Alternative Acceptance Criteria for Flaws in Ferritic ... Acceptance... · The flaw evaluation...

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ALTERNATIVE ACCEPTANCE CRITERIA FOR FLAWS IN FERRITIC STEEL COMPONENTS OPERATING IN THE UPPER SHELF TEMPERATURE RANGE

H. L. Gustin Structural Integrity Associates

Centennial, CO USA

R. C. Cipolla Intertek APTECH

Sunnyvale, CA USA

S. X. Xu Kinectrics

Toronto, Ontario, Canada

D. A. Scarth Kinectrics

Toronto, Ontario, Canada

ABSTRACT

The flaw evaluation rules for ferritic vessels in IWB-3610, IWB-3620 and Appendix A of ASME Section XI are based on linear elastic fracture mechanics techniques and were developed primarily for the irradiated reactor pressure vessel beltline region and other low temperature carbon and low-alloy steel applications in which the material exhibits limited or no ductility prior to failure. There are situations in which ferritic steel components operate in the upper shelf temperature range and therefore exhibit significant ductility and increased flaw tolerance. Application of linear elastic fracture mechanics techniques to these cases can be very conservative. In order to address flaw evaluation of ferritic materials exhibiting upper shelf toughness and high ductility, the proposed Code Case N-749 of ASME Section XI was developed and is currently under committee review. This proposed Code Case provides alternate acceptance criteria for situations in which the component is operating in the upper shelf temperature range and therefore has adequate ductility to allow the use of elastic-plastic fracture mechanics techniques.

INTRODUCTION

The evaluation rules for flaws detected in ferritic components are specified in IWB-3610 and IWB-3620 of ASME Section XI [1]. The flaw evaluation procedures are given in Appendix A of Section XI and are based on Linear Elastic Fracture Mechanics (LEFM). Plasticity effects are accounted for through the use of a plastic zone correction

factor, which is included in the direct calculation of the crack-tip stress intensity factor, KI. These procedures were first introduced into Section XI in the early 1970s and have been used to evaluate and accept flaws in the different classes of components. The flaw acceptance criteria of IWB-3610 and IWB-3620 of Section XI require structural factors to be applied to the stresses acting on the flaw. These stresses include primary and secondary stresses including residual stress, and as such, the full magnitude of these structural factors is applied equally to pressure, thermal and residual stresses in the determination of KI. In some applications, where the thermal and/or residual stresses are significant, unacceptable results are obtained in the assessment for allowable flaw size. This can be overly conservative, in particular when the component is operating in the upper shelf temperature range where the material is ductile and cleavage fracture is not a failure mode. This paper describes the development of proposed Code Case N-749 [2] as an alternative to the LEFM-based flaw acceptance criteria currently given in IWB-3610 and IWB-3620 of Section XI. The development of acceptance criteria for proposed Code Case N-749 follows the rationale of Appendix K of Section XI, which is also based on Elastic-Plastic Fracture Mechanics (EPFM) [3,4,5] applied to upper shelf fracture prevention. Proposed Code Case N-749 provides revised acceptance criteria for use in a J-integral based evaluation of ferritic components operating in the upper shelf temperature range. Following the discussion of the acceptance criteria in the proposed Code Case N-749, the basis for the treatment of weld residual stresses in

Proceedings of the ASME 2012 Pressure Vessels & Piping Conference PVP2012

July 15-19, 2012, Toronto, Ontario, CANADA

PVP2012-78190


these criteria is presented. Application of the acceptance criteria and evaluation procedure to an illustrative example problem is then presented.

NOMENCLATURE

a Flaw depth a0 Initial flaw depth af Final flaw depth at the end of the evaluation period ∆a Amount of ductile crack growth c, m Material constants for a power law fit to the J-integral

resistance curve E Young’s modulus Gj Influence coefficients for the stress intensity factor J J integral J-R J-integral resistance curve for the material J0.1 J-integral toughness for 0.1 inch (2.5 mm) crack

extension Japplied Value of J-integral due to applied loading KI Crack-tip stress intensity factor for Mode I KIe Linear elastic crack-tip stress intensity factor for Mode

I without a plastic-zone correction KIeff Effective crack-tip stress intensity factor for Mode I

with a plastic-zone correction ℓ Flaw length ℓ0 Initial flaw length ℓf Final flaw length at the end of the evaluation period PACC Accumulation pressure, = 1.1 times Design pressure PD Design pressure, = 1.1 times operating pressure PO Operating pressure Q flaw shape factor rye Radius of plastic zone at flaw tip Ri Inside radius of vessel Ro Outside radius of vessel RTNDT reference nil-ductility temperature SF Structural factor on load Sm Design stress intensity t Vessel wall thickness tnom Nominal vessel wall thickness Tc Transition temperature for Charpy energy curve σy Material yield strength σu Material ultimate tensile strength ν Poisson’s ratio

TECHNICAL APPROACH Basis for Use of Elastic-Plastic Fracture Mechanics In the upper-shelf temperature range, materials will exhibit ductile flaw extension prior to flaw instability. It is therefore appropriate to use EPFM analysis methods when evaluating a flaw in a material at upper-shelf temperatures such as normal operating conditions for both PWRs and BWRs. A number of precedents exist in ASME Section XI for the application of

EPFM to materials that exhibit sufficient ductility [1]. Such precedents may be found in Appendix C of Section XI for evaluation of flaws in austenitic and ferritic piping, and in Appendix K of Section XI for assessment of reactor pressure vessels with low upper shelf fracture toughness. Appendix C includes a screening criterion to determine which failure mode a ferritic piping flaw evaluation must evaluate (LEFM, EPFM or Limit Load). For the EPFM failure mode, Appendix C specifies full structural factors on primary stresses, and structural factors of 1.0 on secondary stresses. Another approach is presented in Appendix K of Section XI. In addition to identifying different structural factors for primary versus secondary loadings, Appendix K also provides procedures for performing flaw instability analysis for flaws in reactor pressure vessel materials operating in the upper shelf temperature range. The EPFM approach in the proposed Code Case N-749 is very similar to that in Appendix K, and in that sense the use of the EPFM techniques in the proposed Code Case is not unprecedented. The difference is that the EPFM techniques in the proposed Code Case N-749 are applied to actual flaws, rather than hypothetical flaws as in Appendix K. Definition of Upper Shelf Temperature For the proposed Code Case N-749 to be applicable, it must be demonstrated that the material is operating in the upper shelf range of its Charpy energy curve. The proposed Code Case requires that the operating temperature must exceed the upper shelf transition temperature, Tc, where Tc is defined as

F105RTT NDTc+= (U.S Customary Units)

C3.58RTT NDTc+= (SI Units) (1)

where RTNDT is the reference nil-ductility temperature as defined in A-4000 of Appendix A of ASME Section XI. Using the relationship between lower bound KIc and temperature in A-4200(b) of Appendix A, the upper shelf transition temperature in eqn. (1) corresponds to the fracture toughness KIc reaching upper-shelf conditions as shown in Figure A-4200-1 of Appendix A of Section XI, for an upper shelf toughness of 200 ksi-√in. This value is considered to be the reasonable upper limit on achievable toughness KIc for the ferritic materials governed by Appendix A. If materials with lower upper shelf toughness were evaluated using the same A-4200 (b) relationships, the resulting upper shelf transition temperature T-RTNDT would be lower than the limiting 105 ◦F (58.3 ◦ C). Consequently, using the above definition of Tc (equation (1)) is conservative for ferritic materials with upper shelf toughness KIc less than or equal to 200 ksi-√in. This definition of Tc ensures that the material exhibits sufficient ductility in thick sections to allow the use of EPFM techniques.


Flaw Characterization The flaw geometry is characterized in accordance with Appendix A, Article A-2000 of ASME Section XI to determine current flaw dimensions. The method of flaw characterization is the same as for LEFM evaluation of the flaw. Loads and Stresses All primary stresses from internal pressure and mechanical loads, secondary stresses such as from thermal loads, and peak stresses including highly localized stresses, are to be considered in applying proposed Code Case N-749. This is consistent with the present procedure for flaw evaluation in vessels in IWB-3600 and Appendix A of Section XI. Weld residual stresses are not required to be included. The rational for not including residual stresses is provided later in this paper. Flaw Growth Flaw growth is evaluated in accordance with Appendix A, Article A-5200 of ASME Section XI to determine the end-of-evaluation-period flaw depth af and flaw length ℓf. The method of evaluating flaw growth is the same as for LEFM evaluation of the flaw. Flaw Acceptance Two alternate sets of acceptance criteria are in proposed Code Case N-749. The first set of criteria is based solely on limited ductile crack extension (initiation). These criteria do not credit stable ductile tearing and therefore are conservative. The criteria however offer simplicity in the evaluation process for cases where the material is relatively tough or the applied loads are relatively small. The second set of criteria is based on both limited ductile crack extension and flaw stability in which ductile stable tearing is evaluated. The acceptance criteria ensure that crack extension is limited even for a stable flaw.

FLAW ACCEPTANCE CRITERIA The flaw size at the end of the evaluation period (af, ℓf), is acceptable for continued operation if the applied J-integral, J, satisfies the first set of acceptance criteria for initiation of ductile crack extension. If the first set of criteria is not satisfied, then the second set of acceptance criteria based on initiation of ductile crack extension and flaw stability must be met.

Accep tance Crite ria Bas ed So le ly on Limited Ductile Crack Extens ion The purpose of these acceptance criteria is to limit stable ductile crack extension. Hence, the applied J-integral is compared to the material J-R curve at a crack extension of 0.10 in. (a) For normal and upset conditions (Service Levels A and B), the applied J-integral is evaluated at loads equal to 2.0 times the primary loads and 1.0 times the secondary loads, including thermal stresses. The applied J must be less than or equal to the J-integral of the material at a ductile crack extension of 0.10 in. (2.5 mm). (b) For emergency and faulted conditions (Service Levels C and D), the applied J-integral is evaluated at loads equal to 1.5 times the primary loads and 1.0 times the secondary loads, including thermal. The applied J must be less than or equal to the J-integral of the material at a ductile crack extension of 0.10 in. (2.5 mm). (c) In accordance with IWB-3610 of Section XI, the primary stress limits of NB-3000 of Section III must be satisfied, assuming a local area reduction of the pressure retaining membrane that is equal to the area of the end-of-evaluation-period flaw(s) as determined by the flaw characterization rules of IWA-3000 of Section XI. Accep tance Crite ria Bas ed on Limited Ductile Crack Extens ion and Stab ility The purpose of these acceptance criteria is to allow for stable ductile flaw extension, and to ensure flaw stability. (a) For normal and upset conditions (Service Levels A and B), the following requirements must be satisfied. (1) For ductile crack extension, the applied J-integral is

evaluated at loads equal to 1.5 times the primary loads and 1.0 times the secondary loads, including thermal stresses. The applied J must be less than or equal to the J-integral of the material at a ductile flaw extension of 0.10 in. (2.5 mm).

(2) The flaw must be stable at loads equal to 2.14 times the

primary loads and 1.0 times the secondary loads, including thermal stresses. Flaw stability is evaluated as described below in this paper.

(3) The primary stress limits of NB-3000 of Section III

must be satisfied after stable ductile crack extension.


(b) For emergency and faulted conditions (Service Levels C and C), the following requirements must be satisfied. (1) For ductile crack extension, the applied J-integral is

evaluated at loads equal to 1.25 times the primary loads and 1.0 times the secondary loads, including thermal stresses. The applied J must be less than or equal to the J-integral of the material at a ductile flaw extension of 0.10 in. (2.5 mm).

(2) The flaw must be stable at loads equal to 1.2 times the

primary loads and 1.0 times the secondary loads, including thermal stresses.

(3) For faulted conditions, the total end-of-evaluation-

period flaw depth, including stable ductile crack extension, must be less than or equal to 75% of the wall thickness, and the remaining ligament is not subject to tensile instability. For ligament tensile stability, the internal pressure must be less than the internal pressure to cause tensile instability of the remaining ligament. For the case of the reactor pressure vessel, the methods given in Appendix K of Section XI; K-5300(b)(1) for axial flaws, and K-5300(b)(2) for circumferential flaws, may be used.

For emergency and faulted conditions, the ligament tensile stability evaluation for faulted conditions, including stable ductile crack extension, is required in place of the primary stress limits of NB-3000 of Section III. The ligament tensile stability evaluation for faulted conditions is consistent with the requirements of Appendix K of Section XI.

TECHNICAL BASIS FOR STRUCTURAL FACTORS ON LOAD The LEFM methodology of IWB-3610, IWB-3620 and Appendix A of Section XI treats all loadings on the component equivalently, applying equal structural factors (√10 for normal and upset conditions) to both primary stresses due to internal pressure and mechanical loads, as well as to secondary and peak stresses such as those caused by differential thermal expansion and residual stresses. These loadings are equivalent in their potential to produce fracture in materials only where LEFM is the failure mode, such as reactor pressure vessel beltline materials at low temperatures after significant irradiation embrittlement, and thick, ferritic materials at very low temperatures. However, for ferritic materials in the upper-shelf temperature range, cleavage fracture will not occur, and LEFM is not an appropriate evaluation method. In the EPFM evaluation in Appendix K of Section XI, structural factors on primary loads and on secondary loads are applied as shown in Table 1. The technical basis for the structural factors in Appendix K is provided in Reference [6]. However it must be recognized that Appendix K is not used to evaluate actual

flaws, but rather to demonstrate adequate levels of upper-shelf fracture toughness based on very large hypothetical flaws. Proposed Code Case N-749 addresses evaluation of actual flaws in vessels and other components. Therefore, more conservative structural factors, paralleling those in Appendix C of Section XI, are deemed appropriate. Different structural factors are proposed for limited ductile crack extension and flaw stability. Because in an EPFM evaluation failure is predicted at flaw instability, higher structural factors are applied for this condition while lower structural factors are applied when evaluating limited ductile crack extension. A comparison is made between the structural factors in Appendix K of Section XI with those in proposed Code Case N-749. Specifically, the structural factors for the proposed Code Case were derived from the Appendix K values in order to maintain equivalency in philosophy for treating flaws in ferritic material at upper-shelf temperatures. This comparison is given in Table 2. Because of differences among the various operating conditions, namely normal operation, upset transients, and postulated accidents, different criteria have been developed for normal and upset conditions, and for emergency and faulted condition. This approach is consistent with other evaluation criteria in ASME Section III and Section XI. The technical basis for the structural factors for limited ductile crack extension and flaw stability is described below for normal and upset conditions, and for emergency and faulted conditions. Normal and Upset Conditions

Struc tura l Fac tors fo r Accep tance Crite ria Bas ed So le ly on Limited Ductile Crack Extens ion

For the acceptance criteria based solely on limited ductile crack extension, structural factors of 2.0 on primary loads and 1.0 on secondary loads, including peak stresses due to local stress concentrations, but excluding residual stresses, are required. The structural factor on primary loads is more conservative than the structural factor on primary loads for Level A and B Service Loads in Appendix K of Section XI of 1.15. The structural factor of 1.0 on secondary loads is consistent with that specified for secondary loads in Appendices C and K of Section XI. Since Appendix K is applied to evaluate hypothetical flaws, and the proposed Code Case N-749 is intended for application to evaluate actual flaws detected during inspection, it is appropriate to require higher structural factors in the proposed Code Case. With reference to Tables 1 and 2, the structural factor of 2.0 on primary loads was determined as follows. The structural factor of 1.15 for Level A and B Service Loads from Appendix K of Section XI was multiplied by 1.1 to adjust from accumulation pressure to Design pressure, and by 1.1 again to adjust from


design pressure to operating pressure. An additional factor of √2 on load was applied to account for the factor of 2.0 on the postulated flaw size in Appendix K for normal and upset conditions (see Table 2). The total structural factor is then 1.97, which was rounded to 2.0. Additional margins to account for inspection sizing uncertainty are not required, since such margins are typically included in the qualification of inspection processes. Struc tura l Fac tors fo r Accep tance Crite ria Bas ed

on Limited Ductile Crack Extens ion and Stab ility

For the acceptance criteria based on limited ductile crack extension and flaw stability, structural factors of 1.5 on primary loads and 1.0 on secondary loads, including peak stresses due to local stress concentrations, but excluding residual stresses, are required for limited ductile crack extension. These structural factors are based on committee judgment and provide a reasonable margin on crack initiation in addition to providing a conservative margin against tearing instability. The structural factor of 1.15 for Level A and B Service Loads from Appendix K of Section XI, multiplied by 1.1 to adjust from accumulation pressure to Design pressure, and by 1.1 again to adjust from design pressure to operating pressure, is 1.39. A structural factor of 1.5 on primary loads in proposed Code Case N-749 bounds the Appendix K factor on primary load and forms the basis of the committee selection. The lower structural factor of 1.5 for limited ductile crack extension compared to 2.0 for the acceptance criteria based solely on limited ductile crack extension is acceptable since flaw stability is also evaluated. The structural factors for ensuring flaw stability are 2.14 on primary loads and 1.0 on secondary loads, and are to be used in the assessment of ductile crack extension to ensure that crack extension is stable. Since stable crack extension forms the main criterion for acceptance, the factor of 2.14 was derived in the same manner as before using the structural factor on primary load for flaw stability in Appendix K of 1.25 for Level A and B Service Loads, multiplied by 1.1 to adjust from accumulation pressure to Design pressure, and by 1.1 again to adjust from design pressure to operating pressure. An additional factor of √2 on load was applied to account for the factor of 2.0 on the postulated flaw size in Appendix K. The total structural factor is then 2.14.

Normal and Upset Conditions Structural Factors for Acceptance Criteria Based Solely on Limited Ductile Crack Extension

For the acceptance criteria based solely on limited ductile crack extension, structural factors of 2.0 on primary loads and 1.0 on secondary loads, including peak stresses due to local

stress concentrations, but excluding residual stresses, are required. The structural factor on primary loads is more conservative than the structural factor on primary loads for Level A and B Service Loads in Appendix K of Section XI of 1.15. The structural factor of 1.0 on secondary loads is consistent with that specified for secondary loads in Appendices C and K of Section XI. Since Appendix K is applied to evaluate hypothetical flaws, and the proposed Code Case N-749 is intended for application to evaluate actual flaws detected during inspection, it is appropriate to require higher structural factors in the proposed Code Case. With reference to Tables 1 and 2, the structural factor of 2.0 on primary loads was determined as follows. The structural factor of 1.15 for Level A and B Service Loads from Appendix K of Section XI was multiplied by 1.1 to adjust from accumulation pressure to Design pressure, and by 1.1 again to adjust from design pressure to operating pressure. An additional factor of √2 on load was applied to account for the factor of 2.0 on the postulated flaw size in Appendix K for normal and upset conditions (see Table 2). The total structural factor is then 1.97, which was rounded to 2.0. Additional margins to account for inspection sizing uncertainty are not required, since such margins are typically included in the qualification of inspection processes.

Struc tura l Fac tors fo r Accep tance Crite ria Bas ed on Limited Ductile Crack Extens ion and Stab ility

For the acceptance criteria based on limited ductile

crack extension and flaw stability, structural factors of 1.5 on primary loads and 1.0 on secondary loads, including peak stresses due to local stress concentrations, but excluding residual stresses, are required for limited ductile crack extension. These structural factors are based on committee judgment and provide a reasonable margin on crack initiation in addition to providing a conservative margin against tearing instability. The structural factor of 1.15 for Level A and B Service Loads from Appendix K of Section XI, multiplied by 1.1 to adjust from accumulation pressure to Design pressure, and by 1.1 again to adjust from design pressure to operating pressure, is 1.39. A structural factor of 1.5 on primary loads in proposed Code Case N-749 bounds the Appendix K factor on primary load and forms the basis of the committee selection. The lower structural factor of 1.5 for limited ductile crack extension compared to 2.0 for the acceptance criteria based solely on limited ductile crack extension is acceptable since flaw stability is also evaluated. The structural factors for ensuring flaw stability are 2.14 on primary loads and 1.0 on secondary loads, and are to be used in the assessment of ductile crack extension to ensure that crack extension is stable. Since stable crack extension forms the main criterion for acceptance, the factor of 2.14 was derived in the same manner as before using the structural factor on primary


load for flaw stability in Appendix K of 1.25 for Level A and B Service Loads, multiplied by 1.1 to adjust from accumulation pressure to Design pressure, and by 1.1 again to adjust from design pressure to operating pressure. An additional factor of √2 on load was applied to account for the factor of 2.0 on the postulated flaw size in Appendix K. The total structural factor is then 2.14.

Emergency and Faulted Conditions

Structural Factors for Acceptance Criteria Based Solely on Limited Ductile Crack Extension

For the acceptance criteria based solely on limited ductile crack extension, structural factors of 1.5 on primary loads and 1.0 on secondary and peak loads are required. The factor of 1.5 was based on committee judgment and is a conservative bound to the structural factors for Level C and D Service Loads of 1.0 on primary loads and 1.0 on secondary loads in Appendix K. As an independent verification, the requirements of Appendix F of ASME Section III [7] were reviewed to establish the inherent margins provided in the assessment of net-section plastic collapse under faulted loads. Appendix F Paragraphs F-1331.1(c)2 and F-1341.3 define the allowable stress for the analysis of collapse as 90% of the calculated collapse load where the critical net-section stress is not to exceed the lesser of 2.3Sm or 0.7Su, where Sm is the design stress intensity and Su is the ultimate tensile strength. Therefore, an analysis could use this allowable stress and meet the requirements of the ASME Code. The structural factor implied from this requirement is 1.11. The proposed structural factor of 1.5 bounds this value.

Structural Factors for Acceptance Criteria Based on Limited Ductile Crack Extension and Stability

For the acceptance criteria based on limited ductile crack extension and flaw stability, structural factors of 1.25 on primary loads and 1.0 on secondary loads, including peak stresses due to local stress concentrations, but excluding residual stresses, are required for limited ductile crack extension. The structural factors for ensuring flaw stability are 1.2 on primary loads and 1.0 on secondary loads, and are to be used in the assessment of ductile crack extension to ensure that crack extension is stable. The structural factors were based on committee judgment and are a conservative bound to the structural factors for Level C and D Service Loads of 1.0 on primary loads and 1.0 on secondary loads in Appendix K, as well as the implicit margins in Appendix F of Section III for faulted loads.

RATIONALE FOR NOT INCLUDING RESIDUAL STRESSES IN ACCEPTANCE CRITERIA

The technical basis for not including residual stresses in flaw evaluations using proposed Code Case N-749 consists of two aspects: findings from a literature review of residual stresses in vessel integrity considerations, and reported fracture tests involving residual stresses. Lidbury [8] and Merkle [9] provided comprehensive reviews of the subject of residual stresses in the context of nuclear pressure vessel integrity. Residual stresses in as-welded structures may be up to yield strength magnitude in tension, but they generally relax to 10 to 25% of the yield strength magnitude with post-weld heat treatment (PWHT). Even assuming a high level of initial residual stress, the expectation from the literature review is that the residual stress level only has significant effect on crack initiation, and has no significant effect on the vessel failure load providing upper-shelf behavior is maintained. The flaw sizes that can be safely tolerated are significantly larger than the corresponding initiation flaw sizes, by virtue of the effective increase in toughness occurring with small amounts of stable, ductile crack extension. Provided cleavage fracture is not a possibility, failure is ultimately dominated by mechanical loading as the plastic collapse load is approached. Proposed Code Case N-749 has a specific requirement on the evaluation temperature to ensure fully ductile fracture behavior. Not including residual stresses in proposed Code Case N-749 fracture assessments is consistent with observations from the literature review. Fracture tests involving residual stresses reported in the literature also support not including residual stresses in a proposed Code Case N-749 flaw evaluation. Three tests in the Heavy-Section Steel Technology (HSST) program pertained to the effect of residual stresses on fracture behavior of deliberately flawed experimental vessels containing half-bead, un-stress-relieved, repair welds [9]. The intermediate pressure vessels had outside diameters of 991 mm (39 in.) and wall thicknesses of 152 mm (6 in.). The flaw in vessel V-7A was remote from the repair weld so that the performance of the repair weld could be observed without the influence of a flaw. The flaws in vessels V-7B and V-8 were located within, or close to, the repair welds so that the flaws were subject to the associated residual stress fields. Vessels V-7A and V-7B were tested at an upper-shelf temperature of 87.8°C (190°F). Failure in vessel V-7A was by leakage only, following stable ductile tearing of the fully yielded remaining ligament beneath the deepest portion of the flaw. The repair weld in vessel V-7A was monitored acoustically, but emitted no signals, nor did it show any other signs of damage during the test. The flaw in vessel V-7B, which was geometrically identical to the flaw in vessel V-7A, was located in the fusion zone of the un-stress-relieved repair weld. Failure in vessel V-7B was also by leakage only, and the remaining ligament beneath the flaw in vessel V-7B was fully yielded. The leak pressure was nearly identical to that for


vessel V-7A, suggesting that the residual stresses from the repair weld had no effect on the leak pressure. The test results for vessels V-7A and V-7B demonstrated that residual stresses have no effect on the integrity of the un-stress-relieved repair welds provided the vessel operation is under upper-shelf conditions. The flaw in vessel V-8 was located adjacent to an un-stress-relieved repair weld, at the location of maximum tensile residual stress, and the vessel was tested to fracture at a temperature of -23°C (-9.4°F), which was the RTNDT of the fabrication weld material in which the flaw was placed. At this temperature the residual stresses were expected to, and did have a significant effect on the fracture conditions for the vessel. Vessel V-8 fracture was accurately predicted by linear elastic fracture mechanics with full account of the residual stresses. The test results for vessel V-8 demonstrated that residual stresses have significant effect on the integrity of the un-stress-relieved repair welds when the vessel operation is under lower-shelf conditions. Formby and Griffiths investigated the effects of residual stresses, induced by the application of heating pads, on the fracture load of centre-cracked panels of BS 1501-224 steel [10]. The experimental results of Formby and Griffiths were reviewed in Reference [8]. The results are illustrated in Figure 1, which was taken from Reference [8], which showed the temperature dependence of failure with regard to residual stresses. For temperatures below -180°C, fracture occurred by a brittle mechanism, with residual stresses having a significant influence on the failure load. Above -180°C, ductile mechanisms became increasingly important, with specimens failing by plastic collapse above 0°C. Within the ductile range, the failure load was essentially independent of whether residual stresses were present.

METHOD FOR EVALUATION OF J-INTEGRAL AND FLAW STABILITY

The methods for evaluation of the J-integral and flaw stability were taken from Appendix K of Section XI. Details of the technical basis are provided in References [4,5,6]. Evalua tion o f J -In tegra l Calculation of the J-integral due to applied loads must account for elastic-plastic behavior of the stress-strain curve for the material. When linear elastic fracture mechanics with small scale yielding applies, the J-integral may be calculated using crack-tip stress intensity factor formulae with a plastic-zone correction. Since the plastic-zone size is intended to be representative of the extent of plasticity under actual loading conditions, the plastic-zone correction is calculated prior to application of any structural factors on applied loads. Further

details are provided in the example flaw evaluation given below. Evaluation of Flaw Stability A flaw under ductile crack extension is stable when both (a) and (b) are satisfied. (a) The equilibrium requirement for stable flaw extension is

RJJ = (2)

Where J is the J-integral due to applied loads for the flaw, and JR is the J integral resistance to ductile tearing for the material. (b) The requirement for flaw stability under ductile tearing is

RR JJat

dadJ

aJ

=<∂∂

(3)

where ∂J/∂a is the partial derivative of the applied J-integral with respect to flaw depth, a, with constant load, and dJR/da is the slope of the J-R curve. Under increasing load, stable flaw extension will continue as long as ∂J/∂a remains less than dJR/da. J -In tegra l Materia l Res is tance Curve The use of EPFM for flaw evaluation requires adequate characterization of the J-integral resistance curve for the vessel material. Appendix K of Section XI specifies three methods for selection of the material J-integral resistance curve. (a) A J-R curve may be generated for the material by following accepted test procedures. The J-R curve must be based on the proper combination of flaw orientation, temperature, and fluence level. Crack extension must be ductile tearing with no cleavage. (b) A J-R curve may be generated from a J-integral database obtained from the same class of material with the same orientation using correlations for effects of temperature, chemical composition, and fluence level. Crack extension must be ductile tearing with no cleavage. (c) As an alternative to (a) or (b), an indirect method of estimating the J-R curve may be used, provided the method is justified for the material. Proposed Code Case N-749 permits the same three methods for determining the J-R curve as in Appendix K.

EXAMPLE FLAW EVALUATION


An example flaw evaluation using the proposed Code Case N-749 is presented is this section to illustrate the application of acceptance criteria in the proposed Code Case. Heavy-Section Steel Technology (HSST) Intermediate Vessel Test (ITV) V-8A was analyzed. Test vessel V-8A was one of the tests in the HSST program to study the behavior of flaws under stress states similar to those in full-scale reactor pressure vessels. One particular objective of V-8A test was to demonstrate the fracture behavior of low-toughness material at upper-shelf temperatures [11]. HSST Vessel V-8A Test and Results Details of the vessel V-8A test and results are provided in NUREG CR-4760 [11]. Test data, material properties and other necessary information used in the flaw evaluation are summarized below. Test vessel V-8A was fabricated from a cylindrical test section of ASTM A533 Grade B Class 1 steel plate. The inner radius of the vessel, Ri = 343 mm (13.5 in.), the outer radius, Ro = 495 mm (19.5 in.), and the nominal wall thickness, tnom = 152 mm (6 in.). The vessel had previously been tested as V-8, and the V-8 flaw was removed prior to the V-8A test. A special low upper-shelf toughness weld seam was placed at another location in the vessel. A large external surface flaw with approximately a semi-elliptical profile was placed in the center of the special seam weld in the radial-axial plane, as illustrated in Figure 2. Postweld heat treatment was applied. The flaw dimensions were a flaw depth, a0 = 87.8 mm (3.46 in.), and a flaw length, ℓ0 = 280 mm (11.02 in.). The wall thickness at the flaw location, t = 141.3 mm (5.56 in.). The mechanical properties of the weld metal at 149°C (300°F) were an elastic modulus, E = 200 GPa (29,000 ksi), Poisson's ratio, ν = 0.3, yield strength, σy = 423 MPa (61.4 ksi), and ultimate tensile strength, σu = 525 MPa (76.1 ksi). The material J-R curve of the V-8A special seam weld was determined from JR tests at 149°C (300°F) and fitted to the power-law expression in eqn. (4). The average JR curve for the vessel seam weld of ten specimens was obtained as c = 81.7 and m = 0.318 with JR in kJ/m2 and ∆a in mm (c = 1305 and m = 0.318 with JR in in.-lb/in.2 and ∆a is in.).

( )mR acJ ∆= (4) The vessel was pressurized at temperatures in the range between 146 and 150°C (295 and 302°F). The temperature near the flaw location was between 146 and 158°C (295 and 316°F). The test temperature used in the analysis was 149°C (300°F). Two flaw instability occurrences were observed during the test. The first instability was observed between 138.0 and 140.5 MPa (20.0 and 20.4 ksi). The second instability was observed between 142.6 and 143.0 MPa (20.7

ksi). Once the onset of unstable tearing was confirmed, the pressure was rapidly decreased to prevent a burst. The final flaw depth was af = 101.4 mm (3.99 in.) based on post-test measurements. The ductile crack extension at onset of flaw instability was ∆a = 13.6 mm (0.53 in.). Applicability of Proposed Code Case N-749 To apply the EPFM acceptance criteria in proposed Code Case N-749, the operating temperature must exceed the upper-shelf transition temperature, Tc. The temperature RTNDT for the weld metal is -23°C (-9.4°F) [11]. From eqn. (1), the corresponding Tc is 35.3°C (95.6°F). The test temperature of 149°C (300°F) greatly exceeds the transition temperature Tc. The applicability of the acceptance criteria in proposed Code Case N-749 for test vessel V-8A is confirmed. Applied Load and J-integral Calculation The example calculation was performed for a Level A Service Loading condition. The applied load in the test was internal pressure only. The pressure level used in a proposed Code Case N-749 evaluation for Level A Service Loads is the normal operating pressure. The equivalent Level A normal operating pressure for test vessel V-8A was estimated from the design equation for the vessel nominal wall thickness. Design pressure for test vessel V-8A, PD, was estimated using the equation given below from NB-3324.1 of ASME Section III.

Dm

iDnom P5.0S

RPt−

= (5)

where PD is design pressure, Ri is inner radius, and Sm is design stress intensity. U.S. Customary Units were used in the example calculations. Corresponding values for key results are also provided in SI Units. From ASME Section II, Part D [12], Sm = 26.7 ksi for SA 533 Grade B Class 1 at the test temperature of 300°F. Design pressure PD was calculated as 9.7 ksi using Ri = 13.5 in. and tnom = 6 in. The normal operating pressure, PO, was estimated as 8.8 ksi (60.7 MPa) after applying a dividing factor of 1.1 to the Design pressure. The J-integral due to internal pressure was calculated by using the stress intensity factor with the plastic zone correction for small-scale yielding. The size of the plastic-zone was checked to confirm the small-scale yielding condition. The linear elastic stress intensity factor for the V-8A flaw, KIe, was calculated using the equation given below for an external semi-elliptical surface crack in an internally pressurized cylinder [13]. A constant value of flaw aspect ratio a/ℓ (= a0/ℓ0), was used in the calculation.


2/1

3

3

o2

2

o1

o02

i2o

2io

Ie QaG

Ra4G

Ra3G

Ra2G2

RRRPK

+

+

+

−=

π (6)

where

65.1a593.41Q

+=

(7)

and Ro is outer radius, Q is the flaw shape factor, and Gj are influence coefficients. Tabular values of Gj in Reference [14] were used with consideration of validity limits for Ri/t. The radius of the plastic zone, rye, and the effective stress intensity factor, KIeff, were calculated using eqns. (8) and (9), respectively. The applied J was then calculated using eqns. (10) and (11).

2

y

Ieye

K61r

=

σπ (8)

2/1

yeIeIeff a

r1KK

+= (9)

'EK

J2Ieff= (10)

21E'Eν−

= (11)

In accordance with the proposed Code Case N-749, the plastic-zone correction was calculated prior to application of any structural factors. Weld residual stresses were not included in the J calculation as permitted by the proposed Code Case. Evalua tion fo r Crack Extens ion of 0.1 in . In accordance with the acceptance criteria based solely on limited ductile crack extension in the proposed Code Case N-749, the applied pressure was set equal to 2.0 times the operating pressure to calculate the applied J. The applied J was calculated as Japplied = 653 in-lb/in.2 (114 kJ/m2) for the pressure load of 2.0 times 8.8 ksi, including a plastic-zone correction at the pressure load of 8.8 ksi. The plastic-zone radius was estimated as rye = 0.072 in. based on eqn. (8). Since rye/a = 0.02 and rye/(t-a) = 0.03, the small scale yielding condition is confirmed. The J-integral of the material at a ductile crack extension of 0.10 in. was obtained as J0.1 of 628 in-lb/in.2 (110 kJ/m2). Since Japplied of 653 in-lb/in.2 (114 kJ/m2) is larger than

J0.1 = 628 in-lb/in.2 (110 kJ/m2), the acceptance criteria based solely on limited ductile crack extension are not satisfied. Results are summarized in Table 3. Evalua tion fo r Crack Extens ion of 0.1 in . and Stab ility Alternate acceptance criteria in the proposed Code Case N-749 allow stable ductile crack extension. For the Level A Service Loading condition, the applied pressure is set equal to 1.5 times the operating pressure to evaluate J and compare with J0.1. In addition, flaw stability must be confirmed for an applied pressure equal to 2.14 times the operating pressure. At an applied pressure of 1.5 times 8.8 ksi, the applied J was calculated as Japplied = 367 in.-lb/in.2 (64 kJ/m2), with the plastic-zone correction calculated using the pressure of 8.8 ksi. The Japplied of 367 in.-lb/in.2 (64 kJ/m2) is less than J0.1 of 628 in-lb/in.2 (110 kJ/m2). At an applied pressure loading of 2.14 times 8.8 ksi, flaw stability was evaluated using J-R curve – crack driving force diagram approach [5,6]. The applied J was calculated for a series of flaw depths corresponding to increasing amounts of ductile crack extension, ∆a, ranging up to 0.8 in. The applied J was plotted against ductile flaw extension, ∆a, on the crack driving force diagram to produce the applied J curve as shown in Figure 3. The material J-R curve was also plotted on the crack driving force diagram. The slope of the applied curve is less than the slope of the material curve at the point where the two curves intersect. The flaw is therefore stable at 2.14 times operating pressure. The acceptance criteria based on ductile crack extension and flaw stability are satisfied. Results are summarized in Table 3. Comparis on of Pred ic ted Pres s ure Load and Ductile Crack Extens ion a t Ons e t o f Flaw Ins tab ility with Experiment In the crack driving force diagram, the onset of flaw instability is given by tangency of the applied and material resistance J-integral versus flaw extension curves. From Figure 3, the pressure corresponding to tangency between the two curves was determined to be 2.19 times operating pressure. The applied pressure at 2.19 times operating pressure, or 19.3 ksi (133 MPa) is the predicted pressure at onset of flaw instability. The corresponding applied J is 1073 in.-lb/in.2 (188 kJ/m2), and the predicted ductile crack extension is 0.54 in. (13.7 mm). The predicted pressure loading of 19.3 ksi (133 MPa) and ductile crack extension of 0.54 in. (13.7 mm) at the onset of flaw instability are in good agreement with the measured pressure loading of 20.0 ksi (138 MPa) and ductile crack extension of 0.53 in. (13.6 mm) at the onset of flaw instability. Results are summarized in Table 4. The above calculations are approximate as the calculations are based on linear elastic fracture mechanics with a plastic-zone


correction. The intent of the comparison is to demonstrate the consistency of proposed Code Case N-749 calculations and V-8A test data. A more detailed comparison can be made using finite element modeling with actual stress-strain data.

SUMMARY AND CONCLUSIONS (a) In the upper-shelf temperature range, materials will exhibit ductile flaw extension prior to flaw instability, and it is appropriate to use elastic-plastic fracture mechanics analysis methods when evaluating a flaw under these conditions. Application of linear elastic fracture mechanics analysis methods in the upper shelf temperature range can be overly conservative. (b) The proposed Code Case N-749 of ASME Section XI provides alternate acceptance criteria based on using elastic-plastic fracture mechanics methodologies for evaluation of flaws in ferritic steel components operating in the upper shelf temperature range. (c)The technical requirements in proposed Code Case N-749 are based in part on those in Appendix K of Section XI, which allows the use of elastic-plastic fracture mechanics procedures for reactor pressure vessels with low upper shelf toughness. Structural factors consistent with other provisions in ASME Section XI which allow the use of elastic-plastic fracture mechanics for actual flaws, are required in proposed Code Case N-749. (d) It is expected that the acceptance criteria in proposed Code Case N-749 will reduce the excessive conservatism inherent in flaw evaluation methodologies for ferritic components in ASME Section XI that are currently based on linear elastic fracture mechanics.

ACKNOWLEDGMENTS

This work was conducted under the activities of ASME Section XI Working Group on Flaw Evaluation. The authors wish to acknowledge the support from all Working Group members in developing the proposed acceptance criteria.

REFERENCES 1. “Rules for Inservice Inspection of Nuclear Power Plant

Components, Section XI, Division 1,” ASME Boiler and Pressure Vessel Code, 2010 Edition.

2. “Case N-749: Alternative Acceptance Criteria for Flaws in

Ferritic Steel Components Operating in the Upper Shelf Temperature Range, Section XI, Division 1," ASME Boiler and Pressure Vessel Code, Draft, 2011.

3. Rice, J. R., “A Path Independent Integral and the Approximate Analysis of Strain Concentration by Notches and Cracks,” ASME Journal of Applied Mechanics, Vol. 35, pp. 379-386, 1968.

4. Hutchinson, J. W., and Paris, P. C., “Stability Analysis of J-

Controlled Crack Growth,” American Society for Testing and Materials, ASTM, STP 668, pp. 37-64, 1979.

5. Kumar, V., German, M. D. and Shih, C. F. “An Engineering

Approach for Elastic-Plastic Fracture Analysis,” Electric Power Research Institute, Palo Alto, CA, Report No. NP-1931, July 1981.

6. “Development of Criteria for Assessment of Reactor

Vessels with Low Upper Shelf Fracture Toughness,” Welding Research Council Bulletin 413, July 1996.

7. “Rules for Construction of Nuclear Power Plant

Components, Section III, Division 1,” ASME Boiler and Pressure Vessel Code, 2007 Edition.

8. Lidbury, D. P. G., “The Significance of Residual Stresses in

Relation to the Integrity of LWR Pressure Vessels,” International Journal of Pressure Vessels and Piping, 1984.

9. Merkle, J. G., “Appraisal of Data Concerning Residual

Stresses in Main Pressure Vessel Welds,” ORNL/NRC/LTR-96/15, 1997.

10. Formby, C. L. and Griffiths, J. R., “The Role of Residual

Stress in the Fracture of Steel,” Welding Institute Conference: Residual Stresses in Welded Construction and Their Effects, London, 1977, pp. 359-374.

11. “Test of 6-in.-Thick Pressure Vessels. Series 3:

Intermediate Test Vessel V-8A – Tearing Behavior of Low-Upper-Shelf Material,” NUREG/CR-4760, 1987.

12. “Section II, Part D - Properties,” ASME Boiler and

Pressure Vessel Code, 2010 Edition. 13. Newman, J. C. and Raju, I. S., “Stress-Intensity Factors for

Internal and External Surface Cracks in Cylindrical Vessels,” ASME Journal of Pressure Vessel Technology, Vol. 104, pp. 293-298, 1982.

14. “Recommended Practice for Fitness-For-Service,” API 579,

American Petroleum Institute, Washington, D.C., 2001.


TABLE 1 ASME SECTION XI APPENDIX K FLAW PARAMETERS

AND ACCEPTANCE CRITERIA

PARAMETER

LEVELS A AND B

(K-2200)

LEVEL C (K-2300)

LEVEL D (K-2400)

Flaw Size 1/4 t Up to 1/10 t

1 in. max (25 mm max)

Up to 1/10 t 1 in. max

(25 mm max) SF for Initiation, J < J0.1

On pressure (PACC(1) for

Levels A and B) 1.15 1.0 1.0

On thermal stress 1.0 1.0 1.0 On residual stress na na na

SF(1) for Stable Tearing On pressure (PACC

(1) for Levels A and B) 1.25 1.0 1.0

On thermal stress 1.0 1.0 1.0 On residual stress na na na

J-R Curve Conservative Conservative Best Estimate Ligament Stability

Evaluation No No Yes K-5300(b)

NOTE: (1) Accumulation pressure PACC = 1.1 times design pressure PD, and PD = 1.1 times operating pressure PO.


TABLE 2 COMPARISON OF ACCEPTANCE CRITERIA IN

ASME SECTION XI APPENDIX K AND PROPOSED CODE CASE N-749

Service Level

APPENDIX K ACCEPTANCE CRITERIA(1)

Flaw Size

SF for Initiation SF for Stable Tearing J-R Curve

Ligament Check Approach Pressure Thermal Pressure Thermal

A and B 1/4 t 1.15 1.0 1.25 1.0 Conservative No

Limited Ductile Crack

Extension and Stability

C up to 1/10 t

1 in. (25 mm) max

1.0 1.0 1.0 1.0 Conservative No

D up to 1/10 t

1 in. (25 mm) max

1.0 1.0 1.0 1.0 Best Estimate Yes

Service Level

PROPOSED CODE CASE N-749 ACCEPTANCE CRITERIA(2, 3)

Flaw Size

SF for Initiation SF for Stable Tearing J-R

Curve Ligament

Check Approach Primary Second-

ary Primary Second-

ary A and B Actual 2.0 1.0 na na Conservative No

Limited Ductile Crack

Extension

C Actual 1.5 1.0 na na Conservative No

D Actual 1.5 1.0 na na Conservative No

A and B Actual 1.5 1.0 2.14 1.0 Conservative No Limited Ductile Crack

Extension and Stability

C Actual 1.25 1.0 1.2 1.0 Conservative No

D Actual 1.25 1.0 1.2 1.0 Conservative Yes NOTES: (1) For pressure loading, accumulation pressure is used for Levels A and B (normal and upset) and accident pressure for

Levels C and D (emergency and faulted). (2) For pressure loading, operating pressure is used for Levels A and B and accident pressure for Levels C and D. (3) PD = 1.1PO, PACC = 1.1PD, flaw size margin = 2 (~√2 on load).


TABLE 3 RESULTS FROM EVALUATION OF TEST VESSEL V-8A FLAW

USING PROPOSED CODE CASE N-749

CALCULATED QUANTITY

VALUE

Evaluation for crack extension of 0.1 in.

Japplied for pressure loading (SF = 2.0) 653 in-lb/in.2 (114 kJ/m2)

JR at crack extension of 0.1 in., J0.1 628 in-lb/in.2 (110 kJ/m2)

Japplied < J0.1 ? No

Evaluation for crack extension of 0.1 in. and stability

Japplied for pressure loading (SF = 1.5) 367 in-lb/in.2 (64 kJ/m2)

JR at crack extension of 0.1 in., J0.1 628 in-lb/in.2 (110 kJ/m2)

Japplied < J0.1 ? Yes

Flaw stable at applied pressure loading (SF = 2.14) ? Yes


TABLE 4 COMPARISON OF PREDICTED PRESSURE AND DUCTILE CRACK EXTENSION AT FLAW

INSTABILITY IN TEST VESSEL V-8A WITH EXPERIMENT

PARAMETER

PREDICTION

TEST

Applied internal pressure loading 19.3 ksi (133 MPa) 20.0 ksi (138 MPa)

Ductile crack extension 0.54 in. (13.7 mm) 0.53 in. (13.6 mm)

Applied J integral 1073 in.-lb/in.2 (188 kJ/m2) na


Figure 1: Fracture Loads as a Function of Temperature for Specimens With (+) and Without (ο) Residual Stress;

F Represents Fatigue Pre-Cracked Specimens Without Residual Stress (Test results of Formby and Griffiths [10], as given in Reference [8])


Figure 2: Schematic of Vessel V-8A Showing Location of Flaw Relative to Special Seam Weld


0

200

400

600

800

1000

1200

1400

0 0.2 0.4 0.6 0.8

J (in

.-lb/

in.^

2)

Crack extension, ∆a (in.)

Applied J at 2.19*operating pressure

Applied J at 2.14*operating pressure

Material J_R

Figure 3: Evaluation of Test Vessel V-8A Flaw Using the J-R Curve – Crack Driving Force Diagram

Alternative Acceptance Criteria for Flaws in Ferritic ... Acceptance... · The flaw evaluation...

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