L-09-102, Design Rept for Recirculation Sys Weld Overlay Repairs at Brunswick Steam Electric Plant,Unit 1
| ML20076G109 | |
| Person / Time | |
|---|---|
| Site: | Brunswick  |
| Issue date: | 05/05/1983 |
| From: | Russell Adams, Charnley J, Reeves P NUTECH ENGINEERS, INC. |
| To: | |
| Shared Package | |
| ML20076G103 | List:
|
| References | |
| CPL-09-102, CPL-09-102-R00, CPL-9-102, CPL-9-102-R, NUDOCS 8306140549 | |
| Download: ML20076G109 (75) | |
Text
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CPL-09-102 Revision 0 May 1983 CPLOO9.0102 DESIGN REPORT .
FOR RECIRCULATION SYSTEM ;
WELD OVERLAY REPAIRS AT BRUNSWICK STEAM ELECTRIC PLANT UNIT 1 Prepared for Carolina Power and Light Company Prepared by NUTECH Engineers, Inc.
San Jose, California Prepared by: Reviewed by:
F C & flA,%
dJ. E. Charnley, /P.E. P. E. Reeves Project Engineer Project Ouality Assurance Engineer Approved by: Issued by:
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R. H. Adams, P.E. N. Eng '
Director Project Manager Date: % [. /f83 8306140549 830516 DR ADOCK 05000
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REVISION CONTROL SHEET l
TITLE: Design Report for Recirculation DOCUMENT FILE NUMBER: CPLOO9.0102 System Weld Overlay Repairs at Brunswick Steam Electric Plant Unit 1 J. E. Charnley / Principal Engineer v
A. Rao /Cen sult ant II kkk P. R. Para / Consul' tant I M iN J. R. Tavlor/ Consultant II M
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REVISION CONTROL SHEET (CONTINUATION)
TITLE: Design Report for Recirculation DOCUMENT FILE NUMBER: CPLOO9.0102 System Weld Overlay Repairs at Brunswick Steam Electric Plt Unit 1 APPECTED OCC PREPARED ACCURACY CRITERIA REMARKS PAGElS) REV av/DATE CHECK BY / OATE CHECK sv/OATE 23 0 0 $0 \b 'N S&*O )/$$Sf~fS) ~
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CERTIFICATION BY REGISTERED PROFESSIONAL ENGINEER I hereby certify that this document and the calculations
,o contained herein were prepared under my direct supervision, reviewed by me, and to the best of my knowledge are correct and complete. I am a duly Registered Professional Engineer under the laws of the State of California and am competent to review this document.
Certified by:
} r ;_ 'y J. E. Charnley Professional Engineer State of California Registration No. 16340 Date
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TABLE OF CONTENTS Page LIST OF TABLES vii LIST OF FIGURES viii
1.0 INTRODUCTION
1 2.0 REPAIR DESCRIPTION 4 3.0 EVALUATION CRITERIA 9 3.1 Strength Evaluation 10 3.2 Fatigue Evaluation 10 3.3 Crack Growth Evaluation 11 4.0 LOADS 12 4.1 . Mechanical and Internal Pressure Loads 12 4.2 Thermal Loads 12 5.0 EVALUATION METHODS AND RESULTS 14 5.1 Safe End Evaluation 14 5.1.1 Code Stress Analysis 15 5.1.2 Fracture Mechanics Evaluation 17 5.2 12" Elbow Evaluation 22 5.2.1 Code Stress Analysis 22 5.2.2 Fracture Mechanics Evaluation 24 5.3 28" Elbow Evaluation 26 5.3.1 Code Stress Analysis 27 5.3.2 Fracture Mechanics Evaluation 28 i
5.4 Effect on Recirculation and RHR Systems 31 6.0 LEAK-BEFORE-BREAK 50 6.1 Net Section Collapse 50 6.2 Tearing Modulus Analysis 51 6.3 Leak Versus Break Flaw Configuration 52 6.4 Axial Cracks 53 6.5 Multiple Cracks 54 6.6 Crack Detection Capability 54 l
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4 TABLE OF CONTENTS (Continued)
Pace 6.7 Non-Destructive Examination 55 6.8 Leakage Detection 56
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6.9 Historical Experience 57
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7.0
SUMMARY
AND' CONCLUSIONS 63
8.0 REFERENCES
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LIST OF TABLES Number Title Pace 5.1 Thermal Stress Results 33 5.2 Safe End Code Stress Results 34 5.3 12" Elbow Code Stress Results -
35 5.4 28" Elbow Code Stress Results 36 6.1 Effect of Pipe Size on the Ratio 58 of the Crack Length for 5 GPM Leak Rate and the Critical Crack Length (Assumed Stress s = (Sm)/2 )
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LIST OF FIGURES Number Title Pace 1.1 Conceptual Drawing of Recirculation System 3 2.1 Configuration of Safe End-to-Pup Piece 6 _
Weld Overlay 2.2 Configuration of 12" Elbow-to-Pipe Weld Overlay 7 2.3 Configuration of 28" Elbow-to-Pipe Weld Overlay 8 5.1 Safe End Finite Element Model 37 5.2 Weld Overlay Thermal Model 38 5.3 Thermal Transients 39 5.4 Safe End IGSCC Susceptibility 40 5.5 Typical IGSCC Crack Growth Data 41 (Weld Sensitized 304SS in BWR Environment) 5.6 Safe End Tearing Modulus 42 5.7 12" Elbow Finite Element Model 43 5.8 12" Elbow Tearing Modulus 44 5.9 28" Elbow-to-Pipe Finite Element Model 45 5.10 28" Elbow-to-Pipe Circumferential Weld 46 Residual Stress 5.11 28" Elbow-to-Pipe Crack Growth 47 5.12 28" Elbow-to-Pipe Tearing Modulus 48 5.13 Piping Model 49 6.1 Typical Result of Net Section Collapse 59 Analysis of Cracked Stainless Steel Pipe 6.2 Stability Analysis for BWR Recirculation 60 System (Stainless Steel) 6.3 Summary of Leak-Before-Break Assessment 61 of BWR Recirculation System 6.4 Typical Pipe Crack Failure Locus for Combined 62 Through-Wall Plus 360* Part-Through Crack CPL-09-102 viii Revision 0 nutp_qb
1.0 INTRODUCTION
This report summarizes evaluations performed by NUTECH to assess weld overlay repairs in the Recirculation System at Carolina Power and Light Company's Brunswick Steam Elehtric Plant Unit 1 (Brunswick 1) . Weld overlay repairs have been applied to address leakage and additional ultrasonic (UT) examination results believed to be indicative of intergranular stress corrosion cracking (IGSCC) in the vicinity of the welds. The purpose of each overlay is to arrest any further propagation of the cracking, and to restore original design safety margins to the weld.
The required design life of each weld overlay repair is at least five years. The amount that the actual life exceeds five years will be established by a combination of future analysis and testing.
Leakage was obaerved adjacent to one recirculation inlet safe end to pup piece weld and adjacent to one 12" vertical riser to elbow weld. In addition, cr'ack indications have been detected adjacent to one 28" elbow CPL-09-102 1 Revision 0 i
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to vertical pipe weld. All three of these welds were repaired with weld overlay designs evaluated in this report.
Figure 1.1 shows the safe end and the two elbow to pipe welds in belation to the Reactor Prissure Vessel and other portions of the Recirculation and RHR Systems.
All of the existing Recirculation System material is type 304 stainless steel.
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$n INDICATIONS IN SAFE END TO PUP PIECE WELD
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LOOP B 60 LOOP A D Figure 1.1 CONCEPTUAL DRAWING OF RECIRCULATION SYSTEM b
2.0 REPAIR DESCRIPTION The through-wall cracks and other indications in the existing safe end and elbow weld heat-affected zones have been repaired by establishing additional " cast-in-place" pipe wall thickness from weld metal deposited 360 degrees around and to either side of the existing weld, as shown in Figures 2.1, 2.2, and 2.3. The weld deposited band over the cracks will provide wall thickness equal to that required to provide the original design safety margins. In addition, the weld metal deposition will produce a favorable compressive residual stress pattern. The deposited weld metal will be type 308L, which is resistant to propagation of IGSCC cracks.
The non-destructive examination of the weld overlays consisted of:
- 1) Surface examination of the completed weld overlay by the liquid penetrant examination technique in accordance with ASME Section XI.
- 2) Volumetric examination of the completed weld overlay by the ultrasonic examination technique in accordance with ASME XI.
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- 3) Volumetric preservice examination of the weld overlay and existing circumferential pipe weld by the ultrasonic examination technique in accordance .
with ASME Section XI.
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TYPE 30$L WELD CVERLAY 2.5" MIN -
a a 12", 0.57" THICK WALL. 3M SE PtPE Q.35" MIN ~
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ASWELDED SURPACE ACCEPTA8LE FOR OVERLAY TAPER TRAN$ MONS Figure 2.2 l CONFIGURATION OF 12" ELBOW-TO-PIPE WELD OVERLAY PATENT APPLIED POft CPL-09-102 Revision 0 7 nutg_qh
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Figure 2.3 CONFIGURATION OF 28" ELBOW-TO-PIPE WELD OVERLAY PATENT APPUED POR CLP-09-102 Revision 0 8 nutp_qh l
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3.0 EVALUATION CRITERIA This section describes the criteria that are applied in this report to evaluate the acceptability of the weld -
overlay repairs described in Section 2.0. Because of
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the nature of these repairs, the geometric configuration
, is not directly covered by Section III of the ASME Boiler and Pressure Vessel Code, which is intended for new construction. However, materials, fabrication procedures, and Quality Assurance requirements are in accordance with applicable sections of this Construction Code, and the intent of the design criteria described below is to demonstrate equivalent margins of safety for strength and fatigue considerations as provided in the ASME Section III Design Rules. In addition, because of the IGSCC conditions that led to the need for repairs, IGSCC resistant materials have been selected for the weld overlay repairs. As a further means of ensuring structural adequacy, criteria are also provided below for fracture mechanics evaluation of the repairs.
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3.1 Strength Evaluation Adequacy of the strength of the weld overlay repairs with respect to applied mechanical loads is demonstrated with the following criteria:
- 1. An ASME Boiler and Pressure Vessel Code Section III, Clars 1 (Reference 1) analysis of the weld overlay repairs was performed.
- 2. The ultimate load capacity of the repairs was calculated with a tearing modulus analysis. The ratio between failure load and applied loads was required to be greater than that required by Reference 1.
3.2 Fatique Evaluation The stress values obtained from the above strength evaluation were combined with thermal and other '
secondary stress conditions to demonstrate adequate f atigue resistance for the design life of each repair.
The criteria for fatigue evaluation include:
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- 1. The maximum range of primary plus secondary stress was compared to the secondary stress limits of Reference 1.
- 2. The peak alternating stress intensity, including all hrimary and secondary stress terms, and a fatigue strength reduction factor of 5.0 to account for the existing crack, was evaluated using conventional fatigue analysis techniques. The total f atigue usage factor, defined as the sum of the ratios of applied number of cycles to allowable number of cycles at each stress level, must be less than 1.0 for the design life of each repair.
Allowable number of cycles was determined from the stainless steel fatigue curve of Reference 1.
3.3 Crack Growth Evaluation Crack growth due to both fatigue (cyclic stress) and IGSCC (steady state stress) was calculated. The allowable crack depth was established based on net section limit load for each cracked a 2 repaired weld (Reference 2). The design life of each repair was established as the minimum of either the predicted time for the observed crack to grow to the allowable crack
, depth or five years.
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J 40 LOADS The~1oads considered in the evaluation of the safe end and elbow welds consint of mechanical loads, internal _
pressure, differential thermal expansion loads,:and welding'r'esidual stresses. The mechanicalsloads and internal pressures used in"the analysis are described in Section 4.1, and an explanation of the thermItl transient conditions which cause differential thermal expansion loads is pr'esented in Section 4.2. Welding residual stresses are considered in the crack growth analyses and are described in Nections 5.1.2.2 and 5.3.2'.'2'.[
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composite transient is a startup/ shutdown transient with a heatup or cool down rate of 100*F per hour. The second composite transient consists of a 50*F step temperature with no change in pressure. The third composite transient is an emergency event with a 416'F step temperature change and a pressure change of 1325 psi. In the five year overlay design life, there are 38 startup/ shutdown cycles, 25 small temperature change cycles, and one emergency cycle.
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F7 5.0 EVALUATION METHODS AND RESULTS (,
The evaluation of the weld overlay repairs consist of a code stress analysi. sper Section III (Reference 1) and a .
fracture mechanica evaluation per Section XI (Referenc'e 6).
l 5.1 Safe End Evaluation L ,
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Tre heat affected zone on the piping side of the weld between the N2H recirculation' inlet' safe end extension ,
(pup piece) and the recirculation pipin'g had a small i
through wall crcck. This weld '(12-BR-M4)'is between the a
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i 104 stainless steel. piping and a lov carbon content (0.03 weight percent) stainless steel pup piece. It is close to the NiCrFe safe end (Figure 2.1). Due to the
. proximity of the NiCrFe safe end, a nonstandard weld
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overlay was designed. The weld overlay is shorter than I
i a standard weld overlay on the pup piece side of the
. 12-BR-H4 weld. The short weld overlay was judged to be acceptable based on the low carbon content, and thus high resistance to IGSCC cracking, of the pup piece.
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5.1.1 Code Stress Analysis The ASME Code stress analysis of the cracked and repaired 12-BR-H4 weld was performed with an ANSYS _
(Reference 7) finite element model. The model was based
, on an overlay thickness of 0.20 inch, which is 50 percent smaller than the actual minimum average thick-
. ness of 0.40 inch. The small added weight (less than 25 pounds) of the weld overlay will not significantly change the existing seismic analysis. Figure 5.1 shows i.
3
, the model. The stress in the overlaid safe end due to
/ design pressure and applied moments as described in
) Sections 4.1 and 4.2 was calculated with the finite L
element model.
The weld overlay thermal model was taken to be
, axisymaetrical (Figure 5.2) . The exterior boundary was assumed to be insulated. The temperature distribution in the weld overlay subject to the thermal transients
, defined in Section 4.2 can be readily calculated using Charts 16 and 23 of Reference 8. The maximum through I ( wall temperature difference was determined to be less !
than 2*F for the normal startup cycle, 40*F for the small temperature cycle and 329'F for the emergency transient.
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_ _ _ . - _ _. _ . _ = _ _ _. _ _ _ _ . -. . __ _ _ _ . _ _ . _ _
The maximum thermal stress for use in the fatigue crack growth analysis was calculated as follows:
(Reference 1)
. Ea aT1 Ea oT 2 -
i 2 (1- v) 1-v Where E = 28.3 x 10 6 psi (Young's Modulus) a = 9.11 x 10-6.y-1 (Coef ficient of Thermal Expansion) ar y = Equivalent Linear Temperature Difference ar 2 = Peak Temperature Difference The values of dry , cr2 , and o are given in Table 5.1 for all three thermal transients.
The results of a code stress analysis per Reference 1 are given in Table 5.2. The allowable stress values for Reference 1 are also given. The weld overlay repair satisfies the Reference 1 requirements.
A conservative fatigue analysis per Reference 1 was performed. In addition to the stress intensification CPL-09-102 16 Revision 0 nutagh i
I factors required per Reference 1, an additional fatigue strength reduction factor of 5.0 was applied due to the crack. The fatigue usage factor was then calculated assuming 38 startups, 25 small temperature change cycles ,
and one emergency cycle every five years. The results
~
are summcrized in Table 5.2.
5.1.2 Fracture Mechanics Evaluation Three types of fracture mechanics evaluations we're performed. The allowable crack depth was calculated based on Reference 2. Crack growth due to both fatigue and IGSCC was calculated using the NUTECH computer program NUTCRAK (Reference 9) with material constants and methodology from References 10 and 11. Finally, the ultimate margin to failure for a crack assumed to propagate all the way through the original pipe material to the weld overlay was calculated per References 12 and 13.
5.1.2.1 ' Allowable Crack Depth The allowable depth for a 1/2 inch long crack was determined using Reference 2. Since the crack orienta-tion (axial cr circumferential) was not established by CPL-09-102 17 Revision 0 nutggb
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the UT examination, the allowable depth was calculated for both orientations. The as-built dimensions of the i
repaired safe end were used. The minimum average overlaid wall thickness is 0.97 inch. The allowable crack depth for an axial crack of length equal to or less than'l/2 inch (the maximum length of the heat affected zone) is 75% of the overlaid 12-BR-H4 weld thickness. Thus, the allowable crack depth for an axial crack is (0.75) (0.97) or ,0.73 inch, which is greater than the original pipe thickness.
The allowable crack depth for a 360* circumferential crack in the overlaid 12-BR-H4 weld was determined based on References 2 and 11. The allowable crack depth is 63 percent of the overlaid 12-BR-H4 weld thickness. Thus, the allowable crack depth for a 360* circumferential crack is (.63) (0.97) or 0.61 inch, which is greater than the original butt weld thickness.
Thus, the allowable crack depth is greater than the original weld thickness independent of crack i
orientation.
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5.1.2.2 Crack Growth .
The existing cracks could grow due to both fatigue and stress corrosion. Fatigue crack growth due to the three _
types of thermal transients defined in Section 4.2 was calculated using the material properties from Reference
- 10. The fatigue cycles considered are shown in Figure 5.3. The fatigue crack growth for 5 years was calculated to be less than 0.01 inch.
IGSCC growth requires a susceptible material. The pup piece side of the 12-BR-H4 weld is not susceptible to IGSCC due to its low carbon content. The overlay weld itself is not susceptible due to its high ferrite content and its duplex structure. The piping that was not exposed to high temperature during welding is not susceptible since it was annealed prior to welding.
Thus, the only susceptible material is the heat affected zone on the piping side of the original 12-BR-H4 butt weld and a very thin layer on the pipe side of the 12-BR-H4 weld on the boundary between the piping and the weld overlay (Figure 5.4). The thickness of this sensitized layer is small since the overlay welding I
technique uses low weld heat input specifically to minimize additional sensitization. Thus, the potential CPL-09-102 19 Revision 0 nutgtsh
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1 for additional crack growth in the axial direction is minimized. A conservative calculation of the axial crack growth in this sensitized layer was performed. It was based on an upper bound thickness of the sensitized layer (1/8 inch) and on a state of the art calculation of the ho'op residual stress due to the overlay weld.
The upper bound crack growth law from Figure 5.5 was used. The predicted axial crack growth is less than 0.01 inch in five years. This growth rate is slow enough that any potential error in the residual stress prediction will not invalidate the conclusion that the overlay has a design life of at least five years.
The weld overlay is not susceptible to IGSCC. Thus the maximum depth after five years of a 360' circumferential crack is approximately 0.58 inch. Based on Section 5.1.2.1 a circumferential crack of that size is acceptable.
Thus, the overlay design is acceptable for five years.
5.1.2.3 Tearing Modulus l
The largest size to which the existing crack could reasonably be expected to grow was postulated to be a CPL-09-102 20 Revision 0 nutggh
long circumferential crack of depth equal to that given in Section 5.1.2.2 (0.58 inch). This assumes l
significant growth of the crack in the circumferential direction. A tearing modulus evaluation was then ,
performed for this postulated crack. The applied loads were pressure, weight, seismic, and thermal expansion.
The evaluation was performed using the methodology of Reference 12 with material properties from Reference 13.
The postulated flaw and the results are shown in Figure 5.6. The upper dotted line represents the inherent material resistance to unstable fracture in terms of J-integral and Tearing Modulus, T. The line originating at the origin represents the applied loading. Increasing load results in applied J-T combinations moving up this line, and unstable fracture is predicted at the intersection of this applied loading line with the material resistance line.
Figure 5.6 shows that the predicted failure load is in excess of 3.4 times the normal applied loads. Thus, there is a safety factor on normal loads (including OBE seismic) of at least 3.4, which is well in excess of the CPL-09-102 21 Revision 0 nutggh
safety factor inherent in the ASME Code, even in the presence of this worst case assumed crack.
5.2 12" Elbow Evaluation ,.
The heat affected zone of the weld (12-AR-E2) between the N2E recirculation inlet elbow and the vertical riser has a through wall crack. During the repair, several additional closely spaced indications were discovered in the weld heat affected zone. To ensure a conservative overlay design, it was assumed that a long through wall circumferential crack exists in the 12-AR-E2 weld.
5.2.1 Code Stress Analysis A finite element model of the cracked and weld overlaid region was developed using the ANSYS (Reference 7) computer program. The model was based on an overlay thickness of 0.35 inch, which is 22 percent smaller than the actual minimum average thickness of 0.45 inch. The small added weight (less than 40 pounds) of the weld overlay will not significantly change the existing l
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l seismic analysis. Figure 5.7 shows the model. This l i
figure also shows the material that was removed to represent the crack.
The stress in the overlaid elbow due to design pressure and appliid moments as described in Sections 4.1 and 4.2 was calculated with the finite element model. The thermal analysis was performed in the same manner as for the safe end (Section 5.1.1), with appropriate dimensional changes.
The results of a code stress analysis per Reference 1 are given in Table 5.3. The allowable stress values for Reference 1 are also given. The weld overlay repair satisfies the Reference 1 requirements.
A conservative fatigue analysis per Reference 1 was performed. A fatigue strength reduction factor of 5.0 was applied due to the crack. The fatigue usage factor was then calculated assuming 38 startups, 25 small temperature change cycles and one emergency cycle every 5 years. The results are summarized in Table 5.3.
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5.2.2 Fracture Mechanics Evaluation Three types of fracture mechanics evaluations were performed. The allowable crack depth was calculated _
based on Reference 2. Crack growth due to both fatigue and IGSCC was calculated using the NUTECH computer program NUTCRAK (Reference 9) with material constants and methodology from References 10 and 11. Finally, the ultimate margin to failure for the assumed crack was calculated per References 12 and 13.
5.2.2.1 Allowable Crack Depth The design minimum overlay thickness of 0.35 inch was l established (References 2 and 11) based on a 360*
through wall circumferential crack. The minimum average as-built overlay thickness is 0.45 inch. The allowable crack depth for a 360* through wall crack with an overlay thickness of 0.45 inch is 0.66 inch, which is greater than the original butt weld thickness.
5.2.2.2 Crack Growth The assumed 360* crack is completely surrounded by IGSCC resistant material; the 12-AR-E2 weld, the weld overlay, l
CPL-09-102 24 Revision 0 nutagh
the annealed elbow and the annealed piping. Thus, no -
IGSCC growth will occur. The fatigue crack growth due to 5 years of the cycles shown on Figure 5.3 is less than 0.01 inch. Thus, the assumed crack will not grow _
to an unacceptable size within the next 5 years.
5.2.2.3 Tearing Modulus A tearing modulus evaluation was performed for the assumed 360* through original pipe wall crack. The normal operating loads of OBE seismic, pressure, weight and thermal expansion were applied.
The evaluation was performed using the methodology of Reference 12 with material properties from Reference 13. The postulated flaw and the results are shown in Figure 5.8. The upper dotted line represents the inherent material resistance to unstable fracture in terms of J-integral and Tearing Modulus, T. The line originating at the origin represents the applied loading. Increasing load results in applied J-T combinations moving up this line, and unstable fracture is predicted at the intersection of this applied loading line with the material resistance line.
CPL-09-102 25 Revision 0 nutp_qh
[ ,
Figure 5.8 shows that the predicted failure loads is in excess of 3.4 times the normal operating loads. Thus, there is a safety factor on normal operating loads, including OBE seismic, of at least 3.4, which is in _
excess of the safety factor inherent in the ASME Code, even in tNe presence of this worst case assumed crack.
5.3 28" Elbow Evaluation The heat affected zone of the weld (28-A-15LU-0) between the recirculation pump discharge elbow and the vertical riser has two small axial UT indications. The larger of the two axial indications was estimated to have a depth of approximately 11 percent of the wall thickness. To ensure a conservative overlay design, it was assumed that a small circumferential crack exists in the 28-A-ISLU-0 weld. Thus, two bounding cases were considered; an axial crack of depth equal to 0.14 inch and length equal to 1/2 inch, and a circumferential crack of depth equal to 0.14 inch and length equal to ten times the crack depth (1.4 inches).
t I
l CPL-09-102 26 i Revision 0 s
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5.3.1 Code Stress Analysis A finite element model of the cracked and weld overlaid region was developed using the ANSYS (Reference 7) ,
computer program. The model was based on an overlay thickness'of 0.30 inch, which is 36' percent smaller than the actual minimum average thickness of 0.47 inch. The small added weight (less than 100 pounds) of the weld overlay will not significantly change the existing seismic analysis. Figure 5.9 shows the model. This figure also shows the material that was removed to represent the crack.
The stress in the overlaid elbow due to design pressure and applied moments as described in Sections 4.1 and 4.2 was calculated with the finite element model. The thermal analysis was performed in the same manner as for the safe end (Section 5.1.1), with appropriate dimensional changes.
The results of a code stress analysis per Reference 1 are given in Table 5.4. The allowable stress values for Reference 1 are also given. The weld overlay repair satisfies the Reference 1 requirements.
I CPL-09-102 27 Revision 0 nutggb
, A conservative fatigue analysis per Reference 1 was i
i l performed. A fatigue strength reduction factor of 5.0 was applied due to the crack. The fatigue usage factor j was then calculated assuming 38 startups, 25 small temperature change cycles and one emergency cycle every five year's. The results are summar'ized in Table 5.4.
5.3.2 Fractura Mechanics Evaluation Three types of fracture mechanics evaluations were performed. The allowable crack depth was calculated based on Reference 2. Crack growth due to both fatigue l and IGSCC was calculated using the NUTECH computer program NUTCRAK (Reference 9) with material constants and methodology from References 10 and 11. Finally, the ultimate margin to failure for the assumed crack was calculated per References 12 and 13.
5.3.2.1 Allowable Crack Depth The design minimum overlay thickness is 0.30 inch. The minimum average as-built overlay thickness is 0.47 inch. The Reference 2 allowable crack depth for a 1/2 inch long axial crack with an overlay thickness of 0.47 inch, is 1.3 inches. The Reference 2 allowable CPL-09-102 28 Revision 0 l
nutggh 1
crack depth for a 1.4 inch long circumferential crack with an overlay thickness of 0.47 inch also is 1.3 inches.
5.3.2.2 Crack Growth ,
The existing cracks could grow due to both fatigue and stress corrosion. Fatigue growth due to the three types of thermal transients defined in Section 4.2 was calculated using the material properties from Reference
- 10. The fatigue crack growth for five years of the cycles shown in Figure 5.3 was calculated to be less than 0.01 inch.
IGSCC crack growth was calculated using the upper bound crack growth law shown in Figure 5.5. The residual stress was based on both the original butt weld residual stress and the residual stress from the weld overlay.
The residual stress due to the original butt weld was obtained from Reference 14 and the residual stress due to the weld overlay was obtained from a state of the art calculation. The combined residual stress is shown in Figure 5.10. The crack growth analysis was performed per Appendix A of Reference 6. Crack growth as a function of time was determined for both an axial and a CPL-09-102 29 Revision 0 N h._
circumferential crack of size described in -
Section 5.3. The axial crack was predicted to grow faster than the circumferential crack. The predicted axial crack depth as a function of time is shown in _
Figure 5.11. Maximum crack depth after five years is predicted' to be 0.25 inch which is well below the '
allowable of 1.3 inches. The predicted growth is slow enough that any potential error in the residual stress prediction will not invalidate the conclusion that the overlay has a design life of at least five years.
5.3.2.3 Tearing Modulus A tearing modulus evaluation was performed for a crack of depth equal to 0.50. The normal operating loads of OBE seismic, pressure, weight and thermal expansion were applied. The evaluation was performed using the methodology of Reference 12 with material properties from Reference 13.
t The postulated flaw and the results are shown in Figure 5.12. The upper dotted line represents the inherent material resistance to unstable fracture in terms of J-integral and Tearing Modulus, T. The line originating at the origin represents the applied CPL-09-102 30 Revision 0 nutp_qh
I loading. Increasing load results in applied J-T combinations moving up this line, and unstable fracture is predicted at the intersection of this applied loading line with the material resistance line. _
Figure 5.12 shows that the predicted failure load is in excess of 3.2 times the normal operating loads. Thus, there is a safety factor on normal operating loads, including OBE seismic, of at least 3.2, which is in excess of the safety factor inherent in the ASME Code, even in the presence of this worst case assumed crack.
5.4 Effect on Recirculation and RHR Systems Installation of the weld overlay repairs caused a small amount of radial and axial shrinkage underneath the overlay. Based on measurements of the weld overlays, the maximum axial shrinkage was 5/64 inch (safe end to pup piece).
The effects of the radial shrinkage are lLaited to the region adjacent to and underneath the overlay. Based on Reference 15, the stresses due to the radial shrinkage are less than yield stress at distances greater than 4
! inches from the ends of the overlay. Weld residual CPL-09-102 31 Revision 0 1
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stresses are steady state secondary stresses and thus are not limited by the ASME Code (Reference 1).
The effect of the axial weld shrinkage on the Recirculation and RHR Systems was evaluated with the NUTECH computer program PISTAR (Reference 5) and the piping model shown in Figure 5.13.
The measured axial shrinkage of the 12-inch elbow weld overlay (3/64 inch), the 28-inch elbow weld overlay (1/16 inch), and of the safe end to pipe weld overlay (5/64 inch) were imposed as boundary conditions on this model. Since the ASME Code does not Ibnit weld residual stress, all stress indices were set equal to 1.0.
The maximum calculated stress was less than 4.0 ksi.
The location of this stress is shown on Figure 5.13.
Steady state secondary stresses of 4.0 ksi are judged to have no deleterious effect on the Recirculation or RHR Systems.
CPL-09-102 32 Revision 0 l nutggh l
l
NORMAL SMLL TEMPERATURE EMERGENCY STARTUP PARAMETER' CYCLE M GE CYCLE CYCLE (CYCLE 1) (CYCLE 2) (CYCLE 3) 0 0 EQUIVALENT 2F 32 F 265 F LINEAR TEMPERATURE AT g 0
PEAK 0 8F 640F TEMPERATURE AT g THROUGH 368 PSI 8,840 PSI 72,370 PSI WALL THERMAL STRESS e Table 5.1 THERMAL STRESS RESULTS CPL-09-102 Revision 0 33 nutagh
ACTUAL EQUATION STRESS SECTION III CATEGORY NUMBER N8 ALLOWABLE OR THICKNESS S N/A N/A S, = 16,800 PSI 15,650 PRIMARY (9) p$x 25,200 PSI 24,700 (10) p5 50,400 PSI 5 ARY PEAK CYCLE 1 (22,900)S*
(I'1) (12,900)5 N/A CYCLE 2 CYCLE 3 (129,200)5 USAGE FACTOR N/A 0.03 1.0 (5YR)
- THE FACTOR OF 5 IS THE CONSERVATIVELY ASSUMED FATIGUE STRENGTH REDUCTION FACTOR.
Table 5.2 SAFE END CODE STRESS RESULTS CPL-09-102 Revision 0 34 nutggh
I l
ACTUAL EQUATION STRESS SECTION III CATEGORY NUMBER OR N8 ALLOWABLE THICKNESS -
S N/A N/A S, = 14,400** PSI PRIMARY (9) 15,450 PSI 21,600 PSI (10) 23,820 PSI 43,200 PSI ARY PEAK CYCLE 1 (18,300)5'*
(11) N/A CYCLE 2 (12,900)5 CYCLE 3 (124,500)5 USAGE FACTOR N/A 0.02 1.0 (5YR)
- THE FACTOR OF 5 IS THE CONSERVATIVELY ASSUMED FATIGUE STRENGTH REDUCTION FACTOR.
- LIMITING LOCATION IS IN THE 308L WELD OVERLAY. i l
l t
Table 5.3 12" ELBOW CODE STRESS RESULTS CP L-0 9-10 2 Revision 0 35 nutggb
ACTUAL EQUATION STRESS SECTION III CATEGORY NUMBER OR NB ALLOWABLE THICKNESS S N/A N/A S, = 16,800 PSI PRIMARY (9) 13,200 PSI 25,200 PSI
$gy* (10) 26,200 PSI 50,400 PSI PEAK CYCLE 1 (20,700)S*
(11) N/A '
CYCLE 2 (12,900)5 CYCLE 3 (126,900)5 USAGE FACTOR N/A 0.02 1.0 (5YR)
- THE FACTOR OF 5 IS THE CONSERVATIVELY ASSUMED FATIGUE STRENSTH REDUCTION FACTOR.
Table 5.'4 28" ELBOW CODE STRESS RESULTS CPL-09-102 Revision 0 36 nutggb
l CRACK LOCATION
/
JN w -
s Ns d't:y Ny
's N
, N '
Figure 5.1 SAFE END FINITE ELEMENT MODEL CPL-09-102 Revision 0 37 nutggb
INSULATION WELD ..
, , , , , , , , , , , , , , , , , , , , , , ,,,,, 1?
,, , b$ $N,,,,,, i S\\ \\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\% " l AA 0.57" J 6.58" 5.81"
T
+A i
N N
h h==
~
.:,' ~
k = 10 BTU /hr-ft OF T :;,
, , ' [, , -
I SECTION A-A 1
Figure 5.2 WELD OVERLAY THERMAL MODEL CPL-09-102 Revision 0 38 nutggb
EMERGENCY -
SMALL TEMPERATURE -
$ CHANGE STARTUP SHUTDOWN -
NORMAL - j j OPEP.ATION {
RESIDUAL - '
38 25 5 CYCLES CYCLES _
YEARS __
^
\ ENIIRE 1' SECUENCE CYCLE REFEATS TIME Figure 5.3 THERMAL TRANSIENTS CPL-09-102 Revision 0 39 nutsch
y
. . ..t, s s' i
- s. -
w 1 I
8
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POTENTIALLY IGSbC <*i, i- 'T s t SUSCEPTIBLE MAT 3MIAL .' [LG.th/iRLAY s .'
g
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l SAFE END 1GSCC SUSCEPTIBILITY -
N I
CPL-09-102 '
Revision 0 40 nutRCh
.c -
+
j, u 7,.e.
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s s.. <
i 3 g i L, ', ,
s ,
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T Figure 5.5 s~ a TYPICAL IGSCC CRACK GROWTH DATA
~
3 (WELD SENSITIZED 304SS IN BWR ENVIRONMENT) t ,
\ 4 C'?L-09-102 -
Revision 0 41
. ,N w
/ '
240 -
200 - J = 9,300 c in 2 u, 160 -
M N
- 120 - 3.4 x NORMAL
- S LOADS 5 80 -
3 x NORMAL LOADS I"D 40 - J = 6000 2 IC in 0 . , , , , , ,
0 40 80 120 160 200 240 T
OVERLAY WELD
/ \
ANNEALED MATERIAL ANNEALED MATERIAL WELD / 0.58" RADIUS FLAW l
Figure 5.6 SAFE END TEARING MODULUS CPL-09-102 Revision 0 42 nutgLqh
I e
e
/
\\
y
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1q
~
..++ / \\
lil J
I l
- /
Figure 5.7
- s. / s /*' h c '
ys k0 .
t s
x h
l i
240 -
200 - J e
= 9,300 '" 2 in u, 160 -
M N
{;*. 120 - 3.4 x NORMAL LOADS J!!
! 80 -
3 x NORMAL LOADS I"Ib 40 - J = 6000 2 IC in 0 . . . . . . .
0 40 80 120 160 200 240 T
OVERLAY WELD
/ \
ANNEALED MATERIAL ANNEALED MATERIAL 0.58" WELD RADIUS FLAW Figure 5.8 l
12" ELBOW TEARING MODULUS CPL-09-102 Revision 0 44 nutgg.h
CRACK LOCATION
/
s
/
/
/ /
/
1
& w
~
Figure 5.9 28" ELBOW-TO-PIPE FINITE ELEMENT MODEL CPL-09-102 Revision 0 45 nutggh
40 30 - l 20 -
10 -
b 0 00 d0
$ ELBOW OVERLAY E .
4 E G U
t
-40 Figure 5.10 28" ELBOW-TO PIPE CIRCUMFERENTIAL WELD RESIDUAL STRESS CPL-09-102 Revision 0 46 0
v f:5 .;g SiA ~
'NNNkll n N a N0 4 -t -
? -
L -
?
NB W 1.0 r
...sssss . ... . .
g::. - 4, sssss- <msss ---~ gg; s 0.8 O.6 -
C.
0.4 -
PREDICTED IGSCC GROWTH NEXT 5 YEARS
- 0. 2 -
2 0.0 , , .
j 0.0 1.0 2.0 3.0 4.0 5.0 TIME (years)
Figure 5.11 28" ELBOW-TO-PIPE CRACK GROWTH CPL-09-102 -
Revision 0 47 nutggb
240 - _
I"Ib 200 - J = 11,200 e 2 in
~, 150 -
M
~
~ 3.2 x NORMAL 5 120 - LOADS 2
3 3 x NORMAL LOADS 5 80 -
"b 40 - J = 6000 2 0 . . . . . . .
0 40 80 120 160 200 240 T
OVERLAY WELD
/ \
ANNEALED MATERIAL ANNEALED MATERIAL WELD / 0.50" RADIUS FLAW Figure 5.12 28" ELBOW TO-PIPE TEARING MODULUS CPL-09-102 Revision 0 43
1 1
I HIGHEST STRESS 1.0 CATION %.
~
i - h=~=
=
p.
m., -im ,
)
7 i m
' mp i
m N
= .
ma f k '. '" '
i p '
,,,, g, _
7 == '"
== "
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- i. L rm 8 sa ,..
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'ta" i
1.
f ,,. e '
= .. 'm / 'F. .
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, f fm w -
u.=.
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)
em M""
p"r ass se
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. 7%
. m 3 ==
=
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=
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)=
=
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m -. , , ,
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=
l l
l l Figure 5.13 PIP!NG MODEL CP L-09-10 2 Revision 0 49 l nutg_Q,b l
6.G LEAK-BEFORE-BREAK 6.1 Net Section Collapse The simplest way to determine the effect of IGSCC on the structurai integrity of piping is through the use of a sLmple " strength of materials" approach to assess the load carrying capacity of a piping section after the cracked portion has been removed. Studies have shown (Raferences 11 and 13) that this approach gives a conservative, lower-bound estbnate of the loads which would cause unstable fracture of the cracked section.
Typical results of such an analysis are indicated in Figure 6.1 (Reference 11). This figure defines the locus of limiting crack depths and lengths for circumferential cracks which are predicted to cause failure by the net section collapse method. Curves are presented for both typical piping system stresses and stress levels equal to ASME Code limits. Note that a very large percentage of pipe wall can be cracked before reaching these limits (40% to 60% of circumference for through-wall cracks, and 65% to 85% of wall thickness l for 360* part-through cracks).
CPL-09-102 50 Revision 0 nutqqh l
I
l l
Also shown in Figure 6.1 is a sampling of cracks which have been detected in service, either through UT examination or leakage. In each case there has been a )
comfortable margin between the size crack that was -
observed and that which would be predicted to cause failure u'nder service loading conditions. Also, as discussed below, there is still considerable margin between these net section collapse limits and the actual cracks which would cause instability.
6.2 Tearinq Modulus Analysis Elastic-plastic fracture mechanics analyses are presented in Reference 13 which give a more accurate representation of the crack tolerance capacity of stainless steel piping than the net section collapse approach described above. Figures 6.2 and 6.3 graphically depict the results of such an analysis (Reference 13). Through-wall circumferential defects of arc-length equal to 60* through 100* were assumed at various cross sections of a typical BWR Recirculation System. Loads were applied to these sections of sufficient magnitude to produce net section ihmit load, l and the resulting values of tearing modulus were compared to that' required to cause unstable fracture CPL-09-102 51 i Revision 0 l
l nutggb
(Figure 6.2). Note that in all cases there is substantial margin, indicating that the net section collapse lhmits of the previous section are not really failure limits. Figure 6.3 summarizes the results of _
all such analyses performed for 60* through-wall cracks in terms 'of margin on tearing modulus for stability.
The margin in all cases is substantial.
6.3 Leak versus Break Flaw Conficuration of perhaps more significance to the leak-be f ore-break argument is the flaw configuration depicted in Figure 6.4. This configuration addresses the concerns raised by the occurrence of part-through flaws growing, with respect to the pipe circumference, before breaking through the outside surface to cause leakage. Figure 6.4 presents typical size lbmitations on such flaws based on the conservative, net section collapse method of Section 6.1. Note that very large crack sizes are predicted. Also shown on this figure are typical detectability lLaits for short through-wall flaws (which are amenable to leak detection) and long part-through flaws (which are amenable to detection by UT). The margins between the detectability limits, and the conservative, net section collapse failure limits are CPL-09-102 52 Revision 0 nutggb
substantial. It is noteworthy that the likelihood of l flaws developing which are characterized by the vertical axis shown in Figure 6.4 (constant depth 360* circumfer-ential cracks) is so remote as to be considered _
impossible. Material and stress asymmetries always tend to propagate one portion of the crack faster than the bulk of the crack front, which will eventually result in
" leak-before-break." This observation is borne out by extensive field experience with BWR IGSCC.
6.4 Axial Cracks The recent IGSCC occurrences at Monticello and Hatch I were predominately short, axial cracks which grew through the wall but remained very short in the axial direction. This behavior is consistent with expectations for axial IGSCC since the presence of a '
{ sensitized weld heat-affected zone is necessary, and this heat-affected zone is limited to approximately 0.25 inch on either side of the weld. Since the major loadings in the above net section collapse analysis are bending moments on the cross section due to seismic loadings, and since these loads do not exist in the circumferential direction, the above leak-before-break arguments are even more persuasive for axially oriented CPL-09-102 53 Revision 0 nutggb 1
- 1
cracks. There is no known mechanism for axial cracks to lengthen before growing through-wall and leaking, and .
l the potential rupture loading on axial cracks is less than that on circumferential cracks. -
6.5 Multiple' Cracks Recent analyses performed for EPRI (Reference 16) indicate that the occurrence of multiple cracks in a weld, or cracking in multiple welds in a single piping line do not invalidate the leak-before-break arguments discussed above.
6.6 Crack Detection capability IGSCC in BWR piping is detected through two means: non-destructive examination (NDE) and leakage detection.
Although neither is perfect, the two means complement one another well. This detection capability combined with the exceptional inherent toughness of stainless steel, results in essentially 100% probability that IGSCC would be detected before it significantly degraded the structural integrity of a BWR piping system.
CPL-09-102 54 Revision 0 i
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6.7 Non-Destructive Examination The primary means of non-destructive examination for IGSCC in BWR piping is ultrasonics.
This method has -
been the subject of considerable research and developme'nt in recent years, and significant improvements in its ability to detect IGSCC have been achieved. Nevertheless, recent UT experience at Brunswick 1, Monticello, Hatch 1, and elsewhere indicate that there is still considerable room for improvement, especially in the ability to distinguish cracks or crack-like indications from innocuous geometric conditions.
Figure 6.4, however, illustrates a significant aspect of UT detection capability with respect to leak-before-break. The types of cracking most likely to go undetected by UT are relatively short circumferential or axial cracks which are most amenable to detection by leakage. Conversely, as part-through cracks lengthen, and thus become more of a concern with respect to leak-before-break, they become readily detectable by UT, and I
are less likely to be misinterpreted as geometric conditions.
CPL-09-102 55 Revision 0 l
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I 6.8 Leakace Detection Typical leakage detection capability for BWR reactor coolant system piping is through sump level and drywell -
activity monitoring. These systems have sensitivities on the or' der of 1.0 gallon per minute (GPM) of unidentified leakage (i.e., not from known sources such as valve packing or pump seals) . Plant technical specification Ibnits typically require investigation /
corrective action at 5.0 GPM unidentified leakage.
Table 6.1 provides a tabulation of typical flaw sizes to cause 5.0 GPM leakage in various size piping (Reference 11).
Also shown in this table are the critical crack lengths for through-wall cracks based on the net section collapse method of analysis discussed above. For conservatism, the leakage values are based on pressure stress only, while the critical crack lengths are based on the sum of all combined loads, including seismic.
(Considering other normal operating loads in the leakage l
l analysis would result in higher rates of leakage for a given crack size.) Note that there is considerable i
l CPL-09-102 56 Revision 0 l nutg_qh
margin between the crack length to produce 5.0 GPM leakage and the critical crack length, and that this margin increases with increasing pipe size.
6.9 Historical Experience The above theories regarding crack detectability have been borne out by experience. Indeed, of the approximately 400 IGSCC incidents to date in BWR piping, all have been detected by either UT or leakage, and none have even come close to violating the structural integrity of the piping (Reference 16).
O CF'_-09-102 57 Revision 0 nutgq.h
4 CRACK LENGTH FOR CRITICAL CRACK gjg NOMINAL c
PIPE SIZE 5 GPM LEAK (in.) LENGTH Ac (in.)
4" SCH 80 4.50 6.54 0.688 10" SCH 80 4.86 15.95 0.305 t 24" SCH 80 4.97 35.79 0.139 i
1 ee l
i i
! Table 6.1 EFFECT OF PIPE SIZE ON THE RATIO OF THE CRACK LENGTH FOR 5 GPM LEAK RATE AND THE CRITICAL CRACK LENGTH (ASSUMED STRESS a = Sm/23 CPL-09-102 Revision 0 58 l nutggh
l N
d 8h _.
t l
l l l
t 1.0 "N g O
\
\
\
\
0.8 4 g %
g P, = 6 ksi, Pb=0 5
@@ %'*% ** * ==
I g 0.6 - Pm = 6 kai, Pm+Pb " l38m f-3 0 .4 - 4 Field Cata - Part-Througn Flaws f ,
O Field Data - Laaks 3 @
E Sm = 16.0 ksi
@ af = 48.0 kai 01 -
Values at SECCF 0 - .-
O 0.2 0.4 0.6 0.8 1.0 Fraction of Circumferenca. 0/r l
Figure 6.1 TYPICAL RESULT OF NET SECTION COLLAPSE ANALYSIS OF CRACKED STAINLESS STEEL PIPE CPL-09-102 Revision 0 59 nutech
i m.e
- 2108 5 g ,, a " T h p Header a
? w c punio "
l m./ !
/g \ oissnerve soo -
l T=0 l
250 -
l n 200- (Unsame)
S l
n 4 15G . (5 aWei a 2e = 12aa
. -28 = 2408 3"#
l ' 1 3aa .
( 29 = 18o8 50 - , , , , , , , _ _ ,,, , , , , , , , , , , ,__,
,_,_,,,,,_,,_,__,,,,,_,,,,,,_.,,Jte 2, = -
I a .- . . . . .
50.0 12.5 25.0 82.5 10a.0 137.5 175.0 212.5 250.0 7
l l
Figure 6.2 STABILITY ANALYSIS FOR BWR REClRCULATION SYSTEM (STAINLESS STEEL)
CPL-09-102 Revision.0 60 1 nutech
20 -
60 TitPOUGil CHACK gy PLASTIC LIGAMENT "s
o ta 15- A
$ RANGE OF MATERIAL o TEARING MODULUS RANGE OF APPLIED
$ TEARING MODULUS
$ 10- .
i /
i 0 50 100 150 200 250 TEARING MODULils 1
Figure 6.3 3
SUMMARY
OF LEAK BEFORE BREAK ASSESSMENT i
C 4 OF BWH HEClHCULATION SYSTEM l $
l O
T
\
l
l 2=
\ \
t a PIPE CROSS SECTION 0.7 0.6 - - - - - - - - - --
l 0.5 - I I
I
- 0. 4 - g
<5 l 0.3- l l
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0.0 0 0.2 0.3 0.4 0.5 0.6 0.7
=/r Figure 6.4 TYP! CAL PIPE CRACK FAILURE LOCUS FOR COMBINED THROUGH WALL PLUS 3600 PART THROUGH CRACK CPL-09-102 Revision 0 62 nutech
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7.0
SUMMARY
AND CONCLUSIONS The evaluation of the repairs to the Recirculation System reported herein shows that the resulting stress .l levels are acceptable for all design conditions. The stress le'vels have been assessed from the standpoint of load capacity of the components, fatigue, and the resistance to crack growth.
Acceptance criteria for the analyses have been established in Section 3.0 of this report which demonstrate that:
- 1. There is no loss of design safety margin over that provided by the current Code of Construction for Class 1 piping and pressure vessels (ASME Section III).
- 2. During the design lifetime of each repair, the observed cracks will not grow to the point where the above safety margins would be exceeded.
Analyses have been performed and results are presented which demonstrate that the repaired welds satisfy these l
CPL-09-102 63 Revision 0 0
r-criteria by a large margin, and that the design life of each repair is at lease five years.
Furthermore, it is concluded that the recent IGSCC ,
experienced in the Reactor Recirculation System at Brunswick 'l does not increase the probability of a design basis pipe rupture at the plant. This conclusion expressly considers the nature of the cracking which has been repaired at Brunswick 1, and the likeliood that other similar cracking may have gone undetected. The conclusion is based primarily on the extremely high inherent toughness and ductility of the stainless steel piping material; the tendency of cracks in sucn piping to grow through-wall and leak before affecting its structural load carrying capacity (which indeed was the case in the defects observed at Brunswick 1); and the fact that as cracks lengthen and are less likely to
" leak-before-break", they become more amenable to detection by other NDE techniques such as UT and RT.
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nutggb l
l
l
8.0 REFERENCES
l
- 1. ASME Boiler and Pressure Vessel Code Section III, Subsection NB, 1974 Edition with Addenda through -
Summer 1975. .
- 2. ASME Boiler and Pressure Vossel Code Section XI, Paragraph IWB-3640 (Proposed), " Acceptance Criteria for Austenitic Stainless Steel Piping" (Presented to Section XI Subgroup on Evaluation Standards in November 1982).
- 3. General Electric Design Specification 22A1417, Revision 2.
- 4. General Electric letter G-KB1-1-193, December 30, 1981, " Transmittal of GE Design Memo 170-17 Rev. 1, on Seismic Reevaluation of Recirculation Piping System for Brunswick Units 1 and 2".
- 5. NUTECH Computer Program PISTAR, Version 2.0, Users Manual, Volume 1, TR-76-002, Revision 4, File Number 08.003.0300.
l CPL-09-102 65 Revision 0 nutp_q,b
I
- 6. ASME Boiler and Pressure Vessel Code Section XI,
~
- 1980 Edition with Addenda through Winter 1981.
l
- 7. ANSYS Computer Program, Swanson Analysis Systems, -
Revision 4. .
- 8. Schneider, P.J., " Temperature Response Charts,"
John Wiley and Sons, 1963.
- 9. NUTCRAK Computer Program, Revision 0, April 1978, File Number 08.039.0005.
- 10. EPRI-2423-LD, " Stress Corrosion Cracking of Type 304 Stainess Steel in High Purity Nater - a Compilation of Crack Growth Rates," June 1982.
- 11. EPRI-NP-2472, "The Growth and Stability of Stress Corrosion Cracks in Large-Diameter BWR Piping,"
July 1982.
- 12. NUREG-0744, Volume 1 for Comment, " Resolution of the Reactor Materials Toughness Safety Issue."
CPL-09-102 66 Revision 0 nutp_qh
- 13. EPRI-NP-2261, " Application of Tearing Modulus Stability Concepts to Nuclear Piping," February 1982.
- 14. NUTECH Internal Memo, PCR-8 3-0 03 , " Weld Residual Stress for IGSCC Crack Growth Calculations,"
March 4, 1983
- 15. NUTECH Report NSP-81-105, Revision 2, " Design Report for Rec'irculation Safe End and Elbow Repairs, Monticello Nuclear Generating Plant,"
December 1982.
Nuclear Regulatory Commission, " Status of BWR IGSCC Development Program," October 15, 1982.
I.
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