step improvements in metal-loss assessment critiria …

22nd JTM, 29 April – 3 May 2019, Brisbane, Australia of 14 Leis et al

STEP IMPROVEMENTS IN METAL-LOSS ASSESSMENT CRITIRIA FOR PIPELINES

Brian Leis* B N Leis Consultant, Inc, Worthington, OH, USA Xian-Kui Zhu and Tom McGaughy Edison Welding Institute, Columbus, OH, USA David Aguiar PG&E, San Ramon, CA, USA Shahani Kariyawasam TCPL, Calgary, AB, Canada Laurie Perry PRCI, Chantily, VA, USA * presenting author; [email protected]; 001-614-802-9376

ABSTRACT While ASME B31G has been expanded to provide options that reflect its adaptation to modern higher-strength steels, model uncertainty has remained in the context of the two forms of flow stress still used, and in the form of the bulging factor (BF) and the shape factor (SF). PRCI Project EC-2-7 was initiated to minimize model error entering B31G and Modified B31G with respect to the BF and SF, whereas Project EC-2-6 sought to minimize model error that entered in regard to the reference (or flow) stress. In addition to minimizing model error, this Project sought a reference stress that could be seamless from the early vintage steels and Grades that underlay the calibration of B31G through modern steels in Grades up to X80 (552 MPa).

Following a brief review of the background, this paper first presents the alternative reference stress adapted in EC-2-6 for use with the outcomes of Project EC-2-7. This alternative reference stress is a function of the UTS and the strain-hardening exponent of the steel, and as such has the potential to account for differences in the steels used in pipelines from the early years through the present. Coupled with the analytical results of EC-2-7 in regard to the BF this technology was used to design a full-scale validation test matrix. The objective of that design was to broaden the scope of the current test database to include longer features and also to explore width effects – along with the usual consideration of defect depth.

The results of the full-scale testing indicated that accurate precise predictions were possible with this alternative metal-loss-severity criterion. A similar outcome was achieved when this approach was applied across the available literature full-scale test database. This was achieved with little bias, and much reduced scatter as compared to the corresponding predictions made using B31G and Modified B31G. In practical terms, reduced scatter means less conservatism is needed to achieve the same level of safety, such that rehab or repair that does not affect reduced risk can be avoided. These outcomes indicated that in contrast to the existing models, response to manage such features could be deferred – resulting in a significant reduction in the overall number of required digs and extending the timeframe for those features needing consideration.

The paper first details the alternative reference stress, and its validation based on literature and archival data, and thereafter presents the results of the recent full-scale metal-loss testing that considered longer and wider metal-loss features. Finally, the paper presents the validation database, including consideration of predictions made for a range of alternative models that have emerged to predict the pressure of metal-loss defects.

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1. BACKGROUND In the late 1960s the pipeline industry began work to quantify the effects of metal loss due to corrosion on the failure pressure of transmission pipelines. This early work was empirical in nature. Areas of corrosion in pipes that were removed from service during pipeline rehabilitation were sized and otherwise documented, then end-capped and pressurized to failure. The resulting data provided the basis to understand the failure process, and to empirically establish safe operating limits in the presence of such defects [1][2]. The outcome of that and other work addressing crack-like features [3][4] led to the AGA NG-18 Equation that quantified failure in terms of unstable axial growth of through-wall (TW) defects [5], whose form traces to prior work due to Hahn, Duffy, and others [3][4]. Modifications to the TW defect equation led to a form dealing with quasi-static failure at part-through-wall (PTW) defects [6]. As the toughness increased to a sufficiently large value, these TW and PTW equations become practically independent of toughness. As such, these “NG-18 Equations” could characterize fracture-controlled failure at lower toughness levels, and collapse-controlled failure at higher toughness levels.

Because blunt metal-loss defects typically failed under collapse control [1][2], the toughness independent form of the PTW NG-18 Equation was adapted to quantify the failure pressure for corrosion. As expected for a Level 11 approach to assess defect severity, this corrosion criterion expressed the failure pressure, Pf, in the simple functional form:

Pf = SR ⨯ t/R ⨯ (f (defect geometry)). (1)

The symbols t and R denote the pipe’s thickness and radius, whereas SR denotes a “reference stress” that quantifies the line-pipe’s inherent resistance to failure, and the last term quantifies the severity of the defect as a function of its geometry.

Given the above functional form, as the size of the defect decreases to zero, the value of the term f (defect geometry) must tend to unity, such that the predicted failure pressure must tend to a value that is consistent with that observed for defect-free pipes. Recognizing that the hoop stress at failure for defect-free pipes is close to the ultimate tensile stress (UTS) for a very broad range of line-pipe grades [8][9][10], SR should have a value close to the UTS. In contrast, the reference stress used in the initial empirical fit that gave rise to ASME B31G circa 1984 [2][11][12] was expressed as a function of the actual flattened-transverse-strap yield stress, AYS. That reference stress was termed a “flow stress” and expressed as AYS + 10 ksi (68.9 MPa). This was a simplification of the flow stress determined by Hahn et al [3], whose value based on a best fit to failure data for pipes containing axial slits was 1.04 AYS + 10 ksi (68.9 MPa). But, when ASME B31G was released in 1984, this flow stress was changed to SR = 1.1 ⨯ (SMYS) [13]. When Modified B31G was released in 1989, the value of SR was returned to its original form – except that the AYS was replaced by the specified minimum yield stress (SMYS) [14], i.e., SR = SMYS + 10 ksi (68.9 MPa).

As many different definitions of reference stress have been adopted over the years across the many forms of Equation 1 that have been proposed, Project EC-2-6 evaluated such work with a view to minimize the modeling error that entered B31G and Modified B31G related to the reference (or flow) stress. In particular, this Project sought a reference stress that was:

1) consistent with the form of Equation 1, the available data for full-scale testing of defect-free pups, and the mechanics of their failure; which also would be

2) seamless from the early vintage steels and Grades that underlay the calibration of B31G, through the modern steels in Grades up to X80 (552 MPa).

In parallel, Phase II of Project EC-2-7 developed a test matrix to evaluate the viability of the alternative bulging factor (BF) developed in its Phase I work [15][16]. Phase II of EC-2-7 also compiled the published full-scale test results for other such testing on both defect-free pups as well as for machined metal-loss defects. Following completion of the testing, these data were used to evaluate the utility of 1) the alternative reference stress and 2) the predictive outcome of that reference stress and the alternative BF.

This paper presents the just completed work for Project EC-2-6 [17] and Phase II of Project EC-2-7 [18]. The development and validation of the alternative reference stress is detailed first. Thereafter, the scope of the full-scale metal-loss testing designed to evaluate/validate the alternative BF is presented. As detailed in References 15 and 16, this BF is a function of defect’s depth, width, and

1 API 579 [7] defines levels of complexity in such applications, with 1 being simple equations, 2 reflecting more complex schemes,

like software-based methods, and 3 being complex/case specific, like finite element modeling.


length, in contrast to the Folias-based BF [19][20] used in the NG-18 PTW collapse equation, which as it was for TW slits depended only on the defect’s length. That discussion is followed by presentation of the full-scale test results, and the evaluation of this new technology, in comparison to the published full-scale test results, and a range of other metal-loss criteria.

2. DEVELOPMENT AND VALIDATION OF THE ALTERNATIVE REFERENCE STRESS 2.1 Literature Review Comprehensive reviews of the relevant published literature have been reported recently, with two notable for present purposes: Zimmermann et al. [21] and Zhu and Leis [22]. These reviews evaluated the range of the available criteria that were potentially useful as the basis for an alternative reference stress. Both reviews statistically quantifying their predictability relative to full-scale data in terms of predictive bias, and dispersion for grades ranging from those in early use up through X100. Both reviews independently concluded that the Zhu-Leis (Z-L) criterion [23] was best.

2.2 The Proposed Alternative Reference Stress Because pipelines and defect-free end-capped pipe tests involve a biaxial state of stress it was appropriate to formulate the Zhu-Leis criterion in a multi-axial setting. This was done in terms of an equivalent uniaxial stress based on Tresca [24] and von Mises [25] criteria, and on an average shear-stress criterion, which was defined as the average the maximum shear stress, Ƭmax, and the von Mises effective shear stress, ƬvM. As that the von Mises shear stress is related to the octahedral shear stress, Ƭoct, as ƬvM = (3/2)0.5Ƭoct, the average shear stress is a weighted average of the maximum shear stress and the octahedral shear stress.

For a thin-walled, end-capped line pipe subjected to internal pressure, failure according to the average shear-stress criterion occurs at [23]:

. (2)

where PZL is the Z-L predicted failure pressure, n is the power-law strain-hardening exponent, D is the nominal outside pipe diameter, t is its nominal wall thickness, Dm = D - t is its nominal mean diameter, and Suts denotes the UTS for the pipe steel. This failure criterion reflects the use of an equivalent average shear-stress criterion, a power-law stress-strain constitutive relation, and a corresponding equivalent average shear-strain. Through use of a correlation between n and the yield to tensile ratio, Y/T (e.g.,[26]), this criterion also can be expressed in terms of Y/T. Corresponding forms have been developed for both Tresca and von Mises multi-axial equivalence criteria (see [23]).

Inspection of Equation 2 and its development indicates that the alternative reference stress is:

. (3)

2.3 Evaluation and Validation Given that pipeline-related work has seldom characterized and published values of n along with related data for full-scale pipe-pup tests, results that could be used to directly validate Equation 2, and by inference Equation 3, are limited. Fortunately, that limited database can be supplemented through consideration of pup testing where Y/T is surrogate for n, or where the value of the strain at the UTS is reported (i.e., the uniform strain, eu) as the basis to calculate the true value εu = ln(1+eu). Note in this regard that the uniform true strain is a direct analog for n for power-law hardening steels (via Considere’s construction [27]). This paper makes use of data developed by each of these approaches, the results of which are shown in Figure1 for cases where Y/T was used as a surrogate for n, in Figure 2a for cases where n was determined, and in Figure 2b for cases where eu was reported as an analog to n.

Figure 1 is reproduced from Reference 23, wherein the predicted failure pressure for Tresca, von Mises, and Z-L criteria was compared to the observed full-scale burst pressure for defect-free end-capped pups. In total these data represent 103 full-scale tests primarily involving pipe steels for which the strain hardening exponent n was estimated using a correlation between n and Y/T [26]. The citations for each dataset noted in this figure refer to the list of references in the original paper [23]. Included were 41 tests by Kiefner involving line-pipe grades from Gr B to X65, three tests by Maxey involving X70, two by Mok et al. involving X60, six by Chouchaoui involving X46, 12 by Amano et al. involving X65 and X70, two by Hillenbrand et al. involving X100, eight by Liessem et al. involving X60, X65, X70, and X80, two by Okaguchi et al. involving X80 and X120, nine by Papka et al. involving X120, five by Law et al. involving X60 and X80, and 13 tests by Paslay et al. involving H40-Q125 casing steel. One of these

utsm

1n

ZL SD

t2

34

322P

+

+=

uts

1n

R S34

322S

+

+=


datasets (Kiefner et al) in some cases also reported values of n that could be confirmed from file data, whereas a second among these (Amano et al [29]) also reported values of eu

2. These datasets respectively underlie Figure 2a [28] and Figure 2b [29], wherein the vertical scales differ as compared to that in Figure 1.

Figure 1: Predicted failure pressure for Tresca, von Mises, and Z-L criteria

Figure 1 indicates that the trend for the Z-L predicted failure pressure is consistent with the central tendency for this dataset which ranges from the early lower-strength Grades up through the modern higher-strength Grades. These results also indicate that the von Mises-based solution serves as an upper bound to these data, whereas the Tresca-based solution serves as a lower bound. Scatter is also evident, although this seems to be less for the higher-strength grades that reflect more recent, possibly improved testing practices. Scatter also enters this comparison by way of model uncertainty in the correlation used between n and Y/T. Finally, scatter enters due to uncertainty in the inherent properties and the experimental aspects. That being said, these results do clearly indicate the dependence of the reference stress on n and show that dependence on average is consistent with Equation 3.

a) results of Kiefner et al, versus n b) results of Amano et al, versus εu

Figure 2: Predicted failure pressure for the Z-L criteria

Figure 2a is comparable in format to Figure 1 except that these data are presented relative to a measured value of n that is specific to each of the burst tests. The scatter in this case is much reduced as compared to Figure 1, which in part traces to the reduced number of results, and the absence of scatter that enters by way of model uncertainty in the correlation used between n and Y/T. It is evident from this plot that the Z-L criterion readily predicts the failure pressure, and that the trend inferred for the reference stress given as Equation 3 corresponds to the effects of n on the failure pressure.

2 This notation reflects the engineering strain at the ultimate stress, which is what is usually reported, whereas the

power-law value of n reflects true stress and plastic (≈total) strain at the UTS, as is plotted in Figure 2b. The quantitative difference between eu and εu is relatively small such that eu also could be used as a surrogate.


Figure 2b also is comparable in format to Figure 1 except that these data are presented relative to a value of εu that is specific to each of the burst tests. As noted above, εu is a direct analog to the value of n for power-law hardening materials. As in Figure 2a, the scatter is much reduced as compared to Figure 1. As noted above, this in part traces to the reduced number of results, as well as to the absence of scatter due to model uncertainty in the correlation used between n and Y/T. It is evident from this plot that the Z-L criterion readily predicts the failure pressure, and that the trend inferred for the reference stress given as Equation 3 corresponds to the effects of n on the failure pressure.

2.4 Summary Regarding SR It is evident from Figure 1 that the Z-L criterion and by inference its reference stress broadly tracks the central tendency of the failure behavior for a wide range of grades from Gr B through X120. Figures 2a and 2b indicate a similar correspondence and show that the failure pressure varies in a significant way with the strain-hardening exponent for such grades. These figures also show that the reference stress used in the Z-L criterion closely tracks this failure response for 0.04≤n≤0.22. One can conclude on this basis that the proposed alternative reference stress is validated for most line-pipe steels.

3. DESIGN OF THE FULL-SCALE METAL-LOSS TEST MATRIX AND RESULTS 3.1 Background Work reported in the last JTM [16] sought to minimize the modeling error that entered B31G and Modified B31G through their definitions of the bulging factor (BF) and the shape factor (SF). As reported there, that work (done as part of Phase I of EC-2-7) indicated that the BF was a function of defect’s depth, width, and length, in contrast to the BF used in B31G and Modified B31G which was derived in reference to TW slits and as such depended only on the defect’s length. Figure 3a indicates that the effect of depth alone on the value of the BF is significant for longer defects that are shallower than one-half of the wall thickness. In contrast to this observation, the majority of prior full-scale testing had not considered such features. Figure 3b indicates that for such cases a comparable effect in the context of the failure pressure. Accordingly, a test matrix was designed to evaluate the role of depth at the same time it provided a link to the prior experiments and also explored the role of width. As the focus was the effect for the SF, this work considered flat-bottomed features for which the SF = 1.

a) effect of depth on the BF b) relative failure pressure as a function of BF

Figure 3: Effects of defect depth on the BF and its effect on the failure pressure

3.2 Experimental Design and Test Preparation This full-scale testing used 24x0.492-inch (610x12.7 mm) thick wall recent X70 (L485M) production with a budget sufficient to complete seven tests. It was reported that the pipe joints supplied for this testing were not expanded. The joints supplied were uncoated, from the same heat, and close together in the production sequence.

As the experimental design sought to eliminate the complicating effects of variable depth, and so the role of SF, flat-bottomed features were used (SF = 1). A range of potential feature lengths, depths, and widths was considered based on their failure pressures predicted using values of their BFs as discussed above, coupled with the reference stress given in Equation 3. For purposes of the initial experimental design, as the value of n had not yet been quantified for the test pipes an estimate for it was based on mill data for Y/T for the heat and the n –Y/T correlation noted earlier [26]. Based on those results feature sizes were chosen for the seven tests and machined. The matrix adopted used rectangular isolated features with lengths between 3 ≤ L / (D t)0.5 ≤ 10, and bounding values of width taken at arc angles of 20 and 60 degrees, with depths at d/t of 0.30 and 0.80. The dimensions follow later in detail


– along with the results of this testing. This set of features were centered in pups that were ~3.7 m (12 feet) in length of which the central ~2.44 m (8 feet) was test pipe. This length suffices to prevent any influence of the end caps. While these features were being machined, the shape of the pipe was constrained on the OD and the ID with a view to minimize variation in the depth across the bottoms of these features. Local variations in wall at the ID was not offset by this practice.

Each feature was then documented using a grid of points, then a mix of biaxial and uniaxial strain gages were applied and connected along with other necessary instrumentation. Thereafter, the pups were end-capped, after which the features were speckled for use with two-dimensional digital-image-correlation (DIC) practices, and then pressurized with water to failure using proven practices. Figure 4a presents some of the details of this setup in regard to the strain gaging, whereas Figure 4b shows another feature after its preparation for DIC analysis.

a) feature prepared with strain gages b) gaged feature prepared for DIC analysis

Figure 4: Aspects of the test practices (not to scale)

3.3 Test-Pipe Properties Usual testing to quantify the pipe’s mechanical and toughness properties was done consistent with ASTM E08 and ASTM E23, respectively [30][31]. The chemical composition was also determined (using optical emission spectroscopy [32]). The results of the mill pipe-acceptance testing indicated that the lot of pipes that the test pipes came from satisfied the applicable API 5L requirements for PSL2 X70M pipe [33]. Testing done specific to the pipes are reported here involved strap specimens that were flattened between platens in accordance with API 5L. Because this metal-loss testing involved a reduced section of the full-pipe-wall, this flattened strap testing was done to sample the response of the full-pipe-wall, as well as the response of part-thickness specimens cut from the inner-half of the wall. These part-thickness samples were nominally 0.26 inch (6.5 mm) thick, with this testing done with a view to better quantify the tensile properties in material more typical of where failure would initiate in the metal-loss testing.

Figure 5 shows a profile view of these strap samples whose planforms were geometrically similar, with the OD surface shown up in this view. Note that the full-thickness sample failed on an inclined (shear) plane, whereas the reduced thickness sample failed on a plane normal to the pipe’s surface. Related to this, the reduced thickness sample shows increased ductility through-thickness. The features in both cases showed traits typical of micro-void coalescence throughout these failure planes. But, while both were fully ductile failures, it is clear that reducing the specimen thickness has affected their failure response, with those effects embedded in the properties so measured. Such effects can complicate interpreting the results of tension testing of samples that otherwise have comparable scaled planforms, with such effects being well documented [e.g., [34]] and considered in later discussion.

Figure 5: Comparing the profiles of the tensile strap samples

All tensile testing was done in a servo-hydraulic test frame under constant displacement control at a rate on the order of 0.10-in./min (2.5-mm/min). Load was measured using a load-cell mounted in series


with the sample. A high capacity extensometer was used to monitor strain in the reduced section through ultimate load, after which it was removed. As the part-thickness samples required machining, this was done following flattening. Thus, no secondary bending or asymmetric loading occurred during their testing. The output was tracked via analog methods and converted to a digital format to facilitate analysis. While testing was done in both principal directions of the pipe and is reported in detail in Reference 18, only the transverse results for replicate testing are presented herein.

The full-thickness transverse AYS was 72.8 ksi (501 MPa) with a UTS of 92.0 ksi (634 MPa), such that Y/T was 0.77, with a eu of 0.0974. The corresponding part-thickness testing showed an AYS of 85.4 ksi (588MPa) with a UTS of 91.5 ksi (630 MPa), such that Y/T was 0.93, with a eu of 0.0652. The full-thickness value of n was 0.0747. As in some cases the part-thickness results showed clear bilinear power hardening, the part-thickness value of n reported herein reflects the flow response approaching collapse, which was 0.0584.

Toughness testing using full-size Charpy-vee-notch (CVN) samples was done in both principal directions of the pipe, leading to ductile-to-brittle transition curves, as detailed in Reference 18. Suffice it here to introduce results characteristic of the resistance to axial TW growth, or the C-L direction, with the notch machined on the through-wall face. These results showed fully ductile energy response increased as the test temperature increased, with both directions showing similar upper shelf resistance ranging from 147 to 236 ft-lbs (200 and 320 J) and 100% shear evident above 32°F (0°C). Such rising upper-shelf behaviour is common for modern high-toughness line pipe steels. It follows from these results that ductile failure is anticipated for burst tests done under ambient conditions.

3.4 Full-Scale Test Results Table 1 summarizes the feature sizes and the corresponding failure pressures. As evident therein, these failure pressures, Pf, ranged from 0.337 ≤ Pf (%SMYS) ≤ 1.229. In this context one target of this seven-test experimental design – to cover a broad range of normalized failure pressures – has been achieved. The feature dimensions detailed therein had normalized lengths (i.e., L / (D t)0.5) of 3, 5, and 10, for widths (as arc angle) of either 20 or 60 degrees, and normalized depths (i.e., d/t) of either 0.30 or 0.80. Through use of precise controls during machining, which was facilitated through use of both equipment and by constraining the pipe’s shape, based on measured values these dimensions were consistently achieved.

Table 1 Details of the features tested and their failure behavior

Of interest beyond the earlier discussion of this experimental plan is the observation that is led to differences in the origin of the failure within the feature, as well as regarding whether the defect extended axially (rupture) or did not (leak). Note that this definition of leak and rupture is concerned with whether or not the failure could have propagated axially beyond the constrains of the defect boundary. Further details of the test protocol and the results Details of this testing can be found in Reference 18.

3.5 Predicted Failure Behavior and Comparison with the Full-Scale Results Predictions were made for these seven tests selectively using finite element analysis (FEA) and through use of the BF and reference stress, coupled with the pipe’s properties. Predictions also were made using B31G, Modified B31G, PCORRC [35], LPC-1 [36], BS7910 [37], and DNV RP-F101 [38]. Of these PCORRC makes use of a reference stress that is of the same order as that developed by Equation 3. Given the Corrosion Committee’s desire for an independent assessment, such predictions were also done for the same set of criteria by C-FER. As this testing involved flat-bottomed features, to reduce the bias all predictions involving B31G and Modified B31G replaced their SFs (i.e., the value 0.67 in B31G and 0.85 in Modified B31G) with a value of one. The author’s predictions are detailed in Table 2. Those by C-FER, which were identical, can be found in an Appendix to Reference


18, which embeds their reporting in that Appendix. Consider first predictions for B31G and Modified B31G in contrast to the reference stress and BF discussed earlier herein. Thereafter, the broader scope of predictions will be considered later along with discussion of the literature data for flat-bottomed features in X70 pipe.

Table 2 Predicted versus actual failure behavior for the full-scale tests in Table 1

This tabulation shows that predictions made using the existing criteria lead to a quite large bias. This is evident in the summary lines for this tabulation that show the mean ratio of the predicted to actual failure pressure for B31G was 0.75, while that for Modified B31G was 0.81. Had the SFs specific to these criteria been used, the mean ratio of predicted to actual failure pressures would be much smaller still (i.e., even larger bias). For the SF taken at unity the values of the coefficient of variation (CoV) were 10.4% for B31G, and 10.9% for Modified B31G.

Table 2 indicates that the predictions made using the technology developed and discussed earlier in this paper led to a mean of 1.017, with a CoV of just 8%. It is apparent that the largest predictive error occurs for the only test that sought to identify the effects of width on the current state of this development – which is Test 7 (with an arc angle of 60 degrees). If that test is excluded, the bias decreases as the average trends to conservative at 0.991, with a CoV that drops to just 4.7%. Inspection of the statistics summarized in the lower lines of this tabulation indicates that the bias has been significantly reduced as compared to B31G and Modified B31G. In turn, this corresponds to a significant reduction in the digs and maintenance required to affect the same level of safety. Likewise, a 20% reduction in scatter was achieved by this developing technology, which contributes to the reduced effort needed to achieve the same level of safety. Reduced model uncertainty (i.e., scatter) coupled with reduced bias affords predictions of both failure pressure and defect size that lead to reduced maintenance (digs and rehab) that focus on the features that affect risk. As these predictions limit consideration of features that would not fail in hydrotesting, they affect reduced maintenance without reducing safety.

FEA predictions were also made based on the same set of properties. These predictions were done specific to Tests 1 and 7, which reflect the extremes in the tabulated failure pressures. The FEA predicted failure for Test 1 was 1.233 Psmys, while that for Test 7 was 0.337 Psmys. As Table 2 shows, these predictions were sensitive to the failure location – that is the location within the flat-bottomed feature where failure initiates. It follows that while the current developments dealing with the BF and the reference stress give rise to a significant reduction in the bias and in the scatter, they fall short of the quality achieved using a Level 3 method – which is to be expected. While the present Level 1 predictions are much improved as compared to B31G and Modified B31G, minor disparities remain as evident in this tabulation. These reflect 1) the rudimentary modeling of width effects to this point, and 2) the observation that where failure occurs reflects the flow properties, as well as the planar size and depth of the features. Work has been planned as part of EC-2-10 to address these aspects.

3.6 Full-Scale Testing Summary The full-scale testing was completed with practices that sought to minimize the test variability – but reality in this context is that properties variability remains inherent in all such results. Although the properties testing made clear that both the UTS and n showed scatter, along with some major differences in the flow response, predictions must be made without access to the parameters specific to where failure will occur. Consequently the predictions made using this developing technology affected step improvements in the context of these full-scale tests.

4. EVALUATION BASED ON THE PUBLISHED FULL-SCALE METAL-LOSS TESTING


The present experimental plan was directed at selectively expanding the scope of the current full-scale test database for flat-bottomed features, which is substantial particularly with respect to shorter and narrower features [18][39][40][41][42][43]. These data also can be used to evaluate the improvements outlined herein. As is generally the case, such data do not include details concerning the flow properties, such that the n – Y/T correlation of Reference 26 has been employed in these predictions, with Y/T serving as a surrogate for n. In total, this X70 (L485) database covers 75 tests, which includes the seven dealt with above in Tables 1 and 2. This composite database covers failure pressures that ranged from 0.3 ≤ Pf (%SMYS) ≤ 1.5, with normalized lengths (i.e., L / (D t)0.5) from ~0 to 14, for widths (as arc angle) from ~0 to 75 degrees, and normalized depths (i.e., d/t) from near zero to approaching one. Because the predictions were similar whether done by the authors or by C-FER, the ensuing results reflect the independent trending by C-FER. Their assessment considered model bias and scatter in terms of CoV, as were reported above in Table 2 for the testing detailed herein, with their results presented in graphical form in Figure 6.

Figure 6: Comparison of predicted versus actual burst pressure for the known models (C-FER)

Inspection of these plots and related tabulations presented by C-FER indicates that their analysis showed that the quality of the predictions made by the combination of reference stress and the BF outlined earlier model is comparable in terms of predictive bias and variability to the LPC-1, DNV F101, and PCORRC models. Taken together those predictions were somewhat better than those by BS7910, and as noted above in regard to Table 2, the current work provided a clear improvement versus B31G and Modified B31G (whose predictive outcomes were similar, as is evident in Table 2).

Note that this figure includes results for the BF absent width effects, which reflects the observation that the focus of work to date on the BF has focused on the roles of depth and length. Accordingly, the quality of the predictions achieved for the technology reported herein will improve still further once the clear effects of feature width are incorporated. This point was made by C-FER in their discussion of this work at the PRCI Spring Committee Meeting in Miami (in 2018).

In view of the above it follows that the viability of the present approach has been demonstrated for flat-bottomed defects, with the potential for still further step gains in predictability available by the inclusion of width effects, which is now early in its development. Consideration of SF thereafter will address


‘riverbottom’ effects and provide the basis to quantify the consequences of failure in terms of Leak versus Rupture as well as the scale of the release. Each step along this path affects a reduction in model uncertainty, which in turn opens to a reduction in the scope of digs and maintenance needed to affect the same level of safety. Work has been planned as part of EC-2-10 to address these aspects.

5. DISCUSSION AND VALIDATION OF THE TECHNOLOGY DEVELOPED An alternative reference stress has been presented that is a function of the UTS and the strain-hardening exponent of the steel, which has the potential to seamlessly account for differences in the steels used in pipelines from the early years through the present. Coupled with the analytical results of EC-2-7 related to the BF, this technology was evaluated first in regard to full-scale tests designed to quantify bulging response and its effects at the practical limits of metal-loss features. Thereafter, it was evaluated for a database of flat-bottomed features tested in X70 (L485) pipe steels.

The results of the full-scale testing indicated that accurate precise predictions were possible with this alternative metal-loss-severity criterion. But, consistent with the planning of that small matrix, it was also clear that the current state of development of this BF, whose focus has been on depth and length was inadequate regarding the effects of width. With that shortcoming recognized, the developments of this work were more broadly evaluated against the other popular Level 1 models including as noted earlier BS7910, DNV F101, LPC-1, and PCORRC. This independent assessment contrasted the predictive results of those now long-completed criteria to that achieved by the still developing model of this paper, which couples an alternative reference stress with a BF developed to quantify the response of narrow metal-loss features. Table 3. Results of C-FERs trending

As noted earlier and detailed in Table 3, predictions with the current developing model are comparable to that for the just noted established criteria. This predictive improvement has been achieved in spite of the uncertainty that enters due to the practical need for a surrogate for n affected through a correlation between n and Y/T. It follows in this context that the current developments are validated in conjunction with the n – Y/T correlation by full-scale testing, and a level of predictability that is comparable to other such criteria. The reduced bias and CoVs as compared to the corresponding predictions made using B31G and Modified B31G indicates that their use would open to reduced maintenance and related digs.

Finally, as discussion above regarding Figure 5 reveals, testing was done on part-thickness samples with a view to assess if such work provided a better approach to quantify the response in metal-loss features where failure initiates in material whose processing history differs from that on the pipe’s surface. It was clear from Figure 5 that the full-thickness sample failed on an inclined (shear) plane, whereas the part-thickness sample failed on a plane normal to the pipe’s surface. It was also evident that the part-thickness sample showed increased ductility through-thickness. As each of these differences opens to question the utility of nominal properties based on full-thickness sampling, these observations must be addressed, as follows.

Considering Equation 3, which defines the reference stress for this work, it is apparent that the UTS and n (or eu or Y/T as its surrogate) are the parameters that must be assessed in light of the differences evident in Figure 5. Of these, the UTS has a first-order influence, whereas n modifies it. While limited testing has been done, and the usual variability due to inherent differences in the steel sampled are evident in the flow response, the average full-thickness transverse UTS was 92.0 ksi (634 MPa), with Y/T = 0.77, eu = 0.0974, and n = 0.0747. The corresponding part-thickness testing the UTS at 91.5 ksi (630 MPa), with Y/T = 0.93, a eu = 0.0652, and n = 0.0584.

The observed differences in the values of the UTS for part versus full-wall sampling are well within the scatter typical of the underlying test practice. Likewise, the differences in the measured values of n are within scatter for replicate testing, as are those for eu. In contrast to these trends, the difference apparent in the values of Y/T seems large. In spite of that disparity, the predictive scatter for the 75 predictions that used Y/T as a surrogate for n remained comparable to the best of the other predictive models. In this context, the differences in the failure behavior evident in Figure 5 are not a factor. Based on this limited testing it follows that full-thickness sampling suitably quantifies the resistance to failure within metal-loss features. It should be noted in this regard that the ratio of AYS to SMYS varied significantly, with the highest result occurring in the part-thickness sample – which at least for this testing implies a latent margin of safety for models that are benchmarked to SMYS.

6. BENEFITS ILLUSTRATED: AN APPLICATION TO INTERNAL PITTING CORROSION


As the present paper reflects still developing technology its practical use in integrity management decisions is premature. However, there are details available from prior engineering critical assessments (ECAs) that if sanitized could be reassessed and used to illustrate the benefits that derive in light of the work completed under Projects EC-2-6 and EC-2-7. The scenario chosen involves an ECA to assess the serviceability of a condensate pipeline in service for a national petroleum company involving the PCORRC criteria. This PCORRC-based ECA was supported by select Level3 analyses, and laboratory properties testing and related vessel testing to validate the use of PCORRC for this application. Recall from the above discussion that the reference stress adopted in PCORRC is collapse-based’ and comparable in magnitude to that of Equation 3, including the dependence on n determined from the archived properties testing. As the subject ECA was done several years ago, the rehab and return to service decisions were based on PCORRC. In the event that this ECA indicated that the line could be economically salvaged, it was planned to clean and hydro-retest the line prior to its return to service.

Motivated by aspects of the operation of a critical condensate pipeline that had experienced upsets in dosing the inhibitor, the company had the line cleaned, and then subjected it to inspection that included running a high-resolution metal-loss (magnetic-flux-leakage) tool. That run was selectively validated by digs supported by nondestructive inspections (NDI) that included ultrasonic wall thickness measurements consistent with internal corrosion direct assessment practices. The results of the run indicated the presence of internal pitting where dropout occurred with sufficient residence time to promote corrosion. The aggressive nature of the transported product and the flow regime led to long strings of closely overlapped pitting, wherein the maximum relative depths were similar. Properties testing for sampling at convenient sites indicated the toughness was consistently high enough that failure initiating in these axial strings of pitting would be collapse controlled. Figure 7 plots the results of the post-millennial ILI run specifically for the deeper indications. That figure also presents the acceptance boundaries that corresponded to operation at the maximum allowable operating pressure (MAOP includes the surge allowance), for each of ASME B31G and the corresponding boundary for the current technology as is evident in Figure 7. Of importance for present purposes is that the still developing technology developed in EC-2-6 and EC-2-7 was for this application practically identical to PCORRC for the deeper features, and similar for the shallower features. This is anticipated in that PCORRC shares a reference stress comparable to Equation 3 and develops predictions that differ little from that equation as was evident in Table 3. As such, the predictive improvements evident in Tables 2 and 3 for the current state of this still developing technology are anticipated to parallel that achieved for PCORRC.

Figure 7: ILI indications and evaluation for a condensate pipeline

This figure shows 6 indications lie above the failure boundary for the current technology (and PCORRC), in contrast to the almost 70 features that lie above the B31G criterion. The viability of this decision was supported by the results of select Level 3 analysis, and the related full-scale testing. Features that lie above the PCORRC boundary were reinforced, as were the features that lay at or above d/t = 0.8. As the line was to be proof tested consistent with ASME B31 requirements, to address potential uncertainties in the processes involved the owner/operator also chose to reinforce features that lay on


or immediately below the PCORRC boundary. The line survived the proof test and was returned to service proven safe while also avoiding the digs and reinforcement associated with about 60 features. As the PCORRC boundary was practically identical to that developed using the still developing outcomes of Projects EC-2-6 and EC-2-7, the same benefit in reduced maintenance and rehab would derive from its use in this application.

7. SUMMARY AND CONCLUSIONS This paper has presented an alternative reference stress that has been adapted for use in EC-2-6. This alternative reference stress is a function of the UTS and the strain-hardening characteristics of the steel. It has been validated in reference to full-scale testing on pipe pups in grades that reflect the Grades used in early pipeline construction, through those that have emerged up to and beyond X80 (L482). As it is a function of the UTS and the strain-hardening characteristics of the steel, this alternative reference stress has the potential to seamlessly account for differences among the steels involved from historic through modern construction.

The reference stress has been coupled that with the BF of Project EC-2-7 that currently accounts for depth effects and length as occurs in narrow metal-loss features. The results were of predictions based on that model which were then used to design a full-scale test matrix with the objective of broadening the scope of the current test database to include longer features, and also to explore width effects – along with the usual consideration of defect depth. The results of that testing provided a basis to assess the utility of the just noted model in comparison to the corresponding predictions made using B31G and Modified B31G. Predictions made using B31G and Modified B31G lead to a quite large bias and scatter. The mean ratio of the predicted to actual failure pressure for B31G was 0.75, while that for Modified B31G was 0.81, with a CoV of 10.4% for B31G and 10.9% for Modified B31G. In contrast, the technology developed and discussed earlier led to a mean of 1.017, with a CoV of just 8%. The largest predictive error occurred for the only test that sought to quantify the effects of width on the current state of this model development. If that result was excluded, the average now trends conservative at 0.991, with a CoV that falls to just 4.7% -- which is typical of the scatter inherent in tensile testing [30]. It follows that the bias has been significantly reduced as compared to B31G and Modified B31G. At the same time the scatter has been decreased by the order of 20%.

The viability of predictions based on the reference stress coupled with the BF was also evaluated more broadly in terms of published full-scale metal-loss testing for flat-bottomed features machined into X70 (L485) pipe. This dataset was substantial with 75 such results located. That database covered failure pressures from 0.3 ≤ Pf (%SMYS) ≤ 1.5, normalized lengths (i.e., L / (D t)0.5) from ~0 to 14, widths (as arc angle) from ~0 to 75 degrees, and normalized depths (i.e., d/t) from near zero to approaching one. Predictions were also made using best known more recently developed Level 1 criteria for metal-loss, notably LPC-1, DNV F101, BS7910, and PCORRC. Based on an independent assessment of these results done by C-FER it was apparent that the quality of the predictions made by the combination of reference stress and the BF were comparable in terms of predictive bias and variability to the LPC-1, DNV F101, and PCORRC models. Taken together those predictions were somewhat better than those by BS7910, and major improvement benchmarked to B31G and Modified B31G (whose predictive outcomes were similar). In practical terms, reduced bias and scatter mean less conservatism is needed to achieve the same level of safety, such that rehab or repair that does not affect reduced risk can be minimized. This was demonstrated in regard to the integrity management of a condensate pipeline.

In view of the above it follows that the viability of the present approach has been demonstrated for flat-bottomed defects, with the potential for still further step gains in predictability available by the inclusion of width effects, which is now early in its development. Given this outcome work has been planned as part of EC-2-10 to continue this development.

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