WELD-DEFECTS-PART-A

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Draft for review A - 1 Part A

ASSESSMENT OF PIPELINE GIRTH WELD FLAWS

EXTENSION OF EPRG-TIER 2 GUIDELINES

ON ALLOWABLE FLAW SIZES

Part A –Study of girth weld specific ECAs

1 ENGINEERING CRITICAL ASSESSMENT (ECA) 1.1 Background

Over the past several years, a number of specific Engineering Critical Assessment (ECA) approaches have been developed to evaluate the fitness-for-purpose of pipeline girth weld containing fabrication flaws [1-9]. To date, these methodologies can be applied when the applied stresses in the axial direction remain within their elastic stress-strain response limits. The primary input parameters for a conventional ECA analysis are: flaw type and size, level of applied stress, and toughness of the material containing the flaw. Here, it should be emphasized that an ECA is critically dependent on the ability to reliably find and accurately determine the size and location of any flaw that is present in the weld area. Current ECA methodologies do not provide the desired guidance for the evaluation of the interdependent input factors affecting girth weld integrity. Fortunately, the critical (failure) flaw sizes are normally larger than those predicted by the mathematical model. This is because:

• The prediction assume that the material is homogeneous while the weld metal is

normally stronger than the pipe metal (weld metal mis-match effect) • Current fracture toughness (CTOD) requirements are derived from specimens

simulating full constraint whilst girth weld flaws allow relief of constraint. The reason is that line pipe walls are normally too thin to cause a plane strain condition. In addition, unless the operating temperature is very low, the toughness of pipeline girth weld is normally adequate to prevent elastic fractures.

• flaws are represented by their containment rectangle and, this assumption increases

the level of conservatism for the possible and very likely instance of irregular-shaped flaws.

A better understanding of the preceding factors would make the assessment more accurate. However, these factors still do not take into account all aspects of the problem. Fig A1 gives a general overview of the complexities associated with an ECA. Although existing analyses do not address the many complex interactions, the challenge remains to understand and to find the right balance between the many factors affecting the integrity of pipeline girth welds.

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Fig. A1 – Overview of factors affecting girth weld integrity

Finally, it might be noted that the authorities do not always accept ECA based allowable flaw sizes. Beyond the psychological impact, this problem can only be overcome if the mathematical model used is supplemented with experimental evidence demonstrating the validity of the predicted flaw sizes. For the case where failure is controlled by yield and plastic collapse, such evidence is now available as a result of comprehensive experimental investigations [10-18].

1.2 Standard assessment

A standard ECA approach ensures that the calculated flaw size is acceptable both from a (brittle/unstable) fracture and a stable (plastic collapse or ductile failure) failure viewpoint:

• For the fracture (or toughness dominated) assessment, a fracture mechanics CTOD

based design curve is used to ensure that failure by unstable fracture will not occur for known flaw dimensions, applied stress and CTOD toughness.

• The plastic collapse (or flaw size dominated) assessment prevents failure by local or

overall yielding of the flawed cross section. For that purpose, flaw size and applied stress are compared with a notional flow stress, derived from the tensile properties of the flawed region.

How to measure ?Charpy V or CTOD ?

Practical relevance of CTOD test

Pipe to weld metal YS mismatch ?

Pipe and weld metalproperties

GIRTH WELDINTEGRITY

Inspection techniqueX-ray and / or AUT

Pipe geometry

Minimumtoughness and tearing

behaviour ?

Welding Processes

Level of weld metal yield strength mismatch ?

Surface-breaking vs buried flawsAccuracy in height sizing ?

Interaction rules ?

Diameter and wallthickness

Pipe and weld metal Y/T ratiosin longitudinal direction

Pipe grade

SMAW, GMAW, P-GMAW, FCAW ?(Type of flaws)

Performance requirement: ?

Workmanship and / or ECA based

weld acceptance criteria ?

Allowable Allowable Allowable Allowable flaw size limitsflaw size limitsflaw size limitsflaw size limits

3

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The fracture assessment finds it origin in the early 1970’s. At that time, achieving adequate weld metal toughness properties was a real challenge. With advances in welding technology, brittle fracture is nowadays only a concern for very low temperature applications. Research over the last several years has shown that the primary failure mode of flawed girth welds is by plastic collapse. Moreover, large-scale tests show that flawed girth welds can exceed the pipe metal yield strength and fail at strain levels many times larger than the strain at the yield point [11].

1.3 Input parameters The quantitative nature of an ECA analyses implies a need for quantitative input data. Safe ECA predictions are obtained when the driving force, which causes a flaw to extend, and toughness, which is the resistance of the material containing the flaw(s) to this extension, and the flaw size can be adequately determined. This implies that any assessment procedure modelling the relationship between material properties, applied load, flaw size and toughness represents the actual situation. In deterministic terms, this means that:

• The most onerous applied stresses or strains acting on the flawed weld are

conservatively estimated, • The lower bound toughness of the flawed region can be determined, • The tensile properties of the girth weld meets those of the pipe metal (Weld metal yield

strength mis-match), • The actual flaw dimensions and, in particular, the height of a crack-like flaw as well as

its position can be accurately measured (Flaw sizing).

The above observations are simply meant to say that uncertainties surrounding the values of the input parameters will be transmitted through to the final result. Furthermore, an ECA explicitly assumes that the provisions/assumptions used in the derivation of the calculated/predicted flaw size limit are adhered to during construction while the mechanical and toughness properties across and along the weld are consistent. Therefore, flaw acceptance criteria must be based on “achievable” material properties and proven inspection techniques. Aside from the above requirements, the equation(s) used to calculate the allowable flaw size should be:

• Technically sound as well as reliable, and experimentally validated • Simple and practical • Acceptable to pipeline owners and pipeline operators • Acceptable to relevant jurisdictional authorities

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It should also be emphasized that more accurate solutions can be obtained if realistic flaw interaction criteria for assessing the effect of adjacent flaws and suitable recategorisation procedures of near surface breaking flaws should be used.

1.4 Limitations and challenges Current ECA methodologies have limitations when addressing higher strength pipes or higher strain applications. In particular, with the extension of pipeline applications to high strength pipeline steels (X80 and above), more severe service design conditions, and the introduction of limit state design requirements, current ECA approaches are not equipped to handle this type of problems. For example, for non-elastic design conditions, the level of weld metal yield strength mismatch is a factor that requires due consideration in an ECA analysis. An undermatched girth weld may be a potential problem that limits the overall deformation capacity. However, existing ECA procedures do not provide adequate guidance for this design situation.

2 FAILURE MODES

Girth weld failure caused by a pre-existing flaw can occur in the loading range from brittle fracture (applied stresses below yield - brittle facture) to plastic collapse (overload of remaining ligament or remote pipe section - ductile failure). Between these two extremes failure occurs by contained yielding, commonly indicated as elastic-plastic fracture behaviour). Brittle fracture (elastic or elastic-plastic behaviour) occurs when the yielding at the flaw tip is contained. To prevent brittle fracture the pipeline industry includes in design codes toughness requirements and weld quality inspection criteria. Failure by plastic collapse requires that a lower bound limit on toughness. The point is that flawed girth welds can be made sufficiently tough so that a relaxation of the triaxial state of stress ahead of the flaw tip is easily achieved. Thus, the failure condition of a flawed pipeline girth welds can be described by the tensile (flow stress) properties of the materials in the weld region. In a girth weld containing a part-wall flaw, plastic collapse or ductile failure can occur by Ligament (local) yielding, Global yielding (or Net Section Yielding, NSY) or Overall (or Gross, pipe metal, Section Yielding, GSY), Fig. A2.

• Local (or ligament) collapse occurs when yielding is confined to the ligament ahead of the

flaw tip and the opposite back face (Fig. A2 – left figure). • Global collapse (or Net Section Yielding - NSY) occurs when the complete cross sectional

area containing the flaw becomes plastic. In this case, yielding is uncontained. Yielding of the entire cross section is only possible when the weld is subjected to tensile loading (Fig. A2 – middle).

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• Overall collapse (or Gross Section Yielding – GSY) occurs when the applied stress in the remote cross sections exceeds the yield strength of the pipe. GSY is only possible if the flawed material has strain hardening capabilities. Strain hardening acts to spread the extent of the plasticity away from the flawed cross section (Fig. A2 – right figure).

Fig. A2 - Deformation modes at plastic collapse

For strain hardening materials, the three plastic collapse modes can merge from one mode into the other with increasing load. Local collapse is normally not the limiting deformation mode since the materials used in modern pipelines can easily meet the minimum toughness for collapse by NSY. The occurrence of NSY or GSY collapse depends upon the level of applied stress, the flaw size and location, the level of weld metal yield strength mismatch and the strain hardening capability (or Y/T ratio) of the material containing the flaw. The transition from NSY to GSY depends essentially on flaw size and strain hardening (or Y/T ratio).

2.1 Brittle Fracture Assessment

The assessment of flaw severity depends essentially on the toughness properties of the flawed material whereas the procedures of elastic (KIc-approach) or elastic-plastic (CTOD-approach) fracture mechanics can be applied to assess flaw criticality. The allowable flaw size can be derived from the following generic equation:

Required (available) TOUGHNESS =

(Allowable) Flaw size x Applied remote (failure) stress

A A

Cross Section AA

Overall Overall Overall Overall (GSY)(GSY)(GSY)(GSY)

Collapse Collapse Collapse Collapse

Plastic Deformation

SurfaceFlaw

Global Global Global Global (NSY)(NSY)(NSY)(NSY)

Collapse Collapse Collapse Collapse

Local Collapse

Global versus Overallxxxx Plastic Collapse

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Above a certain level of toughness the (critical) values of stress and flaw size are insensitive to toughness. On the other hand, for new constructions, codes or fabrication documents require the pipe/weld metal toughness levels warranting plastic collapse as failure mode. Thus, flawed girth welds in properly designed pipelines have the ability to deform plastically. The point is that girth welds can be made sufficiently tough so that the failure conditions of flawed girth welds are governed by the flow properties of the material.

2.2 Plastic Collapse Assessment For applied remote stresses below yield, plastic collapse predictions use the flaw dimensions and the remote (nominal) applied stress to calculate the stress level developed in the flawed (net) cross section. This stress level is compared with a notional flow stress derived from the tensile properties of the material containing the flaw:

Several plastic collapse assessment models have been proposed which can be grouped into three categories [19-25]. The first group embraces the models based on a theoretical analysis of flat plates. The second group involves those based on a theoretical analysis of actual pipe geometries. The final group of models is those derived from experimental results [24-25]. Standard plastic collapse assessments assume that the plasticity is confined to the flawed cross section (Net Section Yielding). For design conditions requiring that the applied loads cause plastic deformations in the remote pipe sections, existing plastic solutions cannot be used.

2.3 Brittle fracture or plastic collapse? Conventional ECA approaches do not exclude failure by brittle fracture. This assumption is not very relevant for new pipeline constructions. Flawed girth welds in properly designed pipelines fail by plastic collapse because:

• as a result of the understanding of brittle fracture problem, improvements were made in

weld metals and weld procedures. • the selection of the weld metal is based on the assumption that undetected

(workmanship-type) weld flaws may be present. This assumption requires that the toughness of the weld metal and HAZ regions should be sufficiently high to exclude the possibility of a brittle fracture.

Applied remote (failure) stress =

FLOW STRESS x (Allowable) Flaw size

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• it is now common practice to specify filler metals with a Charpy test requirement to ensure adequate resistance to brittle fracture.

Having excluded the occurrence of brittle fracture, the assessment of a flawed girth welds can be concentrated on plastic collapse. In particular, an ECA for pipeline girth welds can be simplified since the failure conditions are controlled by geometrical factors (wall thickness and pipe diameter) and material properties (pipe and weld metal strengths) rather than by toughness. Stated in another way, beyond a toughness threshold, increased toughness does not increase the flaw size limit. This is because plastic collapse solutions relating the applied stress to the flow stress and flaw size do not include a toughness parameter.

2.4 Threshold toughness for plastic collapse

EPRG (European Pipe Research Group) has used experimental correlations to determine the lower bound Charpy V notch (CVN) impact values for plastic collapse by pipe metal yielding (performance requirement = 0,5% pipe metal strain, GSY). The required CVN impact energies for collapse controlled failure of matched / overmatched welds at minimum operating temperature are [26]:

Lowest individual value: 30 J Mean value (of 3 specimens): 40 J

The toughness requirements are based on correlations between the results of curved wide plate tests incorporating 3 mm deep surface breaking root cracks and Charpy V impact tests. The Universiteit Gent database that was used for the EPRG study included girth welds made in X52, X60, X65 and X70 pipeline steels in thickness ranging from 6,6 mm to 25,4 mm. The correlations have also demonstrated that for a wall thickness less than 12,7 mm, the impact requirement can be reduced to 27 J (minimum individual value: 20 J) [27]

2.5 Correlation between Charpy V and CTOD toughness For practical reasons, the CTOD fracture toughness cannot always be measured. For example, it is virtually impossible to determine a representative CTOD value of girth welds in older pipelines. For this and other similar situation, the Charpy V impact values are normally available. However, to satisfy the possible need for CTOD data and to overcome the practical difficulties, a number of attempts have been undertaken to correlate Charpy V impact energy and CTOD toughness. Contrary to some opinions, reasonable correlations can be found provided that care is taken that both tests sample the same region of the weld deposit [28]. Such a correlation is not yet available for pipeline girth welds. To provide a workable basis for the prediction of CTOD using CVN test data for girth welds, a substantial volume of pipeline specific CVN and CTOD results produced by Labo Soete - RUG were compared with the aim to identify a lower bound correlation, Fig. A3. The plots in Fig. A3 contain comparisons between both minimum (open diamonds) and mean values (solid diamonds). The test results were taken from a database containing a broad range of weld

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Fig. A3 – Correlation between mean and minimum CVN and CTOD properties

(a) CVN < 40 J

(b) CVN > 40 J

CTOD = 0.00118 CVN

0.00

0.05

0.10

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0.30

0 40 80 120

CVN (J)

CTO

D (m

m)

Average values

Minimum values

UM Welds

Labo Soete - G ent

CTODave = 0.0019CVNave - 0.0287

0

0.2

0.4

0.6

0.8

0 50 100 150 200 250 300

CVN (J)

CTO

D (m

m)

Lower bound mini values Lower bound average values

Minimum values Average values

CVN = 40 J

CVN = 40 J (EPRG)CTOD = 0,0473 mm

Labo Soete - Gent

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metal strength and toughness levels (see also Part B) [29]. The CVN and CTOD values were measured at the same temperature while the mean values were determined for at least three specimens. The Charpy specimens were taken from the subsurface weld root region.

At first sight, and as expected, a straightforward correlation between CVN and CTOD cannot be made. The large scatter is due to fact that the data was not grouped by welding process and wall thickness. The other reasons for the poor correlations are discussed in [29]. However, one might conclude that no universal CVN-CTOD correlations can be found. By fitting a lower bound envelope (straight line) to the experimental data, the Charpy V impact energy can be used to estimate the lower bound CTOD toughness of girth welds. The following equations give a conservative description of the correlation between the CVN and CTOD properties of pipeline girth welds:

CVN00118.0aveCTOD = (A1)

0287.0aveCVN0019.0aveCTOD −= (A2)

Eq. (A1) was derived for Charpy impact properties lower than 40 J (Fig. A3a). For Charpy values exceeding 40 J, Eq. (A2) can be used (Fig. A3b) It should be noted that Eqs. (A1) and (A2) gives the same value if the CVN impact energy is equal to 40 J. In using Eqs. (A1) and (A2) the fracture toughness, CVN, or the notch toughness, CTOD, are to be determined at the same test temperature. Thus, where the CTOD test is impractical, the lower bound weld metal CTOD toughness may conveniently be estimated from Charpy impact data using an empirical formulation provided by Eqs. (A1) and (A2). It should be noted here that the equations should not be used for wall thickness exceeding 25.4 mm.

Another expression developed by the Japanese Welding Society yields similar predictions. The Japanese fitness-for-purpose document WES 2805-1980, provides an expression that allows the CTOD to be derived from CVN data [31]. The correlation is based on the results of experiments conducted on plate, weld metal and HAZ tests of various kinds of steels:

tVNC01.0CTOD = (A3)

In this equation, CVNt is the CVN energy at the temperature of

)kgmin(t5YS112T −−+

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where t = plate thickness, YS = yield strength and T = design temperature or CTOD testing temperature (°C). For material with a yield strength of 550 MPa (X80) and a wall thickness of 25 mm to be used at 0 °C, Eq. (A3) gives for a CVN value of 40 J to be obtained at +31 °C a CTOD value of 0.04 mm at 0 °C. This requirement compares well with that derived from Esq. (A1) or (A2).

2.5.1 Charpy versus CTOD testing

For a Charpy V value of 40 J, Eq. (A1) allows concluding that the assessment of flaw in a girth weld with a CTOD toughness of 0,0473 mm can be based on plastic collapse. Using full-scale test results on circumferential part-wall flaws Hopkins proposed a CTOD of 0,05 mm (without safety factor) as the “threshold’ toughness above which failure can be based on plastic collapse [32-33]. A supplementary PD 6493:1991 fracture mechanics calculation by Hopkins for buried flaws of any length in pipeline girth welds having a CTOD toughness of 0,05 mm led to the conclusion that brittle fracture at stresses up to yield strength can be excluded provide their height does not exceed 3 mm [26]. Coote [twi p21] arrived at a similar conclusion at the time he assessed the Canadian full scale test results which were used to validate the CSA Z662 Appendix K ECA approach. In particular he concluded, “…high strength pipe with CTOD toughness exceeding 0.10 mm, maximum allowable imperfection sizes will be limited by the analysis to prevent plastic collapse…”.

2.5.2 Practical conclusion

The previous observations allow concluding that the inexpensive Charpy V test and the CTOD test are equivalent test methods. In addition, the CVN impact requirements of 30 J / 40 J can be used as a simple criterion to define the boundary between fracture and plastic collapse controlled fracture. In particular, the CTOD test is not needed to determine the failure mode if, as demonstrated, an acceptable empirical correlation between CTOD and Charpy can be applied. On the other hand, the possibility of expressing CTOD as a function of CVN is attractive for situation where CTOD data is not available. However, note that CTOD values derived from CVN-CTOD correlations provide lower bound data, and thus lower bound ECA based allowable flaw sizes. Therefore, such correlations cannot always be a substitute for the CTOD test if failure behaviour is toughness dependent

3 PLASTIC COLLAPSE 3.1 Background

The design of pipelines under conditions normally encountered in the oil and gas industry is based on pressure containment, bending and thermally induced stresses. These loads cause hoop and axial/longitudinal stresses. Codes limit the allowable stresses to prevent failure.

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The criteria are stress-based and limit the maximum allowable hoop stress (or design factor) to 72 or 80% of the specified minimum yield strength (SMYS). The maximum axial stress level is normally restricted to 50 % of the SMYS. Using stress-based design criteria means that the pipeline is not exposed to plastic strains since the pipeline industry uses a strain limit of 0.5%. This limit coincides with the strain at which the value of the SMYS is defined.

3.2 Plastic collapse of flawed girth welds The failure characteristics of a flawed girth weld subject to bending can be derived from the failure characteristics of a tensile loaded curved pipe segment. This simplification is based on large scale test results which have demonstrated that tensile loaded curved wide plate (CWP) specimens fail at lower strain and stress levels than full scale bend test pieces do, Fig. A4 [10,18].

Fig. A4 - Comparison of full-scale bend and curved wide plate (CWP) test results Conventional, stress-based, plastic collapse solutions presume that the limit state occurs when the average net section stress or the limit stress for failure reaches some notional flow stress, FS. This assumption is reasonable for large flaws, but it can be unrealistic, i.e. conservative, for “small” flaws in strain hardening materials. The effect of strain hardening, which can be characterised in a simplified way by the yield-to-tensile ratio, Y/T, is counted on to delay the localisation of strain in the net section until remote (pipe metal) yielding is achieved. As shown in Fig. A4, the actual limit stress can exceed the predicted stress for failure. It is further worth of note that the level of weld metal yield strength mis-match also interferes in this deformation process. Currently, little practical guidance can be given to determine the

Full-scale bend vs Curved wide plate tensile test results

0

0.25

0.5

0.75

-70 -50 -30 -10 10 30

Test temperature (C)

Failu

re s

trai

n (%

)

Curved wide plate tests

Full scale bend tests

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combinations of Y/T ratio and weld metal yield strength mis-match on the limit stress for failure.

3.3 Plastic Collapse solutions 3.3.1 Basic equation

Conventional plastic collapse solutions use the flaw dimensions and the remote (nominal) applied stress to calculate the stress in the flawed (net) cross section. This stress is then compared with a notional flow stress, σf or FS, derived from the stress-strain tensile behaviour of the material containing the flaw. For strain hardening materials, the value of the flow stress exceeds the uniaxial yield strength. When the net section stress acting on a flawed pipe segment of arc length, s (or plate width, W), is equal to the flow stress, plastic collapse is obtained for an applied remote (gross) tensile stress given by:

]ts

hl1[FS)remote(appliedpc −=σ=σ (A4)

In this generic equation, σpc is the applied gross tensile stress at plastic collapse, FS is the flow stress, l is the flaw length, h is the flaw height, s (or W) is the arc length and t is the wall thickness. The flow stress is an artificial material property for calculating the strength at plastic collapse of a flawed ductile strain hardening material. The flow stress is at least equal to the yield strength and its value is situated between the yield and ultimate tensile strength values (see also next Section). Eq. A4 shows that the reduction in limit load for collapse is simply proportional to the reduction in cross sectional area caused by the flaw. However, Fig. A5 illustrates that the actual relationship is not linear.

In Fig A5, the predicted failure stresses, using Eq. (A4) (dashed line), is compared with the measured failure stresses (solid and open data points) for a 10 mm thick pipeline steel. As can be seen, the predictions (dashed line) provide a comfortable margin of safety (as expressed by the difference between the predicted and measured failure stresses – data points above the dashed line are safely predicted) for plates failing by GSY (solid circles). Fig. A5 also illustrates that a simple plastic collapse assessment can greatly underestimate the limit load in the case of GSY. When the applied stress reaches the flow stress, the pipe metal has, depending on the flaw size additional load carrying capacity. The margin of safety decreases for NSY (open circles). On the other hand, the predicted failure stresses can be unconservative for materials having a high propensity to stable ductile tearing in the NSY deformation mode [34].

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Fig. A5 – Effect of flaw length on deformation behaviour

3.3.2 Flow stress Several numerical definitions of flow stress have been developed. Some flow stress equations are a function of Yield Strength (YS) and Ultimate Tensile Strength (UTS), while the others depend on yield or tensile strength. The published expressions are based either on theoretical considerations and or on empirical correlations [35]. For general fitness-for-purpose and, in particular, for pipeline girth weld flaw assessments, the flow stress is usually defined as the average of the yield and ultimate tensile strengths, Eq. (A5) or 1,2.YS whichever is less (BS-PD 6493):

2TSYS FS += (A5)

With R = YS/TS, the above flow stress equation can be expressed and characterized in terms of yield strength and the yield to tensile ratio, R:

R2)R1(YS FS += (A6)

Depending on the stress-strain properties of the material, and in particular of the Y/T ratio, the flow stress given by Eq. (A6) will vary in its magnitude.

3.3.3 Flow stress and weld metal yield strength mismatch The differences in yield strength between the pipe metal and weld metal (yield strength mis-

Actual failure curve

500

550

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0 100 200Defect length (mm)

Rem

ote

stre

ss a

t max

. Loa

d (M

Pa))

Shift GSY to NSY

Measured values - NSY

Measured values - GSY

YS

FS

Predicted failure curve

GSY

NSY

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nmatch) influences the plastic deformation behaviour of a flawed girth weld. However, yield strength mis-match in the weld region complicates the prediction of the conditions of plastic collapse. BS 7910-2000, Section 7.2.8, states that in welded joints the tensile properties of the region in which the flaw is located should be used. For HAZ zone regions it should be assumed that the material has the lower of the adjacent weld metal or parent metal properties for application to assessment of flaws in structures. The accuracy of this recommendation can be investigated by comparing the predictions against experimental results. Fig. A6 compares the experimental data with the predictions. In calculating the failure stress for collapse, the flow stresses of both the pipe and weld metals have been used. The experimental data were taken from a curved wide plate (CWP) test database containing 382 results. The test specimens contained a single surface breaking flaw in either an undermatched or an overmatch girth weld (the range of weld metal yield strength mismatch varied from –25 % to 54 %) while the base metal consisted of X52 up to X80 pipeline steel grades [30]. The data points represent, as indicated, the fracture or NSY (failure strain < 0.5 %) and GSY (failure strain > 0.5 %) failure mode. Figs A6a and A6b do not make a distinction between under and overmatched welds. However, these data are re-plotted for matched/overmatched and undermatched welds in Figs. A6a1, a2 (M/OM) and Fig.A6b1 and b2 (UM).

The comparison of the test results (“Measured”) with the calculated (“Predicted”) failure stresses using Eq. (A4), illustrates that the failure stresses are conservatively predicted if the flow stress of the weakest material is used. More specifically, the comparisons show quite clearly that the collapse behaviour of undermatched welds should be assessed on the basis of the flow stress of the girth weld, Fig. A6a. For the situation of flawed overmatched welds, the flow stress of the pipe metal must be used to obtain conservative predictions. In other words, the use of the weld metal flow stress could lead to non-conservative predictions. This also implies that BS7910-2000 provides inaccurate information for the assessment of flaws containted in overmatched weld metal.

3.3.4 Plastic collapse solutions considered

Various formulae have been proposed to predict the collapse or ‘theoretical’ limit load of a flawed material [20,22,23,36]. In an attempt to provide guidance on the selection of on the most appropriate formula for application to flaws in ductile pipeline girth welds, Wang et al [23] have assessed a number of possible solutions that have been developed for various loading conditions, or a combination thereof. These solutions are not considered here for the simple reason that they have not been developed for HSLA/TMCP pipeline steels while experimental validation of their usefulness is rather limited.

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Fig. A6 - Comparison of measured and predicted stresses for plastic collapse of

undermatched and overmatched girth welds (a) input: weld metal flow stress (b) input: pipe metal flow stress

UM and OM welds

400

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400 500 600 700 800

Measured stress (MPa)

Pred

icte

d st

ress

(MPa

)

Failure strain > 0,5 % Failure strain < 0,5 %

FS of Weld metal

(a)

Unconservative

Conservative

UM and OM welds

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Pred

icte

d st

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(MPa

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Failure strain > 0.5 %Failure strain < 0.5 %

FS of Pipe metal

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FS of Weld metal

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Pred

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Failure strain < 0,5 % Failure strain > 0.5 %

FS of Weld metal

OK !

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Pred

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FS of Pipe metal

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Pred

icte

d st

ress

(MPa

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Failure strain > 0.5 % Failure strain < 0.5 %

FS of Pipe metal

(b2)

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The Net Section Collapse (NSC) approach is also excluded in the present review because the NSC analysis assumes that failure occurs once the entire cross section containing the flaw becomes plastic. For pipeline girth welds under bending the NSC approach will give on-conservative predictions because the stresses occurring in pipe vary, depending on the circumferential position, from tensile to compression. For all these reasons, the discussions below are concentrated on experimentally validated collapse solutions. These solutions are the DEN model (used in ERPG-Tier 2) and the modified-ligament instability model (adopted by CSA Z662 Appendix K). As will be discussed in the next Sections, these solutions are simple but conservative since the loading condition considered is uniform tension; i.e. the models assume that a uniform tensile load equal to the greatest tensile stress occurs within a well-defined fraction of the pipe circumference that contains the flaw. This is a conservative simplification of the actual loading that girth weld flaws might experience since the most critical loading for a circumferentially oriented flaw in a pipeline girth weld is bending. DEN collapse model - The model is a “curved” plate solution that is extensively used by Laboratorium Soete, Universiteit Gent, to analyse the deformation behaviour of curved wide plate tests. This model combines Eqs. (A4) and (A6). That is, Eq. (A4) can be converted to the following equation:

]stlh - [1

R2)R1(YS =

pcapplied+

σσ = (A7)

This equation has been used in the derivation of the EPRG-Tier 2 flaw acceptance levels (see Section 4.3). Note further that Eq. (A7) enables the quantification of the effect of YS/TS ratio, wall thickness, level of applied stress (less than or equal to yield strength), weld-mismatch, flaw height (or length) and flaw length (or height) to be assessed. Modified-ligament instability model - The modified-ligament instability model proposed by Coote et al [25,37] is based on the assumption that failure should occur once the plastic zone size in the section containing the flaw extends to 35.4 degrees around the pipe circumference (i.e., 10 percent of the pipe circumference). This model, adjusted by using the results of full-scale bend tests, has been adopted in the Canadian Standard Z662, Appendix K. The details of this model are described in Section 4.2.

3.3.5 Girth weld performance requirement under axial loading Failure of flaw free pipe welds under axial loading requires plastic tensile strains. As discussed, provided the strength of the weld matches/overmatches that of the pipe, girth welds containing (small) flaws can yield (strain > 0.5 %) in sections remote from the flaw (Fig. A5). Performance criterion - For the situation where pipe yielding is a credible event it is practical to require GSY (applied strain ≥ 0.5 %) as a demonstration of satisfactory per-formance. This requirement is conceptually different from a fracture mechanics or NSY based

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plastic collapse based assessment since the GSY failure criterion can be used for fitness-for-purpose assessments without some form of safety (see Section 5.4). Back to the basics - The idea of using the GSY requirement as failure criterion was first explored in the 1960’s. At that time, wide plate test results were compared to a semi-arbitrary pass/fail strain level of 0,5 % (four times yield strain). The 0.5% yield criterion was used to provide an answer for a particular problem related to local strain embrittlement in “old type pressure vessel steels” [38]. The necessity for a consistent but simple and practical engineering means for assessing either or both elastic/plastic or plastic material behaviour in wide plate test performance led Gent University in the early 70’s to the development of the Gross Section Yielding (GSY) concept [39-41]. The concept reflects the idea that when the material at the crack tip can strain harden enough to compensate for the missing cross sectional area in the plane of the flaw, the applied strain can be (uniformly) distributed all along the welded joint prior to failure. GSY and welds - For flawed girth welds, GSY is achieved when the gross/remote section fracture stress exceeds the pipe metal yield strength. It is easier to obtain GSY for overmatched welds than for undermatched ones. For overmatched welds containing large flaws it is not to be excluded that the plastic strains will be confined to the weld metal region. For undermatched welds, the GSY requirement can be satisfied when the weld metal possesses adequate strain hardening capability and provided the flaw size is rather small. Benchmark - The fact that the GSY plastic collapse model has been validated by testing some 800 large-scale specimens, the GSY criterion can be used as the benchmark in the ensuing discussions.

4 REVIEW OF CURRENTLY USED ECA METHODOLOGIES

The girth weld specific ECA methodologies most frequently used are detailed in API 1104 Appendix A, CSA Z662 Appendix K and EPRG-Tier 2 [3-4,8]. Before discussing the approaches, the following points are worth mentioning:

• the methodologies generate an “allowable” flaw size and not a critical flaw size, • detailed guidelines on NDE methods are not given (flaw sizing issue)

• none of the methodologies contain procedures for assessing the effect of weld metal

mismatch on allowable flaw size (material’s issue). The mentioned ECA approaches are discussed in the following Sections. The effects of flaw sizing and weld metal yield strength mismatch on allowable flaw size are discussed in Sections 6.

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4.1 API 1104 – Appendix A

The API 1104 Appendix A approach is based on the original COD design-curve concept. However, an adjustment is made to the flaw size to account for the potential for fatigue-crack growth of the flaw during the life of the pipeline. The amount of fatigue crack growth is calculated by integrating the Paris law expression. API 1104 Appendix A assumes failure by brittle fracture and, therefore, does not include a plastic-collapse check on the allowable flaw size. The total applied strain in the API 1104 Appendix A approach is the sum of the applied strain, as determined from a stress analysis, and the residual strain. The allowance for residual strain is assumed to be equal to 0.2 % strain for all materials. API 1104 Appendix A requires that the CTOD toughness be measured at the minimum operating temperature of the pipeline. In addition, the weld procedure must be qualified to one of two CTOD fracture-toughness values, 0.127 mm (0.005 inch) or 0.254 mm (0.010 inch) CTOD. It is necessary for the measured CTOD to meet the minimum of 0.127 mm (0.005 inch) in order to use the Appendix A procedures. API 1104 Appendix A further stipulates that the flaw length be less than 40 percent of the pipe diameter for flaws less than or equal to 25 percent of the pipe wall thickness in height, and less than four times the pipe wall thickness for flaws between 25 and 50 percent of the pipe wall thickness in height. Flaws exceeding 50 percent of the wall thickness are not allowed.

4.2 CSA Z662 – Appendix K

The Canadian Standard CSA Z662-Appendix K includes a fracture and a plastic collapse analysis. The allowable flaw length is the lesser of the maximum allowable flaw length to prevent brittle fracture and the maximum allowable flaw length to prevent plastic collapse. The brittle-fracture assessment is based on the procedures of PD6493-1980. However, specific adjustments were made to the assessment. Using the results of full-scale test, the CTOD design-curve concept has been empirically modified to account for the apparent lack of consistent safety margins for the case of long, shallow surface flaws [42]. The brittle fracture approach does not require any determination or inclusion of the welding residual stresses in the analysis. The justification for the exclusion of welding residual stresses in the analysis is that the girth weld toughness is normally sufficient to exclude a brittle or cleavage fracture.

The plastic collapse assessment used in CSA Z662 Appendix K is based on a modified ligament instability model in which failure is assumed to occur once the plastic zone size in the section containing the flaw extends past approximately 10 % of the pipe circumference. To achieve a factor of safety on the assessment of flaw size a factor of 2.0 is applied on flaw

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height and 2.5 on flaw length. By applying the above approach to the results of full-scale pipe bend tests, the following linear equation for plastic collapse K was derived, Eq. (A8):

]tDhl18 - 03[1, SMYSpc π

=σ (A8)

where l = allowable flaw length, h = flaw height, D = pipe diameter, t = wall thickness, SMYS = Specified Minimum Yield Strength of the pipe metal and σpc is the applied stress at plastic collapse. In addition, CSA Z662 Appendix K stipulates that the maximum height of the flaw must be restricted to one-half the pipe wall thickness. These flaw size restrictions are imposed as the modified ligament instability model has only been checked experimentally against the results of full-scale bend tests, which had flaws within these limits. Eq. (A8) shows that plastic collapse is related to the SMYS (specified minimum yield strength) and thus neglects the effect of Y/T ratio. The factor of 1.03 is an empirical value that reflects the curve fit effects. The use of the SMYS avoids ambiguity in defining the flow stress while the SMYS is a guaranteed value that is not subject to variability. This choice is partly influenced by the fact that the prediction of the allowable flaw length, l (or L2max in terms of Appendix K), using a flow stress of 1,03.YS rather than the actual yield strength, provides a minimum safety factor of 1,5 on stress for the full scale test data which were used to verify this solution. The CSA approach specifies that the weld toughness shall be determined by CTOD testing at the minimum design temperature. Note that, unlike the API approach, the CSA approach does not require a minimum value of toughness (CTOD). Furthermore, it is stipulated that two weld-metal tensile-test specimens shall be tested and that the yield strength of each specimen shall be equal to or greater than the specified minimum yield strength of the parent pipe material, i.e., the weld metal strength shall match/overmatch the parent material strength. Finally, unlike its American counterpart, API 1104 Appendix A, a fatigue analysis is not considered necessary for the CSA approach since the CSA Z662 Appendix K approach is only applicable to gas pipelines where the cyclic loading on girth welds in gas pipelines is extremely low1.

1 CSA approach for the liquid-pipeline, the maximum height of a flaw is restricted to 25% percent of the pipe-wall thickness

compared to the 50 % restriction imposed in the gas pipeline standard. This additional restriction on the flaw depth for liquid

lines is a guard against fatigue-crack growth. If the operator of the liquid pipeline wishes to allow flaws greater than 25 percent

of the wall thickness in depth but less than 50 percent of the wall thickness, then they must show by analysis that the subject

flaw will not grow by fatigue.

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4.3 EPRG 4.3.1 Background

Prior to the development of the European Pipeline Research Group (EPRG) ECA guidelines on allowable flaw sizes in pipeline girth welds, the discrepancies and limitations between the above standards were carefully assessed [43]. As a result of this study, a novel philosophy for setting allowable fabrication flaw size limits was adopted. EPRG offers the opportunity for an ECA analysis two levels of conservatism. The first level (Tier 2) is a simple assessment method with a built in safety factor. The second level (Tier 3) allows using “User Specific Flaw Acceptance Levels”. Note that the discussion of the Tier 3 approach is beyond the present discussion. Also note that Tier 1 specifies the Workmanship flaw acceptance levels. The EPRG approach is somewhat different than the two previously discussed approaches because flaw assessment according to ERPG-Tier 2 is uniquely based on plastic collapse while API 1104 and CSA Z662 do not exclude brittle fracture. The EPRG-Tier 2 guidelines require a girth weld to meet a minimum toughness (30J minimum / 40J mean), and provided this toughness requirement is met, allowable flaw sizes are derived from pipe yielding as limit load.

4.3.2 EPRG-Tier 2 The basis for EPRG-Tier 2 flaw acceptance is the simple DEN plastic collapse model, Eq. (A7), and the performance criterion described in Section 3.3.4. For pipe yielding (GSY), σpc = YS, Eq. (A7) can be converted to the following equation:

]ts

hl - [1 R2

)R1(YSYS += (A9)

The maximum allowable flaw length for GSY is obtained by inverting Eq. (A9):

hts

R1)R1(l

+−= (A10)

Eq. (A10) assumes that collapse occurs if the pipe yields over a fixed arc length of “s” mm. Curved wide plate test results have shown that conservative predictions are obtained for an arc length, s, of 300 mm [24]. This length corresponds to approximately 10 % of the circumference of a large diameter pipe. Regardless the NDE method used, EPRG-Tier 2 specifies that no flaw shall have a through thickness height in excess of 3.0 mm. That is, it is assumed that the average flaw height of natural flaws is less than 3.0 mm.

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For a YS/TS ratio (R) of 0.87 and an arc length of 300 mm, the maximum allowable length for a fixed flaw height of 3.0 mm as a function of pipe wall thickness, t, using Eq. (A10), is:

l = 6,95t (A11)

The flaw geometry used in the above calculations is assumed to be rectangular. This simplifying assumption, which is in line with the 'containment rectangle' approach used in current flaw assessment, is conservative because natural flaws are irregularly shaped. Therefore, EPRG have decided to set the allowable flaw length at 7t. This length gives a flaw area limit of 7 % per 300 mm length of weld. By using the maximum allowable flaw size, 7t x 3.0 mm2, it is assumed that a Charpy toughness of 40 J average, 30 J minimum is achieved at the design temperature, and the weld metal is matching or overmatching. Note further that EPRG-Tier 2 does not require CTOD toughness testing since the minimum toughness for plastic collapse can be derived from Charpy V impact tests.

4.4 ECA approaches not considered

The above ECA approaches are not the only codified approaches for assessing the serviceability of flaws in girth weld. For example, the Failure Assessment Diagram (FAD) approach, BS 7910-2000 (revised PD6491-1991), API RP579, etc…, can also be used in a fitness-for-service assessment. Those approaches are included in this review because:

• they are general purpose approaches which assess the likelihood of fracture and plastic overload simultaneously

• the procedures are not restricted to the specific case of girth welds.

• they have not yet been, as far as the present authors know, extensively validated for

flaws in pipeline girth welds.

However, the major concern is that the assessment for plastic collapse does not represent the particular configuration of flawed girth welds in large diameter pipelines.

5 INTERPRETATIVE COMPARISON OF THE METHODOLOGIES

Appendix A of API 1104 (further denoted as API) and CSA Z662 Appendix K (CSA) are now, in terms of length of weld assessed, the most widely used methods of performing fitness-for-service assessments. The limited application of the ERPG approach is most likely due the imposed restriction on allowable flaw height (3 mm). Prior to their use, the primary input parameters to be considered in an ECA are: toughness of the region (weld metal or HAZ) containing the flaw, material properties of the weld region,

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applied stress, and flaw type and size. Since current ECA methodologies are deterministic based, the input parameters are chosen on the basis of lower bound values [44-46].

The last but not the least point of note is that the dimensions of flaw (length and height), and its through-wall position must be established by an appropriate non-destructive-inspection technique before an ECA can be used. The use of conventional radiography is adequate for measurement of flaw length but is insufficient for determining flaw height. The use of ultrasonics for determining flaw height is acceptable provided the accuracy of the inspection procedure has been previously established and any potential inaccuracy is included in the analysis.

5.1 Comparison of input parameters

As discussed, API, CSA and EPRG-Tier 2 (further denoted as EPRG) provide guidelines for alternative flaw acceptance criteria for pipeline girth welds, together with requirements of greater stringency for welding. These girth weld specific ECA’s cannot directly be compared to their qualities. Each of these approaches use different failure models and require different input requirements. For discussion purposes, the most important provisions and limitations are compared in Table A1. In examining Table A1, one can easily conclude that each code will generate different flaw size limits depending on the incidental combination of pipe diameter, wall thickness, tough-ness and applied stress. The striking differences are briefly outlined hereinafter. Section 5.2. illustrates the differences by means of by a sample application in which the allowable flaw sizes are determined using the procedures defined in each of the standards. Table A1 illustrates that:

• The approaches use different applied stresses/strains inputs. EPRG is based on remote

yielding at a fixed applied strain of 0.5 %. API and CSA require the user to determine the applied strain (stress). The calculated applied strain is usually elastic (< 0.5 %).

• API includes a residual stress correction that is already embodied in the assessment

curve. Both CSA and EPRG ignore the effect of residual stress in the analysis. The neglect of residual stresses is consistent with the experimental findings that residual stresses do not affect ductile failure.

• The CSA and EPRG approaches, because of their derivation, provide a factor of large

safety. The authors of CSA introduced a margin of safety on flaw size. The conservatism of EPRG is based on the GSY performance criterion.

• The failure criterion of API does not require a plastic collapse assessment. This limitation

is not, as illustrated in Section 2.5, very logical because the API toughness requirements (0.127 mm or 0.254 mm) ensure plastic collapse while the failure model assumes “brittle fracture”. A related point of note is that API ignores the imposed limitation incorporated in the application of the CTOD design curve [47]. Thus, by ignoring a plastic collapse

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API 1104

Appendix A (198??)*

CSA Z 662 Appendix K

(198?)*

EPRG Tier 2

(1994)

APPLIED STRESS/STRAIN User input User input Fixed value

FAILURE CRITERION

- Fracture asssessment

- Plastic collapse assessment

Yes

No

Yes

Yes

No

Yes

FRACTURE ASSESSMENT

- Residual stress/stress

- Applied stress/strain

0.2 %

0.5 % (max)

No

SMYS (max)

No

0,5 %

PLASTIC COLLAPSE ASSESSMENT

- Flow stress

- Model

-

-

1.03 SMYS

Modif. Strip yield

(YS + TS)/2

RD - RUGent

TOUGHNESS TESTING

- CVN testing

- CTOD testing

No

Yes

40 J

Yes

40 J / 30 J

No

WELD METAL TESTING

- Cross tensile

- All weld metal tensile

- YS Matching Requirement

Yes

No

No

Yes

Yes

Yes

Yes

Yes

Yes

TOUGHNESS REQUIREMENTS

- Minimum CTOD

- Minimum CVN

0.127 mm

-

No

40 J

-

40 J / 30J

MAXIMUN FLAW DIMENSIONS

- Length (D = diameter)

- Height (T= wall thickness)

0.40 D

0.50 t

0.10 πD

0.50 t

7 t

3 mm

CORRECTION ERROR IN FLAW SIZING Yes Incorprated No

EXPERIMENTAL VALIDATION

- Full-scale tests

- Curved wide plate tests

Limited

No

Adequate (60+)

No

A few

Many (400 +)

“RATING” Can be

unconservative Very safe Conservative

* = First edition, later revisions

Table 1 – General comparison of ECA methodologies

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assessment, API could be non-conservative if applied to the case of tough welds. In such a case, the resistance of a flawed weld to failure is not related to toughness, but instead, depends solely on the flow stress of the “weakest” material (see Section 3.3.3).

• Each of the ECA approaches is applicable for applied strain up to the onset remote

yielding (GSY). Although strain hardening (or the Y/T ratio) plays a crucial role in obtaining remote yielding, neither API nor CSA address this effect. EPRG solves this issue by using a strain hardening dependent flow stress.

• Both API and CSA require that the welding procedure must be qualified for CTOD

toughness. API specifies a minimum toughness of 0.127 mm (the toughness level of 0.254 mm generates larger allowable flaw size). CSA does not establish a minimum CTOD level. However, the CSA Charpy requirement of 40 J implies that an effective minimum CTOD must exist. EPRG does require CTOD testing. Toughness testing for EPRG (Tier 2) is confined to Charpy testing.

• API assumes that the weld is “matching” if the tensile strength of the composite, as

measured in the cross (transverse weld) tensile test, meets the minimum specified tensile strength. CSA stipulates that the yield strength of the cross-tensile specimen (without the reinforcement in place) is at least equal to the SMYS of the pipe. EPRG requires information on the weld metal yield strength. In addition, ERPG requires matching (overmatching) weld metal is demonstrated by all-weld metal tests.

• API requires the user to establish the accuracy of the NDE technique to be used and thus

to apply a safety factor to the flaw height to account for potential flaw sizing inaccuracies.

• API and CSA give maximum flaw lengths proportional to pipe diameter. The maximum

allowable flaw length of EPRG is proportional to wall thickness

• Flaws with a height greater than one half the wall thickness are not allowed for API and CSA. On the other hand, flaw height in EPRG is restricted to 3.0 mm. The height limitation specified by EPRG is a deterrent to its use if the user has to account for potential flaw sizing errors. These errors are indirectly accounted for in the CSA approach since the allowable flaw size curve embodies safety factors on length and height.

• The reliability of the CSA and EPRG assessment procedures has been established by

extensive series of large-scale tests, and for CSA, by more than 10 years of ‘field’ experience.

The above observations demonstrate that an ECA based flaw assessment can only be applied if the “relevant” input parameters are well defined. Furthermore, it should be emphasized that the welding procedures(s) qualified should ensure that good workmanship could be maintained.

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Another potential limitation common to any of the ECA approaches is that, if the production welders fail to strictly apply the welding procedures previously qualified, the end result could be production welds that have lower toughness values than qualified. This observation emphasizes the fact that setting ‘high’ minimum toughness levels should be avoided.

The other concern is associated with the requirement that the weld-metal yield strength matches (overmatches) the pipe metal strength. In practical weldments the tensile characteristics of the weld deposit differ from those of the base metal. The level of weld-metal strength matching is a significant factor for flaw behaviour. If the yield strength of the weld zone does not exceed the yield strength of the pipe material, then the applied strain might be concentrated in the weld region. This important issue is discussed in Section 6.

5.2 Comparison of approaches

As a further aid in showing the differences between the ECA approaches discussed above, graphical comparisons are presented for a series of sample cases, Fig. A7. Each plots shows the allowable flaw length(s) versus the allowable flaw length(s) for a typical range of pipe diameter and wall thickness combinations. The comparisons are restricted to 0,5% applied strain and relate to the three approaches as applicable. Consequently, the comparisons are focussed on failure by plastic collapse as it is implicitly assumed that the weld metal toughness is adequate to ensure collapse by GSY.

5.2.1 Comparison of API, CSA and EPRG allowable flaw sizes Allowable flaw size curves have been generated for the cases of 40” diameter x 20 mm wall and 20” diameter x 12.7 mm (and x 7.5 mm) wall X80 pipe. The resulting plots presented in Fig. A7 show that the predicted maximum allowable flaw sizes are not consistent. Note also that, both CSA and EPRG include a ‘default’ limitation on flaw length. It should also be recalled that the API and CSA predictions are independent of the Y/T ratio. On the other hand, EPRG limits the Y/T ratio, that is, the EPRG allowable flaw size cannot be applied for Y/T ratios exceeding 0.90. CSA and API do not stipulate such a restriction.

API gives the largest allowable flaw sizes because the CTOD design, used to derive the API curves, does not account for the plastic collapse and allows that high CTOD values give proportionally large allowable flaw sizes. In order words, without additional plastic collapse constraints, API may predict non-conservative estimates of allowable flaw sizes. CWP tests have recently demonstrated this fact [18]. Both CSA and EPRG give conservative predictions for each of the cases considered here. The smallest allowable flaw sizes are predicted by the CSA approach. The difference between API and CSA is very significant for “thick” wall pipe while the difference becomes smaller for “thin” wall pipe. For 7.5 mm wall pipe, API becomes restrictive on allowable length for flaw heights exceeding the average height (3 mm) of a weld bead. In contrast, the limitation on flaw length for EPRG is proportional to wall thickness. Of further note is the fact that CSA is the least sensitive to the effect of wall thickness. Another way of viewing the comparisons is

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that the CSA approach, as applicable, cannot be used for applied strains of yield magnitude. In particular, it cannot be excluded that the maximum flaw size allowed by CSA could be less than that allowed by workmanship standards.

5.2.2 Comparison of the CSA and DEN approaches The maximum flaw sizes preventing plastic collapse using the CSA approach are derived from the collapse model given by Eq. (A8). As discussed, Eq. (A8) incorporates a safety factor on flaw length of 2.5 and a safety factor on flaw depth of 2.0. Stripping Eq. (A8) of these safety factors gives Eq. (A12):

]tD

2h

2.5l18

- [1,03 SMYSσpc π=

]tD847.0

hl - [1 SMYS03.1=σpc (A12)

where l = maximum flaw length, h = maximum flaw height, R = Y/T ratio, D = pipe diameter (D = 2πr, r = pipe radius), t = wall thickness, SMYS = (Specified Minimum) Yield Strength. By comparing Eq (A8) - CSA approach with safety factor - to Eq. (A12) - CSA approach without safety factors - for an applied strain of yield magnitude, it can be appreciated from Fig. A8a that the safety margins on the allowable flaw sizes (CSA Z662 Appendix K approach) are quite large for the case of high stress. Fig. A8b compares the allowable flaw size curves for collapse by GSY using the DEN model, Eq (A10), and the CSA approach without safety factors (Eq. (A12)). The DEN predictions are shown for two Y/T ratios. Note again that the comparisons are for an applied stress of yield magnitude (failure criterion: GSY). Examination of Fig. A8b shows that the differences between the CSA and the DEN predictions become smaller if one removes the safety factors contained in CSA Z662 Appendix K approach. Depending on the Y/T ratio, the DEN collapse criterion generates either less restrictive (Y/T = 0.80) or more conservative (Y/T = 0.90) flaw sizes than those generated by CSA without safety factors. In this connection, it should be recalled that in deriving the EPRG allowable flaw sizes the Y/T ratio was set at 0.90.

Summarised, accepting the criticism of repetition, it must be stated once more that, the mathematical models used to derive the CSA Z662 Appendix K and EPRG allowable flaw sizes, have been ‘calibrated’ by using results of large-scale tests. However, it must also be emphasized that the large majority of these large-scale tests were performed on large diameter pipeline. This means that there is no direct comparison possible for small diameter pipelines. The question whether of similar comparisons for ‘small’ diameter pipe will yield similar conclusions is clearly of great importance and it is hoped that further research will enable this to be clarified.

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Fig. A7- Comparison allowable flaw sizes for three ECA approaches

Pipe Grade X80 - Y/T = 0.90Pipe: 40" x 20 mm - Requirement: Pipe Yielding (GSY)

0

3

6

9

12

0 100 200 300 400 500

Flaw length (mm)

Flaw

hei

ght (

mm

)

CSA Z662

API 1104

EPRG

Y/T independent

Pipe Grade X80 - Y/T = 0.90Pipe: 20" x 12.7 mm - Requirement: Pipe Yielding (GSY)

0

3

6

0 100 200 300

Flaw length (mm)

Flaw

hei

ght (

mm

)

CSA Z662

API 1104

EPRG

Y/T independent

Pipe Grade X80 - Y/T = 0.90Pipe: 20" x 7.5 mm - Requirement: Pipe Yielding (GSY)

0

3

6

0 100 200 300

Flaw length (mm)

Flaw

hei

ght (

mm

)

CSA Z662

API 1104EPRG

Y/T independent

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5.3 Discussion 5.3.1 Performance criterion

The preceding comparisons have been focussed on the ‘base line’ case of GSY or pipe metal yielding. The comparisons have to be placed in another perspective for axial the stresses below yield. Using the DEN model, Fig. A8 shows the effect of applied stress on allowable flaw size for GSY and an elastically strained girth weld. An axial stress of 0.80% SMYS exceeds the actual stress. The actual stress is generally less than half the hoop stress. As indicated, elastically stressed girth welds are assumed to fail by NSY.

Fig. A8 - Comparison of CSA and DEN collapse models

Pipe Grade X70 - Y/T = 0.80 and 0.90Pipe: 40" x 20 mm - Requirement: Pipe Yielding (GSY)

0

3

6

9

12

0 100 200 300 400 500

Flaw length (mm)

Flaw

hei

ght (

mm

)

CSA without Safety Factors

DEN-MODELY/T = 0,80

DEN-MODELY/T = 0,90


0

3

6

9

12

0 100 200 300 400 500

Flaw length (mm)

Flaw

hei

ght (

mm

)

CSA without Safety FactorCSA Z662

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One will note from that Fig. A9 that the GSY failure criterion (applied remote axial stress = SMYS) includes a substantial margin of safety when compared to the case of an elastically stressed girth weld. Furthermore, it is worthwhile noting here the flaw length limit (50 mm) allowed by workmanship acceptance criteria ensures failure by GSY whereas the flaw-size limit for GSY is smaller than for NSY. Fig. A9 can also be used to explain why the incidence of girth weld failure is very low.

Despite the observation that the GSY criterion be restrictive in terms of allowable flaw size, the use of the GSY failure criterion is recommended. The GSY requirement eliminates a number of possible uncertainties, namely:

Fig. A9 - Effect of the applied stress on the calculated flaw size limit

• the GSY failure criterion eliminates the need for detailed analyses to determine the

applied stress in the axial direction. • GSY is a conservative and safe assumption for conventional (elastic) pipeline designs

because the criterion incorporates a quantifiable safety margin for conventional designs while a degree of safety is also provided for stresses arising from sources other than those used in design.

• possible flaw size measurement inaccuracies become less critical with GSY. In

contrast, for NSY flaw sizing might become a point of discussion since the safety margin, or the ratio between the predicted and the critical flaw size, can be small. In passing, it may also be pointed out that any deviation of the qualified properties might reduce the assumed safety margin.

0

2

4

6

0 100 200 300 400 500

Defect length (mm)

Def

ect h

eigh

t (m

m)

Applied stress = 0,80.SMYS

Applied stress = SMYS

Workmanshiplimit

(max. 50 mm)

GSYNSY

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5.3.2 Assessment of small diameter pipelines The collapse solution used by CSA embodies the effect of pipe diameter and was essentially developed for failure by NSY. Moreover, the plastic collapse analysis in Appendix K is very restrictive at stresses of yield magnitude. However, the CSA approach avoids the flaw sizing uncertainties by using a safety factor on flaw height and flaw length in the collapse equation while the flow stress only depends on the pipe metal yield strength. Except for very high values of Y/T ratios, the value of the CSA flow stress is also smaller than that used by EPRG. These differences generate different allowable flaw sizes, but the sample applications shown in Fig. A8 have illustrated that CSA and EPRG provide a comparable margin of safety for large diameter pipelines. The conservatism incorporated into the original CSA collapse solution (Eq. (A8)) can be reduced by replacing the ‘virtual’ value of the CSA flow stress, 1.03 SMYS, by a Y/T depended flow stress and by eliminating the factor of safety on flaw length. For example, substituting Eq. (A6) into Eq. (A8), the modified CSA collapse solution (further denoted MLM-collapse criterion) is given by Eq. (A13):

]tDhl18 - 03[1. SMYSpc π

=σ (A8)

]tD169.0

hl - [1 R2

)R1(SMYSpc

+=σ (A13)

In Figs. A10 and A11, the original CSA and EPRG maximum allowable flaw size curves are compared to the DEN (Eq. (A9) and the MLM collapse model with and without (Eq. (A13) the safety factor of 2.5 on length flaw. The comparisons are made for a representative set of pipe diameter, wall thickness and Y/T ratio combinations. All plots have been generated for failure by GSY. The comparisons of the various curves lead to the following observations:

• the relative position of the curves depends on the Y/T ratio, pipe diameter and wall

thickness.

• as Figs. 10 and A11 indicate, the MLM model with safety factor on flaw size (curve decorated with open circles) gives for the low Y/T ratio case (Y/T = 0,80) a substantial increase in flaw size. For thin wall and high Y/T ratio (Y/T = 0.90) pipe, the increase in flaw size is less spectacular.

• for small diameter pipe, Fig. A11, the differences between the MLM and the DEN-

predictions are due to the difference in the treatment of the extent of yielding. The DEN assessment gives the least conservative predictions.

• the comparison of the relative positions of the flaw curve predicted by the MLM model

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without the safety factor on flaw length (curve decorated with solid circles) and the DEN analyses shows that the DEN model gives conservative results for large diameter pipe. For small diameter pipe, the maximum flaw sizes derived form the MLM model without a length correction are more restrictive than the DEN prediction. This is not surprising since the MLM model is based on the assumption that the extent of yielding at plastic failure is proportional to the pipe diameter. The DEN model uses a fixed arc length of 300 mm irrespective the pipe diameter. Except for the case of large diameter pipe, this assumption could be non-conservative for small diameter pipe.

Fig. A10 - Comparison allowable flaw sizes and collapse models for large diameter pipe.

Summarised, the above observations would justify the use of the MLM collapse model without length correction, Eq. (A13), as it offers improvements over the existing EPRG and CSA collapse assessment procedures. Although a number of practical aspects must be resolved, Eq. (A13) could also be used to reduce the safety margin embodied in CSA Z662 Appendix K approach and to formulate an EPRG solution for predicting the collapse conditions of small


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diameter pipe. However, the latter contention needs to be validated since appropriate large-scale test data is not yet available.

Fig. A11 - Comparison allowable flaw sizes and collapse models for small diameter pipe. 6 GENERAL REQUIREMENTS

Prior to the use the ERPG-Tier 2 based girth weld flaw assessment must be met. A stress analysis is required to determine the maximum applied tensile stress acting on the girth weld. However, since the GSY criterion is the performance requirement one has only to verify that the design stress in the axial direction does not exceed the pipe metal yield strength. The

Weld qualification testing includes tensile testing to determine the yield strength of both pipe and weld metal (see Section 6.1), and toughness testing of the weldment to determine the


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minimum fracture toughness of the weld region. The Charpy V notch toughness tests are to be determined at the minimum design temperature. A related, but very important consideration in applying an ECA is the control of the welding procedure. The essential welding variables have to be carefully monitored with the intention of minimizing the variation in mechanical properties of the welds. Thus, the weld filler metals and weld procedures must be carefully selected in order to produce girth welds with an optimum balance of strength and toughness. A NDT method must be used to be capable of determining the maximum flaw height in order to determine the maximum allowable length of the detected flaw. However, EPRG assumes that the height of individual flaws has a height equal to the maximum depth of the corresponding weld pass. Thus, as presently developed, EPRG cannot be applied to flaw heights exceeding 3.0 mm [48] (see Part B if this report).

6.1 Weld Mechanical properties

It is undesirable for any region of the weld to contain low strength material. Low strength or undermatched welds are prone to concentrate applied deformations [49-52]. Therefore, weld metal testing is required to verify that the actual yield strength of the weld is at least equal to the specified minimum yield strength of the pipe [53-55].

6.1.1 Weld metal yield strength mismatch effects

Specifications assume that the weld metal has the same properties as the pipe metal. This assumption is not correct [56]. On the other hand, current plastic collapse assessments do not specifically address the effect of weld metal undermatching or overmatching on flaw tolerance. A standard ECA approach is focussed on flaws, which are assumed to be located in a matched (or overmatched) weld. This assumption is rarely a problem for the common lower grades of pipeline steels. However, in the event that the axial strains of yield magnitude occur, this assumption might lead to unsafe predictions since undermatched welds tend to localise plastic strains within the weld zone so that failure might occur at strain levels lower predicted strain.

The degree of weld-metal under or overmatching is crucial in a plastic collapse assessment. In particular, the choice between an undermatched or overmatched weld cannot be dissociated from the weldability aspects, the desired margin of safety, the flaw characteristics (surface versus embedded flaws) and the efficiency of the NDT inspection method used. If the weld metal yield strength exceeds the applied stress, elastic designs do not exclude the use of undermatching weld. In this case, however, the user should be aware of the consequences since undermatched welds have a lower flaw tolerance than overmatched welds, Fig. A12. The ‘lower’ flaw tolerance of undermatched welds implies that greater attention needs to be paid to flaw sizing. This implies that undermatched welds require precise, high accuracy NDT inspection. The issue of flaw sizing is discussed in the next Section.

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Conversely, overmatching protects weld flaws from large strains, as plasticity will occur in the pipe material. The shielding effect provided by an overmatched weld allows, as compared to undermatched weld, larger flaws for similar levels of applied remote strain. This effect applies especially for elastic designs. Equally, overmatched welds permit less rigorous flaw sizing requirements. For these reasons, it is attractive to require weld metal yield strength overmatching2. On the other hand, in order to prevent failure at strain levels below yield (< 0,5 %), moderately undermatched welds exposed to plastic straining require a high strain hardening capability (or a low YS/TS ratio).

Fig. A12 – Effect of weld metal yield strength mismatch on allowable flaw size (schematic) 6.1.2 Unintentional undermatch

Even when an overmatched weld metal is selected it is still possible that some girth welds are undermatched [58-61]. This is because the distributions representing the scatter in tensile properties of the base and weld metal can overlap, Fig. 13. Fig. A13 shows that it is easier to obtain overmatching in mechanised GMAW welds while the spread of the yield strength properties is narrower than for SMAW welds. Furthermore, Fig. A13 illustrates that a variable degree of pipe to weld mismatch may be observed in SMAW girth welds. Moreover, for pipe grades X70 and above, the use of cellulosic electrodes does not automatically exclude weld metal yield strength undermatching. Since girth welds in high strength pipelines may unknowingly be undermatched because the pipe material has much higher yield strengths than the SMYS, specific guidelines on achieving matched/overmatched girth welds are needed.

2 Overmatching can be ensured when the minimum weld metal yield strength exceeds the specified minimum yield strength

of the pipe by 60 MPa (SMYS + 60 MPa) [57] and by a proper control of the welding parameters.

Wall thickness = 20 mmY/T = 0.85

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Fig. A13 - Effect of SMAW and GMAW weld metal variation on weld metal mismatch (Hatched area represents the undermatching density)

6.1.3 Pipe Metal Properties

Variability of pipe properties is inherent in the manufacturing process. It depends on many factors, including sophistication of process control, technology applied, and whether the mill is integrated or manufactures steel from scrap. Of most interest, when considering mismatch of girth welds, are the properties in the longitudinal (pipe axis) direction. The tensile properties in the axial direction were and are not often specified. Therefore, historical data is not commonly available from pipe manufacturers. Correlation to hoop direction properties is difficult due to the varying effects described above, and because it depends on other factors such as pipe diameter, whether the pipe is longitudinal or spiral formed, cold expansion, test specimen geometry (round bar vs. full section bar), etc. It is therefore advisable to either specify, or independently measure, the pipe tensile properties in the axial direction, if an accurate measure of weld mismatch is to be determined. The range of yield strength, in any particular order of pipe, can be as high as 100 MPa. This range can increase for thinner wall pipe. A further complication is that Y/T ratio tends to increase as plate thickness decreases []. Moreover, as common large diameter pipe grades are now higher, pipe manufacturers tend to produce similar steel for many pipe grades. The result for “low grade’ pipe is that the range of yield strengths might exceed the standard strength distribution.

0

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470 520 570 620 670Yield strength (MPa)

CEN X70 PIPE METALAve: 545 MPa

Max/min. = 480-600

GMAW (85 tests - 4 projects)

Ave: 635 MPaMax/min. = 580-680

SMAW (127 tests - 15 projects)

Ave: 562 MPaMax/min. = 474 - 638

UM

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6.1.4 Weld Metal Properties

Weld metal strength is affected by process factors such as heat input, travel speed, and welding current [62-65]. These factors will vary from weld to weld due to differences in technique between welders. As well, the width of the weld affects the amount of dilution from the parent metal; the properties will therefore vary in the through thickness direction, and according to the o'clock position around the pipe [59]. Studies have shown that the lowest weld metal strength will be measured at the bottom of the pipe (6 o'clock). The levels of yield strength mismatch are normally quite variable in manual SMAW girth welds. Mechanized GMAW girth welds have an advantage over manual welds because they are more repeatable. The spread of properties is narrower than for manual welds [60]. Weld metal yield strength testing is usually done on 6.4 mm diameter all-weld metal specimens. This provides only a sample from near the middle of the weld, and may not be representative of the strength of the entire weld. A rough estimate of weld metal strength can be obtained from micro hardness measurements [7] and is particularly useful if a lower strength electrode is used for the root pass. Another approach consists of comparing the results of edge notched pipe metal and edge notched transverse weld tensile specimens [8].

6.2 Flaw characteristics

Flaw acceptance based on an ECA analysis requires information on both flaw length and height of the detected flaw, and on its location (depth) within the weld. In other words, it is essential to consider NDT as an integral part of an ECA. Since the actual dimensions and the through-thickness position of a girth weld flaw (surface breaking vs. embedded) affect its significance, it is necessary to consider how flaw sizing, recategorization of an embedded flaw and flaw interaction can be dealt with (Fig. A14). Since girth weld flaws might have complex geometries, there are three areas where the designer should have answers:

• Can X-ray inspection be used in an ECA analysis? • How effective is ultrasonic testing (UT) for girth-weld inspection?

• What are the achievable levels of accuracy in flaw sizing?

Little published literature is available to provide straightforward answers to these questions [66-70]. However, ultrasonic testing (UT) is the only method capable of measuring the ‘invisible” height and depth dimensions at present. Whilst UT is capable of providing an estimate of through-wall flaw dimensions, consideration must be given to the accuracy of

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such measurements. Radiography can be used to measure flaw length. However, radiography is of little use in assessing flaw height from conventional radiographs under field conditions.

Fig. A14 – Factors involved in flaw sizing and flaw characterisation

The hurdle that stands in an ECA is the quest for certainty in UT flaw sizing accuracy. This accuracy of flaw sizing cannot always be quantified. It is not possible to say with complete certainty that a flaw will be larger or smaller than its estimated (measured) size. That is, a ‘unique’ value for flaw height (or depth) sizing inaccuracy cannot be given as this value depends on the welding process, flaw shape and flaw through-wall position, and ultimately on the calculated/predicted flaw size limit. Also, the sizing error might have a larger effect on the accuracy of the predictions than the choice of the collapse model. The required precision may not be possible or necessary [70]. This observation is not meant to place a disproportionate emphasis on flaw height-sizing accuracy because, there are other factors influencing the required accuracy. For example, the level of weld yield strength overmatch can offset the error in flaw sizing.

6.2.1 Sizing effects on allowable flaw size

Significant errors in the prediction of the flaw size limits can be made if the flaw size is underestimated, in this case the conservatism of the ECA analysis is eroded and the margin of safety is reduced. An overestimate of the flaw size will increase the degree of conservatism and may result in an unnecessary repair of flaws that would have been innocuous in reality. Moreover, an ECA analysis based on conservative assumptions can predict allowable flaw sizes smaller than the workmanship acceptance levels. For example, this problem might occur for girth weld flaws in thin wall pipes.

FLAW

DETERMINEINSPECTION ACCURACY

CharacterisationSize/dimensions ?

Type ?

Surface breaking

Embedded

Height

Length

Depth

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Recategorisation(near surface

breaking flaws)

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The user must be address these issues and safety factors should be applied to account for potential sizing inaccuracies. As it is potentially unsafe to assume that a flaw has the dimensions predicted by UT, the inaccuracy in height sizing has to be accounted for by adding a correction factor equal to the inspection error, ∆h, to the calculated flaw dimensions. The direct implication is that the dimensions of the allowable flaw should be smaller than its calculated limit dimensions.

The API approach requires such an approach in that it requires the user to establish in advance the accuracy of the NDE technique to be used, and to apply a safety factor to the flaw height to account for potential inaccuracies. For CSA, the uncertainty in flaw height sizing is directly accounted for in the applied collapse model.

6.2.2 Illustration The effects of a flaw height correction on the calculated flaw size are shown in Fig. A15. The calculated flaw length-height curves (see thick solid lines) apply to surface breaking flaws in 20.0 mm thick pipes (Eq. (A10)). The calculations have been performed for two Y/T ratio values (0.80 and 0.90). These curves delineate the boundary between GSY (flaw length-height combinations below the curve) and NSY (flaw dimensions above the curve) failure behaviour. Furthermore, this simple assessment illustrates that the sizing error is minor problem for low Y/T ratio pipeline steels because the condition for plastic collapse of an overmatched girth weld depends on the flow stress of the weakest material (see Section 3.3.3 – Fig. A6)

The sizing problem can also be approached from another perspective. Historical data and experimental evidence available demonstrates that very large planar flaws/cold cracks in older (X-ray inspected) pipelines have been missed and thus have unwillingly been accepted. Even so, with the use of the best NDT techniques, or a combination thereof, some ‘large’ flaws will be missed. Judged from these observations and the evidence of long service without failure, one can use the rigorous concept of probability or incorporate some elements of the reliability theory in an ECA analysis. However, a serious warning should be given against an indiscriminate use of such methods. Currently, few data exist in sufficient quantity to permit a statistical analysis of the many factors (wall thickness, flaw location, flaw type, flaw profile, welding process, weld bevel preparation, inspection procedure, etc..) affecting flaw-sizing accuracy.

6.2.3 Dimensions of detected flaws

The flaw interaction and flaw recategorisation criteria currently being used are very, if not extremely, conservative and do certainly not reflect the fact that flawed girth welds, made following standard procedures, fail by plastic collapse. Flaw characterisation rules are based on elastic fracture behaviour. Beyond this limitation, the current interaction and recategorisation procedures do not take account of the favourable effects of weld metal yield strength overmatching and weld reinforcement on weld performance.

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Fig. A15 – Sample calculations illustrating the effects of Y/T ratio and flaw sizing error on allowable flaw height (performance requirement: GSY).

Flaw interaction - Existing interaction rules for multiple flaws are highly conservative for girth welds subjected to plastic strains [71]. An associated issue is that UT inspection might find a larger number of planar ’neighbouring’ flaws. On the other hand, since UT is ‘subject’ to inaccuracies in flaw sizing, the application of current interaction rules may cause higher rejection and repair rates. Note that the same comment applies for ‘workmanship based’ flaw accumulation criteria. Flaw recategorisation - Present flaw recategorisation procedures reclassify an embedded near-surface breaking flaw into a surface breaking or through-thickness flaw if the top and/or bottom ligaments fail the ‘theoretical’ local plastic collapse check. By using this (standard) procedure, an innocent near-surface embedded flaw might be redefined as a potential

Pipe Grade X70 - Y/T = 0,80 Pipe: 40" x 20 mm - Requirement: Pipe Yielding (GSY)

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Pipe Grade X70 - Y/T = 0,90 Pipe: 40" x 20 mm - Requirement: Pipe Yielding (GSY)

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‘harmful’ surface-breaking flaw. In other words, current practices lead to a significant number of unnecessary weld repairs, considering the fact that tearing only occurs after extensive yielding. As current flaw characterisation practices lead to a significant number of weld repairs that are, considering the ad-hoc testing experience, not needed, it is recommended to investigate these problems.

6.2.4 Practical ad-hoc solution

It should be underlined that a probabilistic approach is only one amongst several solutions to established reliable allowable flaw sizes. For example, the interaction between the uncertainties related to flaw height sizing and acceptable flaw size can directly be simulated by Curved Wide Plate (CWP) testing. This is a very attractive option for the study of girth welds in pipelines for which little previous experience exists. By testing CWP specimens containing a surface breaking flaw with a height equal to its assumed value (normally 3 mm) plus a 2,0 mm allowance for error in height sizing, the actual failure characteristics can be determined. Although the option of CWP testing has a number of advantages, it must be emphasized that CWP testing is only an intermediate step in the development of a generally accepted ECA solution consistent with the capabilities and limitations of the inspection technique.

7 RECOMMENDATIONS Throughout this review, reference has been made to various aspects of an ECA that are not completely understood. However, the current state of technology is sufficiently advanced to make a major step in the development of reliable allowable sizes for flawed girth welds. This possibility might be of concern or beneficial depending on the viewpoint taken. Emphasis for the foreseeable future should be placed on steady improvements in the ECA approaches by introducing issues not currently addressed. Some of the more important issues are listed below. The non-exhaustive list of recommendations is presented with the expectation that others will have opinions to offer.

7.1 Recommended steps to harness existing knowledge to improve ECA practices 7.1.1 Material properties

��Specify materials, welding processes and procedures ensuring ductile failure behaviour. ��Avoid materials susceptible to scatter in properties. In particular, limit the upper supply

range of pipe yield strength in the axial direction.

��Ensure weld metal yield strength match/overmatch, if possible (the welding process used might be the limiting factor).

��Rationalize the toughness testing requirements (Charpy or CTOD testing?).

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��Establish correlations between specification requirements and the actual toughness and

tensile properties. Lack of reliable material data hampers the efficient use of reliability models to optimise inspection and design.

7.1.2 ECA analysis

��Use an experimentally proven ECA methodology specific to the application in mind. Explore the material properties available, if needed.

��Determine and document the effects of the input parameters and their relative effects on

the calculated defect size limit(s). Incorporate this information into the material and welding specifications.

��Implement the extension of the EPRG-Tier 2 guidelines on allowable flaw sizes.

��Teach engineers how to avoid potentially unsafe ECA assumptions which actually may

reduce girth-weld integrity. This recommendation reflects the fact that many users are not familiar with all details involved in an ECA analysis.

��For critical applications, plastic designs and for situations involving high-strength pipe

establish the critical defect size limits by type (CWP or full-sale bend) testing. 7.1.3 Testing and inspection

��Revise the testing requirements. In particular, design and specify material testing requirements that provide data relevant to the ERPG-Tier 2 approach. This action must be regarded as a prime need in developing a cost-effective ECA methodology.

��Determine the weld metal's yield strength properties and their variability.

��Reduce human and procedural errors by applying simple and transparent testing and

inspection procedures.

��Determine what defect size limit(s) the inspection should be designed to find.

��Tailor the inspection procedures to the location of the defect (surface-breaking vs. embedded).

7.2 Recommendations for future research to improve ECA methodologies 7.2.1 Material properties

��Collect appropriate weld and pipe metal yield and tensile strength distributions to: - Determine the level of weld-metal vs. pipe-metal yield strength mismatch as a

function of pipe grade and welding process/procedure

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- Optimise the reliability models for future use in defect assessments of ‘old’ pipelines

��Develop material and welding specifications for strain-based designs.

7.2.2 ECA analysis

��Develop a multi-level ECA methodology that connects both the pipe-to-weld metal mismatch effects and the inspection capabilities. In particular, create "tailor-made" tolerable defect length-height curves as a function of the performance requirement (stress-based vs. strain-based), weld properties, defect type, well thickness and inspection capabilities.

��Determine the arc length for plastic collapse of small diameter pipelines ��For (very) high strength pipes, quantify the interaction between yield strength, Y/T ratio

and uniform elongation of the pipe metal because the quantification of the deformation and failure behaviour of a high strength steel pipe in terms of Y/T ratio alone may be an over simplification.

��Develop workmanship requirements for AUT inspected girth welds and strain-based

designs. These requirements should be based on simple and transparent fitness-for-purpose principles.

��Develop defect interaction rules for ductile failure behaviour.

��Revise the defect characterization rules for near-surface-breaking defects.

7.2.2 Testing and inspection

��Experimentally establish the minimum Charpy V requirements ensuring plastic collapse of "defective" girth welds in heavy wall pipe (t > 25.4 mm).

��Develop a simple and representative small-scale test method that simulates the constraint

conditions occurring at defects in pipeline girth welds. If possible, the test should allow one to determine the extent of crack-tip ductile tearing in terms of applied load and yield strength/tensile strength ratio.

��Experimentally verify the limit levels of weld-metal yield strength undermatch and

overmatch for stress-based and strain-based designs.

��Establish a database that provides detailed information on the type and the shape of weld defects occurring in field welds. This information would allow engineers to:

- Place the relation between defect profile and defect sizing accuracy in a proper perspective

- Select the most relevant and economic non-destructive testing technique

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- Avoid the overselling of technical non-destructive testing capabilities.

��Develop non-destructive testing requirements that take into account the service requirements (stress-based or strain-based design), defect location (surface breaking vs. embedded) and material toughness properties (undermatching vs. overmatching weld metal strength to pipe strength).

REFERENCES 1. BSI-PD 6493-1991, "Guidance on Some Methods for the Derivation of Acceptance Levels for

Defects in Fusion Welded Joints," British Standards Institution, London, 1991. 2. BS 7910-2000 Guide on methods for assessing the acceptability of flaws in metallic structures,

London, BSI, 2000. 3. API Standard 1104, "Welding of Pipelines and Related Facilities," 17th Edition, September 1999. 4. Canadian Standards Association, CSA-Z662-99, "Oil and Gas Pipeline Systems," Ontario,

December 1996.

5. Australian Standard, AS 2885.2-1995, "Pipelines - Gas and Liquid Petroleum, Part 2: Welding," 1995.

6. API RP 579-2000, “Recommended practice for fitness-for-service”. Washington, API, 2000 7. CEN 12732-2000, “Gas Supply Systems – Welding Steel Pipework – Functional Requirements” 8. Knauf, G. and Hopkins, P., "The EPRG Guidelines on the Assessment of Defects in Transmission

Pipeline Girth Welds," 3 R International, 35 Jahrgang, Heft 10-11/1996, S. 620-624. 9. SINTAP: Structural Assessment Procedures for European Industry. Final Report, July1999,

Contract No BRPR-CT95-0024, Project No. BE95-1426, Report No. BE95-1426/FR/7. 10. Hopkins, P., Demofonti, G., Knauf. G., and Denys, R., “An Experimental Appraisal of the

Significance of Defects in Pipeline Girth Welds”, Paper 20, Proc. EPRG/NG18 Eight Biennal Joint Technical Meeting on Line Pipe Research, Paris, May 14 -17, 1991

11. Denys, R.M., "Fracture Behaviour of Large Diameter Pipeline Girth Welds: Effect of Weld Metal

Yield Strength in 25 mm Thick Pipe", AGA Catalog No 202-011, Vol. I, Febr, 9-94 12. Denys, R.M., "Effect of Weld Metal Matching on Girth Weld Performance, Vol. II," AGA Catalog

No. L51 851-IN1, p. 137, January 1993. 13. Denys, R., 'Fracture Behaviour of Large Diameter Pipeline Girth Welds, Effect of Weld Metal Yield

Strength and Defect Interaction', AGA Report, PRC-Line Pipe Welding Supervisory Committee, Contract Number PR 202-011, May, 24, 1994

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14. Denys, R.M., Lefevre, A.A., Glover, A..A. and Horseley, D.J., "An Approach to Defect Acceptance Criteria for GMAW Girth Welds in High Pressure Pipelines," Conf. Proceedings PRCI 9th Symposium on Pipeline Research, Houston, Texas, 1996,

15. Denys, R.M., Lefevre, A.A., "PRCI report PR 202-9328: Alternative Acceptance Criteria of Girth

Weld Defects in Cross Country Pipelines," American Gas Association, Arlington, Virginia, June 1996.

16. Demofonti G, Fersini M, Knauf G and Schipaanboord W: Defects in girth welds. New test results,

NDT experience, EPRG Guidelines. PRC/EPRG Eleventh Biennial Joint Technical Meeting, Arlington, Virginia, April 7-11, 1997. Paper 17.

17. Schipaanboord, W., Denys, R., Lefevre, A., and Roovers., P “Validation tests to Support the EPRG

– Tier 2 Guidelines for Misaligned gGrth Welds in Large Diameter Onshore Pipes through Full Scale Bend and Wide Plate Tensile Testing”, Conference Proceedings 12th Biennal Joint Technical Meeting PRC/EPRG on Pipe line Research, Groningen, Mei 17-21, 1999

18. Denys RM, Lefevre AA, De Jaeger C and Claessens S: Failure characteristics and defect tolerance

levels of girth welds in large diameter X65/X70 steel pipelines: experimental verification through wide plate testing and comparison with ECA prediction models. In: Denys RM, editor. Pipeline Technology Proceedings of the third international pipeline technology conference Volume I, Brugge, Belgium, May 21-24 2000. Amsterdam: Elsevier Science, 2000, pp.151-67.

19. Kastner, W., Roehrich E, Schmitt W and Steinbuch R, “Critical Crack Sizes in Ductile Piping”,

International Journal of Pressure Vessels and Piping 9 (1981), pp 197-219. 20. Willoughby, A.A., “A Survey of Plastic Collapse Solutions Used In Failure Assessment”, Welding

Institute Report 191/1982, 1982. 21. Milne I, Ainsworth RA, Dowling AR and Stewart AT: Assessment of the Integrity Of Structures

Containing Defects”,CEGB Report R/H/R6 Rev 3, 1986. 22. Miller, A. G., "Review of Limit Codes of Structure Containing Defects," International Journal of

Pressure Vessels and Piping, Vol. 32, 1988, pp.191-327. 23. Wang Y-Y, Wilkowski GM and Horsley DJ: Plastic collapse analysis of pipelines containing surface

breaking circumferential defects, Denys R, editor. Pipeline Technology Proceedings of the third international pipeline technology conference, Volume I, Brugge, Belgium, May 21-24 2000. Amsterdam: Elsevier Science, 2000, pp.191-209.

24. Denys, R., A Plastic Collapse Based Procedure for Girth Weld Defect Acceptance, Proc. Int. Conf.

on Pipeline Reliability, Canmet, Pergamon, Calgary Canada, Pergamon Press, 1992, Vol. II, paper VIII.1-VIII.11

25. Coote, R. I., Glover, A.G., Pick, R.J. and Burns, D.J., “Alternative Girth Weld Acceptance

Standards in the Canadian Gas Pipeline Code”, “Investigation of Plastic Collapse Criteria for Defects in Line Pipe Girth Welds During Bending” , Proc. 3rd. Int. Conf. on Welding and Performance of Pipeline, T.W.I., London, Nov. 1986, paper 21.

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26. Hopkins, P., and Denys, R.M., "The Background to the European Pipeline Research Group's Girth Weld Limits for Transmission Pipelines," EPRG/NG-18, 9th Biennial Joint Technical Meeting on Line Pipe Research, Houston, Texas, May 11-14, 1993, Paper 33.

27. Denys, R.M., "Toughness Requirements for Pipeline Integrity," Proc. of 13th OMAE Conf., Vol.

III, Part A, Copenhagen, Denmark, June 18-24, 1995. 28. Dolby, R.E., “Charpy V and COD – Correlations Between Test Data for Ferritic Weld Metal”, Metal

Construction, Vol. 13, No. 1, Januari 1981, pp. 43-51. 29. Denys, R., “Lowerbound Charpy V Notch Toughness for Plastic Collapse”, To be published. 30. NN. Gegevens bestanden van Laboratorium Soete m.b.t. kleine en grootschalige proeven op

rondnaden in pijpleidingen, April 2001. 31. NN, Method of assessment for defects in fusion welded joints with respect to brittle fracture,

Japan Welding Engineering Society Standards WES 2805-1980, The Japan Welding Engineering Society, Tokyo (1980)

32. Hopkins, P., Pistone, V.and Clyne, A., “A Study of the Behaviour of Defects in Pipeline Girth

Welds, The Work for EPRG”, Conf. Proc. Int. Conf. on Pipeline Reliability, Canmet, Pergamon, Calgary Canada, Pergamon Press, 1992,

33. Hopkins, P., “The Application of Fitness-for-Purpose Methods to Defects Detected in Offshore

Transmission Pipelines”, Proc. Conf. On Welding and Weld Performance in the Process Industry, IBC, London, 27-28 April, 1992.

34. Denys RM, Lefevre AA, De Baets P and Degrieck J: “Effects of stable ductile crack growth on

plastic collapse defect assessments”. Pipeline Technology Proceedings of the third international pipeline technology conference Volume I, Brugge, Belgium, May 21-24 2000. Amsterdam: Elsevier Science, 2000, 69-89.

35. Denys, R., 'Girth Welds In Modern Pipelines, Quality Control and Safety Aspects', Pipeline

Pigging and Integrity Monitoring Conference, Amsterdam, 11-14 april, 1994. 36. Clyne, A. J. and Jones, D. G., "The Integrity of Transmission Pipelines Containing Circumferential

Girth Weld Defects," Pipeline Technology, Eds. R. Denys, Vol. II, 1995, pp. 299-314. 37. Worswick M.J., Coote R.I., and Burns D.J., “Investigation of Plastic Collapse Criteria for Defects

in Line Pipe Girth Welds During Bending” , Proc. 3rd. Int. Conf. on Welding and Performance of Pipeline, T.W.I., London, Nov. 1986, paper 30.

38. Woodley, C.C., Burdekin, F.M. and Wells, A.A., “Mild Steel for Pressure Equipment at Sub-Zero

Temperatures,” Brit. Weld J. 1964, 11, 3, pp.123-136, 1964. 39. Soete, W., and Denys, R.M., “Fracture Toughness Testing of Welds”, Proc. of Conf. on Welding of

HSLA (Micro-alloyed) Structural Steels, Roma, Nov. 1976, ASM, ISBN 087170 0050, pp.63-84

Universiteit Gent


40. Denys, R.M., “Wide Plate Testing of Weldments, Parts I, II and III”, “Fatigue and Fracture Testing of Weldments”, ASTM STP 1058, Eds. H. McHenry and J. Potter, ASTM, Philadelphia, 1990, pp.157-228.

41. Denys R.M., “Toughness Requirements In Transversely Load Welded Joints - An Evaluation

Based On Wide Plate Testing”, The Fracture Mechanics of Welds, EGF Pub. 2 (Edited by J.G. Blauel and K.H. Schwalbe, 1987), Mechanical Engineering Publications, London, pp. 155-189.

42. Mayville, R.A. and Hilton, P. D., “A Study of Plasticity Effecst on the Crack Geometry Correction

Factors Used with the COD-Design Curve”, Fracture Mechanics, 18th Symposium, ASTM STP 945, 1988, pp. 686-698.

43. Roodbergen, A.H., and Denys, R., Limitations of Fitness for purpose Assessments of Pipeline

Girth Welds, Proc. Int. Conf. on Pipeline Technolgy, AIM and Commission of European Communities, Rome, Nov. 1987.

44. Denys, R.M., "Effect of Weld Metal Matching on Girth Weld Performance, Vol. I," AGA Catalog No.

L51 651-IN1, p. 96, February 1992. 45. Scott, P.M., “Interpretative Study of Published and Recent Research on the applicability and

limitations of Current Fracture Prediction Models for Girth Welds,” Canadian Metallurgical Quarterly, Vol. 32, No. 3, pp. 223-237, 1993

46. Wilkowski GM, Olson RJ and Scott PM: State of the art report on piping fracture mechanics

NUREG/CR-6540. Columbus, Ohio, Battelle, 1998. 47. Dawes, M.G., 1985, The CTOD Design Curve Approach: Limitations and Finite Size and

Application, Welding Institute Report 278/1985. 48. Andrews, R.M., Denys, R.M., Muesch. K, “Recent studies to Enlarge the Limits of the EPRG

Guidelines on the Assessment of Defects in Transmission Pipeline Girth Welds”. PRC/EPRG Tenth Biennial Joint Technical Meeting, New Orleans, Louisiana, USA, April 30-May 4,, 2001. Paper 3

49. Denys, R., Difference Between Small and Large Scale Testing of Weldments, An Evaluation of the

Weld Metal Overmatching Effects in Relation to Weld Joint Performance”, The Welding Journal, Research Supplement, N2, Febr. 89, p 33s-43s.

50. Lian, B., Denys, R and L. Van de walle, An experimental assessment on the effect of weld metal

yield strength overmatching in pipeline girth welds , Proc. of the Third Conference on Welding and Performance of Pipeline, TWI, London, November 1986.

51. Denys, R., Lefevre, A.A., Martin, J.T., Glover, A.G., Yield Strength Mis-Match Effects on Fitness

for Purpose Assessments of Girth Welds, PRC/EPRG Tenth Biennial Joint Technical Meeting on Line Pipe Research, American Gas Association, Paper 3, Cambridge, UK, April 18-21, 1995, pp. 1-15

52. Denys R., “Methods for the Assessment of Girth Weld Strength”, Welding Technology Institute of

Australia, Proc. Conference on Welding of High Strength Thin-Walled Pipelines”, Wollongong, October 26, 1995, Paper 7, 1-15.

Universiteit Gent


53. Denys, R., and Lefevre, A.A., "Weld Metal Mismatch, Challenges and Opportunities” Proc. ICAWT

’99 - Pipeline Welding and Technology Conference, Galveston, TX, USA, EWI-AWS, October 26-26, 1999.

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Technology Conference, Oostende, Oct. 15-18, 1990. 56. Denys, R., “Is the Transverse Weld Tensile Test a Reliable Test?”, “Proceedings Pipeline

Technology Conference”, September 11 - 14, 1995, ISBN Number: 0 444 82197, Vol II - pp. 581 – 591

57. Denys R., “How Much Weld Metal Yield Strength Do We Need?”, Second Pipeline Technology

Conference Ostend, September 11-14, 1995, Vol II - pp. 555 - 560. 58. Cousin V., and Baguley R., “Enhancing Pipeline Reliability: A Constractors Perspective”, Canmet,

Int. Conf. on Pipeline Reliability, June 2 5, 1992, Calgary, Canada, Paper II-9. 59. Denys, R.M., Lefevre A.A., A.G. Glover, "Weld Metal Yield Strength Variability in Pipeline Girth

Welds", Proc. of the Second International Conference on Pipeline Technology, Elsevier Publishers, Ostend, 11-14 september 1995, p.591-598.

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Pipelines”, Second Symposium on Mis-Matching of Interfaces and Welds”, Proceeding GKSS Geesthacht, April 24-26,1996, pp. 547-560.

61. Bannister, A.C. and Harrison, P.L., “A Comparative Study of Wide Plate Behaviour of a Range of

Structural Steels Using the Failure Assessment Diagram”, OMAE, 1995, Vol III, Materials Engineering, pp. 65-75.

62. C. Dallam, M. Quintana, “Welding X-80 And Beyond: A Consumable Manufacturers Perspective”,

Pipeline Technology Proceedings of the third international pipeline technology conference Volume II, Brugge, Belgium, May 21-24 2000. Amsterdam: Elsevier Science, 2000, pp.483-498.

63. D.J. Widgery, “Welding High Strength Steel Pipelines - A Consumable Manufacturer's View”,

Pipeline Technology Proceedings of the third international pipeline technology conference Volume II, Brugge, Belgium, May 21-24 2000. Amsterdam: Elsevier Science, 2000, pp.499-508

64. E. Perteneder, H. Königshofer, R. Bischof, “Capabilities And Limitations Of Modern Welding

Consumables Suitable For Girth Welding Of Pipelines - A Producer's Perspective”, Pipeline Technology Proceedings of the third international pipeline technology conference Volume II, Brugge, Belgium, May 21-24 2000. Amsterdam: Elsevier Science, 2000, pp.509-526

65. Blackman SA and Dorling DV: Advanced welding processes for transmission pipelines. In: Denys

RM, editor. Pipeline Technology Proceedings of the third international pipeline technology

Universiteit Gent


conference Volume II, Brugge, Belgium, May 21-24 2000, Amsterdam: Elsevier Science, 2000, 371-387.

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on Mechanized Girth Welds, Proc. Pipeline Technology Conference, Oostende, Oct. 15-18, Paper 14.17, 1990

67. Van Merrienboer H.A.M, Mechanised Ultrasonic Inpsection of Manually Welded Girth Welds

Compared with Radiography, Canmet, Int. Conf. on Pipeline Reliability, June 2-5, 1992, Calgary, Canada, Paper IV14.

68. Coors P., Denys R., ‘Effect of Automated Ultrasonic Inspection Capabilities on the EPRG

Guidelines’, Paper 16, PRCI/EPRG 12th Biennial Joint Technical Meeting on Line Pipe Research, 1999, Groningen.

69. F.H. Dijkstra, J. van der Ent, T. Bouma, “Defect Sizing and ECA Criteria: State of the Art in AUT”,

Pipeline Technology Proceedings of the third international pipeline technology conference Volume I, Brugge, Belgium, May 21-24 2000. Amsterdam: Elsevier Science, 2000, 69-89, pp.483-498

70. Denys, R.M., Lefevre, A.A., De Jaegher, C. and Claessens, S., “Weld Acceptance Criteria”, Final

Report of Study Commissioned by the International Pipe Line & Offshore Contractors Association (IPLOCA), 2000, Gent, Belgium

71. Denys, R., Lefevre. A.A. and De Jaeger, C., “Defect Interaction for Ductile Material Behaviour”,

OMAE98-2058, ”Proceedings of 17th OMAE Conference, Materials Sessions, CD-rom, p. 8, July 5-9,1998, Lisbon, ISBN No 0-7918-1952-3.

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