ML20203N564
| ML20203N564 | |
| Person / Time | |
|---|---|
| Site: | Yankee Rowe |
| Issue date: | 09/30/1986 |
| From: | Bak W, Emerson R, Kulat S ABB IMPELL CORP. (FORMERLY IMPELL CORP.) |
| To: | |
| Shared Package | |
| ML20203N562 | List: |
| References | |
| TASK-03-05.A, TASK-3-5.A, TASK-RR 09-0570-0055, 09-0570-0055-R00, 9-570-55, 9-570-55-R, NUDOCS 8610090258 | |
| Download: ML20203N564 (41) | |
Text
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I LEAK-BEFORE-BREAK EVALUATION FOR THE 5-INCH-DIAMETER BYPASS LINES AT THE YANKEE NUCLEAR POWER STATION l
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Prepared for: l l
YANKEE ATOMIC ELECTRIC COMPANY Prepared by Impell Corporation 2345 Waukegan Road Bannockburn, IL 60015 Impell Report No. 09-0570-0055 Revision 0 September, 1986 I
I 8610090258 DR 861001 ADOCK 05000029 PDR-
IMPELL CORPORATION REPORT APPROVAL COVER SHEET Client: Yankee Atomic Electric Company I Project: Yankee Nuclear Power Station (Rowe) Job Number: 0570-042 Report
Title:
Leak-Before-Break Evaluation of the 5" Bypass Lines Report Number: C9-0570-0055 Revision: 0 The work described in this Report was performed in acccordance with the Impell Quality Assurance Program. The signatures below verify the accuracy of this Report and its compliance with applicable quality assurance requirements.
Prepared By: b %M Scott Kulat, Senior Engineer Date: 9-/9-76 v 'm Date: if /f h6 Robert Emerson, Supervising Engineer Reviewed By: //y /~// _ /t [() L/ #
Date: k/' [6 Wal ter R. Bak, Manager '
A lied Mechanics Section Approved By: //I<<JqN na J) /l, M6 Date: 9//i[%
John J. Gav%la / h" /
Project Manager "
I REVISION RECORD Rev. No. Pre _ pared Reviewed_ _ ___ _ _ _
Approved Revision I
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I Page 2 of 38 I
Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 3 of 38 I TABLE OF CONTENTS Page EXECUTIVE
SUMMARY
4 1.0 INTRbDUCTION 5
2.0 DESCRIPTION
OF PIPING SYSTEM 6 3.0 COMPUTER PROGRNiS 7 4.0 METHODOLOGY AND NUMERICAL RESULTS 9
5.0 CONCLUSION
S 18
6.0 REFERENCES
20 FIGURES 22 TABLES 25 I APPENDIX A: DESCRIPTIONS OF COMPUTER PROGRNis 29 I APPENDIX B:
SUMMARY
OF PIPING STRESS ANALYSIS 33 I
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Yankee Atomic Electric Company Report No.: 09-0570-0055 l Revision: 0 l Page 4 of 38 EXECUTIVE
SUMMARY
A leak-before-break analysis of the 5-inch-diameter main coolant bypass piping at Yankee Nuclear Power Station has been performed in accordance with SEP guidance, using an alternative to the postulated crack size. The results of the analysis are as follows:
No intermediate points exceed the break-location stress criteria. The critical break locations are at the piping terminal ends. Specifically, the hot leg terminal end point is the bounding case for both Level A and Level D analyses.
- Using conservative analysis assumptions, the critical detectable crack un' der normal operating loads is a circumferential crack 3.00 inches long. The leak rate from this crack is 1 gpm.
Under normal operating plus seismic loads, the 3.00 inch circumferential crack is shown to be acceptable by comparisons of J I to J Ig and the applied moment to the limit moment. Local and global stability is demonstrated under I seismic conditions.
The tearing modulus stability analysis demonstrates that unstable crack growth will not occur in the base metal . Additional analyses show that a sufficient margin of safety exists between the extreme load ccnditions and those conditions which would cause unstable crack growth at a weld.
Using a conservative analysis approach, a 3.00 inch circumferential through-wall crack is shown to have negligible crack growth over 40 years, including a seismic event in the last year.
Based upon a review of documented incidents of cracked piping in PWRs, it has been confirmed that cracks grow through-wall rather than along the surface of the pipe.
As part of the alternate safety assessment for pipe rupture, the NRC requires the application of ASME Section XI inservice inspection (ISI) guidelines for Class I pipe, regardless of actual piping class. The main coolant bypass
- lines are currently inspected volumetrically in accordance with Class I guidelines, so no further ISI requirements need to be imposed.
In conclusion, an alternative to the SEP requirement that flaw sizes calculated under normal operating conditions be doubled before evaluating the crack under seismic loads has been justified. This alternative is required due to the low normal operating stresses which simultar:eously makes crack growth unlikely and crack detection through leakage more difficult. The alternative is justified by a !
crack growth evaluation that demonstrates that a large postulated crack will not I grow appreciably under the Main Coolant System (MCS) design conditions. Given this alternative, the 5-inch-diameter bypass lines are shown to satisfy the SEP I
requirements of a leak-before-break evaluation.
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Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 l Page 5 of 38
1.0 INTRODUCTION
Systematic Evaluation Program (SEP) Topic III-5. A requires that the effects of high energy line breaks on structures, systems, and components inside containment be addressed. A leak-before-break (LBB) evaluation has been performed to address this issue for the 5-inch-diameter main coolant bypass piping at the Yankee Nuclear Power Station (YNPS), at Rowe. The evaluation follows the approach described in the NRC document, " Guidance for Resolution of High Energy Pipe Break Locations where Remedial Modifications are Impractical" (Reference 1). The LBB approach is based on demonstrating that I the bypass piping would develop stable, detectable cracks rather than suddenly breaking in a double-ended guillotine fashion.
In accordance with the NRC guidance, the methodology for the leak-before-break I evaluation consists of the following:
Detectability Evaluation The size of crack which would be detectable under normal operating conditions (Level A) was determined for both circumferential and longitudinal cracks at critical break locations.
Integrity Evaluation The detectable crack as determined by the Detectability Evaluation was subjected to Level D loads (normal operating plus seismic conditions). A fracture mechanics evaluation was performed to determine that substantial crack growth does not occur for normal plus seismic loads. In addition, the load carrying capacity of the cracked section was evaluated to ensure that general plastic instability of the piping does not occur.
Extreme Load Evaluation A 90 circumferential through-wall crack was evaluated under extreme I bending conditions which cause local fully plastic conditions at the cracked section. In this evaluation, a tearing modulus analysis was performed to determine if unstable crack growth could occur.
Crack Development The tendency of the piping to develop through-wall rather than long, surface cracks was evaluated based upon a review of prior PWR experience and analysis of pipe cracking.
The 5-inch-diameter bypass lines at YNPS have a conservative design that results in low normal operating stresses. Due to the low stresses in these lines, the amount of leakage from a crack twice the wall thickness in length (2t) is not great enough at all postulated break locations to be assured of I detection. Str;ct adherence to the SEP guidelines of detecting a 2t crack at all break locations may not be demonstrated. Using a conservative I
Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 I Page 6 of 38 methodology, a crack slightly larger than 4t would be required to produce a leakage of 1 gallon per minute (gpm).
This LBB evaluation provides an acceptable alternative to the SEP criteria I that crack sizes calculated under Level A loads be doubled before evaluation under Level D loads. The alternative is reasonable and justified because of the conservative design of the pipe and the fact that the crack evaluated I under Level D loads is larger than the 4t crack which is discussed in
" Guidance for Resolution of High Level Break Locations where Remedial Modifications are Impractical" (Reference 1).
To support this alternative, the following analysis was performed to demonstrate pipe integrity:
Crack Growth Evaluation The postulated through-wall circumferential crack which yields 1 gpm was shown not to grow appreciably over the next 40 years. The analysis was performed based on a very conservative definition of the expected transien ts.
These analyses are presented along with the leak-before-break evaluation in the " Methodology and Numerical Results", Section 4.0, of this report. The analyses are documented in Impell Calculation No. 0570-042-003 (Reference 2).
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2.0 DESCRIPTION
OF PIPING SYSTEM The Main Coolant System (MCS) consists of four similar loops, each of which contains a 5-inch-diameter bypass line that connects the cold leg to the hot leg. The MCS is designed in accordance with ANSI B31.1. A typical main I coolant loop and the corresponding bypass line are shown in Figure 2-1.
Relevant data on the bypass lines is shown below:
Material : A376 Type 304 seamless, butt-welded piping.
Schedule: 160 Thickness: 0.625 in I Outside Radius: 2.78 in Inside Radius: 2.16 in Operating Pressure: 2000 psig Operating Temperature: 510*F (Cold Leg)
I 550*F (Hot Leg) l Design transients for YNPS are shown in Table 2-1. These transients were i conservatively enveloped in the crack growth evaluation performed. l The lines were welded using a two phase welding procedure. The root pass was I completed on an E.B. Type 308 insert ring using a Tungsten Inert Gas (TIG) process while the filler passes were completed by Shielded Metal Arc Welding (SMAW) using E308-15 or E308-16 electrodes.
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Yan'Kee Atcriic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 7 of 38 Review of operating history data indicates that the bypass lines have never been subjected to a water hammer event and are not subject to an environment which produces stress-corrosion cracking. In addition, the lines are not subject to significant vibration which could lead to fatigue failure.
The bypass lines are subject to the inservice inspection requirements of ASME Section XI for Class I piping and have been inspected with no indications of pipe flaws found.
3.0 COMPUTER PROGR#iS Five computer programs, listed below, were used in the leak-before-break evaluation.
- 1) SUPERPIPE
- 2) CRACK I 3) 4)
5)
IMLEAK AKMIN TEAR Brief descriptions of the use of the programs in the leak-before-break analysis are given below. Additional information on the computer programs is given in Appendix A.
3.1 SUPERPIPE SUPERPIPE performs piping analyses under various design conditions. For the leak-before-break analysis, SUPERPIPE has two separate applications.
First, SUPERPIPE is used to determine internal forces and moments of the piping under both Level A and Level D conditions. These loads are used as input when performing the leakage detectability and fracture evaluations as described in Sections 4.1, 4.2, and 4.3 of this report.
Second, SUPERPIPE is used to determine the elastic compliance of the piping system, assuming a fully-plastic section at the postulated crack.
This information is used in " Extreme Load Stability Analysis" described in Section 4.4 of this report.
In the SUPERPIPE analysis, the loads are combined using the same method described in Section 5.4 of NUREG-1061 (Reference 3). For normal loads (FN) and moments (Mn ), and seismic loads (Fs) and moments (Ms),
I the following combinations were used:
=
F lFNj + lFsj M = (M 2 , gj , g 2) U2 lI l
Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 I Page 8 of 38 where:
M1 = lM N1SI M2= lM 1Nl + N Sl I 2NI+ 2 M3 = lM3Nl + lM3SI 3.2 CRACK The CRACK program performs fracture mechanics calculations in accordance with work done by H. Tada and P. Paris (Reference 4). By using linear elastic fracture mechanics combined with a plastic zone correction factor, CRACK determines the crack opening area and stress intensity factor for a crack under Level A and Level D loads as described in Sections 4.2 and 4.3.
3.3 IMLEAK The IMLEAK program calculates the leak rate of pressurized fluids through narrow cracks. IMLEAK is based on References 5 and 6 and has been validated against experimental measurements of leakage through simulated and actual cracks in piping. Using the crack opening trea as determined by CRACK, IMLEAK determines the leakage under Level A conditions as described in Section 4.2.
3.4 AKMIN The AKMIN program uses a flexibility matrix to determine the minimum
'E5 stiffness in a piping line in accordance with work done by Macek, Grubb and Morton (Reference 7). Using moment-deflection relations determined by SUPERPIPE, AKMIN determines the minimum stiffness for the bypass lines. This information is used as input for the tearing modulu:
analysis as described in Section 4.4.
3.5 TEAR I The TEAR program uses the maximum compliance (reciprocal of the minimum stiffness) and piping material properties to compute values for the I tearing modulus, T. As described in Section 4.4, values of T are computed for chosen J-integral values in order to determine an applied J-T curve. This is used in the stability analysis of the pipe under ,
extreme loads, j I ~
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l Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 9 of 38 4.0 METHODOLOGY AND NUMERICAL RESULTS The leak-before-break analysis was separated into five major steps as follows:
- 1) Break Location Determination
- 2) Level A Leak Rate Calculation
- 3) Level D Integrity Evaluation
- 4) Extreme Load Evaluation
- 5) . Crack Growth Calculation The methodology of each of these steps and the computed results are presented in the following Sections. The final Section discusses the tendency for cracks to grow through-wall . A summary of the piping stress analysis is provided in Appendix B.
4.1 Break Location Determination The first step in the leak-before-break evaluation was to determine potential break locations on the bypass lines. Direction was taken from NRC Docket No. 50-255 (Reference 8). Using the Mechanistic Approach for Category I, ASME Class 2 and 3 piping, break locations are postulated at terminal ends and intermediate locations where the stresses for limiting normal and upset conditions as calculated by the sum of equations (9) and (10) of Section NC-3652 of the ASME Code (Reference 9) exceed 0.8 (1.2 S h + SA) where:
Sh = Material allowable stress at maximum temperature SA = Allowable range for expansion stress
= f(1.25 SC + 0.25 Sh )
SC = Material allowable stress at 70 F f = Stress range reduction factor I (NOTE: Equations (9) and (10) of ASME Section III, NC-3652,1977 1
Edition, Summer 1979 Addenda, are equivalent to equations 12a and 13a, I respectively, of ANSI B31.1,1977 Edition, Reference 10).
Based on the analysis performed using SUPERPIPE, there are no intermediate points for which the equation (9) plus (10) stresses exceed 0.8 (1.2 Sh + SA I-A review of the SUPERPIPE analysis revealed that the Level D loads at the ,
terminal ends envelope those at all intermediate points on the four !
bypass lines. Therefore, for the leak-before-break analysis the terminal ends of the bypass lines provide bounding results.
Once the terminal end locations were defined as enveloping, a l determination of the single most critical location followed a two-step l evaluation. First, the leakage through a crack under Level A conditions was considered. Under these conditions the lowest leakage rate for a given crack size occurred at the locations which were subject to the lowest loads. Minimum loading represented the bounding condition since a
Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 10 of 38 larger crack can exist at this location before a leak of 1 gpm occurs.
Using minimum loading criteria, SUPERPIPE shows that the bounding break location under Level A conditions occurs at the bypass line terminal end at the hot leg on Loop 3.
In the second step, break locations were evaluated for the bypass lines under Level D conditions. For these conditions, the critical location is subjected to the highest loads. SUPERPIPE shows again that the location subject to the highest Level D loads was at the bypass line terminal end at the hot leg of Loop 3.
4.2 Level A Leak Rate Calculation This evaluation determines the size of a crack which would leak 1 gpm under normal operating (Level A) conditions. Using pipe dimensions, material properties, and operating loads as determined by SUPERPIPE, the CRACK computer program calculated the crack opening area and stress intensity factor for a crack with a specified length and orientation.
The crack opening area was then input into IMLEAK along with the operating conditions to determine the amount of leakage that would occur through the crack.
As described in Section 4.1, the terminal end at the hot leg of Loop 3 was the critical location under Level A loads. The input and the results of the CRACK and IMLEAK calculations are given in Table 4-1. As shown in Table 4-1, a longitudinal crack under Level A conditions will leak 1 gpm upon reaching a length of 1.85 inches. However, a circumferential crack under Level A condition will not produce 1 gpm leakage until reaching a length of 3.00 inches. Since the largest crack represents the bounding condition, the most critical crack under Level A condition is oriented circumferentially with a length of 3.00 inches.
4.3 Level D Integrity Evaluation I The Level D integrity evaluation consists of two parts. First, fracture mechanics calculations were performed to demonstrate that the crack determined in Section 4.2 of this report is stable under Level D I conditions (normal operating plus seismic loads). Sec0nd, a net section evaluation was performed to demonstrate global stability of the piping.
Details of these two analyses are included in this section.
As discussed in Section 1.0 of this report, a margin of two on crack length was not imposed in this analysis. Both analyses were based on a 3.00 inch long circumferential crack.
4.3.1 Fracture Mechanics Evaluation This portion of the analysis shows that the 3.00 inch I circumferential crack would not grow in an unstable manner under Level D conditions. This was shown by determining the applied J I integral and comparing this to the JIC for Type 304 stainless steel .
Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 11 of 38 I As discussed in Section 3.2 of this report, the CRACK computer program calculates the stress intensity factor, KI , for cracks under specified loads. In this case, Level D loads at the terminal end at the hot leg of Loop 3 were used as input. Other than the increased loads, the input for CRACK in this evaluation differed from the input used under Level A loads in the following way.
'I The flow stress was used instead of yield stress to calculate the plastic zone correction factor arid ultimately the effective crack u
l I size (In Reference 3, the flow stress is defined as the average of the ASME code specified yield and ultimate strengths at the appropriate temperature). For Level A loads, the plastic zones were small, and the yield stress was appropriate for computing the effective crack size. However, for the greater Level D loads, 3 using the yield stress in effective crack size calculations is M overly conservative. For this case, the use of flow stress represents a more realistic situation because it accounts for the strain hardening properties of stainless steel.
I The input and results of the CRACK program for a 3.00 inch circumferential crack under Level D loads are shown in Table 4-2.
The maximum calculated stress intensity factor under Level D loads 3
M was 101,347 psi / inch. Under conditions of plane stress, the applied stress intensity may be converted to the applied J-integral using the following equation:
U
- I E
I (Where E is the modulus of elasticity for the material .) For the applied stress intensity of 101,347 psi / inch and the elastic modulus of stainless steel at 550 F, the applied J integral is 402.0 psi-inch. From " Fracture Behavior of Stainless Steels" as presented in ASTM STP 668 (Reference 11), the critical value of J for stainless steel at 600*F is 2569 psi-inch. From the lower bound weld data presented in Appendix F of NUREG 1061, Reference 13, the critical value for J for the bypass line welds in 960 psi-in. Since the applied J valc? of 402.0 psi-inch is far less than the critical values of J for both the piping base metal and welds, the 3.00 inch circumferential crack will remain stable under Level D loads.
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3 Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 12 of 38 4.3.2 Net Section Evaluation To evaluate global stability of the piping, the limit moment that the uncracked portion of the pipe could withstand was calculated I and compared to the calculated applied moment. Acceptability is demonstrated if the ratio of the limit moment to the applied moment remains greater than one.
The equation for the limit moment from Reference 3 is:
Mr = 4afR2 t (cos y - 1/2 sin 6 )
where: R = mean pipe radius t = pipe wall thickness _
6 = half crack angle a [-
1/2 (yield + ultimate strength) L 7f == 6/2 + Axial Load /(4 t)R af Mf was determined to be 417,475 i n-lbs. At the terminal end of the bypass line at the hot leg of Loop 3, the maximum applied moment is 227,327 in-lbs. , thus the Mf /M ratio is 1.84. Since Mf /M is greater than one, the uncracked portion of the pipe would be able to carry the calculated Level D loads.
I As presented in Sections 4.3.1 and 4.3.2, the 3.00 inch circumferential crack would have local stability under Level D conditions and the crack would not compromise global stability of the piping. This satisfies the requirements of the Level D integrity evaluation.
4.4 Extreme Load Calculation -
This portion of the evaluation demonstrates that the pipe would be stable under the extreme conditions of a 90' through-wall crack subject to fully plastic bending at the cross section. The methodology of the analysis is as follows:
- 1) Using rotation $ load-deflection data at a postulated flexible coupling in SUPERPIPE, AXMIN calculated the minimum stiffness of the bypass loops.
- 2) Maximum compliance was calculated using the equation shown in Section 4.4.2 of this report.
- 3) Using the maximum compliance and standard piping information as input, TEAR calculated the applied tearing modulus, Tapp, that corresponds to applied J-integral, Japp.
- 4) Japp vs. Tapp was plotted on a graph along with J vs. T curves which conservatively bound the material fracture resistance.
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Yankee Atomic Electric Company f Report No.: 09-0570-0055 i
Revision: 0 Page 13 of 38
- 5) A comparison was made between the Japp vs. Tapp and Jmat vs. Tmat curves. The comparison did not demonstrate unconditionally that unstable crack growth could not occur at a weld, so additional analyses were completed to demonstrate that unstable crack growth would not occur.
The detailed approach and results from each step are described in Sections 4.4.1 to 4.4.5. (
4.4.1 Minimum Stiffness Calculation As described in Section 3.1.2 of this report, the SUPERPIPE program was used to determine the flexibility of the piping system with a fully plastic section at the location of the crack. For this analysis, the terminal end nozzles at the main coolant piping were modeled as anchors. Using the rotation information, AKMIN {
calculated the stiffness matrix corresponding to the bypass lines and the minimum stiffness for all four bypass lines. The resultant minimum stiffness was calculated to be 0.1704E7 in-lbs/ rad.
4.4.2 Maximum Compliance Calculation E
The minimum stiffness of 0.1704E7 in-lbs/ rad was used in the following equation:
Cmax = 1 Kmin where: Cmax = Maximum compliance Kmin = Minimum stiffness The maximum compliance of the bypass lines was calculated as 0.58685E-6 rad /in-lbs.
4.4.3 Tearing Modulus Calculation I Using the maximum compliance and piping material properties as input, TEAR calculated the applied tearing modulus for given values of J. For this case, J values of 0.0, 100.0, 1000.0 and 10,000.0 were chosen so that the entire range of possible J values would be covered. Input and results from TEAR for a 90 through-wall crack under extreme load conditions are given in Table 4-3.
4.4.4 J vs T Plot I The Japp vs Tapp curve is graphically depicted along with material J vs. T curves in Figure 4-1. J vs. T material curves were obtained from the following sources.
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Yankea Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 14 of 38 I (a) The J vs. T curve for A 376 Type 304 stainless steel was obtained from EPRI Report NP-2261 (Reference 12).
(b) The J vs T curve for welds was adopted from NUREG-1061 Volume 1 (Reference 13). The J vs. T curve for Tungsten Inert Gas Welds (TIG) envelopes all the identified test data for pipes, forgings and their welds produced by the GTAW or Shielded Metal Arc (SMAW) processes.
For the 5-inch-diameter bypass lines at YNPS, welds were made in I accordance with Welding Procedure Specification No. 9699 (Refere e 14) . This welding procedure specifies that welds be deposited using a combined process of GTAW at the root and SMAW with Type 308 welding material for the remairder of the weld.
Since GTAW and SMAW are the only processes u;ed, and the J vs. T curve for GTAW envelopes both processes, the GTAW based J vs. T curve provides lower bound material fracture resistance for analysis of the welds.
4.4.5 Justification for Acceptance Unconditional stability is demonstrated in an applied tearing modulus analysis when the Japp-Tapp curve always lies to the left of the Jmat-Tmat curve. This condition is true for the Japp vs.
I Tapp curve when compared to the dat vs. Tmat curve for the base material . For the base material, the applied tearing modulus never reaches the critical tearing modulus. This demonstrates that the onset of unstable crack growth will not occur in the base metal .
The Japp vs. Tapp curve intersects the Jmat vs. Tmat curve for the I wel d . This indicates that there is a possibility for unstable crack growth to occur at the weld for high values of applied J-integral . Three additional analyses were performed to quantify the factors of safety which demonstrate margin against onset of unstable crack prowth.
I (1) Both NUREG-1061 (Reference 3) and EPRI Report NP-2261 (Reference 12) state that for a fully-plastic cracked section, the J integral can be estimated by:
J= C f RF $
where: of = flow stress = 41,150 psi at 550'F R= Mean radius = 2.469 in.
F= Sin (6/2) + Cos 6 = 1.09 6= Half Crack Angle = 45' d = Kink Angle
Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 15 of 38 I Both NiiREG-1061 and EPRI Report No. 2261 have assumed a kink angle of one-degree in their evaluation for pipe under limit moment conditions. Based upon an assumed kink angle of one degree, the value of Japp for the fully plastic cracked I section was determined to be 1932.8 psi-in. (The kink angle is the angular discontinuity due to localized deformation at a fully plastic cracked section.)
In Figure 4-1, the intersection point between the Japp vs.
Tapp line and Jmat vs. Tmat curve corresponds to the initiation of unstable crack growth. By inspection, the critical value of Jmat at the intersection point is 3200 ps i -i n .
The margin between plastic conditions and unstable conditions is represented by the ratio of Jmat to Japp. For the values determined in this section of the report, the Jmat/Japp ratio I is 1.66. Therefore, a 66% margin in J values exists between conditions corresponding to a kink angle of one-degree at the cracked section, and conditions which would cause the crack to -
grow in an unstable manner.
(b) As discussed in Section 4.3.2 of this report, the ratio of the limit moment over the applied Level D moment is 1. 4.
I Therefore, moment loads would need to increase 84 over Level D loads before reaching limit moment conditions and full plastic bending.
(c) For additional confirmation that a margin exists to preclude unstable crack growth conditions, an alternate calculation based on energy considerations was performed. This I calculation shows that the energy needed to cause unstable fracture is greater than the energy that could be supplied to the piping by Level D earthquake conditions. The energy required to cause fracture was conservatively estimated as one-half the limit moment times the critical rotation found in the J-T comparison. The kinetic energy of the piping I resulting from the seismic loads was very conservatively estimated assuming the peak spectral velocity occurred at all points in the p' ping system. The calculation shows that the energy required to cause unstable fracture is 32% greater than the energy that the earthquake would provide.
The tearing modulus analysis demonstrates that unstable crack I growth will not occur in the base metal . Also, additional analyses have shown that justifiable margins of safety exist between maximum applied conditions and those conditions which would cause unstable crack growth at the weld.
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Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 16 of 38 4.5 Crack Growth Calculation This calculation demonstrates that a 3.00 inch circumferential through-wall crack could not grow appreciably. The analysis followed a crack growth methodology similar to that presented in Appendix C of the ASME Code,Section XI (Reference 15). The analysis used a stress-intensity factor for a through-wall flaw based on Reference 4 and crack growth data for stainless steel from Reference 16.
Design transients for YNPS are as shown in Table 2-1. For the crack growth evaluation, transients were divided into two categories. The first category includes all Level A and B service transients except for steady state fluctuations. The steady state fluctuations are the second category.
I For the first category, each transient corresponds to an event which imposes temperature and pressure cycles upon the main coolant system. Of these events , the Plant Heatup and Cooldown cycle represents the largest change in conditions. The crack growth analysis was performed under the extremely conservative assumption that every transient in the first I category subjects the bypass lines to the conditions of the Heatup/
Cooldown cycle. In addition, the analysis considered the location on the four bypass lines which had the highest Level A loads. Even under these conservative conditions, the results of the crack growth analysis for all I level A and B transients except steady state fluctuation: was that the 3.00 inch circumferential crack would only grow 0.0020 inch per year or 0.080 inch over the next 40 years.
For the second category, the effects of steady state fluctuations were evaluated. For the steady state conditions shown in Table 2-1, fluctuations in pressure and temperature were translated into effective load fluctuations. These load transients were then used as input for the crack growth analysis. Results of this analysis are that the 3.00 inch circumferential crack would grow 0.0037 inch under steady state fluctuations over the next 40 years.
In order to conservatively evaluate the possible series of transients I causing crack growth, a seismic event was postulated to occur at the end of 40 years. In this situation, a crack would be subject to growth under seismic conditions in addition to growth occurring under normal operating conditions. Using the highest Level D loads as input, the results of the crack growth analysis show that tSe 3.00 inch circumferential crack would extend 0.037 inch during a seismic event.
I In accordance with ASME Section XI, crack growth calculations for each transient should be based on the final size of the crack determined under the previously evaluated transient. For this analysis, performing the I evaluation in this iterative manner is unwarranted because the amount of crack extension for each transient is essentially negligible. Instead, each portion of the crack growth analysis was conducted on a 3.00 inch I
Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: .O Page 17 of 38 crack, and the results were simply added. Considering the very small extensions which occur in each portion of the analysis, the deviation from ASME Section XI methods produce insignificant changes in results.
The summation of crack extensions under all Level A plus. B transients and a seismic event produce a total crack extension of only 0.121 inch.
Therefore, the 3.00 inch circumferential through-wall crack was shown not to grow appreciably over the next 40 years.
4.6 Subcritical Crack Development This section demonstrates that partial-through wall cracks are likely to break through the pipe wall and leak before they will progress around the pipe and cause a significant break. The likelihood of leak-before-break condition produced by a leaking crack is verified for the two conditions that are of major interest; narely, normat operation and large bending loads in excess of those postulated for seismic loading.
For normal operating conditions, there is a large amount of service experience which demonstrate that cracks progress radially through the I pipe wall and result in leak-before-break conditions. As indicated in References 17 and 18, incidents of pipe cracking have been documented at a number of PWR's in the United States. These references discuss pipe that is 4 inches or more in diameter, which includes the pipe size being considered in this evaluation.
The statistics show data with a wide range of crack sizes and piping I systers. The cracks result from various initiation and propagation mechanisms, such as intergranular stress corrosion cracking, thermal f ati ga:2, dynamic loads, and erosion / cavitation. In addition, these various type cracks are exposed to different combinations of stress states, i .e. bending and tension. For all the different conditions that actually occur in service, the cracking data indicate that the likelihood of a significant break is remote and that the dominant behavior for intermediate and large diameter piping is for the crack to grow radially through the wall to produce the leak-before-break condition.
Because accident loadings have a very low rate of occurrence, it is not I possible to use service experience to define crack growth. Instead, the previously performed analyses are used to demonstrate that the leak-before-break condition will be maintained for loads in excess of postulated large accident seismic loads. The study described in Reference 19 defined the ratio of the J-integral in the circumferential
-to-radial direction for a partial-through-wall crack in a pipe. The I analytic results ir.dicate that the value of J in the radial direction, Ja, is always greater than the value of J in the circumferential direction, Jp, for all cctbinations of a/t and circumferential distance around the pipe. This variation is shown in Figure 4-2, which is taken I from Reference 19. The variation of J4 with Ja shown in Figure 4-2
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Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 18 of 38 I demonstrates that there is a strong tendency for 1eak-before-break conditions to exist for loads in excess of large postulated accident seismic loads.
Based on service experience and analysis, it can be concluded that cracks grow through pipe walls and result in leak-before-break conditions for Yankee's operating conditions ranging from normal operation to large accident loads, and for a wide range of tensile and bending loading combina tions. This justifies the leak-before-break assumption of a through-wall crack used in this report.
5.0 CONCLUSION
S A leak-before-break analysis of the 5-inch-diameter bypass piping has been performed in accordance with SEP guidance, using an alternative to the postulated crack size. The results of the analysis are as follows:
5.1 Break Locations No intermediate points exceed the break-location stress criteria. The critical break locations are at the piping terminal ends. Specifically, the terminal end point at the hot leg of Loop 3 is t:1e bounding case for both Level A and Level D analyses.
I 5.2 Level A Leak Rate Calculation Using conservative analysis assumptions, the critical detectable crack under normal operating loads is a circumferential crack 3.00 inches long. The leak rate from this crack is 1 gpm. -
5.3 Level D . itegrity Evaluation Under normal operating plus seismic loads, the 3.00 inch circumferential I crack was shown to be acceptable by comparisons of J I to J IC and the applied moment to the limit moment. Local and global stability were demonstrated under seismic conditions.
5.4 Extreme Load Stability Analysis The tearing modulus stability analysis demonstrates that unstable crack growth would not occur in the base metal . Additional analyses have shown that a sufficient margin of safety exists between the extreme load conditions and those conditions which would cause unstable crack growth at a weld.
5.5 Crack Growth Evaluation Using a conservative analysis approach, a 3.00 inch -circumferential through-wall crack is shown to have negligible crack growth over 40 years, including a seismic event in the last year.
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Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 I Page 19 of 38 5.6 Subcritical Crack Development Based upon a review of documented incidents of cracked piping in PWRs, it I was confirmed that cracks grow through-wall rather than along the surface of the pipe.
5.7 Inservice Inspection As part of the alternate safety assessment for pipe rupture, the NRC requires the application of ASME Section XI inservice inspection (ISI)
I guidelines for Class I pipe, regardless of actual piping class. The main coolant bypass lines are currently inspected volumetrically in, accordance with Class I guidelines, so no further ISI requirements need to be imposed.
5.8 Summary I In summary, an alternative to the SEP requirement that flaw sizes calculated under normal operating conditions be doubled before evaluating the crack under seismic loads has been justified. This alternative is required due to the low normal operating stresses which simultaneously make crack growth unlikely and crack detection through leakage more di f ficul t. The alternative is justified by a crack growth evaluation that demonstrates that a large postulated crack will not grow appreciably I under the MCS design conditions. Given this exception, the 5-inch-diameter bypass lines are shown to satisfy the SEP requirement, of a leak-before-break evaluation.
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Yankee Atomic Electric Company Report Mo.: 09-0570-0055 Revision: 0 Page 20 of 38
6.0 REFERENCES
- 1. " Guidance for Resolution of High Energy Pipe Break Locations Where Remedial Modifications are Impractical" as presented in NRC Docket tb.
I 50-255 dated December 5,1981.
- 2. " Yankee Atomic Leak-Before-Break (LBB) Analysis of 5 inch Diameter Bypass Lines", Impell Calculation 0570-042-003, Rev. O.
- 3. " Evaluation of Potential for Pipe Breaks" as presented in the Report of I the U.S. Nuclear Regulatory Piping Review Committee, NUREG-1061, Volume 3, dated November,1984.
- 4. "Es timation of Stress Intensity Factors and Crack Opening Area of a Circumferential and a Longitudinal Through-Crack in a Pipe", by H. Tada a nd P. Pa ris.
I 5. " Calculation of Leak Rates Through Cracks in Pipes and Tubes." EPRI Report NP-339 5, December,1983.
- 6. " Study of Critical Two-Phase Flow Through Simulated Cracks," Batelle Laboratories, Interim Report - BCL-EPRI-80-1, November 25, 1980.
- 7. Pipe Crack Study Group, " Crack Tearing Stability Analysis Stiffnesses" by I R. W. Macek, R. L. Grubb, and D. K. Morton, as presented in EGaG Letter J AD-177-80, dated July 14, 1980.
- 8. " Evaluation of Effects of Pipe Break on Structures, Systems and Components Inside Containment. Topic III-5. A for the Palisades Nuclear Plant" as presented in NRC Docket No. 50-255 dated December 4,1981.
- 9. ASME Code Section III, Subsection NC,1977 Edition through and including Summer 1979 Addenda.
- 10. ANSI B31.1 " Power Piping",1977 Edition,
- 11. " Fracture Behavior of Stainless Steel" by W. H. Bamford and A. J. Bush as presented in " Elastic-Plastic Fracture", ASTM STP 668, dated 1979.
- 12. " Application of Tearing Modulus Stability Concepts to Nuclear Piping",
EPRI Report NP-2261 dated February 1982.
- 13. " Investigation and Evaluation of Stress Corrosion Cracking in Piping of Boiling Water Reactor Plants", as presented in the Report of the U.S.
I Nuclear Regulatory Commission Piping Review Committee, NUREG-1061, Volume 1, dated August,1984.
- 14. " Welding Procedure Specification for Main Coolant Loop-20 and 24 in.; 4 in. Pressurizer Surge Line; 5 in. Loop Equalizing Lines for Austenitic
Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 21 of 38 I Stainless Steel Piping, Castings and Forgings, Type 304 an 316, Shop and Field Welding." WPS No. 9699 dated July 15, 1958.
l I 15. ASME Code,Section XI,1983 Edition through and including Winter 1985 Addenda.
l 16. D.0. Harris, E.Y. Lim, D.D. Dedhia, " Probability of Pipe Fracture in the Primary Coolant Loop of a PWR Plant," NUREG/CR-2189, Vol 5, U.S. Nuclear {
Regulatory Commission, Washington D.C.
I 17. Pipe Crack Study Group, " Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors" NUREG-0691, Nuclear Regulatory Commission, dated September,1980.
- 18. Pipe Crack Study Group, " Investigation and Evaluation of Stress-Corrosion Cracking in Piping of Light Water Reactor Plants", NUREG-0531, Nuclear Regulatory Commission, dated February,1979.
- 19. "A Plastic Fracture Instability Analysis of Wall Break Through in a Circumferentially Cracked Pipe Subject to Bending Loads" presented by A.
Zahoor and M. F. Kanninen in Trans. ASME, Vol .103, July 1981, p.194-200.
- 20. H. Tada, P.C. Paris, and R. Gamble. June 1979. " Stability Analysis of Circumferential Cracks in Reactor Piping Systems." NUREG/CR-0838, U.S.
I Nuclear Regulatory Commission, Washington, D.C.
21 . Cygna Energy Services. " Seismic Reevaluation and Retrofit Criteria,"
Revision 2, August 16, 1982.
?2. "5" Main Coolant Loop Bypass Piping Analysis," Impell Calculation 0570-042-001, Rev.1.
- 23. " Compliance Analysis of MCL Bypass Piping," Impell Calculation 0570-042-002, Rev. O I
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Yankee Atomic Electric Company Report tb.: 09-0570-0055 Revision: 0 Page 22 of 38 I
I Y l 9' 0 ,'s 0{ c.
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I Figure 2-1 Typical Bypass Line I
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Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 23 of 38 l
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I 7 i i i I 6 -
TUNGSTEN INERT GAS WELD (LOWER B0UND FOR PIPES, PLATES, A 376
, TYPE 304 BASE -
FORGINGS, AND GTAW AND MATERIAL SMAW WELDS MADE WITH TYPE 308 WELDING MATERIAL) 5 - -
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5 14 I
=
5 E3 I }
(UNSTABLE) 2 -
(STABLE) 3 ~
Japp-Tapp I O O
50 100 150 200 250 300 I TEARING MODULUS T I
I Figure 4-1 Applied and Material J-T Curves I
I
Vankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 24 of 38 lI I
I 1.0 ,,
N
'a %
s a X=af [ G E ~
O'8 -
+-
'l ss I I O.6 -
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g
\
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O l 0 0.2 0.4 0.6 0.8 1.0 I b I
I Figure 4-2 Stability of Part-Through Crack under Fully-Plastic Bending I i I
I Yankee Atomic Electric Company Report No.: 09-0570-0055 I Revision: 0 Page 25 of 38 YANKEL-ROWE SERVICE LEVEL A TRANSIENTS EVENTS /YR.
MCP Startup and Shutdown
- RCS Venting 20 Pump Restart Conditions 5 Hot Plant Conditions- 40 Plant Heatup and Cooldown at 50 DEG F/ Hour 5 at 100 DEG F/ Hour 1/4 Unit Loading and Unloading at 25% Hour 0 to 50% Power 10 50% to 100% Power 50 Step Load Decrease 1 Steady State Fluctuations Temperature (+/-0.5 Deg F ) 75000 Pressure (+/-25 PSI) 10000 Primary Side Leakage Tests to 2200 PSl* 10 FW Cycling 100 Core Life Extension (Covered by Plant Unloading and Cooldown Above.)
- RCS Pressure Changes Only I
SERVICE LEVEL B TRAH51EN15 Loss of Load 1/4 Loss of Power 1/2 Reactor 1 rip from full Power Normal RX 1 rip I
6 RX Trip With Cooldown 1 RX Trip With Loss of MCP 2 Excessive FW Flow 3/4 Inadvertent Steamline NRY Closure 1 Table 2-1 YNPS MCS Transients 1
M M M .
TABLE 4-1: LEVEL A LEAK RATE RESULTS i
PRESSURE M0 MENT CRACK CRACK CRACK WIDTH LEAK RATE TEMP. CRACKAgEA
- F PSI LOAD, in-lbs TYPE LENGTH, in in in GPM 2000 12,179 LONG. 1.85 0.342509 E-2 0.18514 E-2 1.112 51 0 550 2000 12,179 LONG. 1.85 0.354470 E-2 0.19161 E-2 1.135 2000 12,179 CIRC. 2.90 0.370501 E-2 0.12776 E-2 0.948 51 0 550 2000 12,179 CIRC. 2.90 0.375643 E-2 0.12953 E-2 0.957 2000 12,179 CIRC. 3.00 0.402968 E-2 0.13432 E-2 1.066 51 0 550 2000 12,179 CIRC. 3.00 0.408620 E-2 0.13621 E-2 1.074 EERC
% 5.E a Es h S Note: Only the axial load due to inte :al pressure is included. o.9EFg
.. a
@O Pn 8 87 E8 E
4!
W M M M M g e TABLE 4-2: LEVEL D FRACTURE EVALUATION RESULTS TEMP. PRESSURE AXIAL MOMENT CRACK CRACK CRACK AREA Ky
- F PSI LOAD, lbs. LOAD, in-lbs TYPE LENGTH, in in2 psi ftTT 51 0 2000 2662 227,327 CIRC. 3.00 0.0568121 100,770 550 2000 2662 227,327 CIRC. 3.00 0.0581655 101,347 l
22:
'a 5.B a ME \
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@0 Pn 8 2M 6g 8
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M M M M M TABLE 4-3: J AND T RESULTS FOR EXTREME LOAD CONDITIONS l
._--- - ===-------- INPUT PARAMETERS --------------------- ----------------------------------- RESUI.TS ----------------------
TEMP. I MAX IMUM HALF ANGLE E FL0id T at at T at T at SLOPE OF psi J a0p0.0 T,!p 100.0 J'Op1000.0 J 0p 10,000.0
'F in4 COW LIANCE OF CRACK STRESS, psi J J-T LINE 51 0 30.02 0.58685 E-6 45* 27.75 E6 41.39 E3 138.92 138.78 137.55 125.22 -0.13 70E-2 550 30.02 0.58685 E-6 45' 25.55 E6 41.15 E3 137.84 137.70 136.47 124.09 -0.1375E-2 OO e<m, O =*= O X CD ORER
-h
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Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 29 of 38 APPENDIX A DESCRIPTION OF COMPUTER PROGRAMS 1
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Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 30 of 38 A.1 SUPERPIPE SUPERPIPE is a comprehensive computer code for the structural analysis and design checking of piping systems. In addition to a wide array of static analysis options, SUPERPIPE performs dynamic response spectral analyses, force time history and acceleration time history analysis.
For response spectral analyses, a ' missing mass' correction is available to account for dynamic effects represented by higher modes not computed in the eigensolution. In addition to the traditional in-phase, single-level analysis I method using enveloped spectra, SUPERPIPE uses a multilevel response spectral analysis technique to analyze piping subjected to dynamic excitations which vary significantly between different anchor / support locations. For appropriate applications, the multilevel method can provide reduced analysis conservati sm.
SUPERPIPE can also perform the dynamic, multilevel excitation option as a time history modal superposition analysis, with seismic loading represented by an acceleration time history record. Force time history analysis options are available with SUPERPIPE using the modal superposition methodology as well as the direct integration method.
A comprehensive series of design checking options are available for Class 1 and Class 2 stress checking, Class 2 break locations, and support load summaries.
SUPERPIPE was benchmarked by comparison with results published by the NRC in I NUREG/CR-1677 for sev:n sample problems. The comparison was performed in accordance with the < of request for additional verification of computer codes used for analysis 01 nuclear piping systems. The verification specifically i
i addressed the response spectrum method of dynamic analysis commonly used in seismic qualification of nuclear piping.
The program has also been thoroughly tested and verified for a comprehensive I set of sample problems, including extensive comparison with several publicly available programs and ASME benchmark problems. All verification analyses have been documented in accordance with established Impell Quality Assurance procedures. SUPERPIPE is widely used and has been audited by many clients.
A.2 CRACK I CRACK is a computer code that performs fracture mechanics calculations for leak-before-break analyses. The code calculates the stress intensity factor and the opening area of cracks in piping.
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Yanke Atomic Electric Company ,
Report No.: 09-0570-0055 )
Revision: 0 '
l Page 31 of 38 The cracks considered are through-wall and may be oriented axially or circumferentially. The input required is the pipe section geometry, crack size, and loads. The loads used for circumferential cracks are axial load, internal pressure, and bending moment. For longitudinal cracks, only internal I pressure is required.
Linear elastic fracture mechanics formulations are used. Since material near I the crack tip yields, a plastic-zone correction to the effective crack length is included. An iterative calculation scheme is used to ensure a stable sol ution . The code is based on the methodology of Tada and Paris, Reference'4, and has been verified in accordance with Impell Quality Assurance procedures.
The stress intensity factor calculated by CRACK is used to determine local stability of the flaw piping. The opening area of the crack is used, along with other data on the piping system, by the program IMLEAK to calculate fluid flow through the crack.
A.3 IMLEAK IMLEAK is a computer code developed to evaluate the leak rate of pressurized fluids through narrow cracks. The code is used in leak-before-break analyses to estimate leak rates from postulated cracks to determine whether the crack is detectable. The program is based on References 5 and 6 and has been verified against experimental data. The program has been verified in accordance with Impell Quality Assurance procedures.
The analytical model in IMLEAK is a modified version of Henry's non-equilibrium two-phase flow model . The model accounts for non-equilibrium effects in the flow due to flashing within the flow path. The model also includes pressure drops due to entrance losses, friction, and fluid acceleration.
The model handles complex crack geometry, including turns in the flow path, variable flow area, and crack surface roughness. The program also contains its own steam properties subroutines.
A.4 AKMIN l AKMIN is a computer program used to determine the minimum stiffness in a ,
piping system. !
The analytical model for AKMIN is based upon work performed by R. W. Macek, R.
L. Grubb, and D. K. Morton for the NRC's Pipe Crack Study Group (Reference 7).
In the program, rotation at a postulated flexible coupling in the piping system is used to develop a 2 x 2 flexibility matrix. The flexibility matrix is then used to determine the stiffness matrix. By determining the smallest eigenvalue of the stiffness matrix the computer program calculates the minimum i stiffness for the piping system. The program has been verified in accordance i with Impell Quality Assurance procedures.
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Yankee Atomic Electric Company Report tb.: 09-0570-0055 Revision: 0 I Page 32 of 38 A.5 TEAR TEAR is a computer program developed to determine the value of the tearing modulus of pipe for an applied J-integral . TEAR is based on the methodology I of NUREG/CR-0838, Reference 20, and has been verified in accordance with Impell Quality Assurance procedures.
Input for the TEAR program consists of pipe dimensions, material properties, and applied J-integral values and the maximum piping compliance. Maximum piping compliance is determined as the reciprocal of the minimum piping stiffness as determined by the AKMIN program. For each specified J-integral I value, TEAR calculates the corresponding tearing modulus, T.
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Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 Page 33 of 38 I
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I APPENDIX B Slt1 MARY OF PIPING STRESS ANALYSIS
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Yankee Atomic Electric Company Report Pb.: 09-0570-0055 .
Revision: 0 l I Page 34 of 38 I A piping stress analysis of the 5" bypass lines was performed to determine the loads and stresses in the piping during normal operating conditions and faulted conditions. The loadings considered were gravity, pressure, thermal expansion, seismic, and anchor motion loads and displacements. The piping evaluation was I performed using the Impell computer code, SUPERPIPE, which is described in Appendix A.
I The Anplified Respanse Spectra (ARS) for the main coolant loop under the Yankee Composite Spectra (YCS) was used for the seismic analysis. Figures B-1 through B-3 show the ARS at the bypass vent nozzle.
The results of the piping analysis are summarized in Table B-1. The equations are taken from the piping reevaluation criteria, Reference 21. The equation numbers correspond to those of the criteria document. The piping analyses performed by Impell are documented in Calculations 0570-042-001, Revision 1, and 0570-042-002, Revision 0, which are References 22 and 23, respectively.
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i Yar.kee Atomic Electric Company Report No.: 09-0570-0055 l Revision: 0 Page 35 of 38 I 1.00 I i.60
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E-79 am. n X DIREETICN -T ds b}" CYCNA f7.[5,/pf ... n. 0 Figure B ARS at Bypass Vent Nozzle, X-direction I
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I Yankee Atomic Electric Company Report No.: 09-0570-0055 l
i I Revision: 0 Page 36 of 38
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owan.. y I 7s h,n, (See p. E-26)
- s. ,n i ... o. . ., 2,3,51 ""*"'"5 15 1 NS, D*, & Vertical sii. u ... . o. ui u ai . . . p ge' O.1384 3-I tm.
ARS m MMOR MDCHANICAL ICUIftCg MO MAIN CODI.RC I.00P UNDER YCS, h'*safra e..u,.ej,8T M CYm A P) .
sei, 12/16/85 ht' 12/19/85 me .
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I l Figure B ARS at Bypass Vent Nozzle, Y-direction I
Yankee Atomic Electric Company Report No.: 09-0570-0055 Revision: 0 I Page 37 of 38 l.83
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g , ,_ '"'" W y.c._ , cYc m e.u/54r n ... a. 0 Figure B l'tS at Bypass Vent Nozzle, Z-direction
Yankee Atomic Electric Company l Report No.: 09-0570-0055 Revision: 0 Page 38 of 38 I
Bypass Line On:
Stress Equation Loop 1 Loop 2 Loop 3 Loop 4 Allowable Maximum Ratio (from Ref. 21) (ksi) (ksi) (ksi) (ksi) (ksi) (Calculated / Allowable)
- 8. 3.1 - A 6.1 7.5 6.5 5.9 15.9 0.47
- 8. 3.1 -B 22.5 21 .0 24.6 22.4 28.6 0.86
- 8. 3.1 -C 4.4 4.3 3.2 4.4 27.5 0.16 I
Table B Sumary of Maximum Bypass Piping Stresses
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Domestic Offices 350 Lennon Lane Walnut Creek, CA 94598 2345 Waukegan Road Bannockburn, IL 60015 -
225 Broad Hollow Road ~
Melville. NY 11747 800 South Street The Watermill Center ~
Waltham, MA 02154
~
333 Research Court Technology Park / Atlanta -
Norcross, GA 30092 International Offices Impell Corporation
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A Division of Combustion Engineering Ltd.
Genesis Centre Garrett Field Birchwood, Warrington ~
WA3 78H United Kingdom -
Impell-France S A.R.L.
10, Rue du Colise'e c 75008 Paris France Affiliate Company Impell Pacific 920 Southwest Sixth Avenue Portland, OR 97204
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Berkeley, CA 94704 Progest S.p.A. (Fiat TTG)
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