ML20064E509
| ML20064E509 | |
| Person / Time | |
|---|---|
| Site: | Monticello  |
| Issue date: | 10/31/1982 |
| From: | Charnley J, Riccardella P, Rich Smith NUTECH ENGINEERS, INC. |
| To: | |
| Shared Package | |
| ML20064E467 | List: |
| References | |
| FOIA-82-514 NSP-81-103, NSP-81-103-R, NSP-81-103-R00, NUDOCS 8301060035 | |
| Download: ML20064E509 (44) | |
Text
.
>e AOVANCE. CoPT NSP-81-103 David Musolf, P.E. Revision 0 Mang.
nuci., suppori s.rvie o.pt.
October 1982 30.1281.0103 Northern States Power Company [
s
~ 414 Nicollet Mail. 8th Floor u nneanoi,s Minnesota 5s401 DESIGN REPORT Telephone (612) 3306764 RECIRCULATION LINE END CAP REPAIR '
MONTICELLO NUCLEAR GENERATING PLANT Prepared for:
Northern States Power Company L-Prepared by: ,
[ NUTECH Engineers, Inc.
_ San Jose, California r
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Prepared by: .: . Reviewed by:
i e IO 22-82 J. E. Charnley, P.E. R. H. Smith Project Engineer Project Quality Assurance Engineer Approved by: Issued by:
IalL /=cX Ln, 2n "
P. C. Riccardella, P.E. N. Eng Senior Director Project Manager lg$r.c 8301060035 821117
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PDR FOIA GHORT82-514 1g g @ h_
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REVISION CObTROL SHEET
.s TITLE: Design Report for Recirculation REPORT NUMBER: NSP-81-103
- Line End Cap Repair, Monticello Revision 0 Nuclear Generating, Plant
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J. E. Charnley / Principal Engineer M N AME / TITLE INITI ALS
-bat / bY P. C. Riccardella/ Senior Director /Pc/2 N AME / TITLE INITIALS S. Kulat/ Consultant I NAME/ TITLE h
INITI ALS H. L. Gustin / Engineer N AME / TITLE INITIALS Y. S. Wu/ Consultant I V6 @f.
N AME / TITLE INITIALS
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REVISION CONTROL SHEET (CONTINUATION) l TITLE: Design Report for RecirculationREPORT NUMBER: NSP-81-103 Line End Cap Repair, Monticello Revision 0 ,
Nuclear Generating. Plant PREPARED ACCURACY CRITE RI A
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1 CERTIFICATION BY REGISTERED PROFESSIONAL ENGINEER _I hereby certify that this document and the calculations contained herein were prepared under my direct supervision, f reviewed by ne, and to the best of my knowledge are correct and complete. I am a duly Registered Professional Engineer under the laws of the States of Minnesota and California and am competent to review this document. Certified by:
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Registration No. 14372 State of California i Registration No. 16340 Il 1 Date 22 h fih1 l l I 1 NSP-81-103 iv I
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b b TABLE OF CONTENTS Page -, LIST OF TABLES - vi LIST OF FIGURES vii
1.0 INTRODUCTION
1 2.0 REPAIR DESCRIPTION 4 3.0 EVALUATION CRITERIA 6 i ~ 3.1 Strength Evaluation 7 3.2 Fatigue Evaluation 7 3.3 Crack Growth Evaluation 8 3 4.0 LOADS 10 4.1 Mechanical and Internal Pressure Loads 10 4.2 Thermal Loads 11 5.0 EVALUATION METHODS AND RESULTS 12 5.1 Code Stress Analysis 12 5.2 Fracture Mechanics Evaluation 14 5.2.1 Allowable Crack Depth 15 5.2.2 Crack Growth 17 5.2.3 Tearing Modulus 20
. 6.0
SUMMARY
AND CONCLUSIONS - 34
7.0 REFERENCES
35 n 1 _ NSP-81-103 V
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e LIST OF TABLES Number Title Page 5.1 Thermal Stress Results - 22 5.2 Code Stress Allowable 22" End Cap 23 5.3 Crack Growth Cases 24 1 H 1
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l LIST OF FIGURES m Number Title Page _. 1.1 Conceptual Drawing of Recirculation ManiEold 3 I 2.1
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Schematic of Weld Overlay 5 m 5.1 ANSYS Model of 22" End Cap Weld Overlay 25 5.2 Applied Stress Profile Through Limiting 26 _ Section 22" End Cap i 5.3 Weld Overl y Therr.L1 Medel 27 ~ 5.4 Thermal Transients 28 5.5 Crack Growth Residual Stress 22" End Cap 29 5.6 Stress Intensity Factor Versus Crack Depth 30 5.7 Crack Growth 22" End Cap 31 5.8 Allowable Crack Depth 22" End Cap 32 - 5.9 Tearing Modulus 22" End Cap 33 n a _ NSP-81-103 vii , j, Revision 0 , Ilut_ech
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1.0 INTRODUCTION
This report summarizes evaluations perfonned by NUTECH
- to assess a weld overlay repair of the end cap to Loop A recirculation manifold weld at Northern States Power Company's Monticello Nuclear Generating Plant. The weld overlay has been applied to address ultrasonic and radiographic examination results believed to be indicative of incipient intergranular stress corrosion cracking (IGSCC) in the vicinity of the weld. The purpose of the 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 the weld overlay repair is at least one fuel cycle. The actual design life will be established by a combination of future analysis and e . . . testing.
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f Three crack indications have been found in the end cap weld heat affected zone. Figure 1.1 shows the recirculation manifold in relation to the reactor pressure vessel (RPV) and other portions of the recirculation system. All three crack indications are located in the 12 o' clock position adjacent to the weld n NSP-81-103 1 : I lPevision 0 . le ... { nut _ech
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-S .) and'all are axial. The largest crack indication is approximately 11 percent of the wall thickness and 1 inch long. J.
I The existing pipe material is ASTM A358, Class'1, Type 304. The existing cap material is ASTM A403, Grade y WP304. 1
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..aI FIGURE 1.1 l CONCEPTUAL DRAWING OF RECIRCULATION ltANIFOLD NSP-81-103 l,' Revision 0 - 3 nut _ech
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3 l 2.0 REPAIR DESCRIPTION 1 The longitudinal crack indications in the end cap weld _ heat affected zone 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 Figure 2.1. The weld deposited band over the longitudinal crack indications will increase the wall thickness to approximately 0.4 inch greater than that which exists in adjacent uncracked piping. In addition, the weld metal deposition will produce a favorable compressive residual stress pattern and the weld metal will be type 308L, _. which is resistant to propagation of IGSCC cracks. t i . M
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e ..; 3.0 EVALUATION CRITERIA This section describes the criteria that ar'e applied in ~ _ this report to evaluate the acceptability of the weld overlay described in Section 2.0. Because of the nature of this repair, 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 on _ strength and fatigue considerations as provided in the i 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 used in the repair.
- As a further means of ensuring structural adaquacy, criteria are also provided.below for fracture mechanics
.,} 'a evaluation of the repair. W J
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> 3.1 Strenath Evaluation Adequacy of the strength of the weld overlay with
! respect to applied mechanical loads is demonstrated with the following criteria:
- 1. An ASME Boiler and Pressure Vessel Code Section III Class 1 (Reference 1) analysis of the weld overlay was performed using worst case loads and using allowable stresses from the Power Piping Code USAS l B31.1.0 (Reference 2).
l 2. The ultimate load capacity of the repair was Li calculated with a tearing modulus analyis. The ratio between failure load and applied loads was required to be greater than~that required by Reference 1. ( u, 3.2 Faticue Evaluation 1 . jj The stress values obtained from the above strength eval-uation were combined with thermal and other secondary I stress conditions to' demonstrate adequate fatigue resistance for the design life of the repair. The L criteria for fatigue evaluation include: [
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- 1. The maximum range of primary plus seco.ndary stress was compared to the secondary strens limits of Reference 1.
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- 2. The peak alternating stress intensity, including
- all primary and secondary stress terms, as well as a fatigue strength reduction factor of 5.0 to
' account for the existing crack, was evaluated using conventicnal fatigue analysis techniques. The total fatigue 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 the repair.
Allowable number of cycles was determined from the
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stainless steel fatigue curve of Reference 1. 3.3 Crack Growth Eval *uation L Crack growth due to both fatigue (cyclic stress) and
) IGSCC (steady state stress) was calculated. The
- r allowable crack depth was established based on net section limit load for the cracked pipe (Reference 3).
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4 The design life of the repair was established as the minimum predicted time for the observed cra'ck indication to grow to the allowable crack depth. L 7 9 l d t-1 1 e E I NSP-81-103 9
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Y. I J 4.0 LOADS The loads considered in the evaluation of the end cap
- weld overlay repair consist of mechanical loads, internal pressure, differential thermal expansion loads, and welding residual stresses. The mechanical loads and - internal pressures used in the analysis are described in r
Section 4.1, and an explanation of the thermal transient conditions which cause differential thermal expansion loads is presented in Section 4.2. Welding residual
*I stresses are considered in the crack growth analyses and are described in Section 4.3.
4.1 Mechanical and Internal Pressure Loads The design pressure of 1248 psi for the recirculation system was obtained from Reference 4. Since the end cap is at the end of *-the recirculation manifcid, there are l no significant dead weight or seismic stresses applied l to it. This is confirmed by the recent NUTECH analysis of the Monticello Reactor Recirculation System piping (Reference 5). E
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s I i 4.2 Thermal Loads Since the end cap is at the end of the reci)culation 7 h manifold, there are no gross section moments due to the thermal expansion applied to it, which is confirmed by s Reference 5. ~! . I The only transient thermal condition defined in Reference 4 that occurs at the end cap is the normal startup and shutdown cycling. The maximum allowable hea*.up or cooldown rate is 100*F per hour.
.s An additional thermal transient was defined in the RPV Design Specification (Reference 6) to account for ~
potential low pressure coolant injection (LPCI) into the
' recirculation system during a loss of coolant accident (LOCA). The thermal transient was very conservatively defined as a step change in water temperature from 546*F to 90*F at a flow velocity of 10 feet per second. One ,, of these LPCI cycles is assumed to occur every five g
years (Reference 7). Also defined in Reference 7 is a 1 thermal transient based on actual plant operation due to the initiation of shutdown ecoling. The shutdown cooling transient is defined as a 50*F step change in water temperature and it occurs 10 times per year. _ NSP-81-103 11,'
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.5 i 5.0 EVALUATION METHODS AND RESULTS The evaluation of the weld overlay consists','of a code
- stress analysis per References 1 and 2 and a fracture
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mechanics evaluation per Section XI (Reference 8). _ 5.1 Code Stress Analysis The end cap region was assumed to be axisymmetric. That is, the axial crack was conservatively assumed to be 360 degrees around the pipe an6 the effect of the sweepolet
..i was assumed to be negligible (based on a shell intersection analysis). The shrinkage of the weld
_, overlay should therefore have a minimal effect on the w sweepolet. A finite element model of the cracked and 7 weld overlayed region was developed using the ANSYS (Reference 9) computer program. The crack depth was conservatively assumed to be 0.15 inch instead of the measured depth of approximately 0.12 inch. Figure 5.1 shows the model. The pressure stress profile for a design pressure of 1248 psi was calculated with this F model. The results are shown in Figure 5.2. The weld overlay thermal model was also taken to be axi-symmetrical (Figure 5.3). The exterior boundary was NSP-81-103 12 '
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i 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 7 I Charts 16 and 23 of Reference 10. The maximum through h
wall temperature difference was determined to be less 41 than 2*F for the normal startup cycle, 40*F for the initiation of shutdown cooling, and 359'F for the LPCI transient. l The maximum thermal stress for use in the fatigue crack l growth analysis was calculated as follows: (Reference 1)
)
m E a LT y + E n AT 2
"
- 2 (1- v) l- v Where:
6 E = 28.3 x 10 psi (Young's Modulus) a = 9.11 xt10-6 .p-l (Coefficient of Thermal Expansion) eT y = Equivalent Linear Temperature Difference
, eT = Peak Temperature Difference 2
1 The values of eT y, cr2 , and o are given in Table 5.1 for all three thermal transients. NSP-81-103 13. l,
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The'results of a code stress analysis per Reference 1 are given in Table 5.2. The allowable stress values for both References 1 and 2 are also given. The weld 1 overlay repair satisfies the Reference 1 requirements even with the use of the more conservative Reference 2 allowable stress values. A conservative . fatigue analysis per Reference I was per-9 Lormed. In addition to the stress intensification factors required per Reference 1, an additional fatigue strengtn reduction factor of 5.0 was applied due to the crack. The fatigue usage factor was then calculated assuming 10 startups and shutdown cooling initiation cycles per year plus one LPCI injection every five years. The results are summarized in Table 5.2. [ 5.2 Fracture Mechanics Evaluation t. Three types of fracture mechanics evaluations were performed. The allowable crack depth was calculated based on Reference 3. Crack growth due to both fatigue 1 and IGSCC was calculated using the NUTECH computer pro-gram NUTCRAK (Reference 11) with material constants and methodology from References.12 and 13. Finally, the ultimate margin to failure for a crack assumed to pro-NSP-81-103 14 :
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. pagate all the way through the original pipe material to the weld overlay was calculated per References 14 and
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15. 5.2.1 Allowable Crack Depth The allowable depth for a 1 inch long axial crack was determined using Reference 3. The dimensions of the unrepaired pipe were conservatively used. Thus, the ] ratio of applied primary stress to Code allowable stress (S ,) was calculated in the following manner: s, Stress Ratio = EE t_ S, P = 1248 psi (Design Pressure) R = 10.951 inches (Outside Radius of Pipe - before overlay) t = .987 inch (Nominal Pipe Thickness - before overlay) S, = 14,427 psi (B31.1)
= 16,900 psi (Section III) i y
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.~. Substitution yields: I Stress Ratio = .96 (B31.1) , ) =
.82 (Section III) h I
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The average value of .89 was used. The nondimensional _, crack length was calculated in the following manner: b Nondimensional Length = (Rt)l/2 L = 1 inch R = 10.951 inches t = .987 inch
. Substitution yields:
n Nondimensional Le'ngth = .3 Thus per Table IWB-3642-1 of Reference 3, the allowable crack depth is 70 percent of the wall thickness. To be conservative, the unrepaired pipe wall thickness was
used. The allowable crack depth is then 0.69 inch.
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. 5.2.2 Crack Growth Theexisting0.12inchdeepcrackcouldgrohduetoboth fatigue and stress corrosion. Fatigue crack growth due to the three types of thermal transients defined in Section 4.2 was calculated using material properties from Reference 13. The fatigue cycles considered are g
U shown in Figure 5.4. IGSCC growth depends on the total steady state stress. I The steady state stresses can be postulated to be high i due to the presence of weld residual stresses. The magnitude of weld residual stresses without the weld overlay is difficult to determine. Reference 16 gives a measurement of the residual stress through the thickness near a 26-inch butt weld. The weld overlay is expected to reduce the residual stresses, but the magnitude is not known. Future.wprk at the Electric Power Research l Institute (EPRI) is expected to increase our under-l' standing of this reduction. To be conservative, another case with through wall bending residual stress equal to 30,000 psi was also considered. 30,000 psi is the ASME j Code room temperature yield stress fo 304 stainless l' steel. Thus three residual stress distributions were n used: V . l [ NSP-81-103 17 U j,, Revision 0 , e
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, A) Zero stress; assuming the overlay process
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completely eliminates any tensile residual stresses due to the original butt weld. l
.. i B) The best estimate; measured residual stress for the
, original butt weld from Reference 16. n C) A worst case; upper bound residual stress distribution for conservative bounding crack growth calculations. 7 These distributions are presented in Figure 5.5. ? O Two IGSCC growth laws were also considered based on the data compiled in Reference 12. Thus, a total of six combinations of residual stress and crack growth law were investigated... The six cases are summarized in Table 5.3. r l
.y Cases Al, A2, B1 and B2 were analyzed using an infinite b length flaw. Cases C1 and C2 (worst case residual 9
stress) were analyzed using a finite size flaw of 2 inch 1ength. The stress intensity factor (K) versus crack S }- NSP-81-103 18,'
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. l j depth (a) is shown in Figure 5.6 for both infinite and
, finite length flaws for the Case C residual stress H distribution. . n . n - Fatigue crack growth due to the cycles shown in Figure 5.4 for cases A1, A2, B1 and B2 assumed a worst case initial crack depth of 0.5 inch. The total fatigue crack growth was 0.02 inch. Fatigue crack growth due to , the cycles shown in Figure 5.4 for cases C1 and C2 assumed a crack depth based on the finite size flaw (K
] versus a) curve in Figure 5.6. The maximum K occurs for
_.) a crack depth equal to 0.3. The total fatigue crack n i growth for five years of operation due to the cycles shown in Figure 5.4 for cases C1 and C2 with an initial crack of 0.3 was approximately 0.01 inch. The predicted IGSCC and fatigue crack growths for all six cases for the neyt five years are presented in Figure 5.7. Cases A1, A2, B1 and B2, which are the most likely to occur, do' not experience significant crack growth for at least five years. Even the most conservative cases (C1 and C2) with worst case residual stress do not grow to an unacceptable size during the firrt five years. F' pevision 0 ' N*
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. The' initial crack sizes for cases A1, A2, Bl and B2 that m
would be necessary to grow (due to both fatigue and IGSCC) to a depth of 70 percent of the unreinforced pipe (.69 inch) in the next five years are shown in Figure
, 5.8. The time scales in Figures 5.7 and 5.8 are years of operation, not real time years. Thus, an initial - crack depth of greater than 0.4 inch would be acceptable for at least 5 years using the most likely residual stress distributions.
The design life of the repair is thus clearly greater than five years, even considering the worst combinations of analytical conditions considered. A more precise
, determinition of the actual design life of the repair will be possible after completion of the EPRI program to % determine weld overlay residual stress reduction noted above..
5.2.3 Tearing Modulus The largest size to which the existing crack could 8 reasonably be expected to grow was postulated to be a 1 inch radius flaw. This assumes growth of the crack in ~ the radial direction completely through the original .c pipe material to the overlay, even though such propaga-tion is not predicted by the analysis of Section 6.2.2. After such propaga, tion, the assumed crack would
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; he completely surrounded by IGSCC resistant material:
s the weld between end cap and manifold, the weld overlay, and the annealed end cap. A tearing modulus evaluation ? , was then performed for this postulated crack. The only applied load is pressure. The evaluation was performed using the methodology of Reference 14 with material properties from Reference 15. The postulated flaw and the results are shown in Figure 5.9. 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 pressure results in applied J-T combination moving up this line, and unstable fracture is predicted at the intersection of this applied loading line with the material. resistance line.' Figure 5.9 shows that the predicted burst pressure is in
, excess of 6500 psig. Thus, there is a safety factor on 1 normal operating pressure of at least 6, which is well in excess of the safety factor inherent in the ASME ~
Code, even in the presence of this worst case assumed crack. b .
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NORMAL INITIATION SHUTDOWN LPCI STARTUP 7 PARAMETER COOLING CYCLE CYCLE CYCLE
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(CYCLE 1) (CYCLE 2) (CYCLE 3) 0 l EQUIVALENT 2F 32 F 290 F 0 LINEAR TEMPEP.ATURE AT r, I 0 PEAK 0 8F 69 F x TEMPERATURE AT 2 THROUGH 368 PSI 8839 PSI 78,817 PSI WALL THERMAL STRESS o I . i a
, Table 5.1 THERMAL STRESS RESULTS NSP-81-103 ,
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I ACTUAL CATEGORY EQUATION STRESS SECTION III B31.1 NUMBER OR NB ALLOWABLE ALLOWABLE THICKNESS S NA Sg = I6,900 PSI Sh =14,427 PSI S c" ', ~ REQUIRED (1) 1.308" 0.826" 0.952" THICKNESS PRIMARY (9) 5,470 PSI
- 25,350 PSI 14,427 PSI PRIMARY + (10) 12,060 PSI 50,700 PSI 49,765 PSI SECONDARY PEAK (11)
CYCLE I (15,150)5 NA NA CYCLE 2 (27,484)5 CYCLE 3 (130,719)5
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USAGE 0.22 FACTOR (40 YR). ~
- FINITE ELEMENT MODEL GJVES 9920 PSI FOR MAXIMUM STRESS INTENSITY.
EQUATION (9) CALCULATES AXIAL STRESS WHICH IN THIS CASE IS NOT LIMITING AS MOMENTS = 0. 9
. Table 5.2 CODE STRESS ALLOWABLES 22" END CAP ,
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CASE f GROWTHLAWh 4 Al A 1.843 x 10-12 g .615* 4.116 x 10-12 g .615" 4 A2 A 1,843 x 10-12 g4.615
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B1 B 4 B2 8 4.116 x 10-12 g .615 4 C1 C 1.843 x 10'I2 K .615 4 C2 C 4.116 x 10-12 K .615
- BEST ESTIMATE EPRI NP 2423-LD JUNE 1982 0.2 ppm DATA.
[ **UPPERBOUNDEPRINk5423-LDJUNE 1982 0.2 ppm DATA. I 3 L Table 5.3 CRACK GROWTH CASES F- t j ' NSP-81-103 .
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/ PREP 7 ANSYS MODEL OF 22" END CAP WELD OVERLAY NSP-81-103 :
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C 6 v., 4-m f5 m 2. t 0 BASE OF .'5 1.0 1.5 DISTANCE CRACK , (INCHES) OUTSIDE OF WELD OVERLAY l l? F Figure 5.2 APPLIED STRESS PROFILE THROUGH LIMITING SECTION 22" END CAP i 4 , NSP-81-103 *
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I - INSULATION WELD
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.I < . F SECTION A-A f Figure 5.3 WELD OVERLAY THERMAL MODEL f .,' NSP-81-103 .
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i l LPCI - l i SHUTDOWN , m COOLING m 5 STARTUP SHUTDOWN -
, NORMAL - i OPERATION I RESIDUAL - l \
50 50 5 CYCLES CYCLES _ YEARS _ _
- g. _
; ENTIRE e 1' SEQUENCE CYCLE REPEATS TIME l
I Figure 5.4 THERMAL TRANSIENTS
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E
.l 30 ~' 27.5 -
20 - A (NO STRESS) 10 - 2 0 ID OD m a
, g S ' '
O B (BEST
" ESTIMATE)
C (WORST 1 i' i b l
-40 Figure 5.5 l- CRACK GROWTH RESIDUAL STRESS .
l' 22" END CAP F~ 4 i NSP-81-103
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.3 70 -
STRESS INTENSITY FACTOR, X VS CRACK DEPTH, a T' 60 - - - -- E = " t = 2 inches
.. . /
- l. f',; T = 1.0 in-f ae /
~fL /
50 - / PIPE DIAMETER = 22 in. /
/
/
/
40 -
/
7 j c - / v / x 30 - / a
/
- /
/
20 - !
/ '
/
, /
i 10-I O,I 0.2 0.3 0.4 0.5 0.6
~
0 0.1 a (inches) (( Figure 5.6 f NSP-81-103
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~' I.50 - OD OF x REINFORCED PIPE 1.20 -
._ 03RkA,Y,,g[,3LpAh - -
N 3t, 0.90 - iE LIMIT FOR ~ b UNREINFORCED 0.60 - PIPE C2 C1 ~' O.30 - CASE A1, A2, B1, B2 ' O.12 t - -' O 1 2 3 4 5 TIME (years) I- % C1, C2 ARE RESPECTIVELY THE BEST ESTIMATE AND UPPER BOUND IGSCC CURVES FOR A FINITE LENGTH SEMI-ELLIPTICAL SURFACE CRACK. FATIGUE CRACK GROWTH OF 0.01 INCH IN 5 YEARS IS INCLUDED. Figure 5.7 CRACK GROWTH 22" END CAP
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- 1. 2 -
I WELD OVERLAY 0.9 .
' A1
; E ALLOWABLE DEPTH E 70% OF WALL B2
- 0.6 - 81 A2
[ k o 0.3 - I 0- , , , , 0 1 2 3 4 5 TIME (years) I b INCLUDES 0.02 INCH OF FATIGUE CRACK GROWTH IN 5 YEARS. Figure 5.8 ALLOWABLE CRACK DEPTH 22" END CAP
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'N .
- el h
240 - Ci
~
200 - N o 160 - J =14,000 I"ID
- C in2 N
5 s 120 -
. :S PRESSURE =
4000 PSI inib y J =6000
~ in2
' 80 - PRESSURE =
? 5500 PSI PRESSURE =
6500 PSI PRESSURE = 6000 PSI
^ PRESSURE = 5000 PSI O , , , , , ,
0 40 80 120 160 200 240 T OVERLAY WELD w N WELD ANNEALED CAP 1" RADIUS FLAW
-- Figure 5.9 TEARING MODULUS 22" END CAP
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SUMMARY
AND CONCLUSIONS
- 1 The evaluation of t,he repairs to the recirculation end cap reported herein shows that the resulting stress 1cvels are acceptable for all design conditions. The stress levels have been assessed from the standpoint cf'
- load capacity of the components, fatigue, and resistance to crack growth. 1 Acceptance criteria for the analysis have been C established in Section 3.0 of this report which demonstrate that: ? L.I - 1. There is no loss of design safety margin over those provided by both the original Construction Code for the piping system (B31.1) or the current Code of Construction for Class 1 piping (ASME Section III). L
- 2. During the design lifetime of the repair, the fT '
observed cracks will not grow to the point where the above safety margins would be exceeded. 6 Analyses have been performed and results are presented ,
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which demonstrate that the repaired weld satisfies these criteria by a large margin, and that the design life of the repair is at least five years.
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- l. 7.0 REFERL!sCES I 1. ASME Boiler and Pressure Ves'Sel Code Section III, Subscetion NB, 1977 Edition with Adden'da through Summer 1978.
- 2. USA Standard gCode f or Pressure Piping, " Power Piping", USAS B31.1.0 - 1967.
1 3.t ASME Boiler and Pressure Vessel Code Section XI, Article IWB-3640 (Proposed), " Acceptance Criteria for Flaws in Austenitic Stainless Steel Piping" (Presented to section XI Subgroup on Evaluation Standards in September 1982).
- 4. " Design Report Recirculation System Monticello Nuclear Power Station", General Electric Document Number 22A2603 Rev. 1.
I~ . L 5. "NUTECH Reanalysis of the Reactor Recirculation [ ! Piping System," Letter to S. J. Hammer from G. A. i Wiederstein, GAW-82-014, File Number 30.2354.0003. l l - NSP-81-103 35 '
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. 6. Purchase Specification for Monticello Reactor Pressure Vessel, General Electric Document Number 21Alll2, Revision 6.
- 7. Telecon between NUTECH (J. E. Charnley and N. Eng) and NSP (S. J. Hammer), " Weld Overlay Repair Program Technical Issues," dated October 20, 1982, File 30.1281.0001.
- 8. ASME Boiler and Pressure Vessel Code Section XI, 1977 Edition with Addenda through Summer 1978.
- 9. ANSYS Computer Program, Swanson Analysis Systems,
_ Revision 3. l
- 10. Schneider, P.J. " Temperature Response Charts", John Wiley and Sons, 1963.
t
- 11. NUTCRAK Computer Program, Revision 0, April 1978, File Number 08.039.0005.
1
. 12. EPRI-2423-LD " Stress Corrosion Cracking of Type 304 Stainless Steel in High Purity Water - a Compilation of Crack Growth Rates", June 1982.
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- 13. EPRI-NP-2472, "The Growth and Stability of Stress Corrosion Cracks in Large-Diameter BWR Piping,"
July 1982.
\
- 14. NUREG-0744 Vol. 1 for Comment, " Resolution of the Reactor Materials Toughness Safety Issue".
] 1 15. EPRI-NP-2261, " Application of Tearing Modulus n Stability Concepts to Nuclear Piping", February 1982.
)
- 16. EPRI-NP-1413, " Measurement of Residual Stresses in Type 304 Stainless Steel Piping Butt Weldments",
_ June 1980. 89 O q 4 M s s 1 NSP-81-103 37 ,
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