tests. The stress-intensity factor ranges were calculated using the equation

S < P

m*x

- V(a/W)/BW2

(1)

3

Here, according to Srawley and Gross ,

Y ( a/W ) = 1. 992-2. 468fc/W)+12.97(a/W)

2

-23.17(a/W) 3 +24.8(a/W)

4

(2)

The uncertainty in the da/dN data is estimated at less than ± 40%.

The da/dN-versus-AK data from this study are plotted in Figs.

22, 23, and 24. Also shown in these figures are bands representing

the spread of published data from other sources4 >5 for similar

steels and welds. As plotted on log-log scales, the base metal

data exhibit a linear trend, having a slope of approximately three.

Greater scatter and higher apparent slopes are exhibited by weld

and HAZ data, possibly due to effects of residual stress.

Analysis of the findings leads to the following conclusions:

(1) At 20°C (68°F), fatigue-crack growth rates are higher

for base-metai specimens than for weld or HAZ specimens. A temperature increase from 20°C to 60°C (140°F)

has no measurable effect on rates for the base metal.

(2) The pipeline-steel base-metal behavior is typical of ferritic steels in general, as indicated by comparison with

the worls of Barsom4 , Maddox5, and Vosikovsky.6

2. H. I. McHenry and R. P. Reed, unpublished report.

3. J. E. Srawley and B. Gross, Eng. Fract. Mechs., 4, 587-589 (1972).

4. J. M. Barsom, J. Eng. Ind., Trans. ASME, 1190-1196 (Nov. 1971).

5. S. J. Maddox, Welding Journal, 401s-409s (Sept. 1974).

6. 0. Vosikovsky, Closed Loop, 3-12 (April 1976).

42
(3) Over the range of AK investigated, the weld and HAZ data

do not exceed the upper bound of Maddox 's^ data for

carbon-manganese steel weldments.

(4) Vosikovsky 6 reported that a hydrogen-charging environment

that simulates stress-corrosion conditions significantly

reduces fatigue-crack growth resistance of base metal

specimens at 20°C. The limited data available for the

trans-Alaska pipeline at 0.1 mA/cm2 imply that the

(unreported) current densities of Vosikovsky were significantly higher.

Fig. 19. Four-point bend specimens before and after testing.

43
RELATIVE

CRACK

LENGTH,

a/W

o I 1 1 1 1 1 1 100 200 300 400 500 600 700

COMPLIANCE, ESC (dimensionless)

Fig. 20. Crack-length/compliance correlations used in crack-growth-rate tests

i .mi

Fig. 21. Base-metal crack-length calibration specimen used in determining

the correlation shown in Fig. 20.

44
FATIGUE

CRACK

GROWTH

RATE.

da/dN

(mm/cycle)

(ksi -in^2

) 5 10 50

STRESS-INTENSITY RANGE.AK (MPa-m'/2

)

Fig. 22. Fatigue-crack growth rates for the base metal.

45
FATIGUE

CRACK

GROWTH

RATE,

da/dN

(mm/cycle)

(ksi • in

1/2

) 5 10 50

STRESS-INTENSITY RANGE,AK (MPa-n/2

)

Fig. 23. Fatigue-crack growth rates for the weld metal.

46

{in/cycle)
FATIGUE

CRACK

GROWTH

RATE.

da/dN

(mm/cycle)

(ksi-in'/2

) 5 10 50

03o

e

Fig. 24. Fatigue-crack growth rates for the HAZ.

47
/
3G. SIMULATED-SERVICE FATIGUE PROPERTIES

A series of fatigue tests was conducted on samples of transAlaska pipeline weldments to determine the fatigue behavior of natural

and artificial flaws under simulated pipeline operating conditions.

The tests were conducted at the Martin Marietta Corporation, Denver

Division.

Test specimens were machined as shown in Fig. 4 from welds 2,

4, and 5. The natural curvature of the pipe was retained throughout

the test section; however, the specimen ends were flattened in a hydraulic press to facilitate gripping in the test machine. Artificial flaws

were introduced into nine specimens and natural flaws were present in

the other two specimens. The artifical flaws were introduced into the

weld by electrical -discharge machining (EDM) a notch approximately 0.4

cm deep by 5.8 cm long and subsequently extending the notch by bending

fatigue to a nominal size of 1 cm deep by 6.4 cm long. Approximately

100,000 cycles of 0 to 275 MPa (40 ksi) outer-fiber bending stresses

were required to precrack the specimens. The actual specimen dimensions

are summarized in Table 7. Flaw-size dimensions and bending fatigue

cycles are given in Table 8 . All artificial notches were introduced

into the outer (cap pass) surface of the weld, either in the center of

the weld (weld metal tests) or at the edge of the weld (HA2 tests).

The tests were conducted in a 4.4 MN (10® lb) - capacity electrohydraulic fatigue test machine equipped with hydraulically operated

friction grips.

The two test-load spectra shown in Table 9 were used. The fivestep spectrum was representative of the anticipated loadings for aboveground welds with a factor of ten on number of cycles for each step.

Use of the five-step spectrum was discontinued when the eight-step

spectrum representative of below-ground welds was supplied by Alyeska;

the eight-step test spectrum includes a factor of four on number of

cycles for each step.

Tests were conducted either in laboratory air at room temperature

or in a simulated-service environment at 60°C (140°F). The simulatedservice environment was an acidic solution (0.004% by wt. H2 SO4 ) in

water, pH = 3.2) with 5 drops per liter of poison (5 g P in 40ml CS 2 ) coupled with an impressed current of 0.1 mA/cm2. This environment was

developed by Battelle

"

1 to simulate the hydrogen-charging conditions that

are most representative of a defect or hole in the protective coating of a buried pipe.

1. G. P. Groenveld and A. R. Elsea, in Hydrogen in Metals , I. M.

Bernstein and A. W. Thompson, Eds. (Amer. Soc. Metals, Metals Park,

Ohio, 1974).

48
The laboratory air tests were conducted at a cyclic loading rate

of 3 Hz. Tests were continued by repeated applications of the load

spectrum until the crack extended through the thickness or until

a large number of cycles were applied. The environmental

tests were conducted at a cyclic loading rate of 0.0167 Hz; the complete spectrum was applied once. After completion of fatigue cycling

the specimen was loaded to failure and the maximum load recorded for

determination of residual strength.

Results are summarized in Table 10. The results indicate clearly

that deep flaws in pipeline welds do not grow appreciably in the anticipated service life of the pipeline - either in laboratory air or in a simulated holiday environment. The net section stress at failure ranged

from 580 to 640 MPa (76 to 84 ksi); these results indicate that the flaws

introduced reduced the load-carrying capacity of the pipe by less than ten

percent. The two natural-flaw specimens both failed Tn the base metal

remote from the weld defect. A photograph of a failed and an unfailed

test specimen is shown in Fig. 25. The fracture surfaces of the failed

specimen are shown in Fig. 26.

Table 7. Simulated-service- fatigue specimen dimensions.

Specimen

No

Weld

Procedure

Flaw Specimen Dimensions, cm(in)

Type Thickness

Width Weld Base

4-4SF 104A Weld 20.13(7.92) 1.85(0.73) 1.46(0.573)

4-9SF 104A Weld 20.14(7.93) 1.80(0.71) 1.47(0.577)

4-6SF 1 04A Natural 20.32(8.00) 2.01(0.79) 1.47(0.579)

4-5SF 1 04A Natural 20.19(7.95) 1.88(0.74) 1.47(0.578)

2-27SF 104A-1 Weld 20.07(7.90) 1.55(0.61) 1.24(0.488)

2-26SF 1 04A-1 HAZ 20.07(7.90) 1.52(0.60) 1.25(0.494)

2-29SF 104A-1 Weld 20.07(7.90) 1.57(0.62) 1.25(0.491

)

2-28SF 104A-1 HAZ 20.07(7.90) 1.75(0.69) 1.24(0.490)

5-1 SF 104A-1 Weld 20.07(7.90) 1.52(0.60) 1.21(0.475)

5-2SF 1 04A-1 HAZ 20.14(7.93) 1.50(0.59) 1.21(0.475)

5-3SF 104A-1 HAZ 20.17(7.94) 1.47(0.58) 1 .22(0.480)

49
Table 8. Pre-crack dimensions and bending-fatigue cycles.

Specimen

No.

Pre-crack

Dimensions,

a

Flaw

mm(in)

l

Bending

Kilocycles

Maximum

Stress, MPa

4-4SF 11.9(0.47) 59.4(2.34) 125 310

4-9SF 13.5(0.53) 63.8(2.51) 90 276

95 310

4-6SF — — — —

4-5SF ... -

-- -- 2-27SF 9.7(0.38) 64.8(2.55) 55 276

2-26SF 10.4 0.41 60.2 2.37 36 276

2-29SF 11.2(0.44) 62.5(2.46) 55 276

2-28SF 10.9(0.43) 68.8(2.71) 102 276

5-1 SF 9.7(0.38) 57.4(2.26) 146 276

5-2SF 9.4(0.37) 59.7(2.35) 95 276

5-3SF 9.4(0.37) 61.0(2.40) 81 276

Table 9. Test-load spectra.

Step Cycles Stress MPa (ksi)

Minimum Maximum

Five-step spectrum:

3

1 50 0 135. (19.6)

2 350 0 104. (15.1)

3 20 0 216. (31.3)

4 550 0 162. (23.5)

5- 200 0 180. (26.1)

Eight-step spectrum

J

3

1 8 36.5 ( 5.3) 154. (22.4)

2 40 -110. (-16.0) -21.4 (-3.1)

3 12 -no. (-16.0) 40.7 ( 5.9)

4 4 -no. (-16.0) 103. (14.9)

5 12 40. ( 5.8) 203. (29.4)

6 96 141. ( 20.4) 229. (33.3)

7 36 141. ( 20.4) 291. (42.3)

8 4 141. ( 20.4) 353. (51.2)

a. The number of cycles for each step is ten times

greater than the number anticipated in thirty

years of pipeline operation.

b. The number of cycles for each step is four times

greater than the number anticipated in thirty

years of pipeline operation.

50
Table 10. Simulated-service-fatigue test results.

Specimen

No.

Crack

Extension, 4a

mm(in)

Spectrum

Steps

Environment Temp. Spectrum

Cycles

Failure

Load

MN(kips)

Failure

Stress

MPa(ksi

)

4-4SF

3

0.76(0.030) 8 Air RT 14 — —

4-9SF 0.13(0.005) 8 Air/serv RT/60°C 1/1 1.21(271) 522(75.8)

4-6SFb — 8 Air RT 25 1.73(390) 580(84.2)

4-5SFb -- 8 Air/serv RT/60°C 1/1 1 .67(375) 562(81 .6)

2-27SFC 4.06(0.16) 5 Air RT 28 0.85(190) 521(75.6)

2-26SFd

3.30(0.13) 5 Air RT 14 — —

2-29SF 0.13(0.005) 8 Service 60°C 1 1 .09(245) 527(76.5)

2-28SF 0.46(0.018) 8 Air/serv RT/60°C 1/1 1.03(232) 533(77.4)

5-1 SF 0.13(0.005) 8 Air/serv RT/60°C 1/1 1.10(248) 544(79.0)

5-2SF 1.78(0.070) 8 Air RT 50 1.07(240) 551(80.0)

5-3F 0.25(0.010) 8 Air/serv RT/60°C 1/1 1.14(257) 566(82.1)

a. Machine malfunction, broke specimen.

b. Base-metal failure.

c. Estimated Pmax . d. Pmax not measured.

51
Fig. 25. Failed and unfailed fatigue-test specimens.

Fig. 26. Fracture surfaces of fatigue-test specimen showing:

(a) crack growth due to simulated-service fatigue,

(b) bend-fatigue precrack, and (c) EDM notch.

52
I
3H. SIGNIFICANCE OF ARC BURNS

The significance of random arc strikes on the base metal of the

trans-Alaska pipeline was evaluated under two "worst-case" conditions.

One condition simulated the maximum credible hoop-stress spectrum while

assuming a defect in the insulation at the location of the burns.

This permitted accelerated propagation of any existing surface crack

by hydrogen charging due to the cathodic-protection systemJ The

second condition simulated a sustained, static, hoop-tensile stress at

the yield point of the steel while assuming hydrogen charging. No evidence of crack propagation into the base metal was seen. Several cracks

were found in the arc-burn area after testing; however, the maximum

depth was 0.95 mm (0.037 in), confined to the interior of the arc burn.

The arc burns were implanted systematically in four 76.2 cm (30 in)

long by 15.2 cm (6 in) wide plates taken from Alyeska pipeline material

that was 1.17 cm (0.462 in) thick. The burn pattern and the burn intensities (times) are given in Fig. 27. Arc burns were implanted using

facilities at Ohio State University. 2 Burn times from 0.1 to 0.5

seconds duration resulted in spot burns 0.37 cm (0.145 in) to 1.2 cm

(0.47 in) in diameter. Drag strikes were simulated by an overlapping

series of spots for total burn lengths of 2.8 cm (1.1 in) to 5.4 cm

(1.8 in). Linked spots were oriented perpendicular to the intended

stress direction during subsequent tests. The appearance of the panels

after burning is shown in Fig. 28.

All panels were radiographed after burning, and no cracks were seen.

Hoop-Stress Fatigue-Spectrum Study

Panels 1-26AF and 1-27AB were subjected to the stress cycles and

the hydrogen-charging conditions shown in Table 11. The panels with

doublers attached to the tensile grips are shown in Fig. 29. To sustain

the tensile load within the arc-burn area, 15 nsn (0.6 in) were removed

from the panel sides. The testing was done at the General Dynamics

Convair Aerospace Corporation, San Diego, California.

The arc bums in the test plates were inspected visually, radiographically, and with dye penetrants for evidence of surface cracking.

Metallographic sections prepared from longitudinal cuts through the

linked spots of 0.1-, 0.3-, and 0.5-second burn times and through spot

burns in the tensile direction failed to show any crack extension from

the burn areas into the base material. Occasional small cracks were

observed to have propagated from the fusion zone into the heat-affected

1. G. P. Groenveld and A. R. Elsea, in Hydrogen in Metals , J. M. Bernstein

and A. W. Thompson, Eds. (ASM, Metals Park, Ohio, 1974).

2. T. J. Natarajan and R. B. McCauley, Welding J., 54, 879-884 (1975).

53
zone as shown in Fig. 30. Such cracks were only observed in the linked

series of spots made at 0.1 -second burn time.

Static-Stress Study

Panels 1-28AB and 1-29AB were subjected to static bending stresses

(arc-burn side in tension) slightly exceeding the yield point of the

material. The panels were subjected to the hydrogen charging conditions

in Table 11, except that the test temperature was 20± 3°C (68°± 5°F).

The method of stressing the plates is illustrated in Fig. 31. Plate

1-28AB was removed after 225 hours of exposure. No surface cracking

was observed visually, radiographically, or by dye penetrants. Subsequent examination of metallographic sections, prepared as described

previously, indicated a few small cracks confined to the burn area of the

0.1-second linked-spot region shown in Fig. 32. However, no cracks traversed the heat-affected zone into the base metal. Therefore, although

some arc burn radiographs show the presence of cracks in the crater, these

do not extend^ into jtne base metal.

Su p p 1 emen ta ry Da t¥ . . _ These test proqrams provided an opportunity to establish a corre- _ lation between arc -burn diameter and the resulting arc-burn depth and

heat-affected-zone depth. These data are shown in Fig. 33, which

includes a histogram of the arc-burn diameters reported in the preliminary Alyeska waiver request (Revision 9/6/76). This study also

provided an opportunity to determine the hardness profiles from the

parent metal through the heat-affected zones and into the fusion zones

for burn times varying from 0.1 to 0.5 seconds. These data are shown

in Fig. 34.

Table 11. Cyclic hoop-stress spectrum.

a,b

Step Min. Stress Max. Stress

108N/m2(103psi

)

No. Cycles

1 0 4.26(61.8) 8

2 0 3.23(46.8) 244

3 0 3.58(52.0) 68

a. Environment: 0.004% by weight H2SO4 in distilled water

plus 5 drops per liter of solution of 5 g phosphorous

in 40 ml carbon disulfide; 60° + 3° C (140° ±5° F);

current density 0.97 A/m2 (90 mA/ft2); specimen is cathode.

o 2

b. Specimen stressed initially in air to 4.26 x 10 N/m

(61.8 x 103 psi) after which the electrolyte was added and

the current impressed. Specimen held 8 hours, after which

the prescribed program was completed at a cyclic frequency

of 0.17 Hz (1 cycle/min) while maintaining the hydrogencharging conditions.

54
Fig. 27. Distribution of arc burns on test plates.

Test-plate/electrode identifications: 1-20AB = 5/32-in Lincoln 65+ shield-arc;

1-27AB = 3/16-in Lincoln 65+ shield-arc; 1-2CAB = 4 mm Phoenix Cel 80; 1-29AB =

5 mm Phoenix Cel 80. Arc-burn times in seconds are shown in Fig. 27.

Fig. 28. Completed arc-burn panels.

U
Fig. 29. (Left) Panels prepared for the

hoop-stress fatigue-spectrum

study.

Fig. 30. (Below) Examples of surface

cracks observed in the arc-burn

region of plates subjected to

maximum credible hoop-stress

spectrum while under hydrogen

charging conditions. All cracks

terminated in the HAZ. Overlapping 0.1-s burns. 200X.

56
Fig. 31. Test setup for static-stress study.

Figs. 32. Surface cracks

observed in the arc-burn

region of plates subjected

to static bending stress

while under hydrogen

charging for 225 h. Overlapping 0.1-s burns. 200X.

£
DEPTH (in)

0 0.02 0.04 0.06 0.08

DEPTH (mm)

Fig. 33. Variation of total arc-burn depth and heat-affected zone

depth as a function of arc-burn diameter.

r o

ARC-BURN

DIAMETER

(in)
Fig. 34

tf

. Hardness profiles across

arc burns of various burn

times and diameters.

O

<

o

05 o

oCNJ

CO

COQ

OC

o

<

GC

>-

CL. o

z

o

<

a

DISTANCE BELOW SUFACE (Mm)

59
31. CHARPY V-NOTCH IMPACT SPECIMEN TESTING

Charpy V-notch impact specimens were prepared from welds made by

each of the two welding procedures used (102A-1 and 104A-1). All Charpy

V-notch impact testing was conducted using instrumented impact testing

equipment that permitted the measurement of impact load versus time to

fracture so that the loads and times required to initiate and propagate

fracture in a Charpy V-notch specimen could be determined. These instrumented

Charpy V-notch tests can be used to determine the "dynamic" fracture

toughness of the test material as expressed in terms of energy (such as total energy to fracture) or stress-intensity (such as K^). The instrumented

Charpy V-notch test has been used extensively to evaluate the dynamic

fracture toughness of metals. However, a standard test method has

not been established but is still under active development

by the American Society for Testing and Materials (ASTM).

Test Materials

Charpy V-notch specimens were obtained from one weldment made using

welding process 102A-1 (weld No. 79907) and from one weldment made using

welding process 104A-1 (weld No. 88119). All specimens were taken so that the longitudinal axis of the specimen was parallel to the longitudinal

axis of the pipe. Four different orientations were used for the specimen

notch as shown in figure 35 to determine the effect of notch position

on the fracture toughness.

Test Variables

The test conditions were varied to include six temperatures, two

types of specimen notches (machined and fatigue precracked) and four

notch positions for each of the two weldments.

Testing Procedure for Instrumented Charpy V-Notch Specimens

All test specimens were conventional, full-sized Charpy V-notch

specimens (0.394 x 0.394 x 2.165 inches) with a 0.079 inch deep 45° V-notch.

Tests were conducted in a standard Charpy impact machine which is

fitted with instrumentation for measuring load and time during

impact ancTfracturfng'of the specimen.

For the specimens with machined notches, the total energy absorbed

was measured from the Charpy impact machine dial and the lateral expansion

and percent shear fracture were determined after fracture. Only specimens

with notch orientation number 1 (crack propagation parallel to the weld

axis) were tested with machined notches. From the load- time curves

obtained with the instrumented Charpy testing machine, the total absorbed

energy (Eq), the energy to initiate fracture (E^), the energy for fracture

propagation (E

p

), and the dynamic yield strength (ay<j) can be calculated

as follows.

60
The total energy En is normally taken from the machine dial reading

may also be obtained from the integrated load-time curve. This value, is

by the area under the Charpy V-notch and a normalized total energy value

is reported. The energy for fracture initiation, E ». is equal to

W

m

-E

c

, where WM

is the area under the load-time curve up to the

maximum load and E<~ is a correction for the compliance interaction of

the machine which is, Ec

* isP

2 C c , where P is the applied load and C c is

the compliance of the system. 1 The energy to propagate fracture, Ep = Eq-Env is obtained by difference. These energy values, Em and lip, are

normalized by dividing by the area A, similar to the normalized total

energy, Eq/A. The dynamic yield strength, ay^, can also be calculated

from the load- time record, as

but

d i v i ded

E~/A

,

u

where W is the specimen height (0.394 inch}, B is the specimen width

(0.394 inch), a is the crack length, and P^y is the load at which general

yielding occurs. The measured and calculated results of all tests on

machine-notched Charpy specimens are shown in Table 12. The measured

results of these tests are shown as a function of temperature in figures

Bfi-41

.

For precracked specimens, a fatigue crack of approximately 0.10

inch deep was developed under the machined notch. From the load

versus time curves, the total energy absorbed (Eg), energy for

crack initiation (EM ) , energy for crack propagation

(Ep), and dynamic yield strength were determined similar to the same

measurements made on the machine-notched specimens. In addition, for

the precracked specimens, the dynamic fracture toughness, K

T .,can be

calculated.

When fracture occurs under linear elastic conditions, the dynamic

fracture toughness, Kj , is calculated according to linear elastic

fracture mechanics assumptions and

6YPm , K

Id

= ~§W“

2 tU

where Pm is the maximum load measured from the load time curve, a is the

crack length, B and W are the specimen thickness and width, and Y is a polynomial function of a/W. Linear elastic fracture applies when the

maximum load, Pm, is less than the load required for general yielding,

P GY-

! W. L. Server and R. A. Wullaert, Dynamic Three-Point Bend Analysis for

Notched and Precracked Samples, Fracture Control Corporation Report No.

76-8.

61
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62
NOTCH

ORIENTATIONS

FOR

CHARPY

V-NOTCH

TEST

SPECIMENS

63

Position

3

Position
ENERGY

ABSORPTION

(FT-LB)

25.0

37.5

50.0

62.5

75.0

87.5

100.0

FIG. 36

RESULTS 0F INSTRUMENTED IMPACT EVALUATION 0F STANDARD CHARPY V-N0TCH

SAMPLES 0F NOTCH 0RIENTATI0N 1.

SPECIMEN IOENTIFICflT I8N

O IQ2R WELDMENT 1.0- 79907-

O

O ©

LO

<X1_

©

-S0.0 25.0 0-0 25.0 50.0

TEST TEMPERATURE (F)

75.0 Too.o

64 14 SEP TO acoa
7.

SHEfiR

FRACTURE

APPEARANCE

0.0

12.6

25.0

37.5

60.0

62.5

76.0

87.5

100.0

FIG. 37

RESULTS 0F INSTRUMENTE0 IMPACT EVALUAT I 0N 0F STANOARO CHARPY V-NOTC

SAMPLES 0F NOTCH ORIENTATION 1.

SPECIMEN IDENTIFICATION

O

0

0

-50.0 -25-0 oTo 25 .0

TEST TEMPERATURE

50.0

( F )

75.0 100.0

G5 14 SEP 7* SCOT
LRTERRL

EXPANSION

(MILS)

25.0

37.5

50.0

62.5

75.0

87.5

100.

0

FIG. 38

RESULTS 0F INSTRUMENTED IMPACT EVALUATION 0F STANDARO CHARPY V-N0TCH

SAMPLES 0F N0TCH 0RIENTATI0N 1.

SPECIMEN IDENTIFICATION

O 102fl WELDMENT 1.0- 79907.

O

U3

CM_

O

o

°-50 .0 -25-0 oTo 25.0

TEST TEMPERATURE

50.0

( F )

75.0 100.0

66 14 8EP 7B OCOfl
FIG. 39 RESULTS 0F INSTRUMENTED IMPACT EVALUATION 0F STANDARD CHARPY V-N0TCH

SAMPLES 0F NOTCH ORIENTATION 1.

°

SPECIMEN lOENTIFICflTiaN

o

2“! + I04fl WELDMENT 1-0- 88119-

87.5

o• Wc"+

+

FT-LB)

62.5 i

+ // +

o-os

)

NQIJ / +

Q>

Cd

CD

OD w

ODr.

aeon

+ /

+

+

ENERGY

25.0 i

in

• Cv)_

+

o

o 1 r- —i 1 i “i -50.0 -25.0 0.0 25.0 50-0 75.0 100-0

TEST TEMPERATURE (F)

6T

14 SEP DC08
X

SHEAR

FRACTURE

APPEARANCE

25.0

37.5

50.0

62-5

75.0

87.5

100.0

nu. ‘tu RESULTS 0F INSTRUMENTED IMPACT EVALUATION 0F STANDARD CHARPY V-N0TCH

SAMPLES 0F NOTCH ORIENTATION 1.

SPECIMEN IDENTIFICATION
LATERAL

EXPANSIGN

(MILS)

25.0

37.5

50.0

62.6

75.0

87.5

100.

FIG. 41

RESULTS OF INSTRUMENTED IMPACT EVALUATION OF STANDARD CHARPY V-NOTCH

SAMPLES OF NOTCH ORIENTATION 1.

oPECIflEN IDENTIFICATION

+ 104A WELDMENT I.Q. 88U9.

+

O

°-50.0

“i 1 1 1 -25.0 0.0 25.0 50.0

TEST TEMPERATURE (F)

75-0 100.0

69

14 SEP 7B SCOT
When fracture occurs after general yielding, the method for calculating

a correct fracture toughness value has not yet been established. However,

two approaches, the equivalent energy method 2 and the J-integral method 3 are used to calculate dynamic fracture toughness after general yielding

has occurred. For the equivalent energy method, the dynamic fracture

toughness, K

d

*, is

where:

Y, B, W, and a are the same as in eq. (1); E^ is the maximum load energy;

and C$ is the elastic sample compliance.

For the J-integral method, the dynamic fracture toughness is calculated

as

K

j d^

(3)

where E is the elastic modulus and

2E

M

J

M

= B(W-a

)

The results for all precracked Charpy V-notch test specimens are shown

in table 13. Values of dynamic fracture toughness calculated by the

linear elastic procedure, Kj d

, where applicable, or by both the equivalent

energy, Kd*, and J-integral method, Kj d

, are presented. In addition, the

critical crack opening displacement (CuD) value is calculated for each or

Kjd value by eq. (2), Section 3E, and by eq. (B-10), ADDendix B. of this

report.

2 F. J. Witt, Equivalent Energy Procedures for Predicting Gross Plastic

Fracture, USAEC Report ORNL-TM-31 72.

3 J. R. Rice, A Path Independent Integral and the Approximate Analysis of

Strain Concentration by Notches and Cracks, Journal of Applied Mechanics,

35, 379 (1968).

70
Table

13.

Results

of

Individual

Instrumented

Precracked

Charpy-V-Motch

Tests

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72
Discussion

The results of the machined notched Charpy V-notch specimens show

that weldments made with welding process 102A-1 may be slightly tougher than

104A-1 wel aments. A comparison of the total energy absorbed for these two

weldments is shown in figure 42. The other parameters measured in the

Charpy V-notch test: percent shear fracture, lateral contraction,

fracture initiation energy (Em), and fracture propagation energy (E

p

),

also do not differ significantly between weldments made with process

102A-1 and weldments made with process 104A-1. These tests show that

these weld metals have a minimum upper shelf Charpy V-notch energy of

65 ft-lb and that they reach this upper shelf value at approximately

100°F. These results are consistent with similar results obtained by the

British Welding Institute and Cranfield Institute of Technology.

For the instrumented precracked Charpy V-notch tests, the calculation

of stress-intensity based on the equivalent energy method (K<j*) gives

essentially the same result as the use of the J-integral method of

calculation ( Kjd) • A maximum difference of less than 4% is found between

the stress-intensity values calculated by these two different methods.

For tests at 68°F and 100°F, weldments made by both the 102A-1 and

104A-1 processes had a variability of less than 50% in the fracture

toughness measured on replicate specimens. This amount of variability

in toughness is not unusual for weld metals as seen by the variation in

Charpy V-notch energy and COD values measured by the British Welding

Institute. 4 At lower temperatures of 32°F, 10°F, -20°F, and -40°F, the

variability in the fracture toughness measured on replicate specimens

increases to a factor of approximately 2 to 1 . This increased variability

is because these tests are at a temperature range below the "upper

shelf" temperature for the Charpy V-notch tests, and therefore, the

fracture toughness results follow the usual pattern of increased variability

when the tests are conducted in the transition temperature range. The

variability of the fracture toughness was not affected by the welding

process (102A-1 and 104A-1) or the orientation of the notch in the test

specimen.

In general, weldments made with welding process 102A-1 had the same

fracture toughness as weldments made with welding process 104A-1. At

10°F, for specimens with the notch oriented in position 2 (part weld

metal, part HAZ), it appears that the 104A-1 process weldment may have a

4 Ref. [H-I], part 3, Appendix [H]

73
ENERGY

RBS0RPT

I

0N

(FT-LB)

12.5

25.0

37.5

50.0

62.5

75.0

87.5

100-0

FIG. 42

RESULTS OF INSTRUMENTED IMPACT EVALUATION OF STANDARO CHARPY V-NOTCH

SAMPLES OF NOTCH ORIENTATION 1.

SPECIMEN IDENTIFICATION

© 102A WELDMENT 1.0. 79907.

+ 104A WELDMENT 1.0. 89119*

O

© ©

©

+

-50.0 25*0 0.0 25.0 50.0

TEST TEMPERATURE (F)

75.0 100.0

74 occa
higher toughness than the 102A-1 process weldment, but the variability

in the data at this temperature does not permit a proper assessment of

this difference.

The effect of the anisotropy of the weld metal on the fracture

toughness was assessed by comparing the toughness of specimens with

notches oriented in positions 2, 3, and 4 at 10°F and -20°F. No measurable

effect of the direction of crack propagation within the weld was found.

However, if the fracture toughness at 68°F, measured with precracked

specimens, is compared with the fracture toughness measured in J-integral

tests (Section 3E), the weld metal is found to have a higher toughness

In the through- thickness direction (notch orientation 3) than in a direction parallel to the axis of the weld (notch orientation 2). This

effect is not as obvious for similar results at lower temperatures.

The fracture toughness for weldments made with process 102A-1 and

104A-1 decreases with temperature as shown in figure 43-. The fracture

toughness value calculated by the J-integral procedure is shown in

figure 443 and a lower bound limit curve is shown.

For convenience in comparing the fracture toughness results obtained

from precracked Charpy V-notch tests with the fracture toughness results

obtained with J-integral tests (Section 3E) and with COD tests, all

results shown in Table 13 were converted to COD measurements using an

analysis described in (Section 3E) and a second analysis described

by Begley (Appendix B). These results will be discussed in greater

detail in Section 6.

Additional parameters that are obtained from the instrumented

precracked Charpy V-notch tests are the total energy to fracture, energy

to initiate fracture, energy to propagate fracture, dynamic yield strength,

and stress-intensity rate. These parameters are not reported here

because they are not used in the fracture analysis for the pipe.

75
(Fig. 43) FRACTURE TOUGHNESS CALCULATED FROM PRECRACKED

CVN SPECIMENS BY J-INTEGRAL ANALYSIS

® SPCCinEN [OENTtriCRTI0N

TEST TEMPERATURE (F)

7G
3J. STRESS-CORROSION CRACKING TESTS

To evaluate the susceptibility to sustained load cracking in an

aggressive environment, specimens of weldments made by both weld processes

102A-1 and 1Q4A-1 were tested in a corrosive environment. The environment

used was a solution of 0.004 wt. pet. of concentrated H 2 S0 4 in distilled

water with an applied current sufficient to produce a current density of

0.625 mA/in 2

. A cathode poison consisting of 2 grams of phosphorus in

40 ml of carbon disulfide was added in an amount of 5 drops per liter of

test solution. This environment has been used extensively 1 to test

the susceptibility of steel line pipe to sustained load cracking and was

chosen to represent the most likely "worst case" environment that a pipe

will experience. A similar environment was used for fatigue tests

(Section 3F)

.

Test specimens approximately H in. x H in. by 10 inches long were

prepared from sections of weldments made by processes 102A-1 and 104A-1.

A notch, similar to that used in Charpy V-notch specimens, was put in

the weld metal in the notch orientation position 3 (fig. 35) so that

cracking would occur in a through-the-wall -thickness direction to simulate

a leak in the pipe wall. These specimens were precracked by fatiguing

and were then loaded in three-point-bending in the test environment.

Tests were conducted at 100°F which is the highest temperature at which

the cathode poison was soluble in the test solution. The testing procedure

consisted of breaking one specimen of each weldment in air, obtaining a load versus deflection curve for this specimen to failure, and calculating

the stress-intensity and net section stress at failure from the maximum

load sustained in this test. Additional test specimens were then loaded

to various stress levels below that required to cause failure in air,

and this load was maintained on the specimen in the test environment for

time periods up to 450 hours. At the termination of the tests in the

environment, the specimens were cooled in liquid nitrogen, fractured in

air, and examined at 50X magnification under a microscope to determine

if any sustained load cracking had occurred.

The results of these tests are summarized in Table 14. No evidence

of sustained load cracking was found in these tests. Since the applied

stress and the stress-intensity used in these tests is well above the

maximum levels expected in the pipeline, it is concluded from these

tests that crack growth due to an agressive environment is not likely to

be a problem in the Trans-Alaskan Pipeline.

*T. P. Groeneveld and A. R. Elsea, Hydrogen Embrittlement Testing, ASTM

SPP543 (1972).

77
Table 14. Results of Sustained Load Testing in Agressive Environment

Weld

Process Environment

Time to

Failure

Stress-Intensi ty

2

ksi/Tn, Kj

Net Section 3 Stress, 0

|y| , ksi

102A-1 air 0 121 209

cathodic 1 charging

230* 73 112

104A-1 air 0 78 138

cathodic

charging

456* 59 106

cathodic

charging

431* 58 94

cathodic

charging

431* 56 99

Test terminated at these times.

x See text for composition.

2 K

t = ^-^5

- /rra" f(a/w); f(a/W) given by G.C. Si h in "Stress- Intensity Factors

1 2Bl*r

Handbook.

3 a w = ——

s- , P-load, S-span, B-specimen thickness, W-specimen width, and

N 2B(W-a)^

a-crack length.

78
4. WELD DEFECT MEASUREMENT ASSESSMENT

4A. DESCRIPTION OF THE ROCKWELL INTERNATIONAL METHOD

OF WELD DEFECT MEASUREMENT

4B. DESCRIPTION OF THE SOUTHWEST RESEARCH INSTITUTE

METHOD OF WELD DEFECT MEASUREMENT

4C. CRITIQUE OF RADIOGRAPHIC DEFECT SIZE MEASUREMENT

METHODS

79
4A. DESCRIPTION OF ROCKWELL INTERNATIONAL METHOD OF WELD DEFECT MEASUREMENT

Introduction

The procedures used to determine the dimensions of defects in transAlaska pipeline girth welds were developed by Mr. Wayne D. Stump and

Mr. John L. Summers of the Rocky Flats Plant, Atomics International Div.,

Rockwell International Corp. The methods were originally employed to

determine the depth of defects in welds for nuclear applications.

Determinations were made at Alyeska facilities in Alaska. All

determinations were witnessed by NBS - Boulder personnel to confirm

that no change in measurement procedure occurred.

Arc Burns

The length and the width of arc burns were measured directly from

radiographs using either a ruler or a low-magnification eye-piece with

a calibrated reticle. Orientation of the defect relative to the girth

weld was traced on translucent paper placed over the burn area and

retained as a part of the documentation.

Planar and Non-planar Weld Defects

The depth (see fig. 45) of an internal defect within the weld metal

was estimated by comparing the radiographic density of the defect to that

of the adjacent non-defective weld metal. Density differences were converted

to depths using comparison standards. These standards were radiographicdensity step wedges reflecting density differences that would be produced

on radiographs of Alyeska-type welds. The density differences corresponded

to defects ranging in depth from 0.254 to 2.54 mm (0.010 to 0.100 in.).

The depth increments in the wedge were 0.254 mm (0.010 in.). Calibration

standards were prepared by radiographing a standard step wedge machined

from pipeline steel. This standard wedge was nominally as thick as the

weld, i.e., plate thickness plus 3.1 mm (0.125 in.). The standard was

radiographed at 225 kVp using the film and lead screens reported to be used

for the original Alyeska radiography. The standard radiographs were then

reproduced on a clear background to facilitate measurement. The comparison

standard is placed adjacent to the defect image on the radiograph and moved

until a match is obtained between the density of the defect and that on one

of the density steps of the standard. Between steps, visual interpolation

is used.

' Photographic density is a quantitative measure of film darkening defined by

the equation D = log (I

Q

/I-|) where D is the density, I is the light intensity

incident on the film ana ii is the light intensity transmitted through the

film. 1

80
"5

• mmm TD

O

c

Ci

s- G

«cion

of

wold

do

foots
The undulating surface of the weld metal causes significant variations

in radiographic density along the radiograph. Therefore, it is necessary

to select a reference standard having a base radiographic density (density

obtained for the full thickness step of the wedge) near the density adjacent

to the defect. In the present study, density adjacent to a defect was

measured by a microdensitometer, and a reference standard having a base

density ± 0.25 units of the desired density was selected. Defect depth

was independently estimated by two technicians. Results in reasonable

agreement, within 0.102-0.127 mm (0.004-0.005 in.), were averaged.

Significant discrepancies suggest a complicated defect (multiple superimposed defects, wormhole porosity, etc.) requiring further study. Five

to fifteen minutes were required to characterize a defect, depending on its

complexity. Demands on eyes necessitated frequent rest periods for the

technicians.

This technique is applicable for determining the maximum thickness

of a defect as projected on the radiographic film. For defects lying at

an angle to the incident radiation, the resulting measurement will be

substantially less than the overall defect height. Correlation with

overall defect height will be reasonably good for incomplete penetration,

elongated slag, gas pockets, spherical and cluster porosity, and for

hollow beads. Height correlation will be only fair for isolated slag

and is likely to be very poor for incomplete fusion and wormhole porosity.

Records

All data were recorded in ink in a bound laboratory data book, which

was dated and initialed by both technicians. In rare cases the technicians

concluded that defects identified as arc burns by Alyeska were actually

defects requiring a depth measurement (e.g., incomplete fusion). The

depth was measured and the discrepancy documented.

Accuracy Assessment

The bias and precision of the Rockwell International (RI) method of

measuring defect depth were evaluated by applying the method to radiographs

of seven control specimens. These specimens were sections of pipeline welds

containing a series of flat-bottomed holes and slots of known depth milled

into the underside of the weld. NBS - Boulder produced test sections from

welds in 1.17 cm (0.462 in.) thick pipe and from welds in 1.43 cm (0.562 in.)

thick pipe, all flush-ground on the bottom side to facilitate hole or slot

depth measurement. NBS - Washington produced a test section from a weld in

1.17 ran (0.462 in.) thick pipe, leaving the bottom side intact. Actual depths

were determined by the NBS precision measurements laboratories in Boulder and

Washington. The uneven underside of the NBS - Washington section required a high-side and a low-side depth reading for each hole. The hole depth was

taken to be the average of these two readings.

82
The artificial defects were distributed both laterally and longitudinally

along the welds as illustrated in Fig. 46. Depth measurements were made by

RI personnel, J. L. Summers and W. D. Stump, using the apparatus and standards

assembled for use on the Alyeska radiographs . Depths of the defects were

unknown to the technicians before testing; however, information on the

radiographic procedures was provided.

The measured defect depth is compared to known defect depth in Table 15.

Deviations from the true depths are qiven in Table 15 and plotted versus

true depth in Fig. 47.

The equation chosen to best represent the data is

D = D. - 0.0028 D* (1) m t t

where D is depth measured by the RI technique and D

t

is the true depth.

The constant 0.0028 was chosen by a least-squares fit. Deviations of the

D values from the fitted curve have a standard deviation of 3 mils for

true depths of 30 mils or less; this standard deviation increases with

depth, reaching 10 mils at D

t = 80 mils. Furthermore, since the slope of

the curve relating D to D

t

deviates appreciably from unity for large

depths, differences in D produce even larger differences in D.. As a result approximate "two-sigma" uncertainty limits for a predicted true

depth are ±

7 mils for measured depths less than about 30 mils; ± 14 mils

for D = 35 mils; ± 18 mils for D = 50 mjls; and ± 36 mils for D = 60 mils.

For measured depths between these values, the higher estimate of standard

deviation is conservative; for measured depths greater than 60, no

information is available. There was no significant effect on either bias

or imprecision due to type of defect (hole or slot), plate thickness, or

source of specimen.

The curve of D vs. D

t

is shown in Fig. 48. This curve can be entered

on the vertical scale using a value of D ; the corresponding estimate of D. is read from the horizontal scale.

It is possible also to approximate the curve of fig. 48 with a polynomial

- 2

D.(corr) = .95 D + .0052

t • m m

This equation gives the corrected depth accurate to within 1 mil for D up

to 60 mils; there are nr iata to support extrapolation beyond this

val ue.

k 83
Fig. 46. Specimens used to determine the bias and

inaccuracy of the Rockwell International

method of measuring defect depth.

84
Table 15* Results of radiographic determinations of defect depth on standard

specimens; all entries in units of 0.C254 mm = 10*3 in; L * D - D true meas.

or 0 0. 0

Specimen

true meas. A Specimen true meas. A

NBS, Boulder: NBS. Boulder:

0.462-in weld; 20.2 18 2.2 0.562-in weld; 55.8 48 7.8

1/32-in wide slots. 56.0 38 18.0 1/8-in dia. holes. 15.0 12 3.0

12.5 15 -2.5 81.0 63 18.0

21.2 20 1.2 12.0 10 2.0

15.5 18 -2.5 81.4 63 18.4

39.0 36 3.0 12.5 13 -0.5

21.0 20 1.0 38.2 28 10.2

• 38.2 43 -4.8 • 54.2 35 19.2

• 55.6 45 10.6 79.3 63 16.3

81.9 65 16.9 18.2 13 5.2

' 39.5 43 -3.5 55.5 45 10.5

56.4 38 18.4 11.4 • 9 2.4

40.5 33

‘ 7.5

* 19.5 20 -0.5

14.3 10 4.3 15.5 13 -2.5

81.5 50 30.

1

37.5 35 2.5

13.7 ' 10 *‘3.2 19.5 20 -0.5

81.5 50 3V; 5 37.5 38 -0.5

. 20.5 15

' ~5.5 55.5 50 5.5

56.0 53 3.0 80.0 70 10.0

80.0 >60 — 37.5 40 -2.5

NBS, Boulder: . 0. 562-in weld; 80.7 60 20.7 N8S, Washington:

1/32-in wide slots. 20 18 . 2.0 0.462-in weld; . 15.5 21 -5.5 • 56.5 38 13.5- 1/16-in, 1/4-in 23.7 19 4.7

12.1 15 -2.9 diam. holes. 9.3 10 -0.2

19.6. 20 -0.4 40.8 33 7.8

15.5 18 -2.5 17.5 IS 1.5

38.8' 36 2.8 4.3 <5 —

19.5 20 -0.5 36.5 32 4.5

* 38.0 43 -5.0 26.3 23 3.3

55 45 10.0 30.4 30 0.4

80 65 15.0 50.2 45 5.2

38.5 43 -4.5 30.1 33 -2.9

55.8 38 17.8 25.5 2S- 0.d

38.8 33 5.8 13.3 U 18

• 12.5 10 2.5 58.9 53 5.9

80.4 50 30.4 40.0 36 4.0

12.7 10 2.7 1 6.0 — -

80.9 50 30.9. i 72.5 63 9.5

20.5 15 '5.5 38.8 43 -4.2

55.0 53 2.0 19.4 ’ 25 • -5.6

81.5 73 . 8.5

NBS, Boulder:

n.4fi?-in weld: 19.0 15

«4.0 Rockwell (Rocky Flats)

71.4 45 32.4

1/8-in diam. holes. 55.5

12.3

37

15

15.5

2.7

0.462-in base metal. 54.6

6.8

52

9

2.5

-2.2

18.4 18 0.4 32.6 30 2.6

11.5 11 0.5 19.8 22 •2.2

38.0 35 3.0 58.8 56 2.i

18.5 20 -1.5 12.9 15 -2.1

38.0 40 -2.0 49.0 45 4.0

# 55.3 50 5.3

38.7 36 2.7 .

-

. 54.5 50 4.5

38.0 38 0.0

12.0 10 2.0

81.7 75 6.7

#

• • • 11.5 10 1.5 9 80.0 85 -5.0

18.3 13 5.3

55.3 50 5.3

80.2 58 22.2

80.7 58 22.7

85
DEP1

>V.easuredDEPTH

true

(10

~

88
100

m

l

CL- UJ

ca

a

LU

QJ

CO

80 -

60

40

20 -

Measured slots and holes used to

fit curve (Table 15)

.

Field radiographic dimension

measurements (Table 16a). +

0

0

*

+

20 40 60 80 100

D

t=TRUE DEPTH (1CT

3

in)

Fig. 48. Bias curve derived from the RI radiographic defect depth

experiment based on laboratory radiographs. Additional

data from depth measurements of material defects (Table 16a

are plotted as circles.

8 ?
The above curve and uncertainty limits apply only to the type of defects

actually considered in the above we 11 -control led experiment. However, it is

felt that these results can be safely applied to many kinds of defects met in

problems. The major exception to this would be those subject to orientation

problems as noted above. The depths of such defects are likely to be underestimated by the RI technique, even with the bias correction. The uncertainty

for such measurements may well be larger than stated above.

The ability of the RI method to measure the overall height of

natural defects appearing in a series of Alyeska radiographs was also

evaluated. As discussed previously, any method of converting defect

contrast into defect depth will underestimate the overall height of a defect inclined to the axis of radiation. A series of 15 natural defects

were identified in six of the Alyeska radiographic films associated with

girth welds supplied to NBS for fracture mechanics study. The RI

interpreters estimated the depths of these defects. NBS identified

sections of the weldments containing the defects and a 0.5 in. long

segment of the weld containing each defect was removed. Each segment

was ground flat and was radiographed parallel to the axis of the weld

to show the defect in profile. Overall defect heights were measured

directly from the radiographs using a calibrated filar eyepiece and

low-power microscope.

Data from this experiment are summarized in Table 16a. The information has also been added to Fig. 48, plotted as circles. As expected,

the RI method underestimates the overall defect height. Although no

correlation is observed between apparent film contrast, AD, and measurement

error, the spread of film contrast values (Table 16A) was generally in

the normal, predicted range. After correction for bias according to

Fig. 48 and Table 15, the measured values had an estimated relative standard

deviation of about 47%. This natural defect experiment gives an added

degree of confidence in the RI method.

88
Table

IGA.

Measurement:,

of

Natural

Meld

Oefects

as

Confirmed

by

Laboratory

Radiography

Projected

Film

D

ens

ity

and

Co

ntrast

Alyeska

Defect

Rl

Difference

l

of

Defect

PenetraOj

D

?

°l"^2

Radio

graph

Height

Measured

Difference

Type

ate

ter

t

In.*

Depth

Thickness-

|

CMCMtOkOtOkOCMCMCMCMCMCMCMCMCM

on ao cn w

©

CNJ

I VO © ® O

co cm to

i i i

CV* co

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vo

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00 to

I

00 co cn o* CM CM cn CO a— VO co

CM CM *-» O o co o VO a— o o

o o © o o e o o o o o o o o o

o o o o o o o o o o o o o o o

cn VO o CO o CO o O o 03 CM CO to

o* CM CO CO CM CO CO CO CO CM o o o o O o o o o o o o O o

o o o o o o o o o © o o o o

o

CMo

I

1

t— CM VO co o cn CO r*. VO a CM

O*. VO CM fO CM kO CM Ov CM VO co CM VO o o c a o o o o ® O o o o o O

o o c o o o o a o o e o o o o

—^ ^ co o o i o

O O e— O* I I r*. i i o O <

CO GO « —* CO

.—

.

—

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CM CM CM CM CM O o o O o to VO vo to ^ ^ VO »— QC cc — Q£ X— GO CO co 00 —

1 CO 1 CO 1 CO 1 co 9 co 91 co CO CO 1 tO o to o to o tO o i O 1 o • O 1 o o co to cn to cn to cn to VO vo r— to e— VO O

co CO •— co e— 00 CM CO CM CD CM CO CM CO *TM“ *co ^co ^G0 >—00 >-* CO 'CO ^-CO^'CD'—’

01 JQ

E

C

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a>

to

to3 CO 4- 01

tf» TO 4-»

-*-» ro OJ T3 0J <U +-» U

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tn G- £23 U. UJ O 31 '

89

*

Was

determined

by

using

a

filar

eye

piece

and

a

0.66X

objective

in

a

microscope,

ca

unit

a

9x10

to

the

-3

millimeters.

**

See

Section

4B

for

a

discussion

of

film

density

and

4C

for

a

discussion

of

contrast
4B. DESCRIPTION OF THE SOUTHWEST RESEARCH INSTITUTE METHOD OF WELD

DEFECT DETERMINATION

The Southwest Research Institute (SWRI) method of measuring weld defect

dimensions was reviewed in a visit to SWRI on August 16, 1976. Discussions

were held primarily with Mr. Sam Wenk. Mr. Wenk discussed the methods used

by SWRI to determine defect dimensions. For the length and width indications,

a plastic scale was used for the measurement. Wenk felt that, in most cases

(as in slag or porosity, for example), these dimensional measurements could

be made to an accuracy 0.010 inch. For a planar defect, such as lack of

fusion, these dimensions become somewhat more difficult.

For the through-wall dimension measurement, a series of radiographic

density measurements or an estimate of density difference as compared to the

density difference for the radiographic penetrameter and shim are used. For

defects covering an area large enough for a densitometer measurement (densitometer aperture is 1 mm dia.), a formula is used to determine the throughwall reduction (R) (see Fig. 49):

where T

D

D

D

D

1

2

3

4

thickness of penetrameter and shim;

radiographic density in area under the pipe wall;

density under the penetrameter, shim and wall;

density in the weld area near the defect; and

density in the defect area.

This equation is derived in Appendix L. The derivation of this

equation is based on the assumption of narrow beam (non-scatter) conditions.

The accuracy of the equation for the assessment of defect depth cannot be

determined without a careful investigation of scatter effects.

Since returning from Alaska in July, SWRI personnel have radiographed

several steel calibration samples in an attempt to confirm their dimension

measurements. These calibrations originally resulted in a uniform correction

factor in which 0.025 in. would be added to all SWRI densitometer-determined

depth measurements. Present estimates by SWRI are that their depth values are

low by varying amounts. Based on recent X-ray calibration tests with the

API 1104 grooved plates, SWRI* has informed NBS that they now estimate the

depth readings obtained using the above formula are low by 50% for internal

surface defects such as incomplete penetration and by 10% for other defect

types. They believe the internal defect measurement should be increased by

50% up to a maximum of 0.025 in. correction. No defect densitometer readings

(on actual field weld radiographs) were taken for defect widths (or diameters)

less than 0.040 in. since the aperture effect was recognized (see Section 4C).

* Telephone conversation October 5, 1976 between E. Ruescher (SWRI) and

H. Berger (NBS).

90
LOCATION

of

DENSITY

MEA

.SI)BFmfk|

TS

Si
There were many defects (an estimated 75%) which were too small for

accurate densitometer measurements using the 1 mm aperture. In these cases,

an estimate of the change in density was made by a visual comparison with the

density change observed for the penetrameter and shim. The accuracy of these

so-called eyeball measurements is undeterminable as the methods obviously

depend on the individual observer.

Discussions with Mr. Wenk covered some of the problems with regard to

these through-wall dimension measurements. Film processing is a variable that

is presently not controlled. The film type also varies. The pipe wall

thickness itself varies (specifications permit wall variations from -5% to

+10%): according to Mr. Wenk, most of the pipe wall runs about 5% high.

All these factors introduce difficulties in interpreting the density

differences. The density measurement (where it can be done) is also difficult.

Mr. Wenk indicated that SWRI had used this density difference approach

(using actual densitometer measurements) in several previous nuclear problems

with fair results. In one case, the technique was used to determine wall

thickness; the radiographic measurements correlated with other measurements

to within 12.5%. This case was not complicated by large thickness changes

as is the weld situation.

In considering the SWRI method, it is important to note that the

comparisons were always made with the penetrameter-shim image— for both

the methods used by SWRI. Therefore, assuming a constant shim-penetrameter

thickness, this technique would be less sensitive to changes in film

gradient (changes described in Section 4C) than the method used by RI.

Accuracy Assessment

An error estimate for the SWRI visual method will not be given. For the

densitometer method, an approach to an accuracy assessment is to consider the

possible variations in the density measurements used in the formula for R.

The logarithmic ratio of numerator and denominator produces a large

sensitivity to small changes in density and in particular to the densities

referring to the defect (D

3

and D,). Empirically determined or published data

lead to estimates for the variability of R. The values of D, and D~ should be

capable of being determined to an accuracy of ±0.03 density units since they are

measured over a relatively large area of reasonably uniform density. Density

D

3

, which can vary due to weld bead thickness changes has been established by

experimental tests of the 3 weld radiographs available for our tests; in these

tests the variation did not exceed ± 0.225 density units.

Two methods were used to determine this D^ variation. The first was a direct scan of several radiographs along the weld bead using a precision microdensi tometer. The other was a calculation of the direct density change by means

of the simple equation 1 (neglecting scatter):

AD = u eff G AX

2.3

1 R - Halmshaw, "Physics of Industrial Radiology," Heywood Books, London (1966).

32
Where: = effective linear attenuation coefficient for the X-ray beam,

G = film gradient (slope of density vs. log relative exposure

curve) obtained from published sensitometry data,

AX = change in thickness of the sample traversed by the X-ray

beam (i. e. weld peak to weld valley difference).

This equation is derived in Appendix L.

The possible change in was determined empirically by scanning a known

size defect on a radiograph with a precision scanning densitometer and comparing this to a similar measurement with a densitometer of the type and aperture

size used by SWRI. The accuracy in was estimated in this way as about

± 0.05 density units for defect sizes

4

1.25 or larger times the aperture

diameter.

These possible variations in density readings were used in the formula

for R to estimate upper values of the determination of R. Two estimates have

been made. In what might be called a typical case, D, and D~ have been allowed

to vary by 0.03 density units, D* by 0.2 and by 0.u5 density units, in each

'case so as to increase the value of R. Values determined by these typical possible

density variations are compared to calculated R values in Table ICb.

Also in the table are estimated likely maximum R situations in which

variations are 0.05 for D-j and D

2

, 0.225 for D-, and 0.1 for D*. These

calculations of R are only estimates as to possible variations.

Even these estimates of density measurement inaccuracy may be too small

particularly for defects that are tear-shaped or sharp and crack-like. In

these cases, the measurement of the D, with an aperture size of 1 mm would be

very difficult, the measurement would normally be much lower than the peak

value. Therefore, these estimated R values given in Table 16b are generally

useful only for round or flat bottomed defects having a projected area on

the film of about 0.050 in. or greater in diameter or width. This might

include defects such as gas holes, spherical porosity, and slag. Defects such

as incomplete penetration are likely to be too narrow for measurement. Defects

such as hollow bead may be tear-shaped. Therefore, care should be exercised

in interpreting the values in Table 16b.

An evaluation of the uncertainty of values of R can be made from an

approximate statistical analysis of the density readings. Although density

varies from film to film, the difference D,-D

?

should be self-correcting with

respect to variability due to film-to-film differences. Using the differences

D-t-D,* the standard deviation of a density reading is estimated to be 0.12

density units. (In addition to errors in density measurements, this may also

include variability due to radiography variables such as X-ray energy, film

processing, scatter, etc.). Notice that it is larger than the "typical"

value of 0.03.

If it is assumed that errors in the four density readings are uncorrelated

,

a standard propagation-of-error calculation indicates that the relative

standard deviation of R is approximately 85%.

In this discussion of the SWRI method it has been assumed that the

thickness (t) of the shim-penetrameter combination is known. Shim variations,

considered as one possibility for film contrast variations discussed in

Section 4C, will also have a bearing on the accuracy of the SWRI method.

93
table

16b.

Film

Density

Data

and

R

Calculations

oc

<+-o

V)

a)3

a

<u

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o *— a> * * w

i *i rtj H zo

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Table

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35
4C. CRITIQUE OF RADIOGRAPHIC DEFECT

SIZE MEASUREMENT METHODS

The assessment of defect sizes from radiographs of the girth welds

on the Trans-Alaska Oil Pipeline depends ultimately on two factors:

(a) the judgment and experience of the radiographic interpreter,

and

(b) the relationship of radiographic film density changes to

pipeline steel thickness.

Both factors are complicated by the varying film density due to steel

thickness variations introduced by the weld bead.

The radiographic procedure in most cases involves the insertion of

the x-ray source into and centered in the pipe and aligned as nearly as possible in the plane of the girth weld. The 70 mm film, lead screens,

and steel penetrameter with shim to simulate the weld bead thickness

used in the Alyeska procedure 1 are placed on the outer circumference

of the weld joint. Two or more strips of film are used on each weld. An

overlap of film is used to assure complete coverage. Processing of the

exposed film is generally done in a Kodak Model M-7 automatic film processor, modified to reduce the speed of the film transport through the

processor. However, our information is that some radiographs from the

field radiography work were hand processed.

Two independent teams and techniques were used to assess the weld

defect sizes considered in the waiver request. As described in this

report, those teams and measurements are referred to as Southwest Research

Institute (SWRI) and Rockwell International, Rocky Flats (RI) 2 (see

preceeding Sections 4A and 48).

It should be pointed out that the evaluation of the pipeline field

radiography is based on information supplied to NBS by the evaluating

teams previously mentioned, the Alyeska radiographic procedure 1

, and

telephone conversations with the Radiographic Services Division of the

Alyeska Pipeline Service Company. The original Alyeska radiographs have

not been critically examined by the NBS evaluation team, nor have the defects

listed on the waiver request been checked independently by this group.

Nevertheless, all information indicates that the radiographs were of

good quality.

A list of weld defect locations on the girth weld is

given in Appendix A. Defect sizes as evaluated by SWRI and RI are

given in Section 8.

1. "Radiographic Examination of Pipe Welds," Alyeska Specification 82.1.

2. J. Landolt, W. Stump, and J. Summers, "Radiographic Determination of

Defect Thickness", Sept. 1976, to be published.

9S
The two major radiographic factors indicated above are evaluated in

greater detail in the following sections.

Judgment

Both sets of film readers, from SWRI and RI have appreciable

radiographic interpretation experience. Visual experiments indicate

that experienced observers provide consistent, but not necessarily

accurate results 3

. With the limited number of observers in this

case, there is no obvious way to assess the accuracy of the visual

measurements as concerned with the Alyeska field radiographs. One would

need to make use of more observers and destructive tests on defect weld

specimens to define the accuracy as it applies to field radiographs.

It is known 4 from visual tests that an observer looking at a small high contrast target, such as the radiographic image of a small

area but relatively deep wormhole porosity image, would tend to underestimate

the greyness relative to that for a larger or lower contrast target.

Therefore, large density differences from defects of larger depths would

tend to be underestimated by this visual comparison. On the other hand,

another factor in the visual comparison method used by SWRI is that the

penetrameter shim image, of good sharpness because those objects are

on the film side of the pipe, would tend to be visually regarded as of

higher contrast than it is. The reader would then tend to overestimate

the depth of the defect being examined. The degree to which these

factors compensate each other is not known.

The eye has a much better capability for matching two shades of

grey adjacent to each other 5

, as used in the RI method of defect

measurement. The finer steps in grey-scale used in the RI method also

lead to a better visual comparison. Although the error of the RI method

as compared to laboratory specimens radiographed under well -control led

conditions has been assessed in Section 4A, the error of the RI method

and other visual comparisons as applied to less controlled radiographs

is another matter. In the field case, the poorer control between radiographic

film density and variations in pipe wall thickness will lead to greater

complications in the assessment.

There is also the matter of the film density value itself. For total

densities greater than about 4, the accuracy of a densitometer measurement

is in question. Some of the measured film densities in the weld area exceed

this value. In addition, limitations on available illuminators make it

3. R. S. Woodworth and H. Schlossberg, "Experimental Psychology, "Henry

Holt and Co., New York, NY (1960) pg. 746.

4. 0. Bryngdahl , J. Opt. Soc. Am. 56_, 811 (1966).

5. R. N. Evans, "An Introduction to Color", John Wiley and Sons, Inc., New

York, NY (1948) pg. 128.

97
difficult for an observer to see density changes in this range. Therefore,

visual measurements also can be influenced by high film densities.

Relationship Between Film Density and Changes in Steel Thickness

The analytical material which follows relates to Eastman-Kodak Type

AA X-ray film. It is our understanding that about 85% of the girth

welds were radiographed with AA film; therefore, our efforts were concentrated

on that film. The particular film related to any given weld radiograph

under consideration for waiver should be determined. If it is AA film,

our analysis should be valid. Other comparable speed x-ray films are

expected to be similar; however such cases should be treated on an

individual basis.

Table 17 lists all of the variables that affect film contrast. One of

the variables in relating film density to changes in steel thickness is the

effective x-ray energy which is related to tube voltage*, it is our understanding that this is not routinely logged in a field environment so it

is unknown in the case of the field weld radiographs. The radiographic

procedure 1 permits x-ray equipment with a range of tube voltage from 140

to 300 kVp. From data supplied by Alyeska, it appears that field radiographs

were taken over the x-ray tube voltage range of 190 to 240 kVp, with the

most likely range being from 190 to 225 kVp.

At the lower voltages, the subject contrast will be greater,

thereby producing a relatively larger change in film density for the same

defect depth dimension. Figure 50 presents typical curves for inspection

of steel at the two extremes of the likely tube voltage range, 190 and

225 kVp. As an example, a defect depth of 0.060 in. will yield a change

in density (aD) of 0.53 density units for 190 kV versus a aD of 0.45

for an x-ray energy of 225 kVp. The slopes of the two curves are 8.81

for 190 kVp and 7.52 for 225 kVp. This represents about a 15% change in

slope.

A second variable concerns the linearity of the log film density

versus metal thickness plot. This linearity aspect is discussed in the

report of consultant R. C. McMaster in the Appendix K. McMaster examined

the linearity of this plot as a means of determining the change in metal

thickness due to a defect. Note the graphical derivation of the SWRI

equation in Fig. K-10 of the McMaster report (Appendix K).

Published Kodak data (Fig. 51) shows significant departures from

linearity at higher densities. These data show the departure of linearity

for densities greater than 3 +

; many of the Alyeska films measured show

densities above this value and even higher than 4. Therefore, this

*Tube filtration used by Alyeska radiographers is of little importance

in this discussion.

+ Kodak data is for a net density of 3.0, etc. The total density including

background, etc., would be about 3.2.

98
Table 17

POSSIBLE X-RADIOGRAPHY VARIABLES

1. X-RAY VOLTAGE

2. PROCESSING

3. SHIM THICKNESS CHANGE

4. FILM TYPE

5. PIPE WALL THICKNESS

6. LEAD SCREENS AND CONTACT

7. SCATTERING EFFECTS

8. SOURCE POSITION

9. FILM TEMPERATURE DURING AND AFTER EXPOSURE.

X 39
RADIOGRAPHIC

FILM

DENSITY

.100
-Is-

<0

in

ro

CVI

A1IS N3Q 13N

/«/

EXPOSURE

(ROENTGENS)

Fly.

51.

Net

film

dpnsity

a^

a

function

of

x-ray

exposure.
departure from linearity, as much as 10% at a film density of 4, would

lead to lower depth measurements using the SWRI equation. However,

linearity changes probably have a minor effect.

Another variable is the actual density measurement for small defect

sizes, defect indications that approximate the size of the densitometer

aperture. For cases in which the aperture is larger than the defect

density region, one will obtain a lower peak density, reflecting the

averaging effect of the aperture (upper curve. Fig. 52). Only when the

aperture diameter becomes smaller than the defect size, will a density

reading reflecting the true peak value be obtained. Therefore, density

measurements on defect indications determined by SWRI can be in error,

depending on the defect indication size. For small defect sizes, the

readings would tend to be low, thereby leading to an underestimate of

the defect depth. Information in the literature 6 indicates that an

aperture comparable in size to the defect width can give a low reading

of about 20% for a rounded versus a flat contoured defect. For defects,

not rounded and which show small deep penetrations, such as a tearshaped hollow bead, this correction factor may be higher than 20%. In

recognition of this problem, SWRI personnel* have informed NBS that

densitometer measurements were not made for defect sizes that approached

the aperture size. An experiment was conducted here to determine what

density corrections might be required for density readings made under

these conditions (see Section 4B).

Variations in film density in the weld bead due to variations in

thickness make it difficult to determine the change in density between a defect and the weld bead. These variations (discussed in Section 4B and

Appendix N) limit defect detectability and measurement. One alternative

approach to compensate for these density variations is to compare the

defect indication density to the lightest part of the adjacent weld

image. This will produce the highest contrast on aD and will lead to an

accurate or overestimate of defect depth.

The most important observation made in regard to film density

variations relates to the changes in film density for the shim-penetrameter

image on the pipe wall. This AD** for a known change in thickness

should generally increase for increasing film density, because the film

gradient increases 7 * 8

. However, information obtained from SWRI (Density

Data sheets copied at DOT) has been analyzed for the variation in aD.

For an assumed penetrameter-shim thickness value of 0.111, Figure 53

shows a very wide spread in the aD observed for this change.

telephone Conversation, October 5, 1976, between E. Ruescher (SWRI) and

H. Berger (NBS).

**aD here means the difference between density values at two points on the

radiograph.

6. E. L. Criscuolo, "Radiography and Visual Perception," Nondestructive

Testing, 2£, 373 (1962).

7. R. Halmshaw, "Physics of Industrial Radiology," Heywood Books, London (1966).

8. "Radiography in Modern Industry," Eastman Kodak Co., Rochester, NY,

14650, Chapter 14. See also Figures 55 and 56.

il

102
<o

CO

q:

LlJ

l—

LlI

o-H

cr

Ixl H

lxl-1

ozhO

coco

CO z

LlI 2: a

LUQZO

UJ • N

CO

UJ

iroH

tr

UJ

CL • < w 0 2 u- ixi Ll

O' 2 LlI “ Q

h*o

LlJ lJL

Ll111

COco

a>>

4-> ra*

C£

CSJ -O

1U 3
in

3

a -'a aouaj8j. sua Ajisuaa

104

Fig.

53.

Film

contrast

measurements

vs.

base

density

for

a

shim-penetrameter

thickness

of

0.111

in

See

text

for

discussions

of

the

lettered

lines.
These same data are plotted in a different way in order to show the

general shape of the distribution in the plot (Fig. 54) of frequency of

occurrence of a particular value of aD.

In the following discussion reference can be made to either plot as

desi red.

The variation of data shown in Fig. 53 has been analyzed as follows.

The average values of (

D

x -D z ) and were determined. This represents

what should be an average condition for the pipeline radiography. At

the value of Di, the gradient for type AA film (automatic processed) was

determined. Then the effective linear absorption coefficient (u --)

was calculated using equation 1 (see Appendix L):

aD

u eff G

D aX O ( 1 )

where:

aD = average value of Di-D 2 observed from Alyeska film data in

Fig. 53.

G- = film gradient for AA film measured for automatic processed

u film at the average Di value;

u rr - effective linear absorption coefficient for the pipeline

eTT

steel;

aX = assumed thickness change = 0.111 in. = 0.282 cm.

Using y eff and actual values of D lt values of aD were calculated

using the above equation. This determined line A on Fig. 53; it

represents the AD value one should obtain for a 0.111 in. thick steel

piece on a 0.562 in. steel pipe wall based on the film response.

That predicted relationship (line A) was broadened by allowing the

linear absorption coefficient to vary 15% (±7.5%), as determined from

Fig. 50, Band B. An additional variation was determined by allowing the

film gradient to vary due to reasonable, possible variations in film

processing. Kodak data on variations of film gradient as a function of

processing variables are given in Figs. 55 and 56 for automatic and hand

processing, respectively . Practically speaking the film gradient only goes

down (contrast decreases) as one departs from recommended automatic processing

temperature (85°F) or hand processing time (5 min. at 68°F), as shown in

Figs. 55 and 56. Curve C, Fig. 53, represents a combination of high x-ray

voltage and less than optimum automatic processing (95°F). Curve D represents a combination of high x-ray energy and less than optimum hand processing

(2 min. at 68°F). Curve E represents a purely mathematical look at the data

points; it is a line that represents a least squares fit to the data points.

105
FREQUENCY

DISTRIBUTION

OF

CONTRAST

MEASUREMENTS

in

CM

CD

in c

.po o

ri E

<23

CM Q

S’O ii Cl

<3

CO

» « o

PO

CM CO CO CM oo

ssNiara jo yjawriN

108

Fig.

54.

Distribution

of

contrast

measurements

of

Alyeska

radiograph
c

«

‘5

o

w

o

£

iZ

Fig. 55. Film Gradient vs. film density eor automatic orocess^ng of

Tyoe AA Film.

107
Fig. 56. Film gradient vs. film density for hand processing of' type AA film
Two additional points, denoted by 1 and 2, have been added to Fig.

53. Point 1 represents the expected contrast one would obtain if the

average radiographic conditions determined from Fig. 53 were applied to

a case in which the shim thickness was considerably larger than its

normal value (0.158 in. instead of 0.096 in. so that the shim plus

penetrameter was 0.173 in.). Point 2 was obtained in a similar way but

allowing the shim thickness to decrease by almost a factor of two (0.058 in.

plus penetrameter equals 0.073 in.). This shows that the large

spread in contrast displayed in Figs. 53 and 54 could essentially be

explained on the basis of changes in shim thickness. Said another way,

the variation in film contrast means that an observer looking only at

the film density change caused by the shim-penetrameter combination

might read a value as small as 0.073 in. or as large as 0.173 in. for

the presumed actual thickness of 0.111 in. These differences almost

cover the extreme values shown in Fig. 53.

Some support for the changing shim thickness explanation of the

varying contrast might be obtained from RI measurements of the depth of

the penetrameter hole. Once the extreme variation in film contrast

was observed, the RI interpreters were contacted in Alaska. They then

began to read the penetrameter hole depth using their comparison method.

In all these cases, the 0.015 in. deep hole was read as 0.010 to 0.012 in.

deep. Of the 31 weld films read in this way, only 10 were among those

for which NBS had density data. One low contrast reading was not considered

because it involved a tie-in weld. Only two other contrast readinqs were far

removed from predicted film contrast values. It turned out that both these lew

and hiaheontrast values were for a single weld (#81142), strongly implying

that there were two different thickness shims on that weld radiograph.

It should also be mentioned that in an experiment conducted at the

Naval Surface Weapons Center* under conditions very similar to the girth

welds discussed here, the radiographic contrast for a shim and penetrameter

0.1 in. thick varied from about 0.6 at 250 kVp to about 1.25 at 140

kVp. These variations are about equivalent to the density changes

reported by SWRI.

Table 17 lists the possible explanations for the film contrast

variations observed in Figs. 53 and 54. Voltage changes, processing

conditions and factor-of-two variations in shim thickness have been

discussed.** The film type is known to change; this discussion is valid

only for Kodak Type AA x-ray film. The pipe wall can vary -5 to +10%;

*£. Criscuolo and D. Polansky, Naval Surface Weapons Center, Silver

Spring, MD, private communication (Oct. 1976).

^Slrim thickness changes

S2.1 permits the choice

crown) but NBS has been

telephone conversation,

f ,\|QC

^ IIUw /

•

may be a possibility (the Alyeska procedure

of shim thickness to approximate the height of the weld

assured that shim thicknesses remained constant

Sept. 29, 1976, J. Willing (Alyeska), and H. Berger

109
extremes in wall thickness for the 0.562 in. wall would be 0.534 to

0.613 in. A change of steel thickness of 0.111 inches on these two

extreme wall thickness values yields a calculated change of about 2% in

radiation transmitted for an x-ray voltage of 225 kVp (which has peak in

the energy spectrum at about 150 kVp). The other variables do not

permit assessment without more information than is presently available.

A recognized possibility is an extreme change in radiographic procedure,

as in the case of a tie-in or repair weld, where the radiograph may be

taken from outside the pipe by a double-wall technique or by the use of

a radioisotope source (Ir-192) instead of an x-ray generator. Where

this was suspected, the density data were not considered in Fig. 53 or

the subsequent calculations.

Since these film contrast variations are an important consideration

in establishing a relationship between defect depth (or steel thickness)

and film density changes, the information is now presented in an additional

form. Figure 57 shows the contrast data plotted as the ratio D 2 /D x rather than the difference. This ratio permits a calculation of the

effective x-ray linear absorption coefficient without having to use an

empirically determined average for a given change in metal thickness.

The y value determined in this way is within 7% of the value obtained from

the averages of the contrast data on Fig. 53.

These film contrast variations would be expected to most influence the

RI measurements since these were taken by observing only the weld-defect

area of the radiograph. The data on natural defects presented in Table 16a

showed that the RI measurements could be related to defect height measurements (obtained from laboratory radiographs) to within a 2a limit of a factor of two. Further support for the RI measurements is presented in

Table 16c, in which comparisons of SWRI and RI measurements on the same

defects are made. These films cover the complete range of contrast values,

yet the RI depth measurements are consistently (except in one case) greater

than those of SWRI.

no
TABLE 16C. Comparison of Corrected SWRI Densitometer and

RI Data on Specific Defects

SWRI Apparent

Wei d

No

Defect

Type

Cal cul ated

R Value

Corrected

Val ue* RI** SWRI^

100 )

Contrast

AD

81430 IF

(in.)

0.014

(in.

)

0.021

(in.)

0.055 262% 0.89

81153 GP 0.045 0.050 0.052 104% 0.49

81209 GP 0.007 0.008 0.050 625% 1.16

81316 GP 0.061 0.067 0.107 170% 0.90

81338R GP 0.024 0.027 0.076 292% 1.01

81434R GP 0.006 0.007 0.027 385% 1.17

81441

R

GP 0.010 0.011 0.033 118% 1.21

81447 GP 0.036 0.040 0.069 168% 1.04

81670 GP 0.018 0.020 0.060 300% 0.97

81674 GP 0.006 0.007 0.027 386% 1.39

W25TRX GP 0.046 0.051 0.086 162% 0.98

W39TRX GP 0.013 0.015 0.074 493% 1.07

81397 ES 0.033 0.036 0.040 111% 0.73

81 710 ES 0.064 0.070 0.046 -152% 0.94

81319 HB 0.008 0.009 0.044 489% 0.87

81643 HB 0.023 0.025 0.043 172% 0.77

IF - incomplete fusion

G? - gas pocket

ES - elongated slag

HB - hollow bead

* corrected as per P. 90

^corrected as per Fig. 48

110 a

LD

0

J


 o

LD

rm

c*

LD

m

LX)

ru

cn

o

o

m

u*)

rru

o

LD

ru

LD

ru

ru

Q

>N

m

cn

'a/

2

a oijcy Xjisusa

1

* *

Film

density

ratio

vs.

base

density

for

the

Alyeska

films
Concl usions

The SWRI methods have been discussed to some extent here and in

more detail in Section 4B. The densitometer method is capable of providing

reasonable accuracy but depends upon accurate film density measurements.

The aperture of the densitometer is an important consideration. For the

aperture used for these defect depth measurements, a defect width less

than 0.050 in. will lead to an inaccurate defect density measurement.

This factor could be improved by the use of a smaller aperture, as in a scanning microdensi tometer, for example. The weld bead introduces

errors in the densitometer measurement due to the undulating nature of

its surface. Estimates of these potential errors have been given in

Table 16b.

The SWRI visual comparison method, used because defect widths were

too small to measure with the densitometer, is difficult to assess and

no attempt has been made.

Large variations in film contrast have been observed for Alyeska field

radiographs. Because records were not available regarding necessary radiographic

information, such as x-ray energy and film processing parameters, it was

not possible to determine the reason for this observed variation in film

contrast. If the variations were caused by x-ray energy or non-optimum

film processing, the accuracy of the RI method could be influenced. If

the contrast changes resulted from such things as inaccurate densitometer

readings or changes in shim thickness, the accuracy of the RI method may

not be affected. Sufficient information for a rigorous assessment is

not available. One would normally state that a film whose contrast

falls within a designated predicted range of contrast could be used for

a depth determination (with additional variability considered). However,

the many combinations of variables that could place a film contrast in

an acceptable range have to be considered. For example, it is possible

that a thick shim could appear as acceptable contrast if it were radiographed at high x-ray energy or the film processed in a non-optimum

manner.

The RI method is potentially capable of good accuracy because the eye

can match shades of grey very well. Although the observed variation in

film contrast should have most influence on the RI method, there are results

that show the RI measurements may be useful. Limited data are available

in which the RI measurements are compared to the SWRI densitometer measurements. These comparisons are instructive because, (1) the measurements

were taken by two different methods and, (2) the SWRI measurements should

not be influenced by film contrast variations. These comparisons show that

the RI depth measurements (modified by the bias curve in Fig. 48).

are larger (except in one case) than the SWRI measurements (modified

according to the SWRI calibration data, p. 90). These comparisons cover

the full range of film contrast observed. Additional data in support of

the RI measurements is given by the measurements of natural weld defects

(Table 15a) from Alyeska field radiographs. It was found that the RI

depth measurements of these natural defects (in terms of height rather

than depth) showed correlation with laboratory radiographic measurements;

the one a limit was 47". Since the RI measurements were used for all

the weld defects under consideration, DoT may wish to expand these

comparisons in order to set limits for the RI measurements.

112
The reader should recall that the original intent of the field

radiographs was to meet the requirements of the API 1104 code. This

requires that weld defects be detectable on the radiograph and that

lengths of defects be capaole of measurement. The variation in film

contrast observed in these tests is not inconsistent with those requirements. There was never the intent to use these films for defect depth

measurements. If that were the case, much better control of radiographic

variables would have been used.

112A

n0 . REMARKS ON PIPELINE DESIGN AND STRESS ANALYSIS

5.1 Introduction

In Sections 3 and 4, it has been established that two major items of

information on girth welds with defects are pertinent to this investigation.

One item concerns the nature and size of an identifiable weld defect, and

the other an estimate of the fracture toughness and other material properties

of the weldment containing a known defect. The former requires a state-of-art

investigation in nondestructive evaluation, and the latter leads to an

experimental and analysis program in fracture mechanics. There is, however,

a third dimension to this technical problem, i.e., the state of stress at any

point along the girth weld under investigation during the 30-year design life

of the pipeline. For the fracture analysis to be useful as a tool for evaluating

a weld with known defects, it is necessary to know not only the maximum stresses

in both the hoop and the axial directions, but also the cyclic nature of these

stresses due to varying operating and abnormal loading conditions. Maximum

stresses are needed in order to identify the "worst loading condition" under

which each known defect is required to be stable, and cyclic stress history,

also known as the stress spectrum, is required in order to estimate the potential

growth of the defect during the pipeline's design life.

The purpose here is not to assess the stress data which Alyeska prepared

for the pipeline. Rather, what is presented in this chapter is motivated

by a precautionary concern over the uncertainties which are always present

in any stress calculations. Since OPSO guidelines [ref. G-4] contain

specific references to safety factors on defect sizes and fatigue life

as well as minimum or representative material properties, it may be argued

that further discussion of adjustment factors for stress, geometric

constants for modeling, and materials parameters is unnecessary. As will become

clear later in this section and in section 7C, however, we have reason to believe

that the physics of the macroscopic behavior of a buried pipeline is not as quantitative and well -understood as that of an above-ground pipeline, partly

because of the uncertain nature of the soil-pipe interaction. Furthermore,

the present theory of fracture mechanics is based largely on the limiting

states of plane stress or plane strain; mathematical modeling of the failure

behavior of an elastic-plastic material is still under intensive development,

so that today the choice of a particular model and an appropriate set of values

fcr its constants are primarily a matter of judgment. This viewpoint will be

substantiated both in this chapter and in section 7C.

5. 2 T'ne Physics of the Stress Problem of a Buried Pipeline

A buried pipeline behaves somewhat differently from one above ground, simply

because the steel pipe interacts with the comDacted granular material backfill

through frictional forces not unlike these experienced by steel piling of the

friction type. Full understanding of the behavior of the steel pipe when buried

underground would require a mathematical model based on tne material characteristics of the oipe steel and the surrounding backfill as well as data on the

geometry of the pipe alignment, and eatable of handling loading cases such as

113
internal pressure, temperature change, earthquake, and differential settlement.

Such a formulation, while technically feasible, is beyond the scope of this

investigation. The purpose of tnis section is to describe qualitatively how the

pipeline might be predicted to behave if we had such a mathematical model.

This knowledge will then be used to interpret Alyeska's computational results,

wnich were based on a model simplified in ways consistent with current engineering

practice. From time to time, elementary calculations are inserted to enhance the

qual i tati ve results of this section, but the reader is advised to regard these

calculations as purely illustrative because they are not intended either to

check or to supplement Alyeska's pipeline design.

For illustrative purposes let us examine a collection of numbers characterizing a typical section of the Alaska pipeline. Let us consider a 2,000-ft

long section of 48 in. outside diameter (O.D.), 0.462 in. wall thickness, and

horizontally buried pipe with a 4- ft depth of cover from top of pipe to final

graded surface. At the two ends of the straight section, a horizontal 45° bend

and a vertical 45° bend are designed to change the pipeline alignment. As a grade X-65 material, the pipe steel is designed for a specified minimum yield

( SMY ) of 65 kips per sq. in. (ksi). For subsequent calculations, we need the

cross-sectional area of the pipe (A = 67 sq. in.), the plane moment of inertia

(I = 19,500 in 4 .), and the radius of gyration (r = 17 in.). The surface area

of contact between the steel pipe and the granular material backfill as specified

in Alyeska drawings (see, e.g. , Alyeska drawing No. AL-00-65 as listed in [H-ll]),

is 12.6 sq. ft. per lineal foot. For convenience, numerical quantities are

rounded off to emphasize the qualitative aspect of this discussion.

Now assume that this section of pipe is undergoing a hydro-test to 95% of

SMY for the maximum stress in the pipe steel. Ignoring the two ends, the state

of strain throughout the middle section of the 2,000-ft pipe is a very simple

one; namely, zero along the pipeline axis, and uniformly distributed hoop strain

along the perimeter of the pipe due to internal pressure. Assuming a Young's

modulus of 30,000 ksi, the hydro-test imposes a hoop stress of 0.95 x 65 = 61.8

ksi, and a hoop strain of 61.8/30,000 = 0.21%. It is a well-known physical

fact that when a material is stretched along one dimension it tends to contract

in the other two dimensions unless restrained from doing so. For steel, the

constraint of zero longitudinal strain introduces a tensile stress along the

axial direction equal to 0.3 x 61.8 = 18.5 ksi, where 0.3 is the Poisson ratio

(assumed). This induces a tensile force along the pipe equal to 18.5 x 67 = 1 ,240

kips which could be provided by the frictional force between a section of the

pipe away from the midsection and the backfill material. To estimate that

frictional force, Seelye [2, p. 3-29] allows a 35 lb/ft^ equivalent fluid

pressure for some common backfill not unlike those encountered in Alaska. A

4- ft cover implies a 6 - ft average depth of backfill around a 48 in. 0.0. pipe

and an average lateral pressure between backfill and pipe equal to 35 x 6

= 210 lb/ft^. Assuming a coefficient of friction between 0.3 and 1.0, tne frictional force can be estimated as between 0.9 and 2.7 kips per lineal foot.

This range of values checks very well with values for similar problems recently

reported by Kennedy, Chow, and Williamson [1]. To be conservative in estimating

the frictional force, we choose to work with the estimate of 2.7 and calculate

the length of pipe to develop the tensile force of 1 ,240 kips to be about

1 ,240/2.7 or, say, 500 ft. This means that for about 500 ft. along each end of

this pipe, the longitudinal tensile stress decreases from 13.5 ksi at the ends
of the 1 , 000- ft mid-section to zero at the two bends, with an estimated

elongation equal to 0.3 x 0.21% x (1/2) x 500 x 12 = 1.9 inches at each bend.

Based on a formula given on p. 6-78 of reference [2], the 45°-bend introduces

a longitudinal tensile force of 630 kips for an internal pressure of 1,210 psi

which corresponds to the assumed hoop stress of 61.8 ksi. This tensile force

develops a longitudinal tensile stress of 630/67 = 9.4 ksi in the pipe, and an estimated bending stress of 10 ksi at the outer extreme fiber of the bend due

to a bearing resistance of 20 kips/sq. ft. in the backfill, over, perhaps, a pipe length^of about 20 ft. We have used the bending stiffness

(Z = 810 i n

3 ) of the pipe in this calculation. We now conclude that a safe

estimate of the longitudinal tensile stress everywhere along this pipe is

about 20 ksi during a hydro-test at the tie-in temperature. This checks very

well with Alyeska's estimate for a different, but comparably severe loading

condition for the same pipe grade and wall thickness, where the longitudinal

stress due to pressure and live and dead loads was estimated to be 19.8 ksi

(see Table 10 of Part 3, reference [H- 1 ] )

.

Similar reasoning can be applied to achieve an understanding of the

behavior of the 2,000-ft pipe if it was laid at 35°F (tie-in temperature )

,

and contains oil at 145°F. The net temperature strain would be 110 x 6.5 x 10

- 6 = 0.07%, where the coefficient of expansion of steel is assumed to be

0.0000065 per °F. With the operating pressure limited to 76% of the hydrotest pressure (i.e., from 0.95 SMY to 0.72 SMY), the induced tensile stress

(axial) would have been about 15 ksi. The temperature stress (axial) for

zero longitudinal strain would be 0.07% x 30,000 ksi = 21 ksi of a compressive

nature. Hence the mid-section of the pipe experiences a longitudinal compression of 21 -16=6 ksi. The net compressive force equals 6 x 67 = 400 kips.

By the Euler buckling criterion, if the backfill provided no resistance, the

critical length of the pipe would equal

(tt 2 X 30,000 x 19,500 / 400)

1/2

= 3,800 in. or about 320 ft.

Clearly, the backfill pressure provides an excellent deterrent to the problem of

instability due to temperature-induced compressive loads on a long buried pipe.

One can then go through a similar illustrative calculation for the two ends of

the pipe to check the stress behavior as influenced by the temperature change

and the backfill-steel frictional force estimated in the last paragraph. The

result is that if a girth weld is located near a bend (6 ft. away from the

point of tangency), there is a strong likelihood that a longitudinal tensile

stress of the order of 10 ksi t 5 ksi may be encountered by some portion of the

weld under normal operating conditions. This estimate is higher than the

comparable Alyeska estimate (3.2 ksi), and may be explained by the difference

in the two approaches underlying the estimates.

5. 3 Variability in Stress Spectra for Girth Welds

The 0PS0 requirements specify that proposed alternative-allowable welddefect sizes and proposed alternative-allowable arc-burn sizes must be supported

by fracture mechanics analyses using the worst-cast fatigue-stress spectrum. An

axial -stress spectrum is used for girth-weld flaws and a hoop-stress spectrum is

used for arc burns. A detailed discussion of the fatigue crack propagation

problem appears in section 7(A). The purpose of this section is to show the

115
variability of stress spectra by comparing Alyeska's worst case spectrum with

two of the weld stress spectra provided to NBS by Alyeska. Once again, the

intent is to show the qualitative difference between Fig. 58 (worst case spectrum)

and Fig. 59 (weld no. 42R, sta. 0192+32, mile post 731, alignment sheet 0013),

or Fig. 60 (Alyeska synthesized spectrum no. 23A, sta. 1111+11, mile post 78,

alignment sheet 126). In all three cases, the stress ranges denote longitudinal

stress differences due to all the pressure cycles, thermal cycles, and earthquake loadings anticipated during the 30-year life-time of the pipeline. Information for Fig. 58 (worst case) came from Part 4, reference [H-l], and that for

Figs. 59 and 60 from references [H-6] and [H-7] . For the NBS fatigue analysis

given in section 7(A), the stress spectrum corresponding to Fig. 60 has been

adopted. If we assume for this qualitative purpose, that the fatigue damage,

when measured by crack growth due to cyclic stresses, is proportional linearly

to the number of cycles and varies approximately as the third power of the stress

range (AK) based on the following equation (section 7(A)):

da

dN

* c (AK)

3

,

then it is possible to estimate that the worst-case spectrum differs from

either of the two specific weld spectra (Figs. 59 and 60) by a factor of

1.75 to 1.0.

5.4 Concluding Remarks

The preceding qualitative analyses of the physics of a buried pipeline

(section 5.2) and of the spectra of cyclic stresses during the 30-year life-time

of the pipeline illustrate the variability one would confront in choosing a set of stresses to represent the states of stress in all welds under investigation. The variability is particularly hard to estimate when the pipeline is

underground. For arc burn problems, there is also a definite cause for concern

in the uncertainty of any stress estimate based on a single nominal thickness

for a pipe wall which, according to Alyeska material specification [H-2], has an

allowable variability of +10% and -5%. Since the stress calculation is only

valid to the accuracy of the hydraulic analysis, uncertainties due to surge

estimate (see reference [3]) and fluid properties should also be accounted for

in fixing a stress level for fracture analysis. These uncertainties on

applied stress will be further discussed in section 7(C) under the title:

Sensitivity Analysis and the Use of Safety Factors.

5. 5 References

[1] Kennedy, R. P., Chow, A. W., and Williamson, R. A., "Fault Movement Effects

on Buried Oil Pipeline," a technical paper presented at 31st Annual

Petroleum Mechanical Engineering and Pressure Vessel and Piping

Conference, Mexico City, Sept. 1976. Preprint available through

first author's office at Engineering Decision Analysis Company, Inc.,

2400 Michelson Drive, Irvine, California 92715.

[2] Seelye, E. E., Data Book for Civil Engineers , Vol . 1, Design, 2nd ed.

Wiley (1951).

[3] Parmakian, J., Water hammer Analysis , pp. 93-100. Dover (1963).

116
NO,

of

Stress

Excursions

NO.

of

Stress

Excursions

NO.

of

Stress

Excursions N8

*<: \

10

2 1

ALYESKA

WORST- CASE SPECTRUM

as shown in Part 3 [H-1]

a

ui mS3 aLr«

(% X

M2

2*

ia t

i

12

I

WELD #42R

STRESS SPECTRUM

as derived from Alyeska document [H-6]

as

Ui

ea

ui

fi4 rvt

•3

33UiX

25 4

;

2M

is

;

J

fa

S3S3

I

'

Alyeska Synthesized Spectrum No. 23A as

Communicated to NBS in August 1976 [H-7]

STRESS RANGE AK (ksi)

ill
6. ASSESSMENT OF MATERIALS PROPERTY DATA

The determination of allowable defect sizes depends on the availability

of reliable mechanical properties data for the weld metal and base

metal. The properties required for the allowable defect size determination

are: elastic modulus, Poisson's ratio, yield strength, fracture toughness,

fatigue-crack growth rate, and environmentally assisted crack growth

rate due to stress-corrosion, hydrogen cracking, and corrosion-fatigue.

For this program the required mechanical properties data were obtained

from measurements made by the National Bureau of Standards (NBS), measurements

made for the Alyeska Pipeline Service Company by the Welding Institute

(WI) and the Cranfield Institute of Technology ( C IT ) , and from information

reported in published technical literature.

Standardized and accepted methods of measurement have been established

to determine some but not all all of these required mechanical properties.

Therefore, the reliability of all data on material properties used in

determination of allowable defect sizes will be assessed in this chapter.

This discussion will include an evaluation of the selection of materials

used for testing, and an evaluation of the test methods used and the

testing parameters chosen.

Selection of Test Samples

All weldments used in the NBS, WI, and CIT test programs were

obtained from actual field welds made on the Trans-Alaskan pipeline.

Therefore, these weldments should be representative of weldments made

according to the prescribed welding procedures and should be typical of

production welds. This use of production welds is particularly desirable

whenever possible because they are more representative of actual welds

than weldments made under "laboratory" conditions using the same procedure.

The NBS testing program was conducted on a group of four weldments

consisting of one made with welding process 104A, one made with process

102A-1, and two made with process 104A-1. The WI-CIT testing program

used 3 weldments made during 1975 and 3 weldments made during 1976.

Each group consisted of one weldment made by each of the three welding

processes 102A-1, 104A-1, and 302A-1.

Although extensive testing was done on each of the weldments available,

the small number of weldments available (10 weldments representing 3

welding procedures) compared with a total of nearly 60,000 girth welds

made on the Trans-Alaska pipeline, does not permit an accurate assessment

of the weld properties to be made on a valid statistical basis. However,

well defined welding procedures and materials were specified by Alyeska

to be used for all production welds. If these welding procedures were

followed and a high level of quality control was maintained during the

welding, it can be expected that the very limited number of weldments

tested should yield material properties data that are representative of

the production welds in the pipeline. The lack of a sufficient sample

118
size to establish a suitable statistical base for assessing the material

prooerties data is one reason that safety factors are required in the

allowable defect size analysis to account for these uncertainties.

^ensile and Elastic Properties

The elastic constants of the pipeline base-metal measured at NBS

were determined with established and accepted measurement procedures.

Elastic constants of the weld metal were not determined but base-metal

properties are considered to be representati ve of the weld metal elastic

properties. All results are consistent with accepted properties for

steels of the type used in the pipeline.

The tensile properties of the base metal and each of the two weld

metals were determined at NBS using well established and accepted testing

procedures. The minimum yield strength measured at 68° F of all weldmetal samples was 59 ksi for the weld made with process 102A-1 and

55 ksi for the weld made with process 104A-1. The average yield strength

of these welds were 60 ksi and 63 ksi, respectively . These values are

lower than expected for welds made with the welding electrodes specified

as AWS E7010-G for the 102A-1 process and AWS E8010-G for the 104A-1

process. Welds made with these electrodes should have minimum yield

strengths of 70 ksi and 80 ksi, respectively . The low yield strengths

determined in these tests may be caused by the small specimen size

chosen (0.14 in. diameter). Weldments made specifically for establishing

the weld metal yield strength allow larger sized specimens to be obtained

on the weld metal yield strength.

Fracture Toughness

The fracture toughness of the metal is the key parameter through

which the fracture mechanics analysis determines the allowable defect

sizes for specified levels of applied stress. For relatively brittle

metals that fracture under conditions where linear elastic fracture

mechanics analysis applies, the fracture toughness specified as the

critical plane strain fracture toughness, K- , can be measured by the

accepted and standardized method described in ASTM Standard E399-72.

From this fracture toughness value for the specified material, and the

specified applied stress, the allowable defect size can be determined

for any structure that behaves according to the assumptions of linear

elastic fracture mechanics.

For more ductile metals that do not obey linear elastic fracture

mechanics assumptions in either the structure of interest or in laboratory

test specimens, a single standardized method of determining the fracture

tougnness has not been agreed upon. The measured value of fracture

toughness is very sensitive to the test specimen size and geometry ana

to the specific choice of when fracture initiates. There are presently

four test methods under development to measure the fracture toughness of

ductile metals that fracture after undergoing substantial amounts of

plasticity. Fracture toughness data using each of these four test

.methods was obtained on the girth welds of interest and was used to

assess the welds and develop allowable defect sizes. The test methods

considered here are: standard Charpy-V-notch impact tests, i nstrumentedprecracked Charpy-V-notch tests, J-integral tests, and crack-openingdisplacement (COD) tests.

1.19
Standard Charpy-V-Notch Impact Test Results The Welding Institute and

Cranfield Institute of Technology have done extensive testing for Alyeska

of standard Charpy-V-notch test specimens taken from welds made during

both 1975 and 1976 using welding processes 102A-1, 104A-1, and 302A-1.

This test program tested specimens with notches located in 8 different

positions within the weld metal or HAZ and over a temperature range

from -40°F to +32°F. These results were submitted to 00T as part 3 of

the supporting information for the waiver request. NBS tested similar

Charpy-V-notch specimens taken from welds made during 1975 using welding

procedure 102A-1 and 104A-1. Only one notch position (position number

1, shown in Figure 35) was used in these tests and a full transition

temperature curve from -40°F to 100°F was obtained. These results are

summarized in Section 31 and in Appendix Q.

The energy absorbed and percent shear fracture determined in the

NBS and the Alyeska tests are similar over the temperature range where

they are comparable. The Alyeska data did not demonstrate an effect of

changes in notch position on the Charpy-V-notch results. The standard

CVN tests conducted by NBS showed that "upper shelf" behavior as demonstrated

by 1 00% shear fracture was not obtained even at 100°F for either of the

weldments tested. The extensive variability in the test results from

both sets of data, due to the heterogeneity of the weld metal and the

fact that most data were obtained within the transition temperature

range, prohibits more specific conclusions about the effect of welding

procedure or notch orientation to be made from the results of these

Charpy-V-notch tests.

Although a fracture toughness value that can be used directly to

determine allowable defect sizes cannot be obtained from standard CharpyV-notch tests, the results of these tests are useful for comparison of

relative toughness levels for weldments tested at various laboratories

.

This is done by assessing the effect of temperature on the ductile to

brittle fracture transition and to determine the effect of impact loading

on the weld metal. The lowest Charpy toughness found in any of the

welds was in excess of 40 ft- lb (average of 3 specimens) at 10°F and in

excess of 25 ft- 1 b at -40°F.

Although the standard Charpy-V-notch test cannot be used to measure

the fracture toughness, empirical correlations have been developed that

permit the estimation of fracture toughness 1 or the allowable failure

stress for piping containing defects 2 when more precise measures of

fracture toughness are not available. Since more reliable tests, such

as COD, instrumented precracked Charpy-V-notch impact tests, and Jintegral, were conducted in this investigation, these empirical

correlations to estimate fracture toughness were not used.

1. R. A. Wullaert, D. R. Ireland, A. S. Tetelman, "Use of Precracked

Charpy Specimen in Fracture Toughness Testing," Fracture Prevention and

Control, ASM, 1974, pg 255-282.

2. W. A. Maxey, "Fracture Initiation, Prpagation, and Arrest," 5th

Symposium on Line Pipe Research, American Gas Assoc., Cat. No. L30174,

Nov. 1974.

120
Instrumented Precracked Charpy-V-Notch Tests Instrumented precracked

Charpy-V-notch tests were conducted over a range of temperature from

-40°F to 100°F on specimens taken from welds made using welding procedures

1024-1 and 104A-1 to determine a quantitative measure of fracture roughness

to establish a lower bound for the determination of maximum allowable

defect sizes under the most unfavorable conditions of high rates of

loading. A measure of the dynamic fracture toughness, K. , or K, ., is

determined from these tests.

Since no "static" fracture toughness, K, , tests were conducted

using i nstrumented-precracked Charpy-V-notch specimens , the assessment

of the effect of dynamic loading on the fracture toughness of the welds

must make use of static fracture toughness measurements made by the Jintegral procedure at the Welding Institute 3 and at NBS 4

. The measured

J-integral values were converted to stress-intensity, K, , using the

equation:

1/2

where E is Young's modulus. These static K. values are plotted on

Figure 61 for comparison with the dynamic fracture toughness values, K, , or K y

I , determined from the instrumented precracked Charpy-V-notch

specimens

.

The results in Figure 61 show that up to 32°F, the dynamic fracture

toughness has a lower bound that is substantially below the static

fracture toughness and, in fact, the dynamic specimens obey the behavior

associated with linear elastic fracture and do not exhibit plasticity

prior to fracture. A minimum dynamic fracture toughness of K, . = 55 ksi

vin. was measured at 32°F. This low fracture toughness suggests that

only very small defects would be permitted if the pipe welds experienced

loading rates as high as experienced in the Charpy-V- notch te sts.

However, loading rates this high, above 2 x 10 5 ksi /in/ sec. are not

normally experienced in large, compliance structures.

For temperatures above 32°F, dynamic loading does not appear to

degrade the fracture toughness of the weld metal. As shown in Figure 61,

at 63 0p

, the static fracture toughness even appears to be less than the

dynamic fracture toughness, K,.. However, this is an apparent rather

than real effect because for Rj , values above 110 ksi in, the dynamic

fracture toughness is affected oy the small size of the Charpy specimens

and an artificially high value is measured. The static, J, tests were

determined with full wall thickness larger specimens and ar§, therefore,

not artificially affected by specimen size effect the correct values of

these oarameters.

3. J. 0. Harrison, "COD and Charpy-V-Notch Impact Tests on Three Pipeline

Butt Welds Made in 1975," The Welding Institute Report LD 22062/5, July

1976, Ref. [H-l] in appendix H of this report.

4. Section 3E, this report.

121
Fig. 6! COMPARISON OF DYNAMIC AND STATIC FRACTURE TOUGHNESS

0 spccihcm iDrHnrtCflrroN

o

122
J-intecral and Crack-Open ing-Di s pi acement (COD) Tests.

The J-integral test method under development by ASTM requires

several specimens to be tested to obtain a single critical Jj

c value

that represents the critical fracture toughness of the metal. This

method is the most widely accepted, but not yet standardized, method for

testing in this country. The J^

c value can be used directly to obtain

allowable defect sizes when the stress in a structure is known.

A simplified, single specimen method was used by the Welding Institute 3 to estimate J. from three-point bend specimens. This simplified method, using

maximum loads dan only be used for specimens that fracture with limited

plasticity so that ductile crack propagation does not occur prior to fracture.

Therefore, the Welding Institute J-integral test results are only obtained at

test temperatures of 10°F and below. Maximum load estimates tend to overestimate J, as crack initiation may begin below the peak of the load-deflection

curve.

_ A_summary of the results of the Welding Institute J-integral tests is

shown in Table 18. As with all the weld metals tested, there is a large amount

o r variability in the Jj c values obtained. However, since only specimens which

fractured without ductile crack propagation could be used to obtain these Jr

values, the results reported in Table 18 are used to establish a lower bound

level of J T for these weld metals.

Ic

The results of the NBS tests to measure Jj c (Section 3E) are also

summarized in Table 18. The NBS tests were conducted with larger specimens

than the tests of the Welding Institute and used the multiple specimen

technique described in Section 3E of this report. This procedure is

considered to be the most accurate presently developed to determine Jj

c because it clearly defines the initiation of fracture even if ductile

crack propagation occurs. However, the NBS values of Jj c obtained at

4°F (457 and 628 in-lb/in 2

) are consistent, within the observed variability

of the tests, with the Welding Institute test results of 283 and 503 inlb/in2 obtained at the nearest comparable temperature of 0°F. Therefore,

the J-integral test procedure used by the Welding Institute seems to be

reliable for determining Jj c when no ductile crack propagation occurs

prior to fracture.

The largest amount of data on the fracture toughness of the weld

metal was obtained using the crack opening displacement (COD) test

procedure originally developed by Burdekin and Stone 5 and subsequently

adopted as a standard test method by the British Standard Institute.

(Draft for Development #19, 1972.) The COD results are the only fracture

toughness results actually used in the calculation of allowable defect

sizes in this report.

The results of COD testing by the Welding Institute and by Cranfield

Institute of Technology 5 included tests on welds made by processes 102A1, 1C4A-1, 302A-1 and were conducted for 8 different notch orientations

over the temperature range from -40°F to 32°F. The heterogenei ty of the

weld metal resulted in a large variability in the COD results so that

the COD values determined varied by as much as a factor of 6. The

minimum value of COD of 4 mils (0.004 in.) at +10°F was chosen by Alyeska

as the lower bound value for the allowable defect size calculations.

5. F. M.

Approach

Anal yses

,

-

Burdekin, D. E. W. Stone, "The Crack Opening

to Fracture Mechanics in Yielding Materials,"

i (2)

,

1566, pp. 145-153. M3

Displacement

Journal of Strain
Table 18. Fracture Toughness Results from J-Integral and COD Tests

Weld Test

J

Ic

6 c, COD (in.

)

Process Temp. (°F) (in-lb/in 2

) Meas

.

1 Calc. 1

NBS Tests -using Compact Tension Specimens

102A-1 68 571 .0048

- 4 628 -- .0046

104A-1 68 457 - - .0036

- 4 457 -- .0034

Welding Institute Tests-using Bend Specimens

1 02A-1 32 _ - .017 ~ — -- .015 - - 10 -- .009 —

... .004 - - 883 .007 -- 1443 .010

1318 .011 --

0 283 .002 - - 503 .004

- 20 291 .003 -- 895 .008 -- 2830 .019 —

-40 169 .0010

222 .0015 —

652 .006 —

104A-1 32 • <as .012 •• — -- .007 =1

10 - - .013 — — — .012 —

0 -- .015 —

-- .014 --

- 20 • - .011 - .. 554 .006 —

- 40 657 .006 — — 652 .006 —

401 .004 • *

1 . Two lowest measured values at designated temperature reported by Welding

Insti tute in report submitted with waiver request.

2. Calculated by conversion: 5 =

c 2 ay

124
It is possible to compare fracture toughness values determined with

COD tests with those obtained with J-integral tests. However, this

requires a conversion based on an empirically developed relationship of

the form:

a = c mo

y

as given in Section 3E of this report. The value of m can range from

1.0 to 2.1 so that in converting from Jj c the resulting value of COD is

not precisely defined. Results of tests by the Welding Institute 3 suggests that for large specimens m = 2.1 is the best value while m = 1.6 is suitable for smaller specimens. Sumpter and Turner 6 show that m

= 2 for three-point bend geometry specimens, such as used by the Welding

Institute for J-integral tests, when substantial yielding occurs prior

to fracture. For specimen geometries other than three-point bend

specimens, the factor m, may be different from 2.0, but has not been as clearly defined by experiment or by analysis such as the use of finite

element calculations. This uncertainty in the value of m makes it

difficult to compare fracture toughness values obtained using the COD

test method with those made using J-integral procedures. However, to be

conservative, whenever converting measured values of Jj c to COD values,

the value of m is chosen as equal to 2.0.

To compare the results of NBS J-integral tests with the COD measurements

made by the Welding Institute and Cranfield Institute of Technology, the

J-integral values were converted to COD values using a value of m = 2.0

chosen because of the large size and constraint present in the NBS

specimens. The lowest values of COD measured for each weld at each test

temperature are plotted in Fig. 62 along with the COD results calculated

from the NBS J-integral test and a lower bound COD curve as a function

of temperature is drawn. A minimum critical COD value of 3.4 mils

(.0034 -m.) at -4 3 F is obtained by converting the Jj c value measured at

NBS to COD using m = 2.0. If the m value is chosen as 1.6 instead of

2.0, the minimum COD is 4.0 mils at -4°F.

The COD results from Alyeska appear to be higher at 32°F than at

+1 0°F . However this increase was not statistically significant. The

comparable results from the NBS J-integral tests at 68°F suggest that

the toughness may not increase, but may still be less than the equivalent

COD value of 4 mils as shown in Fig. 62. It is possible that the COD

measurement made with relatively small specimens (approximately 1/2 by

1/2 inch) exhibit an artifically elevated COD value when extensive

plasticity occurs prior to fracture. This would explain why the estimated

COD values obtained from the NBS J-integral tests at 68°F are considerably

below the Welding Institute COD values determined at 32°F. This effect

of specimen size on the minimum value of CCD is further substantiates by

tne result of one test reDorted by the Welding Institute 2

, where a COD

value of 2 mils was measured at 0°F when a larger (approximately 1/2 by

1 inch) COD specimen was used. This low value of CCD was attributed to

the high degree of plane strain constraint present in the larger specimen.

6. J. D. G. Sumpter, C. E. Turner,

or J c,

,, A STM ST C 501 ,

1 976, p. 1.

"Methods for Laboratory Determination

125
COD

(in)

(Fig. 62 ) FRACTURE TOUGHNESS COD RESULTS

X 102 A —I WELD, J-INTEGRAL TESTS BY NBS

O I04A-I WELD, J-INTEGRAL TESTS BY NBS • 102 A-l WELD, COD TESTS BY, WI

104 A -I WELD, COD TESTS BY, WI

12S
These results suggest that the value of 4 mils used in the fracture

anal/sis is a reasonable value. However, at +10°F, this cannot be

considered a lower bound value if the defect is oriented in a direction

such that the constraint of the structure increases the amount of planestrain at the defect location.

Fatigue Crack Growth Rate Data

The fatigue crack growth rate data used in the calculations of

allowable defect sizes by NBS was obtained from data published by

Maddox" for fatigue crack propogation in carbon-manganese steel weld

metal and heat-affected zones. In confirmatory testing done by NBS (see

Section 3F), the measured fatigue crack growth rates did not exceed the

upper bound of the Maddox data. However, in order to determine the

fatigue crack growth rate for a defect growing by fatigue in the throughthe-thi ckness direction, the NBS tests were conducted using relatively

small specimens (approximately 1/2 inch square in cross section) so that

as the stress intensity range, AK, increased, the range of validity for

using linear .elastic fracture mechanics to calculate AK was exceeded.

Since all crack growth rate data, da/dN, was determined from compliance

calculations, the crack growth rate is overestimated at higher aK

values due to the nonlinear load-deflection test curve. This has the

effect of increasing the slope of the da/dN versus AK curve above those

measured using larger specimens, as seen in Fig. 23 of Section 3F.

However, even when the crack growth rate is artifically elevated due to

these small specimens, the maximum crack growth does not exceed the

maximum crack growth rate measured by Maddox.

Summary of the Assessment of Materials Properties Data

(1) A fracture toughness value equivalent to a COD value of 4 mils (for

planar defect) and 7 mils (for arc burns) at +10°F as used in calculating the allowable defect size curves are reasonable values.

(2) Tests with larger sized specimens indicate that a fracture

toughness equivalent to a COD value slightly less than 4 mils

may apply if the weld metal is sufficiently constrained so

that a condition of plane strain can exist. Such a high

degree of constraint is unlikely to occur in the thicknesses

of metal used in the pipeline.

(3) The dynamic fracture toughness of the weld metal is sufficiently

low that elastic fracture could occur up to temperatures as high as 32°F. However, it is not expected that the loading

rates in the pipeline will reach the levels seen in the dynamic

fracture testing so that this reduced toughness is not appropriate

for evaluating allowable defect sizes in the pipeline. The

only other known source of dynamic loading is associated with

a fast running crack after "pop-in" such as from an arc burn.

This problem has not been evaluated in this report but is not

expected to be a significant problem.

(A) The fatigue crack growth rates used in the calculation of

allowable defect size curves are reasonable upper bound values

for the weld metals of interest.

127
(5) The nominal, rather than minimum, yield strength value of 60

ksi was used in the allowable defect size calculations.

7 S. J. Maddox, Welding Journal, pg. 40 1 S-409S , Sept. 1974.

123
7. FRACTURE ANALYSES

7A. FATIGUE CRACK PROPAGATION

7B. ALLOWABLE DEFECT SIZE ANALYSIS

7C. SENSITIVITY ANALYSIS AND THE USE OF SAFETY FACTORS
'
7A . FATIGUE-CRACK PROPAGATION

The OPSO requirements, Appendix C, stipulate that surface cracks

equal in size to twice the proposed allowable weld-defect or arc-burn

size (in both length and depth) must not grow in size such that stressing to the maximum-credible service stress could cause leakage. The

crack-growth analysis must account for the cyclic crack growth anticipated during 30 years of operation in the most deleterious service environments and temperatures.

The work reported in this section was undertaken to: (1) independently assess the fatigue-crack growth characteristics of weld defects

in the trans-Alaska pipeline, and (2) evaluate the adequacy of the

procedures and assumptions used by Alyeska in their assessment of

fatigue-crack growth.

Analysis Methods

The NBS fatigue-crack growth analyses were conducted using a computer program developed for the NASA Marshall Space Flight Center

J

The computer program is based on linear-elastic fracture mechanics and

consists primarily of an integration routine that calculates crack growth

from an initial defect size to a terminal size -- leakage in this case.

The analysis approach used in this investigation was to integrate the

Paris equation between the limits of initial flaw size, a., and a flaw

95% through the thickness, t:

N

. 95t

/

da

a

i C(AK)

n

(

1 )

where C and n are empirical constants

that describe the crack-growth behavior of the weld defects under

consideration, and AK is the stress intensity range calculated as

follows:

K(at surface) [1.12 + 0.11(2a/z]]<jv- ( 2 )

K(at maximum depth) = [1 + 0. 12(1 -a/£)"]cr/ ^

(M ,)

W u3C< (3)

AK = K/l -R

1. M. Creager, "MSFC Crack Growth Analysis Computer Program," NASA CR143923, September 1975.

130
Q

= 1 + 4.593UM) (5)

1 ' 65

- 0.212(^)

2

y

M back

= f( a /t> a/0 defined in computer program, a = crack depth, l = crack length, Ac = stress range, a = 0.2* offset yield strength, and

R = stress ratio = cr. /cr . y

min max

The integration of Eq. (1) proceeds incrementally through each

step in the fatigue spectrum and continues to repeat the fatigue

spectrum until the crack reaches 0 . 95 t . The output of the computer

program includes crack-length and crack-depth data at periodic intervals in the fatigue life. These data are plotted as a-vs-N curves,

where N is the number of service lifetimes, i.e. number of times the

spectrum has been applied.

Fatigue Data

The OPSO requirements state that: "A maximum fatigue crack

growth rate for the pipeline shall be established by documenting the

fatigue crack growth behavior of the weld metal and heat affected

zone that is representative of the pipeline welds and operating conditions and, in the case of arc burns, representative of the base metal.

The fatigue crack growth rate used in the fracture mechanics analyses

shall be the maximum fatigure growth rate multiplied by an assumed

safety factor of four."

The fatigue data used in Eq. (1) were obtained from the open « literature, proprietary sources, and the NBS test program. Maddox's^

data were used by both NBS and Alyeska for crack growth in a nonaggressive environment. To account for environmental enhancement,

NBS used a factor of 10 on Maddox's data and Alyeska used an upper

bound of the data obtained by Vosikovsky. 3 The Maddox data with a factor of ten provided a conservative upper bound to all the data

reviewed by NBS except the data reported by Vosikovsky. The NBS test

results on small-scale bend specimens (Section 3F) and

large-scale surface flaw specimens .(Section 3G) were reqarded

as sufficient evidence that Vosikovsky ‘s data were not applicable to

the trans-Alaska-pipel ine welds subjected to a representative current

density at a holiday in the corrosion-protection system.

The specific values for C and n in Eq. (1) used in the analyses

that follow are summarized below:

2. S. J. Maddox, Welding Journal, 4015-4095 (Sept. , 1974)

.

3. 0. Vosikovsky, Closed Loop, 3-12 (April, 1976).

131
Macdox: ^

= 4.94 x 10~^°UK)

3

* (factor)

non-aggressive environment factor = 4

aggressive environment factor = 40.

Vosikovsky: ^

* 1.56 x 10 (factor)

factor a 4.

Fatigue-Stress Spectra

The 0PS0 requirements specify that proposed alternative-allowable

weld-defect and proposed alternative-allowable arc-burn sizes must be

supported by fracture-mechanics analyses using the worst-case fatiguestress spectrum. An axial-stress spectrum is used for girth-weld flaws

and a hoop-stress spectrum Is used for arc burns, i.e. the stresses

perpendicular to the length dimension of the flaw govern the growth

behavior. The spectra employed by Alyeska and by NBS are listed in

Table 19.

The worst-case axial spectrum supplied by Alyeska was for

0.462-in thick X-60 pipe operating at 832 psi internal pressure.

Subsequently, a spectrum representative of 0.562-in thick X-65

pipe was provided by Alyeska. The differences In axial stresses were

considered negligible and the spectrum was not used in the analysis.

The axial spectrum Includes all the pressure cycles, thermal

cycles and earthquake loadings anticipated during the 30-year lifetime of the pipeline. These loadings are superimposed on a pipebending stress (tensile In the lower half of the pipe and compressive

in the upper half) due to soil settlement that occurs during the first

five years of operation. The axial spectrum was given to NBS*1n terms

of a computer output that included 383 stress excursions in a sequence

representing 30 years of anticipated operations. These 383 occurrences

were reduced to fatigue cycles using the range-pair counting method

described by Dowling. 4 All cycles with an amplitude less than +2 ksi

were considered non-damaging and were omitted from the spectrum.

Similar occurrences were grouped in the 8-step spectrum shown in Fig.

63. The spectrum included 53 stress cycles, 19 of which were primarily in compression. For analysis purposes, these 19 cycles were

treated as zero-to-tension cycles of equal stress range.

The axial stress spectrum used by Alyeska consisted of 151 cycles

of zero-to-tension stressing. The significant differences between

the Alyeska and the NBS spectra were attributed to the range-pair

counting method, which combines stresses in such a way to give fewer

cycles and higher amplitudes. The Alyeska approach resulted in a

4. N. E. Dowling, Theoretical and Applied Mechanics Report No. 337,

University of Illinois, Urbana (Jan., 1971).

* Provided by Alyeska and described in detail in Section 5.

132
more conservative spectrum because settlement was treated as a distinct cycle of 0 to 361 MPa (52.4 ksi) instead of a shift in the mean

stress. A spectrum severity study, summarized in Fig. 64, was conducted to evaluate the relative severity of the two spectra. The

results indicate that the Alyeska spectrum is approximately 25% more

severe in terms of predicted life than the NBS spectrum.

A hoop stress spectrum representati ve of X65 pipe operating at

maximum design pressure was used by both NBS and Alyeska. Review of a spectrum that included the lower internal pressures typical of Phase I and II operations (the first 30 months) indicated that the 3-step spectrum

was conservative. However, the degree of conservatism has not been

establ ished.

Fatigue-Crack Growth: Weld Defects

Several fatigue-crack growth analyses were conducted using the

worst-case axial-stress spectrum to independently assess the growth

characteristics of weld defects in the trans-Alaska pipeline. The

results indicate that cracks up to halfway through the thickness

grow 0.001 inch or less in one service life when exposed to a nonaggressive environment as represented by Maddox's data with the

required safety factor of 4. In a corrosive environment representative of a holiday in the corrosion protection system, cracks up

to halfway through the thickness grow less than 0.010 inch in one

service life; Maddox's data with a factor of 40 were used for these

calculations. In general, this increment of growth does not noticeably affect the alternative allowable-defect-size curves.

Additional analyses were conducted to evaluate the adequacy of

the procedures and assumptions used by Alyeska in their assessment

of fatigue-crack growth. The analysis procedures used by NBS and

Alyeska are compared in Fig. bb. In this comparison, the Alyeska

curve was taken from Part 4 of the Alyeska submittal and the NBS data

were generated using the same stress spectrum, fatigue data and thickness. The results indicate that the two analysis methods yield

similar results for crack lengths ranging from 50 to 100 mm. The

differences in the two methods are attributed to differences in the

equations for computing AK; small differences are amplified because

growth rate is proportional to the cube of AK.

A study of environmental severity was conducted to compare the

growth rates computed by Alyeska using Vosikovsky's data with those

obtained by NBS using

summarized in Fig. 6b

conservative than the

through the thickness

same flaw grows about

a factor of 10 with Maddox's data. The results,

indicate that Alyeska's analysis is far more

NBS analysis. For example, a flaw halfway

grows 1.2 mm in the Alyeska analysis and the

0.2 mm in the NBS analysis.

133
In summary, the NBS growth predictions for weld defects subject

to the axial-stress spectrum indicate that growth will be limited to

less than 0.010 inch for flaws less than half the thickness in depth

Thus, fatigue is not an important consideration. However, the conservative approach taken by Alyeska can predict substantial growth,

e.g. planar flaws analyzed using Vosikovsky's data.

Fatigue-Crack Growth: Arc Burns

Fatigue-crack growth analyses were conducted using the worstcase hoop-stress spectrum to determine the growth characteristics of

arc burns in the trans-Alaska pipeline. The results of both the NBS

and the Alyeska analyses indicate that substantial fatigue-crack

growth occurs in adverse environments.

Comparisons of spectrum severity and analysis methods are shown

in Fig. 67. The results indicate that Alyeska' s practice of considering all cracks to have infinite length results in only a slight

degree of conservatism even though arc burns are generally short.

The environmental severity of Alyeska's analysis using Vosikovsky's

data results in a 5 mm deep crack growing 2.5 mm in one service life

A similar crack in the NBS analysis grows 1.7 mm.

Since arc burns are generally short, the influence of crack

length on fatigue-crack growth was calculated. The results, presented in Fig. 68, indicate that a 5-mm-deep flaw will grow 1.5 mm

in a service life when the flaw length is approximately twice the

flaw depth and the same flaw will grow 2.5 mm when flaw length is

twenty times the flaw depth.

Due to the severity of the hoop-stress spectrum, it may be

necessary to consider the actual stresses acting on the location of

an arc burn in accordance with 0PS0 Requirement III. The influence

of stress level on fatigue life is summarized in Fig. 69. a 5-mmdeep defect grows 1.7 mm in one service life for the worst-case

spectrum (stress factor = 1.0) and the same defect grows 0.5 mm

when the stress is reduced to 70%.

Conclusions

1. Fatigue is not an important consideration in establishing

alternative allowable weld-defect sizes subject to the pipeline

axial-stress spectrum and an environment that causes a tenfold enhancement in fatigue-crack growth rates.
2. Fatigue causes significant growth of cracks subject to the

pipeline hoop-stress spectrum and a corrosive environment and must

be considered in establishing allowable defect sizes applicable to

arc burns.

3. The Alyeska analysis appears to be more conservative than

the analysis conducted by NBS, particularly in the area of environment-enhanced fatigue-crack growth.

ea a.

CO

CO

CO

400 r- —160

FATIGUE STRESS SPECTRUM

53 Cycles=30 Year Life

L0P,53°C300

L0P,AT=26°C

L0P,AT=53°C -]

200

100

I—Hydrotest

0 -

•100

-Year b

Years 2.3.4

Causes of Stress

• Hydrotest

• Loss of pressure (LOP)

•Temperature drop (AT) following LOP

• Settlement-causes increase in minimum stress

• Earthquake-negligible, not shown

•20

-Years 5 to 30-

OCCURRENCES. N

Fig. NBS-derived axial -fatigue stress spectrum.

135

STRESS

(ksi)
Table 19. Stress spectra used in fatiguecrack-growth calculations.

Number of Minimum Stress Maximum Stress

Cycles (ksi

)

(ksi

)

Axial -stress spectrum(NBS)

a

2 5.3 22.4

10 -16.0 -3.1

3 -16.0 5.9

1 -16.0 14.9

3 5.8 29.4

24 20.4 33.3

9 20.4 12.3

1 20.4 51.2

Axial

•

•stress spectrum(Alyeska)k

2 0 17.1

17 0 14.4

59 0 12.97

2 0 9.05

5 0 17.91

'15 0 8.86

1 0 52.4

10 0 3.86

10 0 0.96

30 0 0.12

Hoop-!stress soectrum

c

2 0 61.75

17 0 52

59 0 46.8

a. Derived by NBS.

b. From Alaska Pipeline Service Company Report

"Fracture Mechanics Study of Buried Field

Girth Welds of the Trans-Alaska Pipeline,

Part IV."

c. Derived by NBS; used by Alyeska and NBS.

136
14

12

10

8

6

4

2

0

14

12

10

8

6

4

2

0

14

12

10

8

6

4

2

0

Alyeska.

t=11 73mm

H 0.5

-0.4

—I 0.3

SPECTRUM SEVERITY STUDY

Axial Stress Spectrum

Maddox, factor of 4

Non-Aggressive Environment

- 0.2

—i 0.1

50 100 150 200

J o

I

i =400mm

l =100mm

l =50mm

—1 0.5

c - 0.4 «

Alyeska (Gurney), / = oo H0.3

ANALYSIS METHOD COMPARISON

Axial Stress Spectrum (Alyeska)

Maddox, factor of 4 applied

Non-Aggressive Environment

t=11.7 mm

- 0.2

0.1

J,

50 100 150 200

ENVIRONMENT SEVERITY STUDY

Axial Stress Spectrum

0.5

1

0.4

NBS

Maddox, factor of 40

*^ 0.3

i 0-2

-I 0.

10 15 20

SERVICE LIFETIMES, N

137

CRACK

DEPTH,
CRACK

DEPTH,

a

(mm)

SPECTRUM SERVERITY STUOY A,NO ANALYSIS METHOO COMPARISON

14 r— Hoop Stress Spectrum

0 1 1 1 L 0 2 4 6 8

INFLUENCE OF CRACK LENGTH ON FATIGUE CRACK GROWTH

0 5

I J 04

HO 3

h 0.2

hO.1

Hoop Stress Spectrum

Vosikovsky s Oata (with factor of 4)

-0.5

- 0.4

- 0.3

- 0 2

- 0.1

-Jo

- 0.5

JJQ.4

-J 02

0.3

0.1

-0

O

<aco

SERVICE LIFETIMES. N

Fig. 67

Fig. 68

ig 69

138
7.B. ALLOWABLE-DEFECT-SIZE ANALYSIS

1 ) Introduction

The OPSO notice for anticipated Petition for Waiver states that,

"alternative allowable defect sizes should be proposed applicable to

each type of defect for which a waiver is being requested. Proposed

alternative allowable weld defect and proposed allowable arc burn sizes

must be supported by fracture mechanics analysis using the worst case

fatigue stress spectrum. For analysis, these defects must be assumed to

be surface cracks equal in size to twice the proposed allowable weld

defect or arc burn size (in both length and depth). These assumed

defects must not grow in size such that stressing to the maximum credible

service stress could cause leakage."

In response to these requirements, Alyeska has proposed allowabledefect-size curves applicable to four types of defects for which waivers

are being requested:

Type 1: Non-planar weld defects

Type 2: Planar weld defects

Type 3: Arc burns on the weld or HAZ

Type 4: Arc burns on the base metal

Since all defects are treated as surface cracks, the principal differences

in the four types are defect orientation, defect location, and applicable

safety factors. Defect orientation determines whether the applicable

stresses are axial or hoop. Defect location is used to establish the

applicable minimum- fracture- toughness value and the appropriate operating

environment (inside or outside the pipe). The applicable safety factors [G-4]

include the factor of two on length and depth for all defects plus an

additional factor of two on estimated depth (by radiography) of planar

defects

.

The objective of the work reported in this section is to independently

calculate allowable-defect-size curves in accordance with the OPSO notice

requirements , and by comparison to evaluate the adequacy of the allowabledefect-size curves proposed by Alyeska. To meet these objectives,

calculations of allowable defect sizes were made by NBS consultants

(G. R. Irwin of the U. of Maryland and J. A. Begley of Ohio State

University) and by McHenry - Read, Fong and Interrante of NBS, and the results

were compared with the all owabl e-defect-size curves submitted to OPSO by

Alyeska. Some limited experimental results are also discussed.

This section first describes the Alyeska and other analytic models

used to evaluate the Alyeska model. To calculate allowable-defect-sizes

from each model requires the use of material -property data representative

of the trans-Alaska pipeline; these data are discussed. Following

calculations of allowable-defect-size curves, proper assessment of the

reasonableness of the Alyeska curve is required. Factors relative to this

assessment are detailed in the discussion.

133
2) Analytic Models

Allowable-defect-size models have been established by Alyeska.

Models have also been developed by Begley, McHenry and Read; Fong;

Irwin; and Interrante and Kiefner. These approaches are described

bel ow.

Alyeska : Alyeska established allowable-defect-size curves by

computing a defect-size curve, subtracting the increment of crack growth

due to fatigue cycling and then applying the required factors of safety

on defect length and depth. A defect-size curve for each type of defect

was calculated using the procedures described in the Draft British

Standard Rules for Derivation of Acceptance Levels for Defects in Fusion

Welded Joints (see Appendix I, Reference 1-3). The model is based on an

empirical correlation between critical defect sizes and applied stress

for through-cracks in welded wide plates. The procedure is to calculate

an "allowable-defect parameter" by

( 1 )

Here the residual stress (aJ is assumed equal to the yield stress

(r x ). The maximum credibl e appl ied stress is a. From the ratio a/t, and

Fi-$. 14 (applies only when a/t < 0.5) of the Draft Standard, several pairs

of the two quantities a/t and a/t are obtained, where a = defect depth,

t = thickness and I - defect length. From these quantities, an a-versus-t

curve can be drawn that the Draft Standard regards as the boundary between

the acceptable and rejectable welds. However, in accordance with the

notice of Petition for Waiver requirements, each curve (4 types of defects)

is subsequently modified by subtracting the increment of crack growth due

to fatigue as discussed in the previous section of this report and by

applying the required factors of safety on defect length and depth.

Begley, McHenry, Read : Calculations of allowable defect sizes are

based on tne recommendations of J. A. Begley as described in Appendix B.

The model has been adapted to crack opening displacement measurements

from a well-known expression regarding the opening of a through-crack

uncer remote tensile stress. In this calculation, a trial defect of a given length and depth is assumed. This defect is allowed to grow over

tne life of the pipeline by adding calculated increments to the depth

and length of the defect for each stress cycle. These increments are

computed using the procedures described in the preceding fatigue-crackgrowth section of this report (Section 7. A.) The defect is checked

140
for crack growth through the wall after each stress cycle. If the defect

grows through the wall, failure is computed. After all stress cycles of

the projected service life have been applied, the maximum credible stress

is applied to determine if the critical COD, 5 , is exceeded. The applied

COD is calculated as follows:

Here z, the effective crack length, is corrected for the plastic zone

size, and a and a are the maximum credible applied stress and the

minimum credible flow stress respectively . Failure is computed if

the applied COD is greater than the critical COD. Residual stress is

accounted for by subtracting 0.001 in. from 6 as recommended by Begley

(Appendix B). If the trial defect is computed to fail, its initial

depth is decreased and the calculation is repeated until an initial

defect depth is reached at which the trial -defect-depth survives. The

length and depth of this surviving defect are then plotted; several such

points determine the accept-reject curve. The relative position of the

curve is strongly dependent on the difference between the applied stress

and the flow strength. The use of 70 ksi as the flow stress is considered

reasonable based on NBS weld process 102A-1 data and on discussions with

Irwin, who developed the equation, and Begley, who recommended its use.

Fong : In order to gain further perspective, a modification of the

Begley-McHenry-Read model in the moderately-deep surface crack regime

(0.1 < a/t < 0.5) is proposed (Appendix P) as an alternative model for

the derivation of allowable defect-size curves. The new model is

essentially the simultaneous extension of the Begley-McHenry-Read' s deep

crack model and a shallow crack model due to Begley-Landes-Wilson. As

reported by Begley in Appendix B, a shallow crack model (a/t < 0.1) for

infinitely long surface cracks may be described by:

( 2 )

(3)

•k where a is the effective depth of the surface crack which is related to

the critical depth a by:

(1.2) (2t)

( 4 )

The modification consists of introducing a new form of J:

141
J ( 5 ) = A E t a (2c - \

sy

). £ > 7

£

y

which J is related to the COD by

J = m<^y

The coefficient A is introduced as a safety factor to insure a noleak and no-burst condition. The factor of 1/2 is inserted to bring

the shallow crack model more in line with the normalized COD design

curve than is the case for the original Begley model. A discussion of

this model appears in Appendix P.

Irwin : Professor G. R. Irwin of the U. of Maryland, serving as an

MBS consul tant, has calculated and reported (Appendix D) a set of critical

flaw sizes. The calculations are based on a plastic-instability model

which features critical ligament strain rather than fracture toughness

as a failure criterion. Bulging and curvature (using shell theory)

effects are considered in the Irwin analysis, therefore, exterior surface

defects are most critical. Buried defects of a given depth can be approximately

30" longer than external surface defects prior to reaching critical size. The

increment of crack growth due to fatigue is negligible for the case considered

by Irwin, i.e., weld defects subject to axial stress.

Interrante, Kiefner : A series of critical defect sizes were

computed by C. Interrante (NBS, Washington) based on the Kiefner equation : 1

a a

1 - a/t

1 - a/tM ( 6 )

where M = (1 + 1.255 ~ - 0.0135 -j-j ) 1/2

(7)

c = i/2 and r = pipe radius. This semi-empirical equation has been

developed by Battelle 2 to model successfully the fracture behavior of

longi tudinally-oriented defects in full-scale samples of line pipe.

Interrante's results, including fatigue and the application of the

required factors of safety, are presented only as allowable-defect-size

curves of arc burns, representative of longitudinal defects.

1. Kiefner, J. F., Maxey, W. A. Eiber, R. J. and Duffy, A. R., ASTM

ST? (1973), pp. 461-481.

2. P oaiasek, R. j. and Eiber, R. J., Battelle report to Alyeska, 10/9/73.

142
3) Material -Property Data

The material -property data needed for the analysis include the

fracture toughness, the fatigue-crack-growth rate, the stress corrosion

threshold, and the tensile properties. Although both 102A-1 and

104A-1 weld process materials were tested, only 102A-1 process defects

are presently included in the Alyeska waiver request. The applicable

OPSO requirements and the sources of all the data used in the various

analyses are reviewed in this section. The values actually used in

the analyses are summarized in Table 20. The relationships of these

values to the data generated by various programs in support of the

trans-Alaska pipeline project are shown in Table 21.

Fracture Toughness : The OPSO notice for the anticipated Petition

for Waivers states that: "a minimum fracture toughness value for the

pipeline shall be established by documenting the fracture toughness

in sufficient notch locations and temperatures for the weld metal and

the heat-affected zone that is representative of the pipeline welds and,

in the case of arc burns, for the base metal. The toughness value used

in the fracture mechanics analyses should be the minimum toughness at

10 °C below the minimum anticipated service temperature.

Alyeska responded to these requirements by submitting fracturetoughness data for nine notch locations in the weld metal, HAZ, and base

metal. Tests were conducted at temperatures ranging from -40 to 0 °C

using the crack-opening-displacement (COD) method. The temperatures of

interest in establishing the allowable-defect-size curves are -12 °C and

-22 °C, which correspond to 10 °C below the minimum operating

temperature (-2 °C) and 10 °C below the minimum winter start-up

temperature, respectively. The toughness values (0.004 in. for nonplanar weld defects and 0.007 in. for all other defects) established on

the basis of Alyeska 's test programs at the British Welding Institute

and at Cranfield Institute of Technology were generally used for both

the Alyeska and the NBS analyses. These COD values are taken as lower

bound values for the toughness at the temperature of interest. However,

as discussed in Section 6, a single lower COD value of .002 in. at 0 °F

was measured when larger sized COD test specimens were used. The one

exception is the case of base-metal toughness applicable to certain

arc burns; the Alyeska value of 0.011 in. for COD was not regarded

as a well-defined minimum since only one data point was used. Therefore,

the Begley, McHenry, Read curve contains a value of 0.007 in. for COD.

This value could be justified on the basis of the use of the minimum

value of COD at -40 °F, for which the Alyeska submittal contains a number of data points. The 0.004 in. COD value used for non-planar

weld defects was confirmed by NBS (see Section 3E), using J-integral

test methods and a comparable notch orientation. For weld process 102A1, the most conservative estimate of COD at 28 °F is 0.0047 in., using m

= 2.0 in equation 3 (Section 3.E.).

Fatigue-Crack Growth : The requirements for fatigue-crack-growth

data and the data used in the analyses are reported in Section 7. A. of

this report.

143
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145
Sustained-Load Failure : The OPSO notice for the anticipated

Petition for Waivers specifies that a minimum threshold for sustainedload failure must be established. Tests conducted by NBS (see Section

3.H. of this report] and by the British Welding Institute for Alyeska

indicate that sustained-load failure does not occur in the pipeline

welds and base metal for service environments representative of the

pi pel ine.

Tensile Properties : The tensile properties of interest in the

analysis are the yield strength, the flow strength and Young's modulus.

The flow-strength value for the weld material used is 70 ksi. This was

determined by averaging the minimum yield and tensile strengths (see

Section 3.D.) at +28 °F, the temperature where the maximum credible

stress is applied to the pipeline. From Alyeska data, supplied by

Thyssen Stahlunion A6 (the supplier of Phoenix Cel 80 electrodes used in

the welding process 102A-1) the average flow strength is 83 ksi with a two standard deviation (2c) value of 7.8 percent. Since in the models

the use of a low flow strength is conservative, and since the Alyeska

and Begley, McHenry and Read models are very sensitive to flow strength

(see Section 7.C.), the lower NBS-determined weld flow strength was

used; a reasonable estimate of the possible variations of this flow

strength is ± 7.8%. The base metal flow strength used is 75 ksi, which

corresponds to grade x 65 pipe.

148
4) Calculations from Models and Comparison with Experimental Results

The alternative all owabl e-defect-size curves submitted by Alyeska and

calculated on the Begley-McHenry-Read, Fong, Irwin and InterranteKiefner models are summarized in Figs. 70-73. The material property data

used to develop the curves are shown in Table 20. Since the various models

are based on differing physical pictures and approximations, material

property data were used which reflect the requirements and analyses on

which the models were based, and on the various safety factors and

requirements specified by OPSO. For example, with the minor exceptions

noted in Table 20, all the analyses used maximum credible stresses of

57.7 ksi for weld defects and 52 ksi for arc burns. It is apparent

from the curves that the different models predict quite different

results in some of the regimes of interest, and the implications of these

differences are discussed in Subsection 5.

Unfortunately, there are only limited large scale test data available

to assist in the substantiation of the analytical models. Fig. 74 displays

the actual data from the limited experiments available with the predictions

of the models plotted without the OPSO safety factors. Included in Fig. 74

are the NBS data obtained from simulated-service fatigue tests with

subsequent stressing to failure. It also includes data from Irwin,

Krishna, Yen 3 low cycle, flat-plate fatigue tests with flaws transverse

to the load axis, stressed to failure or to a chosen maximum stress

greater than 60 ksi. NBS specimens failed at stresses higher than the

maximum credible operating stress used in the above model calculations

(57.7 ksi for weld defects and 52 ksi for arc burns). The observed

Irwin, Krishna, Yen defect range represents specimen defect dimensions

that failed at applied stresses greater than 60 ksi or were stressed to

60 ksi or more without failure. Since the maximum stresses for these

data were larger than the applied stresses used in the model calculations,

the data are not precisely comparable to the calculated allowable defect

size curves.

Unfortunately, however, we do not believe these results provide

adequate information in the region of main interest, where a = 0.15",

z = 6" on Fig. 74. In order for the model to be fully verified, many

more data are needed.

5) Discussion

The purpose of this study is to develop as much independent

information as possible in order that DOT may make a judgement regarding

the proposed Alyeska allowable-defect size curve. In subsections 3 and 4

of this section, we have developed a number of independent models and

reviewed the limited experimental evidence available. In section 7C

a discussion will be presented of how sensitive the results of the various

curves are for nominal variations in the materials properties and stress

parameters.

3. Irwin, G. R., Krishna, V. C., and Yen, B. T., "Flat-Plate Testing of

Part-Through Cracks in Line-Pipe Steel Plates," Fritz Engineering

Laboratory Report No. 373.1, Lehigh University, March 1972.

/ 147
DEFECT

DEPTH

(in.)

WELD DEFECTS (Non

Gas Pockets

Elongated Slag

Isolated Slag

Wormhole Porosity

Hollow Bead

Spherical Porosity

Cluster Porosity

^

Begley -McHenry -Read

0 . 03- /^ Alyeska

0 . 02-

t 1 — \ —»— ( - t

1 i 1—

S3 tsa SI si S3 SI S3 S3 S3

CS s SI S3 S3 S3 S3 S3 S3

• • • • • • • • • — P4 n 3* Lrl CD r* CS G3

DEFECT LENGTH (in.)

-Planar)

Fin. 70. Allowable defect sizes for non-planar weld defects 148
DEFECT

DEPTH

(in.)

a. is-

{

i

t

0 . 1 H-!

0 . 1

3

” :

i

0 . 12- ^

0.10-

0 . 03-

0 . 00”

0 . 07”

0 . 05”

0 . 05”

0.0H”

0 . 03”

0 . 02”

0 . 01 "

a.cntca

s

WELD DEFECTS (Planar)

Incomplete Penetration

Incomplete Penetration (Hi /Lo)

Incomplete Fusion

Crater Cracks

[

\

\

\

\

\

\

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\yv

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v

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i

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4 1 1 1 1 ! 1 \ h SI si si S3 SI S3 S3 S3

S3 S3 S3 SI S3 S3 S3 S3

rvi ri 3“ Li Lfl r-* m . cri

DEFECT LENGTH (in.),

Fig. 71. Allowable defect ^lees for planar weld defects. 149
DEFECT

DEPTH

(in.)

ARC BURNS (Wald Haz.)

0.IH0.13-

(

0.12- !

t

I 0 . 11 + !

0.10- !

i

0.03-

0.00-U

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A

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0 . 02-

0 . 01 -

Interrante- Kiefner

Begley - McHenry - Read

•*••*•«••• ••• !«•«••• •••••••« • •

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Alyeska

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-

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S3 — ri m* 3"’

Lrt* m r- CD

DEFECT LENGTH (in.)

Fig. 72 Allowable defect sizes for arc burns (weld, HAZ).

See Appendix 0 for expanded scale curves

150

3.B0+
DEFECT

DEPTH

(in.)

DEFECT LENGTH (in.)

Fig. 73 Allowable defect sizes for arc burns (base).

See Appendix 0 for expanded scale curves.

151
Table

22.

Calculated

critical

flaw

sizes

at

very

long

defect

lengths

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152
DEFECT

DEPTH

Fig. 74 Comparison of experimental results with analytic models

for partTM through cracked specimens.

153
The first point to be emphasized is that the attempt in our report

to reduce the failure problem to a rational analysis does not supercede

the use of common sense in quality control during the welding process.

In particular, when by visual inspection, a weld does not meet normal

welding practice, the weld should be replaced without further question.

Our assumptions on material property variability, for example, only apply

to good quality material with the natural variability associated with

good practice and normal standards of quality. They would not apply

to a visibly substandard welding procedure.

The failure problem can be broken into several categories:

1) Failure by elastic fracture

2) Failure of the ligament

3) Failure by plastic fracture

1) Failure by Elastic Fracture

Elastic fracture mechanics provides a relatively well established

analytic basis for the three dimensional problem of an elliptical surface

crack. For an allowable'-defect-size curve, calculated on the basis of

elastic fracture weakness, to equal the curve shown by Aleyska, the

fracture toughness of the material would have to be about 25 ksi /nr. Thi

is well below even the fracture toughness measured (under dynamic loading

condition^ where elastic fracture did occur.

However, if defect sizes can only be measured with a factor of n,

and n is used as a safety factor to account for this^uncertainty, then

the required K» could be increased by a factor of n . For example, if

n is 2, «. wouTd have to be 100 ksi /fn. instead of 25 ksi /Tn. to account

for uncertainties in the measured crack length. .

2) Failure of the Ligament

In a two dimensional specimen, as a crack grows through the specimen

the entire load must be taken up by the remaining "ligament". As the

ligament continues to narrow, the stress in the ligament increases, and

failure will ultimately occur. In our notation, this condition is written

a

u ^ 1-a/t (8)

where a is the ultimate tensile strength of the material. For the

nominal Operating stress, a = 57.7, and a = 80 ksi, this condition is

satisfied for a/t well above the asymptote of the Alyeska curve. The

condition is however, very sensitive to variations in the material

parameter a , and to the applied stress. However, for the ranges suggested

in 7C, the Values of a/t fall on the Alyeska curve at long defect lengths.

154
The failure of a ligament (in two dimensions) might occur for a defect whose COD is too large for the fracture criterion to be satisfied.

However, in three dimensions (for lengths small compared to the total

pipe diameter) the stress in the ligament will not grow as suggested by

Eqn. (8), but the ligament will be partially supported by the rigidity of

the pipe beyond the defect length. (See Appendix P).

Hence, because the lower bound of the sensitivity band (Section 7C)

falls at the Alyeska asymptotic value, and because of the three dimensional

constraints, the Alyeska curve seems reasonable relative to simple ligament

failure.

3) Failure by Plastic Fracture

Two dimensional fracture criteria form a special case, because they

give a lower bound to the failure criterion -- the asymptotes of the

curves of Fig. 70-73.

Where the entire structure is loaded beyond yield, then the model

of Begley (Eqn. B-12) applies. Fong's model (Appendix P), and the Alyeska

asymptotic expression are all closely related to the Begley model (Eqn. B-12)

and are based on similar physical ideas. However, safety factors and

residual stress are handled somewhat differently . in these models. They all

show a relatively minor sensitivity to parameter variability.

When conditions are such that the plastic zone of the crack is

small compared to the pipe thickness, then the elastic case already mentioned

applies.

In the intermediate case, where a < a , but where the plastic zone

covers the entire ligament, we do not have^an adequate model, even though

the models of Begley (Eqn. B-12) and Alyeska have been extended empirically into

that region as discussed by Merkle. Various theoretical formulations of this

problem have been made, but have not been applied specifically to this

problem. By estimating the plastic zone size in a rough way, and taking

the average plastic strain in the ligament to be related to the stress

from the stress-strain curve, again in two dimensions, an estimate

(albeit crude) of failure can be made. We find an expression for the

critical defect depth to be

F is the linear hardening coefficient defined by

a = Ee

y

+ F ( e

- ) (10)

4. Merkle, J. G., "Analytical Relations Between Elastic-Plastic Fracture

Criteria", Int. J. Pres. Ves. & Piping, 4_, pp. 197-206 (1976).

155
5

From this, with the parameters a = 57.7, a = 60, F = 3x10 , 6 = .003",

t = 0.462", we find a s 0.1, which with usual safety factor of 2, becomes

about 0.05". Although this is a value above the asymptotic value given

by the Alyeska curve, it is very sensitive to small variations in the

materials properties and stress parameters.

The Begly-McHenry-Read model was developed independently of the

Alyeska model and is a reasonable expression of the present state of

analytical development. This model is, however, very sensitive to the

values of the parameters, especially for intermediate and long cracks,

as shown in Fig. 75.

In sumnarizing the above discussion, clearly it is not possible to

present an allowable-defect-size curve which has a satisfactory quantitative

theoretical basis. Each of the models are presented as estimates of

reasonable all owabl e-defect-sizes, and, except for Irwin's model, are not

based on a sophisticated theory.

In particular, the great sensitivity to variations in operating

and material parameters of some of the models is noted. We have not been

able to come to a clear conclusion that this sensitivity is "real", or

whether it is an artifact introduced in the models. On the one hand,

it seems that the large sensitivity, when it appears, is connected with

assumptions about how to estimate the stress in the ligament. As shown

in Appendix P a proper three dimensional treatment will lower this

ligament stress, and consequently lower the sensitivity. Such calculations

have not been carried out, therefore, a clear answer to the problem is

impossible. On the other hand, we find that the sensitivity is also

related to the very high stress assumed in the pipe in relation to the

yield stress. Since the models are all sensitive to the cross-over from

elastic-plastic to fully plastic conditions, such sensitivity would be

expected to remain "real" in an adequate model, although it might be

smaller. It is then a matter of engineering judgement whether the

Alyeska curve remains a reasonable allowable-defect-size curve in view

of its built-in safety factors, and in spite of the modest overlap

between it and the sensitivity band of the Begley-McHenry-Read curve.

It is the general opinion at NBS that the Alyeska curve is a reasonable

one (except for very short defects — see below). This conclusion

includes those aspects of the problem involving fracture mechanics and

materials properties, and does not include considerations of uncertainties

in the estimate of operating stresses and in the measurement of defect

sizes.

In parallel with the limitations in the models, there is an

inadequate base of large scale testing to complement the analysis.

Experimental results from the NBS and Alyeska test programs were principally

designed to obtain material property data.

In addition to the above general conclusions, we make the following

detailed consents relative to Figures 70-73.

156
DEFECT

DEPTH

(in)

Ficu 7". Effect of changes in the flow stress on allowable

defect size for planar defects (safety factors included).

157
For non-pi anar defects (Fig. 70) the Alyeska curves seem reasonable.

For planar defects (Fig. 71) and for arc burns (Fig. 72) the Alyeska

curves also seem reasonable, except for very deep cracks with lengths

less than approximately 0.5 inches. In this region, all the curves

except thatof Alyeska intersect the depth axis at depths between 0.08 and

0.13 inches, while the Alyeska curves show tolerance for unlimited defect

depth. This deviation is principally caused by the use of only fracture

criteria in the draft British Standards, while the other models consider

leakage as well as fracture criteria. In all regions where defect

depths on the Alyeska curve are above any of the Begley, McHenry, Read,

Fong, Irwin or Interrante, Kiefner curves, the Alyeska curve is then not

considered conservative with respect to pipeline leakage. Arc burns on

the base material (Fig. 73) constitute a special case. Alyeska, Begley,

McHenry, Read, and Interrante, Keifner show close agreement, except at

small defect lengths. Therefore, the Alyeska curve is considered

reasonable, except for short lengths (< 0.5 in.) where the other curves

are more conservative. For lengths shorter than the point where the

Alyeska curve rises nearly vertically (Fig. 73) the Alyeska curve is not

conservative. For longer lengths the curves are all in general agreement,

and the Alyeska curve is reasonable.

Summary Conclusions

The following conclusions include those aspects of the problem

involving fracture mechanics and materials properties, and do not include

considerations of uncertainties in the estimate of operating stress and

in the measurement of defect sizes.

1. The all owabl e-defect-size curve submitted by Alyeska for nonplanar defects (Fig. 70) is considered reasonable.

2. The all owabl e-defect-size curves submitted by Alyeska for

planar (Fig. 71) and arc burn defects (Fig. 72 and 73) are considered

reasonable, except for the region of small defect lengths, where the

other model curves fall below the Alyeska curve.

3. Through the generation of all owabl e-defect-size curves, fracture

mechanics is a procedure which can be used as a basis for making decisions

in accordance with the 0PS0 notice for anticipated Petition for Waiver

requirements . This report indicates that when a more definitive theoretical

analysis is available and is supported by adequate experimental studies,

a high degree of confidence will be possible in delineating acceptable

from unacceptable defects. There are also indications from the work of

Irwin that such a study might conceivably lead to a technical basis for

less stringent requirements on allowable defect sizes.

158
7C. SENSITIVITY ANALYSIS AND THE USE OF SAFETY FACTORS

1 . Introduction

In addition to specifying the use of both (a) worst case load combinations and (b) minimum material resistance parameters, in fracture mechanics

analyses for supporting any weld defect waiver request, OPSO further

stipulated [G-4] that:

"... these defects must be assumed to be surface cracks equal

in size to twice the proposed allowable weld defect or arc burn

size (in both length and depth)."

This implies that safety factors of 2 on both defect length and depth were

in effect adopted by OPSO in order to cover all uncertainties in design,

stress analysis, acquisition and interpretation of laboratory data, inaccuracies in defect sizes, and the use of simplifying assumptions in

modeling the fracture behavior of girth welds. For planar defects, OPSO

imposes an additional factor of 2 on defect depth because of uncertainty

in radiographic estimation of depth.

In deriving the all owabl e-defect-size curves as shown in the last

section, NBS has followed OPSO's choice of defect size safety factors and

its guidelines on material properties and stresses. However, during the

course of investigation it became necessary to choose additional safety

factors or adjustment constants to cover uncertainties in a variety of

technical areas. The purpose of this section is to point out the importance

of the role of safety factors, whether imposed by OPSO or used by NBS.

As observed by participants in an NBS meeting on Safety Factors 1

, the

choice of safety or adjustment factors is never arbitrary. We present in

the following (a) a listing of some important variables in our analysis

with variabilities large enough to be worthy of concern, (b) a sensitivity

analysis of the all owabl e-defect-size curves to indicate how the curves

may shift under reasonable variation of some of the variables listed in

(a), and (c) a listing of additional safety factors or adjustment constants

NBS had to introduce to accomplish its contractual objective as stated

in reference [G-5].

2. Significance of Variables Included in NBS Analysis

A rational basis for conducting a technical analysis of an engineering

problem lies not only in the choice of "relevant" variables for inclusion

in the analysis, but also in judging which variable cannot be analyzed for

the purpose at hand and how the significance of that variable can be

accounted for through the use of an appropriate safety factor. In a reallife project such as the construction of the trans-Alaska Oil Pipeline,

1 See Appendix M for a summary of the proceedings of a one-day meeting

on safety factors held at NBS-Gai thersburg on September 28, 1976.

159
the number of variables can be so high that any analysis, no matter how

careful and complete, may leave out some variables which eventually may

prove to be significant. The objective of this section is to show which

variables in the NBS analysis are sufficiently significant to motivate

a sensitivity analysis, and which ones are important and accounted for

through the use of safety factors. The first part leads naturally to

sub-section 3 where a sensitivity analysis is documented in detail, and

the second part motivates a discussion in sub-section 4 where some of the

important safety factors or adjustment constants used by NBS are described.

To begin with, we observe that the following variables are significant

enough to deserve a sensitivity study:

1) Applied stress (credible variation on worst case: +5%)

2) Yield or flow stress (credible variation: + 10%)

3) Fracture toughness (credible variation on minimum value: -10%)

4) Defect length or depth (acceptable criterion as a dependent

variable: -50% max.

)

5) Residual stress (credible variation: same as for yield stress)

There are a number of variables which are known to be significant for the

technical problem at hand but are treated in a different way in the NBS

analysis because there is not enough data in the literature to support

a rigorous analysis within the. time constraint of this project. The way

to account for them is through the use of safety factors. Aside from

those variables on defect size measurements which have been discussed in

section 4, the list of variables covered by safety factors may be subdivided into four categories as follows:

Category No. 1: Materials Properties : All of our tests were conducted

without the superposition of a biaxial stress state. It is, of course,

known that the girth weld is subjected both to a longitudinal stress and a hoop stress, and it is also known that fatigue and fracture resistance of a material is influenced by the existence of a biaxial stress state2 .

Category No. 2: Stress Analysis : The placement of 62 remotely operated

valves, n check valves and 9 check and manual block valve combinations along

the 800-mile pipeline made the task of hydraulic analysis exceedingly complex,

particularly when extreme loading conditions such as the shut-down of a

2 See, e.g., Kibler, J. J., and Roberts, R. , J. Eng. for INdustry, Trans.

ASME, (Nov. 1970), pp. 727-734; Roberts, R. , and Pothiraj, S., 2nd Int.

Conf. Struct. Mech. React. Tech., Berlin, Sept. 1973, Vol . V - Paper L

8/3; Brown, M. W. , and Miller, K. J., Proc. Inst. Mech. Engrs. (UK), 187,

745, (1973); and Zamrik, S., and Shabara, M. A., ASME Preprint 76-PVP-37,

presented at Int. J. Conf. Petroleum Mech. Eng. and Pressure Ves. & Piping,

Mexico City, Sept. 1976.

. 160
pumping station are investigated. Since check valves are designed to close

instantaneously when a reverse flow is experienced, the possibility of

hydraulic transients is increased when more check valves are installed as in

this case of the Alaska pipeline. Two facts may be of interest for future

consideration: (a) a comparable oil pipeline, also 48 in. O.D., the transAlpine pipeline, which has been in operation for many years, has not a single check valve along its route over mountains and terrains not unlike

those in Alaska; and (b) the hydraulic analysis of the trans-Alaska pipeline was completed assuming a mathematical model where the presence of all

check valves is ignored. An assessment of the surge analysis was not

required of NBS by DOT, and was not attempted.

Category No. 3: Sampling Problem : The selection of welds for laboratory

testing was carried out by a method appropriate for choosing a very small

number of sample units from a relatively homogeneous collection of potential

sample units. In statistical terminology, it was

" purposive selection , in

which the sample is restricted to units thought by someone to be especially

typical of the population or convenient for sampling. Purposive selection

may produce good results when the sample is small, but it is not amenable

to the development of a theory. . ." 3 The selection of test specimens

from each sample weld, for material properties measurements, was necessarily

guided by the required physical dimensions for test specimens rather than

by principles of statistical experiment design. Thus statistical methods

are not available for evaluation of weld-to-weld variability in the field

nor for discriminating between variabilities in material and those associated

with test methods.

Category No. 4: Modeling Problems : In deriving allowable defect

size curves, NBS used mathematical models based on sound physical arguments

and best judgments in approximating the actual phenomena of fatigue and

fracture. Unfortunately, all theoretical models lack full-scale experimental

basis in the range of interest where the defect depth to wall thickness

ratio lies between 0.1 and 0.4. Some discussion of the current state of

art in this specific area appears in Appendix P.

In the next two sections, we shall see how both types of sources

of uncertainty were studied.

3. Sensitivity Analysis

As remarked in the last section (7B) on Allowable Defect Size Analysis,

the defect sizes predicted by the Begley-McHenry-Read model are very

sensitive to a small change in applied stress or flow stress or both in the

neighborhood of defect depths from 0.09 in. down to 0.041 in., at which

the allowable defect length reaches infinity. The range of defect sizes

predicted by this model for small changes in operating and material parameters

3 William G. Cochran, Sampling Techniques , New York, Wiley, 1953, p. 7.
can easily be + 50%, and hence cancel the OPSO safety factor of 2. Using

physical reasoning, this behavior of the model was investigated to determine whether it arose from a mathematical artifact or was expressive of

the true structural behavior. The conclusions were given in Section 7B,

sub-section 5. Fig. 76 shows the variability of both the BegleyMcHenry-Read and the Fong models for weld defects of the non-planar type.

A similar analysis can be performed on planar defects, and the result is

shown in Fig. 77. The Fong model is seen in both cases as less

sensitive than the Begley-McHenry-Read model so far as variability due

to stress, flow stress, and fracture toughness is concerned.

4. Significance of Safety Factors and Adjustment Constants

Used in NBS Analysis

As mentioned in the introductory section, in addition to the use of

OPSO safety factors on defect sizes and fatigue growth rate, there were

numerous occasions when the analysis called for either a safety factor

or an adjustment constant or both to complete a specific task. The final

conclusion of our analysis depends, in part, on how we choose those factors

of safety and adjustment constants. The purpose of this discussion is to

illustrate how uncertainties of variables listed in four categories (see

sub-section 2) are handled through the use of safety factors.

Regarding Category No. 1, Material Properties, we begin by observing

that in addition to a factor of 4 on fatigue growth rate as specified by

OPSO, NBS introduced an additional factor of 10 on growth rate in its

fatigue analysis (see section 7A). The value of 4 is reasonable for fatigue

growth data in general, but because stress corrosion and biaxial stress

can be important factors, an extra margin of safety through the value of

10 was used. Turning to uncertainties due to change in fracture toughness

from weld to weld, it is worth noting that OPSO requires the use of the

minimum value. This led us to waive the use of any additional safety

factor on fracture toughness, but we did introduce in our sensitivity

analysis a variation of -10% to see how sensitive our allowable defect

size curves are to this change.

On Category No. 2, Stress Analysis, we recall that there was a +5%

variation of applied stress in our sensitivity analysis. This is reasonable

considering the fact that the applied stress we use is the worst loading

case of extremely low probability. We also observe that the NBS stress

spectrum for fatigue analysis is not as severe as the Alyeska one. We

believe that our choice is realistic and any variability in the stress

spectra of different welds can be accounted for by the combination of the

values o' 4 (OPSO) and 10 (NBS) in safety factors for fatigue crack growth

rates. On extra stresses, Alyeska documents asserted that the overstress

due to surges will never be more than 10% over the design stress. Our

allowance for this type of uncertainty lies in either a 10% reduction of

the yield stress or a similar amount for the fracture toughness. The

relationship between applied stress and mechanical strength is, in general,

nonlinear, but this discussion in qualitative terms illustrates the scope

of technical consideration that usually goes with the analysis.

162
DEFECT

DEPTH

(in.)

0.15"

0.IH0. 13"

0 . 12+

0 . 11 -

0.10

: : \ :

..

:

: \

*•

: i \ \ • l 1 •

0.53'

: • \

: ? \ \

: : \

%

\ \

c.oe0.07

B.S5i

0.GS0.0H

0.03-

0 . 02-

0 . 01 -

0 . 0OJ

\

r

Irwin

\ ® 1 •

• ® •

l|\ \ © •

! I \

\

! \ \

\

\

\

\

TJ

ca

S3

WELD DEFECTS (Non- Planar)

Gas Pockets

Elongated Slag

Isolated Slag

Wormhole Porosity

Hollow Bead

Spherical Porosity

Cluster Porosity

& \

\ k Begley - McHenry - Read

1*

«• 7.

jf

•ft Fong

zAlyeska

4-

C3

S3

S3

SI

4-

E3

S3

S3

SI

4- 4- C3

S3

S3

S3

SI

S3

4-

S3

S3

4-

E

I

S3 — rM n L/1 LO S3 1

DEFECT LENGTH (in.)

Fia. 7C. Sensitivity Analysis on Allowable Defect Size Curves Cion-Planar)

163
DEFECT

DEPTH

(in.)

I

t Q.IH+i

0 . 02"

0 . 01 -

0 .00-1 + 1 ! + i -H ! + 'Q d a El El El Cl si

S3 a S3 SI El a S3 El El

« • • • • • • • • S3 — rvi m 3“ Lrl to r* m

DEFECT LENGTH (in.)

Fig. 77. Sar.sit j.vltv Analysis on Allowable Dafccu ri/.a Curves (Planar)

3

. 03

+
Field variability of material properties and construction quality

(Category No. 3 on Sampling Problems) are addressed by applying those

safety factors imposed by OPSO such as 2 on defect length and depth and

an extra factor of 2 on planar defect depth. To illustrate how we handle

uncertainties in modeling problems (Category No. 4), the reader is invited

to examine Appendix P where an extra coefficient of 1.5 (= x) was introduced when a shallow crack model was extended to a moderately deep crack

region and a modification of the theory was called for. As more experimental

data and mathematical analysis become available, that factor x is likely

to be adjusted downward.

5. Conclusion

This report summarizes the technical contributions of many investigators.

We have shown that the choice of safety factors and adjustment constants is

an integral part of this investigation. The conclusion we have arrived at

depends not only on the analytical and experimental results of those

investigators, but also on their technical judgments as embodied in the

usual way through the use of safety factors.

165
8. WELD DEFECT SUMMARY

Sources of Information

Defect-size (1) information in this section is from the following sources:

"Presentation and Analysis of Quality Assurance Audit of 1975

Pipeline Welds/Radiography and Alyeska Engineering Technical

Evaluation." (Approximately April 1976).

(?) "Southwest Research Institute Review of Girth Weld X-Rays," Volumes

I, II, and Addenda, S. Wenk and C. Allen (received by Alyeska

July 21 , 1976).

(3) "Welds Requested for Exemption," Docket No. 76-12W, Appendices I and III, submitted by Alyeska Pipeline Service Company. (Revised

September 10, 1976)

.

Defect-location and pi pel ine- variable information in this section is from

the following sources:

(1) Alignment sheets of Alyeska Pipeline Service Company for trans-Alaska

pi pel ine.

(2) Trans-Alaska Pipeline Project Atlas, Volume I, Livengood to Prudhoe

Bay, prepared by M. Baker, Alyeska Pipeline Service Company (June 1976).

Defect-Size Data

Defect-depth measurements and arc-burn length and width measurements

are required for assessing the Alyeska defect-size data. Defect size

measurements are summarized in Tables 23-25. The measurements made by SWRI

and RI ere included in this summary without correction factors. For RI

arc burr depth measurements listed in Table 25, the depths converted from

arc burn widths (using Fig. 33) are included in the RI data. For completeness anc for comparison, all data available to NBS from Alyeska (length

only), Southwest Research Institute (SWRI), and from the September 10,

1976, Docket No. 76-12W submittal are included in these tables. From the

tables, it is apparent that several omissions and discrepancies exist in

these defect data; these are itemized in Table 26 and they include:

(1) Lack of data on defect size from either Alyeska or SWRI to support

the data submitted in the Docket.

(2) Omission of specific defect length or depth data in the Docket.

(3) Failure to use the most conservative value from Alyeska or SWRI

defect-size data base.

166
During the radiographic measurements by RI, changes of a few defect types were

suggested:

Weld #81586, location 81, listed as a gas pocket by Alyeska;

described as an arc burn by RI. It is included in both of these categories

in the present report.

Weld #81639, location 125, listed as hollow bead by Alyeska; described as

both hollow bead and incomplete penetration by RI and SWRI. It is included

in this report in both categories.

Weld #81 590R , location 143, listed as spherical or cluster porosity by

Alyeska; described as an arc burn by RI. This defect is included in this

report in both categories.

Weld #90008, location 71, listed as arc burn by Alyeska; described as a "dull tool gouge" by RI. This defect is treated in this report as an arc

burn.

Weld #W15TRX, location 55, listed as arc burn by Alyeska; is described as

an undercut by RI. This defect is treated in this report as an arc burn.

Weld-Location Data

From the alignment sheets of the trans-Alaska pipeline, supplied by Alyeska,

and from the Docket submittal, data are presented in Appendix A pertaining to

exact location of the welds that contain the waiver-request defects.

There are several items of concern:

(1) All previous discussions with Alyeska had established firmly that the

pipe grade adjacent to waiver-request weld defects was X65; however

from Appendix A it is apparent that the majority of pipe is grade X70.

This change is of concern because maximum credible stresses are

expected to be larger in grade X70. _Also, the worst-case

fatigue spectrum probably has greater stresses in X70

pipe areas. While Alyeska has represented the credible stress as a maximum and the fatigue spectrum (Table 19) as a worst case, careful

consideration of the effects of X70 grade pipe on these stresses

is -required.

(2) Downstream from a pump station, the pipeline is exposed to

operating pressure surges that are not represented in the worst-case

fatigue spectrum (Table 19). The magnitude of the pressure

variations decreases as the distance from the pump station increases.

Unfortunately , there are several defects whose location is close to

pump station number three (see Appendix A). Documentation of the

magnitude of the resulting fatigue spectrum in the pipeline very

near, and downstream to, a pump station is needed for accurate

assessment of those defects.

167
Table 23. Defect sizes of all planar defects.

Weld

#

Radiographic

Location

L E N G T H

Final

Docket

(9/10/76)

. (In)

SWRI

+

(in)

DEPTH

++

RI

(in)

Alyeska

(in)

SWRI

(in)

Final

Docket

(9/10/76)

(in)

PLANAR DEFECTS

Incomplete Penetra tion

90057 2 N.A. 3.50 3.50 N.A. Est. 0.020 0.035

Incomplete Penetration (Hi/Lo)

81627 111 N.A. 2.00 2.25 Est. 0.030 Est. 0.030 0.013

81635*y 140 N.A. N.A. N.A. Est. 0.025 N.A. 0.022

81706 44 3.00 3.00 2.50 Est. 0.040 Est. 0.035 0.028

64 3.00 3.00 3.00 Est. 0.035 Est. 0.035 0.030

90056 0 N.A. N.A. 5.50 N.A. Est. 0.020 0.035

Incomplete Fusion

81408 89 5.00 N.A. 5.00 Est. 0.015 Est. 0.015 0.024

81430** 135 10.00 N.A. 10.00 0.014 0.014 0.046

81448 135 1.50 1.375 1.375 Est. 0.035 Est. 0.035 0.013

81485 17 2.00 2.00 2.00 Est. 0.030 Est. 0.030 0.012 - 81498 124 N.A. N.A. 2.25 N.A. Est. 0.025 0.028

81499 105 N.A. 2.75 2.75 Est. 0.020 Est. 0.020 DNF

81506 20 1.25 1.25 1.25 Est. 0.035 Est. 0.035 0.020

81595 136 2.00 N.A. 2.00 N.A. N.A. 0.020

81639*y 125 N.A. N.A. N.A. N.A. N.A. 0.038

90008** 124 N.A. N.A. 0.388 N.A. Est. 0.060 0.048

I.A. .Not available.

Not included in 9/10/76 Docket No. 76-12W submittal.

Included in 9/10/76 Docket No. 76-12W, Appendix III submittal.

’. Hollow bead reported at same location by SWRI, RI, not incomplete fusion.

See P. 90 for recommended correction factor.

! + See Fig. 48 for recommended correction factor.

NF Did not find.

168
Table 24. Defect sizes of all non-planar defects.

Weld * Radiographic

Location

Alyeska

(in)

TTN GTH

TWRl

(In)

Final

Docket

(9/10/76)

(in)

SWRI

+

(In)

'T'E P T H Final

Docket

(9/10/76)

(In)

RI~

(In)

NON-PLANAR DEFECTS

Gas Pockets

81153 48 0.25 0.15 0.25 0.018 0.018 0.048

81209 129 0.188 0.175 0.188 0.007 0.007 0.043

81316 50 N.A. 0.150 0.188 0.057 0.057 0.075

81321 68 0.375 0.375 0.375 0.015 Est. 0.015 0.053

81338R 23 0.188 0.150 0.188 0.024 0.024 0.060

81404 18 0.25 0.20 0.25 0.024 0.024 0.038

81434R 32 0.25 N.A. 0.25 0.006 0.006 0.025

81441

R

74 0.188 0.125 0.188 0.010 0.010 0.030

31447 57 0.188 0.155 0.188 0.037 0.037 0.056

81565T 112 0.188 0.150 0.188 N.A. 0.039 0.018

81577 125 0.188 0.150 0.188 N.A. 0.016 0.035

81585 0 0.188 0.250 0.188 Est. 0.050 Est. 0.050 0.078

81586“ 81 0.188 0.375 0.188 N.A. 0.039 0.058

81628 140 0.188 0.200 0.188 N.A. 0.015 0.045

81670 70 0.156 0.125 0.312 0.018 0.018 0.050

81674 28 0.219 0.180 0.219 0.006 0.006 0.025

81701 75 0.188 0.188 0.188 N.A. 0.006 0.068

W15TRX 43 N.A. 0.150 0.156 0.038 0.007 0.042

W25RTRX 74 N.A. 0.135 0.156 0.047 0.047 0.065

W39TRX 81 N.A. 0.25 0.25 0.014 0.014 0.058

NON-PLANAR DEFECTS

Elongated Slag

81135 115 4.062 7.00 4.562 Est. 0.020 Est. 0.020 0.037

81180 10 2.25 1.00 2.25 Est. 0.030 Est. 0.030 0.043

81293 * 24 2.50 2.25 2.50 N.A. N.A. 0.038

81397** 111 6.00 0.8125 N.A. 0.033 0.033 0.035

81401 116 5.00 N.A. 5.00 Est. 0.015 Est. 0.015 0.026

81411 145 0.375 N.A. 3.625 Est. 0.015 Est.0.015 0.044

81420 123 2.75 N.A. 3.50 Est. 0.010 Est. 0.010 0.032

81424 15 3.375 N.A. 3.375 Est. 0.010 Est. 0.010 0.035

81451 146 N.A. 1.00 2.25 N.A. N.A. 0.053

81474 105 3.00 3.75 2.50 Est. 0.030 Est.0.015 0.030

122 3.00 2.375 3.00 Est. 0.030 Est.0.015 0.031

81476 125 3.00 3.50 3.00 Est. 0.020 Est. 0.020 0.033

81488 106 4.50 4.00 4.00 Est. 0.020 Est. 0.020 0.036

124 4.50 5.625 5.625 Est. 0.020 Est. 0.020 0.048

81498 135 2.50 2.25 2.50 Est. 0.025 Est. 0.025 0.020

81499 135 3.50 4.25 4.25 Est. 0.020 Est. 0.020 0.034

81503** 123 3.00 3.00 2.50 Est. 0.040 Est. 0.040 0.038

135 3.00 2.062 3.00 Est. 0.040 Est. 0.040 0.031

81508 20 3.00 2.312 3.00 Est. 0.030 Est. 0.030 0.060

81513 5 2.125 2.625 3.25 Est. 0.020 Est. 0.020 0.023

81551 111 3.00 2.875 3.00 Est. 0.035 Est. 0.035 0.039

31562 126 2.812 6.00 2.812 Est. 0.020 Est. 0. CIO 0.035

81618 4 3.00 5.00 5.00 Est. 0.020 Est. 0.020 0.035

81710** 58 3.375 N.A. 3.375 0.064 0.0.064 0.040

90001 C** 108 N.A. 2.25 2.25 N.A. Est. 0.040 0.030

120 N.A. 2.25 2.25 N.A. Est. 0.040 0.033

** Included In 9/10/76 Docket No. 76-12W, Appendix III submittal,

a. Estimated by RI to be arch bum, not gas pocket.

N.A. -Not available.

* See P. 90 for recommended correction factor.

+ See Fig. 48 for recommended correction factor.

1S9
Table 24. Defect sizes of all non-planar defects (continued).

Weld #

LENGTH

Radiographic AlyesRa SWRI Final

Location (In) (In) Docket

(9/10/76)

(In)

DEPTH

SWRI Final

(In) Docket

(9/10/76)

(In)

RI++

(in)

NON-PLANAR DEFECTS

Hollow Bead

81210 122 2.25 2.75 2.75 Est. 0.020 Est. 0.020 0.045

81255 107 2.50 2.375 2.375 Est. 0.030 Est. 0.030 0.069

81307 10 2.625 3.625 2.125 0.013 Est. 0.020 0.043

25 2.625 2.375 0.050 Est. 0.020 Est. 0.020 0.040

81316 11 2.375 3.00 2.375 Est. 0.020 Est. 0.020 0.040

25 2.375 2.375 1.75 Est. 0.020 Est. 0.020 0.030

81319 8 2.125 3.00 2.125 0.008 0.008 0.038

81346 101 0.50 1.50 1.50 Est. 0.020 Est. 0.020 0.032

81350 26 4.00 N.A. 2.25 N.A. Est. 0.030 0.047

125 4.00 7.625 2.25 Est. 0.030 Est. 0.030 0.041

81396 23 2.50 2.75 2.50 N.A. Est. 0.030 0.048

81397*

** 27 N.A. 2.875 2.75 Est. 0.045 Est. 0.025 0.045

81406 20 2.50 N.A. 2.125 Est. 0.020 Est. 0.020 0.032

81407 12 2.25 N.A. 2.25 0.013 Est. 0.020 0.030

26 2.25 N.A. 2.25 Est. 0.020 Est. 0.020 0.035

81431 26 2.625 N.A. 2.625 Est. 0.020 Est. 0.020 0.033

81432 30 0.375 N.A. 2.375 Est. 0.025 Est. 0.025 0.040

81445** 24 N.A. 2.062 2.062 Est. 0.040 Est. 0.040 0.044

81451 24 2.50 2.125 2.50 Est. 0.035 Est. 0.035 0.031

81473 12 2.125 3.25 3.25 Est. 0.020 Est. 0.020 0.037

81500** 135 2.125 2.375 2.375 Est. 0.045 Est. 0.045 0.034

81515 14 N.A. 2.75 2.438 Est. 0.020 Est.0.020 0.048

30 N.A. 2.50 2.312 Est. 0.020 Est. 0.020 0.038

81522** 33 2.375 2.562 2.562 Est. 0.040 Est. 0.040 0.030

81526** 98 0.50 2.25 2.25 Est. 0.040 Est. 0.040 0.036

81532 48 2.188 2.625 2.625 Est. 0.020 Est.0.020 . 0.041

81535 12 2.125 2.875 2.875 Est. 0.020 Est.0.020 0.039

141 2.125 2.375 2.375 Est. 0.020 Est.0.020 0.045

81537** . 120 2.25 2.688 2.688 Est. 0.050 Est. 0.050 0.032

81538** 34 2.375 2.062 2.375 Est. 0.040 Est. 0.040 0.038

81540** 124 2.50 2.938 2.938 Est. 0.050 Est.0.050 0.045

81545** 18 2.25 2.75 3.312 Est. 0.045 Est. 0.045 0.038

30 2.25 3.312 2.75 Est. 0.045 Est. 0.045 0.044

81 548** 30 2.25 2.50 2.50 Est. 0.040 Est. 0.040 0.043

81549** 96 2.25 2.25 2.25 Est. 0.045 Est. 0.045 0.048

81550** 19 2.50 2.625 2.625 Est. 0.040 Est. 0.040 0.041

81552** 28 2.188 2.938 2.938 Est. 0.040 Est. 0.040 0.051

81597 45 2.812 2.50 2.50 Est. 0.020 Est.0.020 0.054

81625 135 2.625 3.00 3.00 Est. 0.020 Est.0.020 0.022

81626 55 2.75 2.875 2.625 Est. 0.035 Est. 0.035 0.030

135 2.75 2.625 2.875 Est. 0.025 Est. 0.035 0.037

81627 103 2.188 2.50 2.188 Est. 0.020 Est.0.020 0.048

111 2.188 2.00 2.50 Est. 0.030 Est.0.020 0.013

81630 15 2.50 2.031 2.50 Est. 0.035 Est. 0.035 0.025

81635^

125 2.50 3.00 3.00 Est. 0.035 Est. 0.035 0.028

142 2.125 2.031 2.125 Est. 0.035 Est. 0.035 0.035

81 639

z 125 2.125 2.031 2.125 Est. 0.035 Est. 0.035 0.038

81642** 15 2.188 2.50 2.188 Est. 0.040 Est. 0.040 0.034

81643** 12 2.50 2.875 2.875 0.023 0.023 0.038

81644 20 2.062 4.375 2.062 Est. 0.020 Est.0.020 0.039

81645 25 3.00 2.625 2.75 Est. 0.020 Est.0.020 0.039

122 3.00 3.625 3.625 Est. 0.020 Est.0.020 0.039

81646** 48 3.25 2.375 3.25 Est. 0.045 Est. 0.045 0.041

81667 42 2.625 2.50 2.625 Est. 0.020 Est.0.020 0.025

54 2.625 2.50 2.625 Est. 0.020 Est.0.020 0.020

81678 48 2.50 2.50 2.50 Est. 0.035 Est. 0.035 0.033

81688 51 2.25 3.50 3.50 Est. 0.020 Est.0.020 0.035

81 707** 120 2.312 2.156 2.312 Est. 0.045 Est. 0.045 0.055

y. Incomplete penetration (Hi/Lo) also reported at same location by RI and SWRI.

z. Estimated by RI and SWRI to be hollow head, not incomplete penetration.

** Included In 9/10/76 Docket No. 76-12W, Appendix III submittal.

N.A.Not available.

See P. 90 for recomnended correction factor.

See Fig. 48 for recommended correction factor.

170
'
Table 24. Defect sizes of all non-planar defects (continued).

Weld 4 Radiograph

Location

ic

LENGTH

Alyeska SWRI

(in) (in)

Final SWRI

Docket (in)

(9/10/76)

Liiii.-

DEPTH

Mna I Docket j.

(9/10/76) RI

'

(in) _(j_n_L

NON-PLANAR DEFECTS

Spherical Porosity, Cluster Porosity

81138 35 0.125 0.047 0.125 N.A. N.A. 0.023

36 0.125 0.047 0.125 N.A. N.A. 0.005

42 0.125 0.047 0.125 N.A. N.A. 0.028

43 0.125 0.047 0.125 N.A. N.A. 0.040

81142 47 0.125 0.075 0.094 0.014 0.014 0.050

69 0.125 0.075 0.125 0.007 0.014 0.045

81203 40 0.125 0.25 0.125 Est. 0.015 Est. 0.015 0.050

68 0.125 0.375 0.125 Est. 0.015 Est. 0.015 0.068

81252 58 0.125 0.050 0.125 0.065 0.076 0.078

81258 47 0.094 0.075 0.094 0.032 0.032 0.048

81299 32 0.219 0.062 0.219 0.034 0.034 0.054

81300 76 0.125 0.075 0.125 0.062 0.062 0,070

81301 115 N.A. 0.175 0.219 6.084 0.084 0.082

81311 59 0.094 0.100 0.094 0.009 0.009 0.055

81332 50 0.125 0.075 0.125 0.023 0.023 0.098

81343 88 N.A. N.A. 0.188 N.A. N.A. 0.073

81364 108 0.188 0.094 0.188 N.A. N.A. 0.055

81399 35 0.156 N.A. 0.156 N.A. N.A. 0.042

81442 97 0.125 0.075 0.125 0.035 0.035 0.067

81522** 72 N.A. 0.188 0.188 Est. 0.015 Est. 0.015 0.057

81544 108 0.125 0.062 0.094 N.A. N.A. 0.035

109 0.125 0.047 0.125 N.A. N.A. 0.045

81572 . 50 0.125 0.075 0.125 N.A. N.A. 0.052

81 590R

a 143 0.125 N.A. 0.125 N.A. N.A. 0.027

81616 70 0.125 0.100 0.125 N.A. N.A. 0.050

81619 98 0.094 0.075 0.094 N.A. N.A. 0.042

81660 62 N.A. 0.062 0.062 N.A. N.A. 0.035

70 N.A. N.A. 0.062 N.A. N.A. 0,032

81673 41 0.625 0.25 0.625 Est. 0.010 Est. 0.010 0.028

46 0.625 0.031 0.625 Est. 0.010 Est. 0.010 0.068

** Included in 9/10/76 Docket No. 76-12W, Appendix III submittal,

a. Estimated by RI to be arc burn, not spherical or cluster porosity.

N.A. m Not available.

+ See P. 90 for recommended correction factor.

++ See Fig. 48 for recommended correction factor.

171
Table

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f
Table 26. Defect-size-data discrepancies,

Special Considerations

Lack of defect-size data

from Alyeska or SWRT to

support Docket data.

Lack of defect-size data

in Docket submittal

.

Alyeska or SWR-I defectsize data greater than

Docket data.

Planar Defects

length depth

Weld - or Number nf Defects

81706

81448

3

81706

Mon-Planar Defects Arc R'j rn

s

length depth length depth

2 24 20 1

1 17 3 4

81 585 -GP 81321 -GP

81 586-GP 81474-ES

81628-GP 81 307-HB

81135-ES 81 397 -HB

81474-ES 81 407-HB

81476-ES

81562-ES

81 307-HB

8131 6 -HB

81350-H8

81397-H8

81406-HB

81 545-KB

81597-HB

81 626-HB

81627-HB

W15TRX-GP.

81644-HB

81645-HB - 81142-SP, CP

81311-SP, CP • 81544-SP, CP

o .

2? = Gas pocket

:S = Elongated slag

'3 = Hoi 1 c w bead

CP - Spherical porosity, cluster porosity

173
9. CONCLUSIONS

The material properties needed for the fracture mechanics analysis

of the trans-Alaska oil pipeline base material and welds, have been

evaluated. Because of the toughness, yield strength and applied stress

in this situation, fracture mechanics is the best available methodology

to predict failure and assess girth welds containing defects.

There is no exact method for analyzing fracture conditions involving

substantial plasticity. Further, in the region of interest for defect

sizes of this project, there are very few experimental results to substantiate any given fracture model. Therefore, several methods were used

to calculate allowable defect sizes.

The conclusions are based on allowable defect size curves calculated

using maximum credible operating stresses and low rates of loading that

are expected to exist in the main line of grade X-65 pipe. Somewhat

higher stresses may be expected in sections of X-70 grade pipe and in the

fatigue spectrum stress near pump stations due to pressure suroes.

Allowable defect sizes calculated using these higher stresses will be

lower than those shown in this report. At the expected loading rates in

the pipeline, the welds behave in a non-brittle manner at all anticipated

operating temperatures. In the unlikely case that a crack of sufficient

size is initiated in a hard region, such as an arc burn,

propagation or arrest of this crack may be governed by the dynamic

fracture toughness of the metal, which is somewhat less than the static

fracture toughness used to calculate the allowable defect size curves in

this report.

For the stress levels in the X-65 grade of main line pipe and for

the expected low rates of loading, the fracture mechanics analyses

described in this report indicate the following:

1. The calculated allowable defect size curve and analysis

submitted by Alyeska for nonplanar defects is considered

reasonable.

2. The calculated allowable defect size curves submitted

by Alyeska for planar and arc burn defects are considered

reasonable, except for the region of small defect lengths

where Alyeska curves allow unlimited defect depth. The

other calculated allowable defect size curves developed

in this report show a maximum allowable defect depth

even for short defect lengths.

174
When good control is maintained over radiographic variables, defect

sizes can be determined with reasonable accuracy. If more precise measurements are required, other nondestructive testing methods such as ultrasonics, should be used to supplement radiographic testing.

In the case of the pipeline girth welds, the radiographs were

taken to meet the present workmanship codes. Precise radiographic

control is not required and, therefore, it follows that defect size

measurements of high accuracy may not be possible. Sufficient

radiographic control must be maintained, particularly in terms of

radiation energy and film processing, to permit the establishment of

a reasonable relationship between radiographic film density and defect

depth (or steel thickness). The available data from the pipeline field

radiographs show a much greater variability for the radiographic

contrast of the supposedly known thickness of the radiographic

penetrameter and shim on the pipe wall than can be accounted for with

all the considered parameters. This variability makes any depth

measurement subject to error.

Three methods were used to measure defect depths from the Alyeska

radiographs. Southwest Research Institute (SWRI) used both a densitometer

method and a visual comparison technique. Rockwell International (RI)

employed a visual comparison method involving only the weld and defect

area. In both SWRI methods, comparisons were made to the film contrast

for the shim-penetrameter image on the pipe wall.

For the SWRI method in which densitometer measurements were used,

the film contrast variability is less important since defect depth

was determined by a ratio calculation which makes use of the penetrametershim image contrast. The errors then will depend on the accuracy with

which the densitometer measurements can be taken. The difficult measurements are those involving the defect image and its contrast with the

weld. The densitometer reading of the defect will vary depending upon

(1) the ratio of the defect image density area and the area of the

densitometer aperture, and (2) the degree of density uniformity in

the aperture area. Since each defect may be different, it is not

possible to assess the accuracy of this method in a positive manner.

Some determinations of defect depth using this densitometer method are

given as illustrations of the changes produced in the defect depth

calculation for typical variations in densitometer readings.

The SWRI visual method is difficult to assess since no controlled

experiment has been done. It is our recommendation that the SWRI visual

depth measurements be used only as a means of checking other depth

measurements such as those being made by RI. For reasons discussed

in this report, we believe the visual method used by SWRI will underestimate flaw depth. The degree of that underestimation has not been

assessed.

175
The visual depth measurements made by RI have been analyzed. Based

on laboratory x-ray calibration data, the RI method tends to underestimate

the defect depth in increasing amounts for defects of greater depth.

However, these corrections are based on laboratory x-ray films. For

the Alyeska films, there are many field radiographs whose contrast falls

outside the contrast limits predicted by the film response and expected

radiographic procedure. This poses a dilemma since the reason for this

change in film contrast is not known. If the contrast variations are

due to extreme changes in x-ray voltage or changes in film processing,

an accurate depth determination by the RI method may not be possible.

However, if the contrast changes were caused by shim thickness variations

or densitometer reading errors, the accuracy of the RI method may not

be affected.

In support of the RI measurements there is information presented to

demonstrate that the RI depth measurements tend to be larger than those

made by the SWRI densitometer method. This is important because these two

sets of measurements were taken by independent approaches and because the

SWRI measurements should not be influenced by film contrast variations.

It is also demonstrated that RI measurements of several natural defects in

welds furnished to NBS by Alyeska, showed correlation with laboratory

radiographic measurements to a one a limit of 47%. Since the RI method

is the only one used to measure depth of all weld defects still under

consideration for waiver, DoT may wish to expand these comparisons in

order to set limits on the RI method.

The length measurements can, in most cases, be made to accuracies

of 0.010 in. Exceptions involve defects which gradually diminish in

depth so that the ends of the defect are difficult to observe on the

radiograph; defects in this category include incomplete penetration,

incomplete fusion and elongated slag. An independent length measurement

should be provided for those defect types. In addition, some of the

large discrepancies noted in defect length should be explained and/or

remeasured. These are probably caused by differences in radiographic

interpretation.

For arc burns, a method is described for determining effective

arc burn depth based on the measurement of arc burn diameter or width,

a measurement possible from the radiograph.

In summary, the fracture mechanics analyses provide a method to

evaluate the performance of the pipeline girth welds in terms of

calculated allowable defect sizes. Defect lengths can be measured

accurately, although some discrepancies have been noted. Arc burn

diameters or widths can be measured and a method for converting that

measurement to a depth is described. For other defect types, the

depth measurement becomes more difficult and is subject to possible

error as described in this report.

176
N 85- 1 14 A IMV. T.,,1

u.5. 3 'ot. o ' ::m» BIBLIOGRAPHIC Data

SHEET

I. PUIU.ICATION OK REPORT NO. NBSIR 76-11542. Gov't Accession

No.

3. Recipient’* Accession No.

4. TITLE AND SUBTITLE Consideration of Fracture Mechanics Analysis and

Defect Dimension Measurement Assessment for the

Trans-Alaska Oil Pipeline Girth Welds

5. Publication Date October 18, 1976

6. Performing Organization Code

7HaLro?S‘

:> 'Berger, John H. Smith, Editors8- Perlovmtag Organ. Repore No.

. PERFORMING ORGANIZATION NAME AND ADORESS

NATIONAL BUREAU OP STANOAROS

DEPARTMENT OP COMMERCE

WASHINGTON. O.C. 20214

10- Proiecc/Task/^ork Unu No. 3000380

11* Contract/Granc No.

12. Sponsoring Organization Name and Complete ‘Address (5 (reef. City, State, ZIP)

Department of Transportation

2100 2nd Street, SW

Washington, 0. C. 20590

13. Type of Report & Period Covered Final

14. Sponsoring Agency Code

IS. SUPPLEMENTARY NOTES

14. ABSTRACT (A 200’word or Jess factual summery ol most st fruftcini information. It document includes a significant bibliography or literature surwey, mention it here.) In anticipation of a request for waivers on defective girth welds in the

Trans-Alaska oil pipeline, OoT requested assistance from the National

Bureau of Standards (N8S) in evaluating the fracture mechanics analysis

and the nondestructive evaluation (NOE) methods used to detect and

determine dimensions of weld defects.

N8S measured the required mechanical properties of the weld metal, developed and evaluated fracture mechanics methods to determine the allowable defect sizes, and evaluated various methods of measuring the size

of defects present in the welds from existing field radiographs.

Results of this Investigation show that the fracture mechanics analysis

can be used to determine the allowable defect sizes concerning the integrity of the pipeline, but that these analyses have not been ex- perimentally verified at this time. Defect dimensions can be determined with sufficient accuracy to be useful In the fracture mechanics analyses

if the radiographs are made under carefully controlled conditions. If

the radiographs are not made with close control, the accuracy of the

defect sizes may not be sufficient to permit their use in establishing

allowable defect sizes.

17. KEY TORD5 (sis to twelve entries; alphabetical order; capitalize only die first letter of the first key word unless a proper name; separated by semicolons

)

Fracture control; Fracture mechanics; Mechanical properties; Nondestructl evaluation (NDE); Radiography; Welding

ve

IS. AVAILABILITY j Unlimited XT For Official Distribution. Do Not Release :o NTIS

Order From Sup. of Doe., U.S. Government Printing Office lashmgton, O.C. 2d *02, SO O-t. Vo. GU*

Order From National Technical Information Service (NTISI

Springfield, Virginia 221*1

19. SECURITY CLASS

(THIS REPORT)

UNCLASSIFIED

20. SECURITY CLASS

(THIS PAGE)

UNCLASSIFIED

21. NO. OF PAGES

310

22. Pric

1.11

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