ML20078F186

Refuel Outage B1R05 RHR HX Nozzles Ultrasonic Indication Summary

ML20078F186
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Site:	Byron
Issue date:	01/20/1995
From:	Hagemann G COMMONWEALTH EDISON CO.
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.

ENCLOSURE BYRON UNIT 1 REFUEL OUTAGE BIR05 RHR HEAT EXCHANGER NOZZLES ULTRASONIC INDICATION

SUMMARY

Prepared By:

G. Hagemann, CECO Byron SEC 9502010356 950120 PDR ADOCK 05000454 P PDR

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1) BACKGROUND l Byron requested relief (Preservice Relief Request NR-13) from the NRC

. during Preservice Inspection for the volumetric examinations for the -  !

Residual Heat Removal. Heat Exchanger (RHR) nozzle to vessel welds. The  ;

basis behind the request was due to the geometric restraints in performing l the examination as a result of the configuration of the nozzle to .

vessel welds. The configuration of the nozzle is such that construction did  ;

not allow for relevant volumetric examination as the installation utilizes an i internal reinforcing pad and a double bevel groove weld with a fillet j reinforcement. The relief request was granted by the NRC and documented -l in their review of the Unit 1 Byron Preservice Inspection Program Docket # l' 50-455, SSER 7 Appendix K.

i Byron requested similar relief for the Inservice Inspection program via  !

submittal ofInservice Relief Request NR-12. This request for relief sought the same exemption for volumetric examinations of the nozzle to vessel welds and inner radius that had been granted during Preservice. This relief  !

was granted for the inner radius inspection but requested that a "best l cffort" volumetric examination be performed on the nozzle to vessel welds. l This is documented in the NRC Safety Evaluatior. of the Byron Units 1 and  :

2 First Ten Year Interval Inservice Inspection Program.

i It should be noted that Braidwood had previously performed these  !

exnminations on their Unit 2 RHR Heat Exchangers in November of 1991 in i response to their like Relief Request and identified the existence of numerous' indications attributed to manufacturing defects. Braidwood solicited Westinghouse Corporation to perform a Fracture Mechanics Analysis in accordance with ASME Section XI IWB-3500 and/or IWB-3600.

The analysis was initiated for both Byron and Braidwood Units 1 and 2 RHR Heat Exchangers Nozzles and subsequently, this methodology was accepted by the NRC for application at both Byron and Braidwood. This analysis (Westinghouse Report MMDT-SMT-062) accepted the existence of flaws up to and including 60% of the thickness of the material at the flaw location, extending 360* and surface connected. The analysis accepted these types of flaws for the entire 40 year service life of the vessel.

During Byron Unit 2 B2R03 Refueling Outage in spring of 1992, several similar indications were identified in all four nozzle to vessel welds for the 2A and 2B RHR Heat Exchangers. These welds were accepted by the )

Westinghouse Fracture Mechanics Analysis previously mentioned, i

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2) INTRODUCTION Pursuant to the NRC response to Inservice Relief Request NR-12, a "best  ;

effort" ultrasonic examination was conducted on the Byron Unit 1 RHR '

Heat Exchanger nozzle to vessel welds.

During Byron Refueling Outage BIR05 the Unit 1 "B" RHR Heat Exchanger Outlet nozzle to vessel weld was scheduled for examination. Ultrasonic  ;

reflectors were identified in excess of 100% DAC (Distance Amplitude ,

Correction) reference levels with definable depth. These indications were  ;

evaluated and determined to be in excess of the acceptance criteria defined :

in IWB-3500. The examination were expanded to include the "B" Inlet nozzle where similar indications in excess ofIWB-3500 were recorded.

Ultimately, all four Unit 1 RHR nozzle to vessel welds were examined and indications in excess ofIWB-3500 were identified. These indications were sized and all were determined to be within the Westinghouse Fracture Mechanics Analysis performed for the Byron and Braidwood RHR Heat Exchangers and were determined to be acceptable for continued service. l These indications are consistent with the indications found at Braidwood -

and at Byron Unit 2 and are consistent with slag inclusions and/or lack of  ;

fusion in the fabrication process and are not service induced flaws. l As shown in this report all indications are within ASME Section XI '

Subarticle acceptance standards provided in IWB-3500 or IWB-3600. Those indications which were found to be acceptable analytically by the Fracture  :

Mechanics Analysis will be monitored by examinations that will be l conducted in future Byron Unit 1 Refueling Outage as required by /.SME Section XI IWC-2420.

3) FABRICATION INFORMATION The subject nozzles are on the tube side of the RHR Heat Exchangers.

These heat exchangers were inanufactured by Joseph Oats Corporation in 1975 in accordance with ASME Section III, NC-3200 Alternate Design Rules for Vessels. The heat exchanger vessel tube side is ASME Class 2 and the shell side is ASME Class 3. The nozzles are 3/8" nominal thickness (actual measured ultrasonic thickness is .400"),13.875 inches outside diameter, SA-240 TP304 rolled plate welded to the 1 inch thick SA-240 TP304 rolled plate of the tube side vessel wall.

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The weld configuration is a double bevel groove weld with an external 3/8" i nominal fillet weld reinforcement. The welding process was performed using the shielded metal arc process with E308 electrodes. Joseph Oat personnel and their fabrication procedures indicate that the groove weld was backgouged at the root pass (es) during the welding process. The inside and outside surfaces of the welded joints were examined by the liquid ,

penetrant method at the completion of welding. The welds were not radiographed during fabrication as due to the joint configuration. ASME Section III,1974 Edition, NC-3352 does not require radiographic  ;

examination of angle joints between the shell and nozzle which exceed 30'. ;

The ultrasonic examinations performed during this outage were the first volumetric examinations of these welds.

4) ULTRASONIC TECHNIQUES The RIIR heat exchanger nozzle to vessel welds were examined using a procedure that required a calibration which would produce an examination that meets the intent of ASME Section XI Appendix III requirements. ,

The examinations were conducted utilizing CECO NDE Procedure NDT-C-2 Revision 19. The examination was conducted utilizing an Ultrasonic Imaging Flaw Detector (P-Scan) which allows accurate interpretations of flaws and offers the advantage of consistent repeatability in the future.

Flaw indications were plotted on the P-Scan at the 50% DAC reference level and were sized to 50% of their maximum amplitude. Indications which exceed ASME Section XI IWB-3500 acceptance criteria will be monitored in future refueling outages as required by ASME Section XI, IWC-2420.

5) ULTHASONIC EXAMINATION RESULTS Attachment A summarizes, for each nozzle, the recordable indications detected during ultrasonic examinations. The flaws were found not be surface connected, however following the methodology of ASME Section XI, IWA-3310(b), the flaws have been considered as surface flaws and are so noted on these tables. The a/t column values were established based on the nozzle wall thickness (.400").

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6) EVALUATION OF ULTRASONIC EXAMINATION RESULTS The Fracture Mechanics Analysis previously referenced in this report was performed by Westinghouse in accordance with the methodology and criteria provided in ASME Section XI IWB-3600. This evaluation developed maximum allowed flaw dimensions for normal operating conditions and emergency and faulted conditions. The results of the Fracture Mechanics Analysis has been used to develop an upper boundary flaw chart for the recorded indications. The complete Westinghouse analysis which developed this criteria is provided as Attachment B to this report. Figure 4-1 in the Westinghouse Analysis shows the developed chart.

All indications from Section 5 which were found to be in excess of the criteria in IWB-3500 have been found to be within the Fracture Mechanics Analysis allowable limits.

7) PRIMARY COOLANT WATER CHEMISTRY The RHR heat exchanger tube side nozzles are exposed to only primary coolant water. The Chemistry requirements for primary coolant water are provided in Section 3/4.4.7 of the Byron Technical Specifications. These limits are 100 ppb dissolved oxygen when the te.nperature is greater than 250*F and 150 ppb for Fluoride and Chloride at all temperatures. A detailed discussion of Stress Corrosion-Cracking Susceptibility is provided in the Westinghouse Fracture Mechanics Evaluation MMDT-SMT-062 Section 3.4 which is provided as Attachment B to this report.

8) CONCLUSION All the ultrasonic examination indications detected on the nozzle to vessel welds to the 1A and IB RHR heat exchangers tube side nozzles have been found to be acceptable in accordance with ASME Section XI IWB-3500 or IWB-3600 as applicable. Therefore, with the prior NRC approval of the Byron /Braidwood evaluation methodology and guidance from NRC Generic Letter 91-18, system operability was not impacted.

Those flaws which have been determined to be acceptable by the Fracture Mechanics Analysis will be monitored in future Byron Unit 1 Refueling Outages as required by ASME Section XI IWC-2420.

5-

6 The ultrasonic examination indications noted are consistent with those noted at Braidwood. Unit 2 and at Byron Unit 2 and are consistent with fabrication flaws rather than service induced flaws.

i These indications are acceptable for continued service without repair as -

defined in the Westinghouse Analysis.  ;

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ATTACHMENT A i l

INDICATION

SUMMARY

l Component: 1RH02AA Weld No: 1RHXN.1 Outlet Nozzle A total of 17 indications above 50% DAC were recorded with 4 of these indications exceeding 100% DAC with definable length. ,

The following is a summary of 'ndications which exceed 100% DAC and exhibit a definable depth and iangth as reported in Ebasco ultrasonic report number 93BY1 'UT-028:

Ind. Loc. Max. Amp. Lgth. a** a/l a/t*

1 3.56-3.90" 100 % .34" .095" .28" 25%

2 4.00-4.36" 112 % .36" .131" .36" 35%

8 32.95-33.25" 200% .30" .170" .56" 45%

9 9 34.37-34.60" 100 % .23" .095" .41" 25%

• alt represents through wall percentages ofindication utilizing the nozzle wall thickness (.400").

Denotes ASME Section XI (Article IWA-3000) characterization (i.e.

a= surface, 2a= subsurface).

Identified flaws were characterized as surface flaws per Table IWB 3514-2.

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Component: 1RH02AA Weld No: 1RHXN-2 i Inlet Nozzle  ;

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A total of 13 indications ame 50% DAC were recordai with 5 of these indications exceeding 100% DAC with definable length. l l

The following is a summary ofindications which exceed 100% DAC and '

exhibit a definable depth and length as reported in Ebasco ultrasonic report .

number 93BY1-UT-029:

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Ind. Loc. Max. Amp. Lgth, a** a/l a/t*

1 .04 .36" 112 % .32" .080" .25" 21%  !

5 14.00-14.60" 100% .60" .095" .16" .25%

I 8 28.60-29.00" 112% .40" .080" .20" 21% i 9 31.15-32.25" 168/o 1.10" .095" .09" 25 %

11 34.90-35.95" 158% 1.05" .110" .10" b0% l f

a/t represents through wall percentages ofindication utilizing the nozzle wall thickness (.400"). ,

• • Denotes ASME Section XI (Article IWA-3000) characterization (i.e. .

a= surface, 2a= subsurface).

Identified flaws were characterized as surface flaws per Table IWB 3514-2.  ;

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Weld No:

Component: 1RH02AB 1RHXN-1 Outlet Nozzle A total of 11 indications above 50% DAC were recorded with 2 of these indications exceeding 100% DAC with definable length.

The following is a summary ofindications which exceed 100% DAC and -

exhibit a definable depth and length as reported in Ebasco ultrasonic report number 93BY1-UT-026-t i

Ind. Loc. Max. Amp. Lgth, a** a/l a/t*  !

1 3.84-4.20" 126% .36" .150 .42"

- 40%

3 6.70-7.70" 126% 1.00" .131" ,13" 35%

• a/t represents through wall percentages ofindication utilizing the  !

nozzle wall thickness (.400").

Denotes ASME Section XI (Article IWA-3000) characterization (i.e.

a= surface, 2a= subsurface). ,

Identified flaws were characterized as surface flaws per Table IWB 3514-2.

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Component: 1RH02AB Weld No: 1RHXN 2 Inlet Nozzle A total of 21 indications above 50% DAC were recorded with 11 of these indications exceeding 100% DAC with definable length.  ;

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The following is a summary ofindicationc which exceed 100% DAC and  !

exhibit a definable depth and length as reported in Ebasco ultrasonic report number 93BY1-UT-027:

Ind. Loc. Max. Amp. Lgth a** a/l a/t*

4 3 5.48-6.20" 112 % .87" .080" .09" 21%

6 16.85-17.20" 100% .35" .080" .35" 21%  :

8 25.95-27.15" 158% 1.20" .150" .12" 40%

9 31.65-32.00" 200% .35" .080" .23" 21%

10 31.50-31.95" 158% .45" .080" .17" 21%

11 30.65-31.00" 100 % .35" .080" .23" 21%

15 36.25-36.57" 100% .32" .060" .19" 16%

16 36.50-36.75" 112 % .25" .080" .32" 21%

17 37.20-39.00" 126 % 1.80" .080" .04" 21%

18 37.61-37.93" 200% .32" .080" .25" 21%

i 19 37.10-38.00" 126 % .90" .080" .09" -21%  !

a/t represents through wall percentages ofindication utilizing the nozzle wall thickness (.400").

Denotes ASME Section XI (Article IWA-3000) characterization (i.e.

a= surface, 2a= subsurface).

Identified flaws were characterized as surface flaws per Table IWB 3514-2.

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4 ATTACHMENT B FRACTURE MECHANICS ANALYSIS 1

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]MDT-SMT-062(92) -

I FRACTURE NECHANICS EVALUATION BYRON AND BRAIDWOOD UNITS 1 AND 2 RESIDUAL HEAT EXCHANGER ,

TUBE SIDE INLET AND OUTLET N0ZZLES Narch 1992 W. H. Bamford H. Jambusaria Y. S. Lee l

WESTINGHOUSE ELECTRIC CORPORATION Nuclear and Advanced Technology Division P.O. Box 2728 Pittsburgh, Pennsylvania 15230-2728 -

. e 1992 Westinghouse Electric Corp.

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TABLE OF CONTENTS

1.0 INTRODUCTION

1.1 Code Acceptance Criteria 1.2 Geometry 2.0 LOADING CONDITIONS. FRACTURE ANALYSIS METHODS. AND MATERIAL PROPERTIES 2.1 Transients 2.2 Stress Intensity Factor Calculations 2.3 Fracture Toughness 2.4 Thermal Aging 2.5 Allowable Flaw Size Calculation 3.0 SUBCRITICAL CRACK GROWTH 3.1 Analysis Methodology '

3.2 Crack Growth Rate Reference Curves i 3.3 Residual Stresses 3.4 Stress Corrosion Cracking Susceptibility 4.0

SUMMARY

AND RESULTS 4.1 Flaw Evaluation Charts Construction 4.2 Conservatisms in the Flaw Evaluation l l

5.0 REFERENCES

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SECTION 1.0 '

INTRODUCTION This fracture mechanics evaluation has been carried out to determine the l

largest size of indications which can he accepted according to the rules of Section XI, paragraph IW8 3600 for the residual heat exchanger inlet and outlet nozzles. The results of this evaluation are presented in flaw evaluation charts in Section 4, and the technical basis.for construction of the charts is contained in the remaining sections. i 1.1 Code Accentance criteria The evaluation procedures and acceptance criteria for indications in  !

austenitic stainless piping are contained in paragraph IWB 3640 of ASME  !

SectionXI.[1] The evaluation procedure is applicable to all the materials l within a specified distance from the weld centerline, vrt, where r - the pipe nominal outside radius and t is the nominal wall thickness. For example, 1st the RHX nozzle, this distance is calculated to be 1.62 inches,'which i encompasses regions of the heat exchanger, as well as part of the RHR line. l All the materials in this region are Type 304 stainless steel.

l The evaluation process begins with a flaw growth analysis, with the {

requirement to consider growth due to both fatigue and stress corrosion I cracking. For pressurized water reactors only fatigue crack growth need be considered, as discussed in section 3. The methodology for the fatigue crack growth analysis is described in detail in section 3. .

i The calculated maximum flaw dimensions at the end of the evaluation period are then compared with the maximum allowable flaw dimensions for both normal operating conditions and emergency and faulted conditions, to determine '

acceptability for continued service. Provisions are made for considering flaws projected both circumferentially and axially.

In IWB 3640 the allowable flaw sizes have been defined in the tables based on maintaining specified safety margins on the loads at failure. These margins are 2.77 for normal and upset conditions and 1.3g for emergency and faulted wnamumamos 1

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conditions. The calculated failure loads are different for the base metal and the flux welds, which have different fracture toughnass values, as discussed The failure loads, and consequently the allowable flaw sizes, in section 2. Allowable flaw sizes for are larger for the base metal than for the welds.

welds are contained in separate tables, in IWB 3640.

- 1.2 Geometry The geometry of the residual heat exchanger is shown in Figure 1-1, with the details of the inlet and outlet nozzles of the tube side shown in Figure 1-2. -

The notation used for surface flaws in this work is illustrated in figure 1-3.

The fracture and fatigue crack growth evaluations carried out to develop the handbook charts have employed the recommended procedures and material properties for stainless steel as prescribed in paragraph IWB 3640 and Appendix C of Section XI.

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SECTION 2.0  ;

LOAD CONDITIONS, FRACTURE ANALYSIS METHODS AND MATERIAL PROPERTIES The loading conditions used in the analyses described herein were taken directly from the equipment specification. The fracture analysis methods are I the most advanced which are now available, and the material properties are the latest available properties contained in the ASME Code.

2.1 Transients and Load Conditions The design transients for the residual heat exchanger are very minimal, because this component operates only during plant shutdown conditions.

Therefore the only transient conditions which it experiences are the startup and shutdown of th- e tem, which coincides with the shutdown and startup of the plant, respectively. The appropriate limiting load conditions for the location of interest are discussed next.

The loading ecnditions which were evaluated include thermal expansion (normal and upsat), pressure, deadweight and seismic (OBE and SSE) loadings. The RHR piping forces and moments for each condition were obtained from the ASME Code Section III calculations previously performed by Sargent and Lundy and Westinghouse for Byron and Braidwood Units 1 and 2 [2-5]. These loads [6]

were compared with the Equipment Specification design loadings for the heat exchanger nozzles (G-679150 Rev.1) and found to be bounded by them. As a consequence of this comparison, the evaluation perfonned using the design loadings, is applicable to Byron and 3raidwood Units .1 and 2. Residual  !

stresses were not used in this porti::n :f the evaluation, in compliance with the Code guidelines. A further discu::::n of residual stresses is contained I in Section 3.2. The stress intensit; ~ 1as were calculated using the following equations:

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where F, - axial force component (membrane)

M,, My , M, - moment components (bending)

A = cross-section area Z = section modulus ,

-The section properties A and I at the weld location were determined based on the minimum pipe dimensions. This is conservative since the measured wall i

thickness at the weld is generally larger. I i

The following load combinations were used.

A. Normal / Upset - Primary Stress Pressure + Deadweight + OBE i B. Emergency / Faulted - Primary Stress .

Pressure + Deadweight + SSE i

j C. Normal / Upset - Total Stress Pressure + Deadweight + OBE + Normal Thermal 1

0. Emergency / Faulted - Total Stress i'

Pressure + Deadweight + SSE + Normal Thermal l

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2.2 Stress Intensit'y Factor Calculations Dne of the key elements of the fatigue crack growth calculations is the determination of the driving force or stress intensity factor (Kg ). This was done using expressions available from the literature. In all cases the stress intensity factor calculations utilized a representation of the actua.1 stress This was necessary to provide the most profile rather than a linearization.

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The stress profile was. represented by a accurate determination possible. i cubic polynomial:

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where x is the cor.rdinate distance into the wall l

t = wall thickness i a a stress perpendicular to the plane of the crack  !

A, = coefficients of the cubic fit I For the surface flaw with length six times its depth, the stress intensity '

The stress factor expression of [McGowan and Raymund [7))C was used.

The intensity factor Ky ($) can be calculated anywhere along the crack front. '

The following ,

point of maximum crack depth is represented by $ = 0. ,

expression is used for calculating Kg (4), where d is the angular location around the crack.

• 0.5 (2-2)

K.($) = (cos8$ + sin 4) m ( AoHo + AH3

• AH* AH) 2 The magnification factors H,($), H (o) 3

((o) and H3 (p) are obtained by the procedure outlined in refernce [8). l The stress intensity factor calculi:- ir a semi-circular surface flaw,

.ie expressions developed by [Raju (aspect ratio 2:1) was carried out Their expres:

'izes the same cubic representation and Newman [8)).

he same result as the expression of of the stress profile and gives pr-  !

[McGowan and Raymund] for the

ratio flaw, and the form of the Raymund]# above.

equation is similar to that of [M -

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The stress intensity factor expression used for a continuous surface flaw was  !

Again the stress profile is that developed by [Buchalet and Bamford [9))*****.

represented as a cubic polynomial, as shown above, and these coefficients as ,

well as the magnification factors are combined in the expression for K, (2-3)

I,= {is AA + **-hFs + fhFs + ha'hF, 1

where F,, F,, F3 , F, are magnification factors, available in [9]. i 2.3 Fracture Touchness The weld at the The residual hsat exchanger is stainless steel type 304. l nozzle was made by the shielded metal are process, as verified by the shop traveller, and the weld procedure referenced therein.

The fracture toughness of the base metal h&s been found to be very high,.even l

at operating temperatures (10), where the J,, values have been found to be 2 Fracture toughness values for weld materials have well over 2000 in-1b/in .

been found to display much more scatter, with the lowest reported values Although the J,, values significantly lower than the base metal toughness.

reported have been lower, the slope of the J-R-curve is still large for these J ge cases. Representative values for J,, were obtained from the results of Landes, et. al. [11], where the following values were obtained, and used in the development of the fracture evaluation methods- 5 i

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[ for shielded metal arc welds: J,, = 990 in 1b/in ,)..e,e 2.4 THERMAL AGING Thennal aging at operating temperatures of reactor primary piping can reduce the fracture toughness of cast stainless steels and, to a lesser degree, Because of the lower operating temperature (400*F) stainless steel weldrents.

of the residual heat exchanger, and the fact that the materials are type 304 ~

stainless (not cast), thermal aging in this component will be negligible.

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2.5 Allowable Flaw size Determination The critical flaw size is not directly calculated as part of the flaw '

evaluation process for stainless steels. Instead, the failure mode.and critical flaw size are incorporated directly into the flaw evaluation technical basis, and therefore into the tables of " Allowable End-of-Evaluation

- Period Flaw Depth to Thickness Ratio," which are contained in paragraph IW8 l 3640.

Rapid, nonductile failure is possible for ferritic materials at low ,

temperatures, but is not applicable to stainless steels. In stainless steel materials, the higher ductility leads to two possible modes of failure, plastic collapse or unstable ductile tearing. The second mechanism can occur when the applied J integral exceeds the JIc fracture toughness, and some l j

stable tearing occurs prior to failure. If this made of failure is dominant, the load carrying capacity is less than that predicted by the plastic collapse f l

mechanism. -

The allowable flaw sizes of paragraph IWB 3640 for the high toughness base  ;

I materials were determined based on the assumption that plastic collapse would be achieved and would be the dominant mode of failure. [However, due to the l reduced toughness of the shielded metal arc welds, it is possible that crack  ;

extension and unstable ductile tearing could occur and be the dominant mode of l failure. This consideration in effect reduces the allowable end of interval

' flaw sizes for flux welds relative to :he austenitic wrought type 304 vessel

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and piping materia 1s, and has been inc:roorated directly into the evaluation ,

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l SECTION 3.0 FATIGUE CRACK GROWTH In applying Code acceptance criteria as introduced in section 1, the final flaw size af is defined as the flaw size to which the detected flaw is calculated to grow at the end of a specified period, or until the next

- inspection time. This section will examine each of the calculations, and i provide the methodology used as well.as the assumptions.

l 3.1 Analysis Methodoloav l

l The methods used in the crack growth analysis reported here are the same as those suggested by Section XI'of the ASME Code. The analysis procedure i involves postulating an initial flaw at specific regions and predicting the growth of that flaw du: to an imposed series of loading transients. The input required for a fatigue crack growth analysis is basically the information  !

necessary to calculate the parameter AKg which depends on crack and structure l geometry and the range of applied stresses in the area where the crack exists. ,

once AK; is calculated, the growth due to that particular stress cycle can be I calculated by equations given in section 2.2 and figure 3-1. This increment

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of growth is then added to the original crack size, and the analysis proceeds  !

to the next transient. The procedure is continued in this manner until all the transients known to occur in the period of evaluation have been analyzed.

5 The only transients considered in the analysis were the startup and shutdown of the RHR system. These transients are spread equally over the design lifetime of the vessel. '

Crack growth calculations were carried out for a range of flaw depths, and e

,three basic types. The first two were surface flaws, one with length equal to '

six times the depth and another with length equal twice the depth. Third was t a continuous surface flaw, which represents a worst case for surface flaws.

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}- 4 3.2 Crack Growth Rate Reference Curves The reference crack growth law used for the stainicss steel was taken from that developed by the Metal Properties Council - Pressure Vessel Research [

' Committee Task Force in Crack Propagation Technology. The reference curve has f

the equation:

= CFS AK8 (3-7)

L where - = crack grcwth rate, inches per cycle l

C = material coefficient (C - 2.0 x 10-18) ,

F = frequency coefficient for loadings (F = 2.0) 2 S = R ratio correction coefficient (S = 1.0 = 0.502 R )-4.0 material property slope (=3.0321) n =

AK = stress intensity factor range, psi /in This equation appears in Section XI, Appendix C (1989 Addendum) for air environments and its basis is provided in reference [12), and shown in figure 3-1. . For water environments, an environmental factor of 2 was used, based on the crack growth tests in PWR environments reported by Bamford [13].

3.3 Residual Stresses Since the residual heat exchanger ves:si to-piping welds were not stress-relieved, residual stresses srs :learly present. For fatigue crack growth analyses, these stresses wer1 ::.uded directly.

l In general the distribution of res- :resses is strongly dependent.on the degree of-constraint of the struer i stiffer the structure the higher the residual stresses. For a thir. 'arge diameter pipe the residual stresses will be lower than a smai ar thick-walled pipe. This has been found by a number of investigatort 1 is general agreement that the

~

3 near the surface, and then distribution of residual stresses WPf126CJ/03277210 \

- . - . - - - ,- _ , -- - -- .. - . . . ._. i

. 1

~

compressive near the center of the wall after which it reverses to become tensile at the outer surface. ,

The residual stresses were taken from work reported by General Electric [14]  ;

and E. Rybicki (15), which included both measurements of residual stress and j finite element calculations. Both approaches were found to be in agreement, and included a range of pipe sizes from 4 inches to 28 inches in diameter.

The stresses were found to peak at the weld, as shown in Figure 3-2 for a 10 inch diameter pipe. the through wall distribution of residual stresses used in this ar.alysis was taken from the work of Rybicki, and is shown in Figure 3- l

3. This distribution is for a 10 inch schedule 160 pipe with a thickness of l 1.125 inches, which is a much stiffer configuration than the 14 inch diameter, 0.375 inch thick junction at the heat exchanger nozzle.

l 3.4 Stress Corrosion Crackino Suscentibility In evaluating flaws, all mechanisms of subcritical crack growth must be evaluated to ensure that proper safety margins are maintained during service.

Stress corrosion cracking has been observed to occur in stainless steel in operating BWR piping systems. The discussion presented here is the technical basis for not considering this mechanism in the present analysis. The resioual heat exchanger tube side nozzles are exposed to only primary coolant water.

For all Westinghouse plants, there is no history of cracking failure in the reactor coolant system loop piping. For stress corrosion cracking (SCC) to occur in piping. the following three conditions must exist simultaneously:

high tensile stresses, a susceptible material, and a corrosive environment.

Since some residual stresses and some degree of material susceptibility exist in any stainless steel piping._ the potential for stress corrosion is minimized by proper selection of a material immune to SCC as well as preventing the occurrence of a corrosive environment. The material specifications consider compatibility with the system'.s operating environment (both internal and external) as well as other materials in the system, applicable ASME Code rules, fracture toughness, tolding, fabrication, and processing. -

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. Environments [12).

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l

i e e j SECTION 4.0 '

l 4.1 Flaw Evaluation Charts Construction I

The acceptance criteria for surface flaws have been presented in Section 1.

For flaw evaluation in stainless steels, only the fatigue crack growth results must be calculated. The allowable flaw depths were determined directly from  !

the tables in IWB 3640.

h i

The first set of data required for surface flaw chart construction is the final flaw size a,. As defined in IWB-3611 of ASME Code Section XI, a, is the flaw depth resulting from growth during.a specific time period, which can be the next scheduled inspection of t'he component, or until the end of design lifetime. -Therefore, the final depth, a, after a specific service period of time must be used as the basis for evaluation.

The final flaw size a, can be calculated by fatigue crack growth analysis, which has been. performed covering a range of postulated flaw sizes, and flaw' shapes.  ;

The crack growth calculational methods have been discussed in Section '

3.

The results of the crack growth calculation showed that growth for a complete range of crack sizes, up to 60 percent of tt;e wall thickness was inconsequential for the entire service life.of 40 years. This was expected, '

since the region sees so few cycles.

The allowable flaw size for stainless steel is obtained directly from tables in paragraph IWB 3640, so the evaluation process is very straight forward.

The allowable flaw size is calculated based on the most limiting transient for i

' all normal operating conditions. Similarly, the allowable flaw size for emergency and faultad conditions is determined. The theory and methodology for the calculation ut the allowable flaw sizes have been provided in Section 2 and Reference 16. Allowable flaw sizes were calculated for a range of flaw shapes.

The two basic dimensionless parameters, which can fully address the characteristics of surface flaw, have been used for the evaluation chart construction. Namely, wetuounsmano 17

r=

o Flaw Length divided by the circumference, t/c o  :

Flaw Depth parameter a/t I where, i

t - wall thickness, in.

a - flaw depth, in.

t -

flaw length, in.

c =

pipe circumference, in.

The flaw evaluation chart for the residual heat exchanger inlet and outlet nozzles is shown in Figure 4-1. The chart has the following characteristics:

o The flaw length / circumference t/c was plotted as the abscissa from 0 to .5.

For values of t/c which exceed 0.5, use the results for t/c

= 0.5.

o ^

The flaw depth parameter a/t was plotted as the ordinate.

o The upper boundary curve shows the maximum acceptable flaw depth based on flaw evaluation, beyond which no surface flaw is acceptable for continued service without repair. This upper bound curve has been determined by the fracture and fatigue evaluations described j

herein. using Tables IWB 3641-5 and IWB 3641-6, for shielded metal '

arc welds.

o i Any surface indication which Falls below the boundary curve will be '

acceptable by the code rules. based on the analytical justification provided herein. However, :'.id-2420 of ASME Section XI requires future monitoring of suci  ;' stions.

A detailed example on the use of thw a for a surface flaw is presented below:

e WPF12MJn32Mh to

Surface Flaw Er== ale  !

Now suppose an indication.is to be evaluated u' sing the charts. For the circumferential orientation:

a = 0.10" t = 6.1" t = 0.40' c = 44.0"

(

The flaw characterization parameters then become:

a/t = 0.250 t/c = 0.139 t

Plotting these parameters on the surface flaw evaluation chart of Figure 4-1, it is quickly seen that the indication is acceptable. i 4.2 Conservatirms in the Flaw Evaluation i

The stress and fracture analysis results presented herein have been structured to be conservative at each step, to ensure that the final result will be conservative.

The stresses appiled to the heat exchanger nozzles were taken from the vessel

' equipment specification loads, which represent bounding loads for the structure. The actual loads for the Byron and Braidwood Units I and 2 heat exchangers [6] are about 60 percent of the design loads The residual stresses used in the analysis were taken from a combination of measurements and analysis for a lo inch schedule 160 pipe. The smaller pipe diameter and larger thickness (1.125 inches) for this pipe mean that the residual stress distribution used here will be very conservative relative to the heat exchanger nozzle.

Since the publication of the flaw evaluation criteria and methodology for stainless steel [16] a number of experiments have been carried out on large frac.ture toughness specimens and full size pipes with both submerged are welds WF1360J/1BTM2310 19 I

I l

._ l and shielded metal arc welds [17]. These experiments have shown that the 1 fracture toughness from these larger specimens is higher than the toughness '

values used in the development of the flaw evaluation methods. Therefore the flaw evaluation results presented here are conservative. ,

A further conservatism is added to this fracture evaluation by using the fracture criteria for a class I piping system for a class 2 component. There are presently no flaw evaluation criteria for class 2 components, but presumably if they were to be developed, smaller margins could be justified, with resulting larger allowable flaw sizes.

The indication depths from the inspections have been compared with the thickness of the pipe, with no benefit taken of the additional thickness resulting from the large fillet weld on the outside surface of the nozzle. As shown in Figure 1-2, this fillet weld is immediately above the indications, and so the actual percentage flaw penetration is smaller than that reported.

O

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0 0.1 0.2 0.3 0.4 0.5 FLAW LD'CTH / CIRCUMFERDfCE Figure 4-1 Flaw Evaluation Chart for Byron and Braidwood Units 1_ and 2

. Residual Heat Exchanger Tube Side Nozzles W 1240J/032772210 21

?

SECTION

5.0 REFERENCES

1. ASME Code Section XI, " Rules for Inservice Inspection of Nuclear Power Plant Components," 19B3 edition (used for updated code allowable limits); 1983 edition, Winter 1985 Addendum (used for flaw evaluation of  :

austenitic stainless steel piping); 1989 edition (used for reference j crack growth curve, stainless steel). '

2. Jambusaria, H., " Residual Heat Exchangers: Braidwood Unit 2,"

Westinghouse Report No. 031804 Rev. O, Jan. 24, 1986.  ;

3. Jambusaria, H., " Residual Heat Exchangers: Byron Unit 1," Westinghouse Report No. 031805 Rev. 0, 2/27/92.

4. Jambusaria, H., " Residual Heat Exchangers: Byron Unit 2," Westinghouse Report No. 031806 Rev. O, 2/27/92.

l

5. Jambusaria, H., " Residual Heat Exchangers: Braidwood Unit 1,"

Westinghouse Report No. 031803 Rev. O, 2/27/92.

6. Letter # BPM #1577 from D. J. Skoza of Commonwealth Edison Company to Janet Bunecicky of Westinghouse Electric Corporation,

Subject:

RHR Heat Exchanger Nozzle Loads, dated 2/12/92.

7. McGowan, J. J. and Raymund, M., " Stress Intensity Factor Solutions for Internal Longitudinal Semi-elliptic Surface Flaw in a Cylinder Under Arbitrary Loading", ASTM STP 677, 1979, pp. 365-380.

1

8. Newman, J. C. J'r. 'and Raju, I. S., " Stress Intensity Factors for  !

Internal Surface Cracks in Cylindrical Pressure Vessels", ASME Trans.,

Journal of Pressure Vessel Technology, Vol. 102, 1980, pp. 342-346.

9. Buchalet, C. B. and Bamfor'd, W. M.. " Stress Intensity Factor Solutions for Continuous Surface Flaws in 1: actor Pressure Vessels", in Mechanics

, ,of Crack Growth, ASTM, STP 590, .:75, pp. 385-402.

WPF1260J/G32772:13 7.

l

10. Bamford, W. H. and Bush, A. J., " Fracture of Stainless Steel," in ',

Elastic Plastic Fracture, ASTM STP 668, 1979. I

11. Landes, J. D., and Norris, D. M., " Fracture Tou0 ness of Stainless Steel Piping Weldsents," presented at ASME Pressure Vessel Conference, 1984.

1

- 12. James, L. A., and Jones, D. P., " Fatigue Crack growth Correlations for ,

Austenitic Stainless Steel in Air," in Predictive Canabilities in i Environmentally Assisted Crackina," ASME publicttion PVP-99, Dec.1985.

13. Bamford, W. H., "Fatighe Crack Growth of Stainless Steel Piping in a  !

Pressurized Water Reactor Environment," Trans ASME, Journal of Pressure '

Vessel technology, Feb.1979. .

14. " Studies on AISI Types 304, 304L, and 347 Stainless Steels for BWR i Application, April-June 1975," General Electric Report NEDO-20985-1, I September 1975.

15. Rybicki, E. F., McGuire, P. A., Merrick, E., and West, J., "The Effect of Pipe Wall Thickness on Residual Stresses Due to Girth Welds," IC101 A$ti, Journal of Pressure Vessel Technology, Vol 104, August 1982.

16. " Evaluation of Flaws in Austenttic Steel Piping," Trans ASME, Journal of Pressure Vessel Technology, Vol. 108, Aug. 1986, pp. 352-366.

l

17. Wilkowski, G. et. al., " Analysis of Experiments on Stainless Steel Flux Welds," Battelle Columbus labs report for USHRC, number NUREG/CR 4878, April 1987.

E I

i tM 1240J/Os2772:10 23

, _,_ , . - . , _ __ _ . .