ML20066K002

to Design Rept for Recirculation Line Safe-End & Elbow Repairs,Monticello Nuclear Generating Plant

ML20066K002
Person / Time
Site:	Monticello
Issue date:	11/30/1982
From:	Charnley J, Eng N, Rich Smith NUTECH ENGINEERS, INC.
To:
Shared Package
ML20066J960	List: ML20066J958 ML20066K002 ML20066K066
References
IEB-82-03, IEB-82-3, NSP-81-105, NSP-81-105-R, NSP-81-105-R00, NUDOCS 8211290349
Download: ML20066K002 (48)

Text

Att:clunent (3)

Director I&E, NRC November 22, 1982 NSP-81-105 Revision 0 November 1982 30.1281.0105 DESIGN REPORT FOR RECIRCULATION LINE SAFE END AND ELBOW REPAIRS MONTICELLO NUCLEAR GENERATING PLANT Prepared for:

Northern States Power Company Prepared by:

NUTECH Engineers, Inc.

San Jose, California Prepared by: Reviewed by:

. 21m l-J. E. Charnley, P.E. R. H. Smith Project Engineer Project Quality Assurance Engineer Approved by: Issued by:

h' OI M

, 4y" P . C. Riccardella, P.E. N. Eng f Senior Director Project Manager 8211290349 821122 PDR G

ADOCK 05000263 PDR nutech

==

REVISION CONTROL SHEET TITLE. Design Report for Recirculation REPORT NUMBER: NSP-81-105 Line Safe End and Elbow Repairs, Revision 0 Monticello Nuclear Generating Plant J. E. Charnley / Principal Engineer O ME / TITLE E INITIALS If.Ccftlan ff %WYCm? %

P. C. Riccard lla/ Senior Director Mh

• NAME / TITLE INITIALS S. Kulat/ Consultant I  ?/) l/,,

NAME/ TITLE INITI ALS H. L. Gustin / Engineer N AME / TITLE INITIALS Y. S. Wu/ Consultant I yf%[

N AME / TITLE INITIALS PREPARED ACCURACY CRITERIA

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REVISION CONTROL SHEET (CONTINUATION) E-" Design Report for Recirculation REPORT NUMBER: NSP-81-105 Line Safe End and Elbow Repairs' Revision 0 Monticello Nuclear Generating Plant PREPARED ACCURACY CRITERIA PAGE(53 REV # SY / DATE CHECK 8Y / DATE CHECK BY / 0 ATE 14 0 bC "l l lI5N 'I l# Y'? 15 . 16 1 17 18 M ' 19 20 se 21 y l/47-g2 22 YbW lI-11-82 23 }l M st-r7-sa 24 25 , 26 27 l 28 I' ' 29 ,

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CERTIFICATION BY REGISTERED PROFESSIONAL ENGINEER I hereby certify that this document and the calculations 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 States of Minnesota and California and am competent to review this document. Certified by: s

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• Registration No. 14372 State of California Registration No. 16340 Date $] lf$ L NSP-81-105 iv Revision 0 nut.qch

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 7 3.1 Strength Evaluation 8 3.2 Fatigue Evaluation 8 3.3 Crack Growth Evaluation 9 4.0 LOADS 11 4.1 Mechanical and Internal Pressure Loads 11 4.2 Thermal Loads 12 5.0 EVALUATION METHODS AND RESULTS 14 5.1 Code Stress Analysis 14 5.2 Fracture Mechanics Evaluation 16 5.2.1 Allowable Crack Depth 17 5.2.2 Crack Growth 18 5.2.3 Tearing Modulus 21 5.3 Effect on Recirculation System 23 6.0

SUMMARY

AND CONCLUSIONS 38

7.0 REFERENCES

39 i NSP-81-105 v Revision 0 nutggh ,

i 9 i .i

i i

{ LIST OF TABLES

Number Title Page i

l 5.1 Thermal Stress Results 26 5.2 Code Stress Allowables 12" Safe End and Elbow 27 l l 5.3 Crack Growth Cases 28 i i i ? ~!

?

l 4 NSP-81-105 vi Revision 0 O j

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

LIST OF FIGURES Number Title Page 1.1 Conceptual Drawing of Recirculation System 3 2.1 Schematic of Safe End to Pipe Weld Overlay 5 1 2.2 Schematic of Elbow to Pipe Weld Overlay 6 5.1 ANSYS Model of 12" Safe End and Elbow 29 Weld Overlay 5.2 Applied Stress Profile Through Limiting 30 Section 12" Safe End and Elbow 5.3 Thermal Transients 31 5.4 Radial Crack Growth 12" Safe End and Elbow 32 5.5 Axial Crack Growth 12" Safe End and Elbow 33 5.6 Tearing Modulus 12" Safe End and Elbow 34 5.7 Actual Geometry 35 5.8 Revised ANSYS Model 36 5.9 Deformed Geometry 37 NSP-81-105 vii Revision 0 nutach-

1.0 INTRODUCTION

This report summarizes evaluations performed by NUTECH to assess weld overlay repairs of recirculation inlet safe end and elbow welds at Northern States Power Company's Monticello Nuclear Generating Plant. Weld overlay repairs have been applied to address leakage and additional ultrasonic and radiographic examination results believed to be indicative of intergranular t 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 one fuel cycle. The amount that the actual design life exceeds one fuel cycle will be established by a combination of future analysis and testing. Leakage was observed adjacent to three safe end to pipe welds (RREJ-3, RRFJ-3, and RRCJ-3). In addition, small crack indications have been detected adjacent to a riser to elbow weld (RRDJ-5). All four of these welds were repaired with the weld overlay design evaluated in this report. NSP-81-105 1 Revision 0 nute_cb

i Figure 1.] shows the safe ends and the elbow in relation to the reactor pressure vessel (RPV) and other portions of the recirculation system. 1 l The existing pipe material is ASTM A358, Class 1, Type 304. The existing safe end material is SA336, Grade F8. The existing elbow material is ASTM, A240, Type 304. 1 i c i d

;                  NSP-81-105                           2 Revision 0

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INDICATICNS IN SAFE CID IN0! CATIONS IN EL3CW TO PIPE 'AE.0 (RREJ-3) INDICATICNS IN SAFE TO PIPE ' WELD (REUJ-5' CiD TO PIPE ' WELD (RRFJ-3)

 ^                                                   INDICAT!CNS IN SAFE C:D fM         TO PI?E 'WEL3 (RRCJ-3)                 g       ,

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l ad l LCCP A l l l Figure 1.1 CONCEPTUAL DRAWING OF RECIRCULATION SYSTEM NSP-81-105 Revision 0 3 nut _ech l ~

2.0 REPAIR DESCRIPTION The through wall cracks and other indications around and to both sides of 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 and 2.2. The weld deposited band over the through wall crack will provide wall thickness equal that required for the adjacent uncracked piping. In addition, the weld metal deposition will produce a f avorable compressive residual stress pattern and the weld metal will be type 308L, which is resistant to propagation of IGSCC cracks. i NSP-81-105 4 , Revision 0 nutg_qh

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6b 3 L oS AS WELDED SURFACE ACCEPTABLE FOR OVERLAY TAPER TRANSITIONS 6.0" MIN , TYPE 308L WELD OVERLAY e 3.38" n =

                                                                                                -0.54" MIN J                     s'                     f                  NOMINAL (TYP.)
                                                                                                                '       -0.74"
                                                                                                                                                                                    /

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TilROUGilWALL RADIAL CRACK (RREJ 3. HRCJ 3 AND RRFJ 3) PIPE CENTERLINE , l , 12" SCllEDULE 80 PIPE SAFE END TO SAFE END TO PIPE WELD NOZZLE WE LD D C Figure 2.1 g SCllEMATIC OF SAFE END TO PIPE WELD OVERLAY y (TilERMAL SLEEVE OMITTED)

e LONG AS WELDED SURFACE RADIUS ACCEPTABLE FOR ELBOW OVERLAY TAPER TRANSITIONS h RADIAL CRACK APPROXIMATELY 90% THROUGH WALL PIPE TO ELBOW WELD

                           ~

a- /

as I

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ci E I L  % 12" SCHEDULE y y 80 PIPE V / ELDO RLAY ' 18 NOMINAL (TYP)

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Figure 2.2 SCHEMATIC OF ELBOW TO PIPE WELD OVERLAY

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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 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. NSP-81-105 7 Revision 0 nutp_qh

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, Class 1 (Reference 1) analysis of the safe end weld overlay repairs was performed using the worst case loads for any recirculation inlet safe end.

An ASME Boiler and Pressure Vessel Code, Section III, Class 1 analysis of the elbow weld overlay repair was performed using the worst case loads for any recirculation inlet elbow.

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 Fatigue Evaluation The stress values obtained from the above strength eval-uation were combined with thermal and other secondary NSP-81-105 8 Revision 0 nutp_gj]

4

                                                               ~

s r*' stress conditions to demonstrate adequate fatigue resistance for the design life of each repair. The criteria for fatigue evaluation include:

1. The maximum range of primary plus secondary stress I was compared to the secondary stress limits of E '

Reference 1. _ L -

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2. The peak alternating stresa 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 conventional fatigue analysis dechniques. 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 desion life of each repair. Allowablenumberofc/cleswasdeterminedfromthe stainless steel fatigue curve of Reference 1. 4 3.3 Crack Growth Evaluation Crack growth due to both fatigue (cyclic stress) and

                          \                                              :'

{ IGSCC (steady state stress) was calculated. The allow-able crack depth was established based on net section l NSP-81-105 9 Revision 0 nut 9_Ch

l 9,.

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f limit load for the cracked and repaired safe end and elbow welds (Reference 2). J The design life of each repair was established as the minimum of either the predicted time for the observed i through wall crack to grow to the allowable crack depth

 +                      or five years.

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4.0 LOADS The loads considered in the evaluation of the safe end and elbow weld overlay repairs 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 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 stresses are considered in the crack growth analyses and are described in Section 5.2.2. 4.1 Mechanical and Internal Pressure Loads The design pressure of 1248 psi for tha recirculation system was obtained from Reference 3. The dead weight and seismic loads applied to the safe end and elbow welds were obtained from the recent NUTECH analysis of the Monticello Reactor Recirculation System piping (Reference 4). The highest loads for any recirculation inlet safe end were applied to the safe end weld overlay. Thus the safe end weld overlay analysis applies to all recirculation inlet safe ends. The highest loads for any recirculation inlet elbow were NSP-81-105 11 Revision 0 nutgLCh

t applied to the elbow weld overlay. Thus, the elbow weld overlay analysis applies to all recirculation inlet elbows. 4.2 Thermal Loads The thermal expansion loads for the highest loaded recirculation inlet safe end and elbow were also obtained from Reference 4 and applied to the weld overlay repairs. The only transient thermal condition defined in Reference 4 that occurs at the safe ends or elbows is the normal startup and shutdown cycling. The maximum allowable heatup or cooldown rate is 100*F per hour. An additional thermal transient was defined in the 'PV Design Specification (Reference 5) to account for potential low pressure coolant injection (LPCI) into the l recirculation system during a loss of coolant accident (LOCA). The thermal transient was very conservatively 1 i defined as a step change in water temperature from 546*F 1 l l to 90*F at a flow velocity of 10 feet per second. One of these LPCI cycles is assumed to occur every five years (Reference 6). Also defined in Reference 6 is a NSP-81-105 12 Revision 0 nut _ec_h

i

   .   +

.. I thermal transient based on actual planc operation due to the initiation of shutdown cooling. The shutdown cooling transient is defined as a 50*F step change in water temperature and it occurs 10 times per year. f i

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5.0 EVALUATION METHODS AND RESULTS The evaluation of the weld overlay repairs consists of a code stress analysis per Reference 1 and a fracture mechanics evaluation per Section XI (Reference 7). 5.1 Code Stress Analysis The weld overlayed regions were assumed to be axisymmetric. That is, a through wall radial crack was conservatively assumed to be 360 degrees around the pipe and one inch long centered on the existing safe end and elbow welds. Thus the assumed crack conservatively envelopes all observed cracks in the safe end and elbow welds. In addition, all analyses were conservatively performed using a weld overlay thickness of 0.50 inch which is seven percent smaller than the actual minimum thickness of 0.54 inch. A finite element model of the cracked and weld overlayed region was developed using the ANSYS (Reference 8) computer program. Figure 5.1 shows the model. The pressure stress profile for a design pressure of 1248 psi was calculated with this model. The results are shown in~ Figure 5.2. NSP-81-105 14 Revision 0 nutp_q, h

4 The ANSYS model was also used for the rapid thermal transients. The exterior boundary was assumed to be insulated. The temperature distribution in the weld overlay subject to the normal start up cycle defined in Section 4.2 can be readily calculated using Chart 16 of Reference 9. The maximum through wall temperature difference was determined to be less than 2'F for the normal startup cycle. The maximum thermal stress for use in the fatigue crack growth analysis was calculated by hand using the method in Reference 10 for the normal startup transient and was calculated directly by the ANSYS model for the LPCI and i shutdown cooling transients. The results are given in Table 5.1 for all three thermal transients. The through wall thermal stresses listed in Table 5.1, which were calculated without the beneficial presence of the secondary thermal sleeve, are classified as peak stresses. Thus the only ASME code limit for them is the fatigue usage factor. The results of a code stress analysis per Reference 1 are given in Table 5.2. The allowable stress values

given in Table 5.2 are based on Reference 1, Article NB 3600 limits. All results apply to the most limiting NSP-81-105 15 Revision 0 nutp_qh

 =

case of either the safe end or elbow weld overlay repair. A conservative fatigue analysis per Reference 1 was per-formed. In addition to the stress intensification 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 calcu' lated 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 Three types of fracture mechanics evaluations were performed. The allowable crack depth was calculated based on Reference 3. Crack growth due to both fatigue and IGSCC was calculated using the NUTECII 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-pagate all the way around the original safe end or elbow material to the weld overlay was calculated per References 14 and 15. All analyses summarized below NSP-81-105 16 Revision 0 nutgch

t

       =

1 apply to the most limiting case of either the safe end or elbow weld overlay repair. 5.2.1 Allowable Crack Depth t The allowable depth for a 1 inch long radial crack was a determined using Reference 2. The dimensions of the limiting section of the safe end repair were used. Thus, the ratio of applied primary stress to Code allowable stress (Sm) was calculated in the following manner: Stress Ratio = PR/t P = 1248 psi (Design Pressure) i R = 6.915 inches (Outside Radius of Overlay) t = 1.187 inch (Minimum Thickness of Pipe plus Overlay) S, = 16,900 psi (Table 5.2) Substitution yields: Stress Ratio = .43 NSP-81-105 17 Revision 0 nutp_cb

The nondimensional crack length was calculated in the following manner: b Nondimensional Length =

                                              .(Rt)l/2 L  = 1 inch R  = 6.915 inches t  = 1.187 inch Substitution yields:

Nondimensional Length = .35 Thus per Table IWB-3642-1 of Reference 2, the allowable crack depth is 75 percent of the wall thickness. The allowable crack depth is then 0.89 inch. 5.2.2 Crack Growth The existing through wall crack could grow due to both fatigue and stress corrosion. Fatigue crack growth due l 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 shown in Figure 5.3. NSP-81-105 18 Revision 0 nutp_qh

The steady state axial stresses in the weld overlay are significantly higher than the hoop stresses. Thus fatigue crack growth will be predominately in the radial direction. The model used to calculate fatigue crack growth conservatively assumed that the through wall crack was a circumferential crack 360* around the weld. Thus the results from this model are conservative compared to the actual case of a short axial through wall crack. The fatigue crack growth law used is given in Table 5.3. Steady state axial stress due to pressure, dead weight, thermal expansion and weld residual stress were considered. The stress due to pressure, dead weight and thermal expansion were obtained from the ANSYS model. The weld residual stress due to the original weld, combined with the through wall crack and the weld overlay repair is difficult to estimate. It was judged that the weld residual stress due to the original weld was significantly reduced by the through wall crack and the weld overlay repair. Thus two bounding weld residual stress distributions were considered. The first distribution was zero residual stress and the second was a through wall bending stress of 36 ksi, with compression on the inside surface. NSP-81-105 19 Revision 0 nute_cb

Fatigue crack growth is not a strong function of steady state stresses such as weld residual stress. Thus the fatigue crack growth for both assumed weld residual stress distributions were similar. The total radial crack growth due to fatigue for five years of operation (Figure 5.3) was determined to be approximately 0.005 inch, which is well below the allowable crack growth of 0.24 inch. The model that was used and the fatigue crack growth as a function of time are shown in Figure 5.4. The existing crack will not grow due to IGSCC into the IGSCC resistant weld overlay. However, it could grow axially due to the average value across the thickness of the steady state hoop stresses. The average value of weld residual stress across the thickness is zero for all three residual stress distributions used in i Reference 10 as well as for the two distributions described above. The hoop stress is caused only by pressure. Two IGSCC growth laws were considered based on the data compiled in Reference 12. The growth laws I are given in Table 5.3. The crack was assumed to be through wall and one inch long as shown in Figure 5.5. l It should be noted that the weld overlay will help NSP-81-105 20 Revision 0 nut.ech

                                                                             = _ _

    =

arrest crack propagation, but the extent of this beneficial effect is not known. For this reason, the crack was conservatively modeled as a one inch long through wall crack in a 0.622 inch thick infinite plate. The plate thickness is that of the pipe without the overlay for analysis purposes. Using this conservative model, the axial IGSCC growth of the crack for five years of operation was determined to be approximately 0.009 inch. When this small increase in crack length is acded to the actual crack configura-tion (Figure 5.5), the crack is still well within the limits of the weld overlay. The IGSCC growth as a function of time is also shown in Figure 5.5. 5.2.3 Tearing Modulus The largest size to which the existing crack could reasonably be expected to grow was postulated to be a 0.75 inch radius flaw. This assumes growth of the crack in the axial direction, even though such propagation is not predicted by the analysis of Section 5.2.2. After such propagation, the assumed crack would be completely surrounded by IGSCC resistant material: the weld between safe end (elbow) and pipe, the weld overlay, and l NSP-81-105 21 Revision 0 t nutggb

the safe end (elbow). A tearing modulus evaluation was then performed for this postulated crack. The applied loads are pressure, dead weight, seismic and thermal expansion. 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 5.6. The upper dotted line represents the inherent material resistance to onstable fracture in terms of J-integral and Tearing Modulus, T. The '.ine originating at the origin represents the appl. d loading. Increasing load 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 5.6 shows that the predicted burst load is in excess of five times the actual loading. Thus, there is a safety factor for normal loads including OBE seismic of at least five, which is well in excess of the safety factor inherent in the ASME Code, even in the presence of this worst case assumed crack. The analogous safety NSP-81-105 22 Revision 0 nutgch

factor for SSE seismic is also well in excess of that required by the ASME Code for low probability events. 5.3 Effect on Recirculation System Installation of the weld overlay repairs will cause a small amount of radial and axial shrinkage underneath the overlay. Based on measurements of a welding mockup, the maximum radial shrinkage will be 3/16 inch and the maximum axial shrinkage will be 1/64 inch. These measured shrinkages are conservative compared to the expected actual overlay shrinkage because the weld overlay thickness of the mockup was 0.9 inch compared to the actual weld overlay thickness 0.54 inch. The effect on the recirculation system of the maximum expected weld shrinkage was evaluated. Figure 5.7 shows the configuration of the safe end, nozzle and thermal sleeve without an overlay. The effect of the shrinkage on the low alloy steel nozzle, the thermal sleeve and the crotch of the safe end was determined by extending the ANSYS model,to include these areas. The model was also extended to the centerline of the 12" riser pipe. The revised ANSYS model is shown in Figure 5.8. NSP-81-105 23 Revision 0 nutgch

The measured shrinkages from the weld mockup were imposed as boundary conditions on this model. The resulting stresses are steady state secondary stresses (similar to other weld residual stresses) and thus are not limited by the ASME Code (Reference 1) . A plot of the deformed model is shown (greatly exaggerated) in Figure 5.9. The most significant results of this analysis are listed below.

1) The calculated longitudinal displacement at the centerline of the 12" riser is 0.004 inch, which induces a stress of less than 1.0 ksi at the sweepolet to manifold weld. There is no radial displacement at this location.

2) The displacements at the weld between the safe end and the low alloy nozzle are less than 0.001 inch and the induced stresses are less than 1.0 ksi.

3) The radial displacement at the crotch of the safe end is 0.005 inch and the axial displacement is less than 0.001 inch. The induced stresses are less than yield stress.

NSP-81-105 24 Revision 0 nutp_qh

4) The maximum compressive strain underneath the weld overlay is approximately three percent which agrees with the imposed radial displacement divided by the 6

radius (3 9

                              = 0.03). All the parts of the safe end and thermal sleeves (ring nut, plate spring, set screws, thermal sleeve and safe end) were fabricated from either 304 stainless steel or Inconel X-750 (References 16 and 17). Both of these materials can withstand compressive three percent steady state secondary strain without significant deleterious effect.

The threads between the safe end and the ring nut will be forced tightly together. It will be almost impossible to unthread the ring nut, however the only time when ring nut removal is required is during safe end replacement. If this becomes necessary, then the ring nut and safe end can both be removed by cutting.

5) All stresses are below yield stress at an axial distance from the centerline of the safe end to pipe weld of greater than 8.0 inches. Thus, the significant effects of the weld overlay are limited to the region within approximately four inches of the ends of the overlay.

NSP-81-105 25 Revision 0 nutp_qh

l NORMA INITIATION SHUTDOWN LPCI PARAMETER COOLING CYCLE ShRTP CYCLE (CYCLE 1) (CYCLE 2) (CYCLE 3) EQUIVALENT 2F LINEAR *

• TEMPERATURE AT y PEAK 0 *
• TEMPERATURE AT 2

THROUGH 368 PSI 6770 PSI 61,730 PSI WALL THERMAL STRESS o VALUES NOT EXPLICITLY DETERMINED AS TRANSIENTS WERE EVALUATED WITH ANSYS MODEL. Table 5.1 THERMAL STRESS RESULTS NSP-81-105 Revision 0 26 q nutgch

ACTUAL EQUATION STRESS SECTION III CATEGORY NUMBER NB ALLOWABLE OR THICKNESS , S NA S,= 16,900 PSI REQUIRED (1) 0.54" 0.50" THICKNESS PRIMARY (9) 23,170 PSI 25,350 PSI (10) 39,110 PSI 50,700 PSI hRY PEAK CYCLE 1 (40,678)5 (11) (29,964)5 CYCLE 2 CYCLE 3 (84,920)5 USAGE FACTOR d.'60 1.0 (40 YR)

• THE FACTOR OF 5 IS THE CONSERVATIVELY ASSUMED FATIGUE STRENGTH REDUCTION FACTOR.

l Table 5.2 CODE STRESS ALLOWABLES 12" SAFE END AND ELBOW NSP-81-105 Revision 0 27 nut,OM l - - -

4 . o GROWTH LAW CASE da or da dT dN FATIGUE

• 2.84 x 10-8 g2.57 4

IGSCC BEST ESTIMATE ** 1.843 x 10-12 K .615 4 IGSCC WORST CASE *** 4.116 x 10-12 g .615

• EPRI NP2423-LD JUNE 1982 8 ppm 0 DATA 2

                    ** BEST ESTIMATE EPRI NP2423-LD JUNE 1982 0.2 ppm 0 DATA 2
                   **" UPPER BOUND EPRI NP2423-LD JUNE 1982 0.2 ppm 0 DATA 2

Table 5.3 CRACK GROWTH CASES NSP-81-105 Revision 0 28 nutpch

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                                    / PREP 7 Figure 5.1 ANSYS MODEL OF 12" SAFE END AND ELBOW WELD OVERLAY NSP-81-105 Revision 0 nut.e_ch

15 1 12 - STRESS DUE TO PRESSURE 9-C 2

         ~

6-G u a 3-

                               .5          1.0           1.5 DISTANCE (INCHES)

L INSIDE BASE OF CUTSIDE OF 0F PIPE CRACK WELD OVERLAY

             -2 Figure 5.2 APPLIED STRESS PROFILE THR00GH LIMITING SECTION 12" SAFE END l     NSP-81-105

! Revision 0 l l 3o q nute_ch

LPCI - SHUTDOWN _ m COOLING l0 ' STARTUP SHUTDOWN - NORMAL - OPERATION f RESIDUAL - I 50 50 5 CYCLES CYCLES YEARS

; ENTIRE 1 SEQUENCE CYCLE REPEATS I TIME 1

Figure 5.3 THERMAL TRANSIENTS NSP-81-105 l Revision 0 31 nutgqh

1.50 1.25- OD OF WELD OVERLAY

                         ~

ALLOWABLE D'DTH

            .5 j'h' CRACK GROWTH S                                 ID OF WELD OVERLAY
                 '0.50-0.25-0              .           .         .       .

0 1 2 3 4 5 TIME (years)

                                        '   _    3.38"    _

WELD OVERLAY PIPE SAFE END WELD CRACK Figure 5.4 RADIAL CRACK GROWTH 12" SAFE END AtlD ELBOW tiSP-81-105 Revision 0 32 nut.e._qh

1.50 1.25-U j 1.00 0

               @  0.75-5 a

0.50-0.25-0 , , , , 0 1 2 3 4 5 TIME (years)

                                    '       3.38" WELD OVERLAY 4

SAFE END PIPE [ l CRACK

                                      ~ LENGTH /2 I

(WELD 1 l l Figure 5.5 AXIAL CRACK GROWTH 12" SAFE END AND ELBOW f NSP-81-105 Revision 0 33 nutp_qh

240-200 -

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4 x NORMAL LOADS 120 - [, JC = 10,000 I"If 80- 4.5 x NORMAL LOADS 40 - JIC = 6000inI"If 5 x NORMAL LOADS 0 , , , , , , 0 40 80 120 160 200 240 T OVERLAY WELD ANNEALED SAFE END 0.75" RADIUS WELD FLAW Figure 5.6 TEARING MODULUS 12" SAFE END AND ELBOW NSP-81-105 Revision 0 34 nutgqh

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8 .'. OS {12" RISER I LOW ALLOY STEEL -

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                                                                                                 . L, '_.m MAXIMUM DISPLACEMENT IS 3/16 INCH 3

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SUMMARY

AND CONCLUSIONS The evaluation of the repairs to the recirculation elbow and safe end reported herein shows that the resulting stress levels are acceptable for all design conditions. The stress levels have been assessed from the standpoint of load capacity of the components, fatigue, and 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 those provided by both the original Construction Code for the piping system (B31.1) or 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 criteria by a large margin, and that the design life of each repair is at least five years. NSP-81-105 38 Revision 0 nutech t

7.0 REFERENCES

1. ASME Boiler and Pressure Vessel Code Section III, Subsection NB, 1977 Edition with Addenda through Summer 1978.

2. 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).

3. " Design Report Recirculation System Monticello Nuclear Power Station", General Electric Document Number 22A2603, Revision 1. .

4. "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 S. Purchase Specification for Monticello Reactor Pressure Vessel, General Electric Document Number 21Alll2, Revision 6. 1 l l l NSP-81-105 39 Revision 0 l l nutsch

           , - - --,_----w_                               _   , _ _ ,                 -y,--

6. 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.

7. ASME Boiler and Pressure Vessel Code Section XI, 1977 Edition with Addenda through Summer 1978.

8. ANSYS Computer Program, Swanson Analysis Systems, Revision 3.

9. Schneider, P.J. " Temperature Response Charts", John Wiley and Sons, 1963.

I

10. NUTECH Report NSP-81-103, Revision 0, " Design Report for Recirculation Line End Cap Repair, Monticello Nuclear Generating Plant."

11. NUTCRAK Computer Program, Revision 0, April 1978, File Number 08.039.0005.

12. EPRI-2423-LD " Stress Corrosion Cracking of Type 304 S' ainless Steel in High Purity Water - a Compilation of Crack Growth Rates", June 1982.

i I i NSP-81-105 40 Revision 0 l l nutp_qb

i

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."

15. EPRI-NP-2261, " Application of Tearing Modulus Stability Concepts to Nuclear Piping," February 1982.

16. Recirculation Inlet Nozzle Drawing, Chicago Bridge and Iron Company Contract Number 9-5624, General Electric VPF-1811-110.

17. Ring Nut and Plate Spring Drawing, Chicago Bridge and Iron Company Contract Number 9-5624, General Electric VPF-1811-2642.

i NSP-81-105 41 Revision 0 nutp_Ch

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