Rothwell and Coote Ostend 2009 paperFINAL

A CRITICAL REVIEW OF ASSESSMENT METHODS FOR AXIAL PLANAR SURFACE FLAWS IN PIPE A.B. Rothwell1, R.I. Coote2 1

Brian Rothwell Consulting Inc., 100 Hamptons Link N.W., Calgary, Canada brian.rothwell@shaw.ca 2 Coote Engineering Limited, 16 Varbow Place N.W., Calgary, Canada cooteb@shaw.ca Abstract Models for determining failure conditions for planar surface defects, such as weld seam cracks and cracks caused by SCC, are necessary for various purposes, most importantly the development of assessment criteria for sentencing flaw indications from in-line and “in-ditch� inspection. The consequences of inaccuracy can be serious, including operational failures or excessive expenditures on excavations and repairs. Numerous methods are available and in use today for the assessment of cracks. This paper reviews the historical background to the earliest approaches to this problem, and then considers a number of more recent methods. In each case, the performance of the methods, in terms of observed against predicted failure stress, is assessed against a database of laboratory and field failures. It is concluded that all of the models considered can perform very well against results from well-characterized flaws of regular profile, but that performance against flaws of irregular profile is much worse and more variable. Methods that can accommodate actual flaw shape and consider fracture controlled and plastic collapse failure modes concurrently perform better overall. The practical implications of the observations for different integrity-related applications are briefly discussed.

INTRODUCTION Since the 1960s, methods have been available for the assessment of planar flaws in pressurized cylinders. The earliest approach developed specifically for pipelines, due to Kiefner [1], was based on flow-stress dependent failure, analogous to the methods used for assessing volumetric flaws such as corrosion. This method is still in use today, for pipe materials that show reasonably good fracture resistance. A modification, aiming to account for toughness-dependent failure, was developed in the early 1970s, and is also still in use today. More recently, several additional methods of analysis have been developed, making use of advances in the understanding of fracture mechanics and of material behaviour that have taken place over the intervening period. These more recent models are of varying degrees of transparency; the data on which they were validated are not always readily available, nor are the statistical properties of the models easy to determine, so that explicit levels of conservatism are hard to achieve. The consequences of insufficient conservatism are obvious and can be very serious; the economic consequences of excessive conservatism can also be high, including excessive numbers of excavations following ILI and excessive numbers of repairs of flaws evaluated on excavation.

Indeterminate levels of model conservatism make it difficult to make rational decisions regarding the integrity-related maintenance of pipelines subject to time-dependent cracking processes such as fatigue and SCC. As a result, there is a strong motivation to gain a better understanding of the characteristics and validation details of the models that are used for this purpose.

1

This paper briefly reviews the background of the Battelle surface flaw equations, developed for the Pipeline Research Committee of AGA in the 1960s and 1970s, and their original validation against failure stresses determined on sharp-notched pipe vessel tests. It then considers three more recent approaches (PAFFC, CorLASTM , and an FAD-based method according to assessment level 2A of BS 7910) that are also relatively-widely used in the pipeline industry. The performance of these models is first assessed against the same data used to validate the Battelle equations. The performance of all four models is then assessed against some more recent field and hydrostatic test failure results involving cracks arising from SCC and other mechanisms. Some conclusions are drawn relative to the factors that influence the effectiveness of the different approaches.

BACKGROUND AND VALIDATION OF THE BATTELLE SURFACE FLAW EQUATIONS The Battelle surface flaw equations were developed explicitly for the pipeline industry in the 1960s and early 1970s [1,2]. The first version to be developed was the flow-stress-dependent (FSD) form, which is based on the assumption that failure occurs when the stress in the remaining ligament, multiplied by a stress concentration factor related to the bulging caused by the presence of the flaw, reaches the flow stress, conventionally defined as yield strength + 69 MPa. This failure condition can be expressed as1

1 Aeff / A0 F

1 Aeff / M T A0

(1)

where F

and

are the failure stress and the flow stress, respectively;

Aeff is the effective area of the flaw;

A0 is the reference area, equal to L times t ; L is the surface length of the flaw; t is the wall thickness; and M T is the Folias factor for a through-wall flaw of the same length, equal to

[1 + 0.6275L2/(Dt) – 0.003375L4/(Dt)2]0.5 for L2/(Dt) ≤ 50 and [0.032L2/(Dt)] + 3.3 for L2/(Dt) > 50.

For a rectangular flaw of depth d , equation (1) can be written as

1 d /t 1 d / MTt

F

(2)

This is probably the most familiar form of the Battelle surface flaw equation. It can also be applied to flaws of non-rectangular geometries, by substituting an equivalent flaw length

Leq

Aeff / d

for L in the Folias factor.

Equation (2) was validated against a data set of 47 pipe vessel tests containing sharp axial surface notches [2]. It is worth showing a graphical representation of the results (Figure 1), in order to understand the predictive capability of the equation and the influence of material toughness

1

This and the following equations are well-known, but are repeated here for convenience.

2

properties. For all calculations, an equivalent flaw length was used that accounts for the radius at the ends of the machined flaws.

600 y = 1.0267x R² = 0.931

All points

Observed failure stress, MPa

500

No low or unknown toughness y = 0.965x R² = 0.835

400

300

200

100

0 0

100

200

300

400

500

600

Calculated failure stress, MPa

Figure 1 Observed failure stress against failure stress calculated using Equation (2) When all 47 points are included, it can be seen that the mean relationship is somewhat nonconservative, and there are a few outliers for which the observed failure stress is significantly less than the calculated value. When the results of tests for which the material Charpy shelf energy (full size equivalent) was less than 40 J, or unknown, are eliminated, the mean relationship becomes very slightly conservative2, and the quality of the correlation is greatly improved.

In order to extend the applicability of this method to lower-toughness materials, Battelle researchers modified Equation (2), essentially by analogy with the form of the equation that had been developed earlier for through-wall flaws [2]. This is most conveniently expressed in terms of a “toughness factor”, a multiplier applied to the right-hand side of equation (2), that can be written as

2

arccos exp

1000 Cv E 4L

2

where Cv is the Charpy shelf energy per unit area in J/mm2 and

and E are in MPa.

Figure 2 shows the relationship between observed failure stress and values calculated using the toughness dependent (TD) equation, for all of the previously-mentioned pipe vessel tests for which the Charpy shelf energy was known (35 tests). By comparison with Figure 1, it can be seen that there are now no non-conservative low-toughness outliers, and the mean correlation is about 9% conservative. There is one conservative low-toughness outlier; however, when all results for Charpy shelf energies below 40 J (full size equivalent) are removed, the mean correlation changes only slightly, though the correlation coefficient increases somewhat. This indicates that, at least for the 2

Here and in subsequent sections, the term “conservative” has been used to describe observed values higher than calculated values. Clearly, there are some applications (such as estimating the size of the largest flaws remaining after a hydrostatic test) when such a situation would be non-conservative. In the application of any of the models discussed, a safety factor will need to be applied that takes into account model error in the appropriate way.

3

range of pipe and flaw geometries and toughness values examined in these tests, the effect of toughness on failure stress is accounted for consistently by the factor shown above3.

It is also clear that, at high levels of toughness, the exponential will tend to 0, so that the toughnessdependent form of the equation approaches equation (2). In fact, as Figure 3 shows, for many of the vessel tests evaluated, the values calculated by the two methods were within a few percent of each other.

600 All points

y = 1.0904x R² = 0.9085

No low toughness points

Observed failure stress, MPa

500 y = 1.0814x R² = 0.9367

400

300

200

100

0 0

100

200

300

400

500

600

Calculated failure stress, MPa

Figure 2 Observed failure stress against failure stress calculated using the toughness-dependent surface flaw equation 600 All points

Calculated TD failure stress, MPa

No low toughness

500 y = 0.9481x R² = 0.9696

400

300

200

100

0 0

100

200

300

400

500

600

Calculated FSD failure stress, MPa

Figure 3 Calculated toughness-dependent failure stress against calculated flow stress-dependent failure stress 3

It has been noted that this factor, because it contains the flaw length but no representation of flaw depth, can become unduly conservative for long, shallow flaws in relatively tough materials.

4

Overall, these data show that, if the Charpy shelf energy is known, the failure stress for the flaw and pipe geometries and materials evaluated could be predicted quite accurately, and usually slightly conservatively, using the TD surface flaw equation. If the toughness exceeded 40 J full-size equivalent, the FSD equation (2) gave predictions that were nearly as good, though the probability of individual non-conservative predictions increased somewhat. In general, for regular, wellcharacterized flaw shapes, the Battelle surface flaw equations perform well, and this method has the advantage of being very transparent and easy to use.

More recently, Kiefner and Associates has developed software known as KAPA [3] that calculates the failure pressure for blunt metal loss defects using the Battelle FSD model and the failure pressure for crack-like defects using the Battelle TD model. KAPA includes iterative effective flaw size calculations, based on actual depth/length measurements, to find the lowest predicted failure pressure (analogous to the procedure in RSTRENG), and thus can be applied to irregular flaws when a detailed profile is available (see below).

VALIDATION OF ALTERNATIVE ASSESSMENT METHODS AGAINST BATTELLE VESSEL TESTS A number of alternative methods for the assessment of axial, planar surface flaws in pipelines have been developed since the Battelle surface flaw equations. In general, they have aimed to take advantage of the improved understanding of fracture mechanics and material behaviour that has been gained since the 1970s to provide more flexible and robust analysis tools. It seems worthwhile to evaluate the performance of these methods against the original Battelle database, since this should provide a useful test of their intrinsic accuracy. Of course, more sophisticated methods generally imply more sophisticated inputs; in particular, material toughness in fracture mechanics terms is generally unavailable, and must be estimated using correlations that may be more or less materialspecific. The error associated with these correlations will contribute to the overall prediction error.

Pipe Axial Flaw Failure Criterion (PAFFC) This method was developed in the 1990s, with the aim of providing a more flexible and accurate approach that would take advantage of advances in the understanding of fracture behaviour that had taken place since the 1970s. The approach incorporates a ductile flaw growth model, developed by Battelle, that can accommodate stable tearing prior to fracture, which can have a substantial influence on the failure behaviour of high-toughness materials under sustained loading [3]. It considers failure due to fracture (using a J-R based approach) and to net section collapse concurrently, failure occurring when one of the limiting conditions is breached. J-R test data will rarely be available to the analyst, so that, in practice, JIc and dJ/da must usually be obtained from correlations with Charpy and tensile test data. The models have been incorporated into a software package (PAFFC) [4], which also contains several alternative built-in J-R correlations, based on mechanical and fracture toughness tests on materials considered representative of different ranges of strength grade and manufacturing era.

In evaluating PAFFC against the original 33 Battelle pipe vessel tests for which the Charpy shelf energy is known4, the equivalent flaw length was used, and J-R correlations considered appropriate for material grades X52 to X65 produced prior to 1975 were applied. As figure 4 shows, the observed results correlate well with the calculated values, indicating that the method is accurate and the correlations are appropriate over the entire range of material properties and flaw sizes included in the Battelle study. The correlation is also very similar to that in Reference [4], though the validation database is not identical (there may be some common test results). The mean of the correlation is 4

Since tensile strength is a required input for PAFFC, two tests for which it was not reported were not analysed.

5

slightly non-conservative, but this is not a concern for application, since it can be corrected where required by the incorporation of a corresponding multiplicative model error term.

Observerd failure stress, MPa

600 y = 0.944x R² = 0.9264

500 400 300 200

100 0

0

100

200

300

400

500

600

Calculated failure stress, MPa

Figure 4 Observed failure stress against failure stress calculated using PAFFC

CorLAS™ As usually applied, this method is also software-based, and addresses fracture and flow-stressdependent failure concurrently. CorLAS™ 2.2 beta version was used to calculate fracture and flowstress dependent failure pressures. This software version is basically the same as the version 2 software [6-9], with the additional capability to perform batch analysis to calculate the fracture and flow stress failure pressures for numerous flaws based on their depth and length. The FSD criterion is equivalent to that of the Battelle surface flaw failure criterion in Equation (1), but includes iterative effective flaw size calculations, based on actual depth/length measurements, to find the lowest predicted failure pressure (analogous to the procedure in RSTRENG). The fracture-controlled formulation is based on a Jc criterion, for the application of which flaws are evaluated as the equivalent semi-ellipse corresponding to the worst case determined by the effective-flaw analysis.

As with PAFFC, there are generally difficulties related to the availability of the fracture input data, and correlations with more usually-available test data must be used. In evaluating CorLAS™ against the Battelle pipe vessel tests, the correlation used for J was Jc = 1000Cv/ 80 where Jc is in N/mm and Cv is the Charpy shelf energy (full size equivalent) in J. It is noted that the correlation for Jc is equivalent to the Kc correlation used by Battelle in developing the toughness-dependent equation. The strain-hardening exponent used was based on a relationship developed for API pipeline steels [8].

The results of the analyses are shown in Figure 5. The overall quality of the correlation is good, and the method again appears to be applicable over the entire range of variables covered by these tests.

The mean is very slightly non-conservative, though there are one or two significantly nonconservative outliers.

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Figure 5 Observed failure stress against failure stress calculated using CorLAS™

Failure assessment diagram (FAD) approaches - BS 7910 Level 2A [10] Whereas, in CorLAS™ and PAFFC, fracture-driven and strength/plasticity-driven failure modes are considered “concurrently but independently”, FAD approaches combine fracture-based and loadbased criteria in a two-dimensional way. A load ratio Lr (applied reference stress/yield strength) and fracture ratio Kr (applied stress intensity factor/material toughness) are calculated for the flaw being assessed, and plotted on the assessment diagram. An assessment curve (the envelope of acceptable fracture ratios and load ratios) is also plotted (see Figure 6); on this curve, the upper limit of Kr is 1.0, at Lr = 0; there is a load-ratio cut-off at a load ratio corresponding to the mean of yield and tensile strengths, which is thus material dependent. It is important to note that the purpose of this approach, unlike those discussed previously, is to determine whether a flaw is acceptable under given combinations of load ratio and fracture ratio, not to determine failure conditions. As a result, there is intended to be a significant (and not precisely defined) level of conservatism inherent in the process.

As far as the load ratio is concerned, an explicit factor of 1.2 is applied in the calculation of the reference stress.

For the current analysis, the load ratio and fracture ratio were calculated according to the applicable equations in BS7910. Kmat was estimated from the Charpy shelf energy using the same equivalence as is embedded in the Battelle TD surface flaw equation Kmat = [1000CvE]0.5 This simple equation (and its equivalent in terms of Jc) proved to be quite acceptable in the application of the Battelle equation and Corlas™ to the Battelle pipe vessel tests, and thus its use, in place of the more general formulations provided in BS 7910, is considered appropriate. The value of flaw length used to calculate the stress intensity factors was the length of a semi-elliptical flaw with the same area and maximum depth as the actual flaw, because the expressions in BS 7910 are based on a semielliptical flaw shape. However, the equivalent flaw length was used to calculate the Folias bulging factor using the expressions described for equation (1).

The applied hoop stress corresponding to an assessment point lying on the assessment curve was determined iteratively for each set of test conditions (pipe and flaw dimensions, pipe material properties). In addition, the calculations were repeated without the factor of 1.2 applied in the calculation of the reference stress.

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1.00 0.90

Unacceptable

0.80

0.70

Kr

0.60

Acceptable

0.50 0.40

0.30 0.20 0.10

0.00 0.00

0.20

0.40

0.60

0.80

1.00

1.20

1.40

Lr Figure 6 FAD according to BS 7910 level 2A The results of these calculations are shown in Figure 7, in terms of observed failure stress against calculated stress, with and without the factor of 1.2. It can be seen that the correlations in both cases are quite good. When the factor on stress is included, the mean relationship is conservative by about 20%, though, perhaps surprisingly, there are three points that fall below the 1:1 relationship. Without the factor, the mean relationship is close to 1:1.

500

y = 1.2053x R² = 0.8769

Observed failure stress, MPa

400 y = 1.0231x R² = 0.8831

300

200 FAD FAD no factor

100

0 0

100

200

300

400

500

Calculated failure stress, MPa

Figure 7 Observed failure stress against failure stress calculated using the FAD approach in BS7910, with and without using the factor of 1.2 in calculating Lr

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APPLICATION OF THE ASSESSMENT MODELS TO REAL FLAWS INVOLVED IN SERVICE AND HYDROSTATIC TEST FAILURES Summary of failures analysed A very limited amount of data is publicly available concerning the failure stresses of real cracks for which all the geometrical details and material properties are well defined. The current paper considers two data-sets. The first relates to a series of fourteen SCC failures that was presented by a CEPA panel in the course of the NEB’s inquiry into stress corrosion cracking [11]; the second comprises eight hydrostatic test and service failures, mainly from SCC, experienced on two 323.9 mm OD oil pipelines operated by Pembina Pipeline Corporation.

For the failures reported by CEPA, the data available are limited to those presented in response to an undertaking, referenced in [11]. Toughness was presented in terms of full-scale equivalent Charpy energy, at a test temperature that is not defined. Actual yield strength and tensile strength were provided, and the flaw sizes are given in terms of length and depth. No information on flaw shape was given, and the cracks have been assumed to be semi-elliptical with the indicated length and maximum depth.

For the Pembina data, much more complete information is available, as shown in Table 1. Failure 1 is an operating failure caused by stress corrosion cracking on an HVP pipeline constructed in 1981.

Failures 2 to 6 are hydrostatic test failures caused by SCC on the same pipeline. Failure 8 is an operating failure in a crude oil pipeline constructed in 1961. Failure occurred at a fatigue crack which grew to critical size during 39 years of service from a hook crack in the electric-welded seam. Failure 7 is a burst test failure of a removed pipe section from the same crude oil pipeline. The failure in that test occurred at one of the weld seam flaws detected by in-line ultrasonic crack inspection.

Mechanical property data were obtained for each failure pipe. The data shown for pipe that failed in the weld seam area or pipe body were obtained from test specimens prepared to provide data representative of the relevant area. The reported Charpy shear area values show that the absorbed energy values, from tests at the operating temperatures, do not represent near-upper-shelf toughness for some pipe. The depths of the flaws where fracture initiation occurred were measured at intervals of 10 or 12 mm along the length of the fracture face for each flaw.

Table 1 Pipe properties and summary of flaw dimensions for Pembina failures

Results of analysis Battelle TD equation The results of the analysis carried out for the CEPA and Pembina data are shown in Figure 8. For the CEPA data, as mentioned, the flaw was assumed to be semi-elliptical; the equivalent flaw was taken 9

to have an effective length Leq equal to π/4 times the reported (surface) length and a maximum depth equal to the reported depth. This appears to be one of the few approximations to a geometrically regular flaw profile that is available to the analyst, in the absence of a measured profile. For the Pembina data, two semi-elliptical flaw shape assumptions were evaluated. The first of these used an effective length defined as above for the CEPA data and the maximum depth, while the second used an effective length based on the actual flaw area. The former is a reasonable choice among available regular flaw shape assumptions when flaw profiles are not known. The latter is the most accurate representation of the flaw profile that can be accommodated in the SF equation without iterative calculations to determine a worst-case effective flaw. The application of KAPA, which allows actual flaw shape to be taken into account in an iterative manner, is discussed below.

450

Observed failure stress, MPa

400

350 300 250 200 150

CEPA, elliptical, effective length

100

Pembina, elliptical, effective length Pembina actual area

50 0 0

50

100

150

200

250

300

350

400

450

Calculated failure stress, MPa

Figure 8 Observed failure stress against failure stress calculated using the Battelle TD surface flaw equation for CEPA and Pembina data It can be seen that, for the elliptical flaw shape assumption, all of the failure stresses are predicted conservatively. No trend lines are shown because there is a very low correlation between observed and calculated values relative to those that were obtained for regular, machined flaws5. The predictions for the Pembina data using the actual flaw area are considerably less conservative than those for the elliptical assumption, as might be expected, but the correlation is not greatly improved, and some non-conservative predictions are now observed. Overall, the results illustrate the difficulty of obtaining consistent predictions of failure stress for real, irregular cracks using the surface flaw equation and simple assumptions regarding flaw profile.

PAFFC Figure 9 shows the relationship between observed failure stress and that calculated using PAFFC. For both CEPA and Pembina cases, the flaw dimensions were input in terms of maximum length and depth. According to the software documentation, PAFFC analyses cracks as semi-elliptical. Thus, for the Pembina cases, an additional flaw length was input, based on the equivalent semi-elliptical area; this was considered to be the most accurate representation of the flaws available, consistent with the fracture mechanics formulations of the model. The correlation selected for determining material

5

In this and the figures that follow, where no trendline is shown, the calculated coefficient of determination was so low that it did not support the hypothesis of a linear relationship.

10

J-R characteristics in each case was selected based on the actual properties, regardless of the period of manufacture of the pipe. This practice is also recommended in the software documentation; in fact, the selection of a correlation that is inconsistent with the material properties produces a warning message. Two of the Pembina cases could not be analysed using maximum flaw length, since the length lay beyond the limit imposed by the software.

It can be seen that, based on maximum flaw length, the predictions for all of the Pembina data are conservative, while some of those for the CEPA data are non-conservative by varying amounts. In fact, the CEPA data fall into two groups. All five of the non-conservative predictions apply to 219.1 mm OD pipes, while the outside diameters of the other pipes are between 508 and 1067 mm.

However, the relative flaw sizes (in terms of L2/Dt) in the 219.1 mm OD pipes were not among the most extreme that were analysed. PAFFC and the underlying models were certainly developed for larger pipe sizes more typical of transmission service, but there does not appear to be anything inherent in the models or in the material property or flaw size data that are available that provides an obvious explanation of the significantly different behaviour of these five pipes.

The use of the equivalent elliptical flaw length for the Pembina cases obviously produces less conservative predictions than using maximum length, but the overall correlation improves only slightly.

Observed failure stress, MPa

500

400

300

200 CEPA

100

Pembina Pembina equivalent area elliptical

0 0

100

200

300

400

500

Calculated failure stress. MPa

Figure 9 Observed failure stress against failure stress calculated using PAFFC for CEPA and Pembina data CorLAS ™ For the CEPA data, flaw size was input in terms of the maximum length and depth, treated as a semielliptical flaw. For the Pembina data, analyses were conducted using several flaw-shape assumptions.

An initial analysis was based on the maximum length and depth, assuming a semi-elliptical flaw shape. A second analysis used the equivalent semi-elliptical flaw length based on actual area. The third analysis used the iterative capability in CorLAS™ to determine the most appropriate characterization of the flaw for each case. The results of the analyses are shown in Figure 10. For the CEPA data, the correlation between observed and calculated failure stresses is reasonably good, with a slope of 1.05 and no seriously non-conservative predictions. For the Pembina data, using the maximum flaw length, the correlation is poor for a regression forced through the origin (no regression line is shown), since the data points imply a significant positive intercept.

11

Observed failure stress. MPa

500

y = 1.1645x y = 1.0982x R² = 0.9003 R² = 0.4162

400 y = 1.0456x R² = 0.6594

300

200 CEPA elliptical

100

Pembina elliptical (maximum length and depth)

Pembina elliptical (effective length based on area) Pembina iterative

0 0

100

200

300

400

500

Calculated failure stress, MPa

Figure 10 Observed failure stress against failure stress calculated using CorLAS™ for CEPA and Pembina data When a length based on the equivalent elliptical area is used, an improved correlation with a slope of 1.10 is obtained. Using an iterative calculation, the correlation is greatly improved; the slope is 1.16, and there are no non-conservatively predicted points.

KAPA Figure 11 shows the results of the failure stress calculations for the Pembina failures using KAPA, which performs iterative effective flaw size calculations based on the Battelle TD equations and the detailed flaw depth profile. The KAPA correlation is somewhat improved compared to the Battelle results for assumed semi-elliptical equivalent flaws, but shows more scatter than the iterative CorLAS™ results (which were also shown in Figure 10).

500

y = 1.1645x R² = 0.9003

y = 1.8215x R² = 0.1209

Observed failure stress ,MPa

400

300 y = 1.2064x R² = 0.4782

200

Battelle elliptical (effective length based on area)

100

KAPA iterative CorLAS™ iterative

0 0

100

200

300

400

500

Calculated failure stress, MPa

Figure 11 Observed failure stress and calculated failure stress using KAPA, Battelle TD and iterative CorLAS™ for Pembina failures 12

FAD For the analysis using the FAD approach of BS7910 Level 2A, the flaw sizes were input as explained in the description of the FAD approach above, and the same iterative procedure was used to determine a stress corresponding to an assessment point lying on the assessment curve, both with and without the factor of 1.2 on the reference stress. As described previously, the fracture toughness was approximated using the same simple relationship based on Charpy energy that was used in CorLAS™ and in the Battelle equations. The results of the analyses are shown in Figure 12.

500

Observed failure stress, MPa

400

300

200 FAD FAD no factor

100

0 0

100

200

300

400

500

Calculated failure stress, MPa

Figure 12 Observed failure stress against failure stress calculated using FAD approach for CEPA and Pembina data All of the calculated Pembina results are conservative with and without the factor of 1.2 on the reference stress, while one of the CEPA points is marginally non-conservative without the 1.2 factor; the scatter is considerable, however, for both sets of results.

FACTORS THAT POTENTIALLY AFFECT THE ACCURACY OF FAILURE STRESS PREDICTIONS There are several factors that affect the accuracy of the predictions of failure stress using any of the methods discussed above. One of the most important is clearly the way in which flaw size and shape are represented. Significant cracks resulting from SCC often involve the coalescence of several smaller cracks, typically resulting in an irregular profile that cannot be represented accurately by any simple geometric shape. The use of a rectangular shape based on maximum length and depth will virtually always result in highly conservative predictions. Obviously, if the flaw profile is known precisely, as occurs in the case of service or hydrostatic test failure analysis, better representations of the flaw size and shape are possible. Under these circumstances, approaches like CorLAS™, that can use detailed flaw shape information to determine a minimum failure stress without overestimating the flaw severity, can produce quite accurate results for irregular flaws.

In many situations, such as assessment of features reported by in-line crack tools or field assessment of surface cracks, the only information about crack shape available is an estimate of the maximum depth and surface length. Although adjustments to the reported depth and length may be needed to account for reporting inaccuracy, the assessment must be made using a best estimate for the maximum depth and surface length. Figure 13 shows the observed failure stress compared to the calculated 13

failure stress for the CEPA and Pembina failures, assuming that the flaws are semi-elliptical with a length equal to the surface length and a depth equal to the maximum depth. The results indicate the performance of the assessment methods, assuming accurate knowledge of the material properties, and accurate values of depth and length.

500 y = 1.5243x

y = 1.4583x

y = 1.1219x R² = 0.6319

Observed failure tress, MPa

400 y = 1.2143x

300

FAD

200

Battelle TD CorLAS PAFFC

100

0 0

100

200

300

400

500

Calculated failure stress, MPa

Figure 13 Observed failure stress compared to calculated failure stress for CEPA and Pembina failures based on semi-elliptical flaws with length = total length and depth = maximum depth Trendlines are shown in Figure 13 to illustrate the average predictions of each method, even though the linear relationship is not significant except for the CorLAS method. All of the methods except PAFFC6 produce results that are conservative, apart from a few points slightly below the 1:1 line for the CorLAS method, which has the smallest amount of scatter. As shown in Table 2, the observed failure pressures are between 93% and 160% of the failure pressures calculated using the CorLAS method and an assumed semi-elliptical flaw shape. The range of ratios is much higher for the other methods.

Battelle TD PAFFC CorLAS FAD min max

1.03 5.51

0.72 2.52

0.93 1.60

1.11 7.36

Table 2 Ratio of observed to calculated failure pressures A further factor that can have significant influence on the failure stress predicted by all analysis methods is the fracture toughness of the material. Actual fracture toughness values are almost never available, though in the case of failure analysis they could usually be determined. As a result, correlations must be used to determine the required fracture toughness, in terms of K c, Jc or J-R, from Charpy energy. There is clearly a degree of scatter associated with these correlations, and the extent of this scatter, and of any intrinsic bias, will affect the accuracy of failure stress predictions. An additional complication is that, at least in the case of the Battelle equations, the Charpy energy that is intended to be used is the upper shelf energy. This is because, for materials similar to those on which the equations were validated, and for a surface flaw, there is expected to be a transition temperature decrease of at least 70°C from a Charpy test to a static, full-scale failure. The original validation of 6

When the five data-points for the 219.1 mm OD pipes are removed, the minimum ratio is 1.18; the slope of the trend-line through the origin becomes 1.69.

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the TD surface flaw equation clearly was based on shelf energy, and the good accuracy of the other approaches that were evaluated against the same data, using the same toughness inputs, confirms that this is also appropriate for these assessment methods. For the assessment of real failures, however, the shelf energy is often unknown, though again it could be determined in the case of hydrostatic test failures and service failures for which undamaged pipe is available. More commonly, assessment is performed on the basis of Charpy energy at a specific temperature (either from original mill certificates or from post-failure tests). In the case of the CEPA tests, for example, no information was provided relative to the test temperature. Some Charpy values appear to be near upper shelf, while others are clearly lower shelf values that would be expected to give very conservative failure stress predictions unless the Charpy transition temperature is extremely high. If shear area values are available, it is possible to estimate shelf energy from energy at temperature, but such estimates can be very inaccurate, particularly if the Charpy tests were carried out on sub-size specimens. For the Pembina tests, the test temperature and shear area were known. Limited analysis based on estimated shelf energy showed a tendency towards reduced conservatism, but little reduction in scatter.

IMPLICATIONS FOR THE PRACTICAL APPLICATION OF ASSESSMENT MODELS IN INTEGRITY MANAGEMENT PROGRAMS Understanding the accuracy of assessment procedures applied to real flaws is critical to their use in integrity management. There are clearly several key elements involved in determining this accuracy: What is the intrinsic accuracy of the procedure?

How accurately are the relevant material properties known? If correlations or estimation procedures are needed in the analysis, how accurate are they?

How accurately is flaw size known? What is the intrinsic accuracy of the inspection method?

In particular, what does the reported flaw depth actually represent, and how should this be input to the assessment procedure?

The preceding sections have considered some of these issues. In particular, it has been shown that all of the assessment methods are capable of predicting the failure stress of regular, sharp, machined flaws in pipe with well-defined properties with good accuracy. Although the slope of the mean correlation varies somewhat between the different methods, a desired degree of conservatism for a given application could be achieved quite easily. For the selection of real cracks from service and test failures that was examined in this study, the position is less positive. Though, with a few exceptions, all approaches gave failure stress predictions that were conservative or nearly so, there was a great degree of scatter in most cases. CorLAS™ gave the best results, particularly when actual flaw profiles could be used to determine the effective combination of flaw length and depth. This method can be applied to failure analysis situations, but in cases where the severity of flaws identified by nondestructive means must be assessed, there is little alternative to applying a conservative representation of flaw length and depth. This implies that the physical meaning and accuracy of indications from inline or other non-destructive evaluation must be well-understood, in order to choose suitable inputs to the assessment procedure.

As discussed above, it is probable that a considerable part of the inaccuracy in the current study arises from inaccurate characterization of fracture resistance. For integrity management programs, it will always be important to have the best possible information on material toughness properties. For older pipelines, limited or no toughness data may be available from mill certificates. Every effort needs to be made to acquire such data when opportunities for sampling arise, and sufficient testing should be carried out to enable Charpy shelf energy to be determined. For new pipelines, Charpy data will be available from mill certificates; in the majority of cases, the Charpy energy at the minimum service temperature will be representative of the shelf energy. For large orders, consideration should be given 15

to acquiring fracture toughness data, such as K, J and dJ/da, on a reasonably representative sample of the overall production, so that material-specific correlations with Charpy energy can be established for future use, should the need arise.

In the evaluations reported in the preceding sections, a prediction of the failure stress greater than the actual failure stress has been considered to be conservative. This is appropriate, so long as the assessment is applied to the acceptability of an existing flaw. However, when the same methods are applied to estimate the maximum size of flaws remaining after a hydrostatic test, such a prediction becomes non-conservative; larger flaws remain than are predicted by the model. Of course, the actual size of a critical flaw under operating conditions will also be larger than predicted by the model, but it is not intuitively obvious that the magnitude of the error will be the same, or that the effect on remaining life estimates under specific flaw growth conditions (e.g. by SCC or fatigue) will be minimal. For this reason, it is important to consider carefully the effect of errors relative to the specific application being considered. The variability of material properties, particularly toughness, may be of particular significance in this context. Toughness may considerably influence the maximum surviving flaw size, while the effect of any related change in critical flaw size near the end of service life may be very limited.

CONCLUSIONS In this paper, several alternative methods for the assessment of sharp axial surface flaws in pipe have been evaluated, starting with the Battelle surface flaw equations that date back to the early 1970s, and including several more recently-developed approaches. All of the methods were first benchmarked against the original Battelle database of tests on pipe vessels containing sharp machined notches, and then tested against two data-sets involving hydrostatic test or service failures from real cracks, for the most part caused by SCC. Overall, the following conclusions can be drawn.

All of the assessment methods performed relatively well on the Battelle database, which involved machined flaws with regular, well-characterized profiles, with material toughness represented by Charpy shelf energy. Mean predictions were generally within a few percent of the observed values, with few outliers.

While all of the methods gave predominantly conservative predictions for the real cracks, correlations between calculated and observed failure stress were generally poor. The wide scatter is considered to be largely the result of the difficulty of representing real crack profiles (where known) in a consistent way within the different models, and to variability in effective fracture toughness arising from generic correlations with Charpy energy, particularly when the shelf energy is not known and energy at an arbitrary test temperature is used.

An exception to the overall pattern appears to be CorLAS™. This method produced an excellent correlation between calculated and observed failure stress for irregular crack shapes when the detailed crack profile was used, and performed significantly better than the other approaches when the cracks were represented as having a semi-elliptical profile using maximum length and depth.

In general, it can be concluded that any of the methods evaluated can give conservative estimates of the failure stress for a given flaw, when an appropriate safety factor that takes account of model uncertainty as well as measurement uncertainty is applied. However, in most cases where detailed flaw shape is not known, the degree of conservatism can be very variable. The results suggest that caution needs to be exercised in the evaluation of remaining life after hydrostatic testing, since estimates of remaining flaw size are subject to considerable error arising from both the assessment method itself and the variability of mechanical properties (particularly fracture resistance) within a 16

pipeline. Corresponding errors when determining the end-of-life flaw size do not necessarily offset the initial error to produce the same estimate of remaining life.

ACKNOWLEDGEMENTS The authors would like to thank the management of Pembina Pipeline Corporation for permission to use the data related to hydrostatic test and service failures that are presented in this paper.

REFERENCES 1. Kiefner, J.F., 1969. Paper G, 4th Symposium on Line Pipe Research, PRCI Catalogue No. 30075e, PRCI, Arlington, Virginia.

2. J.F. Kiefner, W.A. Maxey, R.J. Eiber and A.R. Duffy, 1973, “Failure Stress Levels of Flaws in Pressurized Cylinders”, ASTM STP 536, “Progress in Flaw Growth and Fracture Toughness Testing”, p. 461, ASTM, Philadelphia.

3. “Kiefner & Associates Pipe Assessment”, www.kiefner.com 4. Leis, B.N., Brust, F.W., and Scott, P.M., 1991, “Development and Validation of a Ductile Flaw Growth Analysis for Gas Transmission Line Pipe”. PRCI Catalogue No. L51643.

5. B.N. Leis and N.D. Ghadiali, 1994, “Pipe Axial Flaw Failure Criteria – PAFFC Version 1.0 User’s Manual and Software”. PRCI Catalogue No. L51720.

6. Jaske, C.E., 1996, “CorLAS™ 1.0 User Manual: Computer Program for Corrosion-Life Assessment of Piping and Pressure Vessels,” Version 1.0, CC Technologies Systems, Inc., Dublin, OH, 1996.

7. Jaske, C.E., 2002, "Development and Evaluation of Improved Model for Engineering Critical Assessment of Pipelines”, ASME IPC2002-27027.

8. Jaske, C.E., 1999, “Addendum to User Manual for Version 2.0 of CorLAS ”, CC Technologies Systems Inc., Dublin OH. 9. C.E. Jaske and J.A. Beavers, 2001, “Integrity and Remaining Life of Pipe with Stress Corrosion Cracking”, PRCI Catalogue No. L51928.

10. Anon., BS7910:1999 “Guidelines on Methods for Assessing the Acceptability of Flaws in Metallic Structures”, BSI, London, UK. 11. National Energy Board, 1996, Report of the Inquiry “Stress Corrosion Cracking in Canadian Oil and Gas Pipelines”, Hearing MH-2-95.

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