A Comprehensive Update in the Evaluation of Pipeline Weld ......Girth weld defect acceptance criteria are set and enforced in all pipeline constructions in the U.S. per federal regulations

DRAFT

A Comprehensive Update in the Evaluation of Pipeline Weld Defects

U.S. DOT Agreement No. DTRS56-03-T-0008

PRCI Contact No. PR-276-04503

Authors: Yong-Yi Wang and Ming Liu

Engineering Mechanics Corporation of Columbus

3518 Riverside Dr., Suite 202 Columbus, OH 43221

Publication Date:

November 2004

For internal circulation within PRCI, DOT, and API

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03-G78-20 ii

Table of Contents List of Figures .................................................................................................................................................v Executive Summary..................................................................................................................................... viii 1.0 Introduction..........................................................................................................................................1

1.1 Background ...........................................................................................................................1 1.2 Scope.....................................................................................................................................2 1.3 Structure of the Defect Assessment Procedures....................................................................2 1.4 Terminology and Notation ....................................................................................................3

1.4.1 Pipe Properties ....................................................................................................................3 1.4.2 Girth Weld Properties .........................................................................................................3 1.4.3 Applied Loads.....................................................................................................................3 1.4.4 Defect Dimensions..............................................................................................................3

2.0 Level 1 Assessment Procedures...........................................................................................................4 2.1 Overview...............................................................................................................................4 2.2 Level 1 Option 1 Assessment................................................................................................4

2.2.1 Additional Requirements ....................................................................................................4 2.2.2 Acceptance Criteria.............................................................................................................5 2.2.3 Computation of the Load Level Pr .....................................................................................7

2.3 Level 1 Option 2 Assessment................................................................................................8 2.3.1 Overview.............................................................................................................................8 2.3.2 Additional Requirements ....................................................................................................8 2.3.3 Determination of the Key Components in the FAD Procedure ........................................10 2.3.4 Defect Acceptance Criteria ...............................................................................................11

2.4 Limitations of the Level 1 Procedures ................................................................................12 3.0 Level 2 Assessment Procedures.........................................................................................................13 4.0 References..........................................................................................................................................14 Appendix A Validation of the Assessment Procedures against Full-Scale Bend Tests ............................ A-1

A.1 Background ...................................................................................................................... A-2 A.2 Experimental Database for Validation ............................................................................. A-2 A.3 Validation Process............................................................................................................ A-3 A.4 Results of the Validation .................................................................................................. A-3 A.5 Observation from the Validation against Full-scale Bend Tests ...................................... A-4 A.6 References ........................................................................................................................ A-7

Appendix B Validation of the Assessment Procedures against Curved Wide Plate Tests .........................B-1 B.1 Background .......................................................................................................................B-2 B.2 Overview of the Wide Plate Tests.....................................................................................B-2 B.3 Validation Process.............................................................................................................B-2 B.4 Validation Results against Curved Wide Plate Test Data .................................................B-3 B.5 Observation from the Validation against Curved Wide Plate Test Data...........................B-3 B.6 References .........................................................................................................................B-7

Appendix C Stress Intensity Factor Solution .............................................................................................C-1 C.1 Background .......................................................................................................................C-2 C.2 Parametric Equations.........................................................................................................C-3 C.3 Comparison between Fitted Equations and the FE Results...............................................C-3 C.4 References .........................................................................................................................C-5

Appendix D Plastic Collapse Solution ...................................................................................................... D-1 D.1 Background ...................................................................................................................... D-2 D.2 New Defect Size Correction Factor ................................................................................. D-2 D.3 Comparison with Full-scale Test Data ............................................................................. D-3 D.4 References ........................................................................................................................ D-4

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Appendix E Estimation of Applied Stress from Applied Strain.................................................................E-1 E.1 Assumed Stress Strain Relations.......................................................................................E-2 E.2 Estimation of Strain Hardening Exponent ........................................................................E-2 E.3 Estimation of Y/T Ratio from Pipe Grade or Yield Stress.................................................E-2 E.4 Estimation of Uniform Strain ............................................................................................E-4 E.5 References .........................................................................................................................E-4

Appendix F Incorporation of Weld Strength Mismatch............................................................................. F-1 F.1 Background ....................................................................................................................... F-2 F.2 Determination of Weld Width for Girth Weld .................................................................. F-2 F.3 Suggested Approach for the Treatment of Weld Strength Mismatch ............................... F-3 F.4 References ......................................................................................................................... F-3

Appendix G Example Problem for a Level 1 Option 2 Assessment ......................................................... G-1 G.1 Background ...................................................................................................................... G-2 G.2 Input Data......................................................................................................................... G-2 G.3 Steps to Derive the Defect Acceptance Level .................................................................. G-2 G.4 Comments and Observations............................................................................................ G-5

Appendix H Comparison of Acceptance Criteria...................................................................................... H-1 H.1 Background ...................................................................................................................... H-2 H.2 Comparison of Acceptance Criteria ................................................................................. H-2

Appendix I Limits of Applicability of the Current API 1104 Appendix A Acceptance Criteria ................ I-1 I.1 Background of API 1104 Appendix A ................................................................................... I-2 I.2 Appendix A from the Perspective of the Code Structure ....................................................... I-2 I.3 Limits of Applicability from Analytical and Experimental Work Funded by API ................ I-3 I.4 Limits of Applicability from 1980’s Work Funded by DOT.................................................. I-4 I.5 Observation from Historical and More Recent Work............................................................. I-6 I.6 Limits of Applicability of the Current API 1104 Appendix A Acceptance Criteria .............. I-7 I.7 Recommendation about the Limits of Applicability of API 1104 Appendix A ..................... I-7 I.8 References .............................................................................................................................. I-9

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List of Figures Figure 1-1 Dimension of a pipe with a surface-breaking defect.........................................3 Figure 2-1 Level 1 Option 1 defect acceptance level at various applied load

levels for CTOD toughness equal to or greater than 0.25 mm (0.010 inch) ..................................................................................................................6

Figure 2-2 Level 1 Option 1 defect acceptance level at various applied load levels for CTOD equal to or greater than 0.10 mm (0.004 inch) and less than 0.25 mm (0.010 inch).........................................................................6

Figure 2-3 Relation between Y/T ratio and pipe grade of Eq. (1) .....................................7 Figure 2-4 Schematic overview of the Level 1 Option 2 procedure ..................................9 Figure A-1 Full-scale test data plotted on the FAD of the Level 1 Option 2

procedure. The nominal SMYS were used as the strength input................ A-5 Figure A-2 Full-scale test data plotted on the FAD of the Level 1 Option 2

procedure. The measured yield stresses were used as the strength input. ............................................................................................................ A-5

Figure A-3 Full-scale test data that fall within the defect size and CTOD limitations plotted on the FAD of the Level 1 Option 2 procedure. The nominal SMYS were used as the strength input................................... A-6

Figure A-4 Full-scale test data that fall within the defect size and CTOD limitations plotted on the FAD of the Level 1 Option 2 procedure. The measured yield stresses were used as the strength input. ..................... A-6

Figure B-1 Curved wide plate test data plotted on the FAD of the Level 1 Option 2 procedure. The nominal SMYS were used as the pipe strength input. .............................................................................................................B-4

Figure B-2 Curved wide plate test data plotted on the FAD of the Level 1 Option 2 procedure. The measured yield and tensile strength of the pipe were used as the pipe strength input......................................................................B-5

Figure B-3 Curved wide plate test data plotted on the FAD of the procedure that is an extension of the Level 1 Option 2 procedure. The measured yield and tensile strength of the pipe AND weld were used as the strength input. ...............................................................................................B-5

Figure C-1 Comparison of the fitted equations as a function of defect depth ratio ........C-2 Figure C-2 Comparison of the fitted curves with the original data of Chapuloit for

pipes with D/t of 42. The symbols are from the original data and the curves are from the fitted equations. ............................................................C-4

Figure C-3 Comparison of the fitted curves with the original data of Chapuloit for pipes with D/t of 82. The symbols are from the original data and the curves are from the fitted equations. ............................................................C-4

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Figure C-4 Comparison of the fitted curves with the original data of Chapuloit for pipes with D/t of 162. The symbols are from the original data and the curves are from the fitted equations. ............................................................C-5

Figure D-1 Comparison of the Miller plastic collapse solution and the full-scale test data. The prior and current defect size correction lines are shown...... D-3

Figure E-1 Comparison of linepipe longitudinal test data with the Webster and Bannister correlation equations [2]...............................................................E-3

Figure E-2 Comparison of the relations between Y/T ratio and pipe grades from estimation equations and codes ....................................................................E-3

Figure F-1 Determination of weld width 2H for a typical girth weld geometry ............F-2 Figure G-1 Defect acceptance curve from the example problem with the Level 1

Option 2 procedure ...................................................................................... G-6 Figure G-2 Illustration of the critical points on the failure assessment curve from

the example problem. Points 8 and 9 are on the cut-off line, therefore the acceptable defect sizes for those points are toughness independent.................................................................................................. G-6

Figure G-3 The upper right corner of Figure G-2 .......................................................... G-7 Figure H-1 Comparison of the defect acceptance levels from API 1104 Appendix

A and those of the current procedures with no safety factor on the allowable defect length ................................................................................ H-4

Figure H-2 Comparison of the defect acceptance levels from API 1104 Appendix A and those of the current Level 1 procedures with the recommended safety factor on the allowable defect length ................................................ H-4


Figure H-4 Comparison of the defect acceptance levels from the current Level 1 Option 1 and Option 2 ................................................................................. H-5

Figure H-5 Comparison of the defect acceptance levels from API 1104 Appendix A and those of the current procedures with no safety factor on the allowable defect length ................................................................................ H-6



Figure H-8 Comparison of the defect acceptance levels from the current Level 1 Option 1 and Option 2 ................................................................................. H-7

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Figure I-1 Comparison of allowable flaw size between API 1104 Appendix A and the NBS criteria with an assumed CTOD toughness of 0.005 inch (0.127 mm) ....................................................................................................I-5

Figure I-2 Comparison of allowable flaw size between API 1104 Appendix A and the NBS criteria with an assumed CTOD toughness of 0.010 inch (0.254 mm) ....................................................................................................I-5

Figure I-3 Comparison of allowable defect size among various codes and procedures......................................................................................................I-6

Figure I-4 Comparison of the defect acceptance criteria from the current plastic collapse solution with no safety factor and those of API 1104 Appendix A....................................................................................................I-8

Figure I-5 Comparison of the defect acceptance criteria from the current plastic collapse solution with a safety factor of 1.5 on the defect length and those of API 1104 Appendix A .....................................................................I-9

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A Comprehensive Update in the Evaluation of Pipeline Weld Defects

Executive Summary Girth weld defect acceptance criteria are set and enforced in all pipeline constructions

in the U.S. per federal regulations (CFR 49 Parts 192 and 195). With the increased use of mechanized welding and AUT (Automated Ultrasonic Testing) in new pipeline constructions, alternative defect acceptance criteria based on ECA (Engineering Critical Assessment) principles are frequently used in lieu of the traditional workmanship criteria. The alternative defect acceptance criteria in the current Appendix A of API 1104 have remained largely unchanged since its introduction in the early 1980’s. In the meantime, the characteristics of the linepipe materials, welding processes, and construction practice have evolved since the adoption of the code. The recent surge in the use of mechanized welding/AUT/ECA created a mismatch between the new materials/welding processes and the outdated alternative defect acceptance criteria. Looking ahead, the trend in pipeline construction is moving towards larger diameter and higher strength linepipes, such as X80, X100, and even X120. The characteristic of these ultra-high strength materials and their welding processes make the use of the current Appendix A highly questionable.

This report presents the girth weld defect assessment procedures for stress-based design. The major components of this report are (1) technical basis for the development of the revised girth weld defect acceptance criteria, (2) validation of the acceptance criteria against experimental test data, and (3) recommended structure for the revision of API 1104 Appendix A. The main body of the report is written in such a way that it can be easily turned into code language. The supporting data, both analytical and experimental, are given in the appendices. Examples are given to show the use of the new assessment procedures. Comparisons in defect acceptance criteria are made between the new procedures and the current API 1104 Appendix A.

The new proposed procedures have two options. Option 1 is given as an easy to use graphical approach, whereby allowable flaw dimensions can be determined on the basis of a somewhat more restrictive minimum toughness level. Option 2 provides more flexibility and generally allows larger flaws, at the expense of more complicated calculations.

In comparison to the current API 1104 Appendix A, the major advantages of the newly proposed procedures are:

• Consistent level of conservatism • Inclusion of both plastic collapse and fracture criteria. The current API 1104

Appendix A includes only fracture criterion. • The acceptance criteria are easier to use for the most frequently occurring defects

in modern pipeline construction. • Reduced minimum CTOD toughness requirements, accompanied by tighter defect

tolerance, allows wider application of the alternative acceptance criteria.

03-G78-20 viii

1.0 Introduction

1.1 Background

Girth welds made in field welding conditions often contain some “imperfections.” These imperfections are sometimes referred to as “defects.” Many of these “defects” are a natural occurrence of the field welding processes. Traditionally, the tolerable defect sizes are set by workmanship-based criteria. One of the earliest and perhaps the most widely recognized workmanship criteria is that given in the main body of API Standard 1104 [1]. These criteria are empirically-based and historically proven safe in practice. In most cases, they are not quantitatively related to the severity of the defects in safely maintaining the operation of the pipelines.

Beginning in the late 1970’s and early 1980’s, “alternative defect acceptance criteria” have been implemented in various codes and standards. These criteria are based on fracture mechanics principles. They relate the tolerable defect size with the magnitude of loading in pipelines and material’s resistance to failure. When correctly used, these criteria allow engineers to assess the suitability of the pipelines containing the defects for intended service conditions, or fitness-for-service (FFS). Assessment based on the FFS principles is alternatively referred to as Engineering Critical Assessment, or ECA. The ECA codes that are most frequently used in the North American pipeline industry are API 1104 Appendix A [1], CSA Z662 Appendix K [2], and BS7910:1999 [3]. Although certain parts of API 1104 Appendix A and CSA Z662 Appendix K had their root in PD6493:1980 [4], the defect acceptance criteria vary significantly. The PD 6493:1980 was succeeded by PD 6493:1991 [5] and more recently by BS 7910:1999. A more complete review of the evolution of pipeline ECA procedures is given in Reference [6].

This document takes advantage of the significant progress made in understanding the girth weld behavior over the last two decades. Significantly, the following elements form the basis of this document.

• Historical and recent experimental data, from small specimen to full-scale tests,

• Fundamental fracture mechanics principles as implemented in weld defect assessment procedures for engineering structures,

• Recently published and/or updated pipeline codes around the world, such as EPRG guidelines [7], Australian Standard AS2885 [8], API RP 579 [9], CSA Z662 2003 Edition [2], etc.

• State-of-art research in weld defect assessment, such as the SINTAP procedure [10] and the PRCI GWIS procedure [11,12].

This document presents procedures for the assessment of defects in transmission pipeline girth welds. The assessment procedures are simplified, whenever possible, to address the

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specific needs of this industry without sacrificing the necessary consistency and accuracy of the procedures.

Unless otherwise specified, the defects or imperfections here refer to planar defects.

1.2 Scope

This use of these assessment procedures is restricted to the following conditions:

• Girth welds between pipes of equal wall thickness

• No onerous fatigue crack growth in construction and under service conditions

• No sub-critical crack growth, such as creep and environmentally-assisted crack growth

• No dynamic loading

1.3 Structure of the Defect Assessment Procedures

The assessment procedures is structured in two levels. Level 1 is for stress-based design and Level 2 is for strain-based design. There are two options in each level. Option 1 is the simplified procedure. Option 2 allows for broader applications than Option 1, but at the expense of more complex computation and/or more required input data. The major characteristics of the assessment procedures are given in Table 1.

Table 1 Structure of the Proposed Assessment Procedures

Diameter Grade and Tensile Property

Longitudinal Load

1

Plastic collapse criterion, corrected by the Option 2 procedure

Graphical format. Minimal calculation required.

2

Failure assessment diagram (FAD) format. Extensively updated from the PRCI GWIS procedure

Allow the assessment of brittle fracture, plastic collapse, and the interaction of the two failure modes in a single consistent format. Ability to accommodate new features, such as weld strength mismatch, welds between pipes of unequal wall thickness, if desired.

1 Current work Graphical or tabular format that covers majority of applications

2 Current workComplex multi-variable format, may require computational software for easy application

Feature

No limit1

2 No limit

Range of Applicability

TBD

Applied stress ≤ SMYS.

Applied strain ≤0.50%.

Applied strain > 0.5%

Ass

essm

ent

Leve

l

Basis

Opt

ion

Test data available up to X100

2

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1.4 Terminology and Notation

1.4.1 Pipe Properties D = pipe outer diameter, inch or mm R = pipe outer radius, R ≡ D/2, inch or mm t = pipe wall thickness, inch or mm α = pipe diameter to wall thickness ratio, α ≡ D/t σy, Y = specified minimum yield stress of the pipe material, or SMYS, ksi or MPa σt, T = ultimate tensile strength of the pipe material, or UTS, ksi or MPa σf = flow stress of the pipe material, σf = (σy+σt)/2, ksi or MPa E = Young’s modulus υ = Poisson’s ratio

1.4.2 Girth Weld Properties

1.4.3 Applied Loads σa = applied longitudinal stress, ksi or MPa Pr = normalized applied stress or load level, Pr ≡ σa / σf

1.4.4 Defect Dimensions a = defect depth, inch or mm c = defect half length, inch or mm β = ratio of defect length to pipe circumference, β ≡ c/πR = 2c/πD, η = defect height to crack-depth ratio, η ≡ a/t

a

2c

a

2c

Figure 1-1 Dimension of a pipe with a surface-breaking defect

3

2.0 Level 1 Assessment Procedures

2.1 Overview

The Level 1 assessment procedures represent the state-of-art understanding in the assessment of the significance of pipeline girth weld defects under stress-based design, normally defined as applied longitudinal strain less than 0.5% and applied longitudinal stress less than the specified minimum yield stress of the pipe material. There are two options at this assessment level.

Option 1 is a simplified approach in graphical format. It relies on theoretically sound and experimentally validated plastic collapse criterion when the fracture toughness is sufficiently high. The criterion is modified by the Option 2 approach when the fracture toughness is lower, but sufficiently high to avoid brittle fracture. The Option 1 approach is based on the premises that modern pipeline steels joined using modern welding procedures and consumables usually produce girth welds with good toughness. Consequently, brittle fracture is usually not a concern. The defect acceptance level can be derived from a suitable plastic collapse criterion, provided that certain minimum toughness requirements are met. A notable example of this philosophy is the EPRG Guideline [7].

Option 2 is in the form of a failure assessment diagram, or FAD, which was first proposed in the mid-1970’s [13]. The FAD format has become by far the most widely used defect assessment procedure for a wide range of industries, from the petroleum refining industry (API RP 579 [9]) to the nuclear power generation industry (R6 [14] and ASME Section XI [15]). The FAD format allows the simultaneous consideration of brittle fracture, plastic collapse, and the interaction between those two failure modes (elastic-plastic fracture). The FAD approach is considerably more complex in computation. Furthermore, some proficiency and understanding of fracture mechanics is necessary to ensure the procedure is applied correctly. However, validated computer programs, either from commercial market or developed internally, should greatly facilitate the assessment process.

2.2 Level 1 Option 1 Assessment

2.2.1 Additional Requirements

In addition to the requirements of Section 1.2, the following requirements are necessary at the minimum design temperature.

1. Weld metal strength even- or over-matches that of pipe material

2. No failure in the HAZ (heat-affect-zone) when defect-free welds are tested

3. Applied longitudinal stress no greater than SMYS and the applied longitudinal strain no greater than 0.5%

4. The minimum CTOD toughness no less than 0.10 mm (0.004 inch)

03-G78-20 4

5. The minimum and averaged Charpy values are greater than 30 J and 40 J, respectively.

The requirements of 1 and 2 are considered met if the cross-weld API tensile specimens do not break in the weld or HAZ. The weld reinforcement on both sides of the specimen shall be removed for such tests.

2.2.2 Acceptance Criteria

Two sets of acceptance criteria are given, depending on the fracture toughness of the materials.

When the CTOD toughness is equal to or greater than approximately 0.25 mm (0.010 inch), the critical defect size is largely independent of toughness value. The defect acceptance level is given in Figure 2-1 at various levels (Pr). This acceptance level is derived from the plastic collapse criteria given in Appendix D, with a safety factor of 1.5 on the defect length. If a load level is not given in Figure 2-1, the acceptance level can be obtained by interpolating the adjacent curves or by taking the value of the next higher load level.

At a CTOD toughness equal to or greater than 0.10 mm (0.004 inch) and less than 0.25 mm (0.010 inch), the critical defect size is dependent on toughness values for deep defects, but fully plastic-collapse-controlled for shallow defects. Some examples of this defect depth dependence are shown in Appendix H. The defect acceptance level given in Figure 2-2 is calibrated to a CTOD toughness level of 0.10 mm (0.004 inch). The safety factor on the defect length is approximately 1.5 at the toughness level of 0.10 (0.004 inch), but higher at higher toughness levels.

The total defect length shall be no greater than 12.5% of the pipe circumference. The maximum defect height shall be no greater than 50% of the pipe wall thickness.

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0.0

0.1

0.2

0.3

0.4

0.5

0.000 0.025 0.050 0.075 0.100 0.125Allowable Length / Pipe Circumference

Allo

wab

le H

eigh

t / P

ipe

W. T

.

0.7250.750

0.7750.800

0.825

0.8500.875

0.900

0.9250.950

0.975

P r=0.700

Figure 2-1 Level 1 Option 1 defect acceptance level at various applied load levels for

CTOD toughness equal to or greater than 0.25 mm (0.010 inch)

0.0

0.1

0.2

0.3

0.4

0.5


Allo

wab

le H

eigh

t / P

ipe

W. T

.

P r=0.725

0.7500.775

0.8000.825

0.8500.875

0.9000.925

0.9500.975

0.700

0.5500.6750.650

0.625

0.600

0.575

Figure 2-2 Level 1 Option 1 defect acceptance level at various applied load levels for

CTOD equal to or greater than 0.10 mm (0.004 inch) and less than 0.25 mm (0.010 inch)

6

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2.2.3 Computation of the Load Level Pr

In accordance with the definition given in Section 1.4.3, it is necessary to determine material’s flow stress in order to obtain the load level Pr. The yield to tensile (Y/T) ratio of the pipe material for a given grade is estimated as,

25.275.2121

1/

⎟⎟⎠

⎞⎜⎜⎝

⎛+

=

y

TY

σ

, (1)

where the pipe grade, , is in the unit of ksi. Alternatively, the Y/T ratio may be obtained from Figure 2-3. The background of the Y/T ratio and pipe grade relation is given in Appendix E.

yσ

The flow stress is therefore computed as,

⎟⎠⎞

⎜⎝⎛ +=

TYyf /115.0 σσ , (2)

The load level, Pr, is given as,

f

arP

σσ

= . (3)

The applied longitudinal stress, σa, is obtained from stress analysis.

0.7

0.8

0.9

1.0

50 60 70 80 90 100 110 120

Grade (ksi)

Y/T

Figure 2-3 Relation between Y/T ratio and pipe grade of Eq. (1)

7

2.3 Level 1 Option 2 Assessment

2.3.1 Overview

There are three key components in the defect assessment in FAD format, see Figure 2-4:

1. Failure assessment curve (FAC), 2. Stress or load ratio, Sr or Lr, and 3. Toughness ratio, Kr.

The FAC is a locus that defines the critical states in terms of the stress and toughness ratios. The stress ratio defines the likelihood of plastic collapse. The toughness ratio is the ratio of applied crack driving force over the material’s fracture toughness. It defines the likelihood of brittle fracture.

The exact form of FAC and the computation of stress and toughness ratios depend on the type of structural geometry, defect location, defect size, and material’s strain hardening behavior. Over the years, many solutions have been developed for various structural geometries, defect locations, and material properties. The assessment procedures presented here is specifically developed and validated for pipeline girth welds.

The defect assessment in the FAD format may be used in one of two ways:

1. For a structure with a known defect and applied load level, the significance of the defect, i.e., safe or unsafe, can be determined. This is done by comparing the location of the “assessment point” with the FAC. If the point falls inside the FAC locus, the structure is deemed safe. Figure 2-4 shows how this type of assessment is done.

2. For a structure with known defect location, material property, and applied load level, the critical defect size can be determined. This is almost always done iteratively, and therefore can be time-consuming without the aid of a computer program.

2.3.2 Additional Requirements

In addition to the requirements of Section 1.2, the following requirements are necessary at the minimum design temperature.

1. Weld metal strength even- or over-matches that of pipe material

2. No failure in the HAZ (heat-affect-zone) when defect-free welds are tested

3. Applies longitudinal stress no greater than SMYS and applied longitudinal strain no greater than 0.5%

4. The minimum CTOD toughness is greater than 0.05 mm.

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0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6Lr

Kr

Plastic Collapse

Brittle Fracture Failure Assessment

Curve, Eq. (4)

Unacceptable Region

Acceptable Region

Assessment Point

y

ene

Jdσ

δ = , Eq. (7) dn, Eqs. (8), (9), (10), and (1)

Je, Eqs. (11), (12), and (13)

mat

erK

δδ

= , Eq. (6) δmat, CTOD toughness

σa, stress analysis

c

arL

σσ

= , Eq. (14) σc, Eq. (15)

Cutoff, Eq. (5)

Figure 2-4 Schematic overview of the Level 1 Option 2 procedure

9

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2.3.3 Determination of the Key Components in the FAD Procedure

2.3.3.1 Failure Assessment Curve (FAC)

The FAC is taken from R6 Option 1 [14],

( ) ( )[ ]62 65.0exp7.03.0)14.01( rrrr LLLfK −+−== (4)

The cut-off of the FAC on the Lr axis is at,

yfcutoffrL σσ /= , (5)

where σf is determined by Eq. (2).

2.3.3.2 Assessment Point, Toughness Ratio Kr

When material’s fracture toughness is measured in CTOD, Kr is given as,

mat

erK

δδ

= (6)

where δmat is the CTOD toughness of the material. The elastic component of the CTOD driving force, δe, may be computed as,

y

ene

Jdσ

δ = (7)

The J to CTOD conversion factor, dn, is given as,

882.0119.3169.32

+⎟⎠⎞

⎜⎝⎛−⎟

⎠⎞

⎜⎝⎛=

nndn

( )( ){ }

(8)

TYn t

//1ln005.0/ln ε

=

22.0*00175.0

, (9)

, (10) +−= yt σε

where the pipe grade, , is in the unit of ksi. The elastic J integral is given as, yσ

( )2

2

1/ υ−=

EKJ I

e (11)

baI FaK πσ= (12)

The parameter Fb is a function of pipe diameter ratio, α, and defect length β, and defect height η,

( ) 2806.01983.0906.0791.031.209.1,, mmFb βα

αβηβαηβα +++= (13a)

10

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21 345.0163.000985.0 ηη −−−=m (13b)

22 155.018.200416.0 ηη +−−=m (13c)

Additional corrections apply to Fb for the following defect conditions,

( ) ⎟⎠⎞

⎜⎝⎛ == η

αη

πβαηβα ,80,,, bb FF if ,1.0≥η and ,

αη

πβ 80

> (13d)

,.απ

β 1080>( ) ⎟

⎠⎞

⎜⎝⎛ == η

απβαηβα ,1.080,,, bb FF if ,1.0<η and (13e)

2.3.3.3 Assessment Point, Stress Ratio Lr

The stress ratio Lr is given as,

c

arL

σσ

= (14)

( ) ( )yc σβπηηβπηβπσ ⎥

⎦

⎤⎢⎣

⎡−⎟

⎠⎞

⎜⎝⎛

⎥⎦⎤

⎢⎣⎡ −+=

2sin

2cos05.0385

45.2 05.0< if ηβ (15a)

( )yc σβπηηβππσ ⎥

⎦

⎤⎢⎣

⎡−⎟

⎠⎞

⎜⎝⎛=

2sin

2cos

4 if 05.0≥ηβ (15b)

2.3.4 Defect Acceptance Criteria

The total defect length shall be no greater than 12.5% of the pipe circumference. The maximum defect height shall be no greater than 50% of the pipe wall thickness.

A safety factor of 1.5 on defect length must be given from the critical defect size computed per Section 2.3.3 procedure.

The defect acceptance level is computed iteratively. The following steps may be followed:

1. Select a defect size as a start point. A reasonable start point is a defect with the maximum allowed height, η=0.5, and a small defect length that represents the smallest defect length that the selected AUT procedure can confidently detect.

2. Determine the assessment point in the FAD format per Section 2.3.3.

3. If the assessment point falls in the safe region,

(a) increase the defect length until the assessment point falls on the failure assessment curve. This represents a critical state with the combination of load, material property, and defect size. Make a note of the defect height and length.

(b) Reduce the defect height by a small increment, say ∆η=0.05 (η=0.45). Start from the defect length determined in (a) and increase the defect length until

11

the assessment falls on the failure assessment curve. This represents another critical state with a shallower and longer defect than that determined in (a). Make a note of the defect height and length.

(c) Repeat the process (b) until the shallowest defect height of interest has reached.

4. If the assessment point falls outside the safe region,

(a) Decrease the defect length until the assessment point falls on the failure assessment curve. This represents a critical state with the combination of load, material property, and defect size. Make a note of the defect height and length.

(b) Repeat the steps 3(b) and 3(c).

An example of the procedure is given in Appendix G.

2.4 Limitations of the Level 1 Procedures

Due to the availability of public domain data, the validations of the procedures have been primarily limited to large diameter and large D/t ratio pipes. The fundamental basis of the assessment procedures does not place limits on diameter or D/t ratio. It is prudent, however, that cautions be exercised in applying the procedures to heavy wall and small diameter pipes.

The effects of residual stress are not explicitly considered. It is believed that the residual stress has minimal effects on the defect acceptance criteria provided that (1) the failure mechanism is not time-dependent and (2) the CTOD toughness is greater than the minimum required value of 0.05 mm (0.002 inch). The examples of time-dependent failure mechanisms include, but not limited to, fatigue and stress corrosion cracking.

03-G78-20 12

3.0 Level 2 Assessment Procedures The part is under development.

03-G78-20 13

4.0 References

1 API Standard 1104, "Welding of Pipelines and Related Facilities," 19th Edition, September 1999.

2 Canadian Standards Association, CSA-Z662, "Oil and Gas Pipeline Systems," 2003. 3 British Standard Institute, BS7910, Guidance on methods for assessing the acceptability of

flaws in structures, 1999. 4 British Standard Institution, PD6493:1980, “Guidance on Some Methods for the Derivation of

Acceptance Levels for Defects in Fusion Welded Joints.” 5 British Standard Institution, PD6493:1991, “Guidance on Some Methods for the Derivation of

Acceptance Levels for Defects in Fusion Welded Joints.” 6 Wang, Y.-Y., Swatzel, J., Horsley, D., and Glover, A., “Girth Weld ECA from the Perspective

of Code Revisions in North America,” Proceedings of the International Pipeline Conference 2002, Calgary, Alberta, Canada, September 29-October 3, 2002.

7 Knauf, G. and Hopkins, P., “EPRG Guidelines on the Assessment of Defects in Transmission Pipeline Girth Welds,” Sonderdruck aus 3R International, 35 Jahrgang, Heft 10-11/1996, s. 620-624.

8 Australian Standard, AS 2885.2-1995, “Pipelines – Gas and Liquid Petroleum, Part 2: Welding.”

9 API RP 579, “Fitness-for-Service,” First Edition, January 2002. 10 SINTAP Procedure, Final Version, November 1999. 11 Wang, Y.-Y., Rudland, D., Horsley, D., “Development of a FAD-Based Girth Weld ECA

Procedure, Part I Theoretical Framework,” Proceedings of the 4th International Pipeline Conference, Calgary, Alberta, Canada, September 29-October 3, 2002.

12 Wang, Y.-Y., Rudland, D., Horsley, D., “Development of a FAD-Based Girth Weld ECA Procedure, Part II Experimental Verification,” Proceedings of the 4th International Pipeline Conference, Calgary, Alberta, Canada, September 29-October 3, 2002.

13 Harrison, R. P., Loosemore, K., and Milne, I., “Assessment of the Integrity of Structures Containing Defects,” CEGB Report No. R/H/6, Central Electricity Generating Board, United Kinddom, 1976.

14 British Energy Generation Ltd., Assessment of Integrity of Structures Containing Defects, R/H/R6-Revision 3, 1999.

15 ASME Boiler and Pressure Vessel Code, Section XI, Appendix H, 1992 Edition, July 1992.

03-G78-20 14

Appendix A Validation of the Assessment Procedures against Full-Scale Bend Tests

03-G78-20 A-1

A.1 Background

This appendix provides the validation of the Level 1 Option 2 procedures. Since the plastic collapse criterion implemented in the Level 1 Option 1 procedure is a part of the Option 2 procedure, this validation is also an indirect confirmation of the Option 1 procedure.

The Level 1 Option 2 procedure is substantially taken from the PRCI-funded work of Wang, et al. [1,2]. This document incorporates several improvements made to the initial work. They include:

• Slightly revised plastic collapse criterion,

• Updated stress intensity factor solutions,

• Updated Y/T ratio estimation from pipe grade,

• Updated estimation of strain hardening exponent, and

• Revised conversion factor from J-integral to CTOD.

These improvements represent incremental advances to the initial work. The outcome of this validation was not expected to be significantly different from that of the initial validation. Nevertheless, this validation is the direct confirmation of the exact procedures outlined in this document.

A.2 Experimental Database for Validation

Data from 69 full-scale experimental tests were collected and used as the basis for this validation. All the tests were conducted in bending with artificially introduced defects in the circumferential direction, simulating girth weld defects. These 69 tests represent perhaps the largest test database for large diameter pipelines in the open literature. Among these tests, 54 tests came from full-scale experimental tests conducted at the Welding Institute of Canada (WIC) and the University of Waterloo [3,4]. Most of the tested pipes were API Grade X70 (483 MPa), a few were X65 (448 MPa) and X60 (414 MPa) grades. The pipe diameter ranged from 20 inch (508 mm) to 42 inch (1067 mm). The reported CTOD toughness was in the range of 0.03 to 0.10 mm (0.0012 to 0.0039 inch). The bending moments, stresses, and remote nominal strains at the critical events were reported for many of these tests. The critical events could be brittle fracture, brittle fracture after ductile tearing, buckling, or manual intervention.

In addition to the WIC and University of Waterloo tests, four tests by Erdogan [5] are included in the test database. These tests were conducted on X60 pipes with 20 inch OD and 0.344 inch (8.74 mm) wall thickness. The reported CTODs were 0.554 mm. The other test data include 8 tests by Hopkins on X65 36 inch (914.4 mm) OD pipe [6] and 3 tests by Wilkowski on X60 30 inch (762 mm) OD pipes [7]. The CTOD toughness of the Hopkins and Wilkowski tests were in the range of 0.02 to 0.10 mm.

03-G78-20 A-2

A.3 Validation Process

For each test, 7 parameters were entered into the assessment procedure,

1. Pipe diameter, 2. Pipe wall thickness, 3. Pipe grade or measured yield stress 4. Defect depth, 5. Defect length, 6. CTOD toughness, and 7. Magnitude of the applied stress. An assessment point is produced for each test, following the procedure outlined in

Section 2.3.3. The relative conservatism and accuracy of the assessment procedure is determined by examining the location of the assessment points. A computer program was developed to facilitate the calculations.

A.4 Results of the Validation

The assessment points of all 69 full-scale tests are plotted on the FAD in Figure A-1 and Figure A-2. The nominal pipe grades are used as the pipe strength input in Figure A-1, whereas the measured yield stresses are used as the pipe strength input in Figure A-2. In reference to Eq. (5), the cut-off for the FAC varies by pipe grade (pipe strength). The cut-off for a nominal X70 pipe is shown in the figures. The cut-off points for the entire database vary slight as the grades are in the narrow range of X60 to X70.

Since the assessment points represent actual failure events, points falling outside of the FAC mean the FAC is conservative with respect to the actual failure events. Conservative predictions are obtained for all tests. When the nominal yield strength is used, the median value of the safety factor in stress is 1.66 with a standard deviation of 0.37. The safety factor is the ratio of the experimentally measured maximum stress over the predicted failure stress, with the material property and defect size remaining the same. The median value of the safety factor in stress is 1.51 with a standard deviation of 0.35 when the measured yield stresses of the pipes were used as the pipe strength input.

The database of 69 full-scale tests include defects from as shallow as 6% of the wall thickness to as deep as through-wall defects. The CTOD toughness ranges from 0.02 mm to 0.55 mm. In reference to Section 2.3.4, the tests that are directly relevant to the acceptance criteria are those with defects less than 50% of pipe wall thickness and CTOD toughness greater than 0.05 mm. The maximum defect length in the database is 12% of the pipe circumference, therefore all within the acceptance limits. When the maximum defect size and the minimum CTOD toughness criteria of Section 2.3.4 are applied, 35 tests fall within the limits, or about one-half of the total number of tests.

The assessment points of these 35 full-scale tests are plotted on the FAD in Figure A-3 and Figure A-4. The nominal pipe grades are used as the pipe strength input in Figure A-3,

03-G78-20 A-3

whereas the measured yield stresses are used as the pipe strength input in Figure A-4. In comparison, the scatter of the test data with respect to the FAC is markedly reduced when the limits of Section 2.3.4 are applied. When the nominal yield strength is used, the median value of the safety factor in stress is 1.46 with a standard deviation of 0.28. The median value of the safety factor in stress is 1.44 with a standard deviation of 0.23 when the measured yield stresses of the pipes were used as the pipe strength input.

A.5 Observation from the Validation against Full-scale Bend Tests

1. The assessment procedure is conservative when compared to all 69 full-scale test data. The database contains tests with defect size and CTOD toughness outside the limits of the current acceptance criteria. This demonstrates the robustness of the assessment procedure.

2. When the limitations on defect size and CTOD toughness, as proposed in the recommended acceptance criteria, are applied, the consistency of the assessment procedure shows marked improvement with respect to the test data.

3. It should be noted that no safety factor was applied when the assessment procedure is validated against the experimental test data. The assessment procedure is assumed to predict critical failure events. In the recommended acceptance criteria, a safety factor of 1.5 is applied after the “critical” defect size is determined. This safety factor in defect length is in line with historical recommendations. A higher degree of conservatism than that shown here is preserved when the recommended acceptance criteria are applied.

4. The scatter in the test data with respect to the FAC is perhaps due to the fact that a number of factors affecting the test results cannot be reconstructed from the published papers and reports. These factors include, but not limited to, weld strength mismatch, scatter in CTOD toughness, variation in stress-strain curves of the pipe material and weld metal, etc. It is believed that, for instance, the reported CTOD toughness is the minimum value of a “batch” of tests. It may or may not reflect the CTOD toughness on a specific piece of pipe at the specific defect location. On the other hand, such details are frequently not available if the assessment procedure is applied to existing pipelines. This validation shows that the procedure is conservative, but not overly so, when used with minimum required input data.

03-G78-20 A-4

03-G78-20

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0.0 0.5 1.0 1.5 2.0L r

Kr

Figure A-1 Full-scale test data plotted on the FAD of the Level 1 Option 2 procedure.

The nominal SMYS were used as the strength input.

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0.0 0.5 1.0 1.5 2.0L r

Kr

Figure A-2 Full-scale test data plotted on the FAD of the Level 1 Option 2 procedure.

The measured yield stresses were used as the strength input.

A-5

03-G78-20

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0.0 0.5 1.0 1.5 2.0L r

Kr

Figure A-3 Full-scale test data that fall within the defect size and CTOD limitations

plotted on the FAD of the Level 1 Option 2 procedure. The nominal SMYS were used as the strength input.

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0.0 0.5 1.0 1.5 2.0L r

Kr

Figure A-4 Full-scale test data that fall within the defect size and CTOD limitations

plotted on the FAD of the Level 1 Option 2 procedure. The measured yield stresses were used as the strength input.

A-6

A.6 References

1 Wang, Y.-Y., Rudland, D., Horsley, D., “Development of a FAD-Based Girth Weld ECA

Procedure, Part I Theoretical Framework,” Proceedings of the 4th International Pipeline Conference, Calgary, Alberta, Canada, September 29-October 3, 2002.


3 Pick, R. J., Glover, A. G., and Coote, R. I., “Full Scale Testing of Large Diameter Pipelines,” Proceedings of Conference on Pipeline and Energy Plant Piping, Pergamon Press, 1980, pp. 357-366.

4 Glover, A. G., Coote, R. I., and Pick, R. J., “Engineering Critical Assessment of Pipeline Girth Welds,” Proceedings of Conference on Fitness for Purpose Validation of Welded Construction, The Welding Institute, Paper 30, 1981.

5 Erdogan, F., "Theoretical and Experimental Study of Fracture in Pipelines Containing Circumferential Flaws," DOT-RSPA-DMA-50/83/3, Contract DOT-RC-82007 Final Report to USDOT, September 1982.

6 Hopkins, P., Demofonti, G., Knauf, G., and Denys, R., “an Experimental Appraisal of the Significance of Defects in Pipeline Girth Welds,” 8th EPRG/PRC Biennial Joint Technical Meeting on Line Pipe Research, Paris, 1991.

7 Wilkowski, G. M., and Eiber, R. J., "Evaluation of Tensile Failure of Girth Weld Repair Grooves in Pipe Subject to Offshore Laying Stresses," Journal of Energy Resources Technology, v. 103, March 1981.

03-G78-20 A-7

Appendix B Validation of the Assessment Procedures against Curved Wide Plate Tests

78 03-G78-20 B-1

B.1 Background

This appendix provides further validation of the Level 1 Option 2 procedure, using experimental test data from curved wide plates. Since the plastic collapse criterion implemented in the Level 1 Option 1 procedure is a part of the Option 2 procedure, this validation is also an indirect confirmation of the Option 1 procedure.

This validation is similar to the PRCI-funded work of Wang, et al. [1]. It is useful as this is a direct confirmation of the proposed assessment procedures, which incorporate the improvements to the prior work as listed in Section A.1.

This appendix also covers the validation of the assessment with the option of incorporating the effects of weld strength mismatch. The inclusion of the weld strength mismatch is further discussed in Appendix F.

B.2 Overview of the Wide Plate Tests

The test data were taken from 31 curved wide plate (CWP) tests performed at the University of Gent [2]. The pipe material was a longitudinally welded API 5L X60 pipe with 36-inch (914.4 mm) OD and 11.6-mm (0.457-inch) wall. The averaged 0.5% proof stress in the longitudinal direction was 64.7 ksi (446 MPa). The averaged tensile strength in the same direction was 81.2 ksi (560 MPa). There were ten girth welds made with seven combinations of cellulosic electrodes (AWS Exx10). This offered seven levels of weld strength mismatch ranging from 20% undermatching to 24% overmatching.

The CWPs were cut from welded pipe sections in the longitudinal direction with the girth weld in the mid-length. The girth weld defects were introduced by sharp starter notch and fatigue pre-cracked. The gauge section of the CWPs had a nominal width of 300 mm. All CWPs were loaded in tension until failure. The load, overall deformation, and CMOD (crack mouth opening displacement) were recorded during the tests. The test temperatures were –10, -30, and –50oC.

B.3 Validation Process

The validation process consists of three input options (IO). Each IO represents different levels of available data that might be encountered in practice.

The IO 1 assumes the following input data are available: 1. Pipe diameter, 2. Pipe wall thickness, 3. Pipe grade, 4. Defect depth, 5. Defect length, 6. CTOD toughness, and 7. Applied longitudinal stress.

78 03-G78-20 B-2

In IO 2, the pipe grade is replaced with the measured pipe yield stress and tensile strength.

In IO 3, the measured yield stress and tensile strength of the pipe material and weld metals are both known. The effectiveness of incorporating the weld strength mismatch in the assessment procedures is evaluated in this IO.

B.4 Validation Results against Curved Wide Plate Test Data

The assessment results are given in FAD format in Figure B-1, Figure B-2, and Figure B-3 and in tabulated form in Table B-1. The assessment points are listed for all cases. In addition, the safety factors in terms of applied stress are also listed. The safety factor is the ratio of the experimentally measured maximum stress over the predicted failure stress.

The results of IO 1 are given in Figure B-1. There is a single cut-off of the FAC for all cases, as the cut-off is related to the nominal pipe grade. The median value of the safety factor is 1.23 with a standard deviation of 0.11. All test cases are conservatively predicted, even for the undermatched cases. This is the direct result of using the nominal pipe grade as the strength input. The actual pipe strength was higher. The weld strength mismatch is not considered in this IO.

The results of IO 2 are given in Figure B-2. The single cut-off of the FAC for all cases is determined by the measured yield and tensile strength of the pipe. The median value of the safety factor is 1.12 with a standard deviation of 0.10. Since the measured tensile properties are greater than the SMYS (pipe grade), the assessment procedure predicted higher failure stresses than those obtained if the nominal pipe grades are used. Consequently, some of the undermatched cases are not conservatively predicted. The weld strength mismatch is not considered in this IO.

The results of IO 3 are given in Figure B-3. The cut-off of the FAC is affected by the weld strength mismatch, therefore, represented by a broken line. The median value of the safety factor is 1.12 with a standard deviation of 0.07. One case has a safety factor of 0.99. All others are conservatively predicted. In comparison to other IOs, this one is clearly the best, as it offers the lowest standard deviation and very good overall accuracy. This demonstrates that the accuracy of the procedure is improved with the inclusion of the weld strength mismatch effects.

B.5 Observation from the Validation against Curved Wide Plate Test Data

It should be noted that no safety factor was applied when the assessment procedure is validated against the curved wide plate test data. The assessment procedure is assumed to predict critical failure events. In the recommended acceptance criteria, a safety factor of 1.5 is applied after the “critical” defect size is determined. The recommended acceptance criteria have a higher degree of conservatism than that shown in this appendix.

78 03-G78-20 B-3

78 03-G78-20

1. In comparison with the validation against the full-scale test data, the assessment procedure showed a higher degree of consistency and accuracy. This is reflected in the small values of standard deviation.

2. In the absence of explicit consideration for weld strength mismatch, the assessment procedure is conservative and accurate when the weld strength at least overmatches that of the pipe.

3. When the weld strength mismatch levels were taken into account, the new procedure produced consistent and highly accurate predictions against the experimental results.

4. If the pipe grade is the only known strength input, even or over-matching weld metal is necessary to ensure conservative predictions.

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0.0 0.5 1.0 1.5 2.0L r

Kr

Figure B-1 Curved wide plate test data plotted on the FAD of the Level 1 Option 2

procedure. The nominal SMYS were used as the pipe strength input.

B-4

78 03-G78-20

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0.0 0.5 1.0 1.5 2.0L r

Kr

Figure B-2 Curved wide plate test data plotted on the FAD of the Level 1 Option 2

procedure. The measured yield and tensile strength of the pipe were used as the pipe strength input.

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0.0 0.5 1.0 1.5 2.0L r

Kr

Figure B-3 Curved wide plate test data plotted on the FAD of the procedure that is an

extension of the Level 1 Option 2 procedure. The measured yield and tensile strength of the pipe AND weld were used as the strength input.

B-5

78 03-G78-20

Table B-1 Summary of the Validation against the Curved Wide Plate Test Data

L r Kr

Fact

or o

f Saf

ety

in S

tres

s

L r Kr

Fact

or o

f Saf

ety

in S

tres

s

L r Kr

Fact

or o

f Saf

ety

in S

tres

s

(MPa) (MPa) (mm) (mm) (mm)1 359 461 0.80 0.82 0.26 3.50 48.7 1.19 0.32 1.08 1.11 0.30 0.98 1.29 0.30 1.156 415 527 0.93 0.94 0.24 3.70 45.6 1.25 0.35 1.14 1.16 0.33 1.03 1.23 0.33 1.0911 473 598 1.06 1.07 0.14 3.90 45.0 1.20 0.45 1.09 1.11 0.43 1.01 1.09 0.43 0.9928 533 649 1.20 1.16 0.21 3.40 48.0 1.56 0.45 1.42 1.45 0.42 1.29 1.38 0.42 1.2224 554 646 1.24 1.15 0.27 3.60 48.5 1.53 0.40 1.38 1.42 0.38 1.26 1.33 0.38 1.182 359 461 0.80 0.82 0.37 3.50 27.1 1.12 0.23 1.01 1.04 0.21 0.92 1.21 0.21 1.073 359 461 0.80 0.82 0.37 4.30 48.0 1.25 0.31 1.13 1.16 0.29 1.03 1.38 0.29 1.227 415 527 0.93 0.94 0.30 3.80 25.7 1.27 0.30 1.16 1.18 0.28 1.05 1.25 0.28 1.118 415 527 0.93 0.94 0.30 4.30 49.3 1.29 0.35 1.17 1.19 0.33 1.06 1.27 0.33 1.1212 473 598 1.06 1.07 0.18 4.10 25.0 1.37 0.42 1.24 1.27 0.40 1.12 1.25 0.40 1.1013 473 598 1.06 1.07 0.18 3.70 50.5 1.25 0.41 1.14 1.16 0.39 1.03 1.14 0.39 1.0116 488 586 1.09 1.05 0.28 4.70 50.3 1.51 0.45 1.37 1.40 0.42 1.24 1.36 0.42 1.2017 488 586 1.09 1.05 0.28 4.20 72.4 1.33 0.39 1.20 1.23 0.37 1.09 1.20 0.37 1.0620 498 614 1.12 1.10 0.23 4.80 48.6 1.44 0.48 1.30 1.33 0.45 1.18 1.29 0.45 1.1421 498 614 1.12 1.10 0.23 3.40 73.6 1.45 0.42 1.32 1.35 0.39 1.19 1.31 0.39 1.1629 533 649 1.20 1.16 0.12 4.00 23.3 1.61 0.60 1.46 1.49 0.56 1.34 1.42 0.56 1.2930 533 649 1.20 1.16 0.12 4.30 49.7 1.29 0.56 1.20 1.20 0.53 1.12 1.14 0.53 1.0825 554 646 1.24 1.15 0.13 3.50 22.2 1.58 0.53 1.43 1.47 0.50 1.30 1.38 0.50 1.2326 554 646 1.24 1.15 0.13 4.10 49.8 1.36 0.56 1.25 1.26 0.52 1.16 1.18 0.52 1.104 359 461 0.80 0.82 0.26 3.60 24.0 1.22 0.30 1.11 1.14 0.28 1.01 1.33 0.28 1.185 359 461 0.80 0.82 0.26 3.00 48.5 1.28 0.31 1.16 1.19 0.29 1.05 1.37 0.29 1.229 415 527 0.93 0.94 0.24 4.20 26.8 1.30 0.36 1.18 1.20 0.33 1.07 1.27 0.33 1.1310 415 527 0.93 0.94 0.24 2.50 47.2 1.30 0.30 1.18 1.21 0.28 1.07 1.26 0.28 1.1214 473 598 1.06 1.07 0.14 3.40 24.1 1.41 0.46 1.28 1.31 0.43 1.16 1.29 0.43 1.1415 473 598 1.06 1.07 0.14 3.50 48.5 1.34 0.48 1.21 1.24 0.45 1.11 1.22 0.45 1.0918 488 586 1.09 1.05 0.13 2.90 23.2 1.37 0.43 1.24 1.27 0.40 1.12 1.24 0.40 1.1019 488 586 1.09 1.05 0.13 3.50 49.6 1.36 0.51 1.23 1.26 0.48 1.14 1.23 0.48 1.1122 498 614 1.12 1.10 0.11 3.10 24.5 1.40 0.49 1.27 1.29 0.46 1.15 1.26 0.46 1.1223 498 614 1.12 1.10 0.11 3.80 50.2 1.36 0.58 1.26 1.26 0.54 1.17 1.22 0.54 1.1431 533 649 1.20 1.16 0.06 4.00 47.3 1.35 0.79 1.36 1.25 0.74 1.26 1.18 0.74 1.2227 554 646 1.24 1.15 0.07 5.00 49.0 1.23 0.76 1.26 1.14 0.71 1.17 1.07 0.71 1.12

Median 488 598 1.09 1.07 0.21 3.70 48.0 1.34 0.43 1.23 1.24 0.40 1.12 1.26 0.40 1.12Std. Dev. 65 65 0.14 0.12 0.08 0.56 13.9 0.12 0.13 0.11 0.11 0.12 0.10 0.09 0.12 0.07

Mis

mat

ch R

atio

at U

TS

IO 3, Using Measured Yield and Tensile of

Pipe and Welds

Def

ect H

eigh

t

Def

ect L

engt

h

Test

No.

Wel

d Te

nsile

CTO

D

IO 2, Using Measured Yield and Tensile of

PipeW

eld

Yiel

d

IO 1, Using SMYS of Pipe

Mis

mat

ch R

atio

at Y

ield

B-6

B.6 References

1 Wang, Y.-Y., Rudland, D., Horsley, D., “Development of a FAD-Based Girth Weld ECA

Procedure, Part II Experimental Verification,” Proceedings of the 4th International Pipeline Conference, Calgary, Alberta, Canada, September 29-October 3, 2002.

2 Denys, R. M., “The Effect of Weld Metal Matching on Girth Weld Performance, Volume II – Experimental Investigation,” final report to the Pipeline Research Committee of the American Gas Association, PR-202-922, January 24, 1993.

78 03-G78-20 B-7

Appendix C Stress Intensity Factor Solution

78 03-G78-20 C-1

78 03-G78-20

C.1 Background

Chapuloit, et al., conducted over 200 3-D FE analyses of pipes containing finite-length semi-elliptical defects [1]. The KI solutions were derived for pipes with D/t ratio ranging from 4 to 162 and a/t ratio up to 0.8. The solutions were also calibrated with the flat plate solutions of Newman and Raju and Irwin’s analytical solution of an elliptical crack in an infinite body [2,3]. Overall, Chapuloit’s solutions are perhaps one of the most comprehensive solutions in the published literature for circumferentially-cracked pipes under bending loads.

In the previously published PRCI-funded work of Wang, et al. [4,5], the KI solutions of Chapuloit at the deepest point were fitted to a set of parametric equations. The KI solutions took a similar form as those of ASME Section XI solutions. However, due to the use of a high order polynomial function in the parametric equations, the fitted equations did not give correct KI values at either very small or large a/t ratios. As shown by an example in Figure C-1, the trends at a/t < 0.1 and a/t > 0.6 are not consistent with the original data. Furthermore, the overall fit has some oscillation even within the range of 0.1 < a/t < 0.6 due to the high order polynomial functions.

A new fitting was conducted to remove the polynomial functions. The new parametric equations provide more consistent agreement with the original data as shown in Figure C-1.

0.8

1.0

1.2

1.4

1.6

1.8

0.0 0.2 0.4 0.6 0.8 1.0a/t

Fb

Previous FitCurrent FitData of Chapuloit

D /t=82, β =0.012

Figure C-1 Comparison of the fitted equations as a function of defect depth ratio

C-2

78 03-G78-20

C.2 Parametric Equations

The newly fitted equations are given as follows.

ta

Dc

tD

=== ηπ

βα ,2, (C.1)

baI FaK πσ= (C.2)

( ) 2806.01983.0906.0791.031.209.1,, mmFb βα

αβηβαηβα +++= (C.3a)

21 345.0163.000985.0 ηη −−−=m

22 155.018.200416.0 ηη +−−=m

Additional corrections apply to Fb for the following defect conditions,

( ) ⎟⎠⎞

⎜⎝⎛ == η

αη

πβαηβα ,80,,, bb FF if 1.0≥η and ,

αη

πβ 80

>

( )

(C.3b)

,.απ

β 1080>

8.01.

⎟⎠⎞

⎜⎝⎛ == η

απβαηβα ,1.080,,, bb FF 1.0 if <η and (C.3c)

≤≤ ηThe above equations were fitted to data in the range of 0 , 16242 ≤≤ α , and

αη

πβ

αη

π322

≤≤ .

When the D/t ratio is sufficiently large, the KI solution reverts back to the flat plate solution. Under such conditions, the KI value at the deepest point is no longer a function of defect length. This is essentially the KI solution specified in Eqs. (C.3b) and (C.3c).

C.3 Comparison between Fitted Equations and the FE Results

The accuracy of the parametric fit of Eqs. (C.1)-(C.3) are examined by comparing the KI values from the fitted equations with the original data of Chapuloit. In Figure C-2 to Figure C-4, the comparisons are made for three D/t ratios and various defect sizes. The overall agreement is very good. The points at which the curves turn flat indicate that the conditions of Eq. (C.3b) are satisfied. At larger β values, the parameter Fb remains constant.

C-3

78 03-G78-20

0.0

0.5

1.0

1.5

2.0

2.5

0.00 0.03 0.06 0.09 0.12 0.15

β

F b

η = 0.1η = 0.2η = 0.4η = 0.6

α = D/t = 42

Figure C-2 Comparison of the fitted curves with the original data of Chapuloit for pipes

with D/t of 42. The symbols are from the original data and the curves are from the fitted equations.

0.0

0.5

1.0

1.5

2.0

2.5

3.0

0.00 0.03 0.06 0.09 0.12 0.15

β

F b

η = 0.1η = 0.2η = 0.4η = 0.6

α = D/t = 82



C-4

78 03-G78-20

0.0

0.5

1.0

1.5

2.0

2.5

3.0

3.5

0.00 0.03 0.06 0.09 0.12 0.15

β

F b

η = 0.1η = 0.2η = 0.4η = 0.6

α = D/t = 162



C.4 References

1 Chapuloit, S., Lacire, M. H., and Le Delliou, P., “Stress Intensity Factors for Internal

Circumferential Cracks in Tubes over a Wide Range of Radius over Thickness Ratios,” PVP Vol. 365, ASME 1998, pp. 95-106.

2 Newman, J. C., Jr. and Raju, I. S., “Analysis of Surface Cracks in Finite Plates Under Tension and Bending Loads,” NASA Technical Paper 1578, NASA, Washington, D. C., December 1978.

3 Irwin, G. R., “Crack-extension force for a part-through crack in a plate,” Journal of Applied Mechanics, Vol. 29, 1962, pp. 651-654.

4 Wang, Y.-Y., Rudland, D., Horsley, D., “Development of a FAD-Based Girth Weld ECA Procedure, Part I Theoretical Framework,” Proceedings of the 4th International Pipeline Conference, Calgary, Alberta, Canada, September 29-October 3, 2002.


C-5

Appendix D Plastic Collapse Solution

03-G78-20 D-1

03-G78-20

D.1 Background

Plastic collapse solutions suitable for girth weld defect assessments have been investigated extensively by Wang [1,2]. There are several earlier reviews, e.g., Rosenfeld [3] and Clyne and Jones [4], that targeted pipelines. Miller [5] conducted an extensive review of limit load solutions for a wide range of structural geometries, including pipes.

The plastic collapse solution adopted here follows the recommendation of Wang [6,7]. The basis of the recommendation has been covered extensively in prior publications. The plastic collapse solution is due to Miller [5],

( )⎥⎦

⎤⎢⎣

⎡−⎟

⎠⎞

⎜⎝⎛=

2sin

2cos βπηηβπσσ f

Millerc (D.1)

where is the nominal longitudinal stress at plastic collapse. The Miller solution of Eq. (D.1) was compared with the full-scale test data of Glover [

Millercσ

8,9] and Erdogan [10]. The features of the test database are described in Appendix A. Based on the comparison, a defect size correction factor was proposed [6],

( )⎥⎦

⎤⎢⎣

⎡−⎟

⎠⎞

⎜⎝⎛=

2sin

2cos βπηηβπσσ f

Girthc f (D.2)

where is the nominal longitudinal stress at plastic collapse of a girth weld with the correction factor f. The defect size correction factor f was given as [6],

Girthcσ

05.0141

1ηβ

π⎟⎠⎞

⎜⎝⎛ −+

=f if ηβ ≤ 0.05 (D.3a)

4π

=f if ηβ > 0.05 (D.3b)

The modified Miller solution of Eqs. (2) and (3) is the basis of the revised plastic collapse criterion in CSA Z662 Appendix K 2003 Edition. The acceptance criteria of CSA Z662 Appendix K has a safety factor of 2 on the defect length computed from Eqs. (2) and (3). This safety factor is consistent with historical recommendations.

D.2 New Defect Size Correction Factor

The plastic collapse solution of Eqs. (D.2) and (D.3) worked well in comparison to test data [1,2,6,7]. However, the discontinuity of the first order derivative at ηβ = 0.05 can pose a problem if the current assessment procedures are cast into an optimization procedure. Although this is not an immediate concern for this project, revisions were made to the defect size correction factor.

D-2

03-G78-20

( ) 5.205.03854

ηβπ−+=f if ηβ ≤ 0.05 (D.4a)

4π

=f if ηβ > 0.05 (D.4b)

D.3 Comparison with Full-scale Test Data

The bending stresses at the critical events in the full-scale tests, normalized by the plastic collapse stresses of Eq. (D.1) are shown in Figure D-1. When the stress ratio is greater than 1.0 on the y-axis, the actual failure stress is greater than the predicted plastic collapse stress. In such cases, the Miller solution is conservative. It is evident from Figure D-1 that the Miller solution is less conservative for larger defects. A defect size correction line, Eq. (D.3), was suggested by Wang [6]. The new defect size correction line, corresponding to Eq. (D.4), is also shown in the figure. The new correction line gives lower plastic collapse stress, and therefore, is more conservative than the prior correction line.

It should be noted that no minimum toughness criterion is applied to the test data. Consequently, it was not expected that all test data would fall conservatively above the correction line. The test data are used to set the overall trend.

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

0.00 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0.08

ηβ

σcEx

p / σ

cMill

er

Full-Scale Test DataPrior Defect Size Correction LineNew Defect Size Correction Line

Figure D-1 Comparison of the Miller plastic collapse solution and the full-scale test data.

The prior and current defect size correction lines are shown.

D-3

D.4 References

1 Wang, Y.-Y., Wilkowski, G. M., and Horsley, D. J., “Plastic Collapse Analysis of Pipelines

Containing Surface-Breaking Circumferential Defects,” in Pipeline Technology, Vol. I, R. Denys, Ed., Elsevier Science B. V., May 21-24, 2000, pp. 191-209.

2 Wang, Y.-Y., Wilkowski, G. M., and Horsley, D. J., “Plastic Collapse Analysis of Pipeline Girth Welds,” in Assessment Methodologies for Preventing Failure: Deterministic and Probabilistic Aspects and Welding Residual Stress, Vol. 1 ASME PVP-Vol. 410-1, Edited by R. Mohan, 2000, pp. 3-9.

3 Rosenfeld, M. J., "Serviceability of Corroded Girth Welds," Draft Final Report, PRI Contract No. PR 218-9438, March 31, 1995.

4 Kastner, W., Roehrich, E., Schmitt, W., and Steinbuch, R., “Critical Crack Sizes in Ductile Piping,” International Journal of Pressure Vessel and Piping, Vol. 9, 1981, pp.197-219.

5 Miller, A. G., "Review of Limit Codes of Structure Containing Defects," International Journal of Pressure Vessels and Piping, Vol. 32, 1988, pp. 191-327.

6 Wang, Y.-Y., Rudland, D., Horsley, D., “Development of a FAD-Based Girth Weld ECA Procedure, Part I Theoretical Framework,” Proceedings of the 4th International Pipeline Conference, Calgary, Alberta, Canada, September 29-October 3, 2002.


8 Pick, R. J., Glover, A. G., and Coote, R. I., “Full Scale Testing of Large Diameter Pipelines,” Proceedings of Conference on Pipeline and Energy Plant Piping, Pergamon Press, 1980, pp. 357-366.

9 Glover, A. G., Coote, R. I., and Pick, R. J., “Engineering Critical Assessment of Pipeline Girth Welds,” Proceedings of Conference on Fitness for Purpose Validation of Welded Construction, The Welding Institute, Paper 30, 1981.


03-G78-20 D-4

Appendix E Estimation of Applied Stress from Applied Strain

03-G78-20 E-1

03-G78-20

E.1 Assumed Stress Strain Relations

Sometimes, it is easier to determine the applied longitudinal strain than the applied longitudinal stress. When applied stress is needed for computation, such as for the determination of applied stress in the application of the Level 1 Option 1 procedure, the following process may be followed.

The overall stress strain curve is assumed to take the form that is suggested in CSA Z662, n

y

ayaa EE ⎟

⎟⎠

⎞⎜⎜⎝

⎛⎟⎟⎠

⎞⎜⎜⎝

⎛−+=

σσσσε 005.0 , (E.1)

where is the nominal yield stress (SMYS) and E is the Young’s modulus. There is an unique relation between applied stress

yσ

aσ and applied strain aε , if the strain hardening exponent n is known.

E.2 Estimation of Strain Hardening Exponent

By assuming a pure power stress strain relation, the strain hardening exponent may be estimated as,

( )( ){ }TY

n t

//1ln005.0/ln ε

= . (E.2)

E.3 Estimation of Y/T Ratio from Pipe Grade or Yield Stress

Webster and Bannister examined the correlation of Y/T ratio and yield strength [1]. Two simple relations were produced, one providing upper bound Y/T ratio, the other providing the best fit to the data. The relations were derived from theoretical and empirical considerations, and are applicable to many kinds of structural steels. Mannucci, et al., found the relations to be reasonable for pipeline steels tested in longitudinal direction [2]. The comparison of the linepipe test data and the upper bound and best fit relations is shown in Figure E-1.

E-2

03-G78-20

Figure E-1 Comparison of linepipe longitudinal test data with the Webster and Bannister

correlation equations [2]

0.7

0.8

0.9

1.0

50 60 70 80 90 100 110 120

Grade (ksi)

Y/T

Upper Bound, Webster and BanisterCSA Appendix KEq. 1 of the Current DocumentAPI min Y and T requirements Best Fit, Webster and Banister

Figure E-2 Comparison of the relations between Y/T ratio and pipe grades from

estimation equations and codes

E-3

03-G78-20

The relations for the upper bound and best fit by Webster and Bannister are shown in Figure E-2. Some reference points are added by computing the Y/T ratio from the API 5L minimum yield and tensile requirements. In addition, the plastic collapse criterion of CSA Z662 Appendix K 2003 provides a reference table between pipe grade and flow stress. The implied Y/T ratio may be obtained when the flow stress is taken as the averaged value between yield and tensile strength. The API 5L and CSA Z662 values are also shown in Figure E-2. A new equation in the same format as that of Webster and Banister, but providing the best fit to the API 5L and CSA Z662 Appendix K, is suggested as follows,

25.275.2121

1/

⎟⎟⎠

⎞⎜⎜⎝

⎛+

=

y

TY

σ

. (E.3)

The nominal yield stress is in the unit of ksi. yσ

E.4 Estimation of Uniform Strain

Estimating the strain at the ultimate tensile strength (UTS), often termed uniform strain or tensile strain, can be difficult. It is generally true that the uniform strain is inversely related to pipe grade. The following equation is suggested for grades up to X100 if no other proven estimation procedure is available.

22.000175.0 +⋅−= yt σε

yσ

. (E.3)

The nominal yield stress is in the unit of ksi.

E.5 References

1 Webster, S., Bannister, A., Engineering Fracture Mechanics, Vol. 67 (2000), pp. 481-514. 2 Mannucci, G., Di Vito, L., Malatesta, G., Izquierdo, A., and Cumino, G., “Evaluation of the

Effect of Yield-to-Tensile Ration on the Structural Integrity of an Offshore Pipeline by a Limit-State Design Approach,” Proceedings of the 4th International Conference on Pipeline Technology, May 9-13, 2004, Ostend, Belgium, pp. 1283.

E-4

Appendix F Incorporation of Weld Strength Mismatch

03-G78-20 F-1

03-G78-20

F.1 Background

The effects of weld strength mismatch on the weld integrity have received great attention in the last two decades. There were two international symposia organized by GKSS (Germany) dedicated to this subject [1,2]. With the trend towards using high-strength linepipes, girth weld strength undermatching has become a distinct possibility [3]. It is therefore necessary to consider how girth welds with mismatching welds should be assessed.

Early work by Wang indicated that the effects of weld strength mismatch in ECA procedures of FAD format can be accounted for by scaling the stress ratio (Sr or Lr) [4,5,6]. The newly developed European structural integrity assessment procedure SINTAP has adopted this approach [7]. Therefore, the weld strength mismatch effects can be effectively incorporated into the ECA procedure by providing a correction factor to the plastic collapse solutions,

GirthcMis

Misc f σσ = (F.1)

where is the mismatch corrected plastic collapse stress and is the mismatch correction factor.

Miscσ Misf

Extensive studies have been conducted by researchers at GKSS on the effects of weld strength mismatch on the plastic collapse loads. The results of these studies have been incorporated into a structural integrity assessment procedure termed Engineering Treatment Model, or ETM [8,9]. They did not, however, study pipes with finite length girth weld defects. The geometry that most closely matches the girth welds in pipes is pipes containing fully-circumferential surface-breaking defects. The mismatch correction factor from this geometry was found to provide good approximation by Wang [

Misf10]. The formulae for the

mismatch correction factor are given in the same reference.

F.2 Determination of Weld Width for Girth Weld

One of the key parameters in the mismatch correction factor is the weld width. The original GKSS work assumed the welds are parallel-sided. For a typical girth weld, the weld width (2H) corresponding to the defect depth may be used as the weld width in determining the mismatch correction factor, see Figure F-1.

2Η

Figure F-1 Determination of weld width 2H for a typical girth weld geometry

F-2

F.3 Suggested Approach for the Treatment of Weld Strength Mismatch

A multi-level approach may be taken in the treatment of weld strength mismatch, depending on the availability of weld property data.

Level 1: If the weld tensile property is not known, conservative assessment can be conducted by (1) using the pipe tensile property and (2) ensuring the weld metal strength even- or over-matches the pipe tensile property. This is the default condition assumed in the Level 1 assessment procedure.

Level 2: If the weld tensile property is known, but the weld profile is not known, conservative assumption can be made on the weld width and the assessment can be done by incorporating the weld strength mismatching effects. For undermatching welds, upper bound weld width should be assumed. For overmatching welds, lower bound weld width should be assumed.

Level 3: If the weld tensile property and the weld profile are known, the assessment can be done using the actual properties and dimensions.

Care should be taken when assessing defects on the fusion boundary. The lower of the base and weld metal tensile property should be taken when a single value of tensile property is needed, such as converting stress intensity factor KI to CTOD. The crack tip deformation is dominated by the lower strength material for fusion boundary defects [11]. For weld centerline defects, the tensile properties of the weld metal should be taken.

F.4 References

1 Schawalbe, K.-H., etc., Mis-Matching of Welds, First International Symposium on Weld Metal

Mis-Matching, Luneburg, Germany, April 1993. 2 Schawalbe, K.-H., etc., Mis-Matching of Welds, Second International Symposium on Weld

Metal Mis-Matching, Luneburg, Germany, April 24-26, 1996. 3 D. J. Horsley and A. G. Glover, “Girth Weld Strength Under-Matching in High Pressure

Natural Gas Pipelines,” Second International Symposium on Weld Metal Mismatching, Schawalbe, etc., Eds., Luneburg, Germany, April 24-26, 1996.

4 Wang, Y.-Y., Kirk, M. T., Gordon, J. R., and Pisarski, H. G., “Incorporating Weld Metal Mismatch into Structural Integrity Assessment,” in Pipeline Technology, Vol. 1, R. Denys, Eds., Elsevier Science B. V., 1995, pp. 475-486.

5 Wang, Y.-Y., and Kirk, M. T., “The Effect of Weld Metal Strength Mismatch and Structural Geometry on Failure Assessment Diagram,” Second International Symposium on Weld Metal Mis-Matching, Schawalbe, etc., Eds., Luneburg, Germany, April 24-26, 1996.

6 Wang, Y.-Y., and Kirk, M. T., “A Structural Assessment Procedure for Welded Structures with Weld Metal Strength Mismatch,” ASME PVP Conference, Montreal, July 22-26, 1996.

7 SINTAP Procedure, Final Version, November 1999.

03-G78-20 F-3

03-G78-20 F-4

8 Schawable, K.-H., etc., EFAM ETM 97 – The ETM Method for Assessing the Significance of

Crack-Like Defects in Engineering Structures, Comprising the Versions ETM 97/1 and ETM 96/2, GKSS 98/E/6, Geesthacht, 1998.

9 Schawable, K.-H., etc., EFAM ETM-MM 96 – The ETM Method for Assessing the Significance of Crack-Like Defects in Joints with Mechanical Heterogeneity (Strength Mismatch), GKSS 97/E/9, Geesthacht, 1997


11 Wang, Y.-Y., “Fracture Testing Procedure and Crack-Tip Fields of HAZ Cracks,” 26th National Symposium on Fracture Mechanics, Shilo Inn, Idaho Falls, ID, June, 1994.

Appendix G Example Problem for a Level 1 Option 2 Assessment

03-G78-20 G-1

G.1 Background

The purpose of this appendix is to show the steps needed for the derivation of the defect acceptance using the Level 1 Option 2 procedure. Although the procedure can be applied using “pencil and paper,” it is far more efficient with the aid of a computer.

G.2 Input Data

Pipe outer diameter: D = 36 in (914.4 mm)

Pipe wall thickness: t = 0.75 in (19.05 mm)

Young’s Modulus: E = 3.002×104 ksi (2.070×105 MPa)

Poisson’s ratio: ν = 0.3

SMYS: σy = 70 ksi (483 MPa)

Minimum CTOD Toughness: δmat = 3.9×10-3 inch (0.1 mm)

Applied stress: σa = 63 ksi (434.4 MPa)

G.3 Steps to Derive the Defect Acceptance Level

The following steps serve as an example. Other suitable steps can also be effective.

Step 1 – Start with the deepest (a/t = 0.5) crack allowed

a = 0.3750 in (9.525 mm)

Step 2 – Guess an initial crack half length (usually start with a small value)

c = 0.1969 in (5.0 mm)

Step 3 – Calculate the non-dimensionalized geometry parameters

α = D/t = 48.0

β = 2c/πD = 0.003481

η = a/t = 0.50

ηβ = 0.001741

80η/πα = 0.2653

80×0.1/πα = 0.05305

Step 4 – Calculate Kr

Y/T = 0.8740, Eq. (1), note that the unit of σy must be ksi

σf = 75.04 ksi (517.4 MPa), Eq. (2)

03-G78-20 G-2

εt = 0.09750, Eq.(10), note that the unit of σy must be ksi

n = 22.06, Eq. (9)

dn = 0.7450, Eq. (8)

m1 = -0.1776, Eq. (13b)

m2 = -1.055, Eq. (13c)

Fb = 0.09193, use Eq. (13a) as η > 0.1 and β < 80η/πα

KI = 6.286 ksi⋅in1/2 (218.4 MPa⋅mm1/2), Eq. (12)

Je = 0.001198 ksi⋅in (0.2097 MPa⋅mm), Eq. (11)

δe = 1.274×10-5 in (0.0003237 mm), Eq. (7)

Kr = 0.05690, Eq. (6)

Step 5 – Calculate Lr

σc = 68.57 ksi (472.8 MPa), use Eq. (15a) as ηβ < 0.05

Lr = σa/σc = 0.9187, Eq. (14)

Step 6 – Calculate FAC

KrFAC = 0.6822, Eq. (4), Lr is taken from Step 5

Lrcutoff = σf/σy = 1.072, Eq. (5)

Step 7 – Determine if the data point is on the FAC; and if not, select a new crack length

Lr < Lrcutoff and Kr < Kr

FAC ⇒ assessment point is inside the FAC ⇒ Increase crack length ⇒

c = 0.7874 in (20.0 mm)

Step 8 – Repeat Steps 3 – 7, until the data point is on the FAC

c = 0.7874 in (20.0 mm)

β = 2c/πD = 0.01392

┆

Kr = 0.6259

Lr = 0.9750

KrFAC = 0.6073

Lrcutoff = 1.072

Lr < Lrcutoff and Kr > Kr

FAC ⇒ Outside the FAC⇒ Decrease crack length⇒

c = 0.7469 in (18.97 mm)

03-G78-20 G-3

β = 2c/πD = 0.01321

┆

Kr = 0.6125

Lr = 0.9712

KrFAC = 0.6125

Lrcutoff = 1.072

Lr < Lrcutoff and Kr = Kr

FAC ⇒ on the FAC ⇒ Record the defect size a, c and β, η.

a = 0.375 in (9.525 mm), c = 0.7469 in (18.97 mm);

η = a/t = 0.5, β = 2c/πD = 0.01321

Step 9 – Select a new crack depth: usually decrease the depth by 0.05t

a = 0.3375 in (8.5725 mm); (η = 0.45)

Step 10 – Select an initial crack length: use the one determined in Step 7

c = 0.7469 in (18.97 mm)

Step 11 – Repeat Steps 3 – 8 to find critical crack length for the reduced crack depth

┆

a = 0.3375 in (8.5725 mm), c = 0.8232 in (20.91 mm);

η = a/t = 0.45, β = 2c/πD = 0.01456

Step 12 – Repeat Steps 9 – 11 to find critical crack length for every selected crack depth

┆

a = 0.30 in (7.620 mm), c = 0.9583 in (24.34 mm);

η = a/t = 0.4, β = 2c/πD = 0.01695

┆

a = 0.2625 in (6.668 mm), c = 1.200 in (30.49 mm);

η = a/t = 0.35, β = 2c/πD = 0.02123

┆

a = 0.2250 in (5.715 mm), c = 1.640 in (41.65 mm);

η = a/t = 0.30, β = 2c/πD = 0.02900

┆

a = 0.1875 in (4.763 mm), c = 2.442 in (62.03 mm);

03-G78-20 G-4

η = a/t = 0.25, β = 2c/πD = 0.04319

┆

a = 0.1500 in (3.81 mm), c = 3.953 in (100.4 mm);

η = a/t = 0.20, β = 2c/πD = 0.06990

┆

a = 0.1125 in (2.858 mm), c = 6.366 in (161.7 mm);

η = a/t = 0.15, β = 2c/πD = 0.1126

┆

a = 0.07500 in (1.905 mm), c = 9.598 in (243.8 mm);

η = a/t = 0.10, β = 2c/πD = 0.1697

┆

a = 0.03750 in (0.9525 mm), c = 19.75 in (501.6 mm);

η = a/t = 0.05, β = 2c/πD = 0.3492

Step 13 – Apply the safety factor

Divide all the calculated c and β by the safety factor of 1.5.

Step 14 – Create critical defect size curve, see Figure G-1

G.4 Comments and Observations

From Figure G-2 and Figure G-3, it is evident that Points 8 and 9 are on the cut-off line of the failure assessment curve. The defect acceptance levels are entirely controlled by the plastic collapse criterion. On the other hand, the other points are in the elastic-plastic fracture regime, therefore, are toughness dependent.

03-G78-20 G-5

03-G78-20

Pt.9

Pt.8

Pt.7

Pt.6

Pt.5

Pt.4

Pt.3

Pt.1

Pt.2

0.0

0.1

0.2

0.3

0.4

0.5


Allo

wab

le H

eigh

t / P

ipe

W. T

.

Toug

hnes

s D

epen

dent

Toug

hnes

s In

depe

nden

t

Figure G-1 Defect acceptance curve from the example problem with the Level 1 Option 2

procedure

Pt.1 Pt.5Pt.6

Pt.7Pt.8

Pt.9

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 1.1L r

Kr

FACExample points

Figure G-2 Illustration of the critical points on the failure assessment curve from the

example problem. Points 8 and 9 are on the cut-off line, therefore the acceptable defect sizes for those points are toughness independent.

G-6

03-G78-20

Pt.9

Pt.8Pt.7

Pt.6Pt.5Pt.1

0.0

0.2

0.4

0.6

0.8

1.0

1.2

0.8 0.9 1.0 1.1L r

Kr

FACExample points

Figure G-3 The upper right corner of Figure G-2

G-7

Appendix H Comparison of Acceptance Criteria

03-G78-20 H-1

H.1 Background

The defect acceptance criteria from the proposed procedures are compared with the acceptance criteria in the current version of API 1104 Appendix A for two sample applications.

• X70, diameter = 36 inch (914.4 mm), W.T. =0.50 inch (12.7 mm)

• X100, diameter=48 inch (1219.2 mm), W.T.=0.75 inch (19.05 mm).

For the X70 pipe, the longitudinal applied strain is 0.20%. Using the suggested procedure of Appendix E gives the applied stress of 58.5 ksi. The flow stress is estimated at 75 ksi, using Eqs. (1) and (2). Therefore, the applied stress is at 84% and 78% of the SMYS and flow stress, respectively.

For the X100 pipe, the longitudinal applied stress is at 90% of SMYS, or 90 ksi. The flow stress is estimated at 103.2 ksi, using Eqs. (1) and (2). Therefore, the applied stress is at 87% of the flow stress.

Three CTOD toughness levels were evaluated, 0.10 mm (0.004 inch), 0.127 mm (0.005 inch), and 0.254 mm (0.010 inch). The latter two CTOD values correspond to the defect height curves of Figure A-5 of API 1104 Appendix A, 19th Edition.

H.2 Comparison of Acceptance Criteria

Figure H-1 shows the defect acceptance curves of the X70 pipe from the proposed procedures at various levels of CTOD toughness with no built-in safety factor. The API 1104 curves are included for comparison. A few observations may be made:

• The acceptable length for the long defects in the current API 1104 is barely adequate in comparison to the current procedure. The safety factor in the current API 1104 is higher for short defects.

• The acceptance curves of Level 1 Option 2 are toughness dependent for relatively deep defects. The acceptance curves converge to the toughness-independent curve of Option 1 (high toughness) at different defect depths, depending on the toughness level.

• If the maximum defect height is limited to 50% of W.T., the Level 1 Option 2 curves become almost toughness-independent when the CTOD toughness is equal to or greater than approximately 0.25 mm (0.010 inch).

Figure H-2 provides comparison at a CTOD toughness of 0.254 mm (0.010 inch) with recommended safety factors in the Level 1 procedures. Figure H-3 provides comparison at the CTOD toughness levels of 0.127 mm (0.005 inch) with recommended safety factors in the Level 1 procedures. The low toughness curve of the Level 1 Option 1 procedure is below that of the Level 1 Option 2 procedure. Further comparison between Option 1 and Option 2 is given in Figure H-4 at the CTOD toughness of 0.10 mm (0.004 inch). API 1104 Appendix

03-G78-20 H-2

A cannot be used at this toughness level. The degree of conservatism in API 1104 Appendix A decreases with the decrease of toughness level, as its allowable defect lengths remain unchanged with respect to the toughness reduction.

Figure H-5 shows the defect acceptance curves of the X100 pipe from the proposed procedures at various levels of CTOD toughness with no built-in safety factor. The following observations may be made:

• The acceptable length for long defects in the current API 1104 can be much greater than that of the current procedure. The safety factor for short defects can also be an issue, depending on the toughness level. In comparison with the X70 pipe (Figure H-1), this X100 case shows that the defect acceptance level of the current API 1104 Appendix A is very much questionable. The reduced conservatism of the Appendix A in the X100 pipe, as compared to the X70, is due to (1) a higher level of applied stress and (2) higher Y/T ratio of the material.

• The acceptance curves of Level 1 Option 2 are toughness dependent for relatively deep defects. The acceptance curves converge to the toughness-independent curve of Option 1 (high toughness) at different defect depths, depending on the toughness level.

• A CTOD toughness of 0.25 mm (0.010 inch) or greater is required if the Level 1 Option 2 curve were to become toughness-independent at the maximum defect height of 50% W.T.

Figure H-6 provides comparison at a CTOD toughness of 0.254 mm (0.010 inch) with recommended safety factors in the Level 1 procedures for the X100 pipe. The high toughness curve of Option 1 is identical to that of the Option 2, except at the defect height close to the maximum allowable height (height/wall thickness = 0.5). Figure H-7 provides comparison at the CTOD toughness levels of 0.127 mm (0.005 inch) with recommended safety factors in the Level 1 procedures for the X100 pipe. The low toughness curve of the Level 1 Option 1 procedure is below that of the Level 1 Option 2 procedure. Further comparison between Option 1 and Option 2 is given in Figure H-8 at the CTOD toughness of 0.10 mm (0.004 inch). API 1104 Appendix A cannot be used at this toughness level.

The conservatism of the current API 1104 Appendix A in the application to the X100 girth welds is highly questionable. The degree of conservatism decreases with the decrease of toughness level. The welds and test data beyond X70 were not available when the acceptance criteria in the Appendix A were established. However, the use of grades X70 and above is becoming wide spread in the current and future constructions. There is an urgent need to review and update the acceptance criteria.

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D=36", W.T.=0.50", X70

Applied Strain = 0.20%

Figure H-1 Comparison of the defect acceptance levels from API 1104 Appendix A and

those of the current procedures with no safety factor on the allowable defect length

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those of the current Level 1 procedures with the recommended safety factor on the allowable defect length

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Figure H-4 Comparison of the defect acceptance levels from the current Level 1 Option 1

and Option 2

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those of the current procedures with no safety factor on the allowable defect length

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Figure H-8 Comparison of the defect acceptance levels from the current Level 1 Option 1

and Option 2

H-7

Appendix I Limits of Applicability of the Current API 1104 Appendix A Acceptance Criteria

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I.1 Background of API 1104 Appendix A

The construction of the TransAlaska oil pipeline prompted the development of fracture mechanics based alternative defect acceptance criteria [1]. The alternative defect acceptance criteria to the main body of API 1104 was first published in 1983 in its 16th Edition [2]. Certain parts of the Appendix were loosely taken from PD 6493:1980 [3], which had its origin in the CTOD design curve of Burdekin and Dawes [4]. The CTOD design curve provides a semi-empirical relation between applied strain and the CTOD driving force. The PD 6493:1980 has gone through two major revisions. One was in 1991 when it became PD 6493:1991 and again in 1999 when it became BS 7910:1999.

The alternative acceptance criteria in API 1104 are given in Appendix A in the most recently published API Standard 1104, 19th Edition. However, at the time of its first publication, there were no other appendices to API Standard 1104. Therefore, what is now Appendix A was referred to as the “Appendix.” Since the appendix has remained largely unchanged since its first adoption, the term “Appendix A” is used hereafter to refer to the alternative acceptance criteria since its first adoption.

I.2 Appendix A from the Perspective of the Code Structure

The original CTOD design curve was derived for a tension-loaded wide plate containing a small through-thickness center crack. The defect size is represented by a single parameter, i.e., the defect length. This single-parameter representation of the defect size posed a problem for girth weld defects. At least two parameters, defect height and length, are required to represent the size of a girth weld defect. In PD6493:1980, the principle of “equivalent driving force” had to be employed to evaluate defects that had more than one dimension. A defect of multi-dimension is “equivalent” to the “one dimensional” defect of a center-cracked plate if both defects produce the same driving force. When the deformation is elastic, the value of this “equivalent” defect size can be determined easily. However, this “equivalent” defect size cannot be determined easily under plastic deformation. A family of charts was provided in PD 6493:1980 that allowed the computation of critical defect length and height. CSA Z662 Appendix K adopted this approach, but with some corrections proposed by Glover and Coote based on full-scale tests [5]. In a review by Hilton and Mayville for the Office of Pipeline Safety [1], the correction was found reasonable.

In API 1104 Appendix A, the allowable defect length is 0.4D if the defect height is less than 25% of the wall thickness, or 4t if the defect height is between 25% and 50% of the pipe wall thickness. The allowable defect lengths are independent of the magnitude of the applied stress/strain and material’s toughness, apart from its aforementioned dependence on the allowable defect height. This defect length criterion was inconsistent with the actual procedure of PD 6493:1980. It is shown by a number of investigators and organizations that this criterion produces vastly different degree of conservatism when applied to pipes of different grades and applied stress/strain levels.

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I.3 Limits of Applicability from Analytical and Experimental Work Funded by API

In the early to mid-1980’s API funded work at the University of Kansas (UK) and the Welding Institute of Canada (WIC), in part to support and validate the defect acceptance criteria proposed in API 1104 Appendix A. The work at the University of Kansas consisted of numerical analysis of pipes containing largest defects allowable in the proposed appendix. One defect had a depth of 25% wall thickness and a length of 0.4D. The other defect has a depth of 50% wall thickness and a length of 4t [6]. The work at WIC consisted of one full-scale experimental test of X60 pipe with an external circumferentially-oriented flaw [5]. The pipe had a diameter of 36 inch and a nominal thickness of 0.875 inch (0.9 inch actual). The yield stress at 0.5% strain was 63.85 ksi and the ultimate tensile strength (UTS) was 89.9 ksi, which gave a Y/T ratio of 0.71. The flaw had a depth of 0.118 inch, or 0.13t. The flaw length was 13.75 inch, or 0.38D. It should be noted that this is not one of the largest flaw sizes permitted by the appendix. At a flaw length of 0.4D, the maximum allowed depth is 0.25t. The major conclusions from the UK and WIC work are as follows:

1. The UK team performed numerical simulation of the full-scale experimental test. Good agreement was obtained between the numerical simulation and experimental test. It showed that the numerical analysis provided reasonable approximation to the test.

2. Failure occurred by local buckling near the transitional weld between the X60 test pipe and the X70 carrier pipe in the experimental test. The strain in the pipe at the buckling was greater than 0.5%. Although this strain was greater than the maximum allowed strain of 0.5% in Appendix A, this test did not constitute a critical validation as the flaw size was less than that allowed under the code.

3. The UK work concluded that the “critical condition” is reached when the applied strain was in the range of 0.25-0.30%, depending on whether the CTOD toughness was taken as 0.005 or 0.010 inch. The critical condition was defined as the point when the CTOD driving force reached material’s toughness, i.e., CTOD toughness in this case.

4. A conclusion in a draft report from the University of Kansas [7], but not presented in the final report stated “the flaw acceptance criteria in Figure A5 of API-1104:Appendix A has safety factor of 1.5 or greater against fracture for long shallow crack (a/t=0.25, L/D=0.4) – provided the applied strain due to external loading is limited to 85-90% of the material yield strain. At larger applied strains, the safety factor approaches and then falls below 1.0.” The yield strain was 0.295% in the UK numerical model. This translates to a maximum applied strain of 0.27% before the acceptance criteria became non-conservative. Although the same statement was not present in the final report, the data contained in the final report supported this conclusion.

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I.4 Limits of Applicability from 1980’s Work Funded by DOT

In the mid- to late-1980’s, the U.S. Department of Transportation funded work to provide validation to the proposed alternative acceptance criteria of API 1104. The Arthur D. Little (ADL) work involved primarily literature review [1]. The National Bureau of Standards (NBS) proceeded to develop its own acceptance criteria based on rigorous engineering analysis [8]. Further analysis and experimental tests were done at Lehigh University [9].

A part of the ADL conclusions states: “We recommend that API 1104 Appendix A be altered in two

areas prior to usage as a girth weld defect tolerance standard for pipeline service conditions. First, it should be made consistent with the base standard on the topic of allowable longitudinal pipeline stresses. In addition, the standard should be altered (or new data provided) to address the apparent lack of conservatism in its application to the long flaw problem. Possible approaches for altering the standard in this area include, but not limited to, further restriction on allowable defect length, inclusion of a plastic collapse limit, and/or inclusion of a crack geometry correction to the CTOD Design curve.”

The “crack geometry correction” referred to the flaw length charts in PD 6493:1980 derived from the “equivalent driving force” principle. CSA Z662 Appendix K adopted this approach with some revisions. API 1104 Appendix A did not.

The NBS defect acceptance criteria had the targeted application of the Trans-Alaska pipeline system with API X70 material, 48-inch diameter, and 15.9-mm (0.625-inch) wall thickness. The comparison between its acceptance criteria and API 1104 Appendix A is given in Figure I-1 and Figure I-2 for the CTOD toughness of 0.005 inch (0.127 mm) and 0.010 inch (0.254 mm), respectively. The CTODRS is the CTOD driving force due to welding residual stress. The flow stress σ is an approximate value to the X70 material, and σ is the applied stress (63 ksi, or 90% of SMYS).

In comparison to the acceptance criteria proposed by NBS as shown in Figure I-1 and Figure I-2, the acceptance criteria of API 1104 for long defects are overly generous.

In the ADL work, further comparison of the acceptance criteria was conducted among several codes available at the time, as shown in Figure I-3. It is evident that none of the other codes allowed a defect length as generous as that of API 1104 Appendix A.

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Figure I-1 Comparison of allowable flaw size between API 1104 Appendix A and the

NBS criteria with an assumed CTOD toughness of 0.005 inch (0.127 mm)

Figure I-2 Comparison of allowable flaw size between API 1104 Appendix A and the NBS criteria with an assumed CTOD toughness of 0.010 inch (0.254 mm)

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Figure I-3 Comparison of allowable defect size among various codes and procedures

I.5 Observation from Historical and More Recent Work

In addition to the historical data presented above, Appendix H provides further comparison between the acceptance criteria of the current API 1104 Appendix A and the more advanced and validated acceptance criteria proposed in this document. The following overall observations can be made.

1. At low to moderate strain levels (<0.20-0.25%), the long allowable defect length (0.4D) was justified for the materials at that time. Modern linepipe materials generally exhibit lower strain hardening capacity than older materials. This is particularly true at high strength levels (X70 and above). The degree of conservatism in the current API 1104 Appendix A is progressively reduced when applied to lower strain hardening materials.

2. Based on the evidence collected so far, the validity of the long allowable defect length (0.4D) at the high end of the permissible applied strain (approximately >0.20-0.25%) has not been adequately proven. In contrast, there is a large volume of work dated as early as 20 years ago that demonstrated the potential non-conservatism of the long allowable defects at the high end of the permissible applied strain.

3. The allowable defect length in the current API 1104 Appendix A is independent of (1) applied strain, (2) CTOD toughness, and (3) material’s strain hardening rate. These parameters directly affect the critical flaw dimensions. The “de-coupling” of these input parameters and the allowable defect length in the current API 1104 Appendix A results in vastly varying degree of conservatism.

I-6

I.6 Limits of Applicability of the Current API 1104 Appendix A Acceptance Criteria

Attempts are made here to determine the limits within which the acceptance criteria in the current API 1104 Appendix A may be applied. Such limits are difficult to determine for all possible cases of application. The primary obstacle is that Appendix A does not mention a consistent theoretical, analytical, or experimental basis upon which the actual acceptance criteria were derived.

As shown in Appendix H and this appendix, the acceptance criteria of the current Appendix A tend to be overly generous and potentially non-conservative for long defects (maximum length = 0.4D criterion). At CTOD toughness of 0.25 mm (0.010 inch) or greater, the critical defect size is generally determined by plastic collapse criteria. The curves in Figure I-4 represent the critical flaw size without safety factor at different load levels. These curves were derived from the plastic collapse solution of Appendix D. Figure I-5 provides similar flaw size curves with a safety factor of 1.5 on the defect length. The permissible size for long defects in the current API 1104 Appendix A, at Length/Circumference=0.127 and Depth/W.T.=0.25, is plotted on the same figures. The maximum “safe” load level Pr is estimated at 0.77 without safety factor and 0.73 with the safety factor, respectively. The Pr level of 0.77 results in an applied stress of 85% SMYS and applied strain of 0.25% for an X70 material with the assumed stress strain relation of Appendix E. At the Pr level of 0.73, the applied stress is at 78% SMYS and the applied strain is at 0.18% for an X70 material with the assumed stress strain relation of Appendix E. The maximum ‘safe” applied stress and strain would be higher if the materials have higher strain hardening than the assumed value of Appendix E.

The above analysis confirms the findings of the University of Kansas. The defect acceptance criteria can become non-conservative when the applied strain is greater than 0.20-0.25%. The current Appendix A has a maximum allowable strain of 0.5%.

It should be noted that above limits are determined from (1) the long defect acceptance criteria (maximum length = 0.4D), (2) assumed CTOD toughness of 0.25 mm (0.010 inch) or greater, and (3) a typical X70 material. The limits of applicability would be even more restrictive if the CTOD toughness and strain hardening rate are lower than the above assumed values. Appendix H shows that the acceptance criteria in the current API 1104 Appendix A can be highly non-conservative when applied to X100 welds.

I.7 Recommendation about the Limits of Applicability of API 1104 Appendix A

For many onshore and some offshore constructions, the maximum applied longitudinal stress/strain occurs in pipe laying. The maximum applied stress is in the 80-90% SMYS range and the applied strain is in the 0.20-0.30% range. Consequently, the acceptance criteria of the current API 1104 Appendix A are safe for (1) materials with strain hardening rates that are comparable to that of early-generation X70 and (2) when the applied stress/strain does not go beyond the above mentioned pipe-laying stress and strain. This,

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plus weld strength overmatching that can be easily achieved for grades X70 or lower, may have explained the safety record of API 1104 Appendix A.

With the reduced strain hardening capacity of new TMCP linepipe materials and increased difficulty of achieving a high degree of weld strength overmatching for these high strength materials, the safety or the conservatism of the current API 1104 Appendix A cannot be demonstrated and guaranteed.

It is difficult to define the precise limits of applicability of the current API 1104 Appendix A for all possible scenarios of applications. It is clear that the maximum “safe” applied stress and strain can drop below the typical pipe laying stress and strain for modern linepipe materials. Consequently, there is no significant merit of keeping the acceptance criteria in the current Appendix A. Even within its safety limits, the structure of the current acceptance criteria results in vastly varying degree of conservatism. Furthermore, at a CTOD level below 0.25 mm (0.010 inch), the defect height can be more restrictive than necessary while allowing an unnecessarily long defect. In summary, there is no apparent advantage of using the acceptance criteria in the current API 1104 Appendix A over the acceptance criteria newly proposed in this document.

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Figure I-4 Comparison of the defect acceptance criteria from the current plastic collapse

solution with no safety factor and those of API 1104 Appendix A

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Figure I-5 Comparison of the defect acceptance criteria from the current plastic collapse

solution with a safety factor of 1.5 on the defect length and those of API 1104 Appendix A

I.8 References

1 Hilton, P. D. and Mayville, R. A., “An Evaluation of Girth Weld Defect Acceptance Criteria,”

Final report to U. S. Department of Transportation, Research and Special Projects Administration, November, 1985.

2 API 1104, Sixteenth Edition, “Appendix Alternative Standards of Acceptability for Girth Welds,” API, 1983.

3 British Standards Institution, BS 7910, “Guidance on Methods for Assessing the Acceptability of Flaws in Metallic Structures, 1999.

4 Burdekin, F. M. and Dawes, M. G., “Practical Use of Linear Elastic and Yielding Fracture Mechanics with Particular Reference to Pressure Vessels,” Proceedings of the Institute of Mechanical Engineers Conference, London, May 1971, pp. 28-37.

5 Glover, A. G., and Coote, R. I., “Full-Scale Fracture Tests of Pipeline Girth Welds,” in Circumferential Cracks in Pressure Vessels and Piping-Vol. II, PVP Vol. 95, G. M. Wilkowski, Eds, 1984, pp. 107-121

6 Dodds, Jr., R. H., Attiogbe, E. K., Vargas, P. M., “Crack Opening Displacements in Pipes Containing a Part-Through Circumferential Flaw,” Report No. CRINC-SM-19, the University of Kansas final report to API, September 1987.

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7 Dodds, Jr., R. H., Attiogbe, E. K., Vargas, P. M., “Crack Opening Displacements in Pipes

Containing a Part-Through Circumferential Flaw Subjected to Bending,” SM Report No. 12, The University of Kansas, July 1987.

8 Reed, R. P., Kasen, M. B., McHenry, H. I., Fortunko, C. M., and Read, D. T., “Fitness-for-Service Criteria for Pipeline Girth Weld Quality,” final report to Materials Transportation Bureau, U.S. Department of Transportation, November 1, 1983.


Documents

A Comprehensive Update in the Evaluation of Pipeline Weld ......Girth weld defect acceptance criteria are set and enforced in all pipeline constructions in the U.S. per federal regulations