I llll I iSf STaN0 s tech

All! 04 0 4 fc, 4 5 1,

NATIONAL INSTITUTE OP STANDARDS &

TECHNOLOGY

Research Information Center

Gaithersburg, MD 20899

NBSIR 83-1695

A

FITNESS-FOR-SERVICE CRITERIA FOR PIPELINE GIRTH WELD QUALITY

National Bureau of Standards

U.S. Department of Commerce

Boulder, CO 80303

November 1983

NBS1R 83-1695

FITNESS-FOR-SERVICE CRITERIA FOR PIPELINE GIRTH WELD QUALITY

R. P. Reed

M. B. Kasen

H. I. McHenry

C. M. Fortunko

D. T. Read

Fracture and Deformation Division

National Measurement Laboratory

National Bureau of Standards

U.S. Department of Commerce

Boulder, CO 80303

Prepared for;

Materials Transportation Bureau

U.S. Department of Transportation

400-7th Street SW

Washington, DC 20590

and

Welding Research Council

United Engineering Center

345 E. 47th Street

New York, NY 10017

U.S. DEPARTMENT OF COMMERCE, Malcolm Baldrige, Secretary

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

CONTENTS

Page

ABSTRACT 1

EXECUTIVE SUMMARY 2

1. INTRODUCTION 8

2. DEVELOPMENT OF ALLOWABLE FLAW-SIZE CURVES 14

2.1 Introduction 14

2.2 Fracture Mechanics 16

2.3 Model Evaluation Tests for Surface Cracks in Tension 38

2.4 Model Verification Tests on Large-Diameter Pipes 71

2.5 Allowable Flaw-Size Curves 81

2.6 Summary ' 94

3. SHARP-FLAW DIMENSIONING 96

3.1 Introduction 96

3.2 Inspection, Configuration, and Operation 101

3.3 Theory of Horizontally Polarized Plate Wave Propagation.... 104

3.4 Theory of Scattering of Horizontally Polarized Shear Waves

by Elongated Flaws 108

3.5 Experimental Verification 113

3.6 Detectability Limitations: Long Flaws 116

3.7 Detectability Limitations: Short Flaws 131

3.8 System Performance Limitations and Standards of

Acceptability 139

3.9 Inspection Protocol 142

3.10 Summary 146

4. SIGNIFICANCE OF BLUNT FLAWS 148

4.1 Introduction 148

4.2 Experimental Procedures 148

4.3 Low-Cycle Fatigue Results 160

4.4 Fractographic Analysis 170

4.5 Discussion 172

4.6 Summary 180

5. INTRINSIC LIMITATIONS ON DIMENSIONS OF BLUNT FLAWS 182

5.1 Introduction 182

5.2 Experimental Procedures 183

5.3 Results 184

5.4 Summary 188

i

6. PRACTICAL LIMITATIONS ON BLUNT FLAW CONTENT USING

RADIOGRAPHIC INSPECTION 189

7. CONCLUSIONS 193

8. RECOMMENDATIONS 197

8.1 Implementation of Inspection Methods 197

8.2 Implementation of Fitness-for-Service Criteria 202

8.3 Future Research 207

9. ACKNOWLEDGMENTS 212

10. NOTATION 213

11. REFERENCES 218

12. BIBLIOGRAPHY OF RELEVANT NBS PUBLICATIONS 225

13. APPENDIXES 229

Appendix A - Details of Calculation of Allowable Flaw-Size

Curves 229

Appendix B - General Research on Corrosion 246

Appendix C - Radiographic Evaluation of the Significance of

Weld Defects 267

Appendix D - Defect Characterization and Dimensioning of Cracks

in Welds by the Ultrasonic Diffraction Method 278

Appendix E - Flaw Characterization 300

Appendix F - Ultrasonic Examination of Girth Welds in 40-Inch,

X-65 Line Pipe Specimens 326

Appendix G - J-Integral Analysis of Surface Cracks in Pipeline

Steel Plates 363

Appendix H - Metal! ographic Analysis 370

ii

FITNESS-FOR-SERVICE CRITERIA FOR PIPELINE GIRTH WELD QUALITY

R. P. Reed, M. B. Kasen, H. I. McHenry, C. M. Fortunko, and D. T. Read

ABSTRACT

Criteria have been developed for applying f i tness-for-service analyses to

flaws in the girth welds of the Alaska Natural Gas Transmission System

pipeline. A critical crack-opening-displacement elastic-plastic fracture

mechanics model was developed and experimentally verified. Procedures for

constructing flaw acceptence criteria curves based on this model are provided.

A significantly improved ultrasonic method for detecting and dimensioning

significant weld flaws was developed and demonstrated on pipeline sections.

The probability of crack initiation from blunt flaws was shown to be very low

under severe low-cycle fatigue. Suggestions are offered for technical

implementation of field inspection procedures and for practical implementation

of the flaw acceptence criteria.

1

EXECUTIVE SUMMARY

The utility of f itness-for-service criteria for judging the need to

repair pipeline girth welds containing flaws exceeding that permitted under

prevailing workmanship standards was demonstrated during construction of the

Trans-Alaska Pipeline System (TAPS). Here, fracture mechanics principles

provided a quantitative technical basis for decision making, eliminating the

need for unnecessary repair of many welds while maintaining assurance of the

structural integrity of the pipeline.

The National Bureau of Standards (NBS) provided the Department of

Transportation (DOT) with the technical information for implementing the TAPS

fi tness-for-service approach. The DOT recognized that the general use of

fitness-for-service criteria in general pipeline construction would be

beneficial. In 1978 NBS was requested to initiate a research program leading

to this objective. However, soon after the work began, the DOT requested that

the NBS program concentrate on developing criteria that could be implemented

during construction of the proposed Alaska Natural Gas Transportation System

(ANGTS), which was then in the planning stage. It was the intention, then, to

redirect the program to generalization of the ANGTS research and, indeed, a

portion of the research was redirected in 1981. However, the present report

emphasizes the specifics of the ANGTS line rather than the case of pipelines

in general

Fracture mechanics analysis was used in this research. Fracture mechanics

permits mechanical understanding of the relationship between the maximum

stress and strain imposed on a pipeline during its lifetime, the pipeline and

weldment properties, and flaws of different sizes and configurations. This

understanding required development and verification of a model relating these

2

parameters to weldment performance. Although the basic principles of fracture

mechanics are well understood and widely implemented in the aerospace industry,

their application to pipeline fabrication was complicated by the relatively

high ductility of the materials used in their construction. It was necessary

to understand crack initiation and propagation in the elastic-plastic range

instead of linear-elastic failure typical for materials and structures

displaying lower ductility.

It was assumed that a pipeline would be subjected to predictable

stresses, that it would be constructed from API line-pipe steel having known

minimum properties, and that the pipe would be welded by a process producing

predictable minimum weldment strength and toughness. This design and

materials information was then fed into the theoretical model to develop a

series of curves relating the combination of maximum flaw length and

through-wall depth permissible for the maximum applied load (allowable flawsize curves). Practical use of such curves presupposes the ability to detect

and measure girth-weld flaws, and methods to accomplish this had to be

developed.

Research was directed into three areas:

1. Refinement and experimental verification of the fracture mechanics

model linking flaw size (length and through-wall depth) to pipeline stresses

and weldment strength and toughness . Analytical predictions were compared

with experimental results obtained on surface-cracked tensile panels and

large-diameter pipe tests. The critical crack-opening-displacement (COD)

model (used in the TAPS evaluation) was modified to optimize agreement between

analysis and experiment. Modifications included: superposition of the

elastic and post-ligament-yield solutions for crack-tip-opening displacement,

use of a strip-yield plastic-zone correction on both crack length and crack

3

depth, and use of a curvature correction for pipes. In summary, a fracture

analysis model that relates allowable flaw sizes to the applied stress level

and the weldment strength and toughness was developed, verified experimentally, and used to calculate allowable flaw sizes for proposed operating

conditions of the ANGTS pipeline.

Procedures for the use of elastic-plastic fracture mechanics for the

derivation of allowable flaw-size curves are in the early stages of development.

Further work is recommended to increase confidence in the analytical results.

2. Development of a practical method of determining the through-wall

flaw dimension . The TAPS experience showed that estimation of dimensions from

conventional radiographs was so inaccurate that very large safety factors were

required. This resulted in fitness-for-service criteria that were far more

restrictive than necessary. Initially, efforts were made to refine the

densitometry method of assessing flaw through-wall dimensions. Although some

improvement was realized, it became evident that this approach was

intrinsically deficient in its ability to measure sharp flaws, such as lack of

fusion or cracks, which are the most significant in a fitness-for-service

analysis. A more promising approach was provided by ultrasonic methods, which

are intrinsically sensitive to sharp flaws. Although ultrasonic inspection

systems based on high-frequency piezoelectric transducers have long been used

for flaw detection in welds, incorporation of even the most advanced

automation and signal processing has thus far failed to produce systems

capable of providing the required flaw measuring capability. Therefore, we

tried a new approach — the development of an ultrasonic inspection system

optimized for pipeline use based on electromagnetic transducers (EMATs), which

transmit and receive low-frequency shear/horizontal (SH) waves. Unlike

conventional systems, this approach provides a signal amplitude that increases

4

monotonical ly with through-wall flaw depth. The system can be gated to ignore

flaws of noncritical dimensions and to trigger an alarm automatically when

flaws of significant size are present. Coupling agents are not required

between the transducer and pipe, and reliability is not affected by geometric

factors such as the weld reinforcement. Its accuracy is comparatively insensitive

to operator skill, and a permanent record is produced. The theory underlying

flaw detection and sizing by this new method has been developed and validated

by the inspection of welds containing a variety of artificial and real flaws

at different depths and orientations. The necessary transducer arrays and

associated signal processing equipment have been developed. Field trials and

assessment of reliability are needed to confirm the usefulness of this system

as an adjunct to the f itness-for-service approach to girth-weld quality

assessment.

3. A convincing demonstration of the probably innocuous nature of blunt

flaws, such as porosity, slag, and arc burns . We found that blunt flaws

(porosity, slag) and arc burns do not decrease weldment strength and fatigue

life. Subjecting pipeline welds containing large quantities of such flaws to

severe low-cycle fatigue at -2°C had little effect on crack initiation.

However, should it be desired to include blunt flaws in a fracture mechanics

analysis, metallographic studies of welds containing porosity and slag showed

that their depth is limited to the depth of one weld pass, eliminating the

necessity for through-wall dimensioning. In metallographic studies during the

TAPS program, arc-burn width was related to arc burn depth, thus providing an

easy size characterization procedure for such flaws.

On the basis of these results, three technical options are suggested for

implementing field inspection oriented toward f itness-for-service. 1) The use

of ultrasonic methods for primary inspection combined with traditional radiography for assessing workmanship and for materials and procedure control

5

appears feasible and is the most technically desirable. 2) Conventional

radiography could be used as the primary inspection method, using a simplified

ultrasonic system for measuring flaws. We believe that the EMAT system

developed in this work appears to be the most promising inspection method for

either of these two options; however, other systems of equivalent performance

could be used should they become available. 3) In the event that ultrasonic

inspection systems of demonstrated performance are not available at the time

of construction of a given pipeline, radiography must be relied on

exclusively. In this case, f i tness-for-service criteria can still be used to

assess blunt flaws. The demonstrated inherent limitations on the height of

such flaws would very much simplify this effort by eliminating the need for

complex flaw-measuring procedures. Even under this least desirable option, a

substantial reduction in remedial welding can be anticipated since the

majority of detectable flaws would be of blunt configuration.

We, suggest two ways in which the overall f itness-for-service approach

might be implemented in pipeline construction: A generalized approach could

prequalify designated sections of the line by establishing maximum allowable

flaw sizes based on limits placed on weld-metal toughness, operating stresses,

and other relevant variables. A specific-site approach might be used wherein

the particular conditions existing for a given weldment are used to calculate

safe flaw sizes.

We expect this report to contribute to a rational, technically defensible

approach to assessing girth-weld quality, which will provide the fabricator

with the maximum flexibility to accommodate his particular needs while maintaining pipeline safety. Additional research has been recommended to assist

in generalization of the program's results.

6
v
Specific support for this work was provided by the Materials Transportation

Bureau, DOT, and by the Welding Research Council. The Northwest Pipeline

Company provided materials and consultation. Consultation was also provided

by Nova, an Alberta Corporation.

7

1. INTRODUCTION

Extensive repair to girth welds on the Trans-Alaska Pipeline System

(TAPS) resulted from the discovery that a large number of completed and buried

welds failed to meet the Department of Transportation (DoT) standards 1 related

to girth welds of liquid pipelines, which incorporate the workmanship criteria

2

of the 15th edition of API Standard 1104. The costs of this remedial welding

were very high because pipe sections buried under permafrost and rivers at

remote locations had to be excavated or looped.

At that time it was recognized that many of the flaws outside workmanship

standards were unlikely to endanger the safety of the line and that repairs

might do more damage to the integrity of the line and to the environment than

if the flaws were left unrepaired. Unfortunately, there was no established

technical basis for such a decision. Repairs were therefore begun.

At the same time, fracture mechanics studies were begun to establish a

technical foundation for assessing flaw significance on a f i tness-for-service

3 4

basis as an alternative to workmanship criteria. The results of this work *

brought about federal acceptance of this approach. Consequently, the amount

of TAPS repair welding was reduced and a substantial amount was saved in

5 unnecessary repair. The DOT concluded in 1976 that "fracture mechanics

analysis is acceptable as a basis for granting exemptions from existing

standards in appropriate circumstances, if such analysis produces a convincing

and conservative estimate of structural integrity.

It was anticipated that the need for long-distance transport of oil and

gas would lead to additional use of fi tness-for-service analysis in pipeline

8
V
construction. The proposed Alaska Natural Gas Transportation System (ANGTS)

was to operate at lower temperatures than the TAPS line and would be entirely

buried, placing additional demands on system performance. Furthermore,

since adoption of the API 1104 workmanship standards upon which the DOT

standards are based, pipeline wall thickness and diameter have increased

significantly, increasing the complexity and cost of weld repair. It was

evident that much additional work was required before the fitness-for-service

concept could be efficiently used to facilitate construction of a safe

pipeline at minimum cost.

Research at NBS was initiated in 1978 to address the generic problems

associated with possible amendment of the DOT liquid and gas pipeline federal

regulations to include a fitness-for-service alternative. This included

research on fracture mechanics, radiography, blunt flaws, and corrosion of

pipelines. In 1979, the apparent imminent start of construction of the ANGTS

gas pipeline led the DOT to request a change of focus of the NBS program to

specifically address the operating conditions and materials anticipated for

that line. This report, therefore, concentrates primarily on developing

fitness-for-service criteria for the ANGTS line.

At the time of the TAPS experience, two fracture mechanics models were

3 4

proposed * to assess the conservatism of the Draft British Standards

methodology^ (later published officially by the British Standards

g Institution ), which was used in the Alyeska waiver request to DOT. One was

based on elastic-plastic fracture mechanics principles [critical crack-tipopening displacement (CTOD) model]; the other considered only plastic

instability of the uncracked ligament. Both of these newer models proved less

conservative than the Draft British Standards.

9

In this study we have focused on the extension of the critical COD model.

Modifications relating to crack-tip physics and mechanics were developed and

measurements were performed, leading to an experimentally verified elasticplastic fracture mechanics model. Consideration of crack growth from fatigue

stresses, included in the TAPS assessment, has not been addressed here.

Except near pump statons, cyclic stresses in the AN6TS line were not considered

significant. Enhanced crack growth from environmental interactions, such as

stress corrosion, also was not considered in the AN6TS case, because of this

origin was not consdered likely.

There are serious deficiencies in the use of radiography as an inspection

tool in a fitness-for-service analysis. The important flaw dimension

affecting fracture is the through-wall depth of elongated, sharp flaws. But

radiography provides direct information only on the projected dimensions of

the flaw normal to the plane of the weldment. It is theoretically possible to

relate the change in radiographic density caused by a flaw to the depth of the

3 4

flaw; This approach was used in the TAPS analysis. * However, it became

evident that the combination of uncertainties in the conditions under which

field radiographs are taken and processed, human bias, and the technical

subjectivity of such determinations led to excessive conservatism in

estimating flaw dimensions.

The majority of the repairs remove flaws of a blunt configuration--porosity,

slag, and arc burns. This reflects, at least partly, the sensitivity of

radiography to such flaws and its relative insensitivity to sharp flaws, such

as lack of fusion or incomplete penetration. Yet, from a fitness-for-service

approach, sharp flaws are more significant, since they produce the highest

stress intensities and, consequently, have the highest probability of initiating

fracture. A better inspection system for detecting and dimensioning flaws in

girth welds was required, preferably one having optimal sensitivity to sharp

flaws.

10

The significance of blunt flaws in crack initiation was questioned.

Results of a number of studies had suggested that the presence of such weld

flaws is irrelevant to the performance of structures joined with tough weld

metal, but the credibility and the boundary conditions of this assumption had

not been examined for pipeline girth welds.

It was likely, we thought, that the through-wall depth of blunt flaws

would intrinsically be restricted to the depth of one weld pass, allowing that

dimension to serve as a maximum in a fracture mechanics analysis. Although

such a limitation is frequently assumed, a thorough search of the literature

did not reveal the required documentation. A study was therefore conducted on

inherent dimensional limitations of blunt flaws with the purpose of confirming

or refuting this assumption.

It is instructive to compare the results of this study with those of

other pipeline f itness-for-service approaches that have been suggested:

1. The British Standard specifically related to pipeline welding,

g "Standard for the Field Welding of Carbon Steel Pipelines," is currently

under review. One proposal incorporating fracture mechanics analysis has been

recently discussed by Carne and Harrison.^ They proposed a three-tier

approach. Tier one corresponds to workmanship standards similar to those now

in existence, but incorporating minimum toughness requirements. Tier two

consists of very conservative generalized fracture mechanics analysis,

intended to be applied by field engineers not necessarily specialized in

fracture mechanics. Tier three is applied by a specialist to a specific site

and labeled engineering critical assessment (ECA). Inspection methodology was

not discussed in this approach.

2. Glover and Coote^ pointed out that the British Standards

Institution fracture mechanics guidelines are also being used for ECA of

segments of the new large-diameter gas pipelines in Canada. Their assessment

11

showed that radiographic inspection could not properly and consistently detect

sharp flaws. They found that planar flaws not detected by radiography were

consistently detected by high-frequency ultrasonic techniques, but that sizing

was unreliable. They suggested that both radiography and ultrasonics methods

be used, assuming that flaw depths are restricted to the thickness of one weld

pass. The British Standards Institution Guidelines is therefore applied

on the basis of flaw length and knowledge of the applied axial strain and

weldment toughness. Note that this approach is limited to girth welds made by

an automatic process.

3. The Japan Welding Engineering Society has recently published their

first report, "Method of Assessment for Defects in Fusion-Welded Joints With

12

Respect to Brittle Fracture." This report outlines procedures for using a

f i tness-for-service analysis in specific cases for establishment of allowable

sharp flaw sizes in terms of local weldment strain and COD weld toughness.

Inspection methodology is not addressed.

4. A major revision of the American Petroleum Institute 1104 Standard

for Welding Pipelines and Related Facilities (issued in May, 1983) was

13 discussed by Von Rosenberg and Royer. This generalized f itness-for-service

approach requires calculation of maximum axial design stress and strain and

minimum weldment toughness in terms of COD values of either 0.13 mm (0.005 in)

or 0.25 mm (0.010 in). Fatigue crack growth is also considered. Inspection

techniques are left to the individual companies. This revision appears as

Appendix A, 16th Edition, API 1104 Standards. However, at the time of this

report, the 16th Edition has not been incorporated into the Federal gas and

liquid pipeline safety standards.

12

These programs demonstrate international recognition of the need to

supplement welding workmanship standards with fitness-for-service criteria for

pipelines. Reduced pipeline construction time and costs through the elimination

of unnecessary repairs are expected to accrue from this attention.

In summary, our research has focused on the development of a

comprehensive model to calculate allowable flaw sizes and laboratory

experiments to assess the validity of such models, the assessment of the

significance of blunt flaws in girth weld structural integrity, the measurement of inherent blunt flaw size limitations, and the inspection methods for

sharp flaws.

Our report first discusses the development and experimental verification

of an elastic-plastic fracture mechanics model applicable to sharp flaws. We

then address development of a new inspection technique that is optimized for

the fitness-for-service approach. Research related to the significance of

blunt flaws is then presented.

Finally, we recognize that nondestructive inspection has two roles in

pipeline construction. One relates to quality control--the monitoring of

welder and equipment performance and the quality of consumables used in the

fabrication. The other is acceptance or rejection of a weld on the basis of

fitness-for-service. Although this report concentrates on the latter role,

both roles must be considered in devising an effective inspection system.

Possible inspection options for achieving these goals and for implementation

of the fitness-for-service approach are therefore addressed in this report.

A series of Appendixes provide more detailed information on selected

subjects addressed in the main body of this report.

13

2. DEVELOPMENT OF ALLOWABLE FLAW SIZE CURVES

H. I. McHenry, Y. W. Cheng, R. deWit, J. G. Early, R. B. King, and D. T. Read

2.1 Introduction

Allowable flaw sizes for pipeline girth welds are set forth in Section 6

2

of API 1104 on the basis of workmanship considerations; that is, flaw-size

limits are based on quality levels that can reasonably be expected from a

qualified welder using satisfactory materials, equipment, and procedures. In

1976, the U.S. Department of Transportation permitted the use of alternative

weld quality standards to evaluate certain girth welds in the Trans-Alaska

Pipeline System (TAPS). The alternative standards were in the form of

allowable flaw size curves derived on the basis of a fracture mechanics

analysis.

An allowable flaw-size curve is a plot of allowable flaw depth, a, as a

function of flaw length, i, for a planar (crack-like) flaw as shown

schematically in Fig. 1. Dimension a refers to the through-wall dimension of

the flaw projected in the plane of the weld. A point on the curve represents

the maximum flaw size (i, vs. a) that can be tolerated without impairing the

service performance of the pipeline under worst-case operating conditions.

Flaw sizes that fall below the curve are acceptable because they will not

impair performance, and flaw sizes that lie above the curve must be repaired

because they are potentially detrimental.

The purpose of this part of the program was to develop allowable flawsize curves suitable for use in evaluating ANGTS pipeline girth welds. The

specific objectives of the investigation were: 1) to develop a fracture

mechanics analysis model to relate allowable flaw size to applied stress and

14
HM
Flaw

Depth,

Fig. 1 - Schematic of allowable flaw-size curve.

15

weld toughness, 2) to verify the analytical model experimentally, and 3) to

use the model to calculate allowable flaw-size curves for representative

operating conditions and weld properties.

The achievement of the first two objectives required extensive research

in analytical and experimental fracture mechanics. These results are

summarized herein. Many of the results have been published previously, and

these publications are cited where appropriate in this report. The

achievement o* the third objective is based on the results of the fracture

mechanics studies. The allowable flaw-size curves for representative

operating conditions and weld procedures are presented in Section 2.5,

Allowable Flaw-Size Curves, along with the procedures used to calculate then.

Section 2.5 is self-contained, and thus, readers primarily interested in the

resulting curves can skip the preceding sections.

Fatigue and corrosion contributions to crack growth were not considered

significant in the case of the ANGTS pipeline and are not included in this

analysis. Preliminary corrosion research was conducted early in the program

(Appendix B) for generalized pipeline analysis end fatigue was included on the

TAPS analysis.

2.2 Fracture Mechanics

The principles of fracture mechanics provide the technical basis for the

development of allowable flaw-size curves. Fracture mechanics is the study of

the influence o f loading, crack size, structural geometry, and material

properties on the fracture resistance of structures containing cracks. There

are two main ideas in fracture mechanics: First, fracture occurs when the

driving force for fracture, a function of stress and flaw size, exceeds the

resistance of the material to fracture, referred to as the fracture toughness.

Second, fracture toughness is a geometry-independent material property; that

16
,
is, a simple laboratory specimen and a large structure both fracture at the

same critical value of driving force. The second idea is only directly

applicable to linear elastic (brittle) fracture and to certain cases of fully

plastic (ductile) fracture; however, it is a conservative assumption for

elastic-plastic fracture if the notch constraint in the test specimen exceeds

the notch constraint in the structure. Such is the case for pipeline girth

welds when full-thickness single-edge-notch-bend specimens are used to measure

fracture toughness.

Several fracture criteria are used to characterize the driving force for

fracture and the fracture toughness. These include the stress intensity

factor, K, the J-integral, J, and the crack-tip-opening displacement, CTOD.

Since K is a linear elastic parameter, it was not suitable as a fracture

criterion for ANGTS pipeline girthwelds, which must operate safely at stress

levels near the yield strength. However, K calculations were useful for

determining the elastic component of J and CTOD, and thus, K calculations for

surface flaws are discussed in this section. The main emphasis in this

section is on the analytical work related to the development and experimental

verification of elastic-plastic fracture mechanics models based on the CTOD

concept. Research related to the J-integral concept is summarized in

Appendix G.

2.2.1 Linear Elastic Fracture Mechanics

The stress intensity factor, K, for surface cracks has been the subject

of numerous studies. Recently, the fracture committee of the Society for

Experimental Stress Analysis compared several elastic numerical solutions for

14

a surface flaw in a plate under remote tension and bending. The results of

15

Newman and Raju compared favorably with the "best-estimate" results of the

committee. The results of Newman and Raju are expressed in terms of numerical

17

equations that facilitate computation, and thus, their K solution has been

used in the present study. The K solution for semiell iptical cracks in simple

tension is:

for a/£ <_ 0.5, a/t < 1.0, £/W < 0.5, and 0 <_ 0 <_ tt

where a = stress

0 = is the square of the complete elliptic integral of the second kind

= boundary correction factor, given below

a = crack depth

£ = crack length

t = specimen thickness

W = specimen width

0 = angular position along crack front

The geometric variables (a, £, t, and 0) are illustrated in Fig. 2. In this

study, the K at maximum crack depth (0 = 90°) was of interest.

The numerical expression for Q developed by Rawe and used in Ref. 16 is:

K = o(Tra/Q)^F^ ( 1 )

0=1+ 1.464(2a/£)

1 - 65

( 2 )

The boundary correction factor for tension is:

F b

= [b

1 + b

2

(a/t)

2 + b

3

( a/t

)

4

]fw

(3)

where

hj = 1.13 - 0.09 (2a/£)

b

2

= -0.54 + 0.89 (0.2 + 2a/i)

or

b

3

= 0.5 - 1. 0/(0. 65 + 2a/£) + 14(1.0 - 2a/J>r

(4)

(5)

( 6 )

The finite width correction, f , of Newman^ is used:

( 7 )

18

O’

Fig. 2 - Geometry of surface flaw.

19

Applicability of the linear elastic analysis has been extended to

conditions approaching net-section yielding by correcting for the zone of

plasticity that exists at the crack tip. The idea is that the plastic

material at the crack tip strains without carrying the incremental load;

therefore, in the elastic sense, the crack behaves as if it were slightly

18

longer. The Irwin correction is made by adding the radius of the plastic

zone, r , to the apparent crack length, where:

ry

= (K/o

y

) 2

/2n (8)

and a is the yield strength

19 Alternatively, the plastic zone has been modeled by Dugdale as a thin

strip of plastic material extending from the crack tip in the plane of the

crack. The extent of the Dugdale plastic zone, p, is given by

a/ ( a + p) = cos (tto/2c ) (9)

If the higher order terms of the series expansion of the cosine function

are neglected, Eq. 9 reduces to

P

* it(K/o

y

) 2

/8 (10)

The plastic-zone correction, r , is taken in this report to be equal to p/2

and is approximately 23% greater than the Irwin correction.

As plasticity develops, the crack tip blunts. The extent of this

blunting is the CTOD. The CTOD is calculated for the small-scale yielding

case as follows:

CTOD = K

2

( 1 - v

2 )/Eo (11)

where E = Young's modulus

v = Poisson's ratio

o = flow stress = 4(o + o

u

)

o

u = ultimate tensile strength

K = K from Eq. 1

20

2.2.2 Modified Critical -COD Model

The CODs of a surface flaw in a flattensile panel that occur after the

assuming the surface flaw behaves as a through-thickness crack of the same

length. The opening of the equivalent through-thickness crack is reduced by

closure forces equal to the product of the flow stress times the area of the

uncracked ligament.

In the TAPS evaluation, a model developed from Irwin's original concept,

referred to as the critical-COD model, was used to calculate allowable

4

flaw-size curves. However, there were no experimental data to evaluate the

model. In this investigation, the analytical predictions have been compared

with experimental results, and the critical-COD model has been modified to

optimize agreement between analysis and experiment. Modifications include:

superposition of the elastic and post-ligament-yield solutions for CTOD, use

of a strip-yield plastic-zone correction on both crack length and crack depth,

use of a finite-width correction for flat plates, and use of a curvature

correction for pipes.

The critical-COD model treats the surface crack in a plate as a

through-thickness center crack after the ligament has- yielded. The COD at the

middle of the center crack of length, £, in an infinite plate under a remote

stress, o, is given by

For a surface crack, the COD of Eq. 12 is reduced by the remaining ligament.

The effect of ligament depth can be estimated by considering a closing force

distributed over the crack-face area. Assuming the ligament is yielded, the

total closina force, F , is

c

COD = 2WE ( 12 )

F

c

= t(t - a)o (13)

21

where a is flow stress and is estimated as the average of the yield strength

and the ultimate tensile strength, t is the plate thickness, and a is the

crack depth. Distributing this closing force over the area £t gives a closing

stress, o

c

» on the equivalent through-thickness crack of

o

c

= (1 - a/t)o (14)

This closing force opposes the remote stress, o, and the resultant opening of

the surface crack is then

(15)

To account for the crack opening due to the crack-tip plasticity, the

COD = 2jl(o - o

c )/E

effective crack length, which includes the plastic-zone size correction, r ,

is used in place of £. The resulting expression when Eq. 14 is substituted

into Eq. 15 becomes

COD = [o - (1 - a/t)c]2(£ + 2r

y

)/E (16)

where ryis a plastic-zone size correction with a finite width correction and

is evaluated^* as

sin ( tt£/2W ) / s i n (tt£ /2W) = cos [

(

tt/2 ) ( a

' /a ' )] (17)

where

£^ is the length of the crack measured to the ends of the plastic zone,

defined as £ + 4 r , o'/o' = 1 - t/a(l - a/a), and W is the plate width.

Using CTOD as the fracture criterion, the relationship among CTOD, o, and

flaw size is

CTOD = CTOD(elastic) + COD (18)

where CT0D(el asti c) is given in Eq. 11 and COD is given in Eq. 16. The value

of COD is zero when a is less than ligament yield stress, given by (1 - a/t)o.

Comparison of analytical predictions of the critical -COD model with

experimental results discussed in Section 2.3, was done in terms of CMOD,

( i . e . , the COD at x = 0 in Fig. 2) where:

CMOD = CMOD(elastic) + COD (19)

22

where COD is given in Eq. 16 and CMOD(elastic) is evaluated using the

22

elasticity solution of Kobayashi for cracks with aspect ratios U/a) less

than 10:

CMOD = r Gal

""V^

C °Q + 3^T

C

2° + 37r

C °2 (20)

where G = shear modulus and the coefficients Cqq, C

2 q, and Cq2 are governed by

the geometry and loading conditions. The values for Cqq, C

2 q, and Cq

2 are

23

taken from Kobayashi. Solutions to Eq. 20 are shown in Fig. 3.

For long cracks U/a > 10), the Kobayashi solution is not applicable and

21

the elastic CMOD is calculated using King's simplified line-spring model

(discussed in Section 2.2.3).

2.2.3 Derivation of Yielded-Ligament Model from Line-Spring Model

24

The line-spring model was developed by Rice and Levy for elastic

surface cracks. Extension of the model to elastic-plastic surface cracks was

25 26 speculated upon by Rice and carried out by Parks. For the purposes of

this program, the line-spring model provided an analytical framework for

extending the critical-COD model to the case of surface cracks in pipes rather

than flat plates.

The use of the line-spring model was suggested after the validity of the

critical-COD model was confirmed for the case of surface cracks in tensile

panels (see Section 2.3.2). Thus, the same physical assumptions used in the

critical-COD model were introduced into the line-spring model, and the two

models predict the same results for the flat-plate case. However, the linespring model made it possible to account for a number of complexities that

exist for the case of long surface flaws in pipes, specifically, the elastic

displacements of long cracks, the curvature effect, and plasticity development.

On the other hand, these same physical assumptions resulted in considerable

23

NORMALIZED

CMOD,

NORMALIZED CRACK DEPTH, a/t

Fig. 3 - Theoretical solution for CMOD in a surface-flawed tension plate.

G = shear modulus, 6 = CMOD, £ = crack length, a = crack depth,

and t = plate thickness.

24

simplification of the line-spring model. The derivation of the model for the

case of flat plates is discussed in this section and the extension of the

line-spring model to the case of surface cracks in pipes is discussed in the

following section, 2.2.4.

In the line-spring model, the surface crack is treated as being equivalent

to a through-crack of the same length as the surface crack. However, owing to

the presence of the uncracked ligament, certain closing forces and bending

moments act on the faces of the equivalent through-crack. The bending moments

are caused by the eccentricity of the ligament with respect to the centerline

of the plate thickness. These forces and moments can be related to the

deformation of the ligament through compliance expressions. The place of the

ligament can thus be thought of as being taken by "springs" connected from one

face of the through-crack to the other. The compliance of these springs at a

point along the crack is assumed to be equal to that of an edge crack in plane

strain of the same depth as the surface crack at that point and in a plate of

the same thickness. This is motivated by considering the limiting cases. If

the crack were very long, it would begin to behave like an edge crack in a

thick plate (thus deforming in plane strain) with all the applied load being

carried by the uncracked ligament behind the crack. The other limit is that

of a very deep crack, in which the ligament becomes increasingly thinner and

eventually disappears entirely. In this case, the crack becomes a throuqh-crack

,

and all the load is carried by the material on the sides of the crack. The

actual situation is between these two extremes, and to analyze it approximately,

we assume it is a mixture of the two limiting cases, so that the load is

carried partially by the "edge crack" and partially by the "through-crack," as

shown in Fig. 4. To determine the unknown force, F(x), and moment, N(x), that

25

Fig. 4 - Line-spring model.

26

act on the edge crack at a point x along the crack and by equilibrium also act

on the through-crack faces, displacement compatibility is enforced; that is,

since the "through-crack" and the "edge crack" coexist in the same plate, they

must exhibit the same opening displacement of the crack and the same rotation

of the crack faces. Once the force and moment are known, we can compute any

parameter of interest, such as J or COD, by focusing our attention on the edge

crack.

To apply the line-spring model, expressions relating displacements and

rotations to forces and moments must be set up. It is convenient to introduce

the average tensile stress,

o o

= F/t (21)

and the bending stress,

m = 6M/t 2 (22)

where t is the plate thickness, and F and M are the force per unit length and

moment per unit length, respectively, on the edge crack. In terms of o

Q

and

m, the compliance expressions for the edge crack are

A

_ 2(1 - v)t (a n

o + a 12m) (23)

c E

e

c

= 12(1

E v2)

< a 12°o

+ a 22

m) (?/!)

where the compliance constants a^, a^> and are f°un d by considering the

edge crack loaded with o , m and with A and 0 defined as the additional

o c c displacement and rotation of the ends caused by the presence of the crack.

24

The procedure of Rice and Levy was used to calculate the compliance

constants (a-,), and the results for the a . . as functions of crack depth are

' J ' J shown in Fig. 5.

The mathematical complexities in the line-spring model arise in deriving

the expressions relating displacement and rotation to o o and m in the throughcrack, because the force and moment at any position along the crack -Front.

27

0.0 0.1 02 0.3 0.4 0.5 0.6 0.7 0.8

Normalized Crack Depth, a/t

Fig. 5 - Compliance coefficients , a^» sod a ^2

as a function of crack depth.

28

affect the displacement and rotation at any other point. Enforcing displacement

compatibility results in a pair of coupled integral equations that must be

24 26

solved numerically, * or the line spring may be inserted into a suitably

27 28 modified finite-element program. * Extension of the model into the

elastic-plastic region has been carried out by inverting Eqs. 23 and 24 and

rewriting the result in an incremental form, thus relating do

Q

and dm to dA

c and de

c through a stiffness matrix. At sections in which yielding has

occurred, the "tangent stiffness" is used. This procedure was proposed by

25 262729

Rice and successfully carried out by Parks and Lockett. *

’ A further

25 complication is the spread of plasticity from the sides of the crack. Rice

speculated on the possibility of handling this by using a strip-yield

approach, but this has not been applied previously. If the line-spring is

inserted into a finite element code, plasticity at the sides of the crack can

readily be accounted for by the code.

The following simplifications reduce the model to a purely analytical

one:

1. The surface crack is treated as existing in an infinite flat plate

loaded in remote tension.

2. The actual crack front is replaced by a crack of constant depth, and

displacement compatibility between ligament spring and through-crack is

enforced only at the center of the crack. This reduces the coupled integral

equations to a pair of linear algebraic equations, a simplification proposed

24

in a slightly different form by Rice and Levy. The constant crack depth

used in deriving the model presented here was the maximum depth of the actual

crack, so that the correct depth in the actual crack at the point of most

interest is matched by the model.

29

3. The spring is elastic perfectly plastic; that is, it remains linear

elastic until the force and moment in the spring reach a yield criterion.

4. Spread of plasticity from the sides of the through-crack is approximately accounted for by using an effective crack length of I, + 2 r^, where i is

the crack length and the plastic-zone size correction, r , is obtained from a

strip-yield analysis of a finite-width plate, the only manner in which finite

width enters the model.

A simplified model results. The compliance of the ligament spring is

still given by Eqs. 23 and 24. For an infinite plate loaded in tension by o,

with the average tensile stress, a , and bending stress, m, acting on the

crack faces, the displacement and rotation at the center of the crack are

A

c

= 2l/E (o - o

0

) (25)

„ . -4(1 + v)i

®c ' (3 + v)Et m (26)

Equation 26 is derived by integrating the solution for a point-moment applied

24 on the crack faces.

Forcing equality between A

c in Eqs. 23 and 25 end O

c in Eqs. 24 and 26

results in

[a

ll + */ta •- v

2 )lo

0 + a^m = r £/ 1 ( 1 - v

2 )]c (27)

a 12

a

o + ^ a 22 + 3(1 - v)(3 + v) ] m ' 0 (28)

o

= ao (29)

m = -Bo (30)

a

= U/[t(l - v

' )]}(a 22

+

3(1 - v ) ( 3 + v') )/S (31)

B = U/[t(l - V

2 )]} a 12

/S (32)

S

= {a n

+ - V )..} [a 22 + 3(1 _ v ) (3 + v

)3 a

i 2 (33)

30

The spring remains elastic until satisfaction of a yield criterion is achieved

in the edge crack. Various yield criteria were tried, including an upper

25

bound slip-line solution presented by Rice, a lower bound static equilibrium

condition that assumes yield occurs when a plastic hinge develops in the

ligament, and a simple average stress condition. The average stress condition

was chosen for simplicity and because its results agreed with the experimental

data presented in Refs. 30 and 31 as well as any of the yield criteria tried.

Yield is assumed to occur when the average tensile stress in the ligament, o ,

reaches a closure stress, a , shown in Eq. 14, resulting in

°o ' t

t - a

-o (34)

From Eqs. 29 and 34 the value of applied stress at which ligament yield

occurs, c LY > is determined:

1 t - a

- °LY '

a t

0 (35)

Subsequently, o o remains constant, given by Eq. 34, and m remains constant,

given by Eq. 30 with o = o

L y.

Equations 25 and 26 are still used for determining A

c and ©c

, but after

ligament yield, plasticity at the sides of the crack becomes important and the

crack length, i 9 is replaced by the "effective" length, i + 2r^. The

expression for r^ is

r = U - 0/4 (36)

y P and l is evaluated from:

P

sin (iri, /2W) = sin (ir£/2W)/[cos £(|- - - r

+ 1)] (37)

P

z a oa The plate will continue to carry additional load until the strip yield reaches

the edge of the plate, which occurs when the applied stress reaches the net

section yield stress

°NSY

= °- 1 ' £a /( Wt )] (38)

31

This results directly from equilibrium or from setting £ = W in Eq. 37.

Equations 37 and 38 are valid for finite-width flat plates. If this

model is applied to other situations, such as cracks in pipes, use of these

equations would not be appropriate, and the strip-yield solution for an

infinite plate might be more suitable, for which

*p- */ cosf |

1 (39)

r o and the net-section yield stress becomes simply o.

The values of o

Q

and m are now known for the entire range of applied

stress. From these, both COD and the J-integral at the root of the crack can

be computed. To calculate COD, it must be noted that A

c in Eq. 25 is the

displacement at the centerline of the plate thickness. The CTOD is used as a

fracture parameter, whereas the CMOD is more easily measured and is used for

experimental verification of model predictions. From the geometry shown in

Fig. 6,

CTOD = A + 0 (t - 2a)

c c (40)

CMOD = A + 0 t

c c (41)

An alternative method for calculating CTOD prior to ligament yield is to use

the relation between CTOD and K (Eq. 11). The stress intensity factor, K, is

calculated using the line-spring model, as discussed below. This method was

preferable because accuracy of estimates of K using the line-spring model can

15

be checked using published elastic numerical results. After ligament yield

the incremental value of CTOD was calculated using Eq. 40. This is justified

because the assumption of crack faces opening as parallel lines is accurate

after ligament yield, as has been confirmed experimentally using replication

31 studies. The result is

CMOD = (2WE) [1 + 2(1 + v) 8/ ( 3 + v) - a] (42)

32
mm
Actual Crack

(a)

Fig. 6

(b)

- Calculation of CM0D and CTOD.

33

( 43 )

for a <_ a^y

CMOD = [2(ji + 2r^)/El o + o(t - a/ 1 ) {[2(1 + v)b]/(3 + v)a - 1}

for o^y < o o^sy

CTOD = K

2 (l - v

2 )/Eo = CT0D

e (44)

for a <_ o^y

CTOD = CT0D

e + [2(£ + 2 r^ ) / E 1 [a( 1 - a/t) o] (45)

for o

E y

< o < a^Y

Both CMOD and CTOD increase without bound at o^.

The model presented is for flat plates in tension. To apply it to actual

structures, the value of the nominal membrane stress component normal to the

crack in the vicinity of the crack (i.e., the stress that would exist in that

region if the crack were not present) is calculated and used as the applied

tension, o. The redistribution of stresses caused by the crack is thus

ignored.

The model that has been described in this subsection is the

yiel ded-1 iganent model for flat plates. Further development of this model is

discussed below.

2.2.4 Crack-Opening Displacements in Pipe

The driving force for fracture in pipes differs from that in flat tensile

panels in two ways: First, shells have greater stiffness than flat plates.

Thus, the rotation of the crack faces, and the resulting CMOD, is less in pipe

than in plates, particularity for long cracks. Second, the axial stresses in

pipes are caused by bending or by bending plus tension, whereas, the flat

plates were tested in tension. The bending stresses result in a stress

gradient along the circumference. Therefore, as the plastic zone develops at

the ends of a crack (located in the region of maximum tension), plasticity

spreads into a decaying stress field. Thus, the stress for plastic collapse

34

is greater for pipes in bending than for plates in tension. These factors

were accounted for in the analyses of the large-dianeter pipe tests.

Section 2.4, and in the calculation of allowable flaw-size curves.

The higher stiffness of shells relative to flat plates against rotation

is approximately accounted for in the simplified line-spring model. The

line-spring model treats a surface crack as being equivalent to a throughcrack with closing forces and out-of-plane bending moments acting on its

faces. The magnitude of these forces and moments is determined by enforcing

compatibility of displacement and rotation between the through-crack faces and

the ligament. The compliance of the ligament is determined by considering an

edge crack of the same depth as the surface crack. The modification to the

line-spring model, introduced to model long surface cracks in pipes more

accurately, was to account for the effect of curvature on the relation between

moment and rotation in a pipe. This relation enters the line-spring model

when the compatibility calculation is made. For a flat plate, the relation

between moment on the crack faces and out-of-plane rotation is

0

= -24(1 + v)i,M/[Et 3 (3 + v)] (46)

The compliance is seen to become very large, approaching infinity in the

limit for long crack lengths. However, a circumferentially cracked pipe has

significant rotational stiffness, even if the crack extends completely around

the circumference. For long cracks, the effects of the ends of the crack on

rotation at the crack center may be neglected, and the moment-rotation

relation may be obtained by considering the axisymmet.ric problem of a cylindrical

shell subjected to moments distributed around its end. The relation between

moment and rotation is, then

©

c = -(2R) 1 2 [12( 1 - v

2 )]

3 4 M/(Et 5 2

) (47a)

35

Fig. 7 - Schematic stress distribution of a pipe containing

circumferential crack under bending.

The line spring calculations have been done in terms of bending stress m

rather than moment M. these quantities are related according to (22).

m = 6M/t 2 (22)

Writing (47a) in terms of bending stress m, we have

e

c

= -{(2R) 1 [12(l-v*)]V / (6Et 4

) Jm (47b)

Equation 47 is derived using axis symmetric shell theory. The relation

between M and e

c enters the line-spring model in the compliance calculation by

replacing Eq. 26 with Eq. 47 and then obtaining curvature-corrected results

for Eqs. 27 through 33.

In tensile panels, plastic collapse occurs when the net-section stress

exceeds the flow stress. The collapse mechanism consists of localized

deformation along 45° slip lines emanating from the crack tips to the edges of

the specimen. However, in a pipe subjected to bending or combined tension and

bending, the collapse mechanism differs from that in a tensile panel. A lower

bound estimate for the collapse load is obtained by assuming a plastic

hinge develops at the cracked section, as shown in Fig. 7. Force balance at

the plastic hinge determines the neutral axis location:

e

2

= (i)[P

t

/(2otR)] + (a/t ) (©

1 + tt) (48)

where the angles ©^ and ©2 are defined in Fig. 7, s the resultant of all

applied loads, and R is the pipe radius. Moment balance results in

M = (2oR 2 t)(2sin ©

2

) - (a/t)sin ©^ (49)

where M is the applied moment. For a given crack size, Eqs. 46 and 47 will

give an interaction curve of values of and M corresponding to plastic

collapse.

The special case P^ = 0 is of interest in the present study. For this

case, the collapse moment is given by

M

c

= 2oR 2 t{2cos [(a/t) ©j] - (a/t) s i n ©

^ } (50)

36

which reduces to

M

c = 2oR 2 t[2 - (a/tjGjl (51)

for small crack lengths. The moment at which flow first occurs in an

uncracked pipe is given by

M = irR 2 to (52)

Comparing Eqs. 52 and 51, for small crack lengths the collapse moment is

approximately 27% higher than the moment corresponding to incipient yield.

Thus, the pipe is capable of carrying significantly higher moments before the

CMOD increases without bound than is predicted by the critical-COD model.

A modification to the critical-COD model has been derived that addresses

this objection. The fully elastic contribution to CMOD is calculated as

described above. The additional contribution to CMOD in the post-ligamentyield range, CMOD^y* is computed using a new expression. The expression for

CM0D

ly

is deriving using an approach similar to that used by Heald, Spink, and

32 Worthinton in deriving strip-yield solutions for the COD of finite-width

33 tensile panels. deWit and Smith followed this approach in deriving a model

for a finite-width tensile panel containing a surface crack that gives

analytical results for CMOD^y comparable to those obtained using the

yiel ded-1 igament model.

First the strip-yield expression for CMOD of a center crack in an

infinite plate under tension is considered:

CMOD = 2oi./

(

tt E

)

* 1 n{ 1 + sin (tto/2o)]/[1 - sin(-rra/2o)]} (53)

Equation 53 can be thought of as an expression that increases without bound as

a approaches a, correctly reduces to the linear elasticity expression for CMOD

at small stress levels, and interpolates in between these limits using the

function in braces. The desired expression for CMOD^y should have similar

features with the following exceptions:

37

1. The expression for CMOD^y should increase without bound as the

plastic collapse moment is approached. When axial load is present, proportional loading must be assumed.

2. At small moment levels, the expression for CMOD^y should reduce to

the expression for CM0D

LY

that results when plasticity at the sides of the

crack is negligible.

An expression which meets the above criteria is

CM0D

ly = [2£a/U 2 E)][4 - at/(tR) - n(l - a/t)]*ln[(l + sin Y )/(l - sin y)

(54)

where y = (

tt / 2 )( M - A )/(M - A ), A

q

= ttR 2 t ( 1 - a/t)o, and M

c is evaluated

using Eq. 51. Using Eq. 54, CMOD^y is seen to increase without bound as the

collapse moment is approached. For small moments, the pipe is elastic, the

nominal membrane stress is

o = M/ tt R

2 1 (55)

and Eq. 54 reduces to Eq. 53. The logarithmic function interpolates between

the two extremes. The model described here is called the pipe

yielded-1 igament model.

2.3 Model Evaluation Tests for Surface Cracks in Tension

The critical-COD and flat plate yiel ded-1 i gament models provide an

analytical relationship among the COD, applied stress, flaw size, material

properties, and specimen (or structural) geometry. Both models can be

expressed in terms of the CMOD or CTOD. The experimental program was based on

the assumption that verification of the model in terms of CMOD is sufficient

to verify the models in terms of CTOD. This assumption is required because

CMOD can be readily measured, whereas CTOD cannot.

In this section, the experimental evaluation of the critical-COD and

flat plate yielded-! igament models is discussed for the case of surface cracks

38

in tension. The CMOD's of surface cracks are measured as functions of stress

and strain in tensile panels of API 5LX-70 steel plates and welded pipe

segments. The experimental results are compared with analytical predictions.

2.3.1 Experimental Procedures

The test materials were API 5LX-70 pipeline steel in the form of 15.9-mmthick plate and pipe segments with transverse girth welds. The pipe segments

were taken from 1220-mm(48-in) diameter, 15. 9-mm(0. 625-in) thick API 5LX-70

pipe welded using representative field practices. Metallography and

properties of the base plate and welds are described in Appendix H. Chemical

composition is given in Table 1, along with compositions of materials

considered elsewhere in this report.

Three series of specimens with different surface crack lengths and depths

were tested: one series each for the base metal, the manual welds, and the

automatic welds. The test matrix is shown in Table 2. The base-metal test

specimens were tensile panels (Fig. 8) that were notched in three different

ways: 1) a saw cut of 0.4-mm (0.016-in) width, 2) a saw cut followed by

fatigue sharpening, or 3) an electrical-discharge-machined (EDM) notch of

0.4-mm (0.016-in) width. The weld-metal test specimens were tensile panels

taken from the 1220-mm (48-in) diameter pipe with the weld transverse to the

tensile axis. The pipe curvature was retained in the specimens and the weld

reinforcement was removed to obtain a uniform cross section. The surface

notches, all prepared by EDM, were located on the concave side of the pipe at

the root of the V-shaped weld.

Each of the specimens in Table 2 was loaded in tension at room temperature,

and the CMOD was measured as a function of nominal stress, a, and gage-length

strain, e^. In the base-metal specimens, the e L was taken as the average

strain over a gage length of 305 mm (12-in), except for the 381-mm (15-inch)

39
.
1

I

Table 1 - Chemical Analyses In Weight Percent

C SI Mn P S A1 Cu Cr Ni Mo V Nb N

Base metal

(API 5LX-65)

0.08 0.3 1.45 0.015 0.003 0.039 0.08 0.09 0.03 0.10 0.07 0.036 0.009

Base metal

(API 5LX-70)

0.076 0.2 1.3 0.12 0.14 0.13 0.33 <0.05

Manual weld 0.12 0.17 0.67 0.005 0.010 0.13 1.6 0.031

Automatic weld 0.11 0.43 1.3 0.011 0.005 0.073 0.92 0.053

40

Table 2 - Test Matrix for

the Critical CUU

Ligament Model

the Experimental

Model and Flat

Evaluation of

Plate YieldedSpecimen

Number

Specimen Width, W, mm

Crack Depth, mm

a. Crack Length, fc, mm

PI

*

102 6.65 40.9

P2 102 9.07 43.5

P3 102 5.64§ 31.2

P4 102 6.88§ 48.3

P5 76 10.54 § 31.2

P6 76 4.57 13.5

P7 76 3.56 31.8

P8 76 5.08 14.7

P9 381 3.81 114.3

Al f 76 8.13 29.7

A2 102 5.72 47.5

A3 76 2.49 17.3

A4 76 4.57 11.8

Ml ^ 76 8.13 29.2

M2 102 5.84 47.5

M3 76 4.32 11.8

M4 76 2.16 15.0

* P: base metal

t A: automatic weld

t M: manual weld

§ specimen also used to evaluate the J-integral model

1 mm = 0.0394 in.

41

Front Side Back

Fig. 8 - Test specimen configuration.

42

wide specimen that had a gage length of 610 mm (24-inch). The specimens were

pulled in tension under displacement control with a closed-loop servo-controlled

hydraulic testing machine. The values of CMOD were measured with a clip-on

displacement gage mounted directly onto the crack mouth with the clip-on teeth

about 0.13 mm (0.005-in) into the crack. The was obtained from displacements measured over the chosen gage length with linear-variable-differential

transducers (LVDTs). All signals from electrical resistance strain gages,

LVDTs, the clip-on gage, and the load cell were acquired and stored by a

multichannel minicomputer for on-line and post-test analysis. For observation

of the deformation patterns near the cracks, the specimens were coated with

photoelastic material and instrumented with electrical-resistance strain

gages.

The CODs in specimen P5 were measured by a replication technique. The

specimen was loaded under displacement control to a CMOD of 0.2 mm (0. 0079-in)

A liquid, silicone base, precision-impression material was mixed with a

hardener and inserted into the crack. The replica material was allowed to

harden at constant CMOD for about 15 min. Subsequently, loading was continued

until a CMOD of 0.5 mm (0.02-in) was reached. The replica was removed and the

replication procedure was repeated. The procedure was repeated at CMOD values

of 0.9, 1.2, and 1.8 mm (0.035, 0.047, and 0.071-in). Each of the replicas

was cut along the minor axis of the semielliptical crack and mounted in phenyl

salicylate. The CODs were measured at several distances (x values) from the

crack mouth with a metal lographic microscope. The precision of measurement was

within ±0.03 mm (±0.0012 in) under 100X magnification.

2.3.2 Results and Discussion

For the test matrix shown in Table 2, CMOD has been measured as a function

of applied stress, o, and of gage-length strain, e^. The experimental results

43

of CMOD vs. o are compared with the analytical predictions of the analytical

models discussed previously.

for o _>_ [ 1 — (a/t)]o

Equations 56 state the critical-COD model applied to prediction of crack

mouth opening displacement. For short cracks, the critical-COD and flat plate

yielded ligament models agree. They agree for all cracks after full ligament

yielding.

23

Equation 56a is evaluated using the elasticity solution of Kobayashi

,

(Eq. 20) for cracks with aspect ratios (a/fi,) greater than 0.1. For long

cracks (a/£ < 0.1), the Kobayashi solution is not applicable and the elastic

21

CMOD is calculated using King's simplified version of the line-spring model

(see Section 2.2.3). Also, for long cracks, the plastic zone development

through the thickness before full ligament yielding becomes more important,

and it is accounted for by adding a plastic-zone correction, r to the crack

depth. The Dugdale plastic zone in a center-cracked finite-width panel is

used. In using this solution for an edge crack in the line-spring model,

bending is neglected; the underestimation of r

^ is considered negligible for

present purposes. The addition of the through-thickness plastic zone

correction to the flat plate yielded-! igament model produces a model called

the flat plate yielded-! igament model with through-thickness plasticity

correction.

Equation 56b is used from the onset of ligament yielding,

o _> [1 - (a - t)]a, until net-section yielding occurs. After net-section

yielding, CMOD increases without bound at a constant applied stress and is

shown as a vertical line in the CMOD-vs.-o curve, at o = c (1 - aa/tW). The

CMOD = CMOD(elastic) , for o < [1 - (a/t)]o

CMOD = [2U + 2r )/E] [o - (1 - a/t)a] + CMOD

(56a)

(56b)

44

COD,

mm

DISTANCE FROM CRACK MOUTH, in

0 0.1 0.2 0.3 0.4

80

60

40

20

0

Fig. 9 - COD as a function of location along the crack face

after ligament yielding. Measured with replication

technique. Crack depth = 10.5 mm (0. 413-in); crack

length = 31.2 mm (1.23-in).

45

COD,

mil

important quantity in the application of fracture mechanics is CTOD and not

CMOD, because CTOD is recognized as being related to crack initiation. With

moderate length and deep surface cracks, it is assumed that the rotation of

the crack faces owing to out-of-plane bending is negligible after the ligament

has yielded and the increments of CTOD at the leading edge and CMOD are equal.

Measurements of the crack profile are shown in Fig. 9. At low values of COD,

the replicating material did not penetrate all the way to the crack tip.

Thus, there are more data points at the higher values of COD than at the lower

end. As shown in Fig. 9, the assumption of negligible rotation of the crack

faces owing to out-of-plane bending after ligament yielding is reasonable.

Base-Metal Results . As shown in Table 2, nine base-metal tests were

conducted. The results, summarized in Figs. 10 and 11, show good agreement

between experiment and analysis. Note that the shape of the CMOD-vs-o curve

for the long crack (a = 114.3 mm, 4.5-in) test differs from that of the

shorter crack tests; nonlinearity due to crack-tip plasticity occurs at a

lower stress, and there is a more gradual transition to net-section yielding.

Addition of the Dugdale plastic-zone correction to the crack depth accounts

for this behavior. Thus, the analysis appears to be valid for a broad range

of crack sizes.

The chronological events associated with the CMOD-vs-a curves can be

described as follows: Initial loading produces elastic specimen extension and

the relationship between CMOD and stress is linear. As the stress increases,

the plastic-zone size ahead of the crack tip enlarges. Eventually, the

plasticity spreads to the back face of the specimen; this is termed ligament

yielding. After ligament yielding, the plasticity contributing to the CTOD

tends to occur in slip bands emanating from the crack tip at angles of ±45°

from the plane of the crack, and the nonlinearity of CMOD-vs-o curves becomes

46

STRESS,

ktl

U| '00N3

uiui 'Q0N3

Fig.

10

-

CMOD

vs.

nominal

stress

for

base-metal

specimens.

Symbols

are

experimental

data,

and

solid

lines

are

analytical

predictions

based

on

the

modified

critical-COD

model.

CMOD,

mm

STRESS, ksi

0 40 80 0 40 80 0 40 80 0 40 80

Fig. 11 - CMOD vs. nominal stress for base-metal specimens. Symbols are experimental data and solid lines are analytical predictions

based on the modified critical-COD model.

48

CMOD,

in.

noticeable. Initially, the slip bands are between the crack tip and the back

face. Finally, the net section yields along ±45° slip bands between the crack

and the specimen edges, and the CMOD increases asymptotically to the net-section

yielding stress. These phenomena were revealed by the brittle lacquer on the

front and back surfaces of the specimen. On the front surface. Fig. 12,

values of strain 3.8 mm above the center of the crack were more than two

orders of magnitude lower than those of remote strain. The strain increased

with distance from the crack plane and eventually reached the values of remote

strain.

On the back surface of the specimen, the deformation phenomena were

examined with brittle lacquer and electrical-resistance strain gages. As the

stresses increased, the highest strains were observed at regions about ±45°

from the plane of the crack emanating from the leading edge of the crack. The

ligament yielding stress estimated from a = (1 - a/t)o agreed well with

readings from strain gages at the highest strain regions. After ligament

yielding, the deformation pattern on the back of the specimen was an ellipse

that surrounded the position of the crack until net-section yielding was

approached. Then the pattern developed slip lines to the edge of the

specimen. The strains at the region right behind the crack remained

relatively low compared with the remote strains. This low-strain region could

be seen necking down after extensive net-section yielding, and then the crack

grew to the back surface. The strain patterns revealed by the brittle lacquer

are shown in Fig. 13.

The results in terms of CMOD vs. are shown in Figs. 14 and 15. Two

distinct behaviors are observed. Seven of the specimens exhibited a nearly

bilinear relationship between CMOD and e L>

with the slope sharply increasing

at strain levels slightly below the yield strain, as shown in Fig. 14. The

49

50

Strain

patterns

revealed

by

brittle

lacquer

on

the

front.

Note:

LVDT

ran

across

the

center

of

the

specimen,

and

the

crack

seen

was

only

half

of

the

surface

crack.

(a)

At

low

stress,

no

noticeable

plastic

deformation

ahead

of

the

crack

tip

was

observed.

51

Fig.

12(b)

-

At

intermediate

stress

level,

plastic

zone

size

enlarged

and

52

Fig.

12(c)

-

net

section

yielded.

53

Strain

patterns

revealed

by

brittle

lacquer

on

the

back

surface,

(a)

The

stress

level

was

at

about

ligament

yielding.

54

Fig.

13(b)

-

After

net-section

yielding.

50

xIO

aim 'aoi^io

z

<

ocH

CO

CO o

ON-2«

< CL

a> o

o2

«

0

Scoc< w

1

c

a)

0)a

X

55

Fig.

14

-

CMOD

vs.

gage-length

strain

for

base-metal

specimens

that

exhibited

net-sectionyielding

behavior.

CMOD,

mm

Fig. 15 - CMOD vs. gage-length strain for base-metal specimens that

exhibited net-section yielding followed by gross-section

yielding.

CMOD,

in

bilinear behavior results when the strains that exceed yield only occur in the

net section, that is, net-section yielding. Photoelastic observations and

strain gage measurements indicate that, for net-section yielding, yielding is

confined to slip bands, which extend at approximately 45° angles from the

crack tip to the specimen edges. As shown in Fig. 15, specimens 6 and 8

exhibited similar behavior at low strains, but at high strains the CMOD

reached a plateau value, increasing only slightly with strain. The plateau

develops when strain hardening elevates the flow strength in the slip bands

and strains above yield occur in the gross section, that is, gross-section

yielding. Once gross-section yielding occurs, the remote displacements are

absorbed along the entire test section and CMOD increases slowly with e^.

The bilinear relationship between CMOD and can be modeled at stress

levels below net-section yielding using Eq. 56, providing that the appropriate

relationship between and o has been established. At strains above net-section

yielding, the slopes of the CMOD-vs.-e^ curves are equal to the gage length;

this occurs because all of the remote displacement is transmitted to the crack

tip through the slip bands. Thus, after net-section yielding,

CMOD = CM0D

nsy

+ L(e

l

- e NSY ) (57)

where L is the gage length and the subscript NSY refers to the value of the

quantity at the onset of net-section yielding.

A relationship between and o is necessary to model CMOD vs. e L at

stresses below net-section yielding and to compute CMOD^y anc* f° r

Eq. 57. The is the sum of the elastic strain, o/E, plus the increment of

due to the presence of the crack. The remote displacement. A', in a

center-cracked panel is given by

A' = (2q£/E)[V(£/W)] (58)

For a surface crack. 34

where V(£/W) is given by Tada, Paris, and Irwin.

18

Irwin's equivalent through-thickness crack concept can be used to modify

Eq. 58; that is, o is reduced by the closing stress, o

c

» caused by the

uncracked ligament,

c

c

= o ( 1 - a/t) (59)

To account for the crack-tip plasticity, the effective crack length, ji^,

which includes the plastic-zone size correction, r , is used in place of

The resultant expression becomes

e L

= o/E + [2 ( o - a

c )/E][^eff/L][V(£eff

/W)] (60)

Given Eqs. 56a, 56b, and 60, it is possible to calculate CMOD vs. for

the net-section yielding case. For stresses below and equal to net-section

yield stress, Eq. 56 is used to calculate CMOD and o is converted to using

Eq. 60. For stresses above net-section yield stress, Eq. 57 is used. Comparison of the experimental and analytical results of CMOD vs. shown in Figs.

14 and 15 indicates good agreement.

For specimens 6 and 8, net-section yielding is followed by gross-section

yielding at high strains. The transition from net-section yielding to grosssection yielding, and CMOD vs. for gross-section yielding have not been

modeled. In the present study, the transition occurred at a ratio of crack

area to cross-sectional area of 5%. The area ratio at which the transition

occurs should be higher for materials with more strain hardening.

Weld Metal Results As shown in Table 2, four automatic-weld specimens

and four manual-weld specimens were tested. As shown in Table 3, the yield

and ultimate strengths of the welds were significantly different for both

types of weld and for the base metal. These differences were far less in

transverse-weld tensile tests on 76-mm (3-in) wide, full -thickness tensile

panels with a 305-mm (12-in) gage length. The stress-strain curves obtained

59

using strain gages mounted in the center of the weld at midthickness on both

edges are shown in Fig. 16. For the CMOD-vs.-o model, only the flow stress

values are important: 596 MPa (86.5 x 10 3 psi) for the automatic weld and

557 MPa (80.8 x 10 3 psi) for the manual weld. However, for CMOD-vs.-e

modeling using the two-tensile-bar model (discussed below), the shape of the

stress-strain curve is also important.

The experimental results of CMOD vs. o for the automatic and manual welds

are compared with the analytical predictions in Figs. 17 and 18. The general

behavior trends observed in the experiment are reasonably modeled; however,

the agreement between experiment and analysis is not as good as it was for the

base-metal tests. The correlations are best at low stresses, but the CMOD is

underestimated (specimens A2 and M4 are exceptions) for stresses between the

stress required for ligament yielding and the stress required for net-section

yielding. At still higher stresses, CMOD is overestimated. To model the

observed behavior with Eq. 56b, it would be necessary to use a variable flow

stress. This option is physically sensible, because the stress-strain curves

for the welds exhibit more work hardening than the base-metal curves. The

quality of the correlations is also influenced by an increase in experimental

uncertainty for testing the welded specimens over that encountered in base-metal

tests. The principal sources of error are: 1) bending caused by distortion,

misalignment, and the curved test section, 2) the presence of residual

stresses in the welded specimen, and 3) thickness variations caused by hand

grinding the crown and root of the weld.

The experimental results of CMOD vs. for the automatic and manual

welds are shown in Figs. 19 and 20. The two characteristic trends observed in

the base-metal tests, a bilinear curve for net-section yielding and a plateau

when gross-section yielding occurs, are not apparent in the weld-metal

60

Table 3 - Summary of Weld Metal Tensile Properties

Weld

Type

Test

Temperature,

°C

Ultimate

Tensile

Strength,

MPa

Yield

Strength

0.2% Offset,

MPa

Percent

Elongation in

25 mm

Percent

Reduction

in Area

Automatic +24 549 491 26 69

Automatic +24 820 735 15 * 31*

Automatic +24 800 * 715 21 66

Average 810 725 21 66

Manual +24 560 470 f 28 68

Manual +24 550 460 + 29 64

Average 555 465 28* 66

Automatic -101 955 820 t 61

Automatic -101 910 790 27 63

Average 930 805 27 62

Manual -101 665 58C + 30 65

Manual -101 675 560 + 21* 30*

Average 670 570 30 65

*Slant fracture; other specimens exhibited cup-cone fracture.

Yield point observed; other specimens did not exhibit a yield point.

j<Gage marks not visible on specimen after test,

psi x 10

3 = MPa/6.894

61

STRESS,

MPa

STRAIN

Fig. 16 - Stress-strain curves of automatic and manual weldments, as determined

in transverse weld tensile tests.

62

STRESS,

ksi

CMOD,

mm

STRESS, ksi

0 40 80 0 40 80 0 40 80 0 40 80

Fig. 17 - CMOD vs. nominal stress for automatic-weld specimens. Symbols are experimental data and solid lines are analytical predictions based

on the modified critical-COD model.

63

CMOD,

in

CMOD,

mm

STRESS, ksi

0 40 80 0 40 80 0 40 80 0 40 80

Fig. 18 - CMOD vs. nomi

experimental

based on the

nal stress for manual-weld specimens

data, and solid lines are analytical

modified critical -COD model.

Symbols are prediction

64

CMOD,

in

CMOD,

mm

STRAIN/YIELD STRAIN

to

0.8

0.6

0.4

0.2

0 1 2 3 4

i i i i r

i i i r Automatic Welds

a (mm) / (mm) v -

v 8.13

o 5.72

29.7

47.5 D * 4.57 11.8 *

a

a o 2.49 17.3 — * D — o°°

D

Lna° D

v* D D

D

*o°°

A

o* o *0*0* <j> O o O A A A

o o

- 30x10

-20

-3

0.002 0.004 0.006 0.008

-10

Jo

0.010

STRAIN

Fig. 19 - CMOD vs. gage-length strain for automatic-weld specimens.

/ 65

CMOD,

In

CMOD,

mm

STRAIN/YIELD STRAIN

1.0

3 1 2 3 4

l I I r

l i l l Manual Welds * a (mm) i(mm) -

A 8.13 29.2 ° A

° 5.84 47.5 o

° 4.32 11.8 0

v 2.16 15.0 ?

/ o

o y o

° *

1

o ° °

v V * v w v

- 30x1

0"3

-20

10

0.002 0.004 0.006 0.008 0.010

STRAIN

Fig. 20 - CMOD vs. gage length strain for manual-weld specimens.

66

CMOD,

In

results; however, the net-section yielding and gross-section yielding designations are still appropriate. The CM0D-vs.-e

l curves for net-section yielding

behavior in welded specimens are more parabolic than bilinear; the four

specimens with the largest cracks (Al, A2, Ml, M2) exhibited net-section

yielding. The CMOD-vs.-e^ curves for gross-section yielding behavior in

welded specimens show a very small increase in CMOD with e^; the four

specimens with the smallest cracks (A3, A4, M3, M4) exhibited gross-section

yielding.

The differences in the CMOD-vs.-e^ curves for the base metal and the

welds can be explained by consideration of the higher yield strength and

increased work-hardening rates in the weld. For small cracks, the base metal

starts yielding while the weld is still elastic. Subsequent increases in

stress applied to the weld depend on the work hardening in the base metal.

Simply, the base metal stretches and the weld does not, and consequently CMOD

increases slowly with strain. For the larger cracks (specimens Al, A2, Ml,

M2), 16% to 20% of the cross-sectional area of the weld is cracked.

Consequently, the crack plane yields first and net-section yielding can occur.

Even at the 20% crack-to-cross-sectional-area ratio, the weld work-hardens

sufficiently to cause yielding in the base metal.

For the net-section yielding case, the apportionment of strain in the

base metal and the weld can be modeled with the two-tensile-bar analogy shown

in Fig. 21.

For a given value of remote imposed displacement, A' = e^L, at the ends

of two tensile bars in series, the compatibility condition is

e^Lj + £ 2

I_2 = e^L (61)

where L is the gage length, l_ 2 is the width of the weld, and L

^ = L - Lp (see

Fig. 21). The force equilibrium is

67

(62) [°l(

e

].) ]

* A

l = ^°2 ^ e 2^ * A

;

where o^(e^) and (^(e^) are the stress-strain curves for materials 1 and 2,

and Aj is the gross-section area and f<2 is the net-section area in the plane

of the crack (see Fig. 21). Since the stress-strain curves are nonlinear,

Eqs. 61 and 62 are solved numerically for and Assuming all the strain

in the weld, e£, goes into the crack, then CMOD = £2*-2‘

A comparison of the measured CMOD-vs.-e^ data with the values calculated

using the two-tensile-bar analogy is shown in Fig. 22. The results indicate

that strain partitioning due to the difference in the stress-strain curves of

the base metal and the weld is a usable physical model of weldment behavior.

2.3.3 Summary and Conclusions

Crack-mouth-opening displacements (CMODs) of surface cracks have been

measured as functions of stress and strain in tensile panels of API 5LX-70

steel plates and welded pipe segments. The conclusions drawn from the

investigation are:

1. For CMOD vs. stress, the flat plate yiel ded-1 igament model with

through-thickness plasticity correction agrees well with the experimental

results for the base metal and the welds.

68

A2 = A 1 -at

(a)

(b)

Fig. 21 - Two-tensile-bar model schematics: (a) specimen and (b) analogy.

69

CMOD,

mm

STRAIN/YIELD STRAIN

0 1 2 3 4 5 6

STRAIN

Fig. 22 - CMOD vs. gage-length strain for an automatic weld-metal

specimen exhibiting net-section yielding. Symbols are experimental data; solid lines are prediction from twotensile-bar model; dashed lines are predictions from

simple net-section yielding models.

70

CMOD,

In

2. For strains above net-section yielding, the relationship between

CMOD and strain is complex, depending upon the specimen type (plate or weld)

and crack size.

a. For large cracks in base metal, all of the remote displacement goes

into the crack tip through the slip bands extending from the crack

tip to the plate edges at 45°, that is, net-section yielding. This

behavior was modeled and predictions from the model agree well with

the experiment.

b. For cracks with crack area less than 5 % of the cross-sectional area,

net-section yielding is followed by gross-section yielding in which

the remote displacement is distributed evenly along the length of

the specimen.

c. For large cracks in steel welds, net-section yielding is observed.

The observed CMOD is lower than predicted by the model because the

yield strength of the weld is higher than that of the base metal. A

model, based on strain partitioning, accounts for the observed

behavior and provides better predictions.

d. For small cracks in steel welds, gross-section yielding is observed

throughout the strain range studied.

2.4 Model Verification Tests on Large-Diameter Pipes

An extensive series of full-scale experiments on pipes were conducted by

the Welding Institute of Canada.^ The pipe was API 5LX-70 steel, 914 mm

(36-in) in diameter and 11.5-mm (0.45-in) thick. Pipes containing surface

cracks were tested to failure. The surface cracks varied in length from 58 mm

(2.28 in) to 300 mm (11.8 in) and in depth from 3 mm (0.118 in) to 10.2 mm

(0.40 in). The tests included in this report are summarized in Table 4. The

tests were conducted at temperatures ranging from -90°C to 20°C.

Twelve-meter-long pipe sections were loaded in bending by 20 hydraulic

77

Table 4 - Test Matrix for Large-Scale Tests Conducted at

the Welding Institute of Canada

Test

Number*

Crack

Length, mm

Crack Depth , mmWall

Thickness, rran Temperature

,

°C

2B 60.4 5.6 11.3 20

4 63.5 5.9 11.2 20

6 68.6 7.8 11.1 -90

7 61.0 5.4 11.5 -90

8 76.4 10.1 12.0 -90

9 81.7 8.8 11.6 -50

11 79.0 9.3 11.3 -50

13 59.6 6.2 10.8 -50

14 64.8 5.5 11.0 -5

15 60.4 5.5 11.2 -5

18 300.0 3.0 11.1 20

19 300.0 3.0 11.1 20

* Welding Institute of Canada identification numbers.

1 mm - 0.0394 in.

72

jacks at 610-mm (24-in) spacing. A clip-on gage was used to measure CMOD, and

the jack pressure was recorded. Strain gages were placed at various locations

to determine the strain distribution in the pipes. In the low- temperature

tests, the temperature distribution was measured with thermocouples. The

surface cracks were located in the girth weld or in the base metal. Manual

arc welding was used. The flow strength of the welds was approximately 590

MPa, which slightly undermatched the base metal.

On the basis of beam theory, the maximum membrane stress induced at the

cracked section was calculated to be

o = 30. 6p (63)

where p is the applied jack pressure. The bending moment at the crack section

was

M (in N-m) = 9.96 x 10

6

p (64)

Eqs. 63 and 64 were used in the present work to permit analytical computation

of CMOD as a function of jack pressure.

2.4.1 Results and Discussion

Experimental results for CMOD as a function of jack pressure are shown in

Figs. 23 through 27. The tests on plain pipe, conducted at 20°C, are plotted

in Fig. 23. The remainder of the tests were on pipe sections containing girth

welds. The tests in Fig. 24 were conducted at -90°C, the tests in Fig. 25 were

conducted at approximately -50°C, and those in Fig. 26 at -5°C. The tests in

Fig. 27 involved relatively long surface cracks and were conducted at 20 C C.

The general form of the results is the same for each test: a linear relation

between CMOD and p prevailed at low pressures, followed by a nonlinear region

as yielding occurred in the vicinity of the crack, and finally a vertical

asymptote as the pressure causing collapse at the cracked section was approached.

73

CMOD,

mm

— — Flat plate yielded-ligament model

Pipe model

• Experimental data

JACK PRESSURE, MPa

Fig. 23 - Comparison of model predictions and experimental

results for CMOD vs. jack pressure. Tests were conducted at room temperature.

74

CMOD,

mm

— — Flat plate yieldecMigament model

Pipe model

• Experimental data

JACK PRESSURE, MPa

Fig. 24 Comparison of model predictions and experimental

CMOD vs. jack pressure. Tests were conducted at

results for

-90°C.

75
in
CMOD,

mm

— — Flat pipe yielded-ligament model

Pipe model

• Experimental data

Fig. 25 - Comparison of model predictions and experimental results for

CMOD vs. jack pressure. Tests were conducted at -50°C.

76

CMOD,

mm

— — Flat plate yielded-ligament model

Pipe model

• Experimental data

0 5 10 15 20 25 0 5 10 15 20 25

JACK PRESSURE, MPa

Fig. 26 - Comparison of model predictions and experimental results

for CMOD vs. jack pressure. Tests were conducted at

-5°C.

77

CMOD,

mm

LIne-eprIng modal with plaatlc tone

• Experimental data correction In depth direction;

Flat plate yielded- ——1. Flat plate

ligament model —2. With shell correction

0 5 10 15 20 25 0 5 10 15 20 25

JACK PRESSURE, MPa

Fig. 27 - Comparison of model predictions and experimental

results for CMOD vs. jack pressure. Tests were conducted at room temperature.

78

Analytical predictions of CMOD as a function of jack pressure are compared

with the experimental results described above. In the discussion, the original

version of the yielded-1 igament model. Section 2.2.3, will be referred to as

the flat-plate yielded-! igament model. The modified version, which accounts

for the plastic collapse mechanism of the pipe, discussed previously in

Section 2.2.4, will be referred to as the pipe yielded-1 igament model.

Typically the flat-plate yielded-1 igament model agrees fairly well with

observed values of CMOD until higher jack pressures, at which the model

predictions are higher than the observed values. However, the discrepancy

does not typically occur until higher values of CMOD have been reached. For

example, if it is assumed that the highest CTOD fracture toughness value of

pipeline weld metal that might be used in a f itness-for-service assessment is

0.25mm (0.010-in), then typical surface-flaw geometries such a toughness

value (i.e., CTOD value) would be equivalent to a CMOD of approximately

0.28 mm (0.011-in). The model predictions of CMOD agree favorably with the

experimental results until values of CMOD higher than this are reached. Thus

the flat-plate yielded-! igament model is, in general, sufficiently accurate

over the range of interest. An exception was test 6, Fig. 24, in which the

model predictions were conservative (by a maximum of approximately 40%) at

lower pressures. The reason for this discrepancy has not been explained.

The conservatism of the flat-plate yielded-1 igament model at higher

pressure values appears to be attributable to incorrect modeling of the

plastic deformation in the pipe. The model predictions are vertically

asymptotic at a value of pressure corresponding to the tensile net-section

yield stress, o^y* given by Eq. 38. The model asymptote is typically to the

left of the data asymptote. The modified version of the model, Eq. 54, called

the pipe yiel ded-1 igament model, assumes a more realistic plastic collapse

79

mechanism and seems to agree more favorably with the experimental results at

higher pressure levels. Exceptions are tests 7, 9, 13, and 15 in which the

predictions of the flat-plate yielded-1 igament model agree more closely. This

observed behavior is curious because it is reasonable to expect the data to

adhere more closely to the modified model. Two possible explanations are

inelastic buckling and tearing. Buckling behavior was observed in several of

the tests. Inelastic ovalization during buckling would significantly lower

the plastic collapse pressure. Some ductile tearing was also reported in some

of the tests. The model predictions based on the original crack length would,

therefore, be too low.

In the two tests shown in Fig. 27 the crack length was relatively long

(300mm, 11.8-in). A straightforward application of the flat-plate

yielded-1 igament model to the calculation of CMOD in this case gives poor

results. It is seen that the data become nonlinear much sooner than the model

predicts. This was attributed to the importance of yielding through the

ligament depth in this case, because the center of a long surface crack

receives very little support from the material at the sides of the crack. Use

of the flat-plate y i el ded-1 igament model to compute the elastic component of

CMOD and accounting for through-depth yielding via the plastic-zone size

correction presented in Section 2.3.2 resulted in extremely conservative

curves. Upon examining the line-spring calculation more closely, it was found

that the largest contribution to CMOD was from the rotation of the crack

faces. The increased stiffness of a curved shell against rotation in

comparison with the stiffness of a flat plate must, therefore, be accounted

for. This was done in the manner described in Section 2.2.4, resulting in the

pipe yielded-1 igament model with through thickness plasticity correction.

Prediction of this model are shown in Fig. 27, which agree well with the

experimental values of CMOD.

80
.
2.4.2 Summary

On the basis of experimental results from full-scale tests of pipes in

bending, the flat-plate yielded-1 igament model was found to give predictions

of crack driving force accurate to within experimental scatter or conservative

by less than 20% over the range of practical interest in a f itness-for-service

assessment. At driving force values above this range of practical interest,

the model predictions become more conservative. This has been explained by

introducing a modification to the model that more properly accounts for the

plastic deformation in a pipe, the resulting model was called the pipe

yielded-! igament model. Calculations of COD by the model in its original form

are relatively inaccurate for long surface cracks. However, a modification to

the model has been proposed that greatly increases its applicability to the

long crack case. This final model was named the pipe yielded-! igament model

with through-thickness plasticity correction. Because the model is relatively

simple to apply, yet predicts crack driving forces that compare well with

experimental checks, being accurate or conservative by less than 10% over the

range of practical interest, it is felt to be a useful tool in determining the

driving force for surface cracks under elastic-plastic deformation. This is

an important ingredient in setting allowable flaw-size standards in pipeline

girth welds and other structures.

2.5 Allowable Flaw-Size Curves

An allowable flaw-size curve is a plot of flaw depth as a function of

flaw length, as shown schematically in Fig. 1. Each point on the curve

represents a critical flaw size (length and depth) calculated using an

appropriate fracture mechanics model and specified values for applied stress,

residual stress, fracture toughness, tensile properties end pipe dimensions.

The specified values are conservatively chosen to preclude the possibility of

81

achieving critical conditions in service. In this section, the fracture

mechanics model developed in this program for the calculation of allowable

flaw size curves, the pipe yielded-1 igament model with through thickness

plasticity correction, is described. Guidelines for the selection of stress

levels, pipe and girth-weld material properties, and pipe dimensions for use

in the analysis are presented. Finally, a representative allowable flaw-size

curve is given, and the influence of selected parameter variations on the

position of the curve is shown.

2.5.1 Fracture Mechanics Model

The analytical model used to calculate allowable flaw-size curves is the

pipe yielded-1 igament model wih through thickness plasticity correction. The

fracture criterion is the CTOD:

6 c = CT0D

e + CT0D

ly + CT0DpS (65)

where 6 c = fracture toughness in terms of CTOD

CTOD^ = elastic component of CTOD

CTOD^y = ligament-yielding component of CTOD

CT0D

RE = residual stress component of CTOD

The idea is that fracture occurs when the driving force for fracture, the

CTOD terms in Eq. 65, exceeds the resistance of the material to fracture, 6 .

Values for the fracture toughness, applied stress, tensile properties, and

pipe geometry are specified on the basis of pipeline operating conditions

(e.g., pressure, temperature, imposed displacements) and pipe and weld

requirements (e.g., diameter, thickness, strength, toughness). The remaining

variables in Eq. 65 are the flaw depth, a, and the flaw length, £,. Using the

specified values, Eq. 65 is solved numerically for a as a function of £ using

a Newton-Raphson iteration procedure. The computer program used for these

calculations is given in Appendix A.

82

The elastic component of CTOD is calculated using the simplified linespring model discussed in 2.2.3. The equations used in the calculations are:

CT0D

e

= K

2 (l - v2 )/Eo (11)

K * t

!

(o

0gt

+ mg

b

) (66)

where g

t and g b are given by:

g

t U) = ttC*(1.12 + 6.52c

2

- 12.39c

4 + 89.05c6

- 188.61c

8 + 207.39c10

- 32.05c

12

(67)

g

b (C) = irC*(1.12 - 1.89c + 18.01c

2

- 87.39c

3

+ 241.39C

4

- 319.94c

5 + 168.01c6

) (68)

C = (a+r )/t and o Q

and m are computed using Eqs. 27 through 33. For cracks

longer than 96 mm (3.78 in), curvature must be accounted for in the compliance

calculations. The curvature correction enters the line-spring model by

replacing Eq. 26 with Eq. 47 and making the associated changes in Eqs. 28

through 33. The revised equations are 28c to 33c, in which the suffix c, and

the subscript c in the symbols a

c and s , stand for curvature.

a + m {

+ Km* rmi-v*) ]^! o

12 0 22 72(l-v 2 )t*

o_ = a o

0 c m * - B c°

(?8c)

(29c)

(30c)

a„ = U/[t(l-v 2 )]) {a 22

+ (2R)i 02U-v2 )]

^ /s

72(l-v 2 )t

i

(31c)

B

c - U/Ct.(l-v 2 )]}a12/S = 8 (32c)

S = {a,,+ n/[t(l-v 2 )]} {a 22 + } - a? 11 22 2 72(l-v 2 )t* 12 (33c)

83

In evaluating Eq. 66, the plastic-zone correction, r , is added to the crack

depth. The r^ is calculated as in Eq. 17, but using depth and thickness in

place of length and width, and inserting the appropriate stress ratio:

The plastic zone size calculation is non-linear and self-consistent; both

(29-c) and (69) must be satisfied. The plastic zone enters (29c) through the

context of an elastic treatment of the ligament. As discussed in Section 2.4.1,

yielding through the ligament must be considered in the calculations to obtain

agreement with experiment. This is accomplished by replacing a/t on the

the applied stress reaches a certain value given by Eq. 34; above this stress,

the ligament is fully yielded and CTOD increases according to Eq. 70. Before

the ligament is fully yielded, no plasticity corrections to

a are made.

The ligament yielding component of CTOD is given by:

The residual stress component of CTOD is estimated by assuming that the

residual stresses result from weld shrinkage, and thus, are self-equilibrating.

In this case, one can assume that yield-point stresses act over a distance

comparable to the weld size or pipe thickness. If so, a displacement of

of such a residual stress to the driving force for fracture is to simply

assume that the estimated displacement is all converted to CTOD:

sin (ira/2t)/sin (Tr(a+2r^) )/2t = cos (tto o/2o) . (69)

CTOD^y = [2(* + 2r )/E] [a - (1 - a/t)a]

where r # is the strip-yield plastic zone solution given in Eq. 17.

(70)

CTODrS = Cyt/E (71)

84
mmxM
Evaluating Eq. 71 using the ANGTS pipeline properties listed above, the

result is a CT0D

R

<- of 0.038 mm (0.0015 in).

Calculations of allowable flaw-size curves using the equations listed

above are most conveniently done by computer. Several aids to study of the

details of the calculation are given in Appendix A. The computer program,

used to compute the allowable flaw-size curves for this study, is listed. A

table gives values of a^, a^> ^22* ^i*

and as f unc

't

"’ ons effective

crack depth. Figure A1 shows the behavior of Eq. 69, which relates r^ to o Q

.

The self-consistent plastic zone calculation is accomplished by finding the

plastic zone size for which both (69) and (29-c) are satisfied. As an aid to

visualization of this solution process, Fig. A1 shows (29-c) and (69) plotted

as curves on the same graph. The calculated r^ value occurs where the two

curves intersect. The advantage of doing these calculations by computer is

clear. At each flaw length, several trial depths must be checked before the

critical depth is found. For each trial depth, the plastic zone size must be

calculated anew. Also listed in Appendix A as an example is a hand

calculation of critical flaw depth for a given flaw length. Computer results

were used, instead of guesses, to reach the final step in both the plastic

zone and critical depth iterations. Although the allowable flaw-size

calculations discussed here are complex, the actual computer costs are

insignificant compared to the manpower costs.

2.5.2 Parameters

The relative position of the allowable flaw-size curve depends on the

parameters used to evaluate Eq. 65. For the ANGTS pipeline, the pipe grade,

API 5LX 70, and the pipe geometry, 1220-mm(48-in) diameter and 15.9-mm

(0.625-in) wall thickness, have been specified, but a reliable data base on

the tensile properties and the fracture toughness of the pi*pe and weld are net

available.

85

It is planned to operate the pipeline at -4°C, but the maximum allowable

stress (or strain) normal to the girth welds is not known by the authors. It

has been assumed that the pipeline will be in nearly continuous operation at a

nearly constant pressure and thus, crack growth in service due to fatigue

cycling has been neglected.

For the purposes of this investigation, the maximum credible stress level

is assumed to be 90% of the specified minimum yield strength of the pipe,

434 MPa (63 x 10 3 psi). The following tensile properties are assumed for both

the pipe and the girth welds:

= 482 MPa (70 x 10

3

psi)

d = 517 MPa (75 x 10

3

psi)

o = 551 MPa (80 x 10

3

psi)

E = 208 GPa (30.2 x 1C

6

psi)

v = 0.29

The minimum fracture toughness at the service temperature is assumed to

be 0.13 mm (0.005 in). It should be emphasized that this value exceeds the

minimum results obtained in this investigation, but it is considered to be an

achievable minimum for qualified welding consumables and procedures.

2.5.3 Results and Discussion

Allowable flaw-size curves have been developed using the fracture

mechanics analysis model described Section 2.5.1. For the parameters listed

in Section 2.5.2, the resulting curve. Fig. 28, is considered to be the

current best-estimate of the allowable flaw-size curve for the ANGTS pipeline.

The position of the curve is strongly affected by the parameters selected.

A sensitivity study has been conducted to assess the influence of variations

in fracture toughness, tensile properties, and stress levels on the allowable

flaw sizes. The results are presented in the form of a reference curve

Flaw

Depth,

mm

15.0

12.5

10.0

7.5

5.0

2.5

0.0

Fig. 28 - Best estimated allowable flaw-size curve.

87

developed using the parameters listed in Section 2.5.2 and two additional

curves developed using values of the variable under study below and above the

reference value (Figs. 29-31).

The curves showing sensitivity to toughness variations are shown in

Fig. 29. Note that by increasing the fracture toughness from 0.13 mm

(0.005-in) to 0.25mm (0.010-in), the allowable flaw depths are more than

doubled for flaws up to 125-mm (4.92-in) long. For longer flaws, the increase

in allowable depth decreases to about 30% for flaws that are 500-mm

(19.7-in)long.

The curves showing sensitivity to variations in applied stress are shown

in Fig. 30. Note that a 10% increase in stress results in a substantial

decrease in allowable flaw size. For cracks longer than 125 mm, (4.92-in) the

allowable depth decreases by more than 50% when the stress is increased from

90% of the specified minimum yield strength (SMYS) to 100% of SMYS, that is,

from 434 to 482 MPa (62.9 to 69.9 x 10 3 psi). The reason for this large

sensitivity to applied stress is apparent from consideration of Eq. 70. Note

that the CTOD^y is proportional to a - (1 - a/t)o, and thus it is the relative

difference between o and a that influences flaw size. The same effect is seen

in the curves depicting the influence of flow stress, shown in Fig. 31.

The allowable flaw-size curves developed herein are compared with allowable

flaw sizes permitted under Appendix A 16th edition, of API 1104. Comparisons

were made at the two toughness levels considered by API, that is, CT0D values

of 0.13 mm (0.005 in) and 0.25 mm (0.010 in). The remaining parameters were

those listed in Section 2.5.2. The results are presented in Figs. 32 and 33.

At the 0.13-mm (0.005-in) toughness level. Fig. 32, the allowable flaw-size

curves developed herein are more restrictive than the API allowables at

essentially all crack lengths. The differences are largest for crack lengths

88

Flaw

Depth,

mm

Fig. 29 - Effect of critical COD on the allowable flaw-size curve.

89

Flaw

Depth,

mm

Fig. 30 - Effect of applied stress levels on the allowable

flaw-size curve.

90

Flaw

Depth,

mm

Flaw Length, mm

Fig. 31 - Effect of flow stress on the allowable flaw-size curve.

91

Flaw

Depth,

mm

Fig. 32 - Comparison of allowable flaw-sizes with various

criteria. Critical COD = 0.127 mm.

Flaw

Depth,

mm

Fig. 33 - Comparison of allowable flaw sizes with various

criteria. Critical COD = 0.25 mm.

93
m
greater than 100 mm (3.94 in). For a flaw that is 250-mm (9.84-in) long, the

allowable flaw depths are 2.5 mm (0.098 in) by the present method and 4.0 mm

(0.157 in) by the API method. At the higher toughness level. Fig. 33, the

allowable flaw sizes developed using the API procedures are more restrictive

for crack lengths up to 230 mm (9.06 in) and less restrictive for longer

cracks.

The differences in allowable flaw sizes between the present approach and

those of Appendix A of API 1104, 16th edition, result from differences in the

treatment of ligament yielding, which is not explicitly addressed in Appendix

A of API 1104, 16th edition. The assessment procedure and flaw-size

allowables in Appendix A of API 1104, 16th edition resemble those of

PD6493:1980, from the British Standards Institution^. Section 8.2.3 of

PD6493 : 1980 accounts for ligament yielding using a flaw recategorization

procedure. Flaws deep enough to produce ligament yielding are treated as

through thickness flaws, whose characteristic size is the flaw length, not its

depth. This procedure disallows some flaws, resulting in a more restrictive

allowable flaw sizes. The treatment of ligament yielding in the present study

is more extensive and more specific than that of PD6493 : 1980

.

2.6 SUMMARY

A fracture mechanics analysis model, the pipe yielded-! igament model

with through thickness plasticity correction, has been developed that relates

allowable flaw sizes to applied stress level and weld toughness. For analysis

purposes, the weld flaws have been treated as surface cracks. The fracture

criterion used is the crack-tip-opening displacement (CT0D); that is, fracture

occurs when the CT0D due to imposed strains (the applied CT0D) exceeds the

fracture toughness (critical CT0D) of the girth welds.

94

The validity of the model was confirmed for the case of surface cracks in

tensile panels at stress levels below the net-section yield strength. The

comparison of analysis and experiment was complicated by the fact that CTOD

cannot be readily measured. Thus, it was necessary to make the comparison in

terms of crack-mouth-opening displacement (CMOD). CMOD was measured with a

clip-on displacement gage and is calculated using the flat plate

yielded-1 igament model with through thickness plasticity correction. The

model further evaluated by analyzing the results of a series of tests on

900-mm (36-in) pipe conducted by the Welding Institute of Canada. Comparison

of analysis, using the pipe yielded-! igament model with through-thickness

plasticity correction, and experiment indicated that the line-spring model has

acceptable accuracy for CTOD values up to 0.2 mm (0.0079 in), the range of

interest in developing allowable flaw-size curves.

This pipe yielded-ligament model with through-thickness plasticity

correcton has been used to calculate allowable flaw-size curves for operating

conditions and girth-weld properties representative of the proposed ANGTS

pipeline. The applicability of the model is limited to pipelines with maximum

stresses in the longitudinal direction below the specified minimum yield

strength of the pipe. It is further limited to pipelines where the yield

strengths of the weld and base metal are similar. The required degree of

similarity in yield strengths has not been established, but the desirability

of higher yield strength in the weld metal has been clearly demonstrated. The

use of allowable flaw-size curves can be considered for either general

pipeline use or for site-specific engineering assessment. Allowable flaw-size

curves generated using typical pipeline operating conditions and weldment

properties are more conservative for long flaw lengths than those derived

2 using the procedures of Appendix A of the 16th edition of API 1104.

95

3. SHARP-FLAW DIMENSIONING

C. M. Fortunko and R. E. Schramm

3.1 Introduction

In the evaluation of pipeline girth-weld quality on the basis of rational

fitness-for-service criteria, ultrasonic inspection appears to be a better

method than radiography, which has usually been used. Although radiography is

well suited for assessing compliance with workmanship quality criteria, the

relative insensitivity of radiography to the presence of sharp flaws makes it

much less effective as a fitness-for-service inspection tool than ultrasonic

inspection (cf. Appendixes C and E). Furthermore, even though experience has

shown that some information about the through-wall depth of blunt flaws can be

inferred from densitometric analyses cf radiographic data, recent work has

shown that the reliability of sizing by such methods is very low for sharp

flaws less than 0.25-mm (0.0098-in) thick, even when the most sophisticated

microdensitometry techniques are employed." Densitometry techniques can

provide information only on the through-wall radiographically projected

dimension of a canted sharp flaw; they provide no information on overall

through-wall flaw height. Therefore, such techniques cannot be reliably used

to detect and measure cracks, incomplete fusion, and inadequate penetration.

Conventional, high-frequency ultrasonic methods are sensitive to sharp flaws

and can provide dimensional information under ideal conditions (cf. Appendix D).

However, as described in Appendix F, tests run on deliberately flawed pipeline

welds used in this work showed an inability of state-of-the-art high-frequency

methods to reliably perform the required flaw detection and sizing. This

result confirmed conclusions in the literature that high-frequency approaches

have fundamental limitations that limit their usefulness in a fracture mechanics

37 approach to a f itness-for-purpose analysis of a welded structure.

96

Recent theoretical and experimental investigations have shown that sharp

flaws and cracks in butt welds can be sensitively detected using a new

36-38

ultrasonic inspection technique. ” The new technique differs from most

conventional ultrasonic nondestructive evaluation (NDE) methods in three

39-43

important respects: First, the weld region is insonified using special

ultrasonic probing signals: shear waves that are polarized parallel to the

surface of the weldment (SH waves). Second, the probing ultrasonic signals

are generated and detected using the recently developed periodic-permanentmagnet (PPM) electromagnetic-acoustic transducers (EMATs)

i

n contrast

with the fluid-coupled piezoelectric transducers, the EMATs do not require

intimate coupling and can operate efficiently over most unprepared surfaces.

Third, the ultrasonic frequency of operation of the new inspection system is

substantially lower (0.5 MHz) than the frequency range used in conventional

ultrasonic inspection of ferritic butt weldments (2-10 MHz).

The lowering of the frequency of operation significantly reduces the

sensitivity of the SH-wave EMAT system to minor perturbations in flaw surface

46 ^7

detail: surface roughness and departure from planarity and branching. *'

If the frequency is sufficiently low, the amplitude of the signal scattered by

40 48 49

the flaw can be unambigously related to the principal flaw dimensions. * *

In addition, sharp flaws are more easily detected than blunt flaws of similar

dimensions. Of course, these features are highly desirable from the fitnessfor-service point of view, which emphasizes the importance of the principal

flaw dimensions and assumes that sharp flaws are inherently more critical than

. f1 4,50-52 blunt flaws. *

The qualitative differences between ultrasonic scattering at low and high

frequencies are illustrated in Fig. 34. (It is assumed that the flaw is

planar and relatively long in the direction of the normal to the plane of

97

(a) (b)

Fig. 34 - Ultrasonic scattering by two-dimensional, planar flaws:

(a) high-frequency region; (b) low-frequency region.

98

reflection). At high frequencies (Fig. 34a), most of the incident ultrasonic

energy from the face of the planar flaw is reflected specularly, that is, the

angle of reflection with respect to the face normal equals the angle of

incidence. Also, a significant portion of the incident signal is diffracted

by the tips of the flaw, which act like ultrasonic line sources. If the

surface of the flaw exhibits some roughness, some of the incident signal can

be scattered incoherently (not shown). Furthermore, depending on the

orientation of the flaw and the type of ultrasonic signal used ( i . e . , shear

vs. longitudinal), significant mode conversions can also occur. As a result,

the scattered ultrasonic field is very complicated and it is very difficult to

infer the flaw dimensions from the ultrasonic data.

At low ultrasonic frequencies (Fig. 34b), the scattered ultrasonic fields

are considerably easier to interpret because the specular reflection component

becomes negligible and the tip diffractions are merged. Consequently, the

scattered ultrasonic fields are much less sensitive to flaw orientation and

47

surface roughness. In addition, at low ultrasonic frequencies, sharp flaws

are detected more easily than blunt flaws of similar length and through-wall

depth. The principal advantage, however, of operating at low ultrasonic

frequencies is that the amplitude of the backscattere_d signal increases

monotonically with the flaw dimensions.

In this report, emphasis is placed on the details of the experimental

configurations and procedures. An experimental configuration is described

that permits the ultrasonic interrogation of the entire weld zone from the

outer surface of the pipe. It has a built-in self-calibration feature that

automatically and consistently records the experimental data. This feature is

needed to offset variations in the EMAT electromechanical coupling

99

Digital Control

TTL Control Signal

I

RF Power Amplifier

impedance Matching

Network

Transmitter EMAT

2

Module

Trigger Signal o Oscilloscope

Display

» To Signal Digitizer

and Computer

Bandpass Filter

Receiver Preamplifier

I

Impedance Matching

Network

Receiver EMAT

tr

Fig. 35 - Low-frequency, SH-wave EMAT inspection configuration for

pi pel ine girth welds.

100

efficiencies, which may occur in practice as a result of variations in EMAT

lift-off distance, material composition, pre-existing residual and applied

53

stresses, and magnetic state.

The physical principles governing the propagation of the special SH-wave

probing signals in isotropic plates are discussed in detail. The discussion

is confined to the special case of ul trasonically "thin" plates whose wall

thickness does not substantially exceed the bulk ultrasonic wavelength (7.4 mm

0.29 in, at 454 kHz). This restriction is primarily dictated by the choice of

a theoretical model used to describe the scattering of low-frequency SH-wave

signals by planar flaws of different through-wall depth, position, and orienta

tion with respect to the free surface. However, general scattering theories

are found in Refs. 54 through 56.

Considerable attention is focused on experimental evidence that is used

to validate the main features of the theory. In addition, experiments are

described that establish the inherent sensitivity limitations of the system

for different categories of surface and buried flaws. The results are

summarized in a set of diagrams that are pertinent to a 454-kHz SH-wave EMAT

system operated on plate-like ferritic steel specimens 15.9-mm (0.62-in)

thick. An inspection protocol, applicable to automated inspection of newly

fabricated pipeline girth welds, is also discussed.

3.2 Inspection, Configuration, and Operation

Recently, several ultrasonic systems have been developed specifically

41-44

for the purpose of inspecting pipeline girth welds. However, the performance of the new systems is limited by uncertainties associated with coupling

and the highly complex nature of ultrasonic scattering from flaws when the

flaw dimensions of interest are large compared with the ultrasonic wavelength,

but not large enough to permit high-resolution imaging. The low-frequency,

101

Fig. 36 - Detection of a surface flaw (inadequate

penetration) in a 15.9-mm SMA girth weld

using 454-kHz SH waves.

102

SH-wave EMAT system was then specifically designed to overcome most of the

main shortcomings of conventional ultrasonic systems that use piezoelectric

transducer elements, which typically operate at high ultrasonic frequencies

(2-10 MHz).

Figure 35 shows an experimental configuration that has been found to be

particularly useful in the detection and limited characterization of elongated

flaws in pipeline girth welds. The configuration of Fig. 35 is composed of

two identical SH-wave EMATs that are positioned on the outer surface of the

pipe and aimed colinearly along the normal to the girth weld. The transmitter

EMAT is located closer to the weld than the receiver EMAT. Because SH-wave

EMATs are bidirectional, the transmitter EMAT generates two SH-wave signals of

equal amplitude, which travel in opposite directions along the normal to the

weld. The ultrasonic signal 1, traveling to the left of the transmitter EMAT,

insonifies the weld region, while the ultrasonic signal 2, traveling to the

right of the transmitter EMAT, passes directly beneath the receiver EMAT. In

practice, signal 1 is used to interrogate the weld region, and signal 2 is

used to calibrate the overall transduction efficiency of the system.

If a flaw is present in the weld, a portion of the probing signal 1 is

backscattered in the direction of the receiver EMAT. The backscattered signal

3 arrives at the receiver EMAT after the reference signal 2 because it travels

a longer distance. The two signals, 2 and 3, are shown in the photo of

Fig. 36.

The signals in Fig. 36 were observed when an inadequate penetration flaw

was insonified in a 1220-m (48-in) diameter, 15.9-mm (0.62-in) thick API

5LX-70 SMA-welded pipe section. The center frequency of the ultrasonic

signals was 454 kHz, corresponding to an ultrasonic (shear-wave) wavelength of

7.4 mm (0.29 in). In the photo, the first signal (0-30 ys) is caused by

103

direct electromagnetic interference between the transmitter and receiver

EMATs. This signal carries no useful information. The second signal

(60-100 ys) is the reference signal 2; the third signal (100-140 ys) is the

backscattered signal 3.

It has been determined that the amplitude of the reference signal 2 and

the amplitude and shape of the backscattered signal 3 carry useful information

about the dimensions, orientation, and position of the flaws within the

38 3Q

weld. ’ ~ Although it is not possible to reconstruct the shape of the flaw

and determine its position and orientation from a single low-frequency

ultrasonic measurement (because of poor spatial resolution at low

frequencies), it is possible to obtain much useful quantitative information

about these parameters from a series of independent measurements along the

normal to the girth weld. In particular, by using appropriate signal

processing techniques, it is possible to classify the flaws in terms of their

position within the weld (i.e., surface vs. interior) and then to estimate

38 their principal dimensions. However, to understand the principles of the

ultrasonic data inversion procedures, one first requires a qualitative insight

into the physical mechanisms governing the propagation and scattering of lowfrequency SH waves in isotropic plates.

3.3 Theory of Horizontally Polarized Plate Wave Propagation

When the frequency of operation is sufficiently low, it is often

advantageous to represent the ultrasonic field in an isotropic plate as a 39 54

modal expansion. * This representation is also approximately valid when

the radius of curvature of the pipes is very large compared with the

ultrasonic wavelength and when the observer is located sufficiently far from

the transducer. Although clearly limited, the modal representation provides

valuable quantitative insights that can be used to optimize the design of

low-frequency, SH-wave inspection systems in terms of the important

104

experimental parameters: plate thickness, frequency of operation,

weld-to-EMAT spacing, and the receiver-EMAT-to-transmitter-EMAT spacing.

In an infinite isotropic plate of finite thickness, there are only a

finite number of propagating, horizontally polarized plate waves and an

infinite number of nonpropagating, horizontally polarized plate waves.

Correspondingly, there are a finite number of plate waves polarized in the

plane normal to the plate (Lamb waves) that can propagate and an infinite

57 58 number that are cut off. * In this report, it is assumed for simplicity

that SH-wave signals are not coupled to Lamb waves by elongated weld flaws,

whose long dimension is parallel to the SH-wave polarization direction. This

assumption is strictly true for two-dimensional geometries and is justified in

practice by the inherent insensitivity the of SH-wave EMATs to ultrasonic

waves polarized in the plane normal to the plate (Lamb and Rayleigh waves)

In the coordinate system of Fig. 37, the transverse ultrasonic field

distributions (particle velocity v

n

' and stresses T' ' and T' ') of the n X XZ XV

propagating and nonpropagating SH-wave plate waves are given by:

59

with

vj^(y) = A

n cos (rnry/t)

T^(y) = (' BGA

n

/ w ) cos (niry/t)

T j^(y) = (-n7rGA

n /tjw) sin mry/t

B n

= EU/v s) 2 - (niT/t) 2 ]* for n = 0, 1, ?,...

(72)

(73)

(74)

where: t is the plate thickness, v

$

is the bulk shear wave velocity, G is the

shear modulus, w is the angular frequency, j = /-l, 8 p is the propagation

2 constant, A

n

= jo) 2 /2e^Gtc

n

, and e o

= 1 and e p = \ for n = 1,2,3,...

Along the z direction, the fields of the propagating and nonpropagating

SH plate waves vary as:

v (n) (z,x), T

(n)

' * XZ

and T

(n)

xy

(z,t) ^ exp [- j ( BpZ + toi ) 1 (75)

105

SH 0

T (0)

xy

Fig. 37 - Transverse profiles of SH

Q

and SH-| plate waves.

106

The particle velocity, v , and the two stress field components, T and

A aZ

T , for the two lowest order SH plate waves, the SH and SH n , are shown in

xy o 1 Fig. 37. It is observed that the SH

q

plate wave fields are uniform, but the

SH^ and higher order plate wave fields exhibit sinusoidal variations along the

transverse direction, y. The SH

q

plate wave is also unique in that it carries

no T stress component.

A L.The phase and group velocities, Vp and v^, of the SH plate waves are

given by:

and

v'

n)

= v

s [1 - (f^/f)’]'*

V V = v c g p s

(76)

(77)

where v

$

is the bulk shear-wave velocity, f is the wave frequency and the

"cut-off" frequency of an n*^ SH plate wave is:

f^

n)

= nv

$

/2t (78)

It follows from Eqs. 73 through 78 that each SH plate wave not only

exhibits a different variation along the transverse direction, y, but also

travels at a different velocity along the direction z. This observation is

the basis for the quantitative procedures for flaw characterization and

dimensioning from the ultrasonic scattering data obtained using the inspection

system and experimental procedures described in the preceding section.

In this report, it is assumed implicitly that both the transmitter and

receiver EKATs can only generate and receive the SH

o and SHj plate waves with

39 relative amplitudes of 1 to 1.6. Although the SH

2

, SH^, and SH^ plate waves

can also propagate in a 15.9-mm (0.625-in) thick plate at 454 kHz, they cannot

be detected or generated efficiently by the EMATs. Neverthless, their effect

must be included in any calculation of the ultrasonic scattering by the

flaws.

39

107

3.4 Theory of Scattering of Horizontally Polarized Shear Waves by Elongated

Flaws

To gain a physical understanding of the scattering of SH plate waves by

flaws, it is useful to use the concept of a reflection coefficient, r. Here,

r is defined as the ratio of the amplitudes of the incident and backscattered

ultrasonic signals at the EMAT electrical terminals. For the two-dimensional

case (no variations in v and T* along x and flaws of infinite length), a very

59

useful relationship for r has been derived by Auld.

r = t4(PjP

z )T k J

(vj-T

2

- v

2

-T

1 )- n dS (79)

where n is the inward-directed unit-normal to the surface, Sp, which surrounds

the flaw. Subscript 1 is used to denote the ultrasonic field in the absence

of the flaw, and subscript 2 is used to denote the ultrasonic field in the

presence of the flaw, assuming that the EMATs 1 and 2 in Fig. 34 are used as

transmitters. and are the electrical power levels delivered to the

EMATs 1 and 2, respectively, when used as transmitters.

Equation 77 can be used to gain valuable insights. Consider the case of

a two-dimensional, planar flaw shown in Fig. 38. In this case, Eq. 77 reduces

to:

a r

r = jm/4(P 1 P 2 ) i / (Au rT 2 )-n dS (80)

ISr

where Au is the mode-III displacement jump across the face of the planar flaw

and the integration is carried over only one-half of the surface, Sp.

In general, because several SH plate waves can propagate in a practical

situation, the reflection coefficient r can be rewritten as:

CC 00

r(z,L

tr ) * jw/MPjPj) exp (-jB

n z) a n

r nmbm

exp r.je^z+L^)]

(81)

n=0 m=0

108

Fig. 38 - Planar flaw geometry.

109

where z is the distance between flaw and EMAT, L. is the distance between

tr

EMATs, and the quantity r is defined as: nm

r = / ds . ( 82 ) nm

iSp

the quantities a m

and bm express the relative transduction efficiencies of the

different SH plate waves by the EMATs 1 and 2, respectively, and T

' m ' is the

stress component T associated with the m

th

plate mode.

In practice, and as described in a preceding section, the ultrasonic

scattering data (signal 3 in Fig. 35) is normalized using a reference signal

(signal 2 in Fig. 35). This normalized signal, called the reflected signal

amplitude ratio, R

$

, is given by:

00 00

-j 2-J JLrf eX P (‘jB Z) a r, r r, m

bm eXp[-jB (Z + L, )]

n n nm m m tr M^tr ) - —2yt

a

&

a

£

b

£

eXP + £ tr

(83)

If the quantities a

n

, b

n

, z, and are known and R^ZjL.^) can be measured

at a sufficient member of points, the unknown coefficients, r , can be found

nm

from Eq. 83. Equation 83 can be inverted if a and b vanish for n or

n n m j> N = M. In general, N x N independent measurements are then needed but

this number can be reduced by taking into account the symmetry of the matrix

r^. Thus, if only the SH

q

and SH^ plate waves are generated and detected by

the EMATs (although more SH plate waves can propagate), only three independent

measurements are needed. If signals carried by the SH

o> SH^, and SH^ plate

waves are detected, then only six independent measurements are required. In

practice, more measurements are available, but the size of the rmn matrix is

determined by the number of SH plate waves that can be generated and detected

by the EMATs. The larger the number N, the more information is available for

flaw characterization and dimensioning.

110

Although a detailed treatment of the dependence of the coefficients

on flaw size, shape, orientation, and position parameters is beyond the scope

of this report, it is instructive to consider several cases of practical

importance. For a more detailed treatment, the reader should consult Refs.

36,37, 54-56, 60, and 61.

The magnitudes and phases of the coefficients can be obtained using

numerical methods for the vertically oriented surface and interior planar

39

flaws of infinite length. The case of canted, planar flaws is treated in

Refs. 40 and 56.

The numerical results for the coefficient Tqq for three different flaw

types are shown in Fig. 39 as a function of the flaw through-wall depth, a. To

illustrate the dependence of Tqq on the relative position of the flaw within a

plate, the numerical calculations were carried out for edge (surface-connected),

centered (buried) cracks, and eccentric (buried) cracks. The calculations

were carried out for a case of current practical interest (wavelength,

X , = 7.4 mm, (0.291 in); plate thickness, t, = 15.9 mm, 0.625 in).

It is interesting to compare the results for the relatively shallow

(a < X/2) surface and buried cracks. It is observed that Tqq for the shallow

crack case initially exhibits a greater rate of increase than that for a

shallow buried crack of the same through-wall depth. However, at a ^ X/3 for

the surface crack and a ^ X/2 for the buried cracks, the rate of increase with

a/t is substantially less, and both coefficients tend linearly to 1. Although

the rate of increase of Tqq for a shallow buried crack is initially less than

that for a shallow surface crack, for a > X/2, the results converge asymptotically. As will be seen later, the initial behavior of the coefficients

determines the detectability thresholds of surface and buried flaws.

Ill

PHASE

FLAW THROUGH-WALL DEPTH TO PLATE THICKNESS RATIO

Fig. 39 - Numerical results for the SH-wave plate mode

scattering coefficient, Tqq.

112

It is also interesting to compare the behavior of Tgg with that of

r 01

= r 10

anc* r n* ^he numer'’ ca ^ results for Tq^ and are shown in

Figs. 40 and 41. For relatively shallow, vertically oriented, surface cracks

(a << X/4) the magnitudes of the four scattering coefficients Tgg, Tg^, r^g,

and r n

are approximately equal. However, as the crack through-thickness

dimension, a, is increased, Tgg and approach unity monotonically, but

and r 01reach a maximum at a/t ^ 0.5 and vanish for a/t = 1. In addition, the

signs of the coefficients Tg^ and r^g depend on which side of the plate

contains the flaw. In contrast, the signs of Tgg and are always positive.

These features can be used to classify flaws with respect to their relative

position within the plate.

A comparison of the numerical results for buried, centered, and eccentric

flaws shows that the magnitudes of Tg^, r^g, and are not significant when

a < X/2. By symmetry, Tg^ = r^g = 0 for buried, centered cracks.

Furthermore, the phases of Tgg, r^g, Tg^, and are -90° for very small

through-thickness dimensions and approach 0° for deeper cracks. Note that the

phases for buried and surface cracks approach 0° more rapidly.

An analysis of the above results indicates that significant results about

the flaw characteristics and through-thickness dimensions can be obtained from

the knowledge of the scattering coefficients, r .

3.5 Experimental Verification

Although the main conclusions have been made on the basis of

two-dimensional modeling, they have been found to be in very close agreement

with experiments conducted on flaws of finite length and calibration

standards. Thus, it can be expected that the general results outlined above

can be used as a basis for rational inversion procedures for estimating flaw

parameters from the available ultrasonic data. Because the magnitude and

113

a/t

FLAW THROUGH-WALL DEPTH TO PLATE THICKNESS RATIO

Fig. 40 - Numerical results for the SH-wave plate mode scattering

coefficient, .

114

FLAW THROUGH-WALL DEPTH TO PLATE THICKNESS RATIO

Fig. 41 Numerical results for the

scattering coefficient,

SH-wave plate mode

T

7 75

phase of the scattered signals depend strongly on the position of the flaw

within a plate, classification of flaws with respect to position must be

performed first to select the appropriate set of sizing curves.

The theory of Eq. 83 has been directly compared with experimental results.

Figure 42 shows a comparison of experimental results for surface notches in

flaw-plate calibration specimens with a calculated response, using ag = b

Q

=

1, a^ = bj = 1.6, z = 50 mm, (1.97 in) L^

r

= 210 mm, (0.625 in) f = 454 kHz,

and t = 15.9 mm. A slight correction for the effects of two-dimensional

diffraction was made. Measurements were taken from the left ("triangles") and

right ("squares") of the simulated flaws. It is seen that the experimental

data adhere very closely to the calculated curve. Similar results were

62 63

obtained for other classes of flaws. *

3.6 Detectability Limitations: Long Flaws

In practice, the detectability of particular flaw types in a girth weld

is not determined by ul trasonic-signal-to-electronic-noise considerations.

Instead, the detectability limitations are determined primarily by: 1) flaw

reflectivity as a function of flaw type, size, position, and orientation and

2) the presence of an ultrasonic background caused by transducer sidelobe

radiation and surface irregularities. The ultrasonic background level caused

by transducer sidelobe radiation can be reduced by using proper EMAT design

techniques. However, the ultrasonic background level caused by topographical

irregularities, for example, weld reinforcement, alignment mismatch ("high-low"),

and "drop through," cannot be eliminated because, at low ultrasonic frequencies,

it is not possible to resolve the flaw signals from the unwanted ultrasonic

interferences.

The scattering of SH plate waves by certain surface irregularities (e.g.,

weld reinforcement and alignment mismatch) can be modeled analytically, but

116

REFLECTION

AMPLITUDE

RATIO

FLAW THROUGH-WALL DEPTH TO PLATE THICKNESS RATIO

Fig. 42 - Comparison of experimental and theoretical results for vertically

oriented surface slots (flat-plate specimens 12-24).

117

ultimately the results of such calculations must be verified experimentally.

Here, a combined analytical-experimental approach is used to determine the

principal factors limiting the sensitivity of the SH-wave EMAT system. For

simplicity, the analysis is limited to the case of two-dimensional scatterers.

Physically, this means that only flaws and surface irregularities longer than

the transverse resolution width are considered. (Detectability of sharp flaws

shorter than the ultrasonic beam width is discussed in Section 3.7).

To determine the reflectivities of very long, buried and surface flaws, a

number of flat-plate calibration specimens were prepared. The specimens were

made using 15.9-mm (0.625-in) thick pipeline steel (API-5LX-70) sections that

were 1220-mm (48-in) long and 120 mm (4.7-in) wide. To permit a systematic

study of the effects of flaw position and orientation within a weld, four

different categories of artificial flaws were used: 1) thin surface slots,

2) centered, buried slots, 3) off-center, buried vertical slots, and

4) off-center, canted slots. The surface and off-center slots were designed

to simulate actual weld flaws (incomplete fusion, inadequate penetration, and

cracks) in shiel ded-metal -arc (SMA) girth welds. The width of the slots was

intentionally held below 0.4 mm (0.016 in), which is substantially less than

the bulk shear wave wavelength (7.4 mm (0.291 in) at 454 kHz, the inspection

frequency). The artificial flaw dimensions are summarized in Table 5, and the

characteristic flaw dimensions. A, B, and C, are defined in Fig. 43.

The buried artificial flaws were made by filling the "V-grooves" in

Fig. 43 using a combined tungsten-inert-gas (GTAW), gas-metal -arc (GMA)

welding procedure. To simulate the effect of weld reinforcement, alignment

mismatch ("high-low") and "drop through," additional flat-plate specimens were

prepared. The maximum dimension for the three surface irregularities was

3.2 mm (0.126 in).

118

Table 5 - Artificial Flaw Dimensions

Specimen No. A (mm) B (mm) C (mm) Type of Slot

1 6.37 1.58

2 5.88 2.15 Off-center

3 5.38 1.95 vertical

4 4.88 2.15 si ots

5 3.62 3.12 (buried)

6 2.38 4.22

7 0.00 5.48

8 6.88 0.76 Centered

9 6.58 2.50 vertical

10 5.5 3.9 si ots

11 3.62 7.45 (buried)

12 0 0.50

13 0 0.75

14 0 1.03

15 0 1.1

16 0 1.5 Vertical

17 0 1.75 edge

18 0 2.18 si ots

19 0 2.50

20 0 3.18

21 0 3.78

22 0 5.2

23 0 6.33

24 0 7.47

25 Full penetration weld (no defect)

26 6.07 0.97 0.75 Off-center

27 5.2 2.55 2.15 canted

28 3.33 5.07 4.3 si ots

119

a)

b)

Fig. 43 - Geometry of vertical slot (a) and canted slot (b)

calibration specimens. The slot dimensions are summarized in Table 5.

120

An examination of Eq. 83 shows that the reflection amplitude ratio, R

s

»

is strongly dependent on the flaw-to-EMAT spacing, z, and on the EMAT-to-EMAT

spacing, L

^

. A similar result can be obtained for the reflection amplitude

ratios of the three topographical irregularities: weld reinforcement, alignment

mismatch ("high-low"), and "drop through." For example, in the case of weld

reinforcement, the reflection amplitude ratio can be determined from:

00 00

je!w, ex P (-jez)A b exp[-jB_(z + L. )]

R

s

= 0 w w ^0 m=0

~ n ~" r 1 w ~ n ~'"nnrm ~ tr ' (84)

2t

£=0

a D b 0 exp (-.1B 0 L, )

with

sin(B + B ) W /2 nm

m

(*3 + 6 )W /2

n m w

/b B e e

( n m n m

j (85)

where W and T represent the mean width and thickness of the weld reinforcement,

w w Assuming typical values of Ww

= 20 mm (0.787), Tw= 1.6 mm (0.063 in),

a n = b n = 1, a, = b, = 1.6, and a = b = 0 for n > 2, it can be shown that

U U 11 n n —

the value of R

$

does not exceed 0.12 when t = 15.9 mm (0.625 in) and f = 454

39

kHz. The above estimate has been verified experimentally using actual girth

welds and flat-plate calibration specimens.

Figure 44 shows the variation of the amplitude reflection ratio as a

function of the distance, z (with L = 210 mm, 8.27 in), obtained using a

15.9-mm (0.625-in) thick flat-plate calibration specimen designed to simulate

a SMA weld reinforcement. It is apparent that the background signal ratio

associated with the simulated weld reinforcement does not exceed 1.2, in

excellent agreement with the theory of Eq. 84. This observation is very

significant, since the presence of the weld reinforcement is one of the

principal factors in determining the ultimate sensitivity of this inspection

system.

121

Amplitude

Reflection

Ratio,

Flaw-to-EMAT Distance, z, in _34 56 7 89 10 11

0.5

0.4

0.3

0.2

co tr

0.1

0

0.5

0.4

0.3

0.2

0.1

0

60 100 140 180 220 260 300

—1—1

1

1 1 1 1 1

1

Transmitter

-n \

- 1

Theo

z —

retical Limit (max)

••••• *•••••*

(a)

Transmitter

q\

l

Theoreticol Limit (mox)

*•• •

I

i I I i I*w

(b)

Flaw-to-EMAT Distance, z, mm

Fig. 44 - Detection of simulated weld reinforcement as a function of the transmitter-EMAT-to-welddistance, z. I_ t

= 210 mm.

122

If vertically polarized plate waves (Lamb waves) were used to probe the

weld, the reflected amplitude ratio associated with the weld reinforcement

would be significantly larger. Physically, this is related to the fact

that Lamb waves in a thick plate, which can support several propagating plate

modes, are tightly bound to the surfaces and do not exhibit large displacements

58

in the interior of a plate. In contrast, the SHq plate wave displacements

are uniform and the SHj, plate wave displacements exhibit only one null at the

midplane of the plate (see Fig. 37).

Significant levels of ultrasonic background signals can also be caused by

the presence of an alignment mismatch condition ("high-low") and "drop through"

(at weld root). Figure 45 shows the measured values of the reflection amplitude

ratio, R , at 454 kHz as a function of the distance, z, with L = 210 mm

(8.27 in)and t = 1.59 mm (0.023 in) for a 3-mm (0.118-in) alignment mismatch

and a 3.2-mm (0.126-in) bead at the weld root. It is seen in Fig. 45 that the

presence of a 2.8-mm (0.110-in) alignment mismatch condition causes a

background signal level that is at least twice as large as that for a weld

reinforcement or root bead of comparable thickness. For comparison, a plot

of the coefficient R as a function of z is shown in Fig. 46 for a vertical

surface slot 1.5-mm (0.059-in) deep and 100-mm (3.94-in) long.

An analysis using Eq. 83 in conjunction with calculated values of

r 00* r 01’ r l0’anci r ll

s

*’10ws that buried sharp flaws scatter less efficiently

than surface sharp flaws of the same through-wall depth. Physically, this is

caused by the fact that shallow surface flaws exhibit a larger mode III

displacement jump (in Eq. 80) than buried flaws. This can be verified by

using elastostatic expressions for au in Eq. 82, which are valid when the

A ultrasonic wavelength is significantly longer than the flaw through-wall

40 dimension a. In addition, shallow flaws lying near the midplane of the

plate cannot efficiently scatter the SH plate waves and convert, the SH

^ plate waves to SHq plate waves, and vice versa. This can also be verified directly

Amplitude

Reflection

Ratio,

Flaw-to-EMAT Distance, z, in

Flaw-to*EMAT Distance, z, mm

Fig. 45 - Detection of simulated 3-mm alignment mismatch

(a) and "drop through" (b) as a function of the

transmitter-EMAT-to-weld-distance, z. L.

210 mm.

-tr =

124

Amplitude

Reflection

Ratio,

Flaw-to-EMAT Distance, z, in

CO a:

(a)

(b)

60 100 140 180 220 260 300

Flaw-to-EMAT Distance, z, mm

Fig. 46 - Detection of a 1.5-mm-deep surface slot as a function of

the transmitter-EMAT-to-weld distance, z. (a) Slot in top

surface of plate, (b) Slot in bottom surface of plate.

L^. r

= 210 mm. (Flat-plate specimen 16.)

125
Bam
by noting that Tq^, i^q, and in Figs. 39 through 41 are negligible for

such flaws. As a result, shallow buried flaws are generally more difficult to

detect than surface flaws of similar dimensions.

It follows from Eqs. 83 and 84 and from the experimental data shown in

Figs. 44 and 45 that the relative amplitudes and phases of the flaw and

background ultrasonic signals are characteristically determined by the

distances z and L^

r and by parameters describing the surface irregularity and

the sources of the ultrasonic background. As a consequence, a particular

selection of z and can result in significantly higher sensitivity to

certain categories of flaws. For example, for certain EMAT-to-weld separation

distances, significantly stronger reflections are obtained from flaws located

near the root of the weld than from flaws located in other parts of the

39

weld. As a result, in this particular configuration, flaws lying near the

root of the v/eld may be easier to detect than flaws located in other parts of

the weld. However, this difficulty can be corrected by devising a detection

algorithm that utilizes ultrasonic data from different EMAT positions. In

this way, a more uniform interrogation of the entire weld can be ensured.

(The additional data are also needed for flaw characterization purposes.)

Equation 83 can be used to study the behavior of the flaw reflection

signal as a function of the flaw position within the weld and the EMAT-to-EMAT

and EMAT-to-weld separation distances, and z. For simplicity, it can be

assumed that the EMATs can only generate and receive signals carried by the

SHq and SH

^ plate waves. This simplification is justified by the fact that

the relative amplitudes Uq and a^) of the two plate waves are 1 and 1.6,

respectively. The relative amplitude of the SH^ plate wave is approxi-

. , n ~ 39,63 mately 0.3. ’

126

To emphasize the influence of flaw position within the weld on the

behavior of the flaw amplitude reflection ratio as a function of the separation parameters, z and it is convenient to fix the value of L^

r such that

^0L tr

= e

l L tr*

"^ his corres P on(:)s t0 the actual selection of the EMAT-to-EMAT

spacing in the present experimental setup. Furthermore, it is convenient to

assume that only three categories of sharp flaws need be considered: 1) surface

flaws located near the top surface, 2) interior flaws located near the midplane,

and 3) surface flaws located near the bottom surface of the weldment. The

bottom surface corresponds to the surface opposite to the surface on which

both EMATs are located.

Assuming that the surface flaws, which must be detected, are very

shallow, it can be shown that the magnitudes of the scattering coefficients in

39

Eq. 83, (Tqq, r n , Tq^, and r^g) are approximately equal. However, the

phases of the four coefficients are determined by which surface contains the

flaw. If the flaw is located near the top surface, then the phases of

r 00’ r 01’ r l0’ancl r ll are identical ly equal. However, if the flaw is

located on the bottom surface, then the phases of and are equal and

opposite to the phases of and r^g. Finally, if the flaw is located near

the midplane of the weldment, then Tg^, r^g, and either vanish or are

negligibly small, and only Tgg contributes strongly to the reflected signal.

The double summation in Eq. 83 can be evaluated graphically using a

vector-diagram (phasor) representation. For simplicity, it is assumed here

that the magnitudes of the four scattering coefficients (Tgg Tg^, r^g, and

r n) are unity for the three categories of flaws. (This can be readily shown

to hold in the elastostatic limit, which is valid when the flaw through-wall

depth is a fraction of the ultrasonic wavelength, by using the linear

elasticity solutions for Au

x to evaluate Eq. 82. ) Then, the amplitudes of

127
4
the vectors (phasors) are determined by the SH plate wave amplitudes, a

p and

b

n

, and the relative phases are determined by the differences between the SH

plate wave propagation constants, B n

» and the EMAT-to-weld distance, z.

Figure 47 shows the evolution of the reflected signal amplitude for the

three representative flaw categories as a function of the EMAT-to-weld

distance, z. The vector diagrams for the case of surface flaws can be

compared directly with the experimental results for surface slots shown in

Fig. 46. For convenience, the vector diagrams are evaluated at six

characteristic locations: 1) z = 0, 2) z = tt/2 ( 6q

- Bj), 3) z = tt/(Bq - 6^),

4) z = 3tt / 2 ( Bq - Bj), 5) z = 2tt/(Bq - B^) , and 6) z = 5it/2(Bq - Bj).

Physically, the above locations correspond to the following flaw-to-EMAT

distances in Fig. 46: 0, 60 mm, 120 mm, 180 mm, 240 mm, and 300 mm 0, 2.36,

4.72, 7.09, 9.45, and 11.8 in). The vector diagrams are arranged vertically

in the figure, such that the column on the extreme left corresponds to surface

slots at the top surface, the center column corresponds to buried slots at the

middle, and the column on the right corresponds to surface slots at the bottom

surface.

In Fig. 47, the vectors corresponding to the four terms involved in

evaluating the summation in Eq. 83 are represented by continuous lines. The

resultant, corresponding to the actually observed signal amplitude, is

represented by dotted lines.

A comparison of results shown in Fig. 47 with the experimental data in

Fig. 46 show good quantitative agreement, except at z = 120 mm (4.72 in)for

the case of the surface slot in the top surface of the calibration specimen.

The discrepancy at z = 120 mm (4.72 in) is attributed to the effect of signals

carried by the SH^ plate wave, which has been excluded from the analysis.

Similar results have been obtained for other calibration specimens, and the

39 effect was observed for weld flaws.

128

Figures 48 through 51 show the results of scattering measurements

performed on calibration specimens containing surface and buried slots (see

Table 12, Sec. 8) Figures 48 and 49 show the results for vertical slots at

two particular EMAT-to-weld separation distances z: 65 mm (2.56 in) and

100 mm (3.94 in). Figure 50 shows the results obtained for vertical slots

whose centers coincide with the midplane of the calibration specimen. The

cross section of a typical specimen is shown at the top of the figure). The

dependence of the reflection amplitude ratio on flaw through-wall depth, a, is

shown in the center of Fig. 50. It should be noted that this plot represents

an average of results obtained from a wide range (60-160 mm, 2.36 - 6.29 in)

of separation distances, z. The actual behavior of the flaw-amplitude

reflection ratio with the distance, z, is shown at the bottom of Fig. 50.

Characteristically, the amplitude reflection ratios in both plots (weld-crown

up and weld-crown down) fall off monotonical ly with the distance, z. This

effect is caused by the fact that flaws located near the midplane of the

weldment are not strongly coupled to the antisymmetric SH plate waves. The

monotonic decrease of the reflection amplitude with the distance, z, is caused

by two-dimensional diffraction effects.

Figure 51 shows the results obtained for canted slots located above the

midplane of the calibration specimens. The cross-sectional view of these

calibration specimens is shown at the top of the figure. As in Fig. 50, the

average amplitude reflection ratio is plotted as a function of the throughwall flaw depth, a. The evolution of the flaw signal with the distance, z, is

shown at the bottom of the figure. In this case, note that the amplitude of

the flaw signal does not decrease monotonically with the distance. This

effect is caused by the fact that the canted flaws were not located at the

midplane of the calibration specimen. As a result, they were weakly coupled

to the SH^ plate waves, causing the interference phenomenon illustrated in

Fig. 47.

129

t t

A 0 mm

t 60 mm

120 mm

180 mm

t 240 mm

300 mm

Fig. 47 - Phasor diagram representation of the signals observed for

different positions of the calibration slot within a plate

at different transmi tter-EMAT-to-weld-distances , z.

L tr = 210 mm.

130

The experimental flaw responses, shown in Figs. 48 through 51, can be

used in conjunction with the experimentally established ultrasonic background

levels, shown in Figs. 44 and 45, to determine the inherent flaw sensitivity

limits of the low-frequency SH-wave EMAT system. The flaw sensitivity limits

can be obtained approximately from the intersection points of the horizontal

lines in Figs. 48 through 51, which correspond to the maximum expected

background signal level and the measured flaw response curves. A more

precise, statistical treatment would require additional experimental

measurements and knowledge of the flaw distribution function. Thus, for

vertically oriented flaws extending beyond the transverse resolution width of

the ultrasonic beam, the minimum detectable flaw depths are 0.75 mm for

vertical surface flaws and 2 mm (0.079 in)for vertical buried flaws located

the midplane of the weldment. For buried flaws, which are canted at 30° with

respect to the surface normal, the corresponding minimum detectable

through-wall flaw depth is approximately 3 mm (0.118 in).

3.7 Detectability Limitations: Short Flaws

Because the ultrasonic beam, which insonifies the weld region, has finite

width, it is important to understand the relationship between the factors

determining the transverse resolution of the SH-wave EMAT system and the

inherent detectability limits for flaws of finite length. When the flaws are

very long, their length, i,, can be measured independently of the depth

parameter, a, by scanning parallel to the welding direction. However, when

the flaw is short, then it is not possible to determine i and a unambiguously.

Furthermore, the reflection amplitude ratio, R

s

, is reduced (Eo. 83), thereby

increasing the minimum detectable flaw depth.

To understand the physical principles that determine the transverse

resolution of the low-frequency SH-wave EMAT system, it is necessary to take

131
,
,

\ r~~i I1£ i

- z 210 mm

Flaw Depth, a, in

0 0.1 0.2 0.3

0 2 4 6 8

Flaw Depth, a, mm

Fig. 48 - Experimental calibration curves for sizing very long surface slots

located on the top surface of plate for z = 65 mm and z = 100 mm (flat-plate specimens 12-24).

132

Flaw Depth, a, in

0 0.1 0.2 0.3 0 0.1 0.2 0.3

Fig. 49 - Experimental calibration curves for sizing very long surface slots

located on the bottom surface of a plate for z = 65 mm and z = 1 00 mm (flat-plate specimens 12-24).

133

Transmitter

co c Flaw Depth,a, in

>

Flaw Depth, a, mm

Flaw-to-EMAT Distance, z, in

2 3 4 5 6 7

co o 4 cn

o 02

c

o

o 0

_a> % 0.4

DC

o> *D

|

0.2

Q.

E

< 0

40 60 80 100 120 140 160 180 200

7.5

0.8

Flaw-to-EMAT Distance, z, mm

Fig. 50 - Experimental calibration curve for sizing very long, vertical slots buried near the midplane of a plate and typical variations of the signal level

as a function of the transmitter-EMAT-to-weld distance, z. L = 210 mm. (Flat-plate specimens 8-11.) '- r

134

Transmitter

a> < Flaw Depth, a, mm

Flaw-to-EMAT Distance, z, in

Flaw-to*EMAT Distance, z, mm

Fig. 51 - Experimental calibration curve for sizing very long canted (30°) slots

buried near the midplane of a plate and typical variations of the signal

level as a function of the transmi tter-EMAT-to-wel d distance, z. L tr = 210 mm. (Flat-plate specimens 25-23.)

1 •?£;

E

E

cl

1U

a

(flaws not detected)

AP1 1104 Allowable Flaw Size

(inadequate penetration)

(flaw detection only)

Limited

Sizing

(flaw detection and sizing)

Detection Threshold (0.1)

100 200

FLAW LENGTH, mm

Fig. 52 - Summary of system performance characteristics for vertical

surface flaws.

136

into account the effects of two-dimensional diffraction. This can be

accomplished by generalizing Eqs. 72 and 75 to three dimensions. If the

transverse amplitude is Gaussian, as indicated by experimental evidence, then

the ultrasonic velocity field distribution near the flaw is given by:

v v (x,y,z,r) ;v An (l + j)(2)' = (e

n /2z)* cos (mry/t) x

n exp (-B

p w^x

2 /2xz2

) exp ( - j

n

x

2

/2z ) exp C-j ( - wt)] (86)

where w is the half-width of the ultrasonic beam near the transmitter EMAT.

o Corresponding expressions for the field quantities and can be

obtained from Eq. 86 by applying Hooke's law. Equation 86 can now be directly

substituted in Eq. 82. For elongated flaws, whose surface length, £, exceeds

the through-wall depth a by a factor of 4 or more, and assuming

z << (z + ) , the result is:

r =

where

00 00 ££exp a

n

r rmb

m

exp [ - jBm

(z + L tr )]

n=0 m=0

(87)

r' ^

nm a. nm

Bm J

a7

f

-l

~c

exp (-jB

n xV2z) dx

( 88 )

The integral in Eq. 88 is recognized as a Fresnel integral whose behavior can

be deduced from the Cornu spiral.^ Specifically, the magnitude of the

integral in Eq. 88 increases monotonical ly for i < ( 4ttz/ 6

n

) ^ £ (2 az) 2 and then

converges in an oscillatory manner to an asymptotic value. In writing Eq. 88,

the effect of the exponential profile of the ultrasonic beam, expressed by the

last term in Eq. 86, was neglected. This was done on purpose, to emphasize

the fact that the transverse resolution of the SH-wave-EMAT system, ( 2xz

2

,

137

increases at a significantly lower rate than the ultrasonic beam width,

2xt/-TTOio

. From the above, it can be concluded that flaws that are longer than

(2xz)^, behave essentially as two-dimensional, planar flaws. At 454 kHz and

z = 50 mm (1.97 in), ( 2xz ) ^ £

27.4 mm (1.078 in).

Because the form of the mode III displacement jump across the face of

planar flaw is not generally known, it is not possible to obtain a closed-form

expression for as a function of all parameters describing a planar flaw.

However, when the flaw length, £, is sufficiently short, it is possible to

assume that the phases and amplitudes of the incident ultrasonic fields are

constant over the face of the flaw. This assumption is equivalent to the

Fraunhofer approximation. It is then possible to calculate the scattering

coefficient Tqq approximately, it alone determines the detectability of planar

flaws located near the midplane of the plate.

To evaluate Tqq using Eq. 82, a well-known, closed-form elastostatic

solution, can be used that is valid for elliptical, buried cracks of length £ 50

and'through-wall depth a:

Au ( 0 ) dS

8-TT£a 2

3G<t> ( 1 - a

2 /£ 2

)

1 - v

2 - v

T (0)

'l (89)

where v is the Poisson's ratio and $(1 - a

2 /£ 2

) is the complete elliptical

integral of the second kind.

Equation 89 can be directly substituted into Eq. 82. The result is:

r , * 8*

2 r oo $

-

J 1

‘ i - v /eo\ K

6 o T

\*)

£a 2 (90)

x t<t> ( 1 - a

2 /£ 2

)

The corresponding reflection amplitude ratio, R , for an elliptical flaw,

buried near the midplane of the plate and canted at an angle, ©

n

, with respect

to the plate surface normal, is given by:

138

/, .x 8tt 2 1-v

Sid 2 COS 2 ©

xt ( 2xz ) ^ »(l-a 2 /£ 2

)

n a^ exp

n=o

a

n

b

n [6n

L tr/e

o (z+Ltr

)]i ex P ( -j B n

L

tr )

( 91 )

For the range of experimental parameters stated previously, and £ ^ 25 mm

(0.984 in) a = 2.5 mm (0.098 in), the coefficient R

$

is 0.21. This value is

in excellent agreement with the experimentally observed value in Fig. 50.

Thus, Eq. 19 can be used to estimate the detectability of sharp flaws that are

shorter than (2xz) . To estimate the detectability of shallow surface flaws,

the contribution of the scattering coefficients Tq^, T^q, and must be

included in the estimate of R . Since the function $(1 - a

2 /£ 2

) = tt/ 2 at

a = £ and tends monotonical ly to unity as a/£ tends to zero, the reflection

amplitude ratio, R , for elongated, sharp flaws is directly proportional to £

when 4a < £ < (2xz) 2 .^ Equation 91 does not include a dependence on the beam

width, 2 wq. This follows from the fact that the flaw is insonified by the

2 far-field of the transmitter EMAT. This assumption holds when (2 z/6qWq2)

>> 1, which is clearly satisfied when z > 50 mm (1.97 in), X = 7.4 mm,

(0.291 in) and Wq = 7.5 mm (0.295 in).

3.8 System Performance Limitations and Standards of Acceptability

The model-based results for flaws whose length cannot be resolved can be

combined with the empirically derived results for two-dimensional, planar

flaws that are longer than (2xz)^, to construct a phenomenological model

defining the limits of detectability for surface and buried flaws. It is

convenient to present the detectability limits graphically with flaw depth, a,

plotted as the y-axis and the flaw length, £, plotted as the x-axis. The

detectability limits can then be directly related to current flaw acceptance

standards based on the industry-wide such as those in Section 6 of API-1104

139

4 68

flaw acceptance criteria based on f itness-for-service considerations. *

This is done in Figs. 52 and 53 for vertically oriented surface and buried

planar flaws.

In Figs. 52 and 53 the allowable flaw sizes are defined by length vs.

2

through-wall depth curves, except for Section 6 of the API-1104 Standard,

which does not address the flaw through-wall depth. The allowable flaw-size

curves were based on a yielded-! igament fracture mechanics model using

arbitrary, but realistic, parameters to characterize the toughness of the

68 materials and to estimate the stresses acting on the girth welds. These

curves are used here only to illustrate the relationships between the

capabilities of the SH-wave EMAT inspection system and different flaw

acceptance criteria.

In Figs. 52 and 53, the range of flaw sizes that can be detected is

bounded by the heavy solid lines. The range of flaw sizes that cannot be

detected is indicated by the shaded areas. It is interesting to note that the

maximum flaw length allowable by Section 6 of the API-1104 Standard coincides

with the resolution limit of the SH-wave EMAT system, ( 2 Xz

2 when z = 50 mm

(1.97 in) and the operating frequency is approximately 500 kHz. Shorter flaws

are considerably harder to detect. As a consequence, the 454-kHz SH-wave EMAT

system may also be appropriately used in conjunction with conventional,

workmanship-based acceptance criteria, such as Section 6 of the API-1104

Standard.

It is interesting to observe (Figs. 48 and 49) that the amplitude of

ultrasonic signals backscattered from surface flaws increases monotonical ly in

the range 0 < a < A/3 and reaches a maximum when a ^ A/3. On the other hand,

in the case of buried flaws (Fig. 50), the amplitude of the backscattered

signals increases monotonical ly in the range 0 < a < A/2 and reaches a maximum

value at a ^ a/2. When the above limits are exceeded, the interaction of the

incident ultrasonic fields with the flaw becomes very complicated and the

140

THROUGH-WALL

FLAW

DEPTH,

mm

AP1 1104 Allowable Flaw Size

(incomplete fusion)

(flaw detection only)

Limited Sizing

(flaw detection and sizing)

odel Allowable

Flaw Size

Detection

Threshold (0.1)

100 200

FLAW LENGTH, mm

Fig. 53 - Summary of system performance characteristics for

buried planar flaws.

141

low-frequency (long-wavelength) inversion algorithms for flaw sizing become

inoperative. Thus, a < A/3 for surface flaws and a < A/2 for buried flaws

define the maximum flaw depths that can be sized. Deeper flaws can still be

detected reliably, but cannot be sized unambiguously. These limits are

indicated in Figs. 52 and 53 by horizontal lines parallel to the x-axis. The

hatched areas in Figs. 52 and 53 represent a range of detectable flaws when

the length i < (2xz) 2 and, therefore, cannot be determined independently of

the through-wall depth, a. (At 454 kHz the shear wavelength. A, is

approximately 7.4 mm (0.291 in). For comparison, the average through-wall

depth of one SMA weld pass is approximately 3 mm, 0.118 in).

A consideration of the performance characteristics summarized in Figs. 52

and 53 reveals that the operational wavelength. A, and the EMAT-to-weld

distance, z, must be selected appropriately to assure the detection of flaws

that are significant from the point of view of a particular acceptance

standard for sharp flaws. The resolution of the system is reduced at longer

wavelengths, but the range of flaw dimensions that can be sized is increased.

In addition, the detectability limitations are determined principally by the

mass-loading effects of the weld reinforcement and the presence of alignment

mismatch. Consequently, consideration must be given to limiting the maximum

dimensions of the weld reinforcement and alignment mismatch by employing

appropriate process and tolerance controls.

3.9 Inspection Protocol

The development of the low-frequency SH-wave EMAT system was guided by a

need for a reliable ultrasonic inspection tool that would be capable of

detecting significant planar flaws in pipeline girth welds. The detection of

significant flaws necessarily involves a decision process based on a set of

nondestructive measurements and a set of acceptance criteria. The output of

142

EMAT Array for

translation along girth

Simplified EMAT System for

raster inspection along girth

Fig. 54 - Two possible implementations of a practical SH-wave

EMAT inspection system.

143

Fig. 55 - An inspection protocol for fully automated

inspection of pipeline girth welds.

144

the detection process is always a binary decision, yes or no, and an

associated measure of confidence. The decision is based on models describing

the nondestructive measurements, the failure processes and associated safety

factors, and the knowledge of the statistical flaw distribution.**^

Acceptance criteria for sharp flaws are normally specified in terms of

the principal length and through-wall depth dimensions. In addition, certain

failure models treat surface flaws more severely than buried flaws. As a

consequence, the inspection system must be able to classify the flaws in terms

of their relative position within the weld and to estimate their principal

dimensions. This can be done simultaneously using ultrasonic measurements

from different locations along scan lines that are perpendicular and parallel

to the pipe circumference. Two possible experimental configurations can be

used to accomplish this task; they are illustrated in Fig. 54.

Figure 54a shows the preferred inspection configuration. The

configuration of Fig. 54a uses two arrays of EMATs that are translated

mechanically along the direction parallel to the girth weld. Each array is

composed of a number of independent sets of EMATs that are positioned at

different distances from the weld. A simplified inspection configuration is

shown in Fig. 54b. The simplified configuration uses only one set of EMATs,

as illustrated in Fig. 54, that is scanned along the girth weld in a "zig-zag"

manner. In this case, the measurements are repeated on the opposite side of

the weld. It is evident that both inspection configurations can perform the

scattering measurements that are needed.

A flow chart, exemplifying a possible inspection protocol, is shown in

Fig. 55. Because most girth welds are free of significant sharp flaws, the

inspection protocol involves a preliminary "detection" threshold, which

permits by-passing the decision process involving acceptance criteria. The

145
.

m
protocol assumes that ultrasonic measurements are made, as illustrated in

Fig. 55 and processed to compute the reflection amplitude ratio, R

$

, for each

measurement location. If a signal level is observed that exceeds the

preliminary detection threshold, then the flaw is classified and measured

using the theory developed in this report. The decision concerning the

significance of the flaw can then be made. If the measure of confidence is

low, then additional available information may be included in the decision

process. The additional information may include visual observations, process

control information (welder's notes), radiographic data, and conventional

ultrasonic data. For this purpose, a subsidiary "loop" is included in the

inspection protocol. At the end, the results of the decision processes are

appropriately displayed and recorded. Permanent recording of the preprocessed low-frequency ultrasonic data is also envisioned.

Since the decision processes must include consideration of factors that

cannot be adequately modeled or simulated in a laboratory environment, successful implementation of the proposed inspection protocol will involve considerable

field experience. In addition, independent data describing the statistical

distribution of flaws will be required from destructive assays.

3. 10 Summary

Significant progress has been made in developing a low-frequency

ultrasonic system optimized for inspection of girth welds in large-diameter

pipelines. This system offers the following advantages: 1) rapid inspection

rates, 2) elimination of transducer coupling agents, 3) insensitivity to weld

reinforcements, 4) minimum dependence on operator skill and judgement, 5) full

automatabil ity , 6) tunability of the system to signal the presence of

rejectable flaws based on any established fracture mechanics accept/reject

criterion, and 7) permanent data recording. Other ultrasonic inspection

146

systems provide comparable overall performance. In contrast to existing

systems, the present system is designed to provide the specific information on

flaw content and sizing required by the accept/reject criteria of a

f itness-for-service analysis of flaw significance.

Phenomenological models of the ultrasonic measurement process have been

devised to extract the dimension and position information from the ultrasonic

data from planar flaws. In addition, the inherent performance limitations of

the system have been established theoretically and empirically in a controlled

laboratory environment. The results indicate that the new system offers

considerable advantages over conventional ultrasonic and radiographic methods.

In particular, it uses noncontacting electromagnetic-acoustic transducers

whose efficiencies can be calibrated for each measurement. In addition, it

operates at low ultrasonic frequencies, which reduce the complexity of data

interpretation. At its present state of development, the SH-wave EMAT

inspection system has the following deficiencies: 1) performance has not been

demonstrated with real flaws under field conditions, 2) field proven hardware

remains to be developed, 3) personnel trained in the system are not readily

available, 4) equipment calibration standards remain to be developed, and 5)

sensitivity of the system to detecting tight cracks closed by compression

stresses has not been determined.

147

4. SIGNIFICANCE OF BLUNT FLAWS

M. B. Kasen

4.1 Introduction,

It has thus far been assumed that both sharp and blunt (nonplanar) flaws

would be treated as significant and subject to an accept/reject criteria based

on analytical fracture mechanics principles. But a number of studies have

concluded that flaws such as porosity, slag, and arc burns have a very low

probability of initiating cracks in tough weld metal and may be considered

innocuous in a fitness-for-service analysi s .^-76WBS therefore of

interest to determine if this proposition could be defended in pipeline girth

welds.

4.2 Experimental Procedures

A series of girth welds containing large contents of porosity, slag,

and arc burns were fabricated in 1016-mm (40-in) API 5LX-65 pipe and in

1220 mm (48-in) API 5LX-70 pipe of 15.9-mm (0.625-in) wall thickness. The

objective was to obtain welds having flaw contents equal to or in excess of

the worst case obtainable under field welding conditions. Welds meeting API

1104 workmanship quality were also produced. Automatic and manual (SKA)

processes were represented, with fixturing, consumables, and personnel

reflecting field procedures as much as possible. Welding parameters and joint

designs, as reported by the fabricator, appear in Table 6. Typical welding

electrode compositions are given in Table 7. Chemical composition of the base

and weld metals are given in Table 1 (Section 2). Tensile properties of the

weld metal and typical fracture toughness data on the welds are summarized in

Table 8. A detailed discussion of toughness is given in Appendix H. Flaw

contents obtained are illustrated in Fig. 56.

148

Table

6

-

Welding

Parameters

4_> oooooooooo

c u5vO(ON^r>*t^r^r>.vD ^ rTM* p— r— pTM" p— r— rTM" r~ \\ V, s^, \v» \v. V oooooooooo LD LD LD ID LD LD LD LD LD LD o

0) OOOOOOOOOO c\jrororororororororo

4-> V, Sy Ns \ Ns V*. Ns, \ Ns^ ^ ^•VDVDVOVDVDVDCDVDCD

o C\J C\J C\J C\J C\J OJ CM OJ OJ OJ >

OJ C

T3 -i- o CMOUCMlOlDOVDlOlOlD

+J • rororo.— i— r-i— .— r— i—

n e Vs Vs V. V. Vs V VVs \

0J HJ inininrororororororo

LU Q

D. Cl C. Cl

CD

•r* t— •»— *r- s0—** i~ C. S_ s- Cl Q 0? +J .*_> 4-> 4-> (O CO > 1/1 l/l l/) l/l U

ns w

QJ LDlDlOOOOOOOO

> E

(C Q. r— O O CO CO 00 00 CO ID

i~ -i- r— t— i—

t—

to

CO f— M ro <3- LfUD S CO Ol O

TO CL

O)cn in in in id in

o ro

C- -t-> ro c\j csj csj •— o

<t I— CSJ CVJ CVJ OJ CSJ CSJ o>

#* TD

0)

QJ

Li_ E uo o o o o LD

CL csj in in in nj

QJ *i- ld «a- «a- ro ro

S-

•rT3

QJ

0JCl CO E Ninmcoi— cn r- CL f= r- r— n- r—

QJ •<- >

TO

s1—

LO

V)

ns i— m co in id

Cl.

C\J

149

Wire

diameter:

0.035

in

Automatic

Shielding

gas:

50%

Ar,

50%

CO

Table

6

-

Welding

Parameters

*

+J• oooooooooo

c tOtDvor«^r>^f^.r^pv.rv.vo

•> V.V.'S.W

t oooooooooo

9 inininininininininin U

01 oooooooooo

u «o CJnronfnnnnnn

U 4-> ^IDOVDOOOOt&lO WVV. —WWN

o CM CM CM CM CM CM CM CM CM CM >

Ol c

T5 •»- O

u • CMCMCMIDIDIDVDVOVDID

4-> • nnnr-l-r-r-f-r-iy g

0» ID iftiniftnromnnnn

UJ o

a. o. c. o.

0) -—* <D . l t t. t a

Q. (V 4J4J4J+J «

> U) W (/I w u

to tntOlOOOOOOOO

£ K '-OOCONOOCOCOCOtf

U t-

*-

V)

1/1 ^Nn^irwNcooto

<o o.

0) 0 tn in 0 in 0

u io • • • • • L. 4-> CO CM CM CM *•— O « •— CM CM CM CM CM CM 0

• o0

0) f- r in 0 0 0 0 in

0. cm in in in 9 cm

(L> i“ in *3- PO PO *

• -O0) O

O.

to E NlftP^POr-Ol ^ &. ^ F* f» ^ 0) 1- >

ID

k. »-

I/I V)

<e r- CM PO ^ in ID

0.

CM

150

Wire

diameter:

0.035

In

Automatic

Shielding

gas:

50%

Ar

t

50%

CO,

lable 7 - Typical Welding Electrode Compositions

Element

Process

Shielded Metal -Arc* Automatic*

Mn 0.47-0.51 1.35

Si 0.07-0.10 0.70

Cr 0.12-0.13 —

Ni 1.03-1. /0 1.15

Mo 0.01 —

V 0.02-0.03 —

C — 0.10

S — 0.015

P — 0.015

Values reported by the fabricator in percent by weight

*AWS E8010G

+AWS E705-6

151

Table 8 - Typical Weld Metal Properties

Yield Strength at Toughness, COD, Critical COD, nn (in)

0.2% Offset, 25°C mn (in) Slow, Stable Crack

Weld Type Hardness MPa Psi x 10 3 -18°C 0°C Growth

Automatic HRC 30 725.0 105.5 0.082 (0.0032) 0.094 (0.0037) 0.074 (0.0029) at 0°C

Manual HRB 91 465.0 67.5 0.124 (0.0049) 0.0187 (0.0074) 0.079 (0.0031) at -18°C

152

o>

to oo

*ou

iCT> OcID CC

LT)

inouo

o.X)2 Co O •!- f- 4-> *-> ID

ID L- E 3 o u

+J in 3 X)

«£

O

(/>

-t-> C

ajEi— • J- sz

QJ Cl Cl 03 X S- gj cr> o 0) *i- x: *o

+-> rv

i~ c

•aaj 00

=3 3: 0J>

1/) E 5 O

fO 4-J r— +J 4- o

-Q

T3r“

QJ • •» s QJ

i_ 4- 3 O 4-> U

00 1C r— J- aj 4- > u QJ JC

CL ' 2

ro r— OJ

L. fO f— o U >

“O CL Cl >> O

tt 1— 1—

ro ' LO

•* in

>»

4-> #

«/> cn O •I"* u Li- o

Cl

•O

ID Z co

X3

—

QJ 4-J *-> ID ID L- i— 3 3 D E </i •l- _Q O0 O

JTO

82.5 cm

Drilled

Area 20.3 cm

Drill Sequence

Fig. 57 Drill pattern used to simulate buried porosity in

manual (SMA) welds in API 5LX-70 pipe.

154

Extensometer C

50.8 Gage 1 }

Gage 2

Fig. 58 Specimen and instrumentation used for low-cycle

studies of the significance of blunt weld flaws

initiation. Dimensions are in mm.

fatigue

in crack

155

Fig. 59 - Schematic of hysteresis loops formed during low-cycle, straincontrolled fatigue test program. Crack initiation was defined

as corresponding to a 10% load drop to maintain the desired

strain range. Specimens were broken in tension after a 40% load

drop. Tests were run at strain levels just over yield (+e = +0.0022) and at twice yield (+2e = +0.009).

156

A level of porosity obscuring 12 to 15% of the projected area of the weld

radiographs was desired, since this level represents more than twice the

obscuration level that would be practically allowable based on the probability

of excess porosity interfering with radiographic detection of sharp flaws (see

Section 6). This level was easily obtained during automatic welding by

disturbing the shielding gas during deposition of two internal weld passes.

Several attempts were made to produce this level of obscuration during manual

welding by improper welding technique, excessive drying of the E8010G

electrodes, or by fracturing 'the flux coating of the electrodes. Although a

large amount of porosity was produced in a given pass, the final welds were

found to contain far less than the desired level of obscuration. This flaw

was therefore simulated by drilling a pattern of 3.175-mm (0.125-in) diameter

holes in five successive interior weld passes and overwelding. The procedure

is illustrated in Fig. 57, and the result is illustrated in Fig. 56c.

Slag was produced by improper cleaning of an interior weld pass. As

illustrated in Fig. 56b, this typically resulted in formation of "wagon

tracks" due to slag entrapment from the E8010G electrodes at the sides of the

weld bead. Lying adjacent to the fusion line, this type of slag residue has

the highest probability of contributing to brittle fracture because of the

stress concentration caused by the weld reinforcement. Arc burns were

introduced at random on the heat-affected zones of sound welds after the

welding had been completed. This simulated the worst case, because there was

no annealing of the brittle martensite phase that forms under the burns due to

rapid cooling.

Large, transverse weld specimens, having dimensions and instrumentation

as illustrated in Fig. 58, were prepared from selected regions of the welded

pipes. These specimens were subjected to fully reversed, strain-controlled,

157

low-cycle fatigue, producing hysteresis loops as illustrated in Fig. 59. A

temperature of -2 ±1°C was maintained to simulate the lowest anticipated

service temperature of the ANGTS line. This was accomplished by controlled

metering of liquid-nitrogen vapor into a chamber surrounding the specimen.

The cyclic rate was held to 0.2 Hz or less to minimize internal heating. The

imposed strain was either slightly over weld-metal yield (±0.0022) or

approximately twice weld-metal yield (±0.0045). The total strain range was,

therefore, either 0.0044 or 0.0090. This technique was selected because,

although pipelines are not normally subjected to such severe cyclic loading,

any flaw found innocuous under the imposed conditions can conservatively be

considered innocuous in essentially statically loaded structures.

^A similar

approach has been used to demonstrate the inherently innocuous nature of blunt

75 76

flaws in pressure vessel weldments. *

As cyclic testing proceeded, development of cracks in the weld region

reduced the effective cross-sectional area, causing a drop in the load

required to maintain the selected strain range. Monitoring of the load therefore provided an indication of the number of cyclic reversals required to

initiate cracking in each weldment. A 10% load drop, corresponding to

development of about a 10% cracked area in the specimen cross section, was

used as the criterion for crack initiation. This criterion was selected

because, once initiated, crack-growth rates were relatively rapid under the

severe test conditions, regardless of flaw content. Cyclic straining was

continued until a load drop corresponding to a 40% cracked area was obtained

to permit the influence of flaw type and content on crack propagation to be

studied. Testing was stopped at 3000 cycles if this criterion had not been

reached. Specimens were then broken in tension. The desired strain range was

initially established by strain gages located at the weld, as illustrated in

158 •

Fig. 60 - Hysteresis loop plot from extensometer on a sound manual weld

tested at 0.0090 total strain with weld reinforcement removed.

Specimen No. 95103.

159

Fig. 58, while control of the 4.45-MN (1 x 10^ Ibf) servohydraul ic test

machine was provided by an extensometer spanning the weld region. Figure 60

illustrates the actual appearance of hysteresis loops produced by the

extensometer and load cell during a test of a sound manual weld.

Specimen asymmetry due to pipe curvature produced a bending moment during

compression, tending to cause failure by column buckling rather than in

tension-compression. This was countered by development of the specimensupport system illustrated in Fig. 61 and by ensuring that the specimen halves

were in good alignment prior to a test. The latter was achieved by slightly

bending the specimens to remove the distortion caused by release of residual

welding stresses when the specimens were removed from the pipe.

All the specimens were radiographed prior to testing. Several specimens

were stereoradiographed after achieving the 40% load drop but before final

fracture to determine if useful information could be obtained on the crack

path relative to buried flaws. However, the cracked region could not be

defined with a resolution justifying continuation of this procedure.

Weldments were tested with the weld reinforcement both intact and removed

by flush grinding. Results with the reinforcement intact were of most

interest in interpreting flaw significance, since this condition exists in the

field. Removing the reinforcement forced failure through the flawed region

and provided useful information on flaw interaction during crack initiation

and propagation.

4.3 Low-cycle Fatigue Results

Results of tests conducted with the weld reinforcement intact are

summarized in Table 9. Similar data for specimens tested with the

reinforcement removed appear in Table 10. Results of tests on unwelded parent

160

Fig. 61 - Specimen support system required to prevent

failure in column buckling during low-cycle

fatigue testing. Dimensions are in millimeters.

161

Fig. 61 - Specimen support system required to prevent

failure in column buckling during low-cycle

fatigue testing. Dimensions are in millimeters.

162

Table 9 - Results of Low-Cycle Fatigue Program Conducted with Welds Having

Reinforcement Intact and with Parent Pipe Materials

Flaw Type Specimen No. Strain, ±%

Cycles

10% —to Load Drop m

Sound 1-3543 0.45 114 477

(Automatic) 1-5159 0.45 159 516

1-12103 0.31 585 1005

1-4351 0.22 —

X2-121129 0.22 885 (1)

Porosity 4-0210 0.45 87 500

(Automatic) 4-2028 0.22 3012 4852

X2-6674 0.22 397 * X2-2028 0.22 863

Sound 3-6775 0.45 106 128

(Manual

X7-8492 0.45 75 272

3-8391 0.22 1118 1276

X7-6876 0.22 1358 1933

Slag 5-8290 0.45 110 247

(Manual

5-0816 0.45 30 49

X7-5260 0.45 88 185

5-9098 0.22 1015 1389

5-98106 0.22 760 1236

Porosi ty+ XP-2 0.44 70 81

(Manual

XP-4 0.45 99 127

XP-6 0.22 812 821

XP-8 0.22 1095 1117

Arc Burn 3-2028 0.45 231 242

(Manual

3-2836 0.45 194 295

3-3644 0.22 1159 1500

Arc Burn PA-1 0.45 853 2830

(Parent X-65

Pipe)

PA0.22 (1) • “ “

Parent X-65 P-2 0.45 1740 1913

Pipe P-1 0.22 3124 • — —

*Not achieved after 3000 cycles. Test discontinued.

tArtif icially introduced. See text.

163

Table 10 - Results of Low-Cycle Fatigue Program Conducted with Welds

Having Reinforcement Removed

Flaw Type Specimen No. Strain, ±%

Cycles

10%

to Load Drop

40%

Sound 1-111119 0.34 120 2130

(automatic) 1-1119 0.22 * —

1-0311 0.22 t —

Porosity 4-4452 0.45 155 165

(automatic) 4-1220 0.45 30 49

4-5563 0.45 117 276

4-2836 0.22 1026 1525

4-3644 0.22 397 1505

Sound 3-7583 0.45 81 136

(manual

3-95103 0.45 118 215

3-0008 0.22 1766 2073

5-4351 0.22 2621 t

Slag 5-115123 0.45 26 44

(manual

5-7179 0.45 104 198

5-0008 0.45 250 363

5-6371 0.22 352 442

5-5563 0.22 132 169

Arc Burn 3-1018 0.45 169 189

(manual

*Not achieved after

+Not achieved after

^Not achieved after

2426

1754

3000

cycles.

cycles.

cycles.

Test discontinued due to malfunction.

Test discontinued due to malfunction.

Test discontinued.

7 64

pipe material and of parent material containing arc burns are included in

Table 9. Figure 62 presents the data of Table 10 in graphical form,

illustrating the number of cyclic reversals required to initiate cracking (10%

load drop) as a function of strain range. Figures 62 a through c illustrate

the results with individual flaw types, while all results are combined in

Fig. 62d. Data of Table 10 do not appear in graphical form; however, the

lower bound performance of this group tested with reinforcement removed is

indicated by the dashed line on Fig. 62d.

These data show that some welds containing porosity or slag initiated

cracking at lower cycles than did sound welds; others withstood higher cyclic

loading. A minimum of 30 reversals at twice yield or about 450 reversals at

just over yield were required to initiate cracking, regardless of the content

of porosity, slag, or arc strikes. Comparison of the performance of welded

specimens with that of unwelded plate illustrates that the strain

discontinuity at the fusion line caused by the contour of the weld

reinforcement was the dominating factor in lowering fatigue life in the welds.

Weld failure occurred predominantly at this location, regardless of flaw

content. The failure path was occasionally observed to pass through some of

the porosity and slag. Neither crack initiation nor propagation was affected

by the presence of arc burns in welded specimens. Arc burns initiated

cracking only when they were introduced at the minimum cross section of

unwelded plate. Here, cracking was observed to initiate at the strike after

853 cycles at twice yield (Table 9). No effect was detected at the lower

strain level in arc-struck parent plate.

It has frequently been observed that the number of cycles defining lowcycle fatigue life, N , can be expressed as a function of total strain range.

165

$5O

too>o>N o o I

c\j to

CD

166

Cyclic

reversals

required

to

initiate

cracking

(10%

load

drop)

asa

function

of

strain

range

for

parent

metal,

sound

welds,

and

welds

containing

blunt

flaws.

Results

are

combined

in

(d).

Solid

line

is

lower

bound

for

initiation

in

tests

with

the

reinforcement

intact.

Dashed

line

in

(d)

indicates

lower

bound

for

tests

welded

with

reinforcement

removed.

Ae

t

, by a relationship of the form

Ae

t

N

c

= constant (92)

implying a straight line relationship between log bz^ and log N

c

, having a

slope of -m. Gurney^ has noted that results reported by a number of

investigators for steel weldments of different configurations fall within a

relatively narrow scatter band on such a plot, producing a slope of

approximately -1/3. Figure 63 illustrates that a very similar result is

obtained when the average number of cycles to crack initiation is plotted as

a function of total strain range for manual welds tested with the

reinforcement intact in the present program. Welds containing porosity and

0 3

slag produced a relationship of the form Ae^N * = 0.04. However, the

averaged data from the sound welds (including those containing arc burns)

0 4

appeared to form a different population following the relationship Ae^N

*

0.07. This resulted in a divergence between the sound and flawed weld data at

the higher (±2e ) strain range, suggesting that the high porosity and slag

content of the welds had somewhat reduced the cyclic life to crack initiation

under this test condition. No effect of the flaw content is observed at the

Figure 64, illustrating the average number of cycles to crack initiation

as a function of total strain range for manual welds tested with the reinforcement removed, shows that forcing the failure to occur through the flawed weld

interior rather than at the fusion line resulted in fewer cyclic reversals to

initiate cracking than in tests with the reinforcement intact. In this test

mode, it is seen that slag is a more significant flaw than porosity in

reducing cyclic life to crack initiation. But, even under this severe test

condition, 26 cyclic reversals were required to initiate cracking at twice

yield, whereas 132 cycles were required at just over yield. This suggests

167

TOTAL

STRAIN

RANGE,

Ae

CYCLES TO 10% LOAD DROP, N

c

Fig. 63 - Averages of lifetime to crack initiation following

Eq. 92 for specimens with weld reinforcement intact.

168

TOTAL

STRAIN

RANGE,

A€,

CYCLES TO 10% LOAD DROP, Nc

Fig. 64 - Averages of lifetime to crack initiation following

Eq. 92 for specimens with weld reinforcements removed.

169

that, even if the reinforcement were removed, porosity, slag or arc burns

would pose little danger to the integrity of pipelines fabricated with welds

having the toughness studied in this work.

4.4 Fractoqraphic Analysis

Fractographic examination showed little flaw interaction at the lower

strain level ( ±ey)* As shown in Fig. 65, the cracks in automatic welds

containing a large amount of porosity were observed to propagate as single

crack fronts with negligible initiation from individual pores. The higher

strain (±2^) did cause cracks to form from individual pores in both the

manual and automatic welds, with the cracks coalescing to form the fracture

surface.

The presence of halos around slag stringers and micropores was noted on

the fractured surfaces of all manual welds. This phenomena, illustrated in

78 79

Fig. 66, is described in the literature as "fisheyes" * andis attributable

to hydrogen embrittlement of the region adjacent to pores and surfaces to

which hydrogen has segregated during the welding operation. Their appearance

here suggests that hydrogen segregated in this manner is very stable, because

the specimens were tested one and one half years after welding. The mechanism

of fisheye formation is not well understood; however, lack of fatigue

indications on the fisheye surface is indicative of a pop-in type of crack,

perhaps reflecting a stress-induced diffusion of hydrogen from its segregated

site into the surrounding weld metal during loading. It is generally assumed

78 that fisheyes only form in weld metal stressed beyond yield, however, this

has not been firmly established. Whatever their cause, fractographic analysis

indicated that the fisheyes contributed relatively little to fatigue crack

initiation unless they penetrated the specimen surface. For example, the main

170

cn CD c X •1— X

inX •T— to

X

CD o ^ s tO X Cl o " * O ro E a> c Cl X •—I ro 2 in • o a> c E E

•r- E O

to ro CL E % - X r— W X <o c 3 o •o

CM E •f— CM X > • •T— o, -a + 

u c

to r—

X E

ro

O Eo

TD 0) E

o>

X

X

XX to E

CD CDO

X E

•T—

X

D 2 ro E o •r— lO X X in r*

>> E

4-> *—

•r— •r— ro

</) *» -- O

U E E ro • o r— E in Cl O a> X r—

CD o a>U E X 3 >>

•r— CD

u E 0) •f—

•i

X

X CD

roX ro 3 X X CD E >> •r- O X X X o O ro D X

*o o» CD E U LD Id •r— E O 2 X CD LD

O) o r—

u "O •i

•f— > E X E CD a>

IT3

O

X E •r— X o CD 3 ro X CD O 3 E X rO X

O) •r— VC E X 2 CO

•r—

X 1 X E X X O o CD E

r“ • F— O) X X E E

fO E r—

CD CDOO

rO E X CD CL 01 CL o CD X CO E E X CL ro CD r— r— • J* E X U CD E E

rO

O o E E E E

<_) ro

XX I

LT> VC

CD

171

fatigue crack in the specimen illustrated in Fig. 66b initiated and propagated

in a region remote from the large internal fisheye.

In addition to the fatigue work described above, two specimens each

representing sound automatic welds and welds containing automatic and manual

porosity at levels shown in Fig. 56 were pulled to failure in tension at -2

±1°C. Results, given in Table 11, showed no effect of the flaw content on the

static tensile strength. All welds failed through the heat-affected zone of

the API 5LX-70 pipe.

4.5 Discussion

These results show that the probability of crack initiation from

porosity, slag, and arc burns during low-cycle fatigue was very low in the

manual and automatic pipeline girth welds tested in this program. Although

some reduction in life to crack initiation was noted when welds containing

large quantities of porosity and slag were tested at the highest strain range

(±2e^), the geometrical discontinuity formed by the weld reinforcement was

found to dominate the fatigue failure mode in all tests conducted with the

reinforcement intact.

This is consistent with the findings of others on the relative effect of

the geometrical discontinuity created by the reinforcement as compared with

that of buried blunt flaws. 72,76,77,80

j p a ^ sence 0 f brittle fracture

initiation, the primary structural effect of such flaws is to reduce the

effective weld cross section. However, porosity content causing a 12 to 15&

radiographic obscuration will reduce the cross-sectional area by only about 1 %

in welds of the thickness studied in the present work, ^ whereas the weld

reinforcement itself typically contributes 5 to 8% to the thickness of the

weld region. Therefore, a negligible effect could be predicted for

circumstances where the weld-metal yield strength matches or exceeds that of

172

the parent pipe material. Table 9 shows that this was the case in the present

study.

Even with the reinforcement removed to force failure through the highly

flawed weld metal, the welds showed a high tolerance to fatigue crack initiation

from the blunt flaws. Because strain levels were well above yield, these data

indicate that the probability of crack initiation from such flaws would be

negligible if such welds were subjected to the essentially static loading

conditions of operating pipelines.

Therefore, the results of this study suggest that porosity, slag, and arc

burns are innocuous as fracture initiation sites in girth welds fabricated

with materials of the toughness equal to or exceeding that listed in Table 8.

The validity of this conclusion has not been experimentally demonstrated for

71 72

welds of lower toughness. However, as noted by Harrison, * the rounded,

blunt shapes of such flaws and their inherently small critical through-wall

dimensions (see Section 5) make them size-for-size less harmful than cracks of

equal dimensions. He observes that if the critical flaw size is so small as

to indicate serious risk of fracture initiation from porosity and slag, it is

doubtful that the selected material has a toughness appropriate for the

application. Harrison concludes that porosity and slag required no special

consideration in regard to brittle fracture. This concept has been accepted

by Commission X of the International Institute of Welding (IIW) following the

72 73 recommendations of an IIW Working Group on Significance of Defects. * This

Group, representing Japan, England, and Belgium, based its conclusions on: 1)

the examination of approximately one hundred cases of fatigue by brittle

fracture in ferritic steel weldments, during which no instance of fracture

initiation at porosity or slag was observed, 2) results of tests at many

laboratories demonstrating that such flaws have little or no effect on static

173

failure load, and significantly less effect on fatigue life than the surface

profile and, 3) consideration of theoretical elastic and fracture mechanics

analysis. The group concluded: "Provided materials of adequate notch

ductility or fracture toughness for the particular service application have

been used to tolerate the smallest planar defects which can reliably and

practically detected, porosity, slag inclusions, and surface irregularities,

in amounts up to which they can be clearly identified as the only defects

72

present, will have no significance on failure by brittle fracture".

The present work appears to substantiate these general conclusions with

one exception: The detection of fisheye crack formation from the surface of

entrapped slag in welds made with the E8010G high-hydrogen electrodes suggests

that it might be prudent to reserve judgement on the potential significance of

slag until the conditions for formation of fisheyes are better defined.

Although such cracks were developed during low-cycle fatigue in the present

work, they are found on tensile, bend test and similar fracture surfaces

87 produced at low strain rates. The possibility of their formation as the

result of pipe-laying stresses or other imposed static loads cannot,

therefore, be ruled out at the present time. Should this occur, fisheyes may

constitute a signficant flaw in weld metal of lower toughness than that

studied in this program.

The probability of arc burns contributing to fracture is related to the

probability of crack initiation in the hard martensitic region produced by a

burn and to the probability that such a crack is of critical size for fracture

propagation in the parent material. As a first approximation, the carbon

equivalent (CE) of the parent material might be taken as a measure of the

former probability, because this parameter is controlled in the selection of

pipeline material to limit the martensite content and hardenabi 1 i ty in the

7 74
fp:
heat-affected zone of the weldment. A typical formula for pipeline construction

might be

CE = C + Mn/6 + (Cr + Mo + V)/5 + Ni + Cu/15 (93)

yielding a CE value of 0.414 for the API 5LX-65 material used for the arc-burn

studies. A typical maximum allowable CE value for 16-mm(0. 63-in) thick

pipe would be 0.40; consequently, it is concluded that the arc-burn

significance tests were conducted under a conservative criterion from a

hardenabil i ty point of view. Despite these conditions, no cracking was

observed in any of the arc burns, either before or after testing.

Even assuming cracking was present, the literature suggests that

propagation would be unlikely owing to the high toughness of current pipe

81 materials. Fearnehough and Jones have noted that a number of tests

conducted in the United States and England have demonstrated that defect

failure in modern pipeline materials is primarily controlled by plastic

collapse phenomena and that the critical defect size is, therefore, associated

with tensile properties. Results of the study to develop appropriate

accept/reject curves for current pipeline materials (see Section 2) support

this view in concluding that a model based on ligament yielding is most

appropriate. These considerations suggest that the controlling parameter for

partial wall defects is flow stress on the remaining pipe wall. Lumb and

82 Fearnehough have analyzed a large number of accidental pipeline failures and

data from experimental burst tests by this criteria, noting that in no case

was failure was found to initiate from flaws having depths less than 20% of

the wall thickness for applied stresses equal to the specified minimum yield

strength of the flawed material. This corresponds to a tolerable flaw depth

of about 3 mm (0.12 in) for pipe of the thickness studied in the present work.

The experimentally determined relationship between arc-burn width and depth

175

CD

C

to QJ • E

to to CL *o O

r— to OJ <

3d QJ

X QJ >

to

X QJ ^ . to •a 0 •T— 2; QJ 4- E 0 - i- r— to 0 t—

s- f— 4- to 0

to E 3 QJ to QJ C >D 3

to QJ QJ CD E x X •r— to X X C •r— to

“T— 4-X 4-

to

to

X 1- X! X O c 4-

•r— OJ to JC x

X

QJ

X 0 C

S- •f— c •r— O O

CL Cl O OJ X

i_ s- S --

C

O 3 O

•rX c ‘r— E O t—

X • to to O

-M s- •r— to

10 4— • X

QJ •f—

O

O OJ 3D E -O c QJ •r— to X X X

to c O , s •r— • r— to c

to $- 4— X to -0

QJ to •r- to QJ X T3 c X X *D 0 4-> OJ CD

OJ >

toC 0 X 3 qj •r— QJ (O QJ to to u 2 X QJ c x 3 «£ 3d

•r" OJ

O QJ XI X X

cd OJ (O to

to >> S- •r— L- 3 QJ 4-

to to 4->

U

-O t_> (O -O

X c to 4— QJ

to 3 s- S_ •r- O 4- 3 S- cd X) to 3 c QJ X

r— OJ r— c

“XX • 1— QJ s- to +-> X E QJ c O -r- CD

•r— to 3 U s- CD QJ *o QJ to

•r- c CL

S- •1

QJ to

O 4-

X

X

QJ

X QJ

u

to T5

X 3 "3 X E 0 C

to

U (O CD QJ

s- v.

C

X 0 c •t— X c 0

QJ •1— r— (O to =D to s_ to QJ 3 X aj X 0 QJ jto S- O c QJ

•r—

C QJ

X

u_ •r— Cl S t tO

to

CD

176

(see Fig. 69) shows that by this criterion a 12-mm (0.5-in) wide arc burn

would be acceptable, even if it were treated as a crack of equal dimension.

Such a large arc burn is unlikely to occur in practice, substantiating the

conclusion about the innocuous nature of this type of flaw in pipe materials

that fail by plastic collapse.

Figure 67 combines the results of the present low-cycle, straincontrolled fatigue program on pipeline girth welds tested (reinforcement

removed) with the results of a similar low-cycle fatigue program conducted by

the British Welding Institute (BWI) in their assessment of the significance of

74 75 porosity and slag on the performance of pressure vessel weldments. *

Specimen sizes, flaw contents, and parent-metal yield strengths were very

similar in both programs. The Welding Institute study was conducted at 25°C

over a total strain range of approximately 0.0009 (below specimen yield

strength) to 0.0045 (approximately twice specimen yield strength), with cyclic

strain applied between a preset maximum and zero. In contrast, the present

study investigated flaw significance under higher total strains of

approximately 0.0044 and 0.0090, produced by fully reversed loading at -2°C.

Therefore, the present work extends the Welding Institute data to higher

strain levels and to lower fatigue life. The Welding Institute data reflect

cyclic life to a 60% load drop, whereas the present data reflects a 40% drop;

however, this difference in failure criteria is insignificant in view of the

high crack propagation rates at the higher strain levels.

As noted in Fig. 67, co-plotting of the data suggests a change of slope

in the log Ae^-vs.-log N

c plot in the vicinity of 10 3 cycles where the total

strain range has reached about 0.007. This is consistent with the observation

that the fatigue failure mode changed from propagation of a single fatigue

crack front through the flawed welds at the lower strain range (±e )

177

TOTAL

STRAIN

RANGE,

Act

Fig. 67 - Combined results of the NBS low-cycle fatigue study and results

published by the BWI in a similar study. 74,75 Change in slope of

the number of cyclic reversals to develop cracks in 40 to 60% of

the cross-sectional area, N , as a function of total strain range

is attributed to a change in the failure mode at high strain levels.

All data are from welds with the reinforcement removed.

178

Table 11 - Ultimate Stress for Welds Pulled to Failure

in Tension at -3°C

Ultimate Stress^ Radiographic

Condition MPa psi x nr Obscuration

Sound Automatic 651 94.4 None

648 94.0

Automatic Porosity 665 96.4 *15%

651 94.4

Manual Porosity 675 98.0 *12%

681 98.8

179

to multiple initiation and coalescence of cracks, from individual flaws at the

higher strain range (±2e^). Newman and Gurney have also noted that multiple

initiation of fatigue cracks is usually observed at high stresses, leading to

75

rapid crack coalescence. Thus, the conclusion of Harrison et al . that

extrapolatinginformation on flaw significance established under high cycle

conditions to endurances of 10 2 cycles on a strain basis may be incorrect for

highly flawed welds; a valid extrapolation may be confined to endurances above

10 3 cycles. It also suggests that the constant in Eq. 92 determined from the

Ae^.-vs.-N

c curve above 10 3 cycles may not be a valid indication of failure

ductility on the first loading cycle for flawed welds.

4.6 Summary

The probability of fracture initiation from buried slag and porosity or

from arc burns has been assessed in manually and automatically welded pipeline

girth welds. Highly flawed welds and sound welds were subjected to fully

reversed, strain-controlled, low-cycle fatigue at strain levels just above

weldment yield and at twice yield. The primary criterion for flaw sensitivity

was a comparison of the number of cyclic reversals for crack initiation in the

flawed welds and in the sound welds. Additional information on the

contribution of the flaws to crack initiation was obtained from fractographic

analysis of tests conducted with the weld reinforcement removed.

Porosity and slag had no discernable effect on the cyclic life to

initiation at the lower strain level. There was evidence of some reduction in

cyclic life in the flawed welds, in particular welds containing slag, at the

higher strain level. But, even under this most severe test condition, a

minimum of 30 cyclic reversals was required to initiate cracking. Arc burns

played no part in crack initiation in any of the weldments, regardless of

test condition. A comparison between the performance of flawed welds, sound

180

welds, and unwelded pipe material showed that the geometrical discontinuity

formed by the weld reinforcement dominated the crack initiation process,

regardless of flaw type or content.

Fractographic examination showed no initiation from individual pores at

the lower strain level, but such initiation was observed at the higher strain

level. Fisheye cracks surrounding micropores and slag inclusions were

observed on the fracture surface of all manual welds tested, indicating

localized segregation of hydrogen to such flaws during welding with the highhydrogen cellulosic electrodes. Such flaws appeared to have a minimal effect

on cyclic life to crack initiation, unless they penetrated the surface.

Because the strain levels were above yield and because the flaw levels

were equal to or in excess of those anticipated under worst-case field

conditions, these data indicate that the probability of crack initiation from

such flaws in statically stressed pipelines containing realistic flaw contents

is vanishingly remote. Note that the validity of this conclusion has been

demonstrated at only one level of weldment toughness and hardenabil i ty

Particularly in the case of arc burns, it cannot be assumed that the

conclusions reached in this study can safety be extrapolated to weldments of

lower toughness.

181
.
5. INTRINSIC LIMITATIONS ON DIMENSIONS OF BLUNT FLAWS

M. B. Kasen, G. E. Hicho, and R. C. Placious

5.1 Introduction

Flaws of a three-dimensional, blunt configuration, such as porosity, slag

and arc burns, presently account for the majority of remedial welding under

workmanship criteria. This partially reflects the sensitivity of radiography

to detection of such flaws. Section 4 of this report considered the more

general question of the significance of such flaws, assessing their influence

on the mechanical stability of weldments under fatigue conditions. We here

consider the question of how blunt flaws might be measured, assuming they are

considered significant in a fracture mechanics analysis.

Poor sensitivity and diffuse reflection from the flaw periphery limit the

detection and measurement of the through-wall depth of porosity and slag by

ultrasonic means. Radiography is able to detect such flaws, but gives direct

information only on their projected dimensions. Through-wall flaw depth can

be estimated by scaling the radiographic density of the flaw to a known

density difference--for example, to the density difference caused by a

radiographic penetrameter shim of known thickness or to density differences

produced by a step wedge exposed and developed under the same conditions as

the radiograph of interest. Such an approach was used in applying fracture

mechanics principles to assessing the need for remedial welding of flawed

4 girth welds in the Trans-Alaska Oil Pipeline. However, intrinsic and human

variables plus uncertainties about conditions under which the field radiographs were made introduced large uncertainties into the measurements. This

led to extreme conservatism in assessing blunt flaw significance. Subsequent

84 analysis of the procedures used has shown that some sources of inaccuracy

could be ameliorated by improving control over the radiographic inspection

182

process, and procedures using advanced state-of-the-art electronic techniques

36

could be used to reduce the subjectivity of such measurements.

Unfortunately, such improvements would increase the cost of pipeline

inspection while providing relatively modest gains in measurement accuracy.

It was therefore of interest to consider more cost-effective ways of

assessing blunt-flaw through-wall dimensions. One such approach is based on

the rationale that the depth dimension of porosity and slag should be related

to the radiographically projected dimensions and that the maximum flaw depth

should be intrinsically limited to the depth of the weld pass in which the

4

flaws occur. The Trans-Alaska pipeline work had shown that the depth of arc

burns could be correlated with burn widths measured from field radiographs.

The present work sought to confirm the validity of the principle for porosity

and slag.

5.2 Experimental Procedures

Pipeline segments containing porosity and slag were cut out

circumferentially adjacent to the weld reinforcement. The segments were taken

from the welds fabricated as detailed in Table 5 and were selected to provide

examples of isolated flaws. Samples containing the flaws were then

radiographed normal to the weld and in the plane of the pipe, so as to reveal

both projected flaw length (or width) and flaw depth (through-wall flaw

dimension). Slag in the form of "wagon tracks" in the manual welds required

sectioning of the welds down the centerline and studying each half

independently. The samples of porosity were obtained from automatic welds.

Several methods were investigated for measuring the flaws; the method finally

selected was an image-enhanced television projection of the flaw images at

enlargements up to four diameters. Correlation of flaw dimensions with that

of the weld-pass depth was established by measuring pass depth directly from

183

polished and etched weld cross sections. It was recognized that the true flaw

size is somewhat larger than that revealed by the projected dimensions;

however, the latter was used because it is the criteria used in all flaw

measuring from radiographs, as required for a fracture mechanics analysis.

Thirty-nine specimens containing slag and ten containing porosity were

studied.

5.3 Resul ts

The experimentally determined relationship between the through-wall

depth, D.p, of porosity and slag and the thickness of the weld pass, H, in

which the flaws occur is shown in Fig. 68. Similar data illustrating the

relationship between the radiographically projected dimensions are illustrated

in Fig. 69.

Both techniques confirmed that the through-wall depth of porosity and

slag was essentially confined to a dimension equal to or less than that of the

depth of the weld pass in which they occur. Flaw depth was also confined to a

dimension less than that of the average weld-pass thickness measured between

the plate surfaces. The widths of slag stringers, W

f

, were always less than

slag depth, whereas the length of the pores, L^, was equal to or larger than

their depth. The projected length and depth of pores appeared to be linearly

related for very small pores, but this relationship broke down as the pore

size approached that of the weld-pass depth. Slag depth could not be

predicted from slag width.

Fractures of the porous weld illustrated in Fig. 70 provides additional

evidence that porosity in an underlying pass is not likely to be extended into

a successive sound pass during multipass welding. Here, the succeeding pass

has cut into the porous region of the underlying pass, but the quality of the

successive pass was not affected. The difficulty experienced in producing

184
'
(a)

(b)

FLAW DEPTH, Df ,mm

FLAW DEPTH, Df ,in

FLAW DEPTH, Df ,mm

Fig. 68 - Experimentally determined relationship between the

through-wall depth of porosity and slag and the

thickness of the weld pass in which the flaw occurred.

18 5

WELD

PASS

THICKNESS,

H,

in

U!‘*M ‘H10IM MVld

uuui^M ‘H1QIM MV“ld

E

E

a

i

i-

^ CL jq LU —

Q

I

i OJ3 xj

Q. XJ O c -t-> •r— QJ s_ fn s- SZ +-> .c •c ro to O +J *— o a. c QJ ro r— o ••“J QJ —

-

+-> •r— o JZ fO • +J s_ +J C •r— SI

<TJ Cl •r— > -t-> • r— XJ OJ Cl

QJ >> c to -o QJ

s- fO c XJ • r— o XJ 4- -D fO to •r- &_ to dJ

O) o 2 Cn fO to oc c •f— fO QJ XJ H3

•r— .c n— S- c CL E

E CL 4- (TJ | o

s_ fC *o -M X5 sO) s- 4-) QJ to r— 4-

4-> OJ C sz QJ

o> o 3 u OJ 2 XJ

xj •r- r— +-> c QJ XJ -O fD o c U

>» n3 zn ro 3

r— S- 4- -4-> QJ XJ

r— O cz E O

03 <D • QJ s-

+-> C to .XZ to OJ CL C +-> c +-> OJ sz QJ

oj o CL s_ s- E c •r— QJ CL

•r— QJ to XJ OJ XJ to

S- dJ C s- c • r—

QJ cu r— 3

CL +-> _E r— , , O . X OJ fO -Q s_ O

LU JD XJ 2 — <T3 *

CTi LD

CO

U!‘h ‘H10N31 MV1J

ww‘h ‘H10N31 MVld

u !

f « 3 i 3 WVIQ Nuna 3 dV m ^ <n CNI — o o o o o o

T I I I I Im

186

FLAW

DEPTH.

D

f

,

mm

CD

cd CL

CD o

re

X Cl

0} O) *o X X • o> X X X

3- o re 4— X CO r= 4- re 4- O , CD X re CL 0

•i— iO f— o • « X u CL X c CD r— o o • X O) CO

IT)X CD CD re -

re •n o re CL c X 3

T3U i~

C QJ

X • • 3 4- c o 4- 0) o

CD QJ

X •J— t— X

re QJ re x o

4- t— • •T— O 1—

4-

•r— c • re c o CD -—-- CD

•T— CD re X re c s:

re CL •?

s- x CD 3 • O) c o «

c •r- J- X

O) "O s- CL QJ re QJ i 3

0)

a QJ

X

u QJ •r—

X 3 o >

CD re CD 4- "O c o> s- QJ

•r—

X OJ •r—

X 4->

X 4-

re c •r*

s- o r—

C

XX CD

CD c CD re 3 •r— CD E

re r— 5 Cl

•r— re T^— r— QJ X

T3 4-

X CD f— -t-> •i

QJ re

X % X CD re QJ O re S_

•r- >> o

+-> >>

x E

re

X •r— E *r— CD QJ O CD o X 4-> O i- X 3 s- O

re o Cl c

CL •r— ID — 3 4- CD E O O C 5

&- o o o s- X X Cl QJ CD CD 4- c

re CD •1— QJ C s- i- 4- re s_ re o s- o> x X QJ

0)

r

u

SX re 3 3 OJ 4-

4-> o X s- O X -t-> QJ

re

XX

i- r— 4- c

Ll_ 5 O •I— I o

cd

187

high levels of buried porosity during manual welding (see Section 4) suggests

that this assumption is generally valid.

Results of prior studies on the inherent depth limiations of arc burns

4 are illustrated on Fig. 69c. In practice, the arc-burn depth is taken as the

intercept between the burn diameter (or maximum width) on the ordinate of this

figure and the lower bound of the scatter band depicting total burn depth.

Arc burn length is measured directly from the radiograph.

5.4 Summary

It has been shown by direct measurement on automatic and manual welds

that the maximum through-wall depth of slag and porosity is intrinsically

limited to a maximum dimension equal to the depth of the weld pass in which

such flaws occur. It was found that the depth of porosity or slag could not

be predicted from the projected dimensions of such flaws on a radiograph.

These results confirm that characterizing the average weld-pass thickness

associated with a specified welding procedure can provide an upper limit to

the through-wall dimensions of porosity and slag, should such information be

desired in a fracture mechanics analysis. It is probable that a more precise

measurement would rarely be justified in practice, because it is probable that

most flaws having depths equal to that of the average typical pipeline

girth-weld-pass thickness would likely be acceptable.

These present results complement previous studies demonstrating an

inherent limitation to the through-wall depth of arc burns based on burn

width.

188
I
6. PRACTICAL LIMITATIONS ON BLUN1 FLAW CONTENT

USING RADIOGRAPHIC INSPECTION

M. B. Kasen

Although the studies described in Section 4 of this report indicate that

blunt flaws have a low probability of initiating brittle fracture,

their presence may interfere with detection of significant, sharp flaws during

the inspection process. This is much less of a concern when ultrasonic

techniques are used, because the proximity of blunt and sharp flaws should

enhance rather than diminish the flaw signal. But when radiography is used as

the inspection method, the high radiographic density produced by blunt flaws

can mask the presence of sharp flaws unless limits are placed on the

permissable size and distribution of blunt flaws.

Slag does not present a problem in this context because of its localized

nature in the weld. However, a primary concern is the possibility that

excessive scattered porosity can mask the presence of sharp flaws, such as

lack of fusion, incomplete penetration, or cracks, when radiography is used as

the primary inspection tool. Studies by the British Welding Institute^ have

suggested that obscuration must be limited to less than 10% of the projected

radiographic area to avoid this possibility. The British Standards

Institution has, therefore, recommended limiting the permitted obscuration to

less than 5% for ferritic steel weldments having a Charpy V-notch energy

absorption of not less than 40 J (29 ft-lbf) at the minimum service

temperature. For materials of lower toughness, porosity should be assessed as

o

a planar flaw on the basis of flaw interaction. The Working Group on the

Significance of Defects under Commission X of the International Institute of

Welding has reached a similar conclusion in proposing acceptance levels for

189

73

defects that may fail by brittle fracture. This group recommended accepting

porosity contents up to 5% of the projected area of the radiograph for welds

in ferritic steel having yield strengths up to 482 MPa (70 x 10 psi),

provided that the weld metal has a minimum Charpy energy absorption of 41 J

(30 ft-lbf) at the minimum service temperature. In other cases of lower

toughness, 3 % obscuration would be acceptable. These would appear to be

reasonable figures, because welds containing less than this level of porosity

are easily obtained by good practice, and the probability of 5% obscuration

masking sharp flaws in tough weld metal is very low. Since Section 4 has

shown that there is a negligible effect on crack initiation due to a 13 to 15%

obscuration by porosity, a 3 to 5% level would also appear to be conservative

from a performance point of view.

Several procedures have been suggested for correlating the percentage of

radiographic obscuration due to porosity in welds with porosity content by

7 1 85

volume. * However, such methods are, for the most part, tedious and timeconsuming and provide more detailed information than is required for field

interpretation of weld quality. If a 5% maximum obscuration level is chosen,

a reasonable means of distinguishing between the appearance of a weld

containing porosity at a 3% and at a 5% obscuration level should be

sufficient. One possibility is to compare the radiograph with a series of

sketches illustrating the appearance of porosity distributions covering the

range of interest. Such a series, based on similar sketches currently used in

API 1104 and in other weld-quality standards, is shown in Fig. 71. These

illustrations suggest that, if an acceptance level of 4% were specified, an

interpreter might err one percentage point higher to a 5% level, but would be

much less likely to err 2 % to a 6% obscuration level. This is because the

human eye is much more capable of judging small differences in obscuration at

low obscuration levels than at high levels.

190

The importance of the dot distribution size in the reference standard is

illustrated in Fig. 71. Here, it appears that the percentage obscuration

increases from top to bottom, whereas in fact each sketch has the same five

percent obscuration. This difference in visual perception is caused by a

progressively finer dot distribution from top to bottom among the sketches,

illustrating that reference sketches containing more simulated large pores

than would be likely in the radiographs would give the interpreter a

conservative bias.

191

Simulated Porosity

Obstruction

• • .

. •

• •••«••• • • • . • • • • • • • • . • • • • • • •

•’ •*: • • • • •

•• • i . •

•

• •

• • • • •••»

• • • • •

• • •• . • :

* •

•• .* * • • •• ••••••!

Fig. 71 - Possible development of a radiographic reference standard for establishing

limits to permissible obscuration due to scattered porosity. Series (a)

illustrates the relative ease of distinguishing percentage differences in

obscuration in the 3 to 1 % range. Series (b) illustrates that the eye

is biased toward a higher degree of obscuration as the size of the scattered

porosity decreases. Each of the sketches in (b) contains 5% obscuration.

192

7. CONCLUSIONS

Studies have been conducted to establish f itness-for-service criteria for

large-diameter pipeline girth welds.

1. A fracture mechanics analysis model that relates allowable flaw

sizes to applied stress level and fracture toughness has been developed,

verified experimentally, and used to calculate allowable flaw-size curves for

the proposed ANGTS pipeline. The model is based on a yielded-1 igament

principle and incorporates a trhough-thickness plasticity correction.

Experiments on surface cracks in tensile panels, welded pipe segments, and

large-diameter (900 mm, 36-in) pipes confirmed the validity of this model.

The applicability of the model is limited to pipelines with maximum stresses

in the longitudinal direction below the specified minimum yield strength of

the pipe. It is further limited to pipelines where the yield strengths of the

weld and base metal are similar. The degree of similarity in yield strengths

has not been established, but the desirability of higher yield strength in the

weld metal has been clearly demonstrated. The use of allowable flaw-size

curves can be considered for either general pipeline use or for site-specific

engineering assessment. Allowable flaw-size curves generated using typical

pipeline operating conditions and weldment properties are more conservative

for long flaw lengths than those derived using the procedures of Appendix A of

the 16th edition of API 1104.^

2. Radiography has serious deficiencies as an inspection tool in a

f itness-for-service analysis. Experience and laboratory studies have shown

that it is relatively insensitive to significant sharp flaws in a fracture

mechanics evaluation. Through-wall flaw dimensioning from radiographs has

also been found to be inaccurate. Ultrasonic methods are sensitive to sharp

193

flaws and are more adaptable to measuring flaws. However, conventional

systems are too slow and complex for routine pipeline inspections. In this

study, a new ultrasonic inspection system was developed for this task. The

system utilizes electromagnetic transducers (EMATs) to generate

long-wavelength shear (SH) waves. Unlike conventional techniques, this

approach does not require contact between the ultrasonic transducer and the

pipeline. The long-wavelength, low-frequency signal simplifies flaw detection

because the signal scales monotonically with increasing through-wall dimension

in the range of interest. Specifically, it has been shown that vertically

oriented 0.75-mn(0. 0295-in) deep and 25-mm (0.984-in) long surface flaws and

2-mm (0.782-in) deep and 25-mm (0.984-in) long interior flaws can be reliably

detected and measured. Detectability limits for shorter flaws have been

determined and related to representative workmanship and alternative acceptance

standards. Furthermore, it has been demonstrated that the flaw signal is

relatively unaffected by differences in type, orientation, and roughness of

the flaws. The use of SH waves enhances sensitivity to sharp flaws oriented

perpendicular to the surface, while providing relatively low sensitivity to

blunt flaws. This facilitates detectability of the most significant flaws in

a f itness-for-purpose analysis. The system is much less sensitive to the

presence of weld reinforcements than are conventional systems. It is adaptable

to full or partial automation, and it increases the reliability of flaw

detection and sizing because it depends less on operator skill. At its

present state of development, the SH-wave EMAT inspection system has the

following deficiencies: 1) performance has not been demonstrated with real

flaws under field conditions, 2) field proven hardware remains to be

developed, 3) personnel trained in the system are not readily available.

194

4) equipment calibration standards remain to be developed, and 5) sensitivity

of the system to detecting tight cracks closed by compression stresses has not

been determined.

3. The presence of large quantities of buried porosity and slag, or of

arc burns in manual welds has been shown to have a negligible effect on the

low-cycle fatigue life of the pipeline girth welds evaluated in this study. A

similar result has been shown for porosity in automatic welds. Because the

applied strains were well in excess of yield strengths, and the flaw levels

were in excess of those anticipated under worst-case field conditions, these

results indicate that the probability of crack initiation from such flaws

would be negligible if such welds were subjected to the essentially static

loading conditions of operating pipelines. These results suggest that such

blunt flaws may be considered innocuous as fracture initiation sites in girth

welds fabricated with materials of the toughness studied in this program. The

validity of this conclusion has not been experimentally demonstrated for welds

of lower toughness; however, the conclusion is consistent with that of others

who have addressed the problem of blunt flaw significance in weldment failure.

It is emphasized that these conclusions pertain only to buried porosity end

slag and to arc burns not associated with cracks. Porosity breaking the

surface or arc burns containing cracks should be treated as equivalent-sized

sharp flaws.

For welds having substantially lower toughness than that evaluated in

this program, blunt flaws may be conservatively treated as sharp flaws of

equal dimensions. This work has shown that the maximum through-wall dimension

of porosity and slag is intrinsically limited by the depth of the weld pass in

which they occur or to flaw-width criteria in the case of arc burns. The

195

upper limit to flaw depth provided by these criteria can therefore be conservatively used in a fracture mechanics analysis, eliminating the need for measuring

the through-wall depth of flaws in the field.

Low-cycle fatigue caused fisheye cracks to form adjacent to micropores

and slag in all welds made with E8010G cellulosic electrodes, even though

testing was performed U years after welding. This indicated a stable,

residual hydrogen content in the weld metal. Available data indicate that

such flaws can also occur if static loads approach or exceed weld-metal yield,

suggesting the possibility that such flaws could develop during the pipe-laying

operation subsequent to the inspection process. Except when they penetrated

the surface, such cracks did not appear to affect the fatigue strength significantly, suggesting that they are basically innocuous flaws in weldments of

the toughness tested. However, it cannot be assumed that cracking from this

source will remain innocuous in weld metal of lower toughness. Conditions

under which fisheye cracks develop in high-hydrogen electrode pipeline welds

deserve future study.

196

8. RECOMMENDATIONS

8.1 Implementation of Inspection Methods

The practical implementation of a f itness-for-service inspection system

for a pipeline involves many factors outside the scope of this report.

Nevertheless, it is instructive to consider the technical factors that might

influence the manner in which the inspection system is integrated into an

overall quality control program. Three possible options, based solely on

perceived technical merit, are listed in Table 12.

8.1.1 Option I

The first option assumes that all of the welds will be inspected by an

automated ultrasonic system to provide the basic flaw detection and measuring

capability required for the f i tness-for-service analysis. It is also

envisioned that some fraction of each welder's production be inspected by

conventional radiography to provide control over workmanship, serviceability

of welding equipment, and quality of consumables. Since welds inspected by

radiography would also be inspected by ultrasonic means, radiography would not

be used to establish an accept/reject criterion for weld quality. This option

is technically preferable because it differentiates between the purpose of

each inspection technique and uses each in the area of its greatest strength.

Documentation provided by Section 3 of this report suggests that an

ultrasonic system based on low-frequency SH waves generated by electromagnetic

197

transducers (EMATs) would provide a promising option for flaw detection and

measuring with the best adaptability to automation. Such a system will not

respond to the presence of blunt flaws unless they are very large or locally

concentrated. This is a distinct advantage because, as documented in Section 4

of this report, blunt flaws may be considered intrinsically innocuous, from a

fitness-for-service point of view, if minimum toughness levels are maintained.

There is no possibility of sharp flaws being obscured by the presence of blunt

flaws, because proximity of the two flaw types will enhance the ultrasonic

signal

This option assumes that blunt flaws not detected by the ultrasonic

inspection method can be neglected, based on the work described in Section 4.

Strictly speaking, this assumption has only been validated for weld metal of

the toughness studied in this program; however, the conservatism built into

the present investigation, along with the conclusions reached by others who

have studied the problem, suggests that the assumption would probably be valid

for any reasonable minimum toughness specified for pipeline construction.

Additional conservatism arises from the realization that slag, which this work

has shown to be the most likely of the blunt flaws to initiate fracture, is

relatively easily detected by ultrasonic means due to its solid nature.

8.1.2 Option II

Option II considers the case where Option I is premature for existing

technology. Here, radiography is assumed to be the main inspection system,

and ultrasonic methods are used selectively to measure the through-wall depths

of sharp flaws that are detected by radiography. Flaw length would be

measured directly from the radiographs. This option is less technically

desirable because of the relatively poor sensitivity of radiography to sharp

198

flaws, which are the most significant from a f i tness-for-service point of

view. Again, it would appear that development of a portable, manual variant

of the ultrasonic system described in Section 3 of this report would be very

useful in this option.

Some restriction must be placed on the allowable content of blunt flaws

when radiography is used as the primary inspection tool. Section 6 has

discussed the limitations imposed on porosity content due to obscuration

concerns. The extent to which large individual pores reduce the weld crosssectional area and, therefore, reduce the weldment tensile strength must also

be considered in some circumstances. This is unlikely to be a problem in

girth welds made in heavy wall line pipe where the average weld pass thickness

(and therefore the maximum through-wall dimension of a pore) is a small

percentage of total wall thickness. But. it could become a problem where

thinner wall pipe is used or where welds are fabricated with automatic

processes producing deep, narrow weld passes. Such concerns can be avoided by

limiting the allowable dimensions of individual pores to that of a permissible

sharp flaw based on a fracture mechanics accept/reject criteria, assuming pore

depth to be equivalent to that of the weld pass.

In addition to limits on the quantity and size of allowable porosity,

consideration must also be given to its distribution. A localized cluster of

porosity exceeding the obscuration limit could conservatively be treated as a

single crack of length equal to the maximum cluster dimension and having a

depth equal to that of the average weld pass. Since a linear distribution of

86

porosity is frequently found to be associated with incomplete penetration,

it would be prudent to treat closely spaced groups as an equivalent-sized

single crack.

199

In view of the tolerance of tough weld metal to relatively large sharp

flaws (see Section 2), it is unlikely that applying a fracture mechanics

assessment to the significance of porosity will result in appreciable remedial

welding.

Where it is desired to do so, limitations on the allowable length of

entrapped slag may also be conservatively established by considering the slag

inclusion to be a sharp flaw of the length indicated on the radiograph to have

a through-wall dimension equal to that of the average weld pass depth.

As with porosity, this procedure is not expected to contribute substantially to the amount of remedial welding required where weld metal of

appropriate toughness is used. Should it be desirable to place ajimitation

on acceptable slag width, the present work suggests that a rational limitation

would be the same as for slag depth. Although Fig. 68b (Section 5) shows that

it is not possible to determine slag depth from projected width, it indicates

that the maximum width will not exceed that of the weld pass depth.

Excessively wide slag width may, therefore, be taken as an indication of

an excessively deep weld pass. Note that applying this approach to the manual

girth welds studied in the present work would result in a practical limitation

of 3 mm on slag width and depth--the same as has been proposed as upper limits

on these parameters under the brittle fracture criteria in the British

87

Standard Institution Document PD 6493:1980.

Where it is deemed necessary to analyze the significance of arc burns on

a f itness-for-service basis, they may also be treated as equivalent-sized

sharp flaws having lengths equal to that of the burns and having through-wall

depths scaled from the maximum burn widths by use of Fig. 68c (Section 5). A

precedent for this approach was established by the Trans-Alaska Pipeline

(TAPS) work.®®

200

These suggested limitations on the content of blunt flaws when inspection

is conducted by radiography are summarized in Table 13.

8.1.3 Option III

This option assumes that appropriate ultrasonic methods for sizing planar

flaws are not available and that all inspection criteria must depend on

information obtained by radiographic techniques. This option is least

technically desirable, since sharp flaws are not easily detected and

sharp-flaw depth cannot be accurately measured using radiography. Unlike

blunt flaws, one cannot, in general, assume an inherent limitation on the

through-wall dimension of sharp flaws. The only alternative is to limit sharp

flaw length by the criteria of existing workmanship standards.

On the other hand, the suggested treatment of blunt flaws presented under

Option II and detailed in Table 13 are equally relevant to this option. Since

the majority of current remedial welding is done to remove blunt flaws having

sizes in excess of that permitted by workmanship codes, this proposed

alternative criteria for acceptability of blunt flaws could substantially

reduce pipeline construction costs by providing a more rational approach to

deciding when remedial welding is justified to remove such flaws.

201

8.2 Implementation of Fitness-for-Service Criteria

A fracture mechanics approach to assessing the significance of weld fl^ws

to pipeline integrity can be applied in several ways, depending on the

objective of the user. Its use in determining the technical basis for

requiring remedial welding in the event of code violations revealed during a

postconstruction audit has already been demonstrated during the TAPS program.

However, the full potential of the approach requires applying the principles

before and during pipeline construction to reduce the need for remedial

welding consistent with pipeline structural integrity and safety and, hence,

to reduce construction cost.

The latter philosophy can be implemented in two basic ways: One is to

establish accept/reject criteria for weld flaws before the construction

begins. This would require defining the maximum applied and residual stresses

and strains for designated portions of a line where such judgments can be

effectively made and combining that information with a characterization of the

minimum fracture toughness of the weld metal, as deposited by the consumables

and processes selected for the line. Allowable flaw-size curves based on such

data could then be routinely applied during construction. Procedures assuring

maintenance of minimum weld metal fracture toughness during construction would

probably be required. Consideration might be given to generating a series of

allowable flaw-size curves reflecting differences in weld-metal toughness,

which would permit welds of lower than expected toughness to be accepted,

provided that the increased restrictions on flaw size associated with the

lower toughness were not exceeded. Advantages of this generalized approach

are overall economy and simplicity of field implementation. The criteria

established by this approach may be overly conservative for some portions of

the line; however, this can be minimized by proper selection of line segnents.

202

Alternatively, a specific-site approach may be selected, either in

conjunction with the generalized approach or independently applied. This

involves an engineering critical assessment of welds at highly stressed

locations, locations where the welds are subjected to unusually low

temperatures or where other environmental factors must be taken into

consideration. Since toughness, stress, and allowable flaw size are related

by the fracture mechanics model, allowable flaw-size curves can be constructed

to address any desired set of conditions. It would be necessary to generate

appropriate accept/reject criteria reflecting the specific prevailing

conditions. Such an approach would probably result in requests for variances

from accepted code requirements. Therefore, at the start, it might prove

useful to generate a family of allowable flaw-size curves covering the

variables anticipated during construction. Availability of such a family

might significantly facilitate implementing either the generalized or

specific-site approach.

When qualifying weld metal to a fracture toughness criteria, it should be

recognized that weld metal deposited with high-hydrogen, cellulosic electrodes

will initially have a low toughness due to the high-hydrogen content in the

weld deposit. Toughness will increase with aging time after welding, because

89

the hydrogen content is lowered by diffusion. Therefore, qualifying

conditions should include a delay time subsequent to welding, selected to

represent the toughness at the time of maximum weldment stress in the line.

Since residual hydrogen content is affected by a number of factors, it may be

necessary to characterize the change in toughness with aging time for a

particular welding procedure.

203

Since residual hydrogen content is affected by a number of factors, it may be

necessary to characterize the change in toughness with aging time for a

particular welding procedure.

204

Table 12 Technical Options for Field Implementation

Fitness--for-Service Criteria

NDE Method Flaw Assessment Workmanshi

p

Option Primary Backup Sharp Blunt Criterion

I Automated None £ vs. a None 10% RT

UT (EMAT)

II Conventional Manual £ vs. a Lf vs. RT

RT UT (EMAT)

III Conventional None £ only L f vs ' D

f

RT

RT

*Assume limited to average pass thickness.

205

Table 13 - Suggested Limitations on Allowable Content of Blunt Flaws

POROSITY

Quantity

Less than 5% radiographic obscuration

Size

Equivalent single crack*

Distribution

Clusters: Over 5% obscuration treat as equivalent single crack*

Linear: Treat groups separated by less than twice wall thickness as equivalent single crack*

SLAG

Treat as equivalent single crack*

ARC BURNS

Treat as equivalent single crack*

*Through-wall depth inherently limited to average weld pass depth.

206

8.3 Future Research

8.3.1 Fracture Mechanics

Procedures for the use of elastic-plastic fracture mechanics for the

derivation of allowable flaw-size curves are in the early stages of

development. Further work is recommended to increase confidence in the

analytical results. Future research needs in the areas of analytical model

j development, model verification, and implementation are summarized below.

Analytical Model Development The present analytical model presented is a

relationship among stress, flaw size, and toughness. For applicability to

situations where the strains exceed yield, a relationship among strain, flaw

size, and toughness is needed. Further developments should be incorporated

into the strain-based model: First, the behavior of small flaws in a large

section (such as large-diameter pipe) needs to be modeled to account for

gross-section yielding. Second, the differing stress-strain relationships of

the weld and the base plate should be incorporated into the model. Third, the

influence of residual stresses on fracture at high strain levels should be

assessed.

Model Verification The model verification studies in the present program

were limited to the analysis of results of large-diameter-pipe tests conducted

as part of an independent program. Direct observations of the spread of

plasticity are needed for proper modeling of the plastic-zone circumferentially cracked pipe in bending. Further tests are needed to verify the

curvature correction proposed herein. And finally, verification of the

207

fundamental assumption of the fracture mechanics approach would greatly

contribute to confidence in the analysis, that is, verification that

large-diameter pipe does fracture at a critical value of CTOD measured in a

fracture toughness test.

Implementation . Given an allowable flaw-size curve, an implementation

strategy is needed to ensure that the assumptions used in the derivation are

adhered to during pipeline construction. Improvements in the CTOD testing

method are needed, particularly for evaluating weld-metal toughness. A

statistical sampling procedure is needed to ensure that the yield strength and

toughness of the pipeline girth welds exceed the values used to establish the

allowable flaw-size curves.

8.3.2 Inspection

Recent studies have shown that it is feasible to increase substantially

the detection sensitivities of the SH-wave EMAT system with respect to planar

flaws that are canted with respect to the through-thickness direction.

Examples of such flaws include cracks in the heat-affected zone (HAZ) and

incomplete fusion. To improve the detectability of such flaws, the present

SH-wave EMAT system would have to be augmented by an additional SH-wave EMAT

that would be sensitive to the SH^, SH^, and SH^ plate waves. It can be shown

theoretically that these three plate waves are scattered much more strongly by

90 canted planar flaws than the dominant SHq and SH^ modes. The technology for

91 constructing such transducers is now available.

Another topic that requires further quantitative investigation is the

effect of residual and applied residual stresses on the functioning of the

SH-wave EMATs. It is believed that under certain circumstances, the efficiencies and pure-mode characteristics of the periodic-permanent-magnet

208

EMATs may be significantly degraded. As a consequence, the EMAT performance

characteristics should be verified experimentally using a pipe section welded

by approved procedures and subjected to realistic membrane and residual

stresses. Current understanding of how SH-wave EMATs function on a surface

under stress is very limited.

It is believed that additional work is needed in the area of signal

analysis. For example, accurate estimation of the SH-plate wave scattering

coefficients could be obtained by nonlinear least-squares fitting of the

measured signal amplitudes as function of the distances separating the EMATs

from the weld along scan lines that are perpendicular to the pipe girth.

Then, the confidence levels in the ultrasonic measurements would be improved.

Also, better estimates of flaw shape, cant angle, position within the weld,

and other relevant characteristics would be obtained. The least-squares

fitting approach is particularly appropriate since an accurate, predictive

model for the flaw scattering amplitudes is now available.

Finally, it is believed that construction of a fieldable prototype of the

454-kHz SH-wave EMAT system is necessary. The fieldable system would be

needed to establish the influence of specific environmental conditions and

welding practices on the inherent limitations of SH-wave EMATs and suggest

possible improvements. The prototype system would be evaluated as a limited

section of pipeline that has been fully characterized using other available

nondestructive testing (NDT) procedures, including ultrasonics, radiography,

and penetrant and visual examination.

209

8.3.3 Metallurgy

Metallurgical parameters affecting the toughness of weldments in pipeline

steels deserves further study. The sensitivity of weld-metal toughness to

variations in chemical composition of the filler material, to variations in

weldihg procedure, and to the interaction of such parameters should be better

defined to increase confidence in the assumption of an existing minimum

toughness level in the f itness-for-service analysis. American Welding Society

(AWS) classifications of electrodes used in pipeline construction presently

give electrode producers wide latitude in selecting a chemistry that will

produce weld metal meeting the minimum mechanical properties of the specification. This is because, up to now, weld strength has been the dominating

consideration. For example, the AWS specification for type E8010G electrodes

commonly used in manual welding of API 5LX-65 and API 5LX-70 pipelines

establishes minimums for Mn, Si, Cr, Ni , Mo, and V, but requires that only one

of these elements meet the minimums in order to comply with the specification.

It is expected that this compositional latitude will, in itself, cause a

substantial variation in weld metal toughness; however, the extent to which

this is true is presently unknown. An equal uncertainty exists with regard to

the variation in toughness produced by a given electrode composition at

different heat input levels. These uncertainties now require that each

manufacturer's proprietary electrode be evaluated for toughness under a

specific welding procedure. This is a tedious and expensive process.

Resolving these uncertainties will permit establishment of parameters

necessary for prequalifying electrode types on the basis of toughness, thus

reducing cost while maintaining weldment reliability.

210

The present study indicates a negligible probability of fracture initiation from arc burns in contemporary pipeline materials. However, a relationship between arc-burn significance and toughness, hardenabil ity, and other

metallurgical parameters has not been established. It would be desirable to

do so to provide increased confidence in the findings of the present study.

For example, it would be valuable to determine if any circumstances exist

under which excessive deterioration of heat-affected-zone toughness could

significantly exacerbate the significance of arc burns.

The conditions under which fisheye cracks can form in weld metal deposited

by the shielded-metal-arc process should be determined. It is of particular

importance to determine the probability of their formation as a result of

pipe-laying stresses because, should this occur, significant flaws could be

introduced into the welds after the inspection process. The literature

suggests that this is a distinct possibility, because such flaws have been

most often observed during bend testing where the weld metal has been strained

beyond yield. The study should correlate the probability of fisheye formation

with the hydrogen content of the electrodes and should establish the required

postweld heating conditions required to eliminate their formation, should they

prove to be a problem in pipeline construction.

211

9. ACKNOWLEDGMENTS

This study was supported by the Materials Transportation Bureau (MIB) of

the U.S. Department of Transportation. Mr. Lloyd Ulrich, Alaska Natural Gas

Pipeline Project, (MTB) Office, Alaska Natural Gas Pipeline Projects, was

contract manager. Support for the nondestructive inspection studies was

partially supplied by the Welding Research Council, Dr. Glenn Oyler, Research

Di rector.

We particularly appreciated the cooperation of the Northwest Alaskan

Pipeline Company, Salt Lake City, Utah. The continued support of

Mr. Edward B. Clark of that company was deeply appreciated.

Many people made essential contributions to the present study. We

gratefully acknowedge the cooperation of Nova, an Alberta Corporation, in

providing data from their large-scale test program. The girth welds investigated in this study were fabricated by the H. C. Price Company, Dewey,

Oklahoma, under the direction of Mr. Richard Gwin and Mr. Steve Madden.

Artificially flawed manual welds were prepared by the Lincoln Electric

Company, Cleveland, Ohio, under the direction of Mr. Jerry E. Hinkel . Lowcycle fatigue studies were performed at the Denver facility of the Martin

Marietta Corporation by Mr. John Shepic and Mr. Richard Gibb. Radiographic

support was provided by the Rocky Flats Division, Rockwell International

Corporation, under the direction of Mr. Jack Summers. Calibration specimens

required for the ultrasonic study were prepared by Dr. Jerry E. Jones,

Colorado School of Mines, Golden, Colorado.

Many NBS people contributed significantly to this study. We particularly

acknowledge the contributions of J. David McColskey, Don Harne, Doug Chrisholm,

Mel Jadah, Mike Lauriente, Bill Carpenter, and Dr. Richie Mikesell to the

success of the experimental program. We are grateful to Ms. JoAnne Wilken for

preparation and typing.

212

10 . NOTATION

Roman Letter Symbols

crack depth

relative amplitude of plate wave

relative transduction efficiencies of a plate wave by

transmitter and receiver EMATs, respectively

normalized crack depth

22 edge crack compliance constants

normalization constant for n^ SH plate mode

gross-section area in the plane of the crack

net-section area in the plane of the crack

contour surrounding crack tip at a point along crack fror*t

Cq

2 coefficients in Kobayashi's solution for CMOD of an

elastic surface crack

blunt flaw through-wall dimension

Young's modulus

wave frequency

cutoff frequency of SH plate waves

finite width correction of Newman

force per unit length

boundary correction factor for computing stress-intensity

factor of surface cracks

closing force on the ligament of a surface crack

constant for computing stress-intensity factor of an edge

crack

G

H

J

el

J

P

K

£eff

SL

i P

L

L tr

L

f

4

M

n

N

c

P

P

4

Q

shear modulus

weld-pass thickness

O-integral

elastic part of J

plastic part of J

stress intensity factor

effective crack length (includes plastic-zone correction,

r ) y

crack length

length of crack measured to ends of the plastic zone

gage length

EMAT-to-EMAT spacing

blunt flaw length

width of weld in a fracture specimen

bending stress

moment per unit length

collapse moment of a pipe

inward-directed unit-normal to the crack face

number of fatigue cycles

applied jack pressure

applied load

power level delivered to EMAT

resultant tensile stress

square of elliptic integral of second kind

radius of plastic zone and plastic-zone correction to

crack length

plastic zone correction to crack depth

214

R

R

s R

ell

S

F

t

T

v/

-4—

T

T

xz’

u

V

9

V P

v

s

V X w

w

0

w

Ww

w

f

X

1 x,y,z

pipe radius

scattered signal amplitude ratio

scattered amplitude ratio for an elliptical flaw

surface surrounding a flaw

specimen or plate thickness

mean thickness of weld reinforcement

dynamic stress tensor

dynamic stress field components associated with SH waves

propagating along z

displacement vector

group velocity

phase velocity

bulk SH-wave velocity

particle velocity

strain energy density

half-width of the ultrasonic beam near the transmitter

EMAT

specimen width

mean width of weld reinforcement

blunt flaw width

direction normal to crack front

orthogonal coordinates

distance between flaw and EMAT

215

z

10.2 Greek Letter Symbols

o.B.e angles or angular displacements

wave propagation constant

r reflection coefficient of EMAT ultrasonic signals

6 CMOD

6 C

CTOD

A

1 remote displacement in a center-cracked panel

A edge crack compliances

A

0 1/2 CMOD

Au mode III displacement jump across the face of a crack

e L gage-length strain

AE

t total strain range in low-cycle fatigue

e NSY strain at net-section yield

Ey yield strain

e component of strain

e c edge crack rotational compliance

en angle of elliptical flow with plate surface normal

A wavelength

v Poisson's ratio

p extent of Dugdale plastic zone

a stress

a flow stress

o

c closing stress

o^y applied stress at which ligament yield of a surface crack

occurs

Oj^y net-section yield stress

216

°0 average tensile stress

°u

ultimate tensile strength

°xy °yz

shear stresses

°y yield stress

T time

0 angular position along crack front

$ elliptic integral of the second kind

U) angular frequency

10.3 Acronyms and Similar Abbreviations

ANCTS Alaska Natural Gas Transportation System

API American Petroleum Institute

AWS American Welding Society

BWI British Welding Institute

CE carbon equivalent

CMOD crack-mouth opening displacement

cmod

ly

crack-mouth opening displacement in the post-ligament-yield

range

cmod

nsy

COD

crack-mouth opening displacement at net-section yield strength

crack opening displacement

CTOD crack-tip opening displacment

ctod

e ctod

ly

ctod

rs dot

elastic component of CTOD

ligament yielding component of CTOD;

residual stress component of CTOD

U.S. Department of Transportation

ECA engineering critical assessment

217

11. REFERENCES

1. "Minimum Federal Safety Standards for Liquid Pipeline," Code of Federal

Regulations, Title 49 Transportation, Part 195, Office of the Federal

Register, National Archives and Records Service, General Services Administration (revised 1 Oct. 1975). Available like a journal in libraries.

2. "Standard for Welding Pipelines and Related Facilities," 15th & 16th

Edition, API Standard 1104, American Petroleum Institute, New York (July

1983).

3. "Consideration of Fracture Mechanics Analysis and Defect Dimension

Measurement Assessment for the Trans-Alaska Oil Pipeline Girth Welds,"

NBSIR 76-1154, National Bureau of Standards, Gaithersburg, Maryland

(Oct. 1976).

4. R. P. Reed, H. I. McHenry, and M. B. Kasen, "A Fracture Mechanics Evaluation of Flaws in Pipeline Girth Welds," Weld. Res. Counc. Bull. 245,

(Jan. 1979).

5. A. A. Butchman, "Trans-Alaska Crude Oil Pipeline" (petition for waiver of

girth weld defects). Fed. Regist. 42 119, 31512 (1977).

6. J. W. Barnum, "Trans-Alaska Crude Oil Pipeline" (decision on petition for

waiver of girth weld regulations), Fed. Regist. 41 233 52933-52949

(1976).

7. "Draft British Standard Rules for Derivation of Acceptance Limits for

Defects in Fusion Welded Joints," Document 75/77081 DC, British Standards

Institution, London (Feb. 1976).

8. "Guidance on Some Methods for the Derivation of Acceptance Levels for

Defects in Fusion Welded Joints," PD 6493 (1980), British Standards

Institution, London (1980).

9. "Specifications for Field Welding of Carbon Steel Pipelines," BS 4515

(1969), British Standards Institution, London (1969).

10. M. M. P. Carne and J. D. Harrison, "Proposals for a Fracture Mechanics

Based Criterion for the Acceptance of Defects in Pipeline Girth Welds,"

presented at the American Welding Society International Conference,

Pipeline Welding and Inspections, Houston, Texas (Sept. 21-22, 1982).

11. A. G. Glover, and R. I. Coote, "Alternative Inspection Methods and Defect

Acceptance Standards for Large Diameter Gas Pipelines," presented at the

American Welding Society International Conference, Pipeline Welding and

Inspection, Houston, Texas (Sept. 21-22, 1982).

12. "Method of Assessment for Defects in Fusion-Welded Joints with Respect to

Brittle Fracture," Japan Welding Engineering Society Standard WES

2805-1980, The Japan Welding Engineering Society, Tokyo (1980).

218

13. E. L. Von Rosenberg and C. P. Royer, "Pipeline Welding Standards,"

presented at the American Welding Society International Conference,

Pipeline Welding and Inspection, Houston, Texas (Sept. 21-22, 1982).

14. "A Critical Evaluation of Numerical Solutions to the 'Benchmark'

Surface-flaw Problem," Exp. Mech. , 20, No. 8, 253-264 (1980).

15. J. C. Newman, Jr. and I. S. Raju, "Analyses of Surface Cracks in Finite

Plates under Tension or Bending Loads," NASA Technical Paper 1518 (1979).

Available from NTIS, Springfield, Virginia.

16. J. 6. Merkle, "A Review of Some of the Existing Stress Intensity Factor

Solutions for Part- through Surface Cracks," ORNL-TM-3983, U.S. Atomic

Energy Commission (Jan. 1973).

17. J. C. Newman, Jr., "Predicting Failure of Specimens with Either Surface

Cracks or Corner Cracks at Holes," NASA TN D-8244 (1976). Available from

NTIS, Springfield, Virginia.

18. G. R. Irwin, "Plastic Zone near a Crack and Fracture Toughness,"

Proceedings Seventh Sagamore Ordance Materials Research Conference , Vol

IV., Syracuse University Press, Syracuse, New York, 63-78 (1960).

19. D. S. Dugdale, "Yieldinq of Steel Sheets Containing Slits," Mech. Phys.

Solids , 8, 100-108 (1960).

20. G. R. Irwin, "Characterization of Part-through Cracks in Tension," The

Surface Crack: Physical Problems and Computational Solutions, ed. J. L. Swedlow, ASME, New York, 1-10 (1972).

21. R. B. King, "Elastic-Plastic Analysis of Surface Flaws Using a Simplified

Line-Spring Model," Eng . Fract . Mech . , in press.

22. A. S. Kobayashi, "Crack Opening Displacement in a Surface Flawed Plate

Subjected to Tension or Plate Bending," Proceedings of the Second

International Conference on Mechanical Behavior of Materials , Ameri can Society for Metals, Metals Park, Ohio, 1073-1077 (1976).

23. A. S. Kobayashi, "Crack Opening Displacement in a Surface Flawed Plate

Subjected to Tension or Plate Bending," Document No. D- 180-19446-1 , The

Boeing Aerospace Company, Seattle (1976).

24. J. R. Rice and N. Lew, "The Part-through Surface Crack in an Elastic

Plate," J. Appl . Mech" , 39, 185-194 (1972).

25. J. R. Rice, "The Line-Spring Model for Surface Flaws," The Surface

Crack: Physical Problem and Computational Solutions, ed. J. L. Swedlow,

ASME, New York, 171-185 (1972).

26. D. M. Parks, "The Inelastic Line. Estimates of Elastic-Plastic Fracture

Mechanics Parameters for Surface Cracked Plates and Shells," ASME

80-C2/PVP-109, ASME, New York (1980).

219

27. R. R. Lockett and D. M. Parks, "The Line Spring Model for J Analysis of

Surface-Cracked Plates and Shells," presented at the Second International

Symposium on Elastic-Plastic Fracture Mechanics, Philadelphia (1981).

28. M. Shiratori and T. Miyoshi, "Evaluation of the J Integral for Surface

Cracks," presented at the Second International Symposium on Elastic-Plastic Fracture Mechanics, Philadelphia (1981).

29. D. M. Parks, "Inelastic Analysis of Surface Flaws Using the Line Spring

Model," Proceedings of the Fifth International Conference on Fracture , Pergamon, Oxford, 2589-2598 (1981).

30. R. B. King, Y. W. Cheng, and D. T. Read, "J-Integral Analysis of Surface

Flaws in Pipeline Steel Plates," presented at the Second International

Symposium on Elastic-Plastic Fracture Mechanics, Philadelphia (1981).

31. Y. W. Cheng, H. I. McHenry, and D. T. Read, "Crack Opening Displacement

of Surface Cracks in Pipeline Steel Plates," presented at the ASTM 14th

Symposium on Fracture Mechanics, Los Angeles (1981).

32. P. T. Heald, G. M. Spink, and P. J. Worthinton, "Post Yield Fracture

Mechanics," Mater. Sci. Eng. , 10, 129-136 (1972).

33. R. deWit and J. H. Smith, "Development of Some Analytical Fracture

Mechanics Models for Surface Defects in Plates of Ductile Metals,"

Continuum Modeling of Discrete Systems , University of Waterloo Press,

Ontario, Canada (1980).

34. H. Tada, P. C. Paris, and G. R. Irwin, The Stress Analysis of Cracks

Handbook , Del Research Corporation, Hellertown, Pennsylvania (1973).

35. R. Coote, Nova Corporation, Calgary, Canada', private communication.

36. S. A. Wenk, "Radiographic Video Technique for Sizing Defects in Pipe

Welds," Mater. Eva! ., 39, No. 8, 748-751 (1981).

37. P. A. Doyle and C. M. Scala, "Crack Depth Measurement by Ultrasonics,"

Ultrasonics , 16, No. 3, 164-170 (1978).

38. C. M. Fortunko and R. E. Schramm, "Ultrasonic Nondestructive Evaluation

of Butt Welds Using Electromaqnetic-Acoustic Transducers," Weld. J., 61,

No. 2, 39-46 (1982).

39. C. M. Fortunko, R. B. King, and M. Tan, "Nondestructive Evaluation of

Planar Defects in Plates Using Low-Frequency Shear Horizontal Waves," J. Appl . Phys. , 53, No. 5, 3450-3458 (1982).

40. C. M. Fortunko, "Ultrasonic Detection and Sizing of Two-Dimensional

Defects at Long Wavelengths," Appl. Phys. Lett., 38, No. 12, 980-982

(1981).

41. N. Nielsen, "P-Scan System for Ultrasonic Weld Inspection," Br. J. NDT,

23, No. 3, 63-69 (1981).

220

42. A. deSterke, "Automated Ultrasonic Inspection of Circumferential Welds in

Pipelines," Met. Constr. , 12, No. 3, 42-46 (1981).

43. W. C. Minton, "Ultrasonic Inspection of Pipeline Welds," in Proceedings

1980 Pipeline Welding and Inspection Conference , ed. D. Weir, American

Welding Society, Miami, 65-76 (1980).

44. R. F. Lumb, "Nondestructive Testing of Pipeline Girth Welds in the

1980's," Br. J. NDT , 24, No. 7, 197-207 (1982).

45. C. M. Fortunko, C. F. Vasile, and R. B. Thompson, "Electromagnetic

Transducers for Weld Inspection," presented at 37th National Fall

Conference of The American Society for Nondestructive Testing, Detroit,

Michigan (Oct. 3-6, 1977).

46. C. F. Vasile and R. B. Thompson, "Excitation of Horizontally Polarized

Elastic Waves by Electromagnetic Transducers with Periodic Magnets," J.

Appl . Phys ., 50, No. 4, 2583-2588 (1979).

47. M. deB illy, F. Cohen-Tenondji , G. Quentin, K. Lewis, and L. Adler,

"Ultrasonic Evaluation of Geometrical and Surface Parameters of Rough

Defects in Solids," J. Nondestruct. Eva!

, 4, 244-261 (1980).

48. B. A. Auld and S. Ayter, "Perturbation Method for Analyzing the Effect of

Ultrasonic Echo Returns on Rough Surfaces in Material Cracks and Voids,"

1980 IEEE Ultrasonics Symposium Proceedings., ed. B. R. McAvoy, IEEE, New

York 852-856 (1980).

49. J. M. Richardson and R. K. Elseley, "Extraction of Low-Frequency

Properties from Scattering Measurements," 1979 IEEE Ultrasonics Symposium

Proceedings , ed. B. R. McAvoy, IEEE, New York, 336-341 (1979).

50. B. Budiansky and J. R. Rice, "On the Estimation of a Crack Fracture

Parameter by Long Wavelength Scattering," J. Appl. Mech., 45, No. 6,

453-454 (1978).

51. C. D. Lundin, "The Significance of Weld Discontinuities - A Review of

Current Literature," Welding Research Council Bulletin 222, Welding

Research Council, New York (1976).

52. M. B. Kasen and R. P. Mikesell, "Interim Report on the Significance of

Blunt Flaws in Pipeline Girth Welds," Proceedings of International

Conference on Pipeline and Energy Plant Piping , ed. N. F. Eaton,

Pergamon, Toronto, 329-336 (1980).

53. R. B. Thompson, "The Relationship between Radiating Body Forces and

Equivalent Surface Stresses: Analysis and Application to EMAT Design,"

J. Nondestruct. Eva! . , 1, 79-86 (1980).

54. S. K. Datta, C. M. Fortunko, and R. B. King, "Sizing of Surface Cracks in

a Plate Using SH Waves," 1981 IEEE Ultrasonics Symposium Proceedings, ed.

B. R. McAvoy, IEEE, New York, 863-867 (1981).

221

55. S. K. Datta and A. H. Shah, "Scattering of SH Waves by Embedded

Cavities," Wave Motion , in press. ,

56. S. K. Datta, A. H. Shah, and C. M. Fortunko, "Diffraction of Medium and

Lonq Wavelenqth Horizontally Polarized Waves by Edqe Cracks," J. Appl

Phys ., 53, No. 4, 2895-2903 (1982). ^ '

57. J. Krautkramer and H. Krautkramer, Ultrasonic Testing of Materials , Second Edition, Springer-Verlag, Berlin, 338-343 (1977).

58. W. J. Pardee and R. B. Thompson, "Half-Space Radiation by EMATs," J.

Nondestruct. Eval . , 1, 150-181 (1980).

59. B. A. Auld, "General Electromechanical Reciprocity Relations Applied to

the Calculation of Elastic Wave Scattering Coefficients," Wave Motion, 1,

3-10 (1979).

60. N. N. Egorov and A. V. Kharitonov, "Diffraction of Shear Normal Modes by

a Surface Crack in an Elastic Plate," Sov. Phys. Acoust., 25, No. 1,

34-37 (1979).

61. R. B. King and C. M. Fortunko, "Extended Variational Solution for

Scattering from Flaws in Plates," J. Appl. Phys. 53, No. 5, 3459-3460

(1982).

62. R. B. Thompson, "A Model for Electromagnetic Generation of Rayleigh and

Lamb Waves," IEEE Trans. Sonics Ultrason. 20, No. 4, 340-346 (1973).

63. C. M. Fortunko and R. E. Schramm, "An Analysis of Electromagnetic-Accustic

Transducer Arrays for Nondestructive Evaluation of Thick Metal Sections

and Weldments," in Proceedings 1982 Review of Progress in Quantitative

NDE , eds. D. 0. Thompson and C. E. Chimenti, Plenum, New York (1983), in

press.

64. R. B. Thompson, S. A. Alers, and M. A. Tennison, "Application of Direct

Electromagnetic Lamb Wave Generation to Gas Pipeline Inspection,"

Proceedings 1972 Ultrasonics Symposium, ed. A. deKlerk, IEEE, New York,

91-94 (1972).

65. H. Tada, P. C. Paris, and G. R. Irwin, The Stress Analysis of Cracks

Handbook , Del Research, Hellertown, Pennsylvania, 5. 1-8.1 (1973).

66. M. Born and E. Wolf, Principles of Optics, Pergamon, Oxford, 430-433

(1975).

67. J. C. Newman, Jr., "A Review and Assessment of the Stress-Intensity

Factors for Surface Cracks," in Part-through Crack Fatigue Life

Prediction , ed. J. B. Chang, ASTM STP 687, American Society for Testing

and Materials, Philadelphia, 16-42 (1979).

68. R. B. King, and Y. W. Cheng, Fracture and Deformation Division, National

Bureau of Standards, Boulder, Colorado, private communication.

222

69. J. M. Richardson and A. G. Evans, "Accept-Reject Decisions for Structural

Ceramics: Application to Failure from Voids," J. Nondestruct. Eval., 1,

No. 1, 37-52 (1980).

70. J. D. Harrison, "Low Cycle Fatigue Tests on Welded Joints in High

Strength Steels," Proceedings, Conference on Welded Structures , Welding

Institute, Cambridge (1971).

71. J. D. Harrison, "The Basis for a Proposed Acceptance Standard for Weld

Defects, Part 1: Porosity," Met. Const. Br. Weld. J. , 4, 99-107 (1972a).

72. J. D. Harrison, "The Basis for a Proposed Acceptance Standard for Weld

Defects, Part 2: Slaq Inclusions" Met. Const. Br. Weld. J., 4, 262-268

( 1972b)

' "

73. F. M. Burdekin, University of Manchester, personal communication.

74. C. R. Boulton, "Acceptance Levels of Weld Defects for Fatigue Service,"

Weld. J. , 56, 13-22 (1977).

75. J. D. Harrison, G. L. Archer, and C. F. Boulton, "Significance of Defects

in Pressure Vessels under Fatigue Loading," Tolerance of Flaws in

Pressurized Components, The Institution of Mechanical Engineers, London,

255-270 (1978).

76. Y. Ishii and K. Iida, "An Analysis of Intermediate Cycle Fatigue Strength

of Defective Welded Joints," Trans. Jap. Weld. Soc. , 3, 1-8 (1972).

77. T. R. Gurney, Fatigue of Welded Structures , 2nd Edition., Cambridge

University Press, Cambridge, England (1979).

78. N. Bailey, "Fisheyes, Hydroaen Embrittlement, and Removal," Weld. Inst.

Res. Bull. , 15, 369-372 (Dec. 1974).

79. A. Matting and M. Neitzel, "The Influence of Weld Defects on the Fatigue

Behavior of Welded Constructions," Br. Weld. J. , 14, No. 1, 3-12 (1967).

80. G. L. Archer, "Research on the Significance of Defects in Weld Regions of

Pressure Vessels," Proceedings Conference on Pressure Vessel Standards,

The Impact of Change , The Welding Institute, London, 100-106 (1972).

81. G. D. Fearnehough and D. G. Jones, "An Approach to Defect Tolerance in

Pi pel ines , "paper ERS E. 199, Proceedings Conference on Defect Tolerance

of Pressure Vessels , paper ERS E.199, Institution of Mechanical

Engineers, Great Britain (1978).

82. R. F. Lumb and G. D. Fearnehough, "Toward Better Standards for Field

Welding of Gas Pipelines," Weld. J. Res. Suppl . , No. 54, 63-s-71-s

(1975).

83. R. P. Newman and T. R. Gurney, "Fatigue Tests on \ in. Thick Transverse

Butt Welds Containing Slag Inclusions: 1st Interim Report," Br. Weld. J. , 11, No. 7, 341-352 (1964).

223

84. R. C. Placious, D. A. Garrett, M. B. Kasen, and H. Berger, "Dimensioning

Flaws in Pipeline Girth Welds by Radioqraphic Methods," Mater. Eval., 39,

755-560 (1981).

85. J. Norrish and D. C. Moore, "Porosity in Arc Welds and its Effect on Mechanical Properties," Proceedings Second Conference on the

Significance of Defects in Welds , Welding Institute, Cambridge (1968).

86. J. D. Harrison, "The Basis for a Proposed Acceptance Standard for Weld

Defects, Part 1: Porosity," Metal. Const. Brit. Weld. J., 4, 97-107

(1972).

87. "Guidance on Some Methods for the Derivation of Acceptance Levels for

Defects in Fusion Welded Joints, PD6493, British Standards Institution,

London (1980).

88. R. P. Reed, H. I. McHenry, and M. B. Kasen, "A Fracture Mechanics

Evaluation of Flaws in Pipeline Girth Welds," Wei d. Res . Counc. Bui 1

245 , Welding Research Council, New York (Jan. 1979)

89. P. Tait and D. M. Haddrill, "Fracture Toughness of Some Mild Steel Manual

Metal-Arc Weld Deposits," Weld. Met. Fabr. , 370-375 (1970).

90. S. K. Datta, University of Colorado, Boulder, Colorado (private

communication.

91. R. B. King and C. M. Fortunko, "Determination of In-Plane Residual Stress

in Plates Using Horizontally Polarized Shear Waves," J. App. Phys., 54,

3027-3035, (1983).

224

12. BIBLIOGRAPHY OF RELEVANT NBS PUBLICATIONS

12.1 Fracture Mechanics

"Consideration of Fracture Mechanics Analysis and Defect Dimension

Measurement Assesssment for the Trans-Alaska Oil Pipeline Girth Welds,"

Vols. I and II, NBSIR 76-1154, National Bureau of Standards, Boulder,

Colorado (1976).

Y. W. Cheng, R. B. King, D. T. Read, and H. I. McHenry, "Post Yield

Crack-Opening Displacement of Surface Cracks in Steel Weldments,"

submitted for publication in Fracture Mechanics, Fifteenth

Symposium , American Society for Testing and Materials, Philadelphia.

Y. W. Cheng, H. I. McHenry, and D. T. Read, "Crack-Opening Displacement

of Surface Cracks in Pipeline Steel Plates," to be published in

Fracture Mechanics, Fourteenth Symposium , STP 793, American Society

for Testing and Materials, Philadelphia.

R. deWit and J. H. Smith, "Development of Some Analytical Fracture

Mechanics Models for Pipelien Girth Welds," Fracture Mechanics,

Twelfth Conference , STP 700, American Society for Testing and

Materials, Philadelphia, 513-528 (1980).

225

R. B. King, Y. W. Cheng, D. T. Read, and H. I. McHenry, "J-Integral

Analysis of Surface Cracks in Pipeline Steel Plates," to be

published in Elastic-Plastic Fracture Mechanics, Second

International Symposium , American Society for Testing and Materials,

Philadel phi a.

H. I. McHenry, D. T. Read, and J. A. Begley, "Fracture Mechanics Analysis

of Pipeline Girth Welds," Elastic-Plastic Fracture , STP 668,

American Society for Testing and Materials, Philadelphia, 632-642

(1980).

D. T. Read and H. I. McHenry, "Strain Dependence of the J-Contour

Integral in Tensile Panels," Advances in Fracture Research ,

Pergamon, New York (1980).

R. P. Reed, H. I. McHenry, and M. B. Kasen, "A Fracture Mechanics

Evaluation of Flaws in Pipeline Girth Welds," Bulletin 245, Welding

Research Council, New York (1979).

12.2 Inspection and Dimensioning of Flaws

S. K. Datta, C. M. Fortunko, and R. B. King, "Sizing of Surface Cracks in

a Plate Using SH Waves," Proceedings, 1981 Ultrasonic Symposium , ed.

B. R. McAvoy, IEEE, New York, 863-867 (1981).

C. M. Fortunko, "Ultrasonic Detection and Sizing of Two-Dimensional

Defects at Long Wavelengths," Appl . Phys. Lett. , 38, No. 12, 980-982

(1981).

C. M. Fortunko, R. B. King, and M. Tan, "Nondestructive Evaluation of

Planar Defects in Plates Using Low-Frequency Shear Horizontal

Waves," J. Appl. Phys. , 53, No. 5, 3450-3458 (1982).

226
.
C. M. Fortunko and R. E. Schramm, "An Analysis of Electromagnetic-Acoustic

Transducer Arrays for Nondestructive Evaluation of Thick Metal

Sections and Weldments," Proceedings, 1982 Review of Progress in

Quantitative Nondestructive Evaluation , eds. D. 0. Thompson and

C. E. Chimenti, Plenum Press, New York (1983), in press.

C. M. Fortunko and R. E. Schramm, "Ultrasonic Nondestructive Evaluation

of Butt Welds Using Electromagnetic-Acoustic Transducers," Weld. J. ,

61, No. 2, 39-46 (1982).

M. B. Kasen, G. E. Hicho, and R. C. Placious, "Inherent Through-Wall

Depth Limitations on Blunt Flaws in Welds," accepted for publication

in Weld J.

R. B. King and C. M. Fortunko, "Extended Variational Solution for

Scattering from Flaws in Plates," J. Appl . Phys. , 53, No. 5,

3459-3460 (1982).

R. C. Placious, "Radiographic Inspection of Pipeline Welds," Pi pel ine

Welding and Inspection , American Welding Society, Miami, 51-63

(1980).

R. C. Placious, D. A. Garrett, M. B. Kasen, and H. Berger, "Dimensioning

Flaws in Pipeline Girth Welds by Radiographic Means," Mater. Eval

39, 755-760.

12.3 Significance of Blunt Flaws

M. B. Kasen, "Significance of Blunt Flaws in Pipeline Girth Welds,"

accepted for publication in Weld. J.

227

M. B. Kasen and R. P. Mikesell, "Interim Report on the Significance of

Blunt Flaws in Pipeline Girth Welds," Pipeline and Energy Plant

Piping: Design and Technology , Welding Institute of Canada, Pergamon

Press, Toronto, 329-336 (1980).

M. B. Kasen and R. P. Mikesell, "Effect of Blunt Flaws Studied in

Pipeline Girth Welds," Oil Gas J. (Techno! .

, 155-159

(March 30, 1981)

12.4 General

M. B. Kasen and C. M. Fortunko, "Fi tness-for-Purpose Criteria for Pipeline

Girth Welds," presented at Pipeline Welding and Inspection Conference,

American Welding Society, Houston, Texas (Sept. 21-22, 1982).

R. P. Reed, "Goal: Better Standards for Pipeline Welds," Weld, Des.

Fabr ., 68-71 (1981).

C. Smith, "Putting It All Together--Al askan Pipeline Weld Criteria,"

Dimensions

, 3-5 (1980).

t

228

13. APPENDIXES

Appendix A - Details of Calculation of Allowable Flaw-Size Curves

The following computer program was used to calculate the allowable flawsize curves presented in this report. Each curve required approximately 40 s

of processor time on a modern large-scale scientific computing facility.

229

PROGRAM YI E L IG( INPUT# OUTPUT- /lOOO# TAPE 88-/1 OOQ# TAP El #TAPE2)

REAL NU

NOTATION*

H IS CRACK

IS

IS

IS

IS

IS

cT

R

W

Cl

DEPTH

CRACK HALF LENGTH

PLATE THICKNESS

PIPE RADIUS

PLATE FULL WIDTH

2*C / W

A 1 IS H/T

DELTA2 IS (EFFECTIVE CRITICAL CTOD)

E IS TOUNGts MOOULUS

NJ IS POISSONtS RATIO (*** REAL

STRESS IS APPLIED STRESS

YIELDST IS YIELD STRENGTH

S IS STRES S/YIELDST

AO IS ARRAY OF CRACK DEPTH VALUES

E / W / YIELDST

NUMBER ****)

FLO I S ARRAY OF CRAC K LE NGTH VALUES

COMMON PI# DELTA2# Cl# S

COMMON /GE3M/ H# C # T# R

COMMON /MAT PR/ YIE LOST # E# NU

DIMENSI ON FL ! 1 00 )

A (

00)# L MESS ! 8

PI * 3. 191592659

PRINT 1

FORMAT! * PROGRAM TO CA LOLA TE AND PLOT A VS

1DED LIG AMENT MODEL .*#/ /)

PRINT 2

FORMAT! CRITICAL COD# YIE LC STRENGTH,i ANC

1READ IN .*)

READ (1 #*) CODCR# YIEL DST# STRESS

I F ( E OF l n.NE.O) GOTO 9 999

CODR ES « 0,0015

YSKSI - YIELDST / 1000 • SKSI STRESS / 1000.

L CURVES USING THE YIEL

APPLIED STRESS ARE NOW

SET UP AND MARK THE PLOT

PLOTTING SUBROUTINES FROM THE ADISSPLA*

PROGRAM,

(REG. TRADE MARK)

PACKAGE ARE USED IN THIS

CALL ID ! AD AVE READ# X CALL C3KPRS

CALL NOBRDR

CALL AR E A2D ! 2 . # 2.)

CALL HWSPEC (88#*FILE*

CALL ENDPL! 0)

CALL RESET ( ANOBRDR A

CALL TITLE! AALL0WA3LE

1 ADEFECT LENGTH# IN*

2 ADEFECT depth# INS*

3 6. 0# 6,0)

X 3 853 ) A VS L**#100)

FLAW SIZES**, 100#

230
*

*

"fPli v^ES'

> . W-f ) m

s # u

•*-t-inw
I CALL GRAF (C.O# 5.0 # 25.0# 0.0# 0.1# 0.6)

ENCOOEteO# 410# LHESS) CODCR

t

*10 FORMAT! MCRITIC AL CTOD M#F7.3#M IN.$M)

CALL MESSAG(LMESS#100#1.6# 5.5)

ENCODEC 80#415#LMESS) CODRES

f

l5 FORMAT! ARES. 5 TRS • CTOD RED. » M#F8.4#i< IN.S*»)

CALL MESSAG!LMESS#100#1.6# 5.0)

C

j ENCODE! 80#420#LMESS) YSKSI

420 FORM AT( MF LO W STRENGTH » M#F5.1#M KSI.S#)

CALL MESSAG!LMESS#100#1.6# 4.5)

ENCODE( 80# 430# LM ESS) SKSI

*30 FORMAT! ^APPLIED STRESS *#F5.1#«* K S I • S #

CALL MESSAG!LMESS#10C# 1.6# 4.0)

C

PRINT 10# CCDCR# TIELDST# STRESS

10 FORMATS CRITICAL COD *# F12.6#* IN*#/#

1 YIELD STRENGTH • *#F12.0#* PSI *#/#

I 2 * APPLIED STRESS - *# F12.C#* PSI.*)

C HERE ME SPECIFY A STEEL PIPELINE# 43 INCHES IN DIAMETER

(; E 30.2E6

NU • 0.29

R • 24'.

, W - 2. * PI * R

j T * .62 5

t EFFECTIVE CRITICAL COD IS CRITICAL COD LESS RESIDUAL STRESS EFFECT.

CODCR - CODCR - CODRES

I DELTA2 E * CODC R / V / Y I E LDST

lli I NPTS 100

RELlMAX • 0.15

I S « STRESS / YIELDST

DO 1000 IPT • 1# NPT

Cl - RELLMAX FLOAT (NPTS - IPT) / FLOAT! NPTS

f

C X * Cl / 2.

three candidate values of first trial value of ai are calculated

A1SIMP * 1. - S DELTA2/2./C1

I

I

!

A1LY • 1. - S

A1NSY (1. - S) / Cl

A1NSY * AMINK A1NSY#1.0)

A1TR » AMIN1 (A1SIMP# 0.99999*A1NS Y

DEVTARS 1.0E-06

STEPS2 - 1.00

DEVPREV 1.0E100

DO 30C I IT - 1# 15

DEVVAL • DEV (A1TR)

DEVSIZ » ABS(DEVVAL)

IF(DEVSIZ .GT. DEVPREV) GO TO 280

231

I DEVREL - ABSIDEVVAL/DELTA2)

s IF ( DEVREL • LT. DEVTARG) GOTO 312

A1TRN • AIT R - STEPSZ * DE VV AL /DDE VOA ( A1TR

1FIA1TRN.GT.A1NSY.OR.A1TRN.LT.! A1LY/2.)) GOTO 280

A1TR A1TRN

DEVPREV * DEVSIZ

IF (STEPSZ .LT. 1.0) STEPSZ AHINK 1 . 3*STEPSZ# 0. 5* i STE PSZ*1.0)

GOTO 299

r

8 0 STEPSZ STEPSZ/2.

DEVPREV « 1.0E100

A1TR • 0.75 * A1LY

II _ _ ...»

A F A 1 L A1TR*T

FLFA1L « H * Cl

PRINT 3 06, FLFAIL# AFAIL

'306 FORMAT! A FLAW DEPTH ITERATION FAILED TO CONVERGE AT*,/#

a * LENGTH • W,F7.2## DEPTH * *,F5.2)

1312 A 1 • A1TR

FL(IPT) • W * Cl

A(IPT) T * A1

, IF(A1.GT.C.75) GOTO 1020

1000 CONTINUE

1020 1LAST • IPT

D3 1CAD I 1# IL AST

PRINT 1025# FL ( I ) # A(I)

1025 FORMAT! 2E14. 5)

1040 CONTINJE

C

C CALL CJRVE! FL, A, ILAST,0)

CALL FRAME

CALL ENDPLt-1)

FUNCTION DE V C A 1

REAL NJ

COMMON PI, DELTA2# Cl# S

COMMCN /GE3M/ H, C# T# R

common /matpr/ yieldst, e# nu

H » T * A1

W 2. * PI * R

CC GOTO A 9999 CALL DTNEPL

CALL EXIT

END

J

232

!300

|315

FOR THIS VERSION* ALWAYS USE SIMPLIFIED LINE SPRING

USE OF CURVEATURE IN SLS DEPENDS ON CRACK LENGTH

I F ( H • GT . O.O) GO TO 100

CALL NEKMANCH# C# T# W# F# Q)

SLY 1.-A1

COD-J FACTOR M SET TO 1 IN DENOMINATOR OF TERM ON RIGHT OF NEXT STATEMENT.

DELELAS « SLY*SLY*F*F*PI*H/C/H

IF (S .GT. SLY) GOTO 200

COD-J FACTOR M SET TO 1 IN DE NC MI NAT OR OF

2 ND TERM ON RIGHT OF NEXT STATEMENT.

OEV • 0ELTA2 - S*S*F*F*PI»H/0/V

RETURN

CONTINJE

STRESS • S Y IE LOST

CALL LS J ( Y1 ELDST# E# NU# H# C# T# R # STRESS# 0.0# CTOD#CMOD#

a STAPL Y#STCLOS)

DELELAS • CTOD*E/K/ YIELDST

IF (5TAPLY .GT. 0.0) SLY * STAP LY/ Y IE LDST

IF (STAPLY .GT. 0.0) GO TO 200 DEV • DELTA2 - OELELAS

GO TO 300

CONTINJE

CPI • (2. /PI) * AS I N ( SIN(PI*Cl/2«) / COS ( P I * < 1

. -C 1. -S ) / A1 ) / 2 . ) ) RY1 « 0.5 * (CP1-C1)

DEV • DELTA2 - DELELAS - 2 .* (C1*R Y1 ) * ( S-SL Y)

CONTINUE

K P I T E ( 2# 3 1 0 ) S# STRESS# STAPLY# SLY

FORMAT ( * S • K# E14.5# K STRESS - M# E1A.5# M STAPLY M#

a E14.5# t SLY • ## E14 • 5 ) VRITE<2#315) H# C FOR M AT ( * H ## E1A.5## C • M# Ei4.5)

WRITE(2#320) DELTA2# DELELAS# DEV F0kMAT(# DELTA2 • *# E1A.5# MOEIELAS • ## E14.5# # DEV * ##E14.5)

END

FUNCTION DDEVOA(A)

Y • DEV(A)

AM, * A * 0.99 - 1.0E-03

IF(AM.LT.O.O) AM* 0.0

YM - DEV(AM)

DDE VDA

END

(Y YM) / (A - AM)

233

I

SUBROUTINE NE WMAN (

H

# C # T# W p F p 0 )

I

COMMON PI# DELT42# Cl# S

REAL Ml# M2# M3

CSAV C

C FORMULAS SET UP FOR C > H

IF (H .GT. C) C*

H

ARG CPI/2,) * (2. * C / »> SORT (H/T

)

FV*l./SQRT(COS(ARG)

)

(

ANGULAR FACTOR • FPHI 1 AT ROOT

F PHI • 1.

G FACTOR • 1 AT ROOT

G 1.

M1-1.13-0.09MH/C)

M2—0.54+0. 89/ (0.2 + H/C ) M3«0.5-l./< 0.65 + H/C ) + 14.*< l.-H/C)**24

I

F«(HlH2»(H/T)M2 + fi3*(H/T)*m*PPHI*G*FW

0*1. 41. 464*<H/C)+*1.65

C CSAV

1 RETJRN

END

SUBROUTINE LSJ (3 FLOW# E#NU# H# C# T# R, ST# SB# CT OD#C MOD# ST A PL Y# STCLOS

)

REAL NJ* NU7# NU8, KF

DATA ( PI«3 .14159265)

> SIMPLIFIED LINE SPRING MODEL FOP CALCULATION OF APPLIED CTOD FOR

C SURFACE FLA* AS A FUNCTION OF STRESS FOR ELASTIC CASE WITH ROOT PLASTIC

ZONE CORRECTION.

NOMENCLATURE

HPHTS • PHYSICAL CRACK DEPTH

c H EFFECTIVE CRACK DEPTH

T • PLATE THICKNESS

% CPHYS PHYSICAL CRACK HALF-LENGTH

c C • EFFECTIVE CRACK HALF-LENGTH

• PLATE FULL WIDTH

ST • APPLIED TENSILE STRESS

SB « APPLIED BENDING STRESS

f SFLOV FLOW STRENGTH

{

E » YOUNGti MODULUS

t NU PDISSONtS RATIO

C LLY FLAG FOR LIGAMENT YIELDING

£

ASSUME PROPORTIONAL LOADING. LET B3NT M/S

' BONT * SB /ST

C HERE the calculation begins

I HPHYS H

HSAV - H

CPHYS * C

j

c UY IS FLAG INDICATING LIGAMENT YIELDING. LL Y<0 MEANS NO LY.

I

234

I

LLY —100

*C PROCEDURE TO SEE IF LIGAMENT IS FULLY YIELDED OR NOT

C MAXIMUM EFFECTIVE VALUE OF H IS H (T-H)/2

C AT THIS H VALUE* MAXIMUM CLOSURE STRESS HAS BEEN REACHED

jJc MAXIMUM CLOSURE STRESS* AT FULL LIGAMENT YIELD* IS SCMAX

' C MAXIMUM CLOSURE STRESS OCCURS WHEN SFLOw IS APPLIED ON

f

T HE WHOLE LIGAMENT

SCMAX-SFLOW*(T-H)/T

t MAXIMUM ALLOWED EFFECTIVE H EXTENDS HALFWAY ACROSS THE LIGAMENT

H«H*(T-H)/2.

|c STAPLY IS APPLIED STRESS AT FIRST OCCURENCE OF FULL LIGAMENT YIELD

CALL C ALC I ( H* C* T* R* NU* 02* 03* 04* D* NU7* NUB

STaPLY-SCMAX*D/(NJ7*(04+MJ8)-NU8+03*BDNT)

IF( ST .GT. STAPLY ) GO TO 52

Jc ligament is not fully yielded

C SET STAPLY • -1000. TO INDICATE THAT THE LIGAMENT HAS NOT FULLY

,C YIELDED

STAPLY • -1C00.

*C CALCULATION OF ROOT PLASTIC ZONE STARTS HERE

i C SAVE STARTING VALUE OF H* NO PLASTIC ZONE* AS HSAV

|C H STARTS AT H IS HSAV* AND INCREASES

1C UNTIL THE CORRECT PLASTIC ZONE SIZE

C IS REACHED BY AN ITERATIVE CALCULATION

)C NOMENCLATURE* HSAV IS VALUE OF H BEFORE ENTERING ITERATION

C HS T ART IS VALUE OF H AT START OF ITERATION

C HCI IS VALUE Of H AT START OF CURRENT ITERATION

C CELT A H IS SMALL INCREMENT OF H

,C BEGIN ITERATION

H « HSAV

HSTART-H

DO 50 11-1*30

HCI-H

DELTAH- .01*H

IF (DELTAH .LT. 0.0001) DELTAH * 0.0001

H-H+DELTAH

CALL CALC1 <H*C*T*R*NU* 02 *03* 04* D* NU7* NU8

CALL CLOSURC 02*03* 04* D* NU7* NUB* ST* SB* STC LOS* S BCLOS

C STCLOS* SBC LOS ARE TENSILE ANC BENDING CLOSURE STRESSES FOR H-DELTAH

‘C SAVE THESE VALUES OF TENSILE AND BENDING CLOSURE STRESSES

STCSAV-STCLOS

SBCS AV- SBCL OS

C ALTERNATE PROCEDURE IS TO CALCULATE CLOSURE STRESS FROM THE

C PLASTIC ZONE SIZE. THE IDEA IS TO FCRCE CONSISTENCY.

C CLOSURE STRESS FROM PLASTIC ZONE SIZE IS SCLOSPZ

CALL $CPZ(H*H$TAPT,T*SFLOh* SCLOSPZ)

C SCLOSPZ IS ALTERNATE CLOSURE STRESS FOR H+DELTAH

SCPZSAV-SCLOSPZ

H-H-DELTAH

call CALC1( H*C*T*R*NU* 02*03*04*D*NU7*NU8

CALL CLOSUR (02*03* 04* D*NU7*NU8* ST* SB* ST

L 3 S * S BC LOS

CALL SCPZ(H*HSTART*T*SFLOW* SCLOSPZ)

: I

235

Ic DSTCDH# DSBCDH ARE DERIVATIVES OF CLOSURE STRESS# MOMENT VRT H

dstcdh*(stcsav-st:los)/deltah

j DSBCQH«(SBCSAV-SBCLOS l/DELTAH

OSCPZOH-ISCPZSAV-SCtOSPZ) /DELTAH

C MAKE THE INDICATED CORRECTION TO H

H«H*(STCLOS-S CL OS PZ)/ CDS CPZDH-DSTCDH)

I F ( HCI • LT. O.OOOOOl) GO TO 50

5 I F ( ABS( (H-HCD/HCI) .LT. O.OCl ) GO TO 55

50 CONTINJE

|C ITERATION COMPLETE. H IS EFFECTIVE CRACK LENGTH

J PRINT 51# H# C

51 FORMAT ( # ITERATION ON PLASTIC ZONE SIZE FAILED TO CONVERGE.*#/

t a # h • *#F7.A#* C • *# F7.4)

GO TO 55

CONTINJE

LIGAMENT IS FULLY YIELDED.

HERE W E CALCULATE THE CNOD

THESE WILL BE ADDED TO THE

LIGAMENT MODEL LATER

LLY-100

SBAPLY-BONT^STAPLY

CALL CLQSUR (02#33#34#D

STCLOS • STCLLY

SBCLOS • SBCLLY

ST • STAPLY

SB S3APLY

CTOD ACCRUING AFTER LI&AM

CONTINJE

H«HPHYSMT-H)/2

AND CTOD UP TO LIGAMENT YIELDING

CONTRIBUTION CALCULATED BY THE YIELDED

#NU7#NU8# STAPLY# SBAPLY# STCLL Y# S BCLLY)

ENT YIELDS WILL BE CALCULATED ELSEWHERE

WE NOW HAVE A PLASTIC-ZONE-CORRECTED CLOSURE STRESS STCLOS

A PLASTIC-ZONE-CORRECTED CLOSURE MOMENT SBCLOS

NOW CALCULATE DELTA AND THETA# TO GET COD S

delta«a.*cmst-st:los)/e

USE CURVED PLATE EXPRESSION FOR LONG CRACKS# FLAT PLATE

EXPRESSION FOR SHORT CRACKS

IF (2 • *C .GT. 0. 025*2. 0*PI*R ) GO TO 7C

THETA • (SB-SBCL3S) 8 . * C 1 .NU ) *C / ( 3 . «NU ) / E/T

GO TO 80

THETA . SORT (R/T ) * (2./3.) * (( 3. ( 1 .-NU+NU )

) **0 . 75

* * (SB-SBCLOS) /E

CONTINUE

HERE WE CALCULATE APPLIED K FACTOR# TO GET CTOD

CALL KCALC( STCLOS# SBCLOS# H# T# KF)

CTOD KF*KFM1.-NU*NU)/E/SFL0W

CMuD*DELTA+THETA*T

H HPHTS
'
SUBSOUriNE SCPZ(H#HSTART#T#SFLOw# SCLOSPZ)

THIS SUBROJTINE CALCULATES ALTERNATE CLOSURE STRESS SCLOSPZ

FROM PLASTIC ZONE SIZE

PI 3.1415927

HPLAS»H$TART*2.*(H-HSTART)

IF ( HPL AS .LT. 1.0E-10) €0 TO 10

0-SlN(PI*HSTART/2./T)/SIN(PI*HPLAS/2./T)

I F C 0 • GT. 0.9999) GO TO 10

SCL0SPZ«(2.*SFL0rf/PI)*AC0S<0)

RETURN

SCLOSPZ 0.0

END

SUBROJTINE CALC1(H#C#T#R#NU#

REAL NU# NU7# NUB

DIMENSION X( 13) # Y(19)# Z(13)

DATA (PI * 3.14159265)

DATA*

X ( 1) *1,9761 )#

X(2) 11.487)# , X ( 3)

7. 7086 )#

X(4)*15.0143 )#

X(5)-2B0.121 )#

X(6)—1099.72)#

X(7)-3418.98 )#

02#03#04#D#NU7#NU8)

X ( 8 ) * X ( 9 ) * X(10)

X ( 1 1

X( 12 )*

X ( 1 3

DATA

r c l) * l Y ( 2 ) Y ( 3 ) Y ( 4 ) * Y ( 5 ) * Y (6 ) Y ( 7 ) * Y ( 8 ) Y ( 9 ) Y (10)

Y ( 1 1

Y ( 1 2 ) Y ( 1 3 ) Y ( 1 4 ) Y ( 1 5

Y ( 1 6

Y ( 1 7

Y ( 1 8

Y ( 1 9 )

-7686 .92 ) #

12794.1 )# —13185.)#

7868 .27)#

-1740.25 )#

•124. 136)

.9735)#

-2.2166 )#

21.6051 )#

-69.3133)

196.3 )#

-406.261)#

644.935 )>

-408.957)#

-159.693)#

-988.988)#

4266.55)

—2997.14 )#

-6050.78 )#

•8855.36)#

3515.43)# —11744.1 )#

4727.98)#

•1695.61)#

-845.896 )

237

I

I

II

il

!

DATA

Z(l)*l

2<2)«

2 13)-

ZC4)-

zmZ ( 5 ) • 2C7IZ ( 3 ) Z(9)*

2(10)

Z € 1 1

2 ( 12 ) Z ( 1 3

XI* H/T

02 * 0 . 03-0*

04*0.

IF< XI .

.971 )#

-4.4277 )#

34.4952 )#

-165.732)#

626.393 )#

-2144.46)#

7043.42 )#

-19003.2)#

37053.3 )#

*-52595.5)

•48079.3)# —25980.2 )#

*6334.24)

LT. l.GE-5) Xl-l.CE-5

DO 10 K*l# 13

1 K - 1

02«Q2+X(K)*(X1**(2 .*1 )

03*Q3+Y(K) *( X 1** I

C4*04+Z(K) * ( X 1** I

CONTINUE

DC 20 K*14#19

I*K-1

03*0 3* Y(K) * ( X 1** I

) 20 CONTINUE

02* Xl*Xl*02

? 03* XI* XI* 03

04*X1*XI*04

C IN THE NOTATION OF KINS+S PAPER ON THE SIMPLIFIED LINE SPRING MODEL#

C 02 *A1 1 03* A12 0 4* A22

C NU-NU# NU7»(C/T)*(2/(1-NU*NU) )

f NU7* (C/T)*2./C1.-NU*NU)

, C THE QJANTITY N6 DEPENDS ON THE CRACK LENGTH. FOR CRACKS LESS

jc than 2.5 * of the circumference# we use the flat plate expression.

|C FOR LONGER CRACKS# THE EXPRESSION FOR C'JRVEATURE IS USED

l I F { 2 •* C .GT. .025*2 . *PI*R) GO TO 3C

NU8 (C/T)*2./3./(l.-NU) /(3.+NU)

60 TO 40

1 30 NU8 SORT ( R/T ) (1./6.) / (( 3 . ( 1 .-NU*NU) ) **.25)

40 D*(02+NU7)*(04*NU8)-03*03

END

ll

SUBROUTINE CLQSUR(02#03,04#D#NU7#NU8,ST,SB# STCLOS# SBC LOS ) REAL NJ7, NUB

C CLOSURE STRESS# SIGMA SUB C# IS STCLOS

C CLOSURE MOMENT# M SUB C# IS SBCLOS

STCL03*(NU7*(C4*NJ8)*ST-NUe*03*S5)/D

SBCL0S*(-NU7*03*ST«, N0 8*{02*K07)*5B)/D

END

SUBROUTINE KCALC ( STCLOS# SBCLOS # H#T# KF)

REAL KF

DIMENSION U { 7 ) # V ( 7 > DATA (PI 3*141592654)

DATA (J(l) • 1.12)#

* (U ( 2) • 6,52)#

a (U ( 3) * -12.39),

a (U ( 4 ) • 09.05)#

a ( J ( 5) -188.61),

a ( U ( 6) * 207.39),

a ( U ( 7) * -32.05)

DATA (VCl) • 1.12)#

a ( V ( 2 ) • -1.89),

a ( V ( 3

» 18.01),

A U(4) • -87.39),

a ( V ( 5) « 241.9),

a U(6) * -319.94),

a ( V ( 7) * 168.01)

XI H/T

G 1 * 0.0

G2 • 0.0

DO 10 K* 1# 7

I K - 1

G1 * G1 U (K ) *( XI** ( 2*1 ) ) G2 « G2 VCK)*(X1«*I)

10 CONTINJE

KF * ($TCL0S*G1 SBCL0$*G2)*S0RT(PI*X1)*5QRT(T)

RETURN

END

239

Example allowable flaw-size calculation

1. Input parameters

6 = 0.125 mm

c CT0Dp

s = 0.038 mm

o = 517 MPa

a = 434 Mpa

E = 2.08 x 10

5 MPa

v = 0.29

R = 610 mm

t = 15.9 mm

2. Short flaw, use (28) - (33)

Long flaw, use (28-c) - (33-c)

Dividing line is 2c = 0.025 x 2 it x R = 96 mm

3. Critical condition: 6 = CT0D r + CT0D, v + CT0D dc C t LY KS

Therefore we demand CTOD^ + CTOD^y = 0.087 mm

4. Procedure: for every i , find a so that

CT0D

e + CT0D

LY = 0.087

4.1 Choose sl = 201.1 mm for this example. From 2, above, we see that we

have the long-flaw case.

4.2 Choose a trial a value, 2.75 mm.

(The first trial a value is arbitrary, but the better the choice, the

fewer iterations needed).

4.3 Calcualte CTOD^^ + CTOD^y for this a value.

4.3.1 Check for full igament yielding.

4. 3. 1.1 Using (14),

o

c

= (1 - a/t) a = 427.6 MPa

240

4 . 3 . 1.2 At full ligament yielding, a +2 r = t; r = 6.575mm;

a^^ = a + = 9.32 mm a^^/t = 0.586

To apply (29-c) we need a

c from (31-c), which requires S from (33-c).

In (31-c) and (33-c) we need a^, a^* and * 22

:

From Table Al, a^ = 3.136; a^ = 1.746; 822

= 0.993

Applying (33-c) we find

S = {3.136 + itVn 0 ?Q2\ } (0.0993 +( 2 - 610 )* [12(1-. 29 2

) 7 15.9 ( 1-0.29 j 72 (1 - 292) 15.9*

-(1.746) 2

= {3.136 + 13.809} {0.993 + 0.8018} - 3.0485

= 27.364

Applying (31-c), using values from the previous step where needed, we get

a

c

= 13.809 {0.993 + 0 . 801 8 } / 27 . 364 = 0.9057

Applying (29-c), we find

o

c = 0.9057 X 434 = 393.1 MPa

4.3. 1.3 We note that a < 0 , therefore the ligament is not

0 c fully yielded.

Therefore, we know

ctod

ly = 0

r < 6.575 mm

y

4.3.2 Find the vlaue of r^ for which (29-c) and (69) are satisfied.

4. 3. 2.1 Guess initial value of r^:

let r^ = 4. 164 mm

4. 3. 2. 2 Apply (29-c).

r = 4.164

(a/t)

E ffective

= 0,4355

241

From Table Al, a^ = 0.8989; a^ = 0.5917; 822

= 0.3955 (33-c) gives:

S = {0.8989 + 13.809} {0.3955 + 0.8018} - .3501

= 17.260

(31-c) gives: a

c

= 13 . 809{ 0 . 3955 + 0.8018}/ 17.260

= 0.9579

(29-c) gives: a

Q

= .9579 x 434 = 415.7 MPa

4. 3. 2. 3 Apply (69).

sin(0.2717) _ .2683

sin(l. 0944) .8887

"

arc cos (0.3019) = 1.2641

0.3019

o 416.1 MPa

0

4. 3. 2.4 a

o from (29-c) agrees with a

Q

from (69) to

than 1 part in 1000. This is sufficient for a

calcualtion. Therefore r^ = 4.164 for the trial

a = 2.75 mm.

4.3.3 Apply ( 66 )

We need g and g

b

From Table Al, we have g^

= 2.7187; g^ 1.5561

We already have o Q

= 415.9, averaging results of 4.3

4. 3. 2. 3.

To get m , we apply (32-c), then (29-c) using vlaues calculated above

6 = 13.809 x 0.5917/17.260 = 0.4734

c From (29-c) , m = -0.4734 x 434 = -205.4 MPa

better

hand

a value.

. 2.2 and

we have

242
'
In applying (66), we convert t to meters to retain consistent units:

K = (.0159)* [415.9 x 2.7187 - 205.4 x 1.5561]

= 102.3 MPa/m

4.3.4 Apply (11)

CT0D

e

= 8.913 x 10' 5

M

= 8.913 x 10~^mm

= .08913 mm

4.3.5 Our calcualted CTOD is within 2.4% of the critical

This is sufficiently accurate for a hand calculation.

5. Therefore, for a = 201.1 mm, we have found a = 2.75 mm, and we

point on the allowable flaw-size curve.

vl aue

have one

243

E

E

o>

o

<

ID

Z

o

N

O

h

(0

<

-J

a

TENSILE STRESS ON EDGE CRACK, <

Q, MPa

Figure A1 . Plot of plastic zone radius, r , against tensile stress on equivalent edge crack, cr0 , for two different equations. The intersection

of the two curves is the plastic zone radius for the specified input parameters.

244

Table A1 . Parameters used in calculation of allowable

flaw size curves.

Effective Effective

-4ii 11? A22 ££_

.0100 .0002 .0002 .0002 .1986 .1955 • 0100

.0200 .0006 .0006 .0006 . 281 * .2729 • 0200 . -0300 .0016 .0017 .0017 - 3 * 8 * .3307 -0800

.0100 .0032 .0031 .0929 .*007 .3787 . 0*00

• 0500 .0050 • 00*6 . 00*5 .*503 .4205 • 0500

- 06 P 0 .0073 .0069 .0065 .* 9*4 .*• 8 ? -0600

.0700 .0100 .0093 .0087 . 5*01 .*930 .0700

.0600 .0131 .0122 .0113 .5822 .5255 .0600

• OQPO .0166 .0155 . 01*3 . 623 ? .5863 -0900

.1000 .0209 .0191 • 0175 .6637 .5856 •1000

.1100 • 0256 .0233 .0211 .7038 •6139 .1100

.1200 .0309 .0276 .0251 . 7*39 . 6*1 2 .1200

.1900 .0367 .0328 .02 9* . 76*2 • 6676 .1300

. 1*00 . 0*32 .0363 . 03*0 . 62*6 . 693 * . 1*00

.1500 .05 04 . 0**3 .0390 . 6*59 .71 66 -1500

.1600 .0563 .0508 . 0 *** .9076 . 7* 3 * .1600

.1700 .0669 .0576 .0501 . 95 C 1 .7677 .1700

.1600 . 076 * . 065 * .0562 . 993 * .7917 .1600

.1900 .0667 .0736 • 0626 1.0377 .6155 .1900

.2000 .0960 . 062 * .0695 1.0830 .6392 .2900

.2100 .1102 .0918 .0767 1.1293 .86 26 .2100

.2200 .1236 .1019 • 06 ** 1.1769 . 866 * • 2200

.2300 .1360 .1127 .0925 1.2256 .9101 • 2300

. 2*00 .1537 . 12*3 .1010 1.2760 . 93*0 . 2*00

.2500 .1707 .1366 .1100 1.3277 .9561 • 2500

.2600 .1691 . 1*98 . 119 * 1.3610 .9627 •2600

.2700 .2090 .1639 . 129* 1 • *35 9 1.0077 .2700

.2600 . 230 * .1788 .1398 1.4926 1.0332 .2600

.2900 .2536 . 19*8 .1508 1.5511 1.0593 .2900

.3000 .2787 .2118 .1623 1.6117 1.0861 .3000

.3100 .3056 .2299 . 17 ** 1 . 67*5 1.1137 .3100

.3200 . 33*9 . 2*92 .1872 1.7395 1 . 1*22 .3200

.3300 . 366 * .2696 .2006 1.6070 1 .1716 .3300

. 3*00 .* 00 * .2917 . 21*8 1.6771 1.2020 . 3*00

.3500 .*371 .3150 .2296 1.9500 1.2335 .3500

.3600 .*767 . 3*00 . 2*53 2.0259 1 . 266 ? .3600

.3700 . 519 * .3665 .2618 2 . 10*9 1.3000 .3700

.3600 .5655 . 39*9 .2793 2 . 167 * 1.3351 .3800

.3900 . 615 * .*251 .2977 2 . 273 * 1 .3716 .3900

.*000 .6692 .*575 .3171 2.3635 1 . 409 * .*000

.*100 . 727* .*820 .3376 2.4576 1 . 4*87 .4100

.*200 .7903 .5269 .3593 2.5561 1.4895 .4200

.*300 . 658* . 568 * . 3 e 22 2.6592 1.5319 .4300

.**00 .9321 . 61 C 7 .* 06 * 2 . 767 * 1.5759 . 4*00

.*500 1.0120 .6560 .*321 2.6809 1.6216 .*500

.*600 1.0966 . 70*5 .*593 3 . 000 C 1.6692 .4600

.*700 1.1925 .7566 .*682 3.1251 1.7186 .4700

.*800 1 . 29 ** . 812 * .5187 3.2566 1.7700 .4800

,*900 1 .

05.6723 .5512 3.3950 1.8236 .4900

.5000 1.5255 .9367 .5857 3 . 5*06 1 . 879 * .5000

.5100 1.8565 1 .0060 .6223 3 . 69*0 1 .9376 .5100

.5200 1.7991 1 . 08 C 5 .6613 3.8557 1.9985 .5200

.5300 1 . 9 * *6 1.1608 .7027 4 . C 263 2.0621 .5300

. 5*00 2 . 12*2 1 . 2*73 . 7*69 *. 206 * 2.1289 . 5*00

.5500 2 . 309 * 1 . 3*07 . 79*0 *.3969 2.1990 .5500

.5600 2.5119 l .**16 . 8**3 *.5985 2.2729 .5600

.5700 2.7335 1.5508 .6981 *.8121 2.3508 .5700

.5600 2.9763 1.6690 .9556 5.0366 2 . 433 ? .5800

.5900 3 . 2*27 1.7972 1 . 017 * 5.2798 2.5207 .5900

.6000 3 . 535 * 1.9365 1 . OB 37 5 . 536 * 2.6137 .6000

.6100 3.6575 2.0882 1.1551 5 . 61 C 1 2.7130 .6100

• 62 CO *.2125 2.2535 1.2321 6.1026 2.8191 .6200

.6300 4 . 60*6 2 .

3*1 1 . 315 * 6.4160 2.9330 .6300

. 6*00 5 . 036 * 2.6319 1 .*056 6.7526 3.0555 . 6*00

.6500 5 . 519* 2 . 6*91 1.5037 7 . 11*7 3.1677 .6500

• 6*00 6 . 05*1 3.0662 1.6105 7.5055 3.3305 .6600

.6700 6 . 6 * 9 * 3.3520 1 . 727 * 7.9263 3.4853 .6700

.6800 7.3156 3 . 6**2 1.5556 6.3868 3.6533 .6800

.6900 8.0617 3.9686 1.9967 6 . 885* 3.8362 .6900

.7000 8.9005 4.3302 2.1525 9.4269 4 . 035 * -7000

.7100 9 . 6*67 4 . 73*5 2.3253 10.0230 *.2526 .7100

.7200 10.9178 5.1863 2.5175 10 . 67*0 *.*90* .7200 .7300 12 . 13*9 5.6995 2.7322 11.3869 4.7503 -7300

. 7*00 13.5233 6.2777 2.9725 12.1758 5 . 03*8 . 7*00

.7500 15 . 113 * 6 . 93*1 3 . 2*37 13 . 0*37 5 . 3 * 6 * .7500

.7600 16 . 9*21 7.6622 3 . 5*96 14.0027 5.6880 .7600

.7700 19 . 05*2 6 . 536 * 3.6965 15 . 06*2 6.0623 .7700

.7600 21.5039 9.5220 *.2912 16 . 2*10 6.4728 .7600

.7900 2*. 3575 10.6563 *. 7 *i

17 . 5*72 fc . 9??7 -7000

.6000 27.5959 11.9695 5.2560 18.9988 7. *159 .6000

.6100 31.6175 13.*953 5.6513 20 . 613 * 7 . 956 * .6100

.me . 1 L. 2A39 6.5350 22.4106 Mlli

245
.
APPENDIX B - General Research on Corrosion

J. Kruger and R. P. Wei

This appendix contains research summaries presented in the quarterly

progress reports. This research was conducted prior to the emphasis on ANGTS

and, in some cases, is only partially completed.

246

Department of Commerce

National Bureau of Standards

Washington, D. C.

September 26, 1979

1

TO

Attention: Dr. J. Kruger, Chief

Corrosion and Electrodeposition Section

FROM: Dr. R. P. Wei, Consultant

3148-B East Boulevard

Bethlehem, PA 18017

SUBJECT: Test Methods for Evaluating Environmentally Assisted

Cracking Under Stress

REFERENCE: NBS Order No. 809451 (April 6, 1978)

DOT/OPSO Program - DOT-AS-70039

Background and Scope of Work

Transmission pipelines are exposed to a broad range of chemical

environments, both in terms of corrosive species that are present in

soils (such as carbonates, chlorides, nitrates, etc.) and of aggressive

species that are transported within the pipelines (such as hydrogen,

hydrogen sulfide and water/water vapor as impurities, ammonia, etc.).

These environments, acting in concert with operating stresses (both

static and cyclic stresses) and residual stresses, can cause cracks

to initiate and grow, and result in subsequent failure (leakage or

rupture). In addition to these external environments, hydrogen that

might be present in the steel (introduced during pipeline fabrication

and processing , or by corrosion or cathodic charging during service)

can also lead to cracking. Quantitative information is needed, therefore, to assess the safety and reliability of pipelines during service.

As a part of the referenced DOT/OPSO Program, more specifically

as a part of Task III of the Program on Effects of Environmental and

Stress Conditions on Defect Enlargement, the following tasks were

undertaken

247

(1) Assess existing knowledge in terms of test methods for

evaluating environmentally assisted cracking under stress

(stress corrosion cracking and corrosion fatigue) for transmission pipeline steels and weldments.

(2) Based on (1) and preliminary experiments at NBS, recommend

a testing program to be carried out in part by NBS in second

year of DOT project to develop appropriate test methods for

assessing susceptibility of transmission pipeline steels

and weldments to environmentally assisted cracking under

stress.

(3) Collaborate with NBS personnel to identify research needed

in this area to bridge the gap between test methodology

and prediction of service performance, and to ensure a

proper collection of field data.

This report summarizes work carried out under these three tasks.

It is important to recognize that this work represents only an assessment of the technological base for evaluating the effects of environmental and stress conditions on defect enlargement. Such a technological base is necessary, but by and of itself is not sufficient, for ruling

on waiver requests. As such, questions relating to waiver requests

were not included as a part of this work.

To place this work in perspective, it is useful to briefly consider the test methodology within the context of the overall problem

of service life estimation (or prediction) . Service life estimation

begins with a proper structural analysis, which provides quantitative

descriptions of stresses (including residual stresses) at critical

locations and includes stress analyses of cracks that might be present

initially or that might develop during service. Projected service

248

loading and environmental conditions are then defined, with due

consideration for variations in actual service experience. Characterizations of material properties and responses are then made in

light of the projected service loading and environmental conditions.

Finally, the various information is synthesized through suitable

modeling and used in estimating service life. The reliability of

this estimate reflects the uncertainties in the supporting analyses

and the projected service conditions, and depends on the appropriateness and reliability of the material characterization data. It will

depend also on how well the entire estimation procedure has been

supported by prior service experience and by experimental verification . Suitable test methods for evaluating environmentally assisted

cracking of linepipe steels and weldments under stress, therefore,

must properly reflect the anticipated service for the material and

must be evaluated in terms of the intended end use of the test results.

The choice of test methods often reflects engineering judgements regarding the relative importance of crack initiation as opposed to

crack growth in determining service life, and with respect to failure

mechanisms (e.g., stress corrosion cracking vs. corrosion fatigue).

Distinction must be made also between those tests that are designed

for screening purposes and those that can provide information for

quantitative life estimations, and between scientifically and technologically oriented test methods.

Existing Test Methods and Their Assessment

Current test methods that can be used for evaluating environmentally

assisted cracking of linepipe steels and weldments under stress include:

249

1 . "Constant-strain" or "constant-displacement" tests (e.g.,

U-bend test, C-ring test, bent-beam test, etc.) [1-3].

2. Constant-load tests using smooth and notched tensile specimens [4 ]

.

3. Constant strain rate tests [5-7]

a. Smooth tensile specimens.

b. Precracked cantilever specimens.

4. Fracture mechanics based test’s using precracked specimens

(subjected to either constant load or constant crack

opening displacement) [8-10]

5. Fatigue tests using smooth and notched specimens.

6. Fracture mechanics based fatigue tests using precracked

specimens [11,12].

Test methods 1 to 4 are designed for evaluating stress corrosion

cracking response (i.e., environmentally assisted cracking under "static"

loads) , while test methods 5 and 6 pertain to fatigue or cyclically

varying loads.

Test methods 1, 2 and 5 are considered, by now, to be traditional

methods for evaluating stress corrosion cracking and corrosion fatigue

response of materials, and are often described as "initiation" tests.

These tests measure the time to produce fracture of the specimens that

have been loaded in prescribed manners and exposed to the appropriate

chemical environment. Some of the tests (such as the U-bend test) are

suitable primarily for screening purposes. Other tests provide threshold stresses for stress corrosion cracking and for corrosion fatigue.

These stresses are then utilized in conventional design analysis,

principally to define stress levels below which failures are not to be

expected. Although the tests involve crack initiation and growth,

250

crack growth is not explicitly taken into account in the interpretation

or in the use of the data. While these methods may be suitable for

screening purposes, there are fundamental problems associated with

their use in developing information for quantitative service life estimation. Their reliability and utility as a part of a predictive procedure remain to be established.

Development of test method 3 is based principally on the observation by Parkins and his coworkers that a critical strain rate is

associated with stress corrosion cracking susceptibility [5-7]. The

test methods have been used principally as a research tool and mostly

in short-term tests lasting up to one (1) week. In addition, the

interpretation of results from slow bend tests of precracked cantilever

specimens as "average" crack growth rate versus beam deflection rate

must be seriously questioned [5-7]. Specifically, the beam deflection

rate itself is a function of the crack growth rate (among others) , and

cannot be accepted, therefore, as a suitable measure of the driving

force for crack growth. Because of this fundamental objection and

because of the short term nature of the tests, these methods are not

considered to be suitable for use in developing data for quantitative

life estimations.

Methods 4 and 6 are fracture mechanics based test methods that

are of potential value in providing data for quantitative life estimations. A tentative test method for measuring fatigue crack growth

rates has been released by the American Society for Testing and Materials recently [12] . Recommended test methods for stress corrosion

cracking and for corrosion fatigue are under development, but are not

yet available. These methods presume the presence of cracks and,

therefore, treat the growth of these cracks during service. Two

251

separate approaches are currently in use; one involving cracking threshold and the other involving crack growth kinetics [10] . These methods

have been used with some success for high-strength aerospace materials.

Their utility in terms of the lower strength materials (such as linepipe steels) remains to be established. In addition, it has been

recognized that there are a number of nonsteady-state phenomena associated with environmentally assisted cracking, and that these phenomena

can affect the proper interpretation and the utility of experimental

results [10]. For example, erroneously high values for stress corrosion

cracking threshold or fatigue threshold, or unconservative crack growth

rates can be obtained when the nonsteady-state phenomena are not properly

taken into account during testing and data analyses [10] . Further

development work is needed before these test methods can be used to provide reliable quantitative data for life estimation.

A large number of test methods have been developed over the years

and are in use for evaluating environmentally assisted cracking of

materials under stress. Many of these test methods have been designed

for screening or research purposes, and are not suitable, therefore, for

providing data for use in service life estimation procedures. Other

test methods (such as the fracture mechanics based methods) require

further development and verification before their data can be relied

upon for quantitative service life estimation. It was recommended that

further development and verification of test methods be made, with

particular attention to their applicability for the quantitative evaluation of environmentally assisted cracking of linepipe steels and weldments under stress.

Test Method Evaluation and Development

To develop appropriate test methods for assessing the susceptibility

252

of transmission pipeline steels and weldments to environmentally

assisted cracking under stress, it is useful to consider the types

of cracking problems that might be of concern.

(1) Initiation of crack or crack growth from notch-like

and crack-like flaws.

(2) Propagation of cracks.

If the flaws or cracks are internal to the material or weldment,

service environments are not expected to have an effect on crack

initiation or propagation. If, on the other hand, the flaws or

cracks are near to the interior or exterior surfaces of the line-

, pipes, crack initiation and propagation could be adversely affected

by the external service environment. Internal flaws or cracks that

are exposed to "trapped" hydrogen in the steel or weldment are to

be treated as flaws exposed to external aggressive environments.

Test methods can be devised to simulate and evaluate initiation and

propagation from internal flaws (such as porosities, inclusions, etc.)

and cracks by using inert test environments. For external flaws and

cracks, similar test methods, incorporating the appropriate environments, are to be used. Broad categorization of test methods is

indicated in the following table, where notches are to be used to

simulate flaws (other than cracks)

MODE ORIGIN ENVIRONMENT LOADING

Notches

Inert Static/Cyclic

Crack

Aggressive Static/Cyclic

Initiation Inert Static/Cyclic

Cracks

Aggressive Static/Cyclic

Crack Inert Static/Cyclic

Propagation

Cracks

Aggressive Static/Cyclic

253

Specific choices of test methods, to a large extent, depend on

engineering judgement as to the most probable mode of failure. If,

for example, stress corrosion crack initiation is judged to be

important, then tests of "smooth" or "notched" specimens exposed to

aggressive environments under static loads would be selected. Heretofore, stress corrosion cracking has been considered to be an important

problem for linepipe steels and weldments, and most of the current

test methods therefore reflect this judgement.

To assess some of the current test methods (particularly to

reproduce past experience) , and to explore other test methods and

to examine the importance of other cracking modes, a number of explora-

• . t tory experiments have been carried out on a X-65 steel plate . These

experiments included slow strain rate tests on "smooth" specimens

(performed at NBS-Gaithersburg) [5] and fracture mechanics based

fatigue crack propagation tests (performed at Lehigh University) [12]

at room temperature. The slow strain rate test results are summarized

in the following table, and the fatigue crack propagation results are

summarized in Figures 1 to 3.

Spec. No. Environment

Strain

Rate

Stress

at

Failure

(psi) Strain

Reduction

in area

(pet)

1 C0

3

/HCC>

3

1 X 10“ 6 /s 80,500 0.213 75.6%

2 Air 1 X 10~ 6 /s 78,700 0.218 77.4

3 co

3

/hco

3

1.6 X 10“ 7 /s 81,400 0.185 73.5

4 co

3

/hco

3

1.4 X 10‘ 7 /s 81,300 0.204 74.4

5 Distilled 1 X 10“ 6 /s 79,700 0.179 66.0

6 co

3 /hco

3

7.2 X 10“ 6 /s 81,300 0.212 76.1

"^Material received as 5/8 in

.

thick plate.-

254

STRESS INTENSITY FACTOR RANGE, AK (MPa*/m)

Figure 1: Kinetics of Fatigue Crack Growth in X-65 Steel

in Air at Room Temperature.

V'nrrrnvv'.

TRACK

GROWTH

RATE,

da/dN

(in.

/cycle)

STRESS INTENSITY FACTOR RANGE , AK (ksi/TnT)

Figure 2: Influence of Frequency on the Kinetics of Fatigue

Crack Growth in X-65 Steel in Distilled Water at

Room Temperature.

256

FATIGUE

CRACK

GROWTH

RATE,

da/dN

(in

./cycle)

FATIGUE

CRACK

GROWTH

RATE,

da/dN

(cm/cycle)

STRESS INTENSITY FACTOR RANGE, AK (ksi/in.

)

Figure 3: Kinetics of Fatigue Crack Growth in X-65 Steel

Exposed to Different Environments at Room

Temperature (Frequency = 1 Hz)

257

FATIGUE

CRACK

GROWTH

RATE,

da/dN

(in.

/cycle)

Slow strain rate tests were carried out in three environments:

air, distilled water and carbonate/bicarbonate (CO^/HCO^) solution,

under open circuit conditions, at room temperature. These results

indicate that for these environments, no appreciable stress corrosion

cracking susceptibility (low reduction in area) was observed. Work

in the literature [5,7] suggests that specimens polarized in CO^/HCO^

solutions may be more susceptible to stress corrosion cracking failure,

and such conditions should be examined in future tests.

Fatigue crack propagation tests were carried out at room temperature in air (at 1 and 20 Hz), distilled water (at 0.1 and 1 Hz),

carbonate / bicarbonate solution under open circuit conditions

(CO^/HCO^ at 1 Hz) and high-purity hydrogen sulfide (at 1.33 kPa or

0.19 psia, and 1 Hz) [12]. The specimens are oriented in the longtransverse (TL) direction [12]. Sinusoidally varying loads, at a

load ratio (ratio of minimum to maximum load in a given cycle) of

0.05, were used. Data in air (Figure 1) show little effect of cyclic

load frequency, and are used as reference data. Results in distilled

water (Figure 2) indicate an effect of this environment on fatigue

crack growth. These data are consistent with the response for other

low alloy steels [13-15] , and reflect embrittlement by hydrogen that

is produced by the water-steel surface reactions [15,16]. The observed frequency dependence is consistent with surface reaction control of this embrittlement [15,16]. Crack growth data in all four

environments, at 1 Hz, are shown in Figure 3. These results indicate

severe embrittlement by the low pressure (1.33 kPa) hydrogen sulfide

[19,20], and an inhibiting effect of CO^/HCO^- under open circuit

conditions. The stress intensity (K) levels for environmentally

assisted fatigue crack growth are believed to be well below the

258

threshold K level for stress corrosion cracking [13] . Taken in

conjunction with the recognition that stress fluctuations do occur

in pipelines during normal operation [17,18], environmentally assisted

fatigue crack initiation and propagation must be considered as a

significant potential cause for failures of transmission linepipes.

Based on these and other results, it is recognized that steels

that appear to be immune to stress corrosion cracking (static load)

may be highly susceptible to environmentally assisted fatigue crack

growth (corrosion fatigue) . The discrepancy between the apparently

low stresses associated with field failures of linepipes and the

high stresses required to induce stress corrosion failure in the

laboratory [5-7] suggests that corrosion fatigue should not and

could not be ruled out as a possible cause for failure in service.

As such, environmentally assisted crack initiation and propagation,

under both static and cyclic loading conditions, must be considered

and assessed. A testing program to incorporate both loading conditions is proposed.

Proposed Test Program

To further evaluate test methodology for assessing environmentally

assisted crack initiation and propagation, the following tests should

be considered:

o Slow strain rate tests (monotonic loading) of smooth and

notched specimens to assess conditions for stress corrosion

crack initiation.

o Slow strain rate cycling (fatigue) tests of smooth and notched

specimens to assess conditions for corrosion fatigue crack

initiation

.

Static load (or wedge force load) tests of precracked (fracture

mechanics type) specimens to determine stress corrosion

o

cracking threshold K ( K

j scc

) f° r appropriate environmental

conditions, and to determine the kinetics of crack growth

under these conditions.

o Cyclic load (fatigue) tests of precracked (fracture mechanics

type) specimens to determine the threshold AK for fatigue

crack growth for appropriate environmental conditions, and

to determine the kinetics of crack growth under these conditions. A range of loading variables should be considered;

these include frequency, load ratio, hold-time, waveform, etc.

In determining crack growth threshold and crack growth kinetics, careful distinction should be made between steady-state and nonsteadystate crack growth [10,12], The impact of nonsteady-state growth on

the measurement of rates and threshold K levels must be recognized

[ 10 , 12 ].

To complement data that have already been developed under the

preliminary testing program, the following series of tests are

recommended to complete test method evaluation under the referenced

program

:

(1) Cyclic loading of smooth and notched specimens to assess

conditions for crack initiations in fatigue, and to determine the importance of this failure mode in relation to

the performance of linepipe steels and weldments.

(2) Laboratory evaluation of test methods for determining K lscc

and of conditions for the initiation of stress corrosion

crack growth. Modified wedge-opening-load (decreasing

K) specimens are to be used. Two. initial K levels (30

and 45 ksi/in. or 33 and 49.5 MPa/in) are suggested for

examination. Tests in the polarized and unpolarized con260
i
ditions should be considered.

(3) To assess the aggressiveness of field conditions, and

to establish potential correlation between laboratory

and field data, in-ground tests of modified wedge-openingload specimens (see item (2) above) at NBS test sites are

recommended. Initial K levels that conform to those used

in the laboratory tests are suggested. A minimum of two

different test sites should- be considered.

Summary

A review has been made of currently available test methods for

assessing the susceptibility of transmission pipeline steels and

weldments to environmentally assisted cracking under stress. Preliminary experiments have also been carried out to reproduce past

experiences with some of the test methods, and to explore other

test methods and potential failure modes. ' It is recognized that

the validity and usefulness of any test method depend on its ability

to simulate service conditions and environments, and to reflect the

mode of failure. It is believed that, at this time, no test method

satisfies this requirement in relation to service performance of

linepipe steels and weldments. Even if such methods exist, data

from the tests would represent only a part of the necessary information in considering waiver requests, but would not be sufficient.

The usefulness of this type of data must be established and documented

on the basis of careful correlation with service performance.

This study has also shown that corrosion fatigue (environmentally

assisted fatigue crack initiation and propagation) is a potentially

important failure mode in linepipe steels and weldments. The ability

to initiate and propagate cracks at relatively low stress levels in

261

fatigue suggests that corrosion fatigue might provide a better

explanation for some of the service failures. It is recommended

that recognition be given to the potential for corrosion fatigue

induced failure of linepipe steels and weldments in service, and

that recordings of cyclic loads should be included in the collection

of field data.

References

1. ASTM G 30 - 72, Standard Recommended Practice for Making and

Using U-Bend Stress Corrosion Test Specimens, Annual Book of

ASTM Standards , Part 10, 1975, pp. 703-711.

2. ASTM G 38 - 73, Standard Recommended Practices for Making and

Using the C-Ring Stress-Corrosion Cracking Test Specimen,

Annual Book of ASTM Standards , Part 10, 1975, pp. 751-759.

3. ASTM G 39 - 73, Standard Method for Preparation and Use of

Bent-Beam Stress-Corrosion Specimens, Annual Book of ASTM

Standards , Part 10, 1975, pp. 760-769.

4. Craig, H. L., Sprowls, D. 0., and Piper, D. E., "Stress-Corrosion

Cracking", Handbook of Corrosion Testing and Evaluation , W. H.

Ailor, Editor, Wiley, 1971, pp. 231-290.

5. Parkins, R. N. , "The Controlling Parameters in Stress-Corrosion

Cracking", 5th Symp. Line Pipe Research, American Gas Association,

Cat. No. L30174, 1974, Ul-40.

6. Berry, W. E. , "Stress Corrosion Cracking Laboratory Experiments",

5th Symp. Line Pipe Research, American Gas Association, Cat. No.

L30174 , 1974, pp. Vl-43.

7. Parkins, R. N. and Greenwell, B. S., "The Interface Between Corrosion

Fatigue and Stress-Corrosion Cracking", August/Sept. Metal Science,

1977, pp. 405-413.

8. Brown, B. F. and Beachem, C. D. , Corrosion Science, Vol. 5, 1965,

pp. 745-750.

9. Novak, S. R. and Rolfe, S. T. , "Modified WOL Specimen for Kj scc

Environmental Testing", Journal of Materials, JMLSA, Vol. 4, No.

3, 1969, pp. 701-728.

10.

Wei, R. P. , Novak, S. R. , and Williams, D. P., "Some Important

Considerations in the Development of Stress Corrosion Cracking

Test Methods", Materials Research and Standards, MTRSA, Vol. 12,

No. 9, 1972, p. 25.

262

11. Wei, R. P., "Fracture Mechanics Approach to Fatigue Analysis

in Design", JEMT , Vol. 100, 1978, pp. 113-120.

12 . ASTM E647-78T, "Tentative Test Method for Constant-Load-Amplitude

Fatigue Crack Growth Rates Above 10“® m/cycle". Annual Book of

ASTM Standards , Vol. 10, 1978.

13 . Barsom, J. M. , Corrosion Fatigue , ed. 0. Devereux, A. J. McEvily

and R. W. Staehle, NACE , Houston, Texas, p. 424, 1972.

14 . Gallagher, J. P. and Wei, R. P., Corrosion Fatigue , ed. O. Devereux,

A. J. McEvily and R. W. Staehle, NACE, Houston, Texas, p. 409,

1972 . 15 . Pao, P. S., Wei, W. and Wei, R. P. , Environment-Sensitive Fracture

of Engineering Materials , ed. Z. A. Foroulis, TMS-AIME, New York,

New York, pp. 565-580, 1979.

16. Simmons, G. W. , Pao, P. S. and Wei, R. P., Met. Trans. A, Vol.

9A, August 1978, pp. 1147-1158.

17 . Vosikovsky, 0. and Cooke, R. J. , Int. J. Pres. Ves. & Piping,

Vol. 6, 1978, pp. 113-129.

18 . "Environmentally Induced Cracking of Natural Gas and Liquid

Pipelines", Vols. 1 and 2, Final Report, Contract #DOT-OS-60519 , ASL Engineering, Inc., 495 South Fairview Ave . , Goleta, CA, 33017,

Dec., 1977.

19. Brazill, R. , Simmons, G. W. and Wei, R. P. , J. of Engr. Matls.

& Tech., Vol. 101, July 1979, pp. 199-204.

20. Nelson, H. G. , Effect of Hydrogen on Behavior of Materials , TMSAIME, A. W. Thompson and I. M. Bernstein, eds., 1976, pp. 602-611.

263
”
Subtask 2: Environmental Assisted Cracking Under Stress

Principal Investigator: Dr. Jerome Kruger, (301) S21-2094

Progress : This is the final report on this subtask (formerly Task 3A).

Objectives and Approach

The objectives and approach of subtask 2 were as follows:

(1) Assess existing knowledge in terms of test methods for evaluating

environmentally assisted cracking under stress (stress corrosion

cracking and corrosion fatigue) for transmission linepipe steels

and weldments.

(2) Based on (1) and preliminary experiments at NBS, recommend a testing program to develop appropriate test methods for assessing sus- ceptibility of transmission linepipe steels and weldments to

environmentally assisted cracking under stress.

(3) Identify research needed in this area to bridge the gaD between

test method mcgy and prediction of service performance, and to

ensure a proper collection of field data.

To carry out the assessment of existing knowledge and to participate

in objectives (2) and (3), Prof. R. R. Wei of Lehigh University was engaged as a consultant. Prof. Wei produced preliminary assessment in

1978 and submitted a final report in October 1979. His final report,

which is attached as Appendix A, contained, besides an assessment of

the problem, recommendations for some measurements to be carried cut

at NBS. It also contained the results of corrosion fatigue measurements

carried out at Lehigh.

Results

The following test results were obtained from the preliminary

experiments carried out at NBS and Lehigh:

(1) Slow Strain Pate Tests (N BS) - Tests have been carried out in three

environments, air, distilled water and carbonate/bicarbonate (CO^/HCO^

solutions. * The following table gives the results found for pipe steel

at room temperatures:

264

Spec. No. Environment —

Strain

Rate

Stress

(psi)

at

failure

Strain %

Reduction

in Area

1 co

3/hco3

1 x 10

"6

/s 80,500 *.213 75.6% —2 Air 1 x l0“6

/s 78,700 .218 77.

3- -co

3

/hc°

3

1.6 x 10~7/s 81,400 .185 73.5

-7 ,

. 4 co

3

/hco

3

1.4 x 10 /s 81,300 .204 74.4

5 Distilled H^O 1 x 10

6

/s 79,700 .179 66.0

6 CVHC0

3

7.2 x 10

6

/s 81,300 .212 76.1

These results indicated that for these environments no appreciable stress

corrosion susceptibility (low reduction in area) was observed.

(2) Fracture Mec hanics Fatigue Tests (Lehigh) - Tests on precracked pipe

steel spec inen showed a three-told enhancement of crack growth rate in

distilled water at 1 Hz.

Based on these results and recommendations of the second Wei report,

tests were carried out that produced the following results:

(1 ) Tapered Pin Loaded Fracture Mechanics Specimens

One specimen loaded to 49.5 MPa /m (45 ksi /in) and immersed in 1 N

CO^/HCO^ solution at room temperature. Solution was changed every

month. ^Specimen when broken apart revealed no crack growth. One

specimen loaded to 49.5 MPa /m (45 ksi /in) polarized at 650 mV SCE

in 1 N CO^/’iCO^ . Specimens when broken apart revealed no crack

growth.

<

265

These results and the results from the preliminary tests all indicate

no failures can be detected for static fracture mechanics tests or static and cyclic slow strain rate tests. Only cyclic fracture mechanics

testing as described in Appendix A appears to reveal the possibility of

susceptibility tc failure. It is this sort of testing regime that is

most closely tied to actual pipeline experience.

In addition to the above tests some field tests were initiated in

late 1979. These involved the burial of tapered pin loaded fracture

mechanics specimens at two test sites, one at MBS in a soil, that is

not particularly corrosive and one at Patuxent, Md in a soil that is

highly corrosive and contains sulfides. The specimens buried at these

sites were as follows:

(1) Three specimens buried at NBS site (12/4/79)

Two loaded to 49.5 MPa /m (45 ksi /in)

One loaded to 33.0 MPa /m (30 ksi /in)

(2) Three specimens buried at Patuxent site (12/13/79)

Two loaded to 49.5 MPa /m (45 ksi /in)

One loaded to 33.0 MPa /m (30 ksi /in)

It is planned to remove these specimens in one year after burial to

determine if any crack growth has occured.

Conclusions

Using the above described test results and the assessment by Prof. Wei,

the following conclusions can be drawn:

(1) No test method currently available is capable of simulating service

conditions and environments to allow a reliable assessment of the

effect of environment on the failure by cracking of linepipe steels

and weldmerts.

(2) If such methods were to be developed, data from these tests would

represent only part of the necessary information in considering waiver

requests, but would not be sufficient without establishing a correlation

between this test data and service performance.

(3) Corrosion fatigue (cyclic stressing) is ? potentially important

failure mode in linepipe steels and weldments. Therefore, recordings

of cyclic loads should be included in the collection of field data.

(4) Greater attention should be paid to the role of cyclic stressing

in the environmentally assisted failure (cracking) of linepipe

steel weldments. More research is needed in this area.

256

APPENDIX C - Radiographic Evaluation of the Significance of Weld Defects

M. B. Kasen

The primary questions that must be addressed In any fitness-forpurpose assessment of weld quality are: What flaws are significant in a

given structure? At what dimensions do such flaws become critical,

endangering the integrity of the structure? How good is the chosen NDE

method for detecting and sizing such flaws?

Other than visual inspection, radiography is the most commonly used

method for examining the quality of a weld. While undoubtedly effective

in judging the quality of a weldors performance, it is necessary to

question if radiography is equally effective in assessing the quality of

welds on a fitness-for-purpose basis. Table I presents a breakdown of

the types of weld flaws detected by radiography under field conditions.

Although illustrated for a pipeline, the relative distribution of flaw

types is probably typical of most manual welding operations. Most

significant is the eight-to-one preponderance of blunt or three dimensional flaws. This will surprise no one familiar with radiography,

because the ability of radiography to reveal a flaw is proportional to

the difference in radiographic absorption of the flaw as compared to

that of the surrounding metal. This difference is much higher for

pores, inclusions and arc strikes than for sharp, two dimensional flaws

such as cracks or lack of fusion. Furthermore, the detection of blunt

flaws is Independent of the orientation of the radiographic source

relative to the weld, while the probability of detecting a sharp flaw is

267

highly orientation dependent. It Is easily demonstrated that a through

crack along the fusion line of a manual weldment made with the usual 60°

Included angle joint preparation has a high probability of being missed

by the Interpreter under the best of radiographic conditions. It can

therefore be concluded that the series of welds surveyed on Table I

contain a much higher percentage of sharp, cracklike flaws than was

actually observed.

This creates two problems. The most important is that many cracklike flaws, which have the highest probability of initiating weldment

failure, go undetected. The second is that a substantial amount of

money is spent to repair blunt flaws which may have a very low probability of initiating failure.

The fitness-for-purpose approach to assessing flaw significance in

a given structure has been described in other papers in this series.

One of the initial applications of this approach was to weldments of the

Alyeska oil pipeline. Here it was necessary to assume that all flaws

detected by radiography were significant as crack initiators and that

those exceeding existing workmanship criteria required dimensioning and

assessing by fracture mechanics criteria. This approach was nonconservative in assuming that the radiographs were revealing all the significant flaws and very likely over conservative in having to assume that

blunt flaws were the equivalent of sharp flaws of equal dimension. This

reflected the difficulty of assessing stress concentrations for blunt

flaws. Furthermore, difficulty in assessing the through-wall dimensions

of blunt flaws from the radiographs necessitated the assumption that

such flaws were twice the estimated size.

268

*

CO CM• CO

00 o— o %E^ 2

Cl flj I 5 © © o c © •— X ©

uj

a » O.

£.£0) H TJ

*©© to E s

CQ

O

© 5aE

co CO

a**o # H O• CO• in• o• o• CM• o N CO o CM CO

IP CM CO

-§5a >w E ® = 2 2z -0o

CO rO

CM *

CO$

CO LL

COcaK oQw©C©©aEoo o©©aEoo ©

.Xow

in

co o

G>

©£©

©© 2? x.ooa ©ow r oa

BE

O +*c

jo

CO

©

© o © 0-0

3-°

© > £ © © = © o o o k £ © oa

©©£© 2 2

to <

h- in

CO

Q

CM t- o*aEoH

269

Total

Welds

The current fracture mechanics approach attempts to provide the

basis for a more realistic method. It is assumed that radiography will

very likely remain the primary NDE technique, but it also assumes that

other techniques such as ultrasonics will provide a backup method for

flaw sizing and may eventually supplant radiography as the primary inspection method. The current approach also continues to assume that

two-dimensional, planar flaws are significant in all structures, requiring dimensioning and analysis by analytical fracture mechanics

techniques. But the assumption is also made that blunt flaws may be

innocuous in some structures, depending on the stresses and toughness of

the weld metal, allowing permissible flaw content to be established on

the basis of technical engineering judgment rather than by techniques

requiring precise dimensioning. The advantage of this approach is

significant cost saving if repair of noncritical flaws can be minimized

and a significant increase in structural integrity by emphasizing the

detection of potentially dangerous flaws.

The approach to the blunt flaw portion of this program has been reviewed in the literature.^ In brief, it is assumed that flaws which can

be shown to have a low probability of initiating cracks in welds subjected to stress regimes far more severe than those in service can be

rationally judged to be innocuous under service conditions. The specific approach has been modelled after similar research by the British

Welding Institute in a study of flaw significance in welded pressure

2

vessels. Weldments representing the parent materials, welding processes, procedures and consumables used in the actual structure and

containing much higher levels of porosity, slag and arc strikes than

270

would rationally be permitted under field conditions, are subjected to

fully-reversed, low-cycle strain-controlled fatigue at strain levels

slightly above yield of the weld metal and twice the strain at yield.

Care is also taken to perform the testing at credible field temperatures. Although weldments in pressure vessels or in the pipelines of

current concern are not subjected to service stresses and strains of

this magnitude, any flaw which can be shown to be innocuous under these

severe conditions will have such a low probability of initiating flaws

under essentially static conditions as to justify limiting their content

on pragmatic reasons without incurring the complexities of the analyti3

cal fracture mechanics approach.

In the present work, a comparison is made between the number of

cycles to crack initiation in pipeline girth welds made in AP5LX-65

line pipe by manual (SMA) cellulosic electrodes and by an automatic process. In addition to sound welds meeting the API 1104 workmanship

criteria, very badly flawed welds were prepared containing porosity

levels giving 1 5;'„ radiographic obscuration, continuous slag and arc

strikes imposed after the welding operation. Specimens containing a

4 inch length of weldment were fatigued in a transverse manner at

strains slightly over yield (± 0.22%) and at twice yield (± 0.45%) while

maintaining the temperature at -2 ± 1 °C. Crack initiation was monitored by observing changes in the load required to maintain the strain

range. A 10% drop in load, representing development of a 10% cracked

area, was taken as the criterion for crack initiation. Tests were

continued until a 40% load drop (40% cracked area), was achieved so as

to provide information on crack propagation. Specimens were th:-n broken

in tension for fractographic analysis.

271

The results obtained from 37 specimens representing parent metal as

well as sound and flawed weldments, tested with the reinforcement cap

both intact and machined off, appear In Figs. 1 and 2. From Fig. 1, it

is seen that approximately 800 cyclic reversals were required for crack

development when the weld cap was intact, regardless of whether the

welds were sound or contained very large amounts of porosity, slag or

arc strikes. Even for strain reversals equal to twice yield, about 30

cyclic reversals were required. Removing the weld cap forces failure

through the flawed welds, decreasing the fatigue life to initiation.

But even here, about 150 cyclic reversals were required at just over

yield and about 25 at twice yield for crack initiation.

The data plotted on Fig. 2 relate more closely to field conditions,

as the weld reinforcement cap is intact in all cases illustrated. This

figure also examines the effect of specific flaw types. The performance

of the unwelded, parent material is also illustrated.

The top plot illustrates that the automatic welds containing 15%

porosity by projected area (about 3% by volume) performed equally well

as did welds free of porosity or other flaws. The center plot illustrates that the presence of slag somewhat decreases the cyclic life to

crack initiation; nevertheless, about 750 reversals were required at

just above the yield and about 30 at twice yield to achieve initiation.

The lower plot illustrates that arc strikes had no effect on cyclic

life. In no case were weld failures observed to initiate from such

flaws. Arc striking the parent plate was likewise ineffective in initiating fracture at the ± 0.22% strain level. It was necessary to

strain the arc struck parent material at ± 0.45% before any effect was

noted.

272

o*

oJUm*

••

O

o

in CO *r

o o o o o• o 6 o

• o

oC c

O C cn

4-> to "O Q

J- f— t- in O cn +* c 2 O cn z to f- a:

1/1 (3 l. **— fc. m S- l- O U Q 2 C +J 4- P -D O

> cn cn c

CJ "C 3

1- 4- c o P o <o X- o

< U <0 3

*r- c cn Ci

u r— O *C "3 •*“ O •«- *— C43 2 >»•*> C re Q O U 2 4- __ C cr . *

o O |

4— 3 SI rc *3

O 4- 4-> «— O

i- cn cn <°

CC -o O o

t rex- t_

CJ C • O

jD cn >> >

E <o c •-* a • j 3 c *<- t. L_) O #

C c •<- cn • o

o «— o >, H o o •*- o *- —+• jt-mc-Oi— CSJ </) W •4-» <0 •»“ C- 3 1 IU

f- C. 4-

4- 4-> C 4-> aJ O f- "3 C C <o O C C

>»•»- = C r— >• l- O *»" "3 O

o ie^ vt n c S- E u •*->+-> •»-> E o J- c cn c

3 s- o o a o

to O 4- o u

»v ‘aoNva Nivais ivioi

273

to

o

o

o

o

o

o

K

<

Z

O

<

oco

o

I-co

III .J

O

>-o

4-> O!E

c "o

03S U

re fc- i~ O

U 4- C O)

4-> CTc

to re •— S- re

to c

i- -r- C TO > iCJ +J

l- 10 U '4- •*- O

o c

>> o

O -r4-> 4- <_) o e

3

fc. 4-

C’ £5 O

E3 to z re

CM

Cl

03 •oo

u

"3

O CL

t- Mre jc

CL lO E c

o re u E

to *- re O

S

to

03C-O

>»•—

a

u

c

reE

s_o

4-

J-o

O- o

c3

o

X>

t-o

re o.a

u x:

•i- cr. u cr a -r- CL4-»

tO C3o

cn E

c

t- to c "O

re a

*j S

c

o o

U 4-»

a

x:

4->

tfl a

c

£to*

O'C

a r—a

03 5

C

t- ~c — c3

•c o

**- to O 4-

40 O

*3 V' 3 DNVU NIVU 1S 1 VJ.01

274

These data add credibility to the argument that blunt flaws have a

very low probability of initiating failure in an operating pipeline,

particularly when it considered that the permissible flaw level would be

limited in practice to substantially less than in the test series and

that line is unlikely to see more than one or two deformations to strains

above yield and that these would not be accompanied by reversal. The

strongest case can be made for automatic porosity and manual arc strikes.

Slag presents a slightly more complex problem because fractographic evidence shows that failures with this type flaw were influenced by the development of hydrogen-assisted cracking originating either from the slag

inclusions or from micropores within the weld metal. It therefore

appears that, to the extent that slag contributes to crack initiation,

it does so by providing a surface for hydrogen segregation rather than

from the stress concentration due to slag shape.

Returning to the original questions, available data suggest that

porosity in automatic welds and arc strikes in manual welds are not significant in weld metal having the toughness of that studied. The probability of failure from slag also appears remote; however, the associated hydrogen cracking arising from the use of cellulosic electrodes

needs further analysis. These results suggest that it is probably

unnecessary to dimension such flaws in order to establish a valid fitnessfor-purpose criteria for their content in a weld. Finally, we conclude

that radiography is effective in detecting the presence of such flaws,

but is of very questionable effectiveness in the detecting or dimensioning of two-dimensional, planar flaws of the type that have a much

higher probability of initiating weld failure. It remains for other NDE

systems such as ultrasonics to perform that task.

275

References

1. Kasen, M. B. and Mikesell, R. P., "Interim Report on the Significance of Blunt Flaws in Pipeline Girth Welds", Pipeline and Energy

Plant Piping: Design and Technology . Pergamon Press, 1980, 329.

2. Boulton, C. F., "Acceptance Levels of Weld Defects for Fatigue Service", Welding J ., 56, 1977, 13-22.

3. Gurney, T. R., Fatigue of Welded Structures, 2nd Ed ., Cambridge

Univ. Press, Cambridge, England, 1979.

276

Appendixes B and C of this document are available on request.

277

APPENDIX D - Defect Characterization and Dimensioning of Cracks

in Welds by the Ultrasonic Diffraction Method

S. Golan*

ABSTRACT

The possibility of applying an ultrasonic diffraction method for

dimensioning of crack-like defects in welds was investigated. A feasibility study was carried out and optimum test conditions were established

using a series of test specimens with narrow slits. A series of welded

specimens with in-welded cracks were tested. The possibility of using a quantitative diffraction technique for nondestructive examination of

pipeline welds in field conditions is discussed.

INTRODUCTION

With the conventional ultrasonic pulse-echo technique, flaw size is

related to echo-amplitude. Amplitude, however, being strongly affected

by many factors 1

, is rather insensitive to size. Consequently, the pulseecho technique is not suitable for quantitative measurements. Ir welds,

quantitative evaluation becomes even more complicated by the anisotropy

and inhomogeneity inherent in welded materials and by geometric factors.

The inability of the pulse-echo technique, which is based on reflectivity,

to serve as a quantitative tool provided the impetus for development of

techniques based on other phenomena of interaction between the flaw and

the acoustic field, such as scattering, diffraction, interference, etc.

It seems that the most promising technique to develop into a practical

and economical quantitative testing tool for field work is the diffraction

technique. 1-6 Figure 1 shows an example of ultrasonic sizing of slits.

More such curves are given in reference 1. It is especially useful for

dimensioning of crack-like defects. This technique has three main advantages over some other emerging techniques: flaw size is computed using

only time measurements, which can be made more accurate and more reliable

than amplitude measurements; conventional pulse-echo equipment can be used;

and operators familiar with the pulse-echo technique can easily adopt the

diffraction technique.

* Permanent address, Israel Institute of Metals, Israel. This report was prepared while the author was on a temporary assignment to the NBS Office

of Nondestructive Evaluation.

278

BASICS OF THE DIFFRACTION TECHNIQUE

When an ultrasonic wave impinges on the tip of a crack, waves of

shear (s) and compress ional (P) modes are scattered from the tip in a circular form. Figure 2 shows a photoelastic pattern of the scattered

acoustic field from the tip of a narrow slit in a glass block. From the

time interval between the entering of the incident wave and emerging of

the tip-diffracted wave the depth of the tip can be computed. 1 ' Most

reliable results are obtained when time delay is measured from a conve- niently chosen reference signal. Figure 3 shows a diffraction triangle

formed by the transmitted, the diffracted and a compressional surface

wave which serves as reference. 1 Equation 1 gives the relationship between time delay and depth of the tip when a P surface wave is used as

a reference.

* — =

C

2

-C

2 Y2 + 4 COSa C-j (a+1 ) - C

2 (sina) Y

2

-4 COS a (A+l)

2

-C

2

C rC

2

C = 1

c Av

s

= 0 (

1 )

Angle of incident beam

- Mode coefficients for impinging and diffracted beams

- For a compressional beam

- For a shear beam

- Shear wave velocity

- Compressional wave velocity

A

= i

26

2a

y

- Depth of source of diffraction; nondimensional

- Depth of source of diffraction; nondimensional

- Time delay expressed in units of length

- Distance between transducers

- Depth of source of diffraction.

From this general equation specific equations can be derived for

different modes of transmitted and diffracted waves and for various

diffraction triangles. 1

* 2 If the reference signal is a bottom reflection

and the diffracted and reference signals are maximized (see Figure 4),

the relationship between time-delay and depth for a single or double

transducer technique is given in Equation 2.

y

= T - 6cosa ( 2 )

T - thickness of material.

OPTIMUM CONDITIONS FOR DIMENSIONING OF CRACKS

The diffraction techniques can be used to dimension cracks open to the

front or back surfaces as well as internal cracks, provided a tip-diffracted

signal can be detected. This, however, is quite often very difficult. The

tip-diffracted wave usually has a low amplitude, and high amplification is

necessary to reveal the signal. This, in turn, forms a noisy background which

might blur the tip signal. The noise signals come from various sources;

scattering from grains and inclusions, mode conversion, scattering from the

crack surface, back wall reflection, surface waves, side loops, internal re- flections from the wedge, and electronic noise. Figure 5 shows a typical

oscil loscopic trace in the background of a signal diffracted from the tip of

a fatigue crack. 5 The crack depth can be sized with sufficient accuracy and

reliability only if the tip diffracted signal can be clearly detected and

unambiguously identified among the spurious signals. As has been shown in

this and some other works 1-5 under certain conditions signal-to-noise ratio can be considerably improved and the tip signal can be detected. By analyzing

the oscilloscopic trace, the tip signal can be in many cases identified. 5 ’ 6

Three parameters determine the detectability of a signal; amplitude,

resolution, and dynamic response. Dynamic response is defined as the drop

in amplitude caused by displacing the transducer from its maximum amplitude

position (see Figure 6). When the dynamic response is high, a sharp drop

in amplitude will be observed when the transducer is slightly displaced.

This parameter is a very strong crack indicator.

All three parameters are mode and angle dependent. On a goniometer

the angle-amplitude dependence was determined for the PP mode (incident

and diffracted waves of the compressional mode) 1

. It was found that maximum amplitude is obtained when the angle between the transducers approaches

180° (a+90°).

In former works 1-4 ’ 6 the PP mode was primarily used in order to prevent

interfering signals of the P mode. In later work6 it was found that, by

using the SS mode and angles a in the range of 40° to 50°, higher amplitudes

are obtained. The dynamic response is higher6 because of the smaller angles

and, because of the lower velocity of shear waves, resolution is almost

doubled. By choosing an angle between 40° and 50° interfering compressional

mode signals do not appear. An additional improvement in amplitude is

achieved by using one instead of two transducers. This has additional advantages: it is more convenient to handle one transducer; no alignment

between the transducers is necessary; and the dead zone is smaller, as it

is not limited by wedge geometry. 6 The high amplitude noise which appears

when one transducer is used can be minimized by chosing a wedge which filters

out the internal reflections. 7

230

SIZING OF SIMULATED CRACKS IN WELDS

For the feasibility study a series of specimens with a simulated internal

crack in welds was prepared. The "inwelded crack" was formed by incomplete

penetration of a butt weld. The location, size and orientation of the "crack"

was controlled by the bevel geometry of the welded plates (see Figure 7, 8 and

9). In order to achieve minimum crack width (0.00 to 0.06 mm), minimum deformation and, as closely as possible, a straight line of the tip, the plates

were welded and cooled to room temperature while clamped in a fixture. The

material used was 25.4-mm low-carbon steel plates. Specimens of 22-mm width

were cut from the welded plates. The specimens were machined, polished and

etched to reveal the "crack" and the weld cross section. In order to check the

ability of the diffraction method to dimension normal as well as inclined internal cracks in the as-welded condition, four types of specimens were prepared

as-welded with normal "cracks"; machined to remove the excess of the welding,

with normal "cracks"; and as-welded and machined with 45° "cracks" (see

Figures 8 and 9). Measurements were made with the two- transducer technique,

45° 5-MHz shear waves, equilateral diffraction triangle, and a back-wall

reflection signal as reference. (The 45° angle is within the range of good

detectability). Because of the many advantages of the one-transducer technique

an initial attempt was made to employ it. This technique was, however, abandoned because of high-amplitude interfering signals reflected from the "crack"

surface (see discussion below).

The setup and the transducer arrangement are shown in Figures 10 and 11.

Measurements were taken after the amplitude was maximized by slight lateral

movements of the transducers. Results were verified by measurements from four

different transducer locations, with direct and double-skip beams (see

Figure 12), and from both sides of the specimen. In Table 1 a schematic presentation of the different beam paths and direct and ultrasonic measurement

results are given in x,y coordinates. The coordinate system is shown in

Figure 8. The x coordinate was determined from the positions of the transducers. The scatter of the x measurements was within ±2mm. The x data are arithmetic averages of five measurements. In Fiqure 13 values of depth (y

coordinate) made by direct and by ultrasonic measurements are compared. The

differences are not necessarily deviations but, perhaps, true differences in

the depth of the tip at the outer surface (measured directly) and at the

middle of the specimen (measured ul trasonically)

.

DISCUSSION

Although it was shown in several works 1-5 that accurate and reliable

sizing of cracks with the diffraction technique is feasible, testing of welds

in materials less than 25 mm in thickness under field conditions presents

special difficulties which have not yet been compl etely . resol ved. This can

be seen in Table 1 where, in many places, measurements were not possible.

Some of the difficulties are described below and schematically presented in

Figure 14. Because of the coarser grains, the directional structure and the

microimperfections usually present in weld material, the ultrasonic noise from

the weld is more severe than in the parent material. This might be improved

by using a lower frequency.

281

The advantage of using the one- transducer technique is obvious. This

technique, however, might cause misleading results because of multiple re- flections from the crack surface. Because of the crown of the weld, scanning

of the weld with a direct beam is limited and often not possible and doubleskip beam paths must be employed. This increases noise and attenuation.

The high-amplitude back-wall reflection often masks the low-amplitude

tip-diffracted signal and resolution becomes difficult and often impossible,

especially in thinner material.

Some of the difficulties mentioned above might be diminished by changing

testing parameters such as frequency, diffraction triangle, mode combination,

transducer and wedges. Further investigation to diminish interfering factors

is recommended.

282

Table 1

Ultrasonic Measurements

I B,C

I

1 ’ 2

3

W

N

As welded specimens

Machined specimens

Normal "cracks"

45° "cracks"

Weld imperfection

Measurement not possible

(explained in text)

Specimen

Number

Coordinate Di rect

Measurement

Ultrasonic Measurements (mm)

(mm)

I II III IV Average

1A y

X

9.1

0.5 N N 9.0

1.0

N

9

1.0

y

X 9.4

0

95

0

N N N 9.5

0

y

X

20.2

0

N N N 20.0

0.5

20.0

0.5

w-y

X

7.0

-2.5 N 7.0

1.0

N 7.0

-1.0

7.0

0

fIB y

X 9.4

0

N 10.5

-1.0

9.5

-1.5 N 10.0

-1.2

y

X 20.1

0

20.8

-1.5

20.4

0

N 20.5

0

CO

CO O

O

CM1

283

Table 1

Ultrasonic Measurements

(Continued)

Specimen

Number

Coordinate Direct

Measurement

(mm)

Ultrasonic Measurements

I II III IV

(mm)

Average

3A y

X

11.1

0

N N N

11.0

0

11.0

0

y 19.6 M 19.4 K| 20.0 19.7

X -8.4 n -8.5 IN -8.0 -8.3

w-y Several points from weld imperfections

X

3C y 11.4 10.6 10.7 11.5 10.7 10.9

X 0 0 1.5 1.0 0 0.6

y 18.8 18.7 N 18.5 18.6

X 18.4 8.0

in IN 8.0 8.0

284

Table 1

Ultrasonic Measurements

(Continued)

Specimen Coordinate Direct Ultrasonic Measurements (rnn)

N umber Measurement

(mm)

I II III IV Average

1C y 9.6

N

8.8 9.1 9.5 9.1

X 0 2.0 1.0 0 1.5

y 18.8 19.7

N N

19.0 19.4

X 0 0.5 0.0 0.3

2A y 12.1

N

12.0

N

11.5 11.7

X 0 0 -0.5 -0.3

y

X

18.4

0

N N N N N

26 y 11.7

N

11.2 11.0 11.5 11.2

X 0 0 0 0.5 0.2

y 18.1 18.3 17.4

N

18.5 18.1

X 0 -1.0 1.5 -0.5 0

2C y 11.0

N 9.0

N 9.8 9.4

X 0 1.0 1.5 1.2

y

X

18.0

0

N

17.9

0

N N

17.9

0

w-y

N N

7.0 7.0 7.0 7.0

X 3.0 2.5 3.5 3.0

285

Non-Dimensional

Depth Y

Figure 1. Sizing of Slits With The Double Transducer Technique - PP array (compressional impinging, compressional

diffracted waves).

286

287

Figure

2

.

Photoelastic

Visualization

of

P

Waves

(outer

circles)

and S

Waves

(inner

circles)

Diffracted

From

A

Tip

of

a

Slit

In

A

Glass

Block.

288

Figure

3.

Triangle

of

Diffraction

With A

P

Surface

Wave

As

Reference.

289

Figure

4.

Sizing

Of

Cracks

With

Tip

Diffracted

Waves

(A)

Single

Transducer

Technique

(B)

Double

Transducer

Technique.

Figure 5. Noisy Background of the Oscilloscope Trace. The expanded

Section Contains: (A) A Signal Diffracted From The Tip

Of The Crack (B) Signals From Protrusions On The Crack

Surface (C) A Signal Diffracted From The Tip Of The Notch

and (D) Noise Signals From The Material. (C) Is The ZeroReference Point. Time AT Is Measured From (C) To The Left.

290

0

00

I) <0 o

1 O

« £

5 §

53

> f

CD

0) £ ®

O c r

0)

o

C Q) £

CD —

S o) a

.22 c a)

13 co -o

co co (0 3 3 3

CO CO CO

Lm a) a) a)

> > >

a> a) a)

T3 "U T5

3 3 3

5 S II a o. a

p r r

§ E

E P

C*

T3

c

E E E <<<

< £ CD nsi o'

E

E

>

epruudiue

dpnijidijue

291

292

Figure

7.

Specimens

With

Inwelded

Cracks

293

Figure

8.

Cross

Section

of

As

Welded

Specimens

With

Normal

and

45°

Inwelded

Cracks

294

Figure

9.

Cross

Section

of

Machined

Specimens

With

Normal

and

45°

Inwelded

Cracks

Figure 10. Experimental Set-Up

Figure 11. Transducer Arrangement

29 5

296

Figure

12.

Oscil

loscopic

Trace

of

(a)

Direct

Beam,

(b)

Reference

Signals,

and

(c)

Double-skip

Beam.

The

signal

to

the

Left

and

to

the

right

of

the

tip-diffracted

signal

is

the

back-wall

reflection

with

transducers

shifted

from

their

maximum

amplitude

positions.

Figure 13. Ultrasonic depth measurements compared with direct depth

measurements of "inwelded cracks". Direct measurements

were taken at the outside surface and ultrasonic

measurements in the middle of the specimen.

i

297

Noise from grains and

small imperfections.

Multiple reflections from crack

surface when one transducer is

used.

Large dead zone because of weld

crown when a direct beam is used.

Poor resolution. The strong

backward signal mask the weak

diffracted signal.

Figure 14. Schematic Presentation of Some Problems In Sizing

Cracks in Welds.

298

References

1. Golan, S. , "Sizing of Cracks with Scattered Ultrasonic Waves,"

Proc. of First International Symposium on Ultrasonic Materials

Characterization, National Bureau of Standards, Gaithersburg,

MD, June 1978 (in press).

2. Silk, M. G. and Lidington, B. N., "Defect Sizing Using an Ultrasonic Time Delay Approach," British J. NDT, 17, 33-36

(1975).

3. Silk, M. G. , "Sizing Crack-like Defects by Ultrasonic Means,"

Chapter 2 in Research Techniques in Nondestructive Testing , 3,

R. S. Sharpe, Ed., Academic Press, New York (1977).

-

4. Golan, S. , Baruch, J. and Yaroh, S. , "Sizing of Cracks with

Time Delay of Scattered Ultrasonic Waves," Israel Institute of

Metals - Technion, Internal Report 014-146 (1977).

5. Golan, S. , Adler, L. and Cook, V., "Ultrasonic Diffraction

Technique for Characterization of Fatigue Cracks," to be

publ ished.

6. Y. R. Birchack and C. G. Gardner, "Comparative Ultrasonic

Response of Machined Slots and Fatigue Cracks in 7075 Aluminum,

Materials Evaluation , 34, 275-280 (1976).

7. Krautkramer, Ultrasonic Testing of Materials , Second Edition,

Springer-Verlag, 186-189 (Berlin 1969).

299

APPENDIX E - Flaw Characterization

R. C. Placious

Progress, Aug 1, 1979 - Nov 30, 1979 :

Radiographic activities for this period were directed primarily

toward the evaluation of manual versus automatic processing of radiographs. The measurement standard for defect dimensioning was designed

by, and is presently being fabricated at, NBS/Gaithersburg. The Rocky

Flats group will use this device in connection with pipeline specimens

from the weldments fabricated at H. C. Price Co. in Dewey, OK. With

this arrangement. Rocky Flats will produce four film standards for

defect dimensioning. Further details on this activity are given in the

letter and the memorandum which are appended to this report as Attachments 1 and 2. Work on the assessment of defects in field radiographs

will proceed following completion of the film standards.

Other radiographic work at NBS on existing weld specimens containing

defects is continuing. One of the most serious problems in pipeline

radiography is crack detectability. For example, a series of NBS radiographs showed that a crack as wide as 0.005 in. and having a depth of

about one weld pass is not ordinarily detectable if the crack is approximately parallel to the sidewall chamfer. Such cracks can usually be

detected only if the x-ray beam is properly oriented. How to best

reveal these "hidden" cracks is undergoing continued study since this is

regarded as a critically important part of this project. Multiple

exposures represent one possible answer.

The report presenting guidelines for assessing field radiographs is

still being assembled. There are many possible approaches toward developing a guide for evaluating existing radiographs and suggestions

from OPSR regarding format would be very beneficial. NBS would rather

not generate recommendations which are not useful or which are rapidly

outdated. The intent is that recommendations for useful image-quality

indicators will make it possible for inexperienced individuals to interpret defect size with acceptable accuracy in the absence of full information regarding exposure and processing conditions.

Ultrasonic activities for this period were largely concerned with

the low signal -to-noise ratio inherent in time-delay diffraction measurements. In order to optimize the experimental set-up with respect to

shear-wave angle, a series of measurements were taken to characterize

the forward and back scattering from cracks. The results were that

300

optimum signal is obtained for 47° shear-wave backscatter (pulse-echo).

For forward scatter (pitch-catch, two-transducer technique) the optimum

angles are 47° and 57° for scatter from the top and bottom of the crack,

respectively. However, these amplitudes are well below that for backscatter.

Measurements were taken on intentionally-cracked weld specimens

using the two-transducer technique with a tuned pulser/receiver, 5-MHz,

45° shear-wave transducers, and a 50-MHz oscilloscope. The results were mixed. Where measurements could be taken, the measurements correlated

fairly well with actual size measurements, typically within a few tenths

of a millimeter. However, in many instances the measurements could not

be made due to excessive noise and spurious signals from grain boundaries.

In some of the specimens the weld crown was removed prior to testing.

This appeared to improve the ability to make measurements, but did not

necessarily improve the accuracy of results. Some results are given in

Table 1

.

Some further experiments were performed on more field-oriented

equipment using the pulse-echo (one-transducer) technique. A 5-MHz,

duel-element transducer was used with a portable, broadband, pulser/

receiver with integral digital thickness readout. This system provides

optimum spacial resolution, easily detecting discontinuities within 1 millimeter of specimen walls, and direct readout of depth of discontinuity. Measurements taken on simulated cracks (elox notches) agreed with

actual values within about 0.2 mm on notches from 3 to 10 mm deep in

specimens of carbon steel bar stock and X-65 pipe. These results are given in Table 2.

Plans, Dec. I, 1979 - Jan. 31 , 1980 :

As soon as the radiographic film standards become available, the

Rocky Flats group will be asked to evaluate representative defects, both

artificial and natural. Both field and laboratory radiographs will be

involved. An attempt will be made to also include a second organization

in the evaluation.

Efforts to define better ways of revealing cracks will continue.

In this connection, the primising developments in real-time radiography

will be followed.

Ultrasonic measurements, using both the one-transducer and the twotransducer techniques, will be made on samples containing continuous

lack-of-fusion (samples cut from pipe weld specimen 4-7482). These

measurements will be correlated with actual defect sizes obtained by

sectioning or from direct measurements made on the sides of the samples.

301

Table 1 - Some Results of Crack Sizing Measurements in Weld Specimens

Using the Two-Transducer Time-Delay Technique

Specimen Coordinate Actual Location Measured by Ultrasonics (mm)

Number Location

(mm) I II III IV Average

1A Y 9.1 N N 9.0 N 9.0

(Cracks normal X 0.5 1.0 1.0

to surface, weld Y 9. A 9.5 N N N 9.5

crown on) X 0 0 0

Y 20.2 N N N 20.0 20.0

X 0 0.5 0.5

IB Y 9. A N 10.5 9.5 N 10.0

(Cracks normal, X 0 -1.0 -1.5 -1.2

crown removed) Y - 20.1 20.8 20. A N 20.5 20.6

X 0 -1.5 0 0 -0.8

3A Y 11.1 N N N 11.0 11.0

(Cracks A5° to X 0 0 0

surface, crown on) Y 19.6 N 19. A N 20.0 19.7

X -8. A -8.5 -8.0 -8.3

3C Y 11.0 10.6 10.7 11.5 10.7 10.9

(Cracks A5°, X 0 0 1.5 1.0 0 0.6

crown removed) Y 18.8 18.7 N N 18.5 18.6

X 8. A 8.0 8.0 8.0

Y depth below top surface

X lateral position relative to arbitrary reference

N measurement not possible

302

Table 2 - Results of Ultrasonic Measurements of El ox Notches

in X-65 Pipe

Notch Actual Depth ( mm) Measured Depth (imm)

I II III Avers

A 9.9 10.0 9.7 9.8 9.8

B 7.4 8.0 7.3 7.2 7.5

C 5.3 5.4 5.0 4.8 5.1

D 3.0 2.8 3.2 3.1 3.0

303

u.r. DEPnrjTcvjRWY of cor/minncE

fJational Ouresu of ^Standards

Washington. D C i-0234

August 4, 1979

Mr. Wayne Stump

Mr. O.L. Summers

Rocky Flats Plant

Energy Systems Group

P.0. Box 464

Golden, CO 80401

Dear Wayne and Jack:

We have performed a series of tests with the radiographic exposures

you made for us. I would like to suggest the following relative to

your measurement standard. We now feel quite sure that 10 minute

manual processing gives a film response nearly equal to automatic

processing (Kodak X-OMat). We do not know - the reverse situation - i.e. , what temperature and time cycle in an automatic processor

gives a response similar to 5 or 8 minute manual processing. Rather

than take the time to do this I propose that you proceed with making

up four film standards using a portion of one of the pipe sections

you received from Kasen. Plan on making up these standards according

to the following recipe:

Exposure #1 , 220 kVcp, 5/AA/10 - Automatic processing at recommended

time/ temperature cycle.

Exposure #2, 220 kVcp, 5/AA/10, Automatic processing at reduced

time/ temperature cycle

Exposure #3, 170 kVcp, 5/AA/10, Automatic processing at recommended

time/ temp, cycle.

Exposure #4, 170 kVcp, 5/AA/10, Automatic processing at reduced

time/temp, cycle.

This will yield four standards for defect assessment and hopefully

provide us with information on contrast change due to kilovoltage

and/or processing variations. With the available budget constraints

this is probably about all we can hope to accomplish.

The concept of reduced automatic processing is an attempt to simulate

the film response for manual processing at its reconinended time and

temperature. One might achieve the reduced temperature in the automatic processor (say e.g. 70°F) by disconnecting the heater (or pulling the wall plug) at the close of business one day and processing

the next day before the temperature has risen. John Callinan (Kodak)

says the X-OMat will run OK at 74°F but 70°F might be closer to the

latitude achieved in manual processing.

Attachment 1

304

The necessary film density needed in the 4 film standards is about

2.2 or 2.3 in the lightest area of the image. That is - the density

at point A in my sketch (attached) of the proposed step wedge to be

used. I will now discuss this step wedge.

Nine pieces of .010" shim stock, cut as indicated and containing a series of holes on one side should be placed on a piece of the pipe

from Kasen. This will give us .090”of added material which should

then approximate the average thickness of the weld including the

crown. The contrast produced by the 9 steps and the 9 hole images

are what you will use to develop your film standard. Note that

these contrast changes will be going in the same direction as the

image of a real defect in the weld region.

If this all sounds complicated - or if you see a serious problem

with the procedure - could we plan on discussing it at the St. Louis

meeting? I will be there from Monday noon until Wednesday late

afternoon. I am looking forward to discussing this with you. If

you prefer, I could supply the shim stock step wedge for you.

Si prprpl \r

Robert C. Placious, Physicist

Radiation Physics Division

Center for Radiation Research

Enclosure

Stated

cc: L. Mordfin

M. B. Kasen

D. Polansky

305

SPECIAL

STEP

WEDGE

FOR

R.F.

FILM

STANDARD

ro|!£

O

* <

in

306

UNITED STATES DEPARTMENT OF COMMERCE

National Bureau of Standards

Wj*ihingT:on. D.C. 20234

October 24, 1979

MEMORANDUM FOR Dr. M. B. Kasen X 3

Cryogenics Division

NBS, Boulder

From: Dr. R. C. Placious

Radiation Physics Div.

NBS, Washington, D. C.

Subject: Pipeline Meeting at Boulder, CC on October 25, 1979

Flaw Assessment (Primarily geared to old program)

1.

Rocky Flats is making up 4 film standards to evaluate our field audiographs (made at HC Price and Co. by Cleveland X-Ray).

Details of Film Standards.

(a) 170 kV, Automatic Processing at Standard Cycle

!

b) 220 kV, Automatic Processing at Standard Cycle

c) 170 kV, Automatic Processing at Reduced Temperature

(d) 220 kV, Automatic Processing at Reduced Temperature

These standards will provide data on variations in depth assessment due

to kilovoltage and automatic versus manual processing, (c) and (d) are

„ designed to provide the kind of latitude obtainable in manual processing.

Note: Kasen has copy of letter to Wayne Stump with details of above.

2.

NBS, Washington, D.C. is evaluating effect of film type on defect

assessment. This will be done by use of a scanning microdensitometer.

The completion of (1) and (2) will provide us with much of the information

needed to (a) Provide DoT with the information needed to evaluate a waiver

request; (b) Permit NBS to make recommendations to DoT on improving field

radiographic procedures; and (c) Place limits on accuracy obtainable using

film contrast to assess flaw depth.

3.

Other Activities.

(a) We are trying to determine probabilities of crack detection by

radiography and ultrasonics.

(b) We are cooperating with the ultrasonics group at NBS-Wash. on the

alternatives to radiography.

4.

We have Dan Polansky at NBS now. He brings long experience in radiographic

NDT with him. He is presently very active in high energy real time radiography and is expected to provide valuable help in assessing the capabilities

of this technology in pipeline radiography in the field. He is also

Chairman of ASTM E-7 which has responsibilities for all consensus 'tandards

in NDT.

Attachment 2

307

Ill FLAW CHARACTERIZATION

Subtask 1: Radiography

Principal Investigator: Dr. Robert C. Placious, (301) 921-2201

Progress : This progress report addresses the suggested procedures for evaluating waiver requests on the basis of present state-of-the-art field radiography. We propose a scope for the radiographic record that should accompany

the waiver application. The evaluation of dimensional inaccuracies that arise,

as a result of the radiographic techniques used, can only be made if this information is available. A radiographic record could well take the form shown

in Figures 1 and 2. The detailed information for evaluating flaws defined by

field radiography and stated by the record will certainly include graphs of

the type shown in Figure 3. (Note that the figures are of a preliminary

nature and thus, in the interest of expediency, it was not felt that they

needed to be prepared professionally.) Improved radiographic procedures would

include specifications such as:

a. Shims, if used, shall include actual thickness identification

numbers (as penetramet.ers do) .

b. Portable x-ray sources shall be. referenced to a prototype unit

which has had its output calibrated (kV, mA, time).

c. Elastic belts shall be required to maintain weld/film contact etc.

308

• [JIZ > I KWL- | l y I- I ^ / I l

• -IX ( ' v / \ . (re T.c ^Y^iTnrX) WiTtf /\tj y u^i/CR TcLcqJ^^T^

309 /% /

itjregrf al

PATA

OK

£XT£fsf"taLsoo^ce

‘DATA

o r^.

Rli^

AN>p>

cAPo50f<F

'PATA

PlL^\ T>^t. ClaOS* ^X?iRA7i:>J 'PAT'S:

ScRf E/^S, MeTaLj ^uJo^^e^T o*>

' fU> o r^\ETA LLl c_ .

f°l £TA !—

^

AV£T :;' ’ A. V„

^ TH 1 Cfs”

. AJb$^ P^O^TF A *sfi> PAC^

Ft U & l\S C£ NFf* C>\'l t A 0

- METALLIC -Tyrs ^rAp'R/

X' klaY - -pova/ccl jNjpc^rrv

j~ s

COT POT CJiV * avi AT)

T*i M-TrJ Aj —

298

/=?& 2 <l

310

Vrr 1 A*

Fro A o To[A a'H C / ma *ju a l.

fKOTo /A AT; c — p£u/~Z rs’??K ^ iAore-u.

c H '5/- \/ >T*^y A C/c

L

it

.

MA fJ iM L - c. M

m

1-^rR.V

^ TlnE 'cTfc/'-'lp.

Fl\wJ

C.H ft^AcTERl- LAW A/J/s f V<; ! c;

y i u

.

IA; E lT> AfTO T>ATE

T7Pl of FLAW/

A^T> Q^aLiP^AITtJ

jVOre^p^GTL*^ Afv)X> Qi/At-iRC/Cn^vJ

piAA^ K^^rJAL A^ALVs\S

t— £ nJ A>T pi

j W iT'T H \ *I)’c f i A

Pro c e t>J :<r *£> osE'fr FotO>i *J 5?omi

a

1 6

.

(^S'A'L. *7* \f\) £’’& br ~ x *Dc»‘OS fTO A^£7“‘|.<yY £> TRcT ^ (rQoiPA\£;N'r ussy

tvpe ^ pi p\^j fa c-T^^ fc,\ c/A Li b

/

vr\ cm

A P u RTJ[;c 5 t *£• cz S E*

T

c ,

CTTM^R. 5^ Pr-Q'iZTi L\ & TetSTS

1?eMAlB/< S

311

7^ V JL* /vo ~y

312

Plans: Plans call for a continuation of present efforts to acquire data

which relate to probable inaccuracies in flaw analysis due to x-ray variables,

processing variables, and operator skill and training. These efforts seek to

further define:

a. The contents and the completeness of the radiographic record.

b. The detailed information needed to evaluate the data on the

radiographic record.

c. The changes needed in field techniques or procedure to improve

image quality and, therefore, flaw analysis.

Subtask 2A: Ultrasonics

Progress : The following interagency report has been issued, which summarizes

the NBS work on the ultrasonic time-delay diffraction technique for sizing

defects:

S. Golan, "Defect Characterization and Dimensioning of

Cracks in Welds by the Ultrasonic Diffraction Method,"

NBSIR 80-1983, March 1980.

A copy of this report is appended hereto. Additional copies may be ordered

from the National Technical Information Service, Springfield, VA 22161.

Plans : There are no plans for continuing this work at the present time.

313

Ill FLAW CHARACTERIZATION

Subtask 1: Radiography

Principal Investigator: Dr. Robert C. Placious, (301) 921-2201

Progress : Figures 1, 2a and 2b of the last progress report presented a tentative listing of the essential elements needed to evaluate a pipeline

weld flaw detected on a field radiograph. Figure 3 of that report indicated how the evaluation would be carried out in cases where documentation

of the radiographic procedure was unavailable or only sketchily available.

Efforts during the present reporting period continued to examine the

influences of various x-ray parameters on radiographic contrast. The last

progress report provided data on the variation of contrast with metal thickness, with processing conditions as a parameter. During the present reporting

period the effects of kilovoltage changes on contrast were examined. See the

figure on the next page. A number of graphs of this kind will be needed to

perform a proper evaluation of flaw depth. DoT has now been provided with

two such graphs.

Plans : Efforts to characterize the influences of various x-ray parameters

on radiographic contrast will continue, using data from NBS studies and from

external sources. Problems relating to x-ray film type variability, and to

kilovoltage versus manual processing variability, are among those remaining

to be addressed.

314

DENSITY

CHANGE

vs.

METAL

THICKNESS

(kV

AND

PROCESSING

TEMP.

AS

PARAMETERS

)

A1ISN3Q oiHdvaooiava

315

INCREASE

IN

THROUGHWALL

THICKNESS

(mils)

Ill FLAW CHARACTERIZATION

If

Principal Investigators: Dr. Robert C. Placious, (301) 921-2201

Mr. Daniel J. Chwirut, (301) 921-3646

• # Progress : The NBS interaction with the Rocky Flats group (RF), on radiographic dimensioning flaws, is continuing. (The tests are described in the letter furnished on pp. 28-30 of the Seventh Report).

Radiographs were made of metal lographic samples containing high

porosity and slag from the pipeline section welded at the H. C. Price

Co. (See, for example. Figure 1.) It is anticipated that threedimensional radiography will be required in order to obtain an accurate

analysis of the volume of the flaws in these damples.

An in-house assessment of the ultrasonics effort on this project

has led to the conclusion that the time-delay diffraction technique

holds the most promise for this application. Most of the NBS work

on this technique has been carried out with in-house funds although

some of the results have been included in these progress reports.

An interagency report summarizing the NBS effort was prepared and

»is presently undergoing detailed technical review prior to publication.

Plans : The RF phase of the radiographic effort will be completed and,

upon NBS approval of the exposures, RF will make up film standards for

flaw measurements.

The tomographic information obtained from the three-dimensional

radiographic work will be provided to other memebers of the project

team to facilitate more precise fracture mechanics analysis.

Ultrasonic examinations of samples containing continuous lack

of fusion (which were deferred due to delays in procuring angle wedges)

will be started.

An effort will be initiated to assess the effects of weld-bead

removal, and the relative merits of single vs. double skip techniques,

for flaw depth sizing by the ultrasonic time-delay diffraction approach.

Copies of the NBS report on the time-delay technique will be

furnished as attachments to the next progress report.

316

Fof^Ti

High

Contrast

Film

-Normal

Contrast

Film

F,V- Polaroid Pictures of Radiographs of Metallographic Samples. Quantitative

Determination of Degree of Porosity May Require Special Radiographic Procedure (3 Dimensional Radiography)

317

Ill FLAW CHARACTERIZATION

(Principal investigators: Dr. Robert C. Placious, 301-921-2201,

and Mr. Daniel J. Chwirut, 301-921-3646)

Progress, Aug 1, 1979 - Nov 30, 1979 :

Radiographic activities for this period were directed primarily

toward the evaluation of manual versus automatic processing of radiographs. The measurement standard for defect dimensioning was designed

by, and is presently being fabricated at, NBS/Gaithersburg. The Rocky

Flats group will use this device in connection with pipeline specimens

from the weldments fabricated at H. C. Price Co. in Dewey, OK. With

this arrangement. Rocky Flats will produce four film standards for

defect dimensioning. Further details on this activity are given in the

letter and the memorandum which are appended to this report as Attachments 1 and 2. Work on the assessment of defects in field radiographs

will proceed following completion of the film standards.

Other radiographic work at NBS on existing weld specimens containing

defects is continuing. One of the most serious problems in pipeline

radiography is crack detectability. For example, a series of NBS radiographs showed that a crack as wide as 0.005 in. and having a depth of

about one weld pass is not ordinarily detectable if the crack is approximately parallel to the sidewall chamfer. Such cracks can usually be

detected only if the x-ray beam is properly oriented. How to best

reveal these "hidden" cracks is undergoing continued study since this is

regarded as a critically important part of this project. Multiple

exposures represent one possible answer.

The report presenting guidelines for assessing field radiographs is

still being assembled. There are many possible approaches toward developing a guide for evaluating existing radiographs and suggestions

from 0PSR regarding format would be very beneficial. NBS would rather

not generate recommendations which are not useful or which are rapidly

outdated. The intent is that recommendations for useful image-quality

indicators will make it possible for inexperienced individuals to interpret defect size with acceptable accuracy in the absence of full information regarding exposure and processing conditions.

Ultrasonic activities for this period were largely concerned with

the low signal -to-noise ratio inherent in time-delay diffraction measurements. In order to optimize the experimental set-up with respect to

shear-wave angle, a series of measurements were taken to characterize

the forward and back scattering from cracks. The results were that

318

optimum signal is obtained for 47° shear-wave backscatter (pulse-echo).

For forward scatter (pitch-catch, two-transducer technique) the optimum

angles are 47° and 57° for scatter from the top and bottom of the crack,

respectively. However, these amplitudes are well below that for backscatter.

Measurements were taken on intentionally-cracked weld specimens

using the two-transducer technique with a tuned pulser/receiver, 5-MHz,

45° shear-wave transducers, and a 50-MHz oscilloscope. The results were

mixed. Where measurements could be taken, the measurements correlated

fairly well with actual size measurements, typically within a few tenths

of a millimeter. However, in many instances the measurements could not

be made due to excessive noise and spurious signals from grain boundaries.

In some of the specimens the weld crown was removed prior to testing.

This appeared to improve the ability to make measurements, but did not

necessarily improve the accuracy of results. Some results are given in

Table 1.

Some further experiments were performed on more field-oriented

equipment using the pulse-echo (one-transducer) technique. A 5-MHz,

duel -element transducer was used with a portable, broadband, pulser/

receiver with integral digital thickness readout. This system provides

optimum spacial resolution, easily detecting discontinuities within 1 millimeter of specimen walls, and direct readout of depth of discontinuity. Measurements taken on simulated cracks (elox notches) agreed with

actual values within about 0.2 mm on notches from 3 to 10 mm deep in

specimens of carbon steel bar stock and X-65 pipe. These results are given in Table 2.

Plans, Dec. I, 1979 - Jan. 31 , 1980 :

As soon as the radiographic film standards become available, the

Rocky Flats group will be asked to evaluate representative defects, both

artificial and natural. Both field and laboratory radiographs will be

involved. An attempt will be made to also include a second organization

in the evaluation.

Efforts to define better ways of revealing cracks will continue.

In this connection, the primising developments in real-time radiography

will be followed.

Ultrasonic measurements, using both the one-transducer and the twotransducer techniques, will be made on samples containing continuous

lack-of-fusion (samples cut from pipe weld specimen 4-7482). These

measurements will be correlated with actual defect sizes obtained by

sectioning or from direct measurements made on the sides of the samples.

319

Table 1 - Some Results of Crack Sizing Measurements in Weld Specimens

Using the Two-Transducer Time-Delay Technique

Specimen Coordinate Actual Location Measured by Ultrasonics (mm)

Number Location

(mm) I II III IV Average

1A Y 9.1 N N 9.0 N 9.0

(Cracks normal X 0.5 1.0 1.0

to surface, weld Y 9. A 9.5 N N N 9.5

crown on) X 0 0 0

Y 20.2 N N N 20.0 20.0

X 0 0.5 0.5

IB Y 9. A N 10.5 9.5 N 10.0

(Cracks normal, X 0 -1.0 -1.5 -1.2

crown removed) Y 20.1 20.8 20. A N 20.5 20.6

X 0 -1.5 0 0 -0.8

3A Y 11.1 N N N 11.0 11.0

(Cracks A5° to X 0 0 0

surface, crown on) Y 19.6 N 19. A N 20.0 19.7

X -8. A -8.5 -8.0 -8.3

3C Y 11.0 10.6 10.7 11.5 10.7 10.9

(Cracks A5°, X 0 0 1.5 1.0 0 0.6

crown removed) Y 18.8 18.7 N N 18.5 18.6

X 8. A 8.0 8.0 8.0

Y depth below top surface

X lateral position relative to arbitrary reference

N measurement not possible

320

. Table 2 - Results of Ultrasonic Measurements of El ox Notches

in X-65 Pipe

Notch Actual Depth ( mm) Measured Depth (imm)

1 II III Aver;

A 9.9 10.0 9.7 9.8 9.8

B 7.4 8.0 7.3 7.2 7.5

C 5.3 5.4 5.0 4.8 5.1

D 3.0 2.8 3.2 3.1 3.0

321

U.C DEPACITCVJFWT CF COMMIIHCE

rJutional Qure?u of Stundut ds

Washington. DC i?Q034

August 4, 1979

Mr. Wayne Stump

Mr. J.L. Summers

Rocky Flats Plant

Energy Systems Group

P.0. Box 464

Golden, CO 80401

Dear Wayne and Jack:

We have performed a series of tests with the radiographic exposures

you made for us. I would like to suggest the following relative to

your measurement standard. We now feel quite sure that 10 minute

manual processing gives a film response nearly equal to automatic

processing (Kodak X-OMat). We do not know - the reverse situation - i.e. , what temperature and time cycle in an automatic processor

gives a response similar to 5 or 8 minute manual processing. Rather

than take the time to do this I propose that you proceed with making

up four film standards using a portion of one of the pipe sections

you received from Kasen. Plan on making up these standards according

to the following recipe:

Exposure rl, 220 kVcp, 5/AA/10 - Automatic processing at recommended

time/ temperature cycle.

Exposure $2, 220 kVcp, 5/AA/10, Automatic processing at reduced

time/ temperature cycle

Exposure r 3, 170 kVcp, 5/AA/10, Automatic processing at recommended

time/temp, cycle.

Exposure #4, 170 kVcp, 5/AA/10, Automatic processing at reduced

time/temp, cycle.

This will yield four standards for defect assessment and hopefully

provide us with information on contrast change due to kilovoltage

and/or processing variations. With the available budget constraints

this is probably about all we can hcpe to accomplish.

The concept of reduced automatic processing is an attempt to simulate

the film response for manual processing at its recommended time and

temperature. One might achieve the reduced temperature in tne automatic processor (say e.g. 70°F) by disconnecting the heater (or pulling the wall plug) at the close of business one day and processing

the next day before the temperature has risen. John Callinan (Kodak)

says the X-OMat will run OK at 74°F but 70°F might be closer to the

latitude achieved in manual processing.

Attachment 1

322

The necessary film density needed in the 4 film standards is about

2.2 or 2.3 in the lightest area of the image. That is - the density

at point A in my sketch (attached) of the proposed step wedge to be

used. I will now discuss this step wedge.

Nine pieces of .010" shim stock, cut as indicated and containing a series of holes on one side should be placed on a piece of the pipe

from Kasen. This will give us .090'of added material which should

then approximate the average thickness of the weld including the crown. The contrast produced by the 9 steps and the 9 hole images

are what you will use to develop your film standard. Note that

these contrast changes will be going in the same direction as the

image of a real defect in the weld region.

If this all sounds complicated - or if you see a serious problem

with the procedure - could.we plan on discussing it at tne St. Louis

meeting? I will be there from Monday noon until Wednesday late

afternoon. I am looking forward to discussing this with you. If

you prefer, I could supply the shim stock step wedge for you.

Si ppprpl \r

Robert C. Placious, Physicist

Radiation Physics Division

Center for Radiation Research

Enclosure

Stated

cc: L. Mordfin

M. B. Kasen

D. Polansky

323

special

step

wedge

FOR

R.F.

FILM

STANDARD

toise

j H

324

UNITED STATES DEPARTMENT OF COIYIMlRCE

National Durcnu of Stamfords

W'j'ihmij'on. DC.

October 24, 1979

MEMORANDUM FOR Dr. M. B. Kasen X 3

Cryogenics Division

NBS, Boulder

From: Dr. R. C. Placious <Zc<P

Radiation Physics Div.

NBS, Washington, D. C.

Subject: Pipeline Meeting at Boulder, CC on October 25, 1979

Flaw Assessment (Primarily geared to old program)

1. Rocky Flats is making up 4 film standards to evaluate our field audiographs (made at UC Price and Co. by Cleveland X-Ray).

Details of Film Standards.

(a) 170 kV, Automatic Processing at Standard Cycle

(b) 220 kV, Automatic Processing at Standard Cycle

(c) 170 kV, Automatic Processing at Reduced Temperature

(d) 220 kV, Automatic Processing at Reduced Temperature

These standards will provide data on variations in depth assessment due

to kilovoltage and automatic versus manual processing, (c) and (d) are

„ designed to provide the kind of latitude obtainable in manual processing.

Note: Kasen has copy of letter to Wayne Stump with details of above.

2. NBS, Washington, D.C. is evaluating effect of film type on defect

assessment. This will be done by use of a scanning microdensi tometer.

The completion of' (1) and (2) will provide us with much of the information

needed to (a) Provide DoT with the information needed to evaluate a waiver

request; (b) Permit MBS to make recommendations to DoT on improving field

radiographic procedures; and (c) Place limits on accuracy obtainable using

film contrast to assess flaw depth.

3. Other Activities.

(a) We are trying to determine probabilities of crack detection by

radiography and ultrasonics.

(b) We are cooperating with the ultrasonics group at N3S-Wash. on the

alternatives to radiography.

4 . We have Dan Polansky at NBS now. He brings long experience in radiographic

NDT with him. He is presently very active in high energy real time radiography and is expected to provide valuable help in assessing the capabilities

of this technology in pipeline radiography in the field. He is also

Chairman of AST‘1 E-7 which has responsibilities for all consensus 'tandards

in NDT.

3

O r- uJ

Attachment 2

APPEDNIX F - Ultrasonic Examination of Girth Welds in 40-Inch, X-65 Line

Pipe Specimens

W. C. Minton

326

SOUTHWEST RESEARCH INSTITUTE

• 220 CULEBRA ROAD • POST OFFICE DRAWER 28510 • SAN ANTONIO. TEXAS 78284

0U4HT* ASSURANCE SYSTEMS AND ENGINEERING DIVISION October 2, 1978

TELE X 767579 NUC ENGR SN1 TELEPHONE I51?l CfU Mil

TELECOPIER 482?

TRIP REPORT NUMBER 1

Southwest Research Institute Project 17-5388-001

TRIP TO: National Bureau of Standards

Boulder, Colorado Laboratories

BATES: August 23-26, 1978

SPONSOR: U.S. Department of Commerce

National Bureau of Standards

Washington, D.C. 20234

Purchase Order No. 813887

BY: Willard C. Minton

Technical Consultant

PURPOSE

OF TRIP:

Ultrasonic Examination of Girth Welds in 40-Inch, X-65

Line Pipe Specimens

PERSONNEL National Bureau of Standards

CONTACTED: “

M. Bi Kasen

D. Chwirut

Southwest Research Institute

W. C. Minton

D. A. Gonzalez

R. Harrison

SAN ANTONIO. HOUSTON. CORPUS CHRISTI, TEXAS. AND WASHINGTON. D C.

327

Trip Report No. 1

October 2, 1978

Page 2

1.0 SUMMARY

All or portions of five girth welds in specimens of AO-inch, X-65 line

pipe were examined ultrasonically by manual techniques. Numerous ultrasonic

reflectors were recorded. The recorded information from the examinations

was plotted and attempts made to separate geometric reflectors, such as those

originating at the inside reinforcement or root of the weld, the external

reinforcement or crown of the weld, high-low, and other nondetrimental

conditions from reflectors present in the weld due to flaws which must be

evaluated in accordance with API Standard 110A for Welding Pipelines and

Related Facilities. The attempts were only partly successful with standard,

vide-beam ultrasonic examination techniques.

The examinations were first performed in the as-welded condition.

Following this, a selected portion of one weld was ground flush, first on the inner surface and then on both surfaces, in order to record the effect

of removal of these geometric reflecting conditions. The results clearly

indicated the interpretation difficulties encountered in attempting to dis- « criminate between geometric and detrimental reflectors in the as-welded

condition.

Recommendations are made for improvements in resolution so that geometric

reflectors may be distinguished from detrimental flaws.

A Search Unit Tracking and Recording System (SUTARS) semi -mechanized

ultrasonic examination was performed on three of the test welds. A reduction

in time required to examine a weld was obtained. A preliminary computer

evaluation of a portion of the SUTARS data verified the results of the allmanual ultrasonic examinations.

2.0

DESCRIPTION OF SPECIMENS

The specimens examined were numbered A through 8 and consisted of

section's of AO-inch outside diameter (0D) , X-65 line pipe (average wall

thickness 0.65 inch). Each of the specimens was at least 2 feet in length

on either side of the girth weld, which joined the sections approximately

at the center. Some of the specimens had been welded manually; others had

been welded by automatic methods. Both automatic and manual welds were

examined by the same ultrasonic procedure.

3.0

EQUIPMENT AND PROCEDURES

3 . 1 Manual Ultrasonic Examination

A manual ultrasonic examination was performed in accordance with

328
.
Trip Report No. 1

October 2, 1978

Page 3

Southwest Research Institute (SwRI) Procedure SwRI-NDT-800-58, Revision 0,

a copy of which is attached in Appendix A.

3.1.1 The ultrasonic instruments used were conventional

pulse-echo flaw detectors. Sonic Instruments type FTS Mark I. These

instruments can be operated either from self-contained batteries or from

alternating current line.

3.1.2 The search units consisted of the following:

(a) Straight-beam: 3/8-inch diameter, 5.0 MHz;

(b) Angle beam: 3/8-inch diameter, 2.25 MHz.

These were used with plastic, low-noise angle beam wedges to produce sound

beam angles in steel of 45° +2° or 60° +2°.

3.1.3 Reference blocks used for screen distance calibration

were half-round mild steel test blocks or IIW test blocks.

3.1.4 The basic calibration block used to adjust equipment

for all examinations was .604 inch thick mild steel and contained

side-drilled hole reflectors as shown in Figure 1.

3.1.5 Couplant material was Sonogel II, a commercial product.

3.1.6 The welded specimens were supported on wooden cradles

about one foot off the floor, a position similar to that which would be found

along a pipeline right-of-way prior to burial of a welded line.

3.2 Semi -Mechanized Search Unit Tracking and Recording System

• This consisted of the SwRI-developed equipment shown schematically

in Figure 2. The function of the SUTAR System is to replace the laborious

manual recording of each datum point produced during the ultrasonic scanning

of welds. The equipment consists of a conventional ultrasonic flaw detector

modified so that selected information from the examination can be recorded

in a format suitable for computer evaluation. In addition, the coordinates

of the search unit positions during the scanning of the weld are continuously

.recorded in a similar format.

4.0 MANUAL EXAMINATIONS

4.1 Lamination Scans

A preliminary examination of the material adjacent to the weld

was performed using straight-beam (0°) 5.0 MHz examination frequency. The

purpose of this examination was to detect and record any laminar conditions

329

Trip Report No. 1

October 2, 1978

Page 4

330

FIGURE

1

Basic

Calibration

Block

Trip Report No. 1

October 2, 1978

Page 5

Receiving Array

331

Trip Report No. 1

October 2, 1978

Page 6

In the steel plate from which these sections of pipe were made. Such laminar

conditions can interfere with other ultrasonic examinations conducted from

or through those areas.

4.2 Veld Examinations

The original intent was to use 45° and 60° shear wave examinations

introduced from both directions perpendicular to the weld axis for detection

of flaws in the plane of the weld, and to use 45° shear waves from both sides

of the weld directed circumferentially for detection of discontinuities not

oriented with the weld axis. Table 1 shows the examinations which were

actually completed. Because of the time required to record the extremely

large number of indications exceeding the threshold recording level, it was

not feasible to examine 100 percent of all welds with the scans originally

planned. The data recorded consisted of the search unit position coordinates

and the indicated sound path distance for those indications exceeding 20

percent of the amplitude of the reference level. This reference level was

established by constructing a distance amplitude curve utilizing the sidedrilled hole targets in the basic reference block. The manual examination

ultrasonic data are attached in Appendix B, together with the data sheets

showing the instrument adjustments for each of the scans.

Each of the indications recorded was plotted manually to show

its position in the cross section of the weld. The flaw plots are attached

in Appendix C.

TABLE I Ultrasonic Examination Coverage

Test

Weld

No.

0° Lamination Scan

From* Through*

45° Weld Scan

From* Through*

60 e i From*

Weld Scan

Through

4 90 100 90 100 90 100

5 115 126 - 115 126 — —

6 0 127 0 22 107 117

*7 0 125 1/2 0 125 1/2 120 - 0 - 08

8 0 127 0 127 «... —

Numerals refer to x-ray film stations.

332

Trip Report No. 1

October 2, 1978

Page 7

5.0

SUTARS EXAMINATIONS

Those velds examined by SUTARS received the same scans as the other

manual examinations. In both cases, the search unit vas manipulated by hand

and the data consisting of search coordinates, amplitude of ultrasonic indications and the sound beam path distances were recorded. The difference

with SUTARS is that the data were recorded on a digital magnetic tape recorder

for later processing by means of a computer program. Specimens Nos. 6, 7,

and 8 were examined using SUTARS.

6.0

RESULTS

6.1

Manual Examinations

6.1.1

Test Weld No. A

The lamination scan with 0° sound beam showed six small

laminar conditions in the base material adjacent to the examined portion of

the veld. All six of the laminar conditions were judged too small to affect

the results of the veld examinations appreciably since none vas as large as the sound beam.

As indicated in Table 1, the examination coverage vas

approximately 10 inches of weld extending from radiographic station 90 to

radiographic station 100 . This area was chosen since radiographs previously

made shoved side-wall lack of fusion in this region. Both 45° and 60 ° examinations were conducted from both sides of the weld. At radiographic station

95-1/2, a concentrated examination was performed using both 45° and 60 °.

Attached in Appendix C are cross-section plots of these indications.. The

side-wall lack of fusion indicated on the radiographs was shown with 60 °, but

had surprisingly low amplitude. It was not clearly shown using the 45° sound

beam, however. The majority of the ultrasonic indications proved to be due

to the root bead geometry conditions and external weld crown irregularities.

6.1.2

Test Weld No. 5

The lamination scan with 0° sound beam showed two small

laminar conditions in the base material adjacent to the examined portion of

the weld, which was a 10-inch sector between radiographic stations 115 and

125 . These two laminar conditions were too small to appreciably affect the

results of weld examination.

Weld examination of Test Weld No. 5 was conducted with

45° sound beam only from both directions. Cross-section Pi ots of the ultrasonic indications show high amplitude root area reflectors from the root

geometry and low amplitude indications of the slag condition which radiographs

had shown in this weld. One of the apparent slag indications (Indication

333
.
Trip Report No. 1

October 2, 1978

Page 8

No. 5 on Data Sheet 1306) located at radiographic station 122 was plotted

to Its half amplitude points. This plot indicates that the apparent flaw

is trapped slag between passes near the side vail of the veld.

6.1.3 Test Weld No. 6

The lamination scan vith 0° sound beam showed nine laminar

conditions in the base material adjacent to the examined portion of this weld.

All nine of the laminar indications were small spots, too small to affect the

results of the veld examinations appreciably.

The areas examined in Weld No. 6 consisted of 45° scans from radiographic stations 0 to 22. Also, 43° and 60° scans were performed

in the area from radiographic stations 107 to 117. Cross-sectional plots

of the indications seen with the use of 45° and 60° angle beams are attached

in Appendix C. The plots indicate that a variety of conditions were picked

up by the ultrasonic beam. The 60° scans from both directions showed numerous

geometric indications due to weld crown irregularities, root reinforcement,

and surface irregularities in the parent material as well as flaw indications

such as lack of penetration or side wall lack of fusion. Centered at x-ray

station 112, a lack of penetration flaw visible in the radiograph of this

weld showed a high amplitude indication. However, other indications due to

root bead reinforcement geometric reflector conditions also produced high

amplitude indications in the same region.

6.1.4 Test Weld No. 7

The lamination scan using 0

e sound beam showed 12 laminar

indications in the base material adjacent to the butt weld in this pipe. With

one exception, all of these were only spot indications with negligible effect

on the examination of the weld itself. The exception was from radiographic

stations 18-1/2 through 20-1/8 and extending to slightly over 2 inches from

the upstream side of the weld. A lamination was detected in this area approximately 0.18 inch below the OD. Severity of this lamination was great enough

to prevent weld examination with angle beams from the upstream direction of

this nearly 2-inch-long area. There was also a visible large open surface

lap about 1 foot long near. this position. This visible condition extended

longitudinally on the surface of .the base material upstream from the weld.

Th is weld was scanned completely with a 45° angle beam

around the circumference of the pipe from both the upstream and downstream

sides. In addition, scans with 60° angle beam were conducted from radiographic stations 120 through 0 to 08 from both upstream and downstream sides.

Additional concentrated 60° scans were performed at selected locations in

•this area. To determine the effect of the internal and External weld bead

reinforcement on the ultrasonic results, a small area near radiographic

location 3 was first ground flush on the inside surface and then on the outside

surface. Following each of these grindings, the areas were reexamined to

334

Trip Report No. 1

October 2, 1978

Page 9

determine the effect. These examinations are shown in the plots of indications

in Appendix C. The results show that grinding the beads flush produces a marked increase in the signal amplitude from the known root crack condition

in the area which was ground flush as well as eliminating the interference

of the root and crown reinforcement reflectors.

a A relatively unused technique for crack depth determination

vas tried in the flush-ground area. This technique, called the "precursor

method," consists of determining the difference in path lengths of a sound

beam reflected from the comer formed by the intersection of a crack with

the surface and the weak reflection of that beam due to crack-tip diffraction.

As shown in Figure 3, this difference should be proportional to the crack

depth. The depth thus indicated at radiographic station 3-1/8 was 0.021

inch in this test weld.

6.1.5 Test Weld No. 8

The lamination scan with 0° sound beam showed that this

base material was nearly free of laminar conditions adjacent to the weld.

One small spot—a laminar reflector smaller than the sound beam—was located.

A 45° scan was performed completely around this test weld

from the upstream and downstream sides. Most of the reflectors recorded

proved to be produced by the weld crown reinforcement geometry. A crosssection plot of these indications is attached in Appendix C. Six of the

eight indications which plot approximately mid-wall in the weld actually

can be shown to be due to mode conversion and redirection of the sound beam

to the crown surface, since the indications can be manually damped on that

surface.

6.2 SUTARS Examinations

; V~ ’ The magnetic recording tapes on which the data from the SUTARS

examinations were acquired were retained by SwRI. In accordance with the

•contract, the final processing of these tapes to extract the data and evaluate

the examination information was deferred awaiting authorization of this

effort by the National Bureau of Standards (NBS) . A portion of data for

Test Weld No. 6 was processed and is attached in Appendix D.

7.0

CONCLUSIONS

7.1

Manual Examinations

7.1.1

The time required for complete manual ultrasonic examination of line pipe we3ds proved to be greater than anticipated. The necessity

to consider and evaluate reflectors which produced amplitudes as low as 20

percent of the reference level required that the search unit coordinates and

335

Trip Report No. 1

October 2, 1978

Page 10

A. Components of Sound Beam Reflected and

Refracted from a Notch or Crack

for Determining Crack Depth

- FIGURE 3

336

Trip Report No. 1

October 2, 1978

Page 11

other data from all such reflectors be recorded. Root bead and veld crown

reinforcements—so called geometric reflectors—constituted large and variable

Interference which was present at nearly al'l locations along each of the welds.

7.1.2 Geometric interference occurs at the same transducer . coordinates and together with indications of real defects, and is essentially

inseparable from such defect indications with the conventional wide-beam

transducers which were employed during these examinations.

7.1.3 Removal of weld beads by grinding eliminates the problem

and might prove feasible in special cases or for examinations of localized

areas, but is considered impractical for routine work.

7.1. A A 45° sound beam angle is more sensitive than 60° to

geometric reflectors as well as to most real -defects.

7.1.5 The "precursor" method, where applicable, shows promise

for improved determination of the depth of detected flaws.

7.1.6 Mode conversion interference was not an appreciable

fa ctor with manual welds, but did occur with one of the test welds which

was welded automatically. This appeared to be due to a small internal weld

preparation edge, which served as a mode converting surface. Such a condition

would be avoidable with a slightly different edge preparation.

7.1.7 Collimaticr. , focusing, or other high-resolution techniques

are needed for better separation of real flaws and geometric features in

welds such as these.

8.0

RECOMMENDATIONS

8.1 The SUTARS data acquired during these examinations should be

processed 'completely , evaluated, and compared with the manually acquired

data from each weld examined by both methods.

8.2 Higher-resolution methods,- such as collimated, focused, or concentrated beam transducer techniques need to be developed for better separation

and identification of geometric and detrimental flaw indications.

8.3 Adaptive learning methods offer additional possibilities for

reducing the data to eliminate geometric indications.

Technical Consultant

Department of Research and Development

Quality Assurance Systems and Engineering

Division

337

Appendix A

SOUTHWEST RESEARCH INSTITUTE

fcK

OPERATING PROCEDURE

at

SwRI-NDT-800-58

Revision 0

August 1978

Page 1 of 23

1 MANUAL ULTRASONIC EXAMINATION OF LINE PIPE WELDS

SwRI-NDT-800-58

1.0 PURPOSE

This procedure provides the technical information and detailed steps

- < / necessary to ensure a complete and accurate manual ultrasonic examination of

|^

pressure piping welds and the adjacent base material, in accordance with

applicable Codes.

)’ SCOPE AND APPLICATION

Piping welds and the adjacent base material in the nominal thickness

range of greater than 0.3 to 2.0 inches shall be examined in accordance with

this procedure.

Manual, contact, pulse-echo shear-wave angle beam and longitudinal wave

straight beam ultrasonic techniques shall be utilized for the examination of

piping welds and adjacent base material.

Welds to be examined shall ba circumferential and longitudinal pipe

welds or as specified in the applicable SwRI Examination Plan.

2.1

applicable:

Applicable Documents

The following document forms a part of this procedure as

( 1 ) API STD 1104, Standard for Welding Pipelines and

Related Facilities, Fourteenth Edition.

3.0 RESPONSIBILITY

(1) The Director of the Department of Engineering Services, Quality

Assurance Systems and Engineering Division, shall be responsible

for the preparation, review, appro /al, ana control pf this procedure.

(2) The examiner shall be responsible for implementing the requirements of this procedure.

CODE AND PROCEDURE REQUIREMENTS

Not applicable.

S^wRl Fc.rm QA J-1 338
-
Appendix A

SOUTHWEST RESEARCH INSTITUTE

OPERATING PROCEDURE

SvRI-NDT-800-58

Revision 0

August 1978

Page 2 of 23

D PERSONNEL AND EQUIPMENT

5.1 Personnel Certification

Personnel performing examinations in accordance vith this

procedure shall be certified in accordance vith SvRI NQAP 11-1, "Special

’rocess Control."

5.2 Reference Block

Reference blocks used for screen distance calibration and

'v's‘ verification shall be either of the following:

= B-

(1) SwRI Half-Round (Sketch 1)

(2) AWS Type DC (Sketch 2)

(3) Ilk7 (Sketch 3) or DSC

The SvRI Half-Round or AMS Type DC reference block shall be of

the same basic material' as the production material, e.g.

,

carbon steel.

stainless steel or

5.3

Basic Calibration Block for Circumferential, Longitudinal,

Support Attachment, and Branch Pipe Connection Welds

Side-drilled basic calibration hole reflectors shall be placed

in a block manufactured from material equivalent to the production naterial.

The calibration material shall be determined by the production piping material

to which' the search unit is applied.

The basic calibration block thickness shall be determined by

the thickness of the production piping material to which the search unit is

applied. When a basic calibration block of the same thickness as the

production piping material is not available and where the production piping

*:s- material thickness is 1 inch or less, the basic calibration block thickness

shall be within 25% of the production material.

C An example of a basic calibration block is attached in sketch U.

Swrfti form QA 3)

339

Appendix A

SOUTHWEST RESEARCH INSTITUTE SvRI-NDT-800-58

Revision 0

OPERATING PROCEDURE

August 1978

Page 3 of 23 *

£

O

Flat basic calibration blocks or blocks of essentially the same

v curvature as the part to be examined may be used when contact surface curva- tures are greater than 20 inches in diameter.

A single curved basic calibration block may be used to estab- • lish sensitivity calibration for examinations on contact surfaces in the range

of curvature from 0.9 to 1.5 times the basic calibration block diameter, when

contact surface curvature is 20 inches in diameter or less.

2 d 5.4 Search Units

(1) The search unit size shall be selected according to the

following table:

(a) Straight Beam : 3/8" or 1/4" round

(b) Angle Beam : 3/8" x 3/8", 3/8" round or 1/2" x 1/2", 1/2" round

(2) The exit point of the sound beam and the actual refracted

. bean angle of shear-wave search units shall be determined

on an IIU or DSC block. The exit point shall be marked

on the search unit wedge.

(3) The nominal search unit frequency for the examination of

carbon steel piping shall be 2.25 MHz.

(4) For examination of circumferential piping welds, search

unit wedges shall be fabricated to produce 45® ±2° and

60® ±2® refracted shear-waves.

(5) For examination of longitudinal piping welds, search unit

wedges shall be fabricated to produce 45® ±2® refracted

shear-waves.

5.5 Ultrasonic Instrument

The examiner shall select an appropriate ultrasonic 'instrument

from the fellcuing:

(1) Sonic FTS Mark I

(2) Sonic FTS Mark II

(3) Sonic UWA Mark III

I F cm m Cj A* 3 i 340

Appendix A

SOUTHWEST RESEARCH INSTITUTE

OPERATING PROCEDURE

SvRI-NDT-800-58

Revision 0

August 1978

Page 4 of 23

CE

»v

f*

i'H

1

|i

n

I

Each instrument shall be aligned and display a current alignment

calibration sticker.

5.6 Couplant

(1) Sonogel, Ultragel, Echogel or other commercial couplants

shall be used vhen performing ultrasonic calibrations and

examinations in accordance vith this procedure.

(2) Couplant materials used for examinations shall be the

same as used for the calibration.

6.0 CALIBRATION METHOD

The complete ultrasonic examination system calibration shall be performed prior to the examination.

NOTE

The ’’REJECT” control shall be maintained in the ”0” position

during calibration and examination.

6.1 Calibration

6.1.1 Straight-Beam Distance Calibration

In all cases, the screen distance chosen shall be the shortest

applicable size to include at least l/4t beyond the thickest production

material to vhich the search unit is applied.

Observing back reflections from the applicable reference block,

adjust- "MATERIAL CALIBRATION," "RANGE," and "DELAY" controls of the instrument

to obtain the required linear sound path distance, displayed along the screen

baseline.

Screen distance calibration shall be selected from the

following: 1.0", 2.5", and 5.0".

6.1.2 Ar.ple-Seam Distance Calibration

In all cases, the screen distance chosen shall be the shortest applicable size which will include at least 1/8 vee path past the

anticipated examination range.

» of mi ua 3- i

341

Appendix A

SOUTHWEST RESEARCH INSTITUTE

OPERATING PROCEDURE

SwRI-NDT- 800-58

Revision 0

August 1978

Page 4 of 23

n

v

Each instrument shall be aligned and display a current alignment

calibration sticker.

5.6 Coup 1 ant

l (1) Sonogel, Ultragel, Echogel or other commercial couplants

shall be used when performing ultrasonic calibrations and

examinations in accordance with this procedure.

(2) Coup 1 ant materials used for examinations shall be the

^

same as used fo-r the calibration.

6.0

CALIBRATION METHOD

The complete ultrasonic examination system calibration shall be performed prior to the examination.

NOTE

The "REJECT” control shall be maintained in the "0” position

during calibration and examination.

6.1 Calibration

6.1.1 Straight-Beam Distance Calibration

In all cases, the screen distance chosen shall be the shortest

applicable size to include at least l/4t beyond the thickest production

material to which the search unit is applied.

Observing back reflections from the applicable reference block,

adjust "MATERIAL CALIBRATION," "RANGE," and "DELAY" controls of the instrument

to obtain the required linear sound path distance, displayed along the screen

baseline.

Screen distance calibration shall be selected from the

following: 1.0", 2.5", and 5.0".

6.1.2 Ar.ple-Becm Distance Calibration

In all cases, the screen distance chosen shall be the shortest applicable size which will include at least 1/8 vee path past the

anticipated examination range.

V*MI f uni, (jA -3 1 342

Appendix A

SOUTHWEST RESEARCH INSTITUTE SwRI-NDT- 800-5 8

•

• Revision 0

August 1978

tg OPERATING PROCEDURE

Page 5 of 23

Observing radius echoes from the applicable reference block,

adjust "MATERIAL CALIBRATION," "RANGE," and "DELAY" controls of the instrument

to obtain the required linear sound path distance, displayed along the screen

baseline.

sizes:

Screen distances shall be selected from the following

2.5", 5.0" and 10.0".

i ¥

*

.it

- When a single instrument is used for both 45® and 60® examinations, the screen distance calibration shall be conducted in the followmanner:

(1) The screen distance size shall be determined by the angle

beam search unit requiring the longer examination range.

(2) Position the 45* ±2* search unit on the appropriate

reference block, observe and record all verification

block entries as required on the appropriate SvRI Instrument Calibration Record.

(3) Position the 60® ±2° search unit on the appropriate

reference block and, without changing the "MATERIAL

CALIBRATION," "RANGE," or "DELAY" controls, observe and

record all verification block entries as required on the

appropriate SwRI Instrument Calibration Record.

(4) No attempt shall be made to compensate for any slight

delay difference observed between 45® ±2° and 60® ±2®

screen distance calibrations. This difference, if any,

instead shall be considered when resolving indications.

6.1.3 Angle-Beam Distance Amplitude Correction

If a curved block is used, distance amplitude correction (DAC) curves for the examination of circumferential welds shall be constructed by

utilizing the responses from the holes oriented perpendicular to the axis of

the basic calibration block for butt welds and parallel to the axis for

longitudinal seam welds.

6. 1.3.1 Material 1 Inch or Less in Thickness

45® and 60® DAC

win cj.'t 3-i

343

Appendix A

SOUTHWEST RESEARCH INSTITUTE

OPERATING PROCEDURE

SwRI-NDT- 800-58'

Revision 0

August 1978

Page 6 of 23

(1) Position 45® ±2® search unit to obtain maximum response

from the hole and position producing the highest

amplitude from the following:

Hole Vee Path Positions

1/2T 2/8, 6/8, 10/8

(2) Adjust instrument gain controls to obtain the primary

reference response at 75% ±5% of*FSH and mark this

amplitude on the screen. The gain controls shall not b

adjusted once the primary reference response has been

established. - .

(3) Position the search unit to obtain maximum response fro

the remaining positions and mark each amplitude on the screen.

(4) Join these points with a smooth curved line, the length

of which shall not extend more than 1/8 vee path

beyond the last qualified calibration point.

(5) Repeat steps (1) through (4) using a 60® ±2® search

unit.

NOTES

If the configuration of the weld is such that a 45® ±2®

full node examination covers the entire weld volume and

the 60® ±2® search unit sound bean is directed into the

weld root on the straight pass, indications detected

beyond the 6/8 vee path while conducting 60° ±2°

examinations shall not be recorded.

If the configuration of the weld is such that a 45® ±2®

vee path examination fails to cover the entire weld

volume and the 60® ±2® search unit sound beam is directed into the weld root on the straight pass, and a 60 c ±2® full vee examination covers the entire veld

volume, indications detected in the calibrated area shall be recorded.

*full screen height

• •

344

SOUTHWEST RESEARCH INSTITUTE

OPERATING PROCEDURE

SvRI-NDT-800-58

Revision 0

August 1978

Page 7 of 23

EXCEPTIONS

If the configuration of the weld is such that the

60° ±2° search unit sound beam is not directed into

the veld root on the straight pass, a 14/8 vee path

calibration shall be accomplished with a 45® ±2°

search unit. Indications detected in the calibrated

area shall be recorded.

If the configuration of the veld is such that a 45* ±2® 14/8 vee path examination and a 60® ±2® 10/8

vee path examination fail to cover the entire weld

volume, a 14/8 calibration shall be accomplished with

both a 45® ±2® and a 60® ±2® search unit and indications detected in the calibration area shall be

recorded.

6. 1.3. 2 Material Creater Than 1 Inch But Not Over 2 Inches in

in Thickness

45® and 60® DAC

( 1 ) Position the 45® ±2® search unit to obtain maximum

response from the hole and the position producing

the highest amplitude from the following:

Hole

1/4T

3/4T

45® Vee Path Positions

7/8, 9/8

3/8, 5/8

( 2 )

(3)

(4)

Adjust instrument gain controls to obtain the primary

reference response at 75% ±5% of FSH and mark this

amplitude on the screen. The gain controls shall not

be adjusted once the primary reference response has

been established.

Position the search unit to obtain maximum response from

the remaining positions and mark each amplitude on the screen.

Join these points with a smooth curved line, the length

of which shall not extend more than 1/8 vee beyond

the last qualified calibration point.

WvRi » C«M»1 UA 3 )

345
mmm
Appendix A

SOUTHWEST RESEARCH INSTITUTE SwRI-NDT-800-58

Revision 0

• August 1978

[£!W,

Uy OPERATING PROCEDURE

Page 8 of 23

to

to

i

; 1

(5) Repeat steps (1) through (4) using a 60* ±2* search unit

utilizing the following positions:

Hole

1/4T

3/4T

60* Vee Path Positions

7/8

3/8, 5/8

EXCEPTIONS

6.2

6 . 2.1

If the configuration of the weld is such that a 45* ±2'

full vee examination fails to cover the entire weld

volume a 9/8 vee calibration shall be accomplished with

a 60* ±2* search unit.

If the configuration of the weld is such that the

60* ±2* search unit beam is not directed into the weld

root, a 13/8 vee calibration shall be accomplished

with a 45* ±2* search unit.

Calibration Verification

Sweep Range and DAC Curve Verification

<3

Sweep range and DAC curve calibration shall be verified on the

appropriate reference and basic calibration block:

(1) At the start of a series of examinations

(2) With any substitution of search unit cable

(3) With any substitution of power source

(4) At least every 4 hours during the examination

(5) At the finish of a series of examinations

(6) At any time when, in the opinion of the examiner, there

is doubt as to the validity of the calibration.

6.2.2 Calibration Chances

(1) Perform the following if any point on the DAC curve has

decreased more than 20% or 2 dB in amplitude, or any

point on the DAC has moved on the sweep line more than

52 of full screen width:

S*Pl f o»m QA 3 ] • .

346

Appendix A

SOUTHWEST RESEARCH INSTITUTE

OPERATING PROCEDURE

SvRI-NDT- 800-58

Revision 0

August 1978

Page 9 of 23

6.2.3

calibration:

ri

(2 )

(a) Void all examinations referring to the calibration In question and performed after the last

valid calibration verification.

(b) Conduct a new calibration.

(c) Reexamine the areas for which examinations have

been voided.

Perform the following if any point on the DAC curve has

increased in amplitude more than 20% or 2 dB:

(a) Correct the calibration.

(b) Reexamine all indications recorded since the last

valid calibration verification.

(c)

Re calibration

Enter proper values on a new SvRI Examination

Record.

Substitution of any of the following shall be cause for re-

(1) Search unit (shoe /transducer)

(2) Couplant

• (3) Ultrasonic instrument

(4) Examination personnel.

6.3 Calibration Data

Calibration data shall be recorded on the appropriate SwRI

Instrument Calibration Record, and shall include but not be limited to:

(1) Serial number of the basic calibration block used

(2) Type and serial number of the reference block used

(3) .Type, manufacturer, and serial number of the ultrasonic

instrument

V»R* f orm OA j-1

347

Appendix A

SOUTHWEST RESEARCH INSTITUTE

rc\

©

&

OPERATING PROCEDURE

SwRI-lTDT-800-58

Revision 0

August 1978

Page 10 of 23

(A) Type* size* bean angle* and serial nunber of the

search unit

(5) Nominal search unit frequency

(6) Couplant (which shall be the sane type as used in the

actual examination)

(7) Signature and ultrasonic certification level of the

examiner conducting calibration

(8) 'Date calibrated

(9) Time of calibration and verification (s)

EXAMINATION

IU

KOTE

Examination data shall be recorded as applicable on the

appropriate SvRI Examination Record (examples attached).

7.1 Examination Areas

7.1.1 Circumferential and Longitudinal Butt Welds in Pining

Welds may be examined from the inside or outside surface

of the pipe. Base material shall be examined for a distance of It (nominal

thickness of pipe) from the fusion line on each side of the weld.

Surface Condition

The contact surfaces shall be free from weld spatter,

roughness, or other conditions which interfere with free movement of the

search unit or impair the transmission of ultrasound.

7.3 Indication Length Zero Reference (Lo) Location

If not already marked with a cere reference location, areas to be examined in accordance with this procedure shall have an "Lo" marked

.in accordance with SvRI practice: normally top center of a butt weld or upstream end of a longitudinal seam weld.

Fi OA J )

348

Appendix A

SOUTHWEST RESEARCH INSTITUTE

OPERATING PROCEDURE

SwRI-HDT-800-58

Revision 0

August 1978

Page 11 of 23

7.4 Scanning Parameters

When practicable, scanning shall be performed at a minimum

- ^. gain setting of 2 times the reference level sensitivity.

Instrument gain setting for scanning shall be determined

on the basic calibration block as follows:

c JV,

(1) Manipulate search unit on the basic calibration block

to obtain a signal from a calibration reflector

(side-drilled hole) of 50% FSH.

(2) Add 6 dB of gain by adjusting the instrument fine

gain control only. Observe the signal amplitude.

(3) Add 6 dB of gain by adjusting a combination of both

fine and coarse instrument gain controls. Observe

the signal amplitude.

(4) Add 6 dE of gain by manipulating the instrument 6 dB

switch, if present. Observe the signal amplitude.

(5) Choose the method above which yields a signal response

closest to 100% FSH.

(6) The method chosen shall then be used during the valid

calibration period for all scanning at 2 times the

reference level sensitivity.

The examiner shall ensure that the signal response of the

method chosen is within ±2 dB of 100% FSH.

Instrument gain settings for scanning shall be recorded on the appropriate SwRI Examination. Record.

The search unit movement rate for scanning shall not exceed

6 inches per second.

7.4.1 Scanning

v Scanning overlap shall be a minimum of 10% of the search

e / unit piezoelectric element dimension perpendicular to the direction of

|, scan. Scanning performed in a direction perpendicular to the weld shall be

5 rV adequate to examine the weld and It of base material from the fusion line

of the weld.

SywRl f oim QA J l

349

Appendix A

SOUTHWEST RESEARCH INSTITUTE

&

OPERATING PROCEDURE

SwRI-NDT- 800-58

Revision 0

August 1978

Page 12 of 23

7.4.2 Base Material Lamination Scan

A lamination scan using straight beam longitudinal waves

shall be performed before the angle-beam examination. This examination

shall cover as much as practical of the area through which the angle beam is

later to be passed. Screen distance calibration for this examination shall

be conducted in accordance with Paragraph 6.1.1. Scanning sensitivity

shall be as required to maintain the first back reflection at an amplitude

between 50% and 90% of FSH.

Intermediate echoes having an amplitude equal to or greater

than 50% of the initial back reflection and accompanied by a 50% loss of back 5 reflection shall be recorded.

¥

'll

To record an intermediate indication, a back reflection signal shall be obtained from an indication- free area and the instrument gain

control adjusted until this signal is at 75% ±5% of FSH; then record the inter m

'mediate indication when its amplitude * is equal to 50% of the initial back re- flection and accompanied by a 50% loss of back reflection. If total loss

^ - of back reflection accompanies the intermediate echo, the area of total loss

^of back reflection shall be recorded on the appropriate SvRI Examination

Record.

7.4.3 Thickness Measurements

In conjunction with the lamination scan, thickness measurements shall be taken adjacent to Lo for longitudinal and circumferential

welds (on the centerline of the weld if feasible and at one point in the

base material on both sides of the weld). If these measurements cannot be

taken adjacent to Lo, then the location of the measurements shall be recorded

on the appropriate SwRI Examination Record.

Screen distance calibration for this examination shall be

conducted in accordance with Paragraph 6.1.1. Measurements shall be taken

by placing the 0° search unit on the appropriate position on the examination

surface and observing the position of the back wall reflection on the

instrument screen. These measurements shall be recorded on the appropriate

SwRI Examination Record.

7.5 Examination of Circumferential and Longitudinal Butt Welds in

Piping

7.5.1 Angle-Beam Examinatfon for Indications Parallel with the Weld

Angle beam examinations for circumferential welds shall be

accomplished using 45 ±2° and 60° ±2° refracted shear waves from both sides

of the weld. For this examination, the sound-beam shall be directed

f Orn, On j ]

350

Appendix A

SOUTHWEST RESEARCH INSTITUTE

ret

• OPERATING PROCEDURE

SwRI-NDT- 800-58

Revision 0

August 1978

Page 13 of 23

perpendicularly into the weld to detect indications which are parallel to

the weld. Calibration for these examinations shall be in accordance with

Paragraphs 6.1.2 and 6.1.4.

' Angle beam examinations for longitudinal welds shall be

accomplished using a 45° ±2° refracted shear wave from both sides of the

weld. For this examination, the sound-beam shall be directed perpendicularly

o' Or into the weld to detect indications which are parallel to the weld.

7.5.2 Angle-Beam Examination for Indications Perpendicular to the

Weld

An angle-beam examination for indications perpendicular

(transverse) to the weld shall be conducted on each weld using a 45° ±2 e shear

wave. This examination shall be conducted by placing the search unit adjacent

to the weld with the beam directed along the weld. The entire length shall

be scanned with the search unit sound beam directed in this manner. The

search unit shall then be turned 180° and the scan repeated. Calibration

for these examinations shall be in accordance with Paragraphs 6.1.2 and

Postexamination Cleaning

Arrangements shall be made with the customer for postexami- ^ nation removal of couplant materials.

Jh

8.0 RECORDING CRITERIA

Ultrasonic reflectors producing a response 50% or greater of the reference level shall be recorded on the appropriate ~SwRI Examination Record.

Indications producing a response 20% or greater of the reference level

and suspected by the examiner to be other than geometrical in nature shall be

recorded and investigated by a Level II or Level III examiner to the extent

necessary to determine the shape, identity, and location of the reflector.

s

H

I

Indications 100% or greater of the reference level shall be investigated

by a Level II or Level III examiner tc the extent necessary to determine the

sho^s , identity , and ^ocaitca of .me reflector.

Indications 20% or greater of the reference level investigated and

found to be other than geometrical in nature shall be reported to the

customer for evaluation.

Any scanning limitations shall be recorded.

SwKi | mm 3.

351

Appendix A

352

000*1

Appendix a

{

Revision 0

August 1978

Page 15 of 23

353

1

.000

Appendix A _ SwRI-NDT-800-58

Revision 0

August 1978

Page 16 of 23

O

inco

•

!

354

SKETCH

2

Appendix A SwRI-NDT-800-58

Revision 0

August 1978

Page 17 of 23 .

355

Appendix A

SvRI-NDT- 800-58

Revision 0

August 1978

Page 18 of 23

O

l-i u

(O

356

Sw.R.I

SONIC

INSTRUMENT

CALIBRATION

RECORD

Appendix A SwRI—NDT-800-58

Revision 0

357

Aim

fcl-{l

III

<114 I

•»“S

“II

iil*Ol

Appendix A SwRI-NDT-800-58

Revision 0

August 1978

Page 20 of 23

358

Appendix

A SvRI-NDT-800-58

Revision 0

August 1978

Page 21 of 23

T

3e

IN]

©z ZOP<•u su.J,

Li

CLto

-JLJ*

WELD

LENGTH •

I*. O

l*j o<a

LJ o

•J z o to e* &< i8 r> croo

LJ

cr z

lJ

L-

»-<

CL

=>

COX

cr<

!jo

->

L-zo o

LJc zo

h* c owC<

•n

O

kwo LOCATION

u>52U 5o

cru

SEARCH UNIT

LOCATION

DATE

<5.

w

w Z . 2

v* o

-J o

&.TP / ^ o

LJ

crDCO<

CO o tJ IE<X

LJ

crV1

LJzXo

u- zo

CVi * zo2 zo

LJ2 XK Q.z> too

Cl • Zo 1 >< -<u Zo CO w

‘-J

-J

LJ>

f—<z ouL-z

LJ c

ffi •J

LJLJ

CO

Cl2

LJ_J

L»z o o Z fSI CO < O

L- j

-J cr=>o

%nO

fi» i E<

LJ

CD

H-3

CD oo

irCl e >

LJ

cr -J

>--Js

C-JC 2

LJ

to

CD< zo

¥-

CSf £

LJ

L-<

L- <

cr

CD

Z>

CO\

LJ>

LJ .

EVEL COO

CL

J

COo

CO

CO LJ

L-LO

LJ5

L-z

CO z

CO

_l

UJOZ

cr

arw

Cl2 w £

CO oLao zo

CL £ 5

fw

tr

COC * 2 o<

LJ

-j <

LJ o c<o

INO

wp

iL <1 >o uLJ c

L. LJz5 ^Socl

COXc< <2 3>

LJ

Oc <X

LJ <X

lJ <X

LJ go 2

LJz <X

lJ >

LJ

c.

359

Appendix A SV7RI-NDT-800-58

Revision 0

August 1978

Page 22 of 23

360

^PPendix A _ SvRI-:jdt- 800-5S

Revision 0

August 1978

Page 23 of 23

361

/

fZ •— *E 8

ec. c O c

•=

v

i:

p Mac JC 3

S.-B I!

*K±i, 3o Xw — ac —e .£

-o £ 8 c -g

V)

O n

tr o C *9 a *° o "2 uz> .a • "5 E *» 2 ©

00 o g

.s Jz£ o »u £ C

ti i H

c.

— <'*. J5.eS

ia **» 33 u <. *8

a

•g

O « s - E c >, _ o ^ 2 E

-~

a s .2 £ 2 — o got> *“ X vt £ E

j! 3 — *o £ o X £ 8 Cl

c.

_ o £ 0 c

JtO « E " B c « « u 0*8

i. ”0 e ± X c £ K 8 E

e.EXP

*0c

ill

lit *® c P P i "II g a ."2 2

'-5 JO o - • 8. 8 £

*C

p 5 £ = £ **» o _ o S*

X “

ec* .©s

JZ

1/5 "5 3 c ^ if« £ ~ c s I ;O &K

.* UJ o “ a

X *N *2 I -E 2 -e

I"X

-scoa >£ 0 w >. 3

«£ 2 3 . w 0 V

IA

W .2 0 2 n

L.O

Cl

wP 8P

CL

X (A ,5 « a

0.

K

ec £ M

*5

W cC9 t > Z 3 C P X CL .3 V «0 0 p

a s- X t

VI

X

X •a*

X&3

Ui Xu

w 5 £ £

IE c

ec *gE e< a

ee

ecu 1O. j; £*“

O ^ •00

t/T pK

00 # c X5 M 0 0 JJ * C X 4= X

Cl t>

ec JT "S 3 •O ka t/5 e u

ec4- 8 £ £ 3 a C PMP P

ec

Ui Pc p

«c

u */> & do B£ Z.

X P —

H

©sV

BC

il8a

•ca

*5I " 3 a £ •a

QC P © -E H X 8*2

ecc

.©ao

Cla

•£ .1 S £>.* 2 | 8 w o

.E

r £ €|r c g «8 3 — E P O E t g o 3 ” o

<« IA

_ 111 X eC g « - 8 o w

*5Vo

*©ca c « •=

'a cl o 2 .<£. a

t? 33 £ a o

«A 3 © J2 g & a ” E ^ e. o

o- 2 « £ c o. .2 a a —

i'E 0> 0) • 8 £ - £

*n w. t = O R in . BC

W!

O

— Cog O O *s "° > c ^ o ^ U O S O E > 8 8 J E

'-5

II II II II II II W - H * 3 (C oor

u £ oa

J*8 3 w «. O C o £ 8

•=

*= 1

55 Z a a . o c 3

.£ c. — o « 8 c c K I*

*C 4)« § e J ? « ! ^

Wa£ £ |** JC o £ ^1

“•O « c ii J

*« .2

©

.102

— CL ” ^ a w ” ” o c i. a J

.2 ^ c «> a s5 !?

Cl -C

M

rl.

V jg

a

j:

«- £ 8 - .s

4> E T c O ac O o — •3 -3 — « 5 .3? »« K "• V g l> 3 8 £ w c ~ £ ia a w I 3 B

.2 -o £ •- 3

!: 1 x K 3

15 c ^ «- 5 Xp 8 CO

ccbm t *8 a \ T? W s> 8 0 1 cE

ec •c :•= s -a

S*

H

r© 1 e&

p v> -o Le i IS- 28a o — XT _ O c a, 0 - O w- a S £ .0. " g a 1 8 l

i-gc £ c » " o. S

•Si 1 .HT -C C £ E O—O > 5 o. C o. O 8

-» *2

>% JS

c.2 >^ o ^ W O IA 3 « V5 ^ 2“ g t Sl a o o a a «- L. O O « 3 CL a c

L.

O

BC > O •"

x: ac I *3

X3

•«» •<Q

*t33) acsQQ

«L g5 05C

”.~b ? s <

15

k_

o>

CD

_c

•8

J 1 0; ^ X

CO

XLO

05

v- o

<DO " 03

u- « 

O W

CO Q_

CO

— > CD

0)

<DC

*5

D.

CLc T9&

Cl.

V)Q

M

oc ^ O JS > * W 3 x S oa o c —

CLoo «^5- « g r >

I- £

g- t

wU 5 0 „ ^ O (A h-

•8

.8 S

.1 § o E 8 | t ? „ s .> £ 8

*o

_a o ® *C ^

•A 2 IA -- ac 3 = V.£

Clt oooa .5

IA ^ —

Cl lt u» JC 2

J= ^

x:

jc 2f wL. d 0£ L. ^ *£ . E t A 0

ill - C5 Is-

.'S 00 := c 3o Xo

JC -3 a- U x E c «

QJ LCU

Q-> t> &f J —CJ C _s

Q-> o S _ 2 I ct; 2W

Cla

O BC .2 '“

.2 .- 3 £ "3 CLa<°-iS 1 w w

L. > u i£ » 362

Matcriah

research

engineers,

physicist,

and

supervisory

niclallurgist,

respectively.

Fracture

and

Deformation

Division.

National

Bureau

of

Standards.

Boulder,

Colo.

80.10.1.

^he

ilalic

numbers

in

brackets

refer

In

the

list

of

references

appended

to

this

paper.

APPENDIX G - J-Integral Analysis of Surface Cracks

in Pipeline Steel Plates

Richard B. King, Yi-wen Cheng, David T. Read, and Harry I. McHenry

363

o ac o X x

1 a

s _ « ~z £ E 3 “ = I

J §

E t r>

« tr •*-* qj (/? ©. — — J: •* « .tr — • l_ u <* P" re 5

j“ c ^

.2 y c

re 5 .E '5 o i

g. -

1/5 0> 7 J JC D.

re 3

U 5

o w

c

E il 2 _• o —

o

| « i

.& s

c->

c-> re X

£ ^

w *

O x

» 3 O 4/5 S-g

ilCl O

E E

i — 75

t- X &C C£ £

£ i

S-5

V >, x o

=

K

0 2

II

1 2

I 2 *“ O

0 o

8> 8

7 £ b T? £ O — C

c J2 g e. E —

a. ©

w £

11

(9 o 2 —

© .£ x . u a j c 2 .E ^ °

o

'5 v —

* 2 u .£ 5 C b fil w -s 1 g

b — re a o • 2 *5 d a

g *

« o 2 £ (SO

. .£ R

O .b b 7y -

•c &) •=s:so Ji

.52 c V

.£ u

(5 ‘E

re E *“ x X o

fc w ~

c

g *5 2 o

C. 3

I* (0 c > £ o

# e

*

‘o

c. =

Jr o

3 O —

•5 "C "o J U H

E — 5 j< ?

e fc v *2

g « o .3

ac £ > J 1*5

*> e x B o .ac -?•&’§>

o a c

C *

n _ .52 t«&ST

2 *3

* s

re 1 -s it

t £ 7 7 o o E E n o

.2P £

"a

o o

— .•30

.f-5 B

O «

^ £ “ S>

>*» ^3 eoc 5 C B _

«*1 *2 2P 2 .£

.2 - x " £

B © — O

w

n X c

3 0 0

.£ E — -

K 2 B O

£

3 b

C O

a . 3

a < *s B

o .= O O

Jtf o

£ — B M

o c

.£ O

£ o £ •S = x

.5 B

X c |5'5E_^ —’ aj — *3 X

2

C 4> ^ g

S « I a u

O D

4/5 Cl ^

o — o !J

<2 c“

= o

•2 9 S £

-d a

&> re ii** *c

re *" £ £

7 "S i 2

. c 2 o

C u 4

o «— £ i . •o .-£ u g.

g I

o o

o

2 S. J i

O u

o -o g g

E — £ ® ^ —

3 —

£ -c

a. S .a ^

•

^ T O -O £"* O 3 C c

ao P 5 -

*P S O O

S O

3 3 O

J*5- — O W> w X •” e — 3

^ o 3 w> — S O _

0 S 12 = *o _ e

* 1

=' 3 O

ec w X 3

s-'i.

1) 1/5 .£ o — _s 3 ”

e _c g 2

B CJ

•3 •D 2 «

ac » o .o

•T 2

“* >% O X X W

"* X

% =e

— 3

•S g

O X o

3 u O Cl g O

2 ^ & g ^ &

"' O

o -=

ac3

£X

B

£ „ o — 3

c *° 7 ^ ii (SOB

“ I 2

B 3 *

O “ b B oc O

9^.£ c ® E.i

£ -3

111

O-o

ac ^3OX

S' g

.. « o — ^ .£ B _

S l

< »

o

*o

«£ 2 3 B 3 JS O 3

ac v

c E X o

o £ B c

(/> 3

7 g ajE

in i 2

E x

3

g ^

B *5 —

£ M E O 3 £ X O *“

** "o *o 3X0 c X 3

.111

5 a -c £ g

-o O X 2 X X B ~ 18 Jt 3 o> O = X «

W H o *»s

1

*5?© x ^

TJ re i !

6

b *2 &

3 2 -e

•= B « 2 ac b 3 s o B x O

jsc a £ o “» 2 o x

0^ U *

^ o r o .5 — re *r X J5 ^

3 b n “go. o E 3 £®o

M E ^ "33

£ g

-2 £ S 5 *° o £

O o 3 X -o 2 1 B

3 C -

.Si b O

isl

1113 ^

8 a- c

£ o O

c — K x 2 id X ac “ £ »; o .£ =

f re - -o .£ 0-30

g S* E

« E

I = s «

= t

3 ac c s

CB 3 3

o *x ” SS 2 g E 52 _3 X' X 3 X =

c11 2fc*3 —

Ow "O

01>

ac x

o3 o

.£ o3

o o

ac ac =X t/i * g

W i/i fc O

j= b 3 £ > £

's ’e -C .3 a

B g

•C v. O _

3 B v.

g

<2 2 2.

Bb

° =

&>3 ^ s U5 w « -c lax 0-3

8 SI

o — X -D C

Z B 1

3

B 8 s re *3 V -2 ^2 c LPl a

CC o 5 1 2

*

.52 JmT £ gb

s" 0

O 0>

o — a o

—

3 o>M

3 = E B p

ac « -S

«« *; a 3

. B 5 -2 s g

o —

o o

e r re t/ a .a, a x

7 8

=

£ |2

B O

* O

0> ac X 3 — O

3 E

> B

o £

e f

jc o> 2 « O g B a

_ s

ac g

3

3

ac TM

-E 8

.Si c

•u w

2 £

a . £- *B .

" — V o> o 8 -S 7 8

£ *

*rr u O

C

8 3

.a 5

0 &

1 1o

<N ^

>-> fc; B | g

3 P C

re c —

g .a I

g.^ 5

s! I

- X C.

re c «

a- &> CJ w ^ ^

ap.i ^ c re o 52 S X o 3

U b — B O .C ^ X = £

o C “ — Co X — X

0 ,0 7 3 o 1 *!

2 O

=

5 -«

_s

C c re a « U! c b ac c " oB

5 E =

•3 3 c a .2 c §|c.

"

o> .3hr

-c */> ^ £

oj re

*2 Cj

>3 CL . 3 » 3

I cl « ~ O

£ .0 x

S.*g g

0 -a « 1 ££

E « c

I

u •Il

a ac “ 3 “ !S 3 O r- c X ac 2 o .E a E 2

.E 2 <£

-It'S IF«« w

oX v

2 £

*re 0 x X V

c — 3

3 X

E — 3

£ — O3

u

# c 3 ^ E x

e b

3 £

fcP g rQ

3 5 o C 4/5 3 ^

< re «

c H

o a g .a £ _ x 3

£ re 3

c .s o o

o x x a

= 4 a a

a . “<

E x < o = a o £ £ O 3 - 3 o .E

0-^3

o o c U 2 o 5 ^ = O 4^ U

Z- cW £ "D

CJ wl

.ss r p ^ « - 2 E c f- 5_ xX 3

E * i o -g = x X o — 3 X

££) ^ ^

<E p — C L- 4T .3 cl re ka fiJ

U

re £

•u

"E.

CL. reo

reE

N © O V ^ T3 *- •* ^ S a o

c

•" M C. 8 t “

0 E c O C e o

f© I 7 \f© 2 3 < x « c ^ e

V

# >

t - g §

"re>X B _ 0

c c © © u u a

« rEJ

e>c

3 O

0 _ 1 7

364

if

I*5 cC2 E £ i " ©

.5 o c ac 1 a»

C3 DC K .£ « ^ O W £ ©£

>% "c? "E£

.£ >

vs OS 0 © 1 i

OS

C OS

VE . B " g X to OS C OS X

ill o "O — C — B ^ V

U Q.

k > |B ESr vs clbj —

;3

© -

"O OS

X

OS .*>*' O. k *• '££

i!x ^ os

wX

_oKXasas

CQ

tore xob 8

oj vs re 2 as o Cl *o E a.© < O Js . ** oc

as ^ Cl n| c

e. a.

F

_ c_

CQ CO

u

,-

V £ £-

U

K c n a>

*Z

— « .S2 £> • —

a J? 2 h « t £ «

CO

.© E

DC£

QJ

QJO

X>

a>>

CQ * !z g c « 8 !T ©

’i ^ S

ll! o © o © g e C *3 1/5

ac 5 *-

&> gE

<L> IT

w o -* 8 © W «C x> r&) IIas Xc

vs« £ «3 “ u -O 5 b c 2 5 « M Al * C

’5c je £ &

CQM’S &

.5 J? 3 v s 2 x c © § B ’O js DoS 2 .2 Z 2 a E > — .s t; .E ii ‘x _ — *3 8^5 £ s &

j< •© < 8 " OS 2 v x o X —

.. *3 8*8 - .Si S = ts 8

i2

B .2 £ E o

I*

g t> £ •*o c 2

&> o * £ © o

CO= i

S-8 • B X

«> £E o 8

*" vs X c

re I O as C as Cl X c

asE ~

fc 2

Cl .Si K .— % © - 8 g T< £ c as&£

Cls oS

c.•owX

*r9

V!eUs

BO .E x

w£iQ - « £ 2 ° « x « ^ .g ~ 8 I i S 3 c U

<£

M

D

b

r e 8 •« E B ! S |1

"o .= ' c .« !

QC - 1 I J 2 —

Sb

© c .2 5.1 O E N

as *o X cB

•o .. *c= as

VS Oi 2 c 3

•O ~

as

b 5 £oX X vs u

.!£

X E >

asE

<- *S 3 &1•« E

*D

©

w « w "

CO T5 S = 8 E

CO CO IS3 u m LC — CQ S|

CL >

>% D>UJ

*S *S "B E "C _ ^ B

— ^ •© *o 2

°- = .a & B c E - £

os *- « *o c X c

*©

a (QB

C

J & &

||5 * o £ o £

.5 - o i

o> . *0

k'

t> JD8 8 £ « < I 5 ^

rc•if 23 -* B re •o 85 i ^ ^

S-.SP ^ U. i S 8 "

os 5 B

k- -© © OS X

re ** X CL

CL = ©51- r ^ 0> I g E £ E L. >> 3 vs o ® 1 8 « 3 OS

re

x

© f—

as ^

CD * X

00

X

12 «

to. as © as

as

X

J 2

Cl vs o0)

as

as c

UO|JD*J|Q SutOOti »)»td

M%I

(A8w

365

•4UbU

-J

horizontal

portion

of

the

contour,

t/v

is

zero.

The

shear

stresses o

vz

andan.

are

ncpligihlc

in

comparison

with

the

applied

tensile

stress an„

so

the

only

FIG. 7—

Typical

specimen

used

far

experimental

evaluation

of

the J

integral.

Contribution

is

from

the

traction

term

E £

ii

£ £ to C.

i — §i - g

QJ O. — ^ r= ^ S

r- cj E 3

J E

at E

= wo

ac _• e ®

R

x5

£ 'O

•5 ^

•3 E

c o

CQ £

E ac E =

X "O 0 "o £ J2 » « £

6

QJ V

P p

.E re p *E5

o © £ £

2 *T 7 j p x

' « -J 2 © — . 8

i< £ !

*f?|

£ 5 = g f-5 t X f” E 2 N n

P E s — £ • x W fN ^ £ £ w ©

£> h c £

i X

1 x :J f

f— P R

~so

£ a £

2 * C c S g. g

w

II QJ *3

1

«C QJ

S

e •= = •=

< n

® o ^

C (9 — X >

ac p ps 3 O * £ « x

ii! ac £ p

3 *“ u

P QJ X £ x —

O QC C

.1-8 1

1||

"Ob X 8

88

E OC *3 CJ — e .2 x

S 'j t”

£ £

« re l~i s C.X *" 8

re * £ V

0 * o-l.

g

J8 E c

•2 “ fc* |

g ^ « 8

3 .£ 2 «

«o « r

-

8

2 £*g

gj « «

» g ^ do c g

= =

e.-e - « O

-* *3 u £ X g J r* ox •< c v J

o

QJ

Cl. V)CJ X

ac c

R QJ

fa c. £

®

R .* C R

C &

« . “"* o

at iL

ftj •»

• £ s g

V

<

i/o'

,*> o

IA

-O JD

~ * |l

- I

0 Z

I!cq c

> —

«- w

•&.a Is

- s

CQ

•a 1= x

c g

.E 2 c

R C

o. c

" R i/i

.2 *

x -g i I

51 =5

r I

£ 2 ^

2_w c

8JatR

ac

— o

. u

•o — 3 g

*> E

C

s ^

w •=»

** w o X

c

** R OC

~e —w £s O

» .g 3— X« © « s 5

E i

c S

Hi 85 •“ ? ^ x —

u x -2 u •- c % 2 S «

* * ^ w "S u> o> r CL S'ils

o « o

c , *o o u c •— 3 g

W>

k. t.

V X

X X

H J

= 2 « = „ X5 LL k

. s X J! — .£ & ~

_ fc =-^ =

o « *> u — *C

ac « *o

5 # C ^ CQ *> ^ Qj qj «L *5 x 1 .S * . 8 3

« x •£ S ac £ "K x H £? £ U. X! £

3 p " £ C X *D £

88 R ~

o- » c £ £i a-g

*. R X

QJC

Ii

g

5 -! « e

_ ? -a tlil

E

=

8 2 " "C >'> c p — « c

QJ P X •—

«e-clc e D

- £

ac o .E x

13*-

E. ^

K S

c

E

E

E

CQ&|

m sr CJ .s •? -

S R

«JS QJ £ .E 3 QJ p -

•5 w

p i/t r r

C. £

QJ •*-

l'£ « i

oc> C 2

c Cj c

•5 .§ ^ 1 e ^ -c g. s.«- _ QJ £

O X P w - 3

8 £ E

>, R «. T

3 i

2 rig

SI*

3 v v CL.> £

E — o R ** R X E __ —

" R E *c 5 o 8

1 p £

s

O- R = e.

QJ * £ "x

ii

» o X c o a O —

c n R X

cQJ r£

z .a

ac

_c

‘E

QJo.oXQJRl_O

Tsc

QJE£

QJC

UO

If- S. c

o ^r p .. X x

g H E "r

332 P 5 .« > ac >, " .£ s

j?© £QJ |QJ " X %

i f _. QJ *D 2 c 2. r .a p

ei

8 R

CQ CQ 4. W

x? ts

QJE ©

E

-0E

QJy U U E X E

QJ

(JE

c

QJ

fa R

tr

QJ w C

•3 X E E QJ O ,0 toac

QJ O P

ac u. s (I P cj C *3 0 0

O QJ * -c

o> X

*w ^

5

S £

>.X B. P § -O X *

QJ R C "O

E £ 3 R E E

J«•» 1

C "O — CJ ll

ffg.

.R QJ

g .E E w> p X

« 5

& 2

15 t

£ x

£ *

E - £ E

QJ R E

c O

P " |

.§

QJ T3

tk 3

*5

X R

£

m o

ac

°- II ^ Ee

£ £

R £

EL|

« I

3 2

e qj

QJ •e^ s R .2 c . P p «o EX

8. K

QJ

QJ

£

Cl z

3

E

» o X n

>

1/5 |

QJ £V

£ 2

ac «

R " X P

•8 H

UJ

E ^' c

; ji p x

H

jlE

U ^

c

j« x

a M E fc.Ce UJ C

cZ

i/5

^ ^

i/S i/S or)

c a ^

ia sc e

Nl ^ fS| - * -

fj

P

= E QJ *“ T5

QJ R £ &

•o ac £ = = £ - R R —

3

- X e i—

QJ ~

g_ ac Cl e

C P

£> 1

CQ CJ QC CQ E &

£ X R

CQ TO £ u - R QJ -O 2 S .£ 5

£ E « 2

x p X QJ R R C.X

QJ ac x

.c H

'5 . R x

2 R

r- C.

* C

p 2 ac .0 2 c £ -

QJ

U5 s

-* O u

K CQ *u X> i-o

0) -c

QC

CQocc

CQ < J

H

CQ= >g

o £

c —

R —

V, « £ 8 g r

s p

£

c

•o P

QJ R

|1

£ x

«il

t =

3 R

p E X QJ O R

-o CL 3 .a E o

accr

2

QJ rS&

E

o >%

•*.IA ^

g 3

O

E

2 £

qjc ^Cl

*0+

a*"

o xQJ x £ R R

C 3

51

L. QJ S £

Ifi C 3 JS

.a S

i 5 *3 1

R o 3 ?

CT © qj e:

.25 0

T3 « g &

c Qj

ac

2 ^

R

> c

QJ R |J S * iiw IQ s I

S « £ i 3 £

-rg

f

I

3 P

R J2 — D. ^ <ji p £

2

1 g 3 o- x

c 8 £ X 3 „ O X. 3

x E _ £

£

Cl O

E

QJ o U c CQ C

1 8

R U ^ X

o _

QJ O*-* R T3

c

© 8

E &

8 QJ .s £

*® w_ 8 JS to

CQ CQ or w o 2

E « ^

p £ _• 2-5 £

•©

QJ = x | 8

ac ^ c

c 2 qj | E £

p at o) J = 8 x X -2 p s 0

QJ » Jr qj 2 is

xs cl

qj TO l.r H E

E

OJCl

ac o

.E x

l -

Cl p £

— — E 3 —

- 2 =

R *3 g

R

QJ

E

o3

R

I 8 — »

QJ ~ X C *" QJ E E . o^ r

t £rt> Cl O

rf

1 •— >

ac *0 >J

X QJ i

. a «

O ^ CJ

•mmQ. . QJ 9 J£

Q> £

QJ C -O

*> CQ c C CL CQ QJ QJ _g .-s -O X M QJ

fc-L C W

8 P 3 X

R .E 2

« o _ D. «. _ a. c *s R QJ c E - < £ o 3 2 to &

oE £ x

o 2

QJ 8 “

•o 'C R 3

3 8

X E

£ ac n £ o _ R P C = R a _ c C CQ ^

o 'u "C

CQ ^ CJ -2 qj -r in

« 8

C R r

E sC“ P X R § * £

P c £ ^ J

0 0 3 ^ g S

•g s! R R •— 6 X

“ Jx

" S' « - a

O C E c » X 1 ® -

.3 °-£ 1 to R ^ P

c

QJ o

X 5P =

o c x

•2 2’5

5.1 E 2 = ^

« < -

§

X

E*SK g

£

S..E I

UJ •o — - Qj re - (- 5 Q C R X> la 1 iu

X -X

P

© *£= 3R

E “ O

p Sc x r u.

5r o • W ©£*

0 ~ «

3 X § o. BC 5

E £ e

.8c -J" =

E 1 £

QJ QJ C

x ac c H 2 p

. O .R S ~ QJ = £ -5

Cl R X

E © g

8 R £

'E &C *E .£ E

•o 3 £

p £ C ^ QJ "r r £

-§ x R-

* U) f S - 1 g£

1 Ije

o .P £

q.2 “

T3 R ^J S “

ac ? Jr at c 5 *

£ R p

p a. ~

2

0 >< 2-i

-0

QJ -S -o iil

•S-2-1 ^ CQ

QJ “ X R"

p u

8 «

£

2

x ” o fc .2 <£ 5. w Cl £

R 3

•5 3

1

.2 o. ll

CQ CQ » g 3

= x « ^ QJ a v w e -c o P ^ K o£ C C to - »£^ Cl •O

QJ ^*“

°

P k.3

8 2

x o

QJ U

« © 3 C £ 3

R P P R

i' QJ

X _ R

,E 2 6

_ Sb p

75 £ £

. C *T

QJ O

£ p

. R

•o E

£

o 2 X R

ac R r

.53 8-5

£ X R H 3 t

Q “ £

>J £ QJ o. 5

£ w <

p

*0 £ "3 l

0 2

§f«S

s R ^

> "O ta.

88 *3

QJ P QJ -o-o

R « 1

ft> *3 £ p

p X

.E E J. Z3 u.£ IE

3 O

2 2

o QJ C. .= B P

2 x

•3

QJ X

QJ 3

£

3

•3

QJg

CL

QJ

IA IA •cw w H o r

Qj QJ QJ

CQ =5 W

Cj CQ 7 >

*£, T3 I 2

I S ^ CQQJ .E E ^ J8 0 ^ ^ CQ c JZ

QJ —

E *9

JJ Q> 3 2 s i

1 ^ C QJ

>o x>

urU

CQ

Q>

(J

CQ •c3t/3 •9C

3

Ofc- CQ "5

fc= ac

QJ

E

QJ£

QJCl '

E =

o '5

QJ Sc p £

£ J

‘5 «J E 2

« ac ac .E

QJ C X ac •5 £

t_ c 3

*“ O QJ 3 £ c

V c p 3

£ .-5 J= o

qj ur *3 CJ *i r — *9 3 a I 2

§ &

*> ^ >

366

ExpeHmcnlal

Results

and

Discussion

*u-rt<3 ivjteiut- r I C

x.r ~ c •ri

'nco d -K n £ I. 3 — d c _ x

k n C

c 3

_C —

X X d

C *

1/1

d C x —

E « s . K d <*

s si

•1 I n p ac • &> t» j; c x £ x

X *d d

x -g £

Si ac u) g.s = S t-|

= l

a

d .cc M “ s a

IS3

.oXK

>

9

"3c

eU

x

5

£

E

E

9

C/5

w — r~ ^ ^ at) 1 H

x £

2

Mi £

d X

.£ &

-ed > X V — X

© C X aV

J£ X

I

> I •=

Q£> vs £ £

g-§

>

8PS x

± =

rea

T3aXItE

3

C

Xe

a

£

"TM g •« E S»' a « a n

c. re

i/>

*8

x « a £

re

ec x c a

2 5

-D

fi a £ £ a

re I "

£ 3 x c

i

a 1ee x

£ — I- ««

V

X **a

0£

©

> sr

e

"® 5 a

.2 a w E

k. J<_ >£ =

.£ 'Z

acc

'E

'rec

ca

5o E 15 x •? e o. & re V

c

_cX £

re © — x e

•? 1

x x

re a_ & ST v a £ o£ c = O X

•a g a X x J- C O. = V

t£ £

c

•3 >»

Cl X .* X

Si X re t;c.

re a a

u

*C S

5 5

c E 5 S. ©~ Xd

C X i£

t X

3V) •*. re < a ^

E

is. C

re —

X 1

— 3

u> C. W x x

c 2 re c o re x ap

Ji X

to o •Si

= Xw

« X

= X

si> d

£

E

= i

2 ^

U

i

=' d .2

d £J K 3

"*5 d

fec ir^r £ J

51

1/5 fij « £

, i> t- S E £

.- © is. i_ l|re w

ac .Si = ~

c ir S “T w o

K •

-C tr il

£ d

O’ X

x =

re 3

d' *

-E tf U ^

d —

S Xd X ^

'I fc G x

£ |

c i X d

.2 Z

d

£ U

x ^

C. d

re 3

c. •

*C J£ d t £ I d "re ^ —

e d

* c_ r 2 d

ac £ £

= '5 s « cc i

g fr-fe ^ X

•najit pti/dde tnua* f jof n/nt»j pju»>uu»dx» pup ruaiyipud ppoui fo uoruodwo^—f -Qy

• dN ‘SS 3 UXS »dH ‘S 63 UXS *dH ‘SS 3 UXS

i*p (•hj*uju*o»3 • Y

i*pouj i*d<»a^*uy — I®®IS OZ-XTS WV

367

This

work

was

supported

by

Che

U.SoQ 3

acu

£Xn

inv

.c

"3

c. cuo

Q. XuE

cioliS l X Cl

g * 5 V) E6 ^e x v

Cl 5 f 3 w c X

© ft— £ ^ To *E £5 CJ . tf:

u »© 5 o .= J=

HZ >

c. <0 E c 5 *K .2 ^

&) co

&>w •-

CO ^COG.

*a o « 6 ^1

•c -c c ^ n «/> £ ^ c * C5 i c

•2 8 « 3

Ji’g

cl i CL C n — E "2 x

«j £«

•O

C o —

I-: 6 S ^ <0 C 2

\Z *© •— S

•©

a i * | c

.o«Xo W8EmeX

v £ 2? ^ * c II ^ g | -| JS« X 8 8 £*3

CL

</> U)

a § |

.•= s Cn

ta H

a £2 ^ Cl

X

C 0

.a e® 3

x> * Z .E fc c

.=

— "O CL

a oc

Cl • H

*5e£ k .C_

— ~O Cl c 2 „ X D TM C 60 ft- •£

*•.c -o E g e ft ^

ft- 53

T* V Cl £

‘-5

H t

Cl

reE£a « E^ S 5s © w "O J- c

to CO

t> -o • -c •— ,2P ~ « LZc w Q |i|

SJ--S£ % U ? 2*d

.z

u c

TS •— “ n C . „d x g

ft c 5 Z T-£ e e g “ E 3 ftC 1 if*

"Salt c isi’O 2

*5 « c © 2 »i a . * 5*1 S ©X £ | 8 Q£

Cl = < 01 C B fi C X > T> “ C . OJ © w .& c ” S x s g

co «© t re f

-5 t ^ c « *? "o £ Sc®

"•go

VE I* ci "re lC •- ft £ gsi iic £ | CiS>i

•uc

co £

CO

CO JC

O"V©*

O’ W

•oc

CO «c

•c

©

II© 4-o

368

stress

at

which

net

section

yield

occurs

is

denoted

by

on$y

and

is

calculated

from

equilibrium:

etc « ^ cj “c fi. e |.§ — £ <« l|s -2 -ic .a

fc rex 8 £ *

Cl

C

C = W c

re .2

X > . t r c x ” re u S -X > OO'SL ^ c

»"> x: (i

SI § - £c.E u ez <* .£ oc 3

ins5 .E w

t> .2 -C

&!

C 2 = s C ure ^ e ~ * S ^j| */; c> *o c •» j: X S c®- > wi cr r '« = w £ “ re w £ • > c | 2 v .£

ss >>i— *c

wEg

Cl xAorreQ.

Ta

reo.

Xb > Cl •XT Cl 6 »

II •£ ^ iuE£ + c! 3

re £ 4> V X X H H

.£

75£ ^ j2

•c7; J5 75>

&>

i> 5 «® £ 2 e *|i« £ w

ai

a.

—

_ <C|?« W

i!g|

ric « I*

c.

— * °

fcg

x e I 8 c .,

cr =

u:' |x« - e

a. c ^

Si

w

Cl ^

Cl Cl

re

x

S-c

55 ?

oe

w

C "C IIc. E c ; #5* ~ o c x c

*- o

.£ cc ^ K £ 3 | 2 * X

!« eo X — o.«c o c x *o «o

*©

II

C t> 3 — u . •c « h o E.£ ^ i ? 1 5 C >* C O X

.2 c-g

re ci

C r=i c X Cl

•• 7 '•' x c 2

ac.£ kT

•-

— c 2 x s e _ c - c ~ c ^ cs C E t C V w

<5 2 t«

ffi c at n e

•g 1 | - < g

«: 2 R *o & I 5> LTl I

OC ^ .5 E

=>

‘| 2 * |2fi

«?•£ £ c — = C- t ^ - ta

C

x - .2 *21 1• * f

ii £ ^

(5

C—

C ^ 5 •; iti c - re *

•5 - c c ?-S <5< c X c ^ I S £ JtoS $ re ~ ®c | ^ t ^ c £

-< £ « 4C t C«Jl

** f,> c x s

•SB

•fecaI£ak8a c O •- • ,£ w *- 0* t r c x

r • c ? U x ^ s I i £ p |

s| I • cr< r p X ^ .= O'-riis o- Bc

of

a ? |di 5

i= °”o

c- ESI I £ “ i ;ll = l ' $ "C ^ il^ll ^ I c

.l

s.; s - ’i&Zi |

-a I a c i • t u ^ t

fNIv = < £

?.e

h

l. : u ^ £ ? .£ s

*5

X

y ^ 00, ! en

CD : • c • ^' Q - a : w •J.S 8 g

.E 4 jj

BE ire ^ S “K 5 ’^ I X c — • Ot CM c a a K &

pi j p

•» • «Ni Cf £ « re S - re E . Jt 5 >c « < <n ire z i£ t a =

*DSs

a. £ ^ s 7^ gU c x ° ^ u » x ? u u* < s j C e “ X

*C ^>

Eg ^ X• X 5 £ s sV w

.5 ^

<S- - x ~ r - .1 S S ' k at 2 u re ire = r* v c Z. 2: s t c .g

>-, " Cl •t - ~ g _ <

ire

— •

-• 8

“*

J*S at

T

—. i-J t — ; > n

as c. £ k

oc £ .5

aS x S

ire £U

ci

a c c c

re ire at : £ 2 £ S c r c c

(0 SC

Cl u. —

ci* I i £ £ c C £ C J£

re ci * ”

•£ ° g

ex .£ x £ ‘re 2X

"?» ClE

1> X

•o *«5. o

•DVEC “ |Sc u

.a t; 11 eo C X = N r e P’S " c

^.E &>

U

*“ -b ~ i o ^ .£ & £ c A — . K S Jg S-E£ - n w

4. 4» « “ i 3 : s

Xd.e £ ^ c X r- <q 4» e xs

QC T5s»5

75 2o , w

Uj

v3

CL

L I

oe&+

«c

_

w ? .2^2 § Pi I ! “ fe

Cl

ii o

.E -£ M vre S C

V> " o O

•c

>z:

•u ii C 3 « X

I!^ u

gg ^ ^ o w v = 5 « *3 S E4. — U

if^ >- # O X X fe 2 2 X ac ii

i!S ° X L> CB wC o 0> £ E <&t -6 t!

369

APPENDIX H - Metalloqraphic Analysis

H. I. McHenry

Metallographic examinations were carried out on sections representative

of the automatic gas-metal-arc (GMA) girth weld and the manual shielded-metal

arc (SMA) girth weld. As shown in Fig. 1, the automatic weld contains seven

passes, including a root pass, a hot pass, fill passes, and the reinforcement

pass. The weld is narrow with the fusion line lying generally normal to the

plate surfaces (except for the reinforcement pass). The base plate microstructure, shown in Fig. 2, is heavily banded and consists of alternate layers

of polygonal ferrite and nonpolygonal ferrite (acicular ferrite or low-carbon

bainite). The banded microstructure originating from chemical segregation in

the ingot is a result of the plate rolling operations prior to pipe

fabrication. The plate midthickness region (Fig. 2b) exhibits thicker

nonpolygonal ferrite layers than regions near the plate surfaces.

Ihe manual SMA weld, shown in Fig. 3, contains eight passes, including a

root pass, a hot pass, fill passes, and the reinforcement pass. Starting at

the root pass, successive weld passes become wider, emd the fusion line lies

at a substantial angle to the plate surfaces. A region containing a large

slag particle can be seen just below the reinforcement pass. Throughout the

cross section of the base-metal microstructure, shown in Fig. 4a, is a heavily

banded region (Fig. 4b) located at the plate midthickness. A large

delamination can be observed (Fig. 4) in the midthickness of the base plate.

This delamination extends from the weld fusion line through the heat-affected

zone (HAZ), terminating in the base metal (Fig. 5a) and occurs along the

boundary between the polygonal and nonpolygonal ferrite phases (Fig. 5b).

370

SLAG

PARTICLE

Fig. 1 - Profile section of automatic GMA girth weld in plate Wl-1. Cross

section shows base metal, HAZ, and weld metal. Plate rolling

direction is- . Etch: 5% nital. Magnification: 5X.

371

THROUGH-WALL

FLAW

DEPTH,

mm

AP1 1104 Allowable Flaw Size

Sumnary of system performance characteristics for vertical surface flaws.

372

Fig. 2 - Photomicrograph of base plate A from plate Wl-1. Light etching phase

is polygonal ferrite; dark etching is nonpolygonal ferrite. Plate

rolling direction is — . (a) Near outside plate surface.

(b) Near plate midthickness. Etch: 5% nital. Magnification: 200X.

373

. 3

- Profile section of

section shows base

direction is

manual SMA girth weld in plate W4-2.

metal, HAZ, and weld metal. Plate rol — Etch: 5% nital. Magnification:

Cross

1 ing

5X.

374

Fig. 4 - Photomicrographs of base plate B from plate W4-2.

Light etching phase is polygonal ferrite; dark

etching phase is nonpolygonal ferrite. Plate

rolling direction is-« . (a) Near outside

plate surface, (b) Near plate midthickness.

Etch: 57o nital. Magnification: 200X.

375

Manual SMA Weld, Plate W4-2

Fig. 6 - Profile schematics of automatic and manual girth welds.

377

Fig. 5 - Photomicrograph of base metal delamination in plate W4-2.

(a) End of delamination in base metal; magnification: 100X.

(b) Location of delamination along boundary between polygonal

ferrite and nonpolygonal ferrite phase; magnification: 500X.

Etch: 5% nital

.

Table 1 - Hardness Data Summary

Average HRA Values

Weld

Type

Base Metal HAZ* Weld

A Side B Side A Side B Side Metal

Automatic 57 (HRB 93) 57 (HRB 94) 62 (HRC 23) 62 (HRC 23) 65*5 (HRC 30)

Manual 57 (HRB 93) S6H (HRB 92) 60 (HRC 20) 60 (HRC 20) 56 (HRB 91)

*Converted from Vickers hardness numbers.

378

1 . Hardness Tests

Profile sections from each girth-weld type containing weld metal, HAZ,

and base metal were prepared for hardness surveys. Profile schematics of the

two types of girth welds identifying the weld-pass boundaries are shown in

Fig. 6. Vickers hardness measurements and Rockwell A Scale hardness measurements were taken throughout the base metal, HAZ, and selected individual weld

passes. The average hardness numbers were converted, where necessary, to the

Rockwell A Scale, HRA, for comparison. The hardness values for individual

weld passes are indicated in Figure 6 and the summary of all hardness results

are given in Table 1.

The average hardness of the HAZ in the automatic weld is intermediate,

lying between the lower base-metal hardness and the higher weld-metal

hardness. For the manual weld, the average HAZ hardness is somewhat higher

than the base-metal or weld-metal hardness. Furthermore, the average HAZ and

weld-metal hardness values for the automatic weld are higher than the average

values for the manual weld. These observations are consistent with the

observation that the automatic weld is generally narrower than the manual

weld, indicating either lower average heat input, higher average welding

rates, or a combination of both, resulting in less tempering of the automatic

weld HAZ and a higher cooling rate for the automatic weld metal.

2. Tensile Properties

The tensile properties of the weld metal from both weld types were

determined at -101°C and +24°C. The strength properties at intermediate

temperatures were determined by interpolation between the two temperatures.

Standard ASTM subsize round tensile specimens, 0.635-cm diameter with a

2.54-cm gage length, were prepared such that the reduced gage length section

contained only weld metal with the specimen axis parallel to the weld axis.

379

The gage-length section of specimens from the automatic weld generally sampled

weld passes 4 and 5, whereas specimens from the manual weld generally sampled

weld passes 5 and 6.

The tensile tests at -101°C were carried out inside an environmental

chamber in which the atmospheric temperature was controlled by balancing

cooling from liquid nitrogen and heating from electric resistance heaters.

The specimen temperature was measured with a thermocouple attached to the

gage-length section, and the test temperature was controlled to ±3°C during

the tests. The test machine crosshead speed was 0.13 mm/min up to yielding

and 0.76 mm/min after yielding. The tensile properties for each girth-weld

type and of the API 5LX-70 base material are summarized in Table 2.

The manual -weld test specimens were taken from two welds; the automaticweld test specimens were all taken from a single weld. The results of the

tensile tests for the manual-weld specimens from both plates at the two test

temperatures indicate good weld uniformity, and thus the measured tensile

properties can be used as representative of each weld type. The average

ultimate tensile strength and yield strength of the automatic weld increased

15% and 11%, respectively, as the test temperature was lowered from +24°C to

-101°C. For the manual weld, the average strength properties increased about

21% for the same change in test temperature. The tensile ductility

parameters, percent elongation and percent reduction in area, were not found

to be strongly temperature dependent for these two test temperatures.

3. Fracture Toughness

It was anticipated that these weldments would have sufficient plasticity

so that slow, stable crack extension would occur within the temperature range

of interest, and therefore, elastic-plastic fracture toughness parameters

would be needed for the flaw-size analyses. Crack-tip-opening displacement

380

v

Table

2

-

Tensile

Properties

of

the

Test

Materials

at

Room

Temperature

Material

Yield

Strength

Tensile

Ultimate

Elongation

Reduction

Notes

0.2Z

Offset,

Strength,

In

25

mm,

of

Area,

MPa

MPa

%

%

a

a>

0)

0)u

0)60

CO

t-> 0)><

ON vO

vO

SOvO

SO

CM CO

Csl CM

ON m o < m

m m 00

CO £

CM

COm

n

co>

aie

V

soCO X)

ON m

SO

•o- CM

«—

1

i-H O SO

CO V o U 1 4J > —I *H C/0 a* o 4-* O

e rH 00 a i"" CO W & u 0) 3 o

CO C CO 4-1 C/0 CO ce 3 9 2 GQ ac < < <

J=

4J60C

0)

l» 4J

00 •v

i—

I

0)

u

a>a

Cl. 3

0) X

H

«

381

(CTOD) tests were carried out over the temperature range of +24 to -101°C.

The temperature dependence of the fracture toughness was determined for the

weld metal and HAZ regions of both weld types. An estimate of the lowest

temperature at which reproducible, slow, stable crack growth would occur was

made from the temperature dependence data and a multiple specimen resistance

1 2 curve approach * and was used to determine the critical CTOD at the

initiation of stable crack growth.

All fracture tests were carried out in three-point bending on specimens

of the preferred geometry, as described in British Standard BS5762 : 1979 . ^ In

the weld-metal specimens, the machined notches were oriented parallel to the

weld axis and centered totally within the weld. For the HAZ specimens, the

machined notches were oriented parallel to the weld axis and centered so that

the final crack location after precracking would place the central portion of

the fatigue crack front in the HAZ.

The introduction of fatigue precracks into weld metal and HAZ specimens

presents a special problem because of the requirements on the straightness of

3 the fatigue crack front. Typically, the transverse residual stress pattern

associated with multipass welds in plates and pipe (e.g., from surface to

surface through the plate thickness or pipe-wall thickness) changes from

tensile stresses near each surface to a balancing compressive stress in the

interior. Fatigue crack growth in this stress environment often results in

two areas of growth near the surfaces in the residual tensile stress region

and little or no growth in the residual compressive stress region, as shown in

Fig. 7a. There is some evidence that measured fracture toughness properties

result from specimens exhibiting this bimodal fatigue crack shape are higher

than from specimens with a straighter crack shape, particularly in the elastic4 plastic toughness region.

382

b.

Fig. 7 - Fractographs of COD specimens illustrating the effect of compression

pretreatment on fatigue crack growth, (a) Without compression pretreatment; type E8010 weld metal; test temperature = +24°C (+75°F).

(b) With compression pretreatment; type E8010 weld metal; test

temperature = -45°C (-50°F). Magnification: 7X.

383

To reduce the effect of nonuniform fatigue cracks, the ligament region

ahead of the notch in all CTOD specimens was compressed about 1% following the

5 6

procedure of Dawes. * The actual load necessary to cause a 1% plastic strain

in the through-thickness direction was within 5% of the calculated load.

Several studies with weld metal specimens have shown little effect of the

4 5 compression treatment on the measured crack opening displacement. *

After the compression pretreatment, fatigue precracks were grown at 24°C

by sinusoidal loading of each specimen at a frequency of 20 Hz with maximum

and minimum loads of 6.7 kN and 0.1 kN. The final ratio of crack length to

specimen width was about 0.50. The average number of cycles required was

25,500 for the automatic-weld specimens with a range of 22,300 cycles to

29.000 cycles. For the HAZ specimens, the average number of cycles was 24,660

with a range of 21,000 cycles to 27,000 cycles. For the manual weld and HAZ

specimens, the averages were 21,130 cycles and 20,940 cycles, with ranges of

20.000 cycles to 22,500 cycles and 19,000 cycles to 22,000 cycles,

respectively. The shapes of the resulting fatigue cracks were generally

straight except near the surfaces, with little evidence of the bimodal

behavior, as can be seen in Fig. 7b.

The crack-opening-displacement procedure followed was based on British

Standard BS5762 : 1979 . Each specimen was loaded in displacement control at a

rate of 0.25 mm/min. The instantaneous load and displacement, as measured by

a clip gage attached to the specimen surface, were continuously recorded. The

temperature was monitored by a thermocouple attached to one of the screw-in

knife edges on the specimen. At the conclusion of each test, the specimen was

rapidly unloaded and heat-tinted to mark the extent of crack growth. Each

specimen was then cooled in liquid nitrogen and broken, and the amount of

crack extension was determined.

384

The CTOD value for each test was calculated using the relations:

K

2 (l - v

2

) 0 . 4 ( W - a )V

CTOD = + 2 (1)

2 a E 0.4W + 0.6(a + h ) and y c

K = YP/(tW*) (2)

where: Y = stress intensity coefficient at a/W

P = load

t = specimen thickness

W = specimen width

V = plastic component of clip-gage opening displacement

h£ = distance of clip gage from specimen surface

The interpretation of the calculated value of CTOD depends on observations

of the load-displacement behavior and the type and extent of crack growth at

the test temperature. The results of the analysis of the temperature dependence

of the fracture toughness are summarized in Tables 3 and 4 and Fig. 8. The

weld-metal data for both welding processes show a regular progression in crack

growth behavior from low test temperatures to high test temperatures: for

example, unstable or brittle fracture or pop-in without evidence of slow,

stable crack growth at the lowest temperatures through unstable brittle

fracture or pop-in after some slow, stable crack growth at intermediate

temperatures, to only slow, stable crack growth through the maximum applied

load at the highest temperatures. The higher-strength automatic weld did not

exhibit slow, stable crack growth until the test temperature was increased to

-18°C. The lower-strength manual weld exhibited slow, stable crack growth at

-46°C.

The HAZ data for both welding processes showed a similar regularity in

the crack growth behavior at the lowest and highest test temperatures: no

stable crack growth at the lowest temperature and only slow, stable crack

385

Table 3 - Temperature Dependence of Fracture Toughness for the

Automatic Weld

Specimen

Number Type*

Temperature,

°C

a/W CTOD

Criteriont

W133 W -73 0.5042 6 = 0.028

W135 W -46 0.5287 6 = 0.076

W128 w -18 0.5388 6 = 0.094

W1214 u -18 0.5215 6 = 0.089

u W129 w -18 0.5189 6 u = 0.079

W139 w 0 0.4971 6 = 0.084 max W138 w 21 0.5101 6 = 0.122 max

W134 w 21 0.5093 6may = °- 104 max

W1216 HAZ -101 0.5018 6 C = 0.018

W123 HAZ -101 0.4959 6 C

= 0.041

W121 HAZ -73 0.4854 6 C = 0.061

W1218 HAZ -18 0.5064 °- 437

W1212 HAZ 0 0.5133

«max

* °' 262

W1213 HAZ 0 0.5053 6 = 0.508 max W1210 HAZ 24 0.5080 6u = 0.442

max

* WM = weld metal

;

HAZ = heat-affected zone

t where : = CTOD at unstable fracture or onset of arrested brittle crac qrowth without evidence of slow crack growth.

6„ = CTOD at unstable fracture on onset of arrested brittle crac qrowth with evidence of slow crack qrowth.

6 max

= CTOD at maximum applied force.

386

Table 4 - Temperature Dependence of Fracture Toughness for the Manual Weld

Specimen

Number Type

* Temperature,

°C

a/W CTOD

Criteriont

W434 WM -101 0.5044 6 c = 0.010

W436 WM -101 0.5050 6 c

= 0.053

W431 WM -73 0.5143 6 c = 0.036

W432 WM -46 0.5086 s u

= 0.097

W439 WM -18 0.5200 6 u

= 0.084

W4117 WM -18 0.5033 6 max

= 0.229

W438 WM 0 0.5173 6 max

= 0.226

W433 WM +21 0.5054 6 max

= 0.236

W412 HAZ -101 0.5026 6c = 0.053

W4110 HAZ -101 0.5132 6C

= 0.018

W415 HAZ -73 0.5971 6c

= 0.114

W4114 HAZ -18 0.5074 6max

= 0.127

W4113 HAZ -18 0.5142 6c

= 0.076

W417 HAZ -18 0.5292 6u

= 0.155

W418 HAZ 0 0.5117 Su = 0.188

W4111 HAZ 0 0.5121 5max

= 0.432

W419 HAZ 24 0.5159 'Wx

= 0.249

W411 HAZ 21 0.5057

^nax

= 0.226

* WM = weld metal ; HAZ = heat- affected zone

T where: 6 = CTOD at unstable fracture or onset of arrested brittle crack

growth without evidence of slow crack growth.

= CTOD at unstable fracture on onset of arrested brittle crack

growth with evidence of slow crack growth.

fi max

= ^TOD at maxlmum applied force.

387

CTOD.mm

CTOD.mm

Fig. 8(a) - Temperature dependence of crack-opening displacement for the

automatic GMA girth weld.

Automatic Weld

Fig. 8(b) - Temperature dependence of crack-opening displacement for the

HAZ from the automatic GMA girth weld.

388

growth through maximum applied load at the highest temperatures. However,

considerable variability in crack growth behavior was observed at the intermediate temperatures, with less correlation between the presence or absence of

slow, stable crack growth and temperatures.

The greater variability in the HAZ results is probably associated with

the geometrical relationship between the weld and specimen notch orientation.

The welds are single-V type so that when the fatigue crack is grown, only the

central portion of the crack (about 25% of the crack front) lies in the HAZ.

The remainder of the fatigue crack lies in the weld metal on one side and the

plate base metal on the other side. Thus, the resulting or controlling

fracture toughness is variable because it depends on the detailed shape of the

fatigue crack, the resulting stress intensity due to the applied load, and the

expected difference in toughness in each of these three regions.

Using the temperature dependence of the CTOD toughness parameter, as

shown in Tables 3 and 4, an estimate was made of the lowest temperature at

which consistent and reproducible slow, stable crack growth would occur. At

this temperature, arrested brittle crack growth or pop-in would not be

expected. These temperatures were estimated to be 0°C for the automatic weld

and -18°C for the manual weld, and the temperatures for each HAZ were

estimated to be -18°C and 0°C, respectively.

A multiple-specimen resistance curve approach was used to determine the

critical CTOD, 6 .., at the initiation of stable crack growth. The crackopening-displacement procedure followed was based on Appendix A of British

Standard BS5762: 1979.* Each fatigue precracked specimen was loaded in displacement control at a rate 0.25 mm/min to various values of clip-gage

displacement to obtain various amounts of crack, 6, for each increment of

slow, stable crack growth. Then each specimen was rapidly unloaded and heat389

tinted to mark the extent of stable crack growth. Each specimen was cooled in

liquid nitrogen, fractured, and the amount of crack extension measured. The

results of these tests are summarized in Tables 5 and 6 and Figs. 9a-d. The

6.j representative of the initiation of stable crack growth was taken as the

value of crack-tip-opening displacement at zero crack extensions, Aa = 0. The

values of 6^ for the automatic and manual girth welds are 0.074 mm at 0°C and

0.079 mm at -18°C, respectively, and the HAZ values of 6^ are 0.14 mm at -18°C

and 0.17 mm at 0°C, respectively.

The measured 6 of 0.083 mm at 0°C for the automatic weld is consistent

max

with the calculated 6.. of 0.074 mm because of the lack of a strong temperature

dependence of the CT0D Fig. 8a and the very shallow slope of the

CTOD-vs. -stable crack extension curve Fig. 9a. The 6 of 0.023 mm at -18°C

for the manual weld is consistent with the calculated 6^ of 0.079 mm because

of the stronger temperature dependence of the CT0D (Fig. 8b) and the steeper

slope of the CTOD-vs. -crack extension curve (Fig. 9b). Although considerably

greater scatter was observed in the HAZ data, the same general relationship

was observed when comparing the calculated 6^ to <$ max

*

390

Table 5 - Critical CTOD for Initiation of Slow, Stable Crack Growth

for the Automatic Weld

Specimen

Number Type

* Temp.,

°C

P,

kN a/W

Aa, mm

CTOD, mm

W122 WM 0 17.1 0.5018 0.431 0.076

WI26 WM 0 18.1 0.5006 0.356 0.097

W124 WM 0 17.4 0.5060 0.660 0.099

W1217 WM 0 17.5 0.5041 0.533 0.107

W1215 HAZ -18 17.9 0.5081 0.152 0.109

W1310 HAZ -18 18.9 0.5059 0.086 0.163

W1311 HAZ -18 18.3 0.5113 0.401 0.190

W1312 HAZ -18 19.4 0.5054 0.142 0.196

* WM = weld metal

;

HAZ = heat- affected zone

9

t

391

Table 6 - Critical CTOD for Initiation of Slow, Stable Crack Growth

for the Manual Weld

Specimen

Number Type*

Temp. , °CP,

kN a/W

4a

,

mm

CTOD,

mm

W4119 WM -18 13.92 0.5092 0.066 0.094

W4118 WM -18 14.41 0.5042 0.122 0.119

W4116 WM -18 14.54 0.5137 0.277 0.142

W4I3 WM -18 14.52 0.5079 0.246 0.167

W414 HAZ -18 14.12 0.5140 0.361 0.185

W4120 HAZ 0 14.14 0.5117 0.129 0.150

W4112 HAZ 0 14.79 0.5169 1.85 0.178

W4I15 HAZ 0 15.50 0.5131 0,. 081 0.231

W4121 HAZ 0 16.06 0.5087 0.193 0.300

* WM = weld metal

;

HAZ = heat- affected zone

t

392

CTOD,

mm

Automatic Weld

Fig. 9(a) - Crack opening displacement for the initiation of slow,

stable crack growth for the automatic GMA girth weld

at 0°C (32°F).

393

CTOD,

mm

Manual Weld

Fig. 9(b) - Crack-opening displacement for the initiation of slow, stable

crack growth for the manual SMA girth weld.

394

CTOD,

mm

Manual Weld

Fig. 9(c) - Temperature dependence of crack-opening displacement for the weld metal

in the manual SMA girth weld.

395

CTOD,

mm

i

Manual Weld

HAZ

Fig. 9(d) - Temperature dependence of

from the manual SMA girth

crack-opening displacement for the HAZ

weld.

395

4. References

1. "Methods for Crack Opening Displacement (COD) Testing," BS5762 British

Standards Institution, London (1979).

2. J. D. Harrison, "The State-of-the-Art in Crack Tip Opening Displacement

(CTOD) Testing and Analysis," Report No. 7302.10/80/210.2, The Welding

Institute, Cambridge, England (1980).

3. G. Kaufman, "Developments in Fracture Mechanics Test Methods Standardization" ASTM STP 632, American Society for Testing and Materials,

Philadelphia, 11- (1977).

4. H. G. Pisarski and M. G. Dawes, "Measurement of Crack Opening Displacement in Weldments with Particular Reference to Offshore Structures,"

Fracture Mechanics , Rome, Italy (June 1980).

5. M. G. Dawes, "Fatigue Precracking Weldment Fracture Mechanics Specimens,

Met. Constr. and Brit. Wei d. J. , No. 2, 61-65 (1971).

6. M. G. Dawes, "Contemporary Measurement of Weld Metal Fracture Toughness,

Weld. J. 55, 1052-1057 (1976).

397

;.n4A (REV- 2-»C) U.s. DEPT. OF COMM.

BIBLIOGRAPHIC DATA

SHEET (See instructions)

1. PUBLICATION OR

REPORT NO.

2. Performing Organ. Report No.

NBSTR 83-1695

TITLE AND SUBTITLE

Fitness-for-Service Criteria For Pipeline Girth Weld Quality

AUTHOR(S)

R p Rp P H ] m R , Kasen , id

I—McHe.pry,. and-X—hL-Eontunko

PERFORMING ORGANIZATION (If joint or other than NBS. see instructions)

NATIONAL bureau of standards

DEPARTMENT OF COMMERCE

WASHINGTON, D.C. 20234

7. Contract/Grant No.

8. Type of Report & Period Covered

SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street. City, State, ZIP)

United States Department of Transportation

Research and Special Programs Administraton

Office of Pipeline Safety Regulation

SUPPLEMENTARY NOTES

|~

"1 Document describes a computer program; SF-185, FIPS Software Summary, is attached. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bibliography or literature survey, mention it here)

Criteria have been developed for applying f i tness-for-service analyses to flaws in

e girth welds of the Alaska Natural Gas Transmission System pipeline. A critical

ack-opening-displacement elastic-plastic fracture mechanics model was developed and

perimentally verified. Procedures for constructing flaw acceptance criteria curves

sed on this model are provided. A significantly improved ultrasonic method for

tecting and dimensioning significant weld flaws was developed and demonstrated on peline sections. The probability of crack initiation from blunt flaws was shown to

very low under severe low-cycle fatigue. Suggestions are offered for technical

plementation of field inspection procedures and for practical implementation of the

aw acceptance criteria.

KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolon s) Elastic-plastic fracture mechanics; fi tness-for-service; nondestructive

evaluation; pipeline; pipeline steel; weldment flaws

AVAILABILITY

[^1 Unlimited

| | For Official Distribution. Do Not Release to NTIS

j | Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C.

20402.

| | Order From National Technical Information Service (NTIS), Springfield, VA. 22161

14. NO. OF

PRINTED PAGES

15. Price

uscomm-dc e043-peo
. — - 7 /