ML20092N654

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Discusses Fitness-for-Svc Evaluation of Radiographic Indications in Piping Sys at Peach Bottom Atomic Power Station... Re Fracture Analyses of Welds W/Unacceptable Radiographic Indications,Per LER 2-83-24/1T
ML20092N654
Person / Time
Site: Peach Bottom  
Issue date: 06/22/1984
From: Logue R
PECO ENERGY CO., (FORMERLY PHILADELPHIA ELECTRIC
To: Murley T
NRC OFFICE OF INSPECTION & ENFORCEMENT (IE REGION I)
References
NUDOCS 8407030394
Download: ML20092N654 (4)
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{{#Wiki_filter:I-PHILADELPHIA ELECTRIC COMPANY 2301 M ARKET STREET P.O. BOX 8699 PHILADELPHI A. PA.19101 (215)841-4000 ^ ~~ June 22, 1984 Docket Nos. 50-277 50-278 s .an= Dr.- Thomas E. Murley, Administrator Office of Inspection and Enforcement U.S. Nuclear Regulatory Commission Region I ~', 631 Park Avenus King of Prussia, PA 19406 FinalReportofthe-FractureMecha$ics

SUBJECT:

Analyses for Weld Acceptability on'the Weld . Imperfections Identified by Philadelphia Electric Compdny' in Licensee Event Report 2-83-24/1T

REFERENCE:

1. Letter to Dr. ,0. E. Murley, NRC, from W. T.: Ullrich,'PEco, dated March 30, 1984 (Preliminary Report of the Fracture Mechanics Analjaes Jor. Weld Acceptability) 2.- Letter to Dr. T. E. Murley,.NRC, from W. T. Ullrich, 'PEco,' dated April 10, 1984 (Supplemeni to Prelin.inary Report of the Fracture Mechanics Analyses'for. Weld Acceptability)-

Dear Dr. Murley:

The attached " Fitness-for-Service Evaluation of Radiographic Indications in Piping System 3 at the Peach Bottom Atomic Power Station - Units';2 and - 3", prepared by' Stone and Webster Engineering,Corpocation under contract to Philadelphia Electric Company, is' the final report on the fracture analyses of the welds whose radiographs were discovered'to contain unacceptable indications as reported in Licensee Event Report 2-83-24/lT Attachment to Dr. 7. E. Murley from M. J. Cooney, PECo, dated November 7, 1983. ,,,.ng

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e L Dr. Thomas E. Murley June 22, 1984 Page 2 As stated in the referenced March 30, 1984 letter, Philadelphia Electric Company discovered that certain Class I piping radiographs were improperly read by Eastern Testing and Inspection, Inc. (ETI) during previous outage modification work. A review of the 131 welds that had been radiographed by ETI in safety-related systems discovered twenty-three welds with unacceptable indications. These unacceptable indications are located in four systems of the plant as follows: Peach Bottom Unit No. of Welds System 2 3 Scram Discharge Volume 2 1* Reactor Water Cleanup 2 5 Core Spray 2 2 Feedwater 3 8 Scram Discharge Volume 3 3 Reactor Water Cleanup 3 1 Core Spray

  • Note:

This weld and'its associated piping will be replaced as part of the Recirculation Pipe Replacement modification which is in progress for Peach Bottom Unit 2. Subsequent to these findings, conference calls were made to the Regional Office of the NRC reporting on the preliminary fracture mechanics evaluation performed on the ' worst case' weld imperfections in each of the four systems. These preliminary findings revealed that the weld imperfections would not propagate to an unacceptable condition within the design life of the plant (40 years) for either the Scram Discharge Volume, Core Spray, or Reactor Water Cleanup Systems. The pre iminary fracture analyses performed on the Feedwater System, ".ich were based on code allowable stresses rather than the act Sl stresses, discovered that the ' worst case' feedwater system imperfection was suitable for 240 cycles (approximately 18 years of cycles). Prior to submission of the March 30, 1984, preliminary report, the NRC requested that Philadelphia Electric Company perform microdensitometer readings of the radiographs to confirm the size and depth of the weld imperfections as identified by a Philadelphia Electric Company NDE Level III. The

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\\ v i 4 Dr. Thomas E. Murley June 22, 1984 J Page 3 microdensitometerar adings1werem perfbrmedonradiographsoffive welds that were selected by-PECo and the NRC (af ter NRC film review, per Inspection Notice:50-277/84-05; 50-278/84-05, T. T. ~ Martin, NRC,.to-S. L.. Daltroff, PECo, dated February 17, 1984). 1As stated'in~the re'ferenced April 10, 1984 letter, Philadelphia Electric Company'has concluded, based-on densitometer readings of these five radiographs, that the defect characteristics observed by the NDE Level III are valid. The attaciled report describes' each of the twenty-three welds, the basis'of?the fatigue analysis and the findings of these analyses for all these welds. The findings of the report are summarized therein in Table '2, 'Results of Fatigue Analysis'. Code calculations indicate that the number of loading cycles that. the weld with the worst defect can sustain m safely, considerably exceeda the estimated total number of i cyclic events'f'or each' system over the plant lifetime. On this premise, Philadelphia Electric Company concludes that the weld defects within these' systems are acceptable for the design life .of the plant.:r i Submission of this' final report completes the corrective actions as described in Licensee Event Report 2 24/lT-Attachment dated November 7, 1983 (Dr. T. E. Murley, NRC, from M. J. Coon'ey, PECo). Should you require additional information, please do not' hesitate to contact us. s \\ .Very t y s, ch ~ R. H.'Logue, 96perintendent Nuclear Services ONuclear Genera _ on Division . Attachment cc: A. R. Blough, site Inspector l NRC Document Control Desk 1 l l t. -t l

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. 1-- PHILADELPHIA ELCCTRIC COMPANY . PEACH. BOTTOM ATOMIC POWER STATION UNITS 2 AND 3 FITNESS-FOR-SERVICE EVALUATION OF RADIOGRAPHIC INDICATIONS IN PIPING SYSTEMS ' DOCKET NUMBERS 50-277 AND 50-278 Submitted to l 2 THE UNITED STATES NUCLEAR REGULATORY COMMISSION l-f June 1984

r I 8 e

SUMMARY

During the period between 1980 and 1982 Philadelphia Electric Company (PECO) performed modification work on several piping systems at Units 2 and l 3 of Peach Bottom Atomic Power Station. These systems contained a number ~ of welds which were radiographed and accepted at that time. PECO re-examined these radiographs in 1983, and concluded that some radiographic indications in the scram discharge volume, reactor water clean-up, core Stonc & spray and feedwater startup bypass systems were not acceptable. Webster Engineering Corporation evaluated these indications using a fitness-for-service approach based on the methods and requirements of ASME Boiler and Pressure Vessel Code, Section XI,1983 Edition, and Winter 1983 Addenda. 'The indications detected during re-examinations of the radiographs were identified by PECO as weld cracks, lack of fusion, lack of penetration, and slag inclusions. An analysis of.the effect of these crack-like flaws indicated that fatigue crack growth is the only relevant failure mechanism under the operating conditions. A fatigue analysis was performed using the + -methods of linear elastic fracture mechanics consistent with Section XI of ASME Code. Several conservative assumptions were made in the analysis. The major ones are-1) che flaws are modeled as cracks with a postulated depth j significantly greater than.the reported depth, 2) all applied stresses act concurrently, 3) welding residual stress is tensile and has its maximum value at-the location of the flaw, 4) all flaws could grow regardless of their initial size, s The results of the analysis showed that the number of loading cycles t required to propagate the postulated flaws to an unacceptable limit considerably exceeds the estimated total number of cyclic events over the plant 1 lifetime. l i l l f 4 BD1-599510-80F an. - ~ - - - - - - - - -, - --

1 e l TABLE OF CONTENTS 4 Section Title Page 1-1 1 INTRODUCTION. . 2-1 2 TECHNICAL ANALYSIS.

2.1 DESCRIPTION

OF THE SYSTEMS . 2-1 . 2-1 2.1.1 Reactor Water Cleanup System. . 2-2 2.1.2 Scram Discharge Volume System . 2-3 2.1.3 Feedwater Startup Bypass System . 2-3 2.1.4 Core Spray System. 2.2 EFFECT OF INDICATIONS ON FITNESS FOR SERVICE. .2-4 2.3' FATIGUE ANALYSIS. . 2-5 . 2-5 2.3.1 General 2.3.2 Stress Intensity Factor. . 2-6 2.3.3 Crack Growth....... . 2-7 2.3.4 Material Constants. . 2-9 . 2-10 2.3.5 Acceptance Criteria 2.4 ANALYSIS AND RESULTS. . 2-11 2.4.1 Reactor Water Cleanup System. . 2-11 2.4.2 Scram Discharge Volume System . 2-11 2.4.3 Feedwater Startup Bypass System . 2-12 . 2-12 2.4.4-Core Spray System 2.5 DISCUSSION. . 2-13 3 CONCLUSIONS . 3-1 4 REFERENCES. . 4-1 7 BD1-599510-80E i

p- 's t.i LIST OF TABLES - Table Title 1 Locations and Sizes of Radiographic Indications 2 Results of Fatigue Analysis LIST OF FIGURES Figure Title 1 RWCU Pipes and Welds 2 RWCU Pipes and Welds 3 SDV Pipes and Welds 4 SDV Pipes and Welds 5 SDV Pipes and Welds 6 FW Pipes and Welds 7 CS Pipes and Welds 8 CS Pipes and Welds 9 CS Pipes and Welds 10 CS Pipes and Welds 11 CS Pipes and Welds 12 CS Pipes and Welds 13 Typical Fatigue Stress Cycles t BD1-599510-80E 11 e-e- v,w--, p

i i SECTION 1 INTRODUCTION 4 During the period between 1980 and 1982, several systems were modified at Units 2 and 3 of Peach Bottom Atomic Power Station (PBAPS). Modifications included a number of pipe welds. These welds were radiographed and accepted at that time. Philadelphia Electric Company (PECO) re-examined these radiographs in 1983, and concluded that some radiographic indications 'in the scram discharge volume, reactor water clean-up, core spray and feedwater startup bypass systems were not acceptable. PECO asked Stone & Webster Engineering Corporation (SWEC) to evaluate the effect of these indications on the performance of these systems. SWEC evaluated the indications using a fitness-for-service approach based on the methods and requirements of ASME XI. This report summarizes the results of these evaluations; details are given in References 1-8. The report discusses the effect of indications on different postulated rupture modes with an emphasis on fatigue crack growth that could be initiated by the indications. 4 BD1-599510-80 1-1

s. s SECTION 2 6 TECHNICAL ANALYSIS e

2.1 DESCRIPTION

OF THE SYSTEMS i Basic characteristics of the systems containing the unacceptable indications are briefly described in the following four subsections as r given in Ref. 9 and 10. A detailed description of the indications and their locations are given in Section 2.4. Reactor Water Cleanup System.(RWCU) - Modification 686 2.1.1 i The RWCU provides for the continuous mechanical and chemical filtration and demineralization of a portion of reactor water that is not circulated Reactor water is removed from the through the condensate demineralizers. suction side of the recirculation loop via shutdown cooling suction piping or from the bottom drain line on the vessel. This flow is cooled in the regenerative and nonregenerative heat exchangers, filtered and demineralized, and returned to the feedwater system through the regenerative heat exchanger. i The RWCU system is designed to support the following reactor operations: 1. Normal'or Reactor Power Operation 2. Startup Operation 3. Blowdown Operation 4. Refueling Operation 5. Hot Standby Operation The welds of Modification 686 are located on the RWCU recirculation pump suction piping inside the Primary Containment (See Figures 1 and 2). The following process parameters are applicable to these welds (maximum values): t 2-1 BD1-599510-80A = -. - - - _

e Temperature ( F) Pressure (psig) Flow Operation Initial Final Initial Final (gpm) Frequency ~ Reactor Power *) 522 531 1018 1035 477 6 Startup 100 522 0 217 411 4/ year Blowdown 100 545 0 1018 189 4/ year Refueling (or shutdown) 522 545 1000 364 4/ year 545 375 364 4/ year 375 330 364 4/ year 330 100 0 364 4/ year Hot Standby 531 545 1000 1018 364 4/ year

  • )

Not used in fatigue calculations due to small temperature and pressure changes, resulting ino K<K dKth (see Section 2.3). I The maximum temperature used in the fatigue calculations was 545F. The maximum pressure used in the calculations was 1135 psig as given in Reference 10. The pipe material is ASTM A312 TP-316L and the weld material is 308L. Pipe size and other characteristics are given in Table 1. 2.1.2 Scram Discharge Volume (SDV) System - Modification 655 The scram discharge volume is used to limit the loss of and contain the reactor vessel water from all the drives during a scram. This volume is provided in the scram discharge header. During normal plant operation, each discharge header is empty (temp. = 70 F) and the drain and two vent valves are open. Upon receipt of a scram signal, the drain and vent valves close. During a scram the control rod drives inject 3.34 gal (at 280 F) each into the SDV through approximately 100 ft of 3/4 in. and 12 ft of 1 in. sch. 80 stainless steel pipe in about 30 sec. Unit 2 SDV is 6 in. sch. 80 pipe (775 gal volume plus 53 gal Instrument Volume) and Unit 3 SDV is 8 in. sch. 80 pipe (930 gal volume plus 53 gal Instrument Volume). Subsequently the SDV will continue to fill at a rate between 185 gal / min. and 550 gal / min. depending on the condition of the drive seals until SDV pressure equals reactor pressure. When the scram signal is cleared (usually less than 20 min.) the SDV scram signal is overridden and the SDV is automatically drained. The locations of the modification welds containing indications are shown in Figures 3, 4 and 5. The pipe material is ASTM A106, Gr. B and the weld material is E 7018. Pipe sizes and other characteristics are shown in Table 1. BDI-599510-80A 2-2

n 2.1.3 Feedwater Startup Bypass System (FW) - Modification 381 The feedwater startup valve provides a flowpath around the reactor feedpump (RFP) discharge gate valve to allow control of feedwater to the reactor for startup, shutdown and hot standby operations. The objectives are to provide better control of reactor water level and minimize thermal cycling + of feedwater nozzles on the reactor vessel. Normal Startup Startup valve (AO-8091) and associated piping is placed in service when reactor pressure reaches about 600 gsig (flow is about 250-300 gpm and feedwater temperature is 100 to 120 F). The startup valve remains in service until reactor power reaches about 10-15 percent power, which is prior to placing any of the feedwater heaters in service. Flow conditions at this point are 1500 to 2000 gpm at a temperature of 100-120 F. At this point the startup valve (AO-8090) and the startup block valve (MO-8090) are closed and all feedwater flow to the reactor passes through the RFP discharge gate valve (MO-2149C). Feedwater tenperature begins to increase as the heat cycle feedwater heaters are placed in service. The temperatures typically increase from 100 F to a maximum of 376 F (full power) over a time period of at least 24 hours. However, the feedwater startup bypass system temperature does not exceed 120 F. The latter temperature is used in the calculations. Normal Shutdown Normal shutdown is much the same as the startup, except the events occur in reverse order. Feedwater temperature decreases as the power level is gradually reduced and the heat cycle is removed from service. The startup valve would only be glaced in service after feedwater temperatures have -dropped to below 150 F. Hot Standby The startup valve could be used during hot standby conditions to maintain reactor vessel water level (feedwater temperatures 100-120 F). The assumed number of transients is: Startups and Shutdowns 4/ year Hot Standby 9/ year The locations of the welds containing indications are shown in Figure 6. The pipe material is ASTM A106 Gr. B, the weld material is E7018. Pipe sizes and other characteristics are given in Table 1. 2.1.4 Core Spray System (CS) - Modification 389 The purpose of the core spray system is to prevent overheating and melting of the fuel during a loss-of-coolant accident. BD1-599510-80A 2-3

i j The core spray system consists of two completely independent spray loops. [ The equipment for each loop consists of two 50 percent core-spray pumps, a j sparger ring, a spray nozzle, and the necessary piping, valves, and j instrumentation. Each pump will take water from the suppression chamber by (' suction and will spray the water through the sparger ring into the plenum chamber above the core. g; The welds in question are located between the core spray isolation valves j and the reactor vessel (Figures 7 through 12). This portion of the system is without flow during all conditions except loss-of-coolant accidents and { system flushing during shutdown. In case of no flow conditions, with reactor operating, for the piping between the isolation valve and the reactor normal operating pressure is 1025 psig, and design pressure is 1135 psig. The temperature gradient along the pipe is such that the temperature greater than 6 ft from the reactor vessel nozzle is drywell ambient (135 F) and less than 6 ft is 575 F. During full injection the system is at a pressure of 127.7 psia, a temperature of 210 F, and a flow of 6250 gal / min. The pipe material is ASTM A312/TP316L and the weld material is 308L and 309L. Pipe sizes and other characteristics are given in Table 1. p 2.2 EFFECT OF INDICATIONS ON FITNESS FOR SERVICE t t Indications detected during the re-examination of the radiographs for the j four systems were identified as cracks, lack of fusion, slag inclusions and t lack of penetration. According to para. IWA-2110 and A-1100 of ASME XI { lack of fusion or lack of penetration flaws are classified as cracks. The locations and sizes of indications as reported by PECO (Ref. 10) are shown l in Table 1. Cracks and crack-like flaws can cause failure of metal com-j ponents through the following mechanisms: rupture caused by unstable crack propagation, [ a. b. fatigue crack growth resulting in unstable rupture or leak, I enhanced corrosion (stress corrosion cracking (SCC) and/or c. j corrosion fatigue). The failure modes, except for SCC, can be analyzed using the methods of fracture mechanics described in Appendixes A and C of ASME XI. SCC is not f j

a. problem in the cases analyzed in this report for the following reasons:

{ i I ASTM A106 Gr. B pipe and E7018 weld metal used in the SDV and l a. feedwater systems are carbon steels which are not sensitive to SCC in the reactor water environment b. Reactor Water Clean-up system and Core Spray System are made of low carbon stainless steel (316L) and 308L and 309L weld metals which are not sensitive to intergranular SCC due to low susceptibility of these materials to sensitization during welding.. Chloride induced transgranular SCC is not considered as a problem since the water chemistry in both systems is controlled. l l BD1-599510-80A 2-4 l i

. s ASME XI requires fracture mechanics analysis of indications if they exceed Most of the indications certain size relative to the component thickness. shown in Table 1 would not have to be analyzed applying the ASME XI criteria if the reported indication depth were used. However, the size of Due to inherent indications was determined by interpreting radiographs. uncertainties of such interpretations, for the purpose of the analysis we assumed that the depth of the postulated cracks is 1/16 in, for all indications with reported depth up to 0.025 in. and 1/8 in. for all indications with reported depth gred er than 0.025 in. and up to 0.060 in. When this is assumed me-; indications listed in Table I have to (Ref. 5). be analyzed. References 5 through 8 show that the initial size of the postulated flaws This can be concluded from the fact that cannot cause unstable rupture. the fatigue analyses in these references show that the welds containing the postulated flaws must be exposed to a large number of cycles before reaching a critical size. Therefore, the only failure mechanisms that has to be analyzed is fatigue crack growth. The effect of environment on fatigue crack growth was taken into account by using the appropriate The fatigue analysis is materials constants (see Subsection 2.3.4). presented in the next section. 2.3 FATIGUE ANALYSIS 2.3.1 General As pointed out in the previous section, the rules of ASME XI consider the indications analyzed in this report to be cracks. The most convenient way to analyze the effect of such flaws on the initiation of fatigue cracks and their propagation is through the use of the methods of linear elastic fracture mechanics (LEFM). This particular application of LEFM is now Its theoretical background and practical universally accepted. applications are described in numerous books and articles (see for example Refs. 11 and 12). The rate of fatigue crack growth can be expressed by the Paris equation: da = C,(aK )" (1) g dN where a is the crack size (the depth for surface and edge cracks) N is the number of cycles C, and n are material constants K is the stress intensity factor range: y K = (K ) max - (K ) min (2) g y y Where (K ) max and (K ) min are values of the stress intensity factor, K, y calculated for the maximum and minimum stress during the fatigue cycle. y y The stress intensity factor used in fatigue calculations almost always corresponds to the crack opening mode (or mode I, where crack surface 2-5 BD1-599510-80A l

t displacement is normal to the crack plane). A general expression for the mode I stress intensity factor is: y = FS 8a' (3) K and 4 K = FAS 8a (4) y where S is the applied stress, and AS = (max. stress) - (min. stress) F is a parameter that depends on the shape and size of the crack and geometry of the body containing the crack. The above methodology is accepted by ASME XI (see Appendices A and C of ASME XI - Winter 1983). Fatigue analysis of the indications discussed in this report is consistent with the requirements of ASME XI (see Refs. 5 through 8). The following subsections summarize the details of these calculations 2.3.2 Stress Intensity Factor To calculate the stress intensity factor range, AK, in Eq. (4) it is y necessary to know the crack size, stress range during a cycle, and the i parameter F. As stated in Section 2.1 it was assumed for the purpose of the analysis that the depth of the postulated cracks is 1/16 in, for all indications i with reported depth up to 0.025 in, and 1/8 in for all indications with j '~ . reported depth greater than 0.025 in, and up to 0.060 in. The length of the postulated cracks corresponds to the reported length (See Table 1). 1 All the indications were modeled as semielliptical cracks of depth a and length L. ~The orientation of the postulated cracks was based on the orientation of the reported indications. In most cases the plane of the indication was normal to the axis of.the pipe (circumferential orientation)..When the orientation was non-circumferential both orientations, circumferential and longitudinal, were considered in the analysis. The following cyclic stress contributions are evaluated where applicable: i 1. Stresses due to the restraint of the free end displacement caused by. thermal expansion. 1 2. Internal pressure stress. 1 i 3. Transient through wall thermal stress due to fluid temperature l change (the linear part (AT )). 3 t These stress contributions are conservatively considered to act f concurrently. The stress components normal to the face of the indications are evaluated. l -These stresses are separated into membrane and bending components. t I i BD1-599510-80A 2-6 l i l

~ Since the welded joints containing indications were not post-weld heat treated welding residual stresses act on the postulated cracks. These stresses are static and do not directly affect AK value in Eq M). y However, a static stress affects the mean stress value, which has an effect on crack growth rate, da/dN. As will be shown later in Section 2.3.4 this indirect effect of residual welding stresses was included intthe a calculations. The total stress intensity factor range that includes effects of all the stresses can be written as: AK = (aS, + H A S ) F (5) 7 b Where S is the uniform tension stress S" is the outer-fiber bending stress Q is the crack shape factor. Factors H, F and Q are given in Ref. 13 for elliptical surface cracks in flat plates. The values of these factors depend on: o a/t = crack depth / plate thickness, a/L = crack depth / crack length, o o f = angle defining the position at the crack tip o (f = 90 at the tip of the crack, and 0 at the point where the crack intersects the surface). type of loading (uniform tension or bending) o Similar solutions are available for cracks in cylindrical bodies (Ref. 14). A comparison with Eq (5) showed that differences between flat plate and cylindrical solutions were insignificant when those solutions were applied to the cases discussed in this report. It was decided to use Eq (5) since factors H and F are given in Ref. 13 in an analytical form, while those for ' cylindrical bodies are tabulated, and thus less convenient for numerical integration of Eq (1) (See 2.3.3 below). 2.3.3 Crack Growth The analysis usually begins with an evaluation of the initial stress If the latter is below a threshold level,6K intensity factor range. there will be no crack growth. A K isamaterialpropertyandthevaIu'es t th for the steels in question have been reported in the literature. In a few cases the reported AK values, obtained under the conditions similar to th those discussed in this report, were higher than some of the calculated K values. But due to uncertainties in measuring AK and limited amount of th data it was assumed that the calculated AK values in all the cases exceeded the reported threshold values, so it was necessary to determine the crack growth rate by solving Eq. (1). BD1-599510-80A 2-7

The usual method is to calculate the number of cycles needed to grow a crack from its initial depth, a, to the final depth, a, by integrating g g Eq (1): a i N= da (6) J C, (eK)"

  • i This integral can be solved analytically if factor F in Eq (4), or factors H and F in Eq (5) do not change when a crack grows. However, these factors are dependent on crack depth and other related parameters, (see 2.3.2), so the direct integration is possible only for a small crack growth, when the factors will be practically constant due to a small change in the crack depth.

In this work piece-wise integration was applied. The integral was solved for a small crack increment, Aa=a -a takes place of The number of cycles, N, needed ko; where a 3 3 cause crack growth a in Eq (6). a 3 f was recorded. The new crack value was then obtained by addinga a to af resulting in the new initial crack depth a, = a +Aa, This was again 4 integrated in the same range a to obtain I corfesponding number of cycles, N. The process was then repeated until the final crack depth, a, was f 2reached, and the sum of all the cycles is obtained: ~. n a N=N ? da (7) iI1 y C, (AK)"

  • i (The final crack depth was based on the acceptance criteria described in 2.3.5).

The above integration process was performed in this work in two ways. 1) A simplified calculation method based on the following assumption was used: a) the maximum stress intensity factor value is at the maximum depth point of the crack and, b) the crack depth / length ratio stays constant during the i growth ("self similar growth").

2) A more accurate method that takes into account the variation of the stress intensity factor along the crack front was also applied. This variation depends on the a/L ratio and the relative magnitude of the bending load. For a/L = 0 to approximately 0.25 the maximum stress intensity factor occurs at the maximum depth point; for cracks with a/L> 0.25 exposed to bending the maximum value of the stress intensity factor is at the point where the crack intersects plate surface.

From this point of view the assumption made in the simplified calculation method that the stress intensity factor is constant and corresponds to the value at the maximum crack depth point is conservative, at least when a/L ratios are less than 0.25. As can be seen in Table 1 a/L is less than 0.25 in most cases. However, conservatism of the assumption about the "self similar growth" and its effect on N was tested by performing calculations taking into account change of the stress intensity factor along the crack front. As will be shown later thin more refined method gave higher N values in most cases, so the simplified method is usually more conservative, at least for the cases discussed in this report. BD1-599510-80A 2-8

r 2.3.4 Material Constants Eq(1) includes two material constants, C and n, which depend on type of material, en g nment and the ratio of tee minimum to the maximum stress in a' cycle (or /Kmax). In this work we discuss two materials: carbon 4 steel and austenitic stainless steel. The materials constants, C and n, for ferritic steels in air and in light-water reactor environment are given 'in Appendix A of ASME XI. Fig. A-4300-1 of that Appendix shows fatigue [ crack growth curves for reactor water environment at different values of R = min stress / max. stress. Data for austenitic stainless steels were I obtained from Ref.'15. These data were developed from tests performed in boiling we.ter reactor environment. As mentioned in 2.3.2 welding residual stresses affect the value of the mean stress, which has an effect on fatigue crack growth curves. Since it is an important factor it will be discussed here in some detail. This discussion is based on Ref. 16 which gives an excellent treatment of this subject. The fatigue terms and their algebraic relationships are defined below and in Figure 13: Smax - Smin (8a) Amplitude

Sa

Smax + Sein (8b) Mean stress = Se = 2 R - 8"I" - min (8c) - Smax K,,, y Smax is the sum of external and internal (residual) stresses. When Smax is less than the yield strength of the material, R is calculated from Eq (8c) and an appropriate da/dN curve is selected from Fig A-4300-1 of ASME XI or another source. However, the tensile residual stress in the weld region of an as-welded structure can be as high as the yield stress (YS). Therefore, during the first cycle Smax will exceed YS. In order to preserve the internal stress equilibrium, the excess stress over YS, equal to Smax - YS, will be redistributed to the areas that are still elastic. (This argument applies to an elastic-ideal plastic body, which does not strain harden; a more realistic case is discussed later.) When the external loads are removed, or become compressive, elastic unloading occurs throughout the affected areas resulting in a modified residual stress distribution. The maximum magnitude of the remaining residual stress is now equal to Sein = YS - Sext., where Sext is the sum of external stresses. Reapplication of the external stresses results in Smax = YS and after removal of the external stresses Sein'is again equal to YS - Sext. So after the first cycle the weldment will be subjected to pulsating type of cycling between YS and the difference between YS and Sext. For example, if the stress cycle based on external stresses is S to c, the nominal stress range is between kS and YS-($. The actual stress ranke will b S - (-S ) = SI+S 3 + S ), that is: 2 (Smax - Smin) = YS-YS - (S3+S =S3+S2 BD1-599510-80A 2-9

a L So the stress ratio with residual stresses is: b ~(b S3+S2 (9) Sein

  • 1*S)*

2 ~

  • Smax YS YS Without residual stresses:

6 S (10) which is a smaller number than R with residual stresses. Since fatigue crack growth is faster for higher R values welding residual stresses may have a significant effect on the number of cycles calculated from Eq (6). The above discussion is based on the assumption that plastic deformation is ideal, that is, that there is no strain hardening. To account for the strain hardening effect we used the flow stress instead of YS in Eq (9). Flow stresa is given by: YS + UTS 73, (11). 2 Where UTS is ultimate tensile strength. In this analysis, the cyclic j AccoIdinglf),, where H accounts for nonuniformity of + HS stress is expressed as (S the bending stress. ys. (Sm + HS ) b R= 73 2.3.5 Acceptance Criteria For carbon steel under normal operating conditions the acceptance criterion of IWB-3612 (a) of ASME XI y(K, [ (12) K 3 was used as suggested in paragraph IWB-3620. Here K is the maximum stress intensity factor for normal conditions for T and K is the crack arrest toughness. The maximum stress the f13w size a intensityfactorwascdiculatedusingallstresses(cyclicandstatic), f including the seismic loads. However, due to the redistribution of I residual stresses the maximum stress is always equal to the flow stress (see 2.3.4 above), so the acceptance criterion depends on the material properties (flow stress and fracture toughness), pipe geometry and crack depth'(a ). f All the indications were reported to be in the weld metal; therefore fracture toughness of the weld metal, SFAS.1 E7018, was used in the analysis for all indications at the inner surface (ID). Due to the weld configuration, indications at the outer surface (OD) could grow into the Therefore, for base metal, if they are near the weld-base metal interface. these indications fracture toughness of the base metal was used. The temperature difference, T-RTNDT, was estimated and the crack arrest BDI-599510-80A 2-10 [ -~ -~y,m-_..,m._,m,,- .-,_m.m,-.

o a fracture toughness, K was found from Fig. A4200-1 of ASME XI - Thecrif$r,ionexpressedbyEq(12)wasmodifiedinthose Winter 1983. cases where its application would allow deep cracks. The acceptable number of stress cycles was calculated on the basis that the crack would not grow value was less than l deeper than half of the wall thickness, even if the K7 that allowed by Eq (12). (See also the following discussion on stainless steel.) For austenitic stainless steel the acceptance criterion is given in IWB-3640 of ASME XI (1983 Winter Addenda). According to this criterion, the maximum allowable flaw depth to thickness ratio is 0.75. However, we use a more conservative criterion of a/ f 0.5, which is in agreement with i the present NRC's staff criterion that crack depth should not exceed two-thirds of the code allowable limit. (Reference 17). 2.4 ANALYSIS AND RESULTS Using the above described methods, the crack depth of the propagating postulated crack was evaluated as a function of a number of the load cycles Calculations were for constant and variable crack depth to length ratios. performed for all the welds, but results are shown only for the worst case flaws when more than one flaw is postulated in a given weld. The results are presented in the following subsections. Details are given in Refs. S through 8. 2.4.1 Reactor Water Cleanup (RWCU) System The RWCU system of Unit 2 contains one pipe to existing penetration weld with two indications (Table 1). The RWCU system of Unit 3 contains three welds with five indications (Table 1). Two of these welds are pipe / 90 deg elbow girth butt welds and the third is a girth butt weld joining two 45 deg L.R. elbows. This system The results of is exposed to RPV pressure cycling and expansion stresses. the crack growth analysis are shown in Table 2. Using the acceptance criterion described in 2.3.4 above, it was found that at least 7,900 cycles are required for a postulated flaw 0.0625 in deep to grow to the maximum allowable depth of 0.22 in., i.e., to 50 percent of the thickness. 2.4.2 Scram Discharge Volume (SDV) System Unit 2 contains three welds with a total of six indications (Table 1). Two i of the welds are girth butt welds joining 90-deg L.R. elbows to straight pipe, and the third is a girth butt weld joining straight pipe members. j \\ Five of Unit 3 contains 8 welds with a total of 19 indications (Table 1). the welds are girth butt welds joining 90-deg L.R. elbows to straight pipe, two join 45-deg elbows to straight pipe, and one is a butt weld joining and 8 in. x 12 in, weldolet to a tank. i BDI-599510-80A 2-11 l

The following cyclic stresses were considered in the analysis of this system: 1. Stresses due to the restraint of the free end displacement caused by thermal expansion. 2. Internal pressure stress. 3. Transient through wall thermal stress due to fluid temperature change (the linear part (6T D. 3 These stress contributions are considered to act concurrently, s The fatigue crack growth analysis showed that in the worst case (a 0.125 in. deep postulated flaw) it would take more than 7,580 cycles to exceed the acceptance criterion given in Eq. (12). (See Table 2.) The K y value in Eq. (12) for the flaws at the inner surface was based on fracture, toughness of the weld metal. 200 63 ksi [ K Ia Imax = = = 2.4.3 Feedwater Startup Bypass System The Unit 2 bypass contains two welds with four indications (Table 1). One weld is at a girth butt weld joining a welded end valve to the large end of a 10 in. x 8 in. schedule 100 concentric reducer. The other weld is a pipe / elbow girth butt weld. The cycling stresses acting on these indications are generated by the pressure and expansion. Postulated cracks with an initial depth of 0.0625 in. (reported depth was 0.010 to 0.015 in.) are expected to grow due to the cyclic loading into the weld metal. The acceptance criterion expressed by Eq (12) was applied using fracture toughness of the weld metal to calculate the K, value. Due y to a lower operating temperature, K, is less than in 2.4.2: g I [. = 39.8 kai Imax=55 Based on these results it was shown that the postulated cracks would remain within the allowable limits of Eq (12) for at least 27,000 cycles. (See Table 2.) 2.4.4 Core Spray (CS) System The Unit 2 CS contains five welds with eight indications (Table 1). Three of these welds are pipe-to-pipe girth butt welds, one is a pipe to elbow weld and the fifth is a girth butt weld adjacent to the RPV nozzle safe end (on the pipe side). The Unit 3 CS contains one weld with one indication (Table 1). The weld is a pipe / elbow girth butt weld. BD1-599510-80A 2-12 i

The CS system does not experience any operating transients (cyclic stress) The CS system experiences RPV pressure cycling and due to CS operation. free end expansion stresses caused by RPV expansion / contraction while the CS pipe remains essentially at ambient temperature. The depth of the reported indications varied between 5 and 60 mils. Initial postulated crack depths of 0.625 and 0.125 in. were assumed in this analysis. Using the acceptance criterion for stainless steel (the maximum allowable flaw depth equal to 50 pct of thickness) it,was found that the minimum number of cycles was 21,000 in Unit 2 and 3,950 cycles in Unit 3 (Table 2). 2.5 DISCUSSION Indications detected in the examined welds were identified as cracks, lack of fusion, lack of penetration, and slag inclusions. It was shown in Section 2.2 that fatigue crack growth is the only failure mechanism that these indications could cause under the operating conditions. The fatigue analysis described in previous sections is based on a number of conservative assumptions. To understand better our conclusions relative to the fitness-for-service evaluation of the analyzed piping systeras we will first summarize those assumptions. The depth of the postulated flaws is 2 to 6 times greater than 1. the depth of the reported indications. This increases the factor of safety in Eq (12) by about 1.4 to 2.6 times. The postulated flaws are assumed to be located at the point of 2. maximum applied stress. For example, thermal expansion produces bending moment resulting in a through-diameter stress distributien from a maximum tensile stress to a maximum compressive stress. It is assumed that the postulated flaw is at the point of the pipe circumference exposed to the maximum tensile stress. 3. It is assumed that all applied stresses act concurrently. 4. Welding residual stress is assumed to be tensile and has its maximum value at the location containing the flaw. Since the maximum stress is assumed to be the flow stress: FS = TS + YS = (Applied Stress) + (Redistributed Residual Stress) 2 this can result in an assumption that residual stress is above the YS value if applied stresses are low. 5. It was assumed that all the postulated flaws would grow, although in at least some cases the calculated AK vaines were lower than a reasonable estimate of AKth' 6. Crack growth was analyzed by two techniques: BDI-599510-80A 2-13

In almost all the cases the a/L = constant and a/L variable.latter more realistic technique, ga of cycles, but in the final evaluation, the lower results were used. The following table gives the minimum number of loading cycles for propagation of postulated cracks to an acceptable limit as calculated in 4 Refs. 5 through 8 and reported in 2.4 above. Lowest Calculated Estimated Number of Cycles Number of Cycles Unit to Acceptable Limit per year over 40 years System 28 1100 3,700,000 Reactor Water Cleanup 28 1100 3 7,900 15 600 2 8,090 Scram Discharge Volume 15 600 3 7,580 13 520 2 27,000 Feedwater Startup Bypass 13 520 2 21,000 13 520 Core Spray 3 3,950 The calculated number of cycles by far exceeds the estimated number of cycles. 2-14 BD1-599510-80A

SECTION 3 CONCLUSIONS Based on the results obtained in this work the following is concluded s Radiographic indications observed in welds in the reactor water cleanup, scram discharge volume, feedwater startup bypass, an 1. core spray systems may cause fatigue crack growth. The number of loading cycles required to propagate postulated ' flaws to an acceptance limit considerably exceeds the estimated 2. total number of cyclic events over the plant lifetime. 3-1 BD1-599510-80a

  • c FIGURES AND TABLES Welds analyzed in this report are identified by three groups of numbers:

(Modification) - (Unit) - (Weld) 686-3-7 represents Hodification 686 in Unit 3, Weld No. 7. For example: Modification numbers are related to the systems as follows: Modification No. System 686 RWCU 655 SDV 381 FW 389 CS The welds in Figures 1 through 12 are identified by the same code, excep that in some cases only the ?.ast number (Weld No.) is shown in the square box. To identify the analyzed welds compare Tables 1 and 2 with the figures. BD1-599510-80G

SECTION 4 REFERENCES F 1. SWEC Calculation 14006.24-NP(B)-1 2. SWEC Calculation 14006.24-NP(B)-2 r 3. SWEC Calculation 14006.24-NP(B)-3 4. SWEC Calculation 14006.24-NP(B)-4 5. SWEC Calculation 14006.24-MED-1 6. SWEC Calculation 14006.24-MED-2 7. SWEC Calculation 14006.24-MED-3 8. SWEC Calculati'n 14006.24-MED-4 o 9. Letter E. A. Clymer (PECO) to L.B. Hirst (SWEC), January 30, 1984 10. Weld Characteristic Summary - MOD 1366, PBAPS, Units 2 and 3, Eastern Testing and Inspection 11. D. Brock, Elementary Engineering Fracture Mechanics. Nordhoff Int. 1978. 12. S. T. Rolfe and J. M. Barsom, Fracture and Fatigue Control in Structures, Frentice-Hall, Inc., 1977 13. J. C. Newman, Jr. and I. S. Raju, Engineering Fracture Mechanics, vol. 15, No. 1-2, pp 185-192, 1981 14. F. Delale and F. Erdogan, J. Appl. Mech., Trans. ASME, vol. 49, pp 97-102, 1982. 15. R. Huet et al., SCC of 304 Stainless Steel in High-Purity Water, EPRI Report NP-2423-LD, 1982 i 16. T. R. Gurney, Fatigue of Welded Structures, Second Edition, Cambridge ' University Press, (p. 233)

17. Memorandum from W. V. Johnston, NRC to F. J. Miraglie, NRC, " Safety Evaluation of the Confirmation Order Inspection and Repairs on the l

Dresden Unit 3 Reactor Coolant Piping Systems"'(p. 3 of the Attachment), March 13, 1984

18. ASME Boiler and Pressure Vessel Code, 1983 Edition and Winter 1983 l

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, O SWEC REPORT 14 006.24 -JIB)-133 i h 0g7 1*. n [FW-389-2-3 ~ N j l2-A B-14-6 38 N o FW-389-2-4 U l2-AS-14-62 A 8 I n s' I b .a 4 5' V [ ./ O l2-A 8-14-6 2 B U FW-389-2-5 Il FW-389-2-6 >i5 $E DET All "A" U 379* 2Y[ 2 NOTES: I

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