ML20141M328
| ML20141M328 | |
| Person / Time | |
|---|---|
| Site: | Comanche Peak  |
| Issue date: | 05/18/1992 |
| From: | William Cahill, Woodlan D TEXAS UTILITIES ELECTRIC CO. (TU ELECTRIC) |
| To: | NRC OFFICE OF INFORMATION RESOURCES MANAGEMENT (IRM) |
| References | |
| RTR-NUREG-0737, RTR-NUREG-0797, RTR-NUREG-737, RTR-NUREG-797, TASK-2.D.1, TASK-TM TXX-92246, NUDOCS 9208110344 | |
| Download: ML20141M328 (68) | |
Text
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i,# war.r,,;;; Log # TXX-92246
"""" *l"llllll File # 10010 7tlELECTRIC May le, 1992 William J. Cahlli, Jr.
Gwup We Presiden a
U. S. Nuc1Sar Regulatory Commission Attn: Document Control Desk Washington, DC 20555
SUBJECT:
COMANCHE PEAK STEAM ELECTRIC STATION (CPSES) - UNIT 2 DOCKET NO. 50-t'6 NUREG-0737, ITEM II.D.1 - PERFORMANCE TESTING OF RELIEF AND SAFETY VALVES REF: 1) NUREG-0797 Supplement No. 21 "CPSES Units 1 & 2 Safety Evaluation Report,' Sectior, 22.2, Item II.D.1, dated April, 1989
- 2) NUREG-0797 Supplement No. 22, *CPSES Units 1 & 2 Safety Evaluation Report," Section 22,2, '
Item II.D.1, dated January, 1990
- 3) TU Electric letter logged TXX-4849 from Mr. W. G. Counsil to Mr. V. S. Noonan dated June 13, 1986
- 4) TV Electric lettcr logged TX-6398 from Mr. W. G. Counsil to USNRC dated April 15, 1987 Gentlemen:
In reference 1, the NRC staff reviewed and accepted the Unit 1-pressurizer-safety and relief line thermal hydraulic analysis. and stated that this issue t remained as a confirmatory issue for Unit 2. In reference 2 the NRC staff described TU Electric's commitment to provide the Unit 2 ana'ysis or demonstrate that the Unit I analysis adequately represents tie Unit 2 conditions prior to Unit 2 fuel load, The Unit 1 analysis was provided by reference 3 and revised by reference 4-in response to NRC staff questioris. The attached Unit 2 analysis was prepared by revising reference 3 to incorporate _the concerns resolved by ;
reference 4 and to incorporate Unit-2 specific data. ;
Revision bars are placed in the right margin.to identify the changes. L The same methodology, processes, and _ computer programs were used inL the Unit 1 analysis and the Unit 2 enalysis. The hydraulic forcing = function input used i for Unit 1 was also used for' Unit 2. The piping system layouts, geometrics, 4 and components are nearly identical in Units 1 and 2 with only small differences in the support system configur$ tion. Based on these unit similarities, the Unit 2 specific analysis results are similar o Unit 1.
F. f. fl F- f' Q 9908110344 920518 Ok PDR ADOCK 05000446 m.oun strat L.n.st Dattas, Texas 75201 /// de./
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TXX-92246 Page 2 of 2 The attached analyses is intended to provide >the information-needed to-allow--
the NRC to close the confirmatory item in reference 2 and thus close TMI Item 11.D.1 for CPSES Unit 2.
If there are any questions. please contact Mr. Jerge Rodriguez (214-812 8323) of the TO Electric Nuclear Licensing Department.
Sincerely,.
Wil'iam J. Caiiill. Jr.
P t D. R. Woo'dlan Docket Licensing Manager JLR/grp Attachment-c - Mr. R. D Martin, Region IV.
Resident inspectors, CPSES (2) -
Mr.- B. E. Holian, NRR.
ATTACHMENT 1 RESPONSE TO NRC REQUEST FOR ADDITIONAL INFORMATION (RAI)
DATED JULY 5, 1985 NUREG-0737, ITEM II.D.1 - PERFORMANCE TESTING OF RELIEF AND SAFETY VALVES OUESTIONS RELATED TO SELECTION OF TRANSIENTS AND INLET FLUID CONDITION OUESTION 1 The submittal identifies a steam discharge flow condition.as a limiting event for the safety valves at th!.s plant (loss of-load for maximum pressurizer prsssure and locking rotor for maximum pressurization rate). The submittal does not discuss whether single failures that could occur-after the initiating event that would result in the dynamic forces on the safety relief valves being maximized were considered.. Present a discussion that shows how single failures that would result in maximum dynamic forces on the safety relief valves were considered in selecting the limiting transient.
EEEPONSE The limiting events for dynamic forces on the safety valves were-evaluated in EPRI NP-2296-LD, " Valve Inlet Fluid Conditions for Pressurizer Safety and Relief Valves in Westinghouse-Designed Plants," dated March 1982. This report was transmitted-te Mr. H.
R. Denton, NRC, from Mr. D. P. Hoffman, Consumers Power Company, on behalf of the participating utilities on Sept. 30, 1982. This report discusses in detail-how the limiting events were selected from licensing FSAR transients, extended high pressure injection
' events, and cold overpressurization transients and provides information on single failures. The following is a summary of the effects of single failures onl determination of the limiting events based on the above report and the Chapter 15 of the CPSES FSAR.
- 1. Licensing Basis Transients--Consistent with regulatory requirements, the FSAR transients are generally censervative and evaluate single failures' effects on-several parameters including pressurizer pressure and pressurization rate. Peak pressure occurs within.a few-seconds of transient initiation. Single :f ailures within the engineered' safeguards-system _would have'little, or no effect, on the pressurization rate or peak pressure observed.- For the loss of. load and-locked rotor event (which bounds other loss of flow events), the-peak pressure and pressurization rates are maximized by 1
, assuming the following: 1 I
1 1
PAGE 1
4
-No Reactor Trip on Turbine Trip
-No Steam Dump
-Pressurizer spray and heaters are off
-No Rod Control System operation
-Safety Injection does not operate
-Auxiliary Feedwater does not operate
-Control Systems function only if it increases the severity of the event.
For the main feedline rupture event (which also bounds the small steam break event with respect to the peak primary pressure), a sensitivity analysis was performed to determine the break conditions causing the worst case results. In addition, the analysis was performed at Engineered Safety Features design power level and no credit was taken for the following:
-Pressurizer power operated relief valves (PORV's)
-Pressurizer Spray
-Turbine-driven Auxiliary Feedwater Pump
-Protective logic signals from:
-High pressurizer pressure
-overtemperature N16
-High pressurizer level
-High containment pressure The loss of normal feedwater event with off-site power bounds the same event accompanied by.a loss of offs!te power (station blackout). To maximize the dynamic effects on the safety valves, the analysis was performed at Engineered Safety Features design power level and the following were assumed:
-Worst single failure in the auxiliary feedwater system
-Secondary system power operated relief valves and condenser steam dump do not operate
-Pressurizer PORV's operate to decrease margin to water relief through the valves
-Normal reactor control systems do not operate For FSAR transients _which result in reactivity and power distribution anomalies, the events are generally terminated or mitigated by the reactor protection system prior to challenging the pressurizer safety valves. Single failures that cause an anticipated transient without trip are beyond the scope of NUREG-0737, Item-II.D.1.
- 2. Extended High Pressure Injection Events--Events which could cause an extended high pressure injection event and subsequent passage of steam or water through the safety valves include steam line ruptures (which bounds an accidental depressurization of the secondary system) and spurious initiation'of'the high pressure safety injection system (SIS). For the steam line rupture, the SIS is initiated on a low pressurizer pressure which then begins to refill and repressurize the pressurizer. In terms of surge flow and liquid temperature at the valve inlet, this event is bounded PAGE 2
by the spurious initiation of high pressure SIS at power.
The spurious initiation of high pressure SIS assumes the following:
-No direct reactor trip on SI signal
-All protective systems unavailable except low pressurizer pressure reactor trip
-No control systems operable
- 3. Cold Overpressure Events--The cold cverpressure events are described in detail in NP-2296-LD. As discussed in the report, the PORV with the lowest setpoint is assumed to fail. In this condition, the second PORV can adequately mitigate the pressure rise. Therefore, prior to requiring operation of the safety velves, both PORV's must fail.
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OUESTION 2 In the PORV operability discussions on low temperature overpressure transients, variable fluid conditions (steam or water) and temperatures (saturated to subcooled) were identified.
To assure that the relief valves operate under all cold overpressurization events, include the nitrogen bubble case in the discussion. Also, identify the test data that demonstrate operability over the entire range of conditions. Confirm that the high pressure steam tests demonstrate valve operability for the low pressure steam case for both opening and closing of the relief valve. In addition indicate the set pressure of the PORV for cold overpressure protection.
RESPONSE
EPRI test conditions for the PORV's were chosen based on expected inlet fluid conditions. Comanche Peak's* cold overpressure (COP) event, including the nitrogen bubble case, was considered when a formulating the enveloping EPRI test conditions. Maximum temperature and pressure conditions that can be achieved at the PORV inlet occur for steam bubblo operation. Since pressure is normally maintained below the PORV setpoint, the maximum eteam-and saturated liquid pressure maintained in the pressurizer during startup and shutdown operations in anticipation of the COP event would occur at the PORV-setpoint. PORV setpoints are given below. Tests were limited but designed to confirm operability over a full range of expected inlet conditions. Steam, steam-to-water, and water flow tests were conducted. Results of these can be found in EPRI Report EPRI NP-2670-LD, Volume 8.
Although the steam tests were conducted only at the higher pressure, it is expected that satisfactory operation would also
., result at the less severe lower pressures. This can be seen by the successful low pressure, low temperature water tests.
The PORV setpoint are as follows:
PORV A PORV B TEMPERATURE PRESSURE TEMPERATURE PRESSURE DEGREES F PSIG DEGREES F PSIG 0 505 0 445 70 (Note) 505 70 (Note) 445 190 505 150 445 210 520 190 435 230 560 210 440 250 640 230 465 260 670 260 565 280 685 350 570 380 685 380 570 450 2350 460 2350 NOTE: 70 Degrees F is not used as a calibration point.
- Unit 2 PAGE 4
OUESTION 3 , ;g The'EPRI/ Marshall-block valva tests were: performed with:-the valves -in: a- horizontal _ position -_(valve stems-! vertical) .- ___Ider.tify
- the - orientation of the Comanche- Peak- 2 block valves. . If:they-are-oriented in any direction other-than-horizontal, provide detailed-
~
information on how the EPRI_ data z was extrapolated _to assure operability of.the block valve in-the. plant specific-orientation.
RESPONSE
As required by design drawings and vdrified by~.as-_ built inspection,.the block valves are installed horizontally'(valve stems vertical).
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OUESTION 4 The submittal referenced EPRI Safety Valve Test Nos. (929, 931a, 932, 1406, 1411, 1415 and 1419) as being suitable for evaluating the operability of the Comanche Peak- 2 valves. The EPRI test l blowdowns exceeded the 5% value given in the valve specifications. These increased blowdowns occurred with typical plant ring settings (-71 or -77, -18 relative to the bottom of
, the ring disk) that are much different from the-ring settings specified in the submittal (i.e., -250, -18 relative to the level position). Based on a review of the EFRI test'results, the Comanche Peat ring settings will produce larger blowdowns (i.e.,
greater than 10%). The higher blowdowns could cause a rise in pressurizer water level such that water may reach the safety valve inlet line and result in a steam-water flow situation.
Also the pressure might be sufficiently decreased that adequate cooling might-not be achieved for decay heat removal. Discuss these consequences of higher blowdowns if increased blowdowr' are expected. In order that the suitability of Comanche Peak 2 safety valve ring setting may be evaluated, provide the factory i ring settings relative to the bottom of the ring disk or relative to the upper limit of ring travel. -
RESPONSE
The previously provided guide ring settings for the safety valves were referenced to the highest locked position rather than relative to the " LEVEL" position (bottom of the disc ring). The following are the ring settings relative to " LEVEL" consistent with the EPRI tested valve settings of -71 to -77 notches (Guide Ring) and -18 (Nozzle Ring):
Valve Valve Nozzle Guide Tag Serial Ring Ring Number Number (Notches) (Notches) 1-8010A N56964-00-0070 -18 -103 1-8010B N56964-00-0069 -18 - 99 1-8010C N56964-00-0006 -18 -100 2-8010A N56964-00-0008 -18 - 82 2-8010B N56964-00-0007 -18 - 94 2-8010C N56564-00-0068 -18 -100 The ring = settings in the table above are all typical ring settings for the model of valve, as are the EPRI tested valve ring settings. The difference in individual settings is due to-manufacturing. dimensional tolerances of the individual valve components. The EPRI valve and the CPSES valves were all tested at Crosby using the same procedure to establish the ring setting to obtain the specified blowdown characteristics. These blowdown characteristics were verified in the'EPRI test program.
Additionally, since the CPSES valve inlet pressure drops are equal to that in the EPRI test configuration (See Question 7),
the CPSES valve stability and blowdown are expected to be consistent with that in the test.
PAGE 6
- OUESTIONS RELATED TO-VALVI OPERABILITY L
QUESTION 5
- The submittal does not address the-expected blowdown'and'
! corresponding valve performance" for' thc' Comanche Peak 2 ~ Crosby 4
, 6M6-safety valves at the plant iring' settings.c-The submittal
- states that the ring setting for the safety. valves;are:-250, -18__
- 1 relative to.the level position. The submitta1Lstates1thatithe. -
l comparable EPRI tests were Test Nos.L 929,:431a,_932,;1406, 1411, i 1415, and 1419. The ring settings;for these tests were;-77'or
- -71, -13 relative to the bottom.of chc-ring disk. - Explain.how-the expected values for backpressure and-blowdownLeorresponding '
to the Comanche-Peak 1 ring. settings-were extrapolated or calculated from' test data with:such-different ring settings, and l identify the values for. backpressure;and blowdown so; determined.
s
RESPONSE
i Refer.to the Response to Question.4. TheLexpected blowdown and.
corresponding valve performance:for the CPSES-safety _ valves.are-
, expected to_compara'le a to the EPRI: test.
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OUESTION 6 Thermal expansion of the pressurizer and piping will induce loads on the safety and relief valve piping and on the pressurizer nozzles. This thermal expansion would also-induce loading on the inlet flange of a-safety or relief valve at the time the valve is required to lift. Evaluate the effects that this loading may have on valve operability.
RESPONSE
The loads induced on the safety and relief valves tested by EPRI exceed the loads for the Comanche Peak
- safety and relief valves.
The maximum moment tested for the 6M6 valve was during test 908 l and was 298.75 in-K. The laggest moment predicted for the safety valve inlet at Comanche Peak is 172.42g. This demonstrates functionability for the Comanche Peak' safety valves. [
, Likewise, a bending moment of 43.0 in-K was induced in the inlet of the Copes-Vulcan PORV test valve per EPRI 64-CV-174-2S. The largest moment predicted for the PORV inlet in the Comanche Peak
- valves is 21.625. . This demonstrates functionability of tha Comanche Peak
- relief valves.
- Unit 2 l_
! PAGE 8
OUESTTON 7 The EPRI guide for application of test results to plant specific evaluations suggests that the inlet piping pressure drops for the Crosby 6M6 EPRI test valves be compared to the calculated Comanche Peak- 2 Crosby 6M6 inlet piping pressure drop as a means of assessing valve stability. Provide the pressure drop calculations and the assessment of valve-stability. If alternate pressure drop calculations were performed, provide a detailed explanation, and a detailed valve stability assessment.
RESPONSE
The safety valve inlet piping pressure drop calculation has been performed in accordance with the "EPRI PWR Safety and Relief Valve Test Program Guide for Application of Valve ~est Program Results to Plant-Specific Evaluations," Interim Report, July 1982. The calculation is available ,b CPSES for NRC review and the following table provides a summary for the bounding valve piping configuration:
EPRI CPSES TIjlLT CALCUIATI.Q AP OPENING 263 PSI 255 - 269 PSI ,
AP CLOSING 181 PSI 152 - 158 PSI As can be seen in the table, the CPSES pressure difference is comparable to that in the EPRI test. -The only value for which the CPSES calculated pressure difference is higher than the EPRI test condition is when a very conservative L/D valve is used in the calculation. This difference.is insignificant-(<2%).and is within the uncertainties of the methodology, L/D data, and mathematical round-off error.
As stated in the EPRI guide,Lif the-calculated pressure difference is less than the test data,Lthe in-plant-valve would be expected to have performance at least as stable as the test valve; therefore, the CPSES valves are expected'to have stability comparable to or better than the tested valve.
PAGE 9
OUESTIONS RELATED TO THERMAL HYDRAULIC ANALYSIS OUESTION 8 The adequacy of the thermal hydraulic analysis could.not be-verified since it is not presented in tne submittal. Provide detailed information on the program used so that the methodology for generating fluid parameters can be-evaluated. _ Identify parameters such as timestep, valve flow area, pressure ramp rate, choked flow junction, and node spacing and discuss the rationale for their selection. Provide detailed information on how the program or methodology was verified for this application.
RESPONSE
The ITCHVALVE Computer Program was used for the thermal hydraulic analysis. (See Appendix A)
The adequacy of the. thermal-hydraulic analyses can be verified by the comparison of analytical and test results for thermal-hydraulic-loadings in safety valve discharge-piping forL EPRI Tests 908 and 917 presented in Appendix A,_Section A.l.2.
In that evaluation, node spacing _and time-step size were_ selected on the basis of stable solutions of the: characteristic equations and natching of test data. The sefety valve full open-flow area of 0.022_ft was_used in the model. This_ area is.slightly _
smaller than the Crosby'M-orifice area of 0.025 ft: _for_the tested valve, but results in a good-analytical' match of the tested fully open valve flow rate. Appropriate water temperatures were used. All pertinent data, including friction factors, loss factors and flow areas 1were-based upon representative calculations and the. system' layout. ~Modeling of the water was conducted with the water seal upstream of the valves prior to transient intitiation. .Atjtime = 0+, the transient was initiated and thenwater slug position:was-analytically calculated during1and subsequent-to valve opening. .;
The Comanche: Peak Unit 2 plant specific thermal-hydraulic--. l l analysis was conducted based upon-the same approach 1as used for the comparison _to test data. Node spacing and time-steo_ size-were consistent with values utilizac_'in the-comparison. Valve flow areas were selected based upontactual valve 1 data Vith appropriate margins'appliedito account-for-flow rate:
uncertainties. All pertinent data, including = friction factors, loss factors and' flow-areas were based upon representative
.. calculations and the system layout. ModelingLof_the_ water slug from.a-temperature profile, considering initial location:and movement post-transient ^ initiation, was_consisteit.with-the: ,
comparison study. The pressurizer' pressure was1 held constant' i through'the transient attinitial values (see. Response 11). ?
Choked flow is checked-internally;and automatically every; time-step to:ensureJthe proper formulation is applied at every i
-ficw path. 1 i
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- _ _ _ _ _ _ - - _ _ _ _ - - _ - - __ 2
OUESTION 9 Discuss whether multiple valve actuations were considered in the thermal hydraulic analysis. The maximum loading on the piping typically occurs under a multiple valve actuation condition during which the valves open in sequence. The experience of EG&G Idaho indicates that the maximum loading occurs when the sequence of opening is such that the initial pressure waves from opening of the safety valves reach the common header downstream simultaneously. Additionally, if a PORV is discharging with flow in the common header, the piping loads could be sicnificantly affected. Provide justification that sequential cpaning of the valves under multiple valve actuation conditions was considered.
RESPONSE
Two valve opening cases were addressed: 1) the three safety valves opening simultaneously and discharging without PORV flow and 2) the two PORV's opening simultaneously without safety valve flow. The three safety valves are identical and have the same set pressure (i 1 percent). It was, therefore, assumed for the analysis that all three safety valves open simultaneously without PORV flow. Because of similarity, the two PORV's were also assumed to open simultaneously without safety valve flow.
Maximum common header (area of piping common to both safety and relief valve discharge piping) forces theoretically could be expected when valve sequencing is such that the initial pressure waves from valve opening recch a common downstream junction simultaneously. Based upon engineering judgment:
- 1. The simultaneous opening of the safety valves results in practically simultaneous peak loads at the safety valves common branch point. The_ peak forces occur within approximately 60 milliseconds of each other. As a result, no significant impact in the common header region due to safety valve discharge is expected if the valve sequencing is adjusted such that the peaks of the initial pressure waves reach a common downstream header point simultaneously.
- 2. The total lengths of effective piping between each valve outlet <rd the common junction point are not exactly the sama. The likelihood of the valve phasing being such to compensate for the'different lengths is very small; therefore; the peaks of the initial pressure waves from valve opening, either safety or-relief, would not reach a common downstream junction at exactly the same time.
- 3. -There is a significant amount of piping and' dynamic supports between the valve outlets and the common point. In the unlikely event that increased loadings from this common point to the relief tank were to occur, the effects would be limited primarily from near the common point to the relief tank. Significant isolation of the common region from the upstream region PAGE 11
because of thefsupport configuration exists.-
Therefore, the operability-andLintegrity of-the-valves, the inlet lines toithe valves, or the nozzles'on-the pressurizer would not be jeopardized.
Considerable margin exists between:the conservatively calculated:-
- maximum stresses and the-allowable stresses for the-safety valve-discharge event. - Table A '7 in- Appendix A illustrates this for
'the upstream-piping-for the faulted load combinations .-
Additionally, for the downstream piping for the loading.
combination (P + DEADWEIGHT + SAFETY VALVE DISCHARGE' , the maximum conservatively calculated-stress is less than 80% of the allowable value.
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PACE 12
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OCESTION 10 -
, Identify the program or methodology for calculating the fluid forces-for the structural analysis. Discuss the accuracy of the results and the procedures used-to qualify the program or methodology.
RESPONSE
See Appendix A, Sections A.1.1 and A.1.2.
l PAGE 13
8 OUESTION 11 Identify the i..itial conditions for the safety and relief valve thermal-hydraulic analyses. Describe the method used for treating valve resistance in the analyses and report flow rates corresponding to the resistances used. Because the ASME Code requires derating of the safety valves to 90% of actual flow capacity, the safety valve analysis should be based on a flow rating equal to 111% of the flow rate stamped on the valve, 4
unless another flow rate can be justified. Provide further information explaining how derating of the safety valves was
- handled and describing methods used to establish-flow rates for the safety valves and PORVs in the thermal hydraulic analysis.
RESPONSI
, The initial conditions for the safety valve water slug discharge casa included:
P (Upstream) ' 2575 psia h (Steam,Up tream) = 1110 Btu /lb i
h (Water,Upriream) =Enthalpy based upon a temperature profile consistent with EPRI safety valve dis-change Test 917, i.e., approximately 300* F-at the valve inlet and saturation temperature at the steam-water interface.
(An insulation arrangement was-implemented that resulted in acceptable heating of the
, loop seals).
P (Downstream) = 18.0 psia The pressurizer conditions were held constant for the transient at 2575 psia and 1110 Btu /lb.
The initial conditions for the relief valve slug discharge case-included:
P (Upstream) = 2350 psia h (Steam, Upstream) = 1162.4 Btu /lb.
T (Water, Upstream) = 150 F-P (Downstream) = 18.0 psia The pressurizer conditions were held constant for the entire
- transient at 2350 psia and 1162.4 Btu /lb.
I PAGE 14-
Safety and relief valves are-modeled as two-way junctions.' The, pressure drop across the valve, provided the system:is sub-cooled is~given by:
1 P =
joCD v where P = pressure-drop.
CD = discharge coefficient- =-f(Cv) j6 = fluid density v = velocity-through the valve In the case of choking at the valve, the velocity-at the valve-orifice area is set at the sonic velocity., Upstream and- -'
downstream boundary conditions are iteratively set'to conserve mass and energy. Choked flow is. internally checked to ensure the proper formulation is applied.-
The nominal steam flow rating :for thel CPSES. Crosby ' safety valves
-(orifice _ size'6M6) at 2500_ psia-is' 420,000:lb/hr. - TheEminimum analytically determined steam flow through each of 5the safety-valves ic greater than 503,50011b/hr. This is equivalent- to: a flow of 120 percent oferated. ,;The maximum: expected. steam flow-through the Copes-Vulcan PORV's 31s 210,000' lb/hr.' - ~ Values greater than 291,000 lb/hr were analytically calculated.-
of 139 percent of rated.
Thiscis a_ flow ,
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OUESTIONS RELATED TO STRUCTU.RAL ANALYSIS l OUESTION 12 The submittal does not present details of the structural analysis. To allev for a complete evaluation of the methods used and results obtained from the structural analysis, please provide reports containing at least the following information:
(a) A detailed description of the methods used to perform the analysis. Identify the computer programs used for the analysis and how these programs were verified.
(b) A description of the method used to apply the fluid forces to the structural model. Since the forces acting on a typical pipe segment are composed of a net, or " wave", force and opposing " blowdown" forces, describe the methods for handling both types cf forces.
(c) A description of methods used to model supports, the pressurizer and relief tank connections, and the safety valve bonnet assemblies and PORV actuator.
(d) An identification of the load combinations performed in the analysis together with the allowable stress limits.
Differentiate between load combinations used in the piping upstream and downstream of the valve. Explain the mathematical methods used to perform the load combinations, and identify the governing codes and standards used to determine piping and. support adequacy.
(e) An evaluation of the resul?- vt the structural analysis, including identification of overstressed locations and a description of modifications, if any.
(f) A sketch of the structural model-showing. lumped mass locations, pipe sizes, and application points of fluid forces.
(g) A copy of the structural analysis report.-
EESPONSE (a) See Appendix A.
-(b) For each piping segment unbalanced or " net" forces were calculated.-~The hydraulic forcing functions were then simultaneously applied to the appropriate segment of'the structural model' . The axial extension from the balancing forces (opposing " blowdown" forces) on each end of the structural segments was considered in tne evaluation.
However, this effect for this particular application was found to be negligible relative to the net unbalanced l forces. Referring to structural analyses comparisons to test results presented in Appendix A for Tests 908 and 917, maximum support and pipe loads compared well l
1 PAGE 16 i
e
I with test results. Good comparisons of the maximum ,
displacement values downstream of the safety valve were also seen.
(c) The structural supports were modeled in suf ficient detail to analytically represent the system. The shock suppressors and struts were modeled by inputting a stiffness in series with the piping. Calculated stiffness values were utilized.
All supports were linear and a linear overall system analysis was conducted.
A simplified model was used to represent the pressurizer.
The pressurizer nozzlos and pipe connections were represented vith appropriate pipe properties. The downstream piping terminated at the relief tank flange where the model was anchored.
The valve bonnet assemblies and the relief valve actuators were modeled as extended masses displaced from the pipe centerline. The valve weight and center of gravity were selected from the valve drawings. The stem properties-(diameter and thickness) were then selected to represent the y valve frequency (d) For analysis purposes, the governing code for the piping qualification is the ASME Boiler and Pressure Vessel Code Section III, 1977 Edition, with Addenda to and including Summer 1979. See Appendix A, Pages A-9 to A-17 for a discussion of the stress analysis.
The governing code for the support qualification is the ASME Boiler and Pressure W-.r.el Code Section III-Division I, subsection.NF, 197% Edition, with addenda to and including Winter 1979.
(e) See Appendix A, Pages A-18 thru A-24 . No overstresses exist l for either upstream or downstream piping. No modifications l are required .
(f) Figures A-1 and A-8 in Appendix A illustrate the lumped l mass locations and by number the application points of fluid forces for the safety valve dis: m case. The piping upstream of the safety valves is ...:h Schedule 160 ,
(numbers 1-6, 29-34, and 38-43 on the figures). The piping between the safety valve outlets and the common header (numbers 7-9, 35-37 and 44-46 on the figures) is 6-inch Schedule 40. The piping in the header pipework is 12-inch Schedule 80 (numbers 10-27 on the figures).
Figures A-2 and A-9 in Appendix A illustrate the lumped l mass locations and by number the appl.ication points of the fluid forces for the relief valve discharge case. A large fraction of the piping upstream of the relief valves is 6-inch Schedule 160 (numbers 1-5 on the figures). The remainder of the upstream piping and the piping between the valve outlets and the common header is 3-inch Schedule 160 PAGE 17
1 4
L piping (numbers 6-14 and 34-41 on the~ figure). The piping
' in the header is 12-inch Schedule 80 -(numbers 15-32 on the figures).
(g) The summary report of the structural analysis is included in Appendix A.
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PAGE 18
f OUESTION 13 According to results of EpRI tests, high frequency pressure oscillations of 170-2o0 Hz typically occur in the piping upstream of the safety valve while loop seal water passes through the valve. An evaluation of this phenomena is documented in the Westinghouse report WCAP 10105 and states that the acoustic pressures occurring prior to and during safety valve discharge are below the maximum permissible pres are. The study discussed in the Westinghouse report determined the. maximum permissible pressure for the inlet piping and established the maximum allowable bending moments for Level C Service Condition in the inlet piping based on the maximum transient pressure measured or calculated. While the internal pressures are lower than the maximum permissible pressure, the pressure oscillations could potentially excite high frequency vibration modes in the piping, creating bending moments in the inlet piping that should be combined with moments from other' appropriate mechanical loads.
Provide one of the following:
(a) a comparison of the expected peak prestures and bending moments with the allowable values reported in the WCAp
, report, or (b) justification for other alternate allowable pressure and bending moments with a similar comparison with peak L
pressures and moments induced in the plant piping.
RESPONSE
The piping system response for Comanche Peak Unit No. 2 including l the safety valve loop seal region, is due to frequencies less than 100 HZ. The frequency of the forces and moments in the 170-260 HZ range potentially induced by the pressure oscillations is significantly greater than this frequency. The upper limit of significant frequency content for similar systems is also much less than this (170-260 HZ) range. Industry data indicates that only frequencies of 100 HZ or lass are meaningful. The EPRI data confirr.s this. -Consequently, no significant bending moment .
, during the pressure oscillation phase of the transient will occur.
1 Pressure stresses based upon a pressure of 2485 psig were included with the bending moments resulting from the deadweight and the safety valve discharge piping loads. Because of the time phasing of the pressure-oscillation (during water slug discharge tnrough the safety valve) and the discharge piping loads (subsequent to water slug discharge through the valve) this. term and moment tern were not added. They do not occur coincidentally. A comparison of the intensified bending moments from the stress evaluation and the allowable moment presented in WCAP-10105 shows that all values are below the allowables.
Specifically, the maximum allowable moment from Table 4-7 of ;
WCAP-10105 for 6-inch Schedule 160 piping for an internal pressure of 5000 pai is 516 in-kips. For the combination
~
pressure plus deadweight plus safety valve water slug discharge, the maximum stresses occurred at nodes 3151, 3151, 3150, and 314<0 respectively for the butt weld, long radius elbow branch connection, and CRUN. The moments corresponding to these components are 187.29, 187.29, and 187.38 in-kips.
PAGE 19 l
t' 1
OUESTIONS RELATING TO PORV CONTROL CIRCUITRY OUESTION 14-NUREG 0737, Item II.D.1 requires that the plant-specific-PORV-~
control - circuitry be qualified for design-basis ; transients -and i accidents. -Provido'information which demonstrates that'this -
! requirement has been fulfilled.
RESPONSE
i
- Both PORV's have been provided.with remote manual control
! capability from the Control Room (CR), and thefHot Shutdown Panel
- (HSP) in addition to the automatic actuation circuitry described-
- in the CPSES FSARES.4.ll, 5.4.13, and-7.6.8.
- The following components of the PORV control circuitry are j qualified for the events shown
- -
I The documentation to support the. qualification of the belou equipment-
- is availsble for insnaction at-CPSES.
EOUIPMENT EVENTS-l 1. Copes Vulcan air-operated fail . LOCA, ? Post-LOCA,- MSLB, closed relief valves FWLB, SSE*
. 2. Limit Switches LOCA' Post-LOCA, MSLB, FWLB, SSE~
F
- 3. Solenoid Valves _LOCA,. Post-LOCA, MSLB
i <
- 5. Control Switches & Control Bds. ** ,:SSE -
- 7. Transfer. Switches- HELB, SSE:-
- 9. Relays ** ,?SSE 5
-: HELB depending on:
locationE SSE'
' NOTES: * --Although environmentally.qualifiedifor-LOCA' conditions,'its operation is!notLassumed, f The1 valve is qualified forca. seismic event to -
- maintain Tits L pressure ; retaining 1 function f only.
i
- The control: board,; control >boardiswitches, and
- - relays are located'in thefcontrol' room-and'are-
- -not; subject tofan' adverse environment.
PAGE,20 i
i.
. _ . . . 4 - _ . . . . . .- .. ., . . . . _ . -m _,. . . . . , . _ . . . . . . - - - - ~ . . - - c-, , ~.
S . wir 4, 3- A m - L y 9 f
d 4 S
I~
i e
I' t
s d
1 f
n a
o i
APPENDIX A d
it 4
4 A
a i
g- .
i k'
J d
a N 1 yp'+ ie .A. .- w -= r e W"- e t- W *18
PRESSURIZER SAFETY AND RELIEF LINE T'!?i.UA710N SUWARY REPORT TEXAS UT!L'71ES GENERATING C0liPANY COPANCHE PEAK - UNIT 2 ,
s
.]
(
)
f i
A.1 F5Em'R12?* 11rrTY AND REttF LINE ANittSJ1 rmmms tvs PresswMer Safety and Relief Valve (PSARV) discharge piping system for pressurized water reactors, located on the top of the pressuriter, provides overpresswa protection for the'reactw coolant system. - A water- seal is saintained . upstream of seen pressuriser safety and relief valve to prevent.
a steam interface at the valve seat. This',veter seal gractically eliminates the possibility of vslve leakage. Vnile this arrar4 ment 1 maximizes the plant availabi.11ty, the water slug, driven by high system pressure .upon actuation 'af the valves, generstes severe hydrsulic shock loads on the piping and supports.
i A-1
W l
a d
e Under NURIG 0737,Section II.D.1, "Perfomance Testing cf FER and NR ,
Relief and Safety Valves", all operating plant licensees and applicants are required to conduct testing to qualify the reactor coolant rista relier and safety valves under expected operating conditions for design-basis transients and accidents. In addition to the qualification of valves, the '
functionsbility and structural integrity of the as-built discharge piping
- and wp;crts must also be demonstrated on a plant specific basis.
In response to these requirments, a progre for the performance testing of
' NR safety and relief valves was formulated by the Electric Power Research Institute (EPRI). The rimary objective of the Test Progra was to provide full scale test data confitning the functionability of the reactor coolant syste power operated relief valves and safe'.y valves for aspected a
operating and accident ocnditions. The second objective of the Fogra was to obtain sufficient piping thermal hydraulic load data to semit confituation of models which may be utilized for plant ur.ique analysis of safety and relief valve discharge piping systas.
The method of analyses described in the following sections is consistent with the findings of the atormentioned EPRI Safety and Relief Valve Test f Progr a.
9 9
t j
1 l
A-2 l
_ y - - , , , _ . r f e v
1 4
j .
5
.i i
l 4
i j . A.l.1 FLU'i HYDRAt12C NODE 1. ,
l t s-1 When the pressurizer pressure reaches the set pressure (2,500 psia for f .;afe ty valve and 2,350 pais for a relief valve) and the valve opens, the j high pressure staan in the pressurizer forces the water in the water seal' :
!' loop through the valve and down the piping systs to the pressr. sc relier
. tank. For the pressurizer astety and relief piping system m .lytical i hydraulic models, as shcwn in figures A 1 and A-2 were e.evered to 4
l represent the conditions described above. ,
h
- The otsputer code ITQWALVE was toed to perform the transient hydraulic analysis for the systas. This program uses the Method of Charse,teristics f generate peruseters as a ftrction of. time.
l approach tc fluid
! One-dimensional Guid flow calculations applying both - the implicit and j explicit characteristic methods are performed.. Using this approach, the
! piping network is input as a series of single pipes.. The network is generally ' joined - Mether at one or more - places try two or three-way .
! junctions. .Each of the' single pipes has associated with it friction j facter s, angles of elevation and ficW areas.
o I.
- .aenation equations can be' converted to the follcuing characteristic 4
squetiens:
d2 l
g = Y+c
?
I dP dV 4c g
- ec g . c(F + pgcess) g 8E i
i- .
.A-3
a e r
4 T
d g' e Y=C 1
dP dV
-C(F
- pgC050) 4c '
j g - oc g e ah
'E 2 - th/so
} C
- an 1 4
30 *7 za variable of length measurunent
- s time 7
V: Guid velocity -
- ca sonic velocity i Ps pressure I os Guid density i F new resistance i ga gravity i e angle off vertical 1
i Js conversion facter for converting pressure units to equivalent l heat totits h: enthalpy '
j q"'s rate of heat generation per unit pipe length i,
ice computer program possesses special provisions to allow analysis of
! valve cpening and closing situations.
R uid acceleration inside the pipe generates reaction forces on all i segments of the lines that are bounded at either end by an elbow or bend.
Reaction forces resulting fran Q uid pressure and momentum variations are.
calculated. These forces can be expressed in terms of the nuid properties avsilable from the transient hydraulic analysis, performed using program ITCHVALVE. The acmentum equation can be expressec ;A . vector form as:
A-4 P
.,ye- _
, . , , , , , - - -- - y9 , ,
e -
e e
^*
F gy .
c h t e
v sV dv +
c, eV(V
- ndA)
From this equation, the total force on the pipe can be derived:
rg (1 - ces eg) ,y r2 II
- 883 *2} nW 8'" *2 N ,8end 2 F,4p, . g g sin sg at . lend 1 Ic
+ straight hdi c pi pe j
A: piping flow area y volume F: force ra radius of cur-nture of appropriate elbow a s angle of appropriate elbow W2 mass acceleration 9, a gravitational constant 111 othe.r tenas are previously defined.
Unbalanced forces are calculated for esch straight sognent of pipe fran the The time-histories of pressurizer to the relief tark using progran FORFUll.
these forces are stcred on tape to be used for the subsequent structural aralysis of the pressurizer safety and relief lines.
A.1.2 COMPARISON TO EPRI TEST RESUt.T5 -
Piping lead data has been generated from the tests conducted by EFRI at the ccmbustion Engineering Test Facility. Pertinent tests simulating dynamic '
opening of the safety ve.1ves for representative caneercial upstre n The resulting downstream piping leadings envirorments were carried out.
__ __ _ _ _ _ _ __._ __ __ _1 AcdL
~ . . - . - . .
a and res;:enses were measwed. Upstrean envircrzents for particular valve opening cases of importance, which envelope the ccamercial scenarios, are:
A, ce!* water eineMarne re11med tv sta= - steam between the pressure source rad the loop seal - cald loop seal between the staac and the valve.
B. Het water et aew --re re11med tv stem - steam between the pressure source and the loop seal - hot loop seal between the staan and the valve. .
C. Stee di ne***te - steun between the pressure source and the valve.
Specific thermal hydraulic and structural ana1,yses have been completed for the Combustica Engineering Test Configuration. Figure A-3 . illustrates the placement of force sensurement sensors at the test site. Figurs. A-4 ar4 A-5 illustrate a ecmparison of the thermal hydrsu11cally calculated results using the ITQNALVE and .FORF1JN canputer prograns versus experimental results. For' test 908, the cold wat.ar discharge follcwed by steam case, Figure A-4 1.11ustrates the force time history of the long vertical run (VE32/VE33) inusediately downstream of the safety valve.
For test 917, the hot water discharge fo11 cued by steam case, Figure A-5 C
1.11ustratos the thermal bydrmulically calculated and the.e1,erimentally determined force time history for (WE32/VE33) . Although not presented here, ca parisons were also ande to the test data svailable for safety valve discharge without a loop seal (steam discharge). '
The application of the ITONALVE and FORFUN :camouter programs for-calculating the fluid-induced loads on the piping downstream of the safety and relief valves has been demonstrated. Although not presented here, the I
'A-6 l y
1 i l 3 .
i I !
- i
+
i i espetuity has also been shown by c.irect cce;arison to the sclutions cf casssical probisms.
?
'he appliention of the structural empWer programs (discussed in Section i
1 A,1,3 for calculating the system response has also been demonstrated.
Structural models representative' cf the Cabustion Engineering Test j
Configuration were developed. Figure A-6 and A-7 ulustrate, ,
! respectively, s cca;mrison of the structural analysis results and the experimental results at Socation (VE32/WE33) for test 908 and test 917, i respectively.
5 i
i 1
$ i 0
'?
i 4
L 4
a i
s t
W-A-7
. _ . _ - . . - . _ _ . _ - . _ . . .___-.3. . -
l l '
! l i
< i I ;
} }
l i :
! J !
i ~!
1 e
i
! A.l.3 VALVE YHRUST ANALY313 I 1 The safety and relief lines were modeled statically - and ' dynatically . '
l The asthmatical i_
(seinsically), >
model used in the sei mic analysis tsas_ modified for the valve' thrust f The :
l analysis to represent' the safety ' and : relief ' valve discharge. +
l time-history bydrmulic foross detemined try FORFUN were, applied to the piping system itap mass points. The dynamic solution for the valve thrust ~ l l was nbtained try using s-modified-predictor-corrector-integration technique
~
- and normal mode theory. ,
! he tine-history solution wss _ found using progra FUFM3* he input to i
this program consists of natural frequencies, norinal --modes, and applied j
forces . The natura1= frequencios and normal modes for the modified- [
l- I pressurizer safety and relief line dynamic model t#ere detemined with the
[ he time-history displsosaent response was stored on - t
[ WE37DYN progr e .
magnetic tape for later use in amputing the total systa response ldue to
[ the valve thrust conditions.- The time-history displacementsof. the FIXFN3
[ '
l progrsa were used as imput .to the WESTDYM2* progr e ito detemined the' l time-history laternal forces and ' deflections at-each and of the piping i
i elments.. For tnis calculation, the displacanents were9 treated as imposed
-deflections-en the pressurizer astety and relief line masses. 'he sclution b
was stored on tapa for'1 ster use in the piping stress evaluation-and piping .
1 l
-support load evaluation.
l time-history internal forca.1 and displacements of the WESTDYN2,pr'ogram - b
- The i
were used as input to the POSDYN2
- program to detemine the assimus E forces, asents and displacements that exist at each end of the~ piping elements and the anximum loses for piping . supports. The:results from' progre POSDYN2*
- Subroutines-of Program WESTDYN
- ~
-_ . . ..--. _ .- _., _ - _ _ ,- ..._ _A-8 .
. . - . - . , . _ . - . _ _ _ _ _ _ _ _ _ ~ _ _ ~
a are saved on TAPEla for future use '.n piping stress analysis and support 16ad evaluation. .
A.1.4 lE'DICD OF STRESS EVALUATION .
In order to evaluate the pressurizer anfety and relief valve piping, appropriate load ca binations and acceptance criteria were developed. The load ombinations and seceptance criteria are identical to those recommended by the piping st-4ttee of the PWR PSARV test progra end are , cut 11ned in Table A-1 and A-2 with a definition of load attreviation ;rovided in Table A-3.
A .1. 4.1 PRIMARY STRESS EVALUATION .
In order to perfons a primary stress evaluation in accordance with the rules of the Code , definitions of stress ocabinations era required for the nonnal, upset, emergency and faulted plant conditions. Tables A-1 and A-2 illustrate the allowable stress intensities for the ap;ro;riate cabination. Tshle A-3 defines all pertinent ta.,as.
A.1. 4 .2 DESIGN CONDITIONS The piping minista wall thickness, t,, is calculated in scoordance with the Code. The actual pipe minimaa wall thickness meets the Code requirenent.
-The canbined stresses due to primary loadings of pressure, weight, and design mechanical loeds calculated using. applicable stress intensity factors must not exceed the allowable limit. The resultant ament, M g, due
. to loads caused by-weight and design mechanical loads is c21culated uair4-
' the following equation:
A ,
9
2 2
. M- . M S
,i , ,A,g Kyt 7wt /WL i-
) -
1/2
+ M Z
+ S2
- wt M*WL)l a
w.r .
4 M, ,My ,M r a deakeight noennt components g
M- ,H ,M e d4 sign sachanical -load moment components -
zggg:
- WL #DML The anximum stresses due to pressure, weight, and DK.~ in the piping syste are reported in Tables A-4 thru A-12.
t C
A-10 4
e A.1.4.3 UPSC COCITZNS W.e combined stresses due to the priawry loadings of pressura, weight, OBE seimic, and relict valve thrust loadings calculated using the applicable strass intensity facters saast not exceed the allwahles. The resultant ancun acments, M g
, due to Iceds caused by these loadings are calculated as belw.
Ter seimic and relief valve thrust loading:
\2 ,[g . g2 .
g,2, , \2 M< . [,v(,2s. I a , _g
,,, y x .
1/2
. [ ,I 2 /2 } 2 q,2I .
,2 I
wt ,
\ ggg 4,j_
1-1 l
A-11
I I
i i
I where
,M ,M . deadweight moment components St I M*wt #wt r
t
. inertial CBE moment components '
M*0BE'#0BE' M 3M 08E
^
- ,M ,H . relief line operation moment component:, ,
l M*$0T U
$0T U
$0Tg A.l.4.4 DOCDCT CONDITIONS ,.
The cabined stresses due to primary loadings of pressure, weight and I safety valve thrust, using applicable stress intensification facters, must not exceed the allowatie limits. -The magnitude of the resultant m: ment, Mg , is calculated fra the ament components as shown below:
+ \2 . /M . $2 ./M .
2 u2 .
M, . ,Z i
/M*$0T g M*wt i I Ys0T E
,#wt L l 3 307E - wt
.k / .( / \ -
wh.re l
4 M, . de.6.eight moment com,onent,-
w t-, M,wt, M,wt R #
50T g
$0Tf*# SOTg_
I A-12. ;
. - . . . . _ , . .. . - - .. a
1 1
] i J
l
- A.1.a.5 FA11TED CCNDITIONS 1,
i Tne cecbined stresses due to primary loadings of presstre, weight, SSE and !
a SOT , using applicable stress intensification factors aust -not exceed the j T
4 allcwable limits. For the resttit. ant ament loeding, M 3, the SSE and SOTy i
aments are cambined using thc aquare-root-cf-the-sta-cf-the-squares -(SR$$)
i addition and added absolutely with deadweight for each ament component. .
j (M,, My , M,). The angnitude of the resultant ament, gM , is miculated fra J the three moment components, as shown belcw:
- M / 2 . 2 v2 . 12 ,
I. '
- $0T ,*$3E M"wt i
2 2 U2 , 2 !
}
+ [ + M + M
- wt i $0T Y #SSE j
Q 2 2 1/2 2' 1/2 l
+ [M 50T3
+ M 2
35E
+ M 3 !
l 7 wt'
- .ere 1
I
.M Z = deadveight moment components M*w t.'M #wt wt ,
. inertial SSI imment components
' M,S$t.M#55E , M*$$t
,M . maximum of 50Ty or W moment g- components:
M,30T $0T , M,$0T '
F F g A-13 i
. . . - .- - . - .. . . . ~ . - .- .. , , . . .
- i i _
i 1
l 3 Fcr the safety and relief piping, the faulted condition load cabinatien of yessure, weight, and valve thrust is considered as given im Table A-1 l
~
and A-2 and defined in Table A-3 . The pipe break iceds (MS/INPB or LOCA) can be isscred for the PSARV rystas. These loads have very little is;act on the gressurizer safety and relief systan when empered to the i loading conditions discussed in this report.
A.l.4.6 SICONDARY STRESS EVALt!ATION
- The cambined stresses due to the secondary loadings of thermal, pressure, and deadweight using appliesble stress intensification facters must not exceed the allowable liatt. For the resultant ament.1 ceding, M ,g theM ,
aments are cabinert aa shown below:
2 . .2 , ,y 2 -- U2 M
I. M
-M g# , y# g* MAX h!N MAX NN MAX MIN
,M ,M 3 . anxisus thermal moment considering all thennal' cases X #WX MAX including normal operation.
M ,My .M = minimus tharsal moment considering all thermal cases
- including norsi;1 operation i
This, M, g
is then substituted into the appropriate egostions_ of the ap;0.icable code.
l A-14.
.1 _ 2 . . u _. ._ .a . . _ . . _ _ . - . . _ _ , .
1 TABLE A-1 e
4.0 CCMBI. NATIONS AND Ar%FTANCE CRITERIA UPSTREAM or V ALVES - FOR FF.ESS Le DELif VALVE ?!/TM L@ $UPPOM'S Piping Allowable Plant /Syste Stress intmenity en rme< ra cwitten Lnad crehir.mtien ret _-.+ 4 en
- N 1.5 5, 1 Normal '
Upset N + CSE + SOTy 1.83,/1.5sy 2
Emergency N + SUT g 2.25 S,/ 1.8 sy 3
N + SSE + SCffF 30 5, 4 Faulted NOTES: for SOT definitions and oWer load abbreviations.
(I) See Table A-3 Use SRSS for cabining dynanic iced ruponses.
(2)
The bounding number of valves (and discharge seqwnee if setpoints (3) are significantly different) for the applicable system cprating transient defined in Table A-3 abould be used.
but (4) Verification of functional cambility is not required,-
allowable .loods imd accelerstions for the safety-relier valves
)
must be met.
(5) When (2) two allowable stress intensities, are identified, the less of the (2) .two is applicabic.
1 1
A-15 i
TABLE A.2 LOAD COMBI. NATIONS AND ACCEPTANCE CRITIRIA FOR PRESSURIIER SAFITY AND RII.III VALVE PIPING AND AUPPSPTs sritMICALLY DESIGNED DOWNSTREAM PORTION Plant /Systen Piping A11cwable cm b4 utten Deerattne tenettien L an d cm bi nn tt e r. stress Intensity ,
1 Norinal N 1.0 Sh 2 Upset N + 5079 1.2 (
3 Up.et N . Ost . m U
'85 8 4 Emergency N + SOT l'O b E
s rault.e .
N . sSc . 30Tr 8*S n NOTES:. (1) This t.able is applicable to the seimnically designed portion of dcunstrean non-Category I piping (aM supports) necessary to isolate the. Category I portion fra the non-seim4cally designed piping response, and to assure acceptable valve loading on the disearge nozz.le.
i (2) See Tahle A-3 for SOT definitions and other load abbrwiations.
(3) The bounding number of valves (and discharge sequence if setpoints are significantly different). for the applicable systen operating .
transient defined in Table A-3 should be used.
(4) Verification of functional rarseility is not required, but I allowable loads and accelerations . for the safety-relief valves must be met.
(5) Use 3RSS for cambining dynamic load responses.-
A-16
- e a i
__________.m.______.____.___-.-_m_ _ _ _ _ _ _ _ _
_ -. _ __ m.__ _ _ _ _ _ _ - _ _ _ ~ . _ _ . _ . . - _ _ _ _ _ . _ . _ _ _ _ _ _ . _ . . _ _ _
e
)
} I i l
- i I
I i
i i
I Tag.g A-3 ,
MFTWITTONS OF i nin ARRRFV(M u ,
i
! N s Sustained loads during normal plant. operation 1
! SUT Sywtasoperatingtransient 4 1
Relier valve discharge t~:asient (1). ;
507 l 9.
l Sofg a safety valve disaharge transient (1), (2)
- SOf7 Man (307 g ;M g ); or M ti m C W .
! 08E Otersting basis earthquake i ;
- SSE a Safe shutdown earthquake S a Basic anterial allowohle stress at anxima (het) taperature h
- 1. -
! S : Allwahle design stress intensity a ,
u 1
! (1) May also include trar.sition ' flow, if determined that ~ required operating' procedures could lead to'this condition.
i m Although cern in nur.l.ar- st - ,,1y do.1 n tran.i.nts.(for -
example, loss of.lood) which are alssaitied as upset conditions any actusta the safety valves,' the entrasely law naber of actual-l safety valve'actustions in operating pressurized' water reactors '-
a
- 4 justifies the mergency oondition fram the ASE design philosopny1 and a stress analysis viewpoint.
1 t
[-
4 i
4 t
t 4
k' ..
. ; A.17;;
r,-___......_,...._,,_.__.._.~....,_._.~,-,m....,_...__
i PRESSURIZER SAFETY AND RELIEF LINES (S.P. 2 053)
TABLE A-4 Class 1 (Upstream) Primary Stress Summary Normal Condition COMBINATION- P + DWT Node Piping Maximum Allowable Stress Point Component Equation #9 3.0S.
j Stress (ksi) (ksi) 4390 Butt Weld 9.237 24.225 4031 Long Radius Elbow 9.042 24.225 5000 Branch Connection 9.831 24.225 9010 CRUN 15.165 24.225 4272 5-D Bend 8.148 24.225 4151 Reducer 14.185 24.225 l
TABLE A-5 Class 1 (Upstream) Primary Stress Summary i
Upset Condition COMBINATION: P + DWT y SRSS (OBE. SOTel Node Piping Maximum Allowable Stress Point Component Equation #9 (1.8 S or 1.5 S,)
- Stress (ksi)
Jksi) l 5020 Butt Weld 14.665 26.850 5041 Long Radius Elbo v 17.532 26.850 5000 Branch Connection 18.922 26,850 J
[ 9010 CRUN 15.;73 26.850 4272 5-D Bend 12.537 26.850 l
4151 Reducer 18.997 26.850
.n
[
- The smaller of the given allowable is to be used.
A-18 1
.- , , , , ~ .
PRESSURIZER SAFETY AND RELIEF LINES (S.P. 2-053)
TABLE A-6 Class 1 (Upstream) Primary Stress Summary Emergency Condition COMBINAI.LQN: P + DWT + SOTS Node Piping Maximum Allowable Stress Point C<>mponent Equation #9 (2.25 S. or 1.8 S,)
- Stress (ksi) (ksi) 3151 Butt Weld 16.124 32.220 3151 Long Radius Elbow 22.318 32.220 3150 Braneh Connection 26.538 32.220 3140 CRUN 16.124 32.220 I 4272 5 D Bend 7.307 32.220 4151 Reducer 13.811 32.220
- The :maller of the given allowable is to be used.
TABLE A-7 Class 1 (Upstream) Primary Stress Summary Faulted Condition COMBINATION: P + DWT + SRSS (SSE. SOT,)
- a. .-
Node Piping Maximum ' Allowable Stress Point Component Equation #9 3.0 S.
Stress (ksi) (ksi) l 5020 Eutt Weld 20.170 48.450 1031 Long Rajius Elbow 27.975 42.450 3150 Branch Connection 30.280 48.450 5010 CRUN :0.780 48,450 4272 5 D Bend 18.214 48.450 4151 Reducer - 22,725 48.450 A-19
_ - _ _ _ _ _ _ J
J' PRESSURIZEP. SAFETY AND RELIEF LINES (S.P 2-053)
TABLE A 8 Class NNS (Downstream) Primary Str.ss Summary Normal Condition COMBINATION: P + DWT Node Piping Maximum Allowable Stress !
Point Component Equation #9 1.0 S, j Stress (ksi) (ksi) l 6020 Butt Weld 5.020 15.900 1311 Long Radius Elbow 5.366 15.900 1440 Branch Connection 5.657 15.900 l l
1300 CRUN 4.819 15.900 l TABLE A 9 Class NNS (Downstream) Primary Stress Summary Upset Condition j COMBINATION: P + DWT + SOTy 3
i Node Piping Maximum Allowable Stress Point Component Equation #9 1.2 S.
4 Stress (ksi) (ksi) 4420 Butt Weld 6.005 19.080 1311 Long Radius Elbow 6.059 19.080 1440 Branch Connection 6.711 19.080 4421 CRUN 6.004 19.080 I
A-20 l
l l
l
l i
PRESSURIZER SAFETY AND RELIEF LINES (S.P. 2-053)
TABLE A 10 Class NNS (Downstream) Primary Stress Summary Upset Condition COMBINAT,10N: P + DWT + SRSS (OBE. SOT.1 4
4 Node Piping Maalmum Allowable Stress i
Point Component Equation #9 1.8 Sg Stress (ksi) (ksi) 4420 Butt Weld 16.626 28.620 i 1311 Long Radius Elbow 12.328 28.620 3430 Branch Connection 13.611 28.620 j 4421 CRUN 16.627 28420 l TABLE A-Il Class NNS (Downstream) Primary Stress Summary Emergency Condition j
COMBINATIQN: P + DWT + SOTc l Node Piping Maximum Allowable Stress i
Point Component Equation #9 1.8 S.
St;ess (ksi) (ksi) 2350 Butt Weld 18.313 28.620 1311 Long Radius Elbow 16.788 28.620 1440 Branch Connection 20.553 28.620 4 1250 CRUN 18.379 28.620
_._g.
d A-21
1 e PRESSURIZER SAFETY AND RELIEF LINES (S.P. 2-053)
TABLE A 12 Class NNS (Downstream) Primary Stress Summary j Faulted Condition 4
COMBINATION: P + DWT + SRSS + ( SSE. SOT,)_
i Node Piping Maximum Allowable Stress a
Point Component Equation #9 2.4 S, Stress (ksi) (ksi) 3 4420 Butt Weld 74.305 38.160 1 1311 Long Radius Elbow 21.500 38.160 1440 Branch Connectior 24.007 38.160 i
4421 CRUN 24.302 38.160 i
1 A
i A-22
PRESSURIZER SAFETY AND RELIEF LINES ( S.P. 2453 )
Summary of Tangue Evaluation TABLE A 13 Relief Line Thermal Piping Maximum Maximum Allowable Maximum Section Component Equation 12 Equation 13 Stress 3.0 S. (ksi) Cumulative Stress (ksi) Stress (ksi) Usage Factor Steam CRON 4.5 13.7 48.3 .01 6 Inch Buttweld 4.5 14.8 48.3 042 Buttweld 9.2 25.6 39.6 .30 (nonle)
Elbow 11.1 21.3 48.3 .01 Branch 20.0 35.6 48.3 .067 (6x6x3/4)
Brarch 21.1 45.6 48.3 .I11 (6x6x3 Weldolet)
Steam CRUN 7.8 20.6 48.3 .01 3 Inch Buttweld 7.4 21.2 48.3 .031 Elbow 13,1 31.8 48.3 .01 Reducer 20A 36.0 48.3 .09
! (6x3) l Water CRUN 5.4 18.9' 48.3 .047 Buttweld 11.9 23.8 48.3 .91 (Valve) 7.4 34.6 48.3 .047' l 5D-bu 1
A-23
- - - ..- . - . ~ . _ .. . .. . --
4 4
i PRESSURIZER SAFETY AND RELIEF LINES ( S.P. 2 053 ) '
Sununary of Fatigue Evaluation TABLE A-14 l Safety Line Thern) ! Piping Maximum Maximum Allowable Maximum
. Section Component Equation 12 Equation 13 Stress 3.0 S. (ksi) Cumulative
- Stress (ksi) Stress (ksi) Usage .
- Factor Steam CRUN 11.0 19.5 48.3 .01
-i Buttweld 11.1 20.6 48.3 .01 i
l Buttweld 22.5 36.3 39.6 .30
! (nonle)
J- Elbow 27.0 35.3 48.3 .01 I
Water CRUN 5.4 20.6 48.3 .021 Buttv eld 5.4 21.6 48.3 .06 Buttweld 6.5 35.8 48.3 .98 (flange) i 12.8 37.2 48.3 '
.05 Elbow
- Branch 31.6 48.0 48.3 .99 (6 x 3 x 3/4) f k
i i
4 s
l 1
I i
j A-24 ,
l
TABl.E A-15 Tabulation of Worst-Case Stress vs. Allowable Stress Representative Supports From Figure A-12 i
Worst Standard Component i Worst Stressed
-( Strut or Snubber ) l Structural Component
.; . Support No. Suppon Type Actual lead Allowable Load L.C." Type Actual Allowable L.C. "
RC-2-I l5-415-C76R Struts attached to 35.9 kips 67.5 kips E 2 Interaction = 0.98 1.0 F frame RC-2-115-408-C76K Snubber attached 2.33 kips 15.0 kips U 2 Interaction = 0.96 1.0 F to frame
,a t-Structural Component Tynes: Load Combination Case:
Suppo:t Member F - Faulted 2 - Weld .
E - Emergency 3 - Base Plate / Anchor Bolt U - Upset
, 4 - Richmond Insen/ Threaded Rod 4
4
TABLE A-16 Tabulation of Worst-Case Stress vs. Allowable Stress Representative Supports From Figures A-10 and A-Il Worst Standard Component $ Worst Stressed
( Strut or Snubber ) Structural Component j Support No. Support Type Actual L7ad Allowable Imad L.C. " Type Actual Allowable L.C. "
RC-2-115-419-C66K Snubbers attached to 17.39 kips 21.00 kips F 2 26.M ksi 27.09 ksi F frame I RC-2-115-420-C66R Rigid frame -
g 3 Interaction = 0.89 1.0 F g RC-2-115-423-C66R Rigid frame -
3 Interaction = 0.63 LO F RC-2-115-429-C56K Snubber attached to 2.47 kips 72.45 kips F ,
2 0.283 inches 0.134 inches F(1) frame Structural Component Types: " Load Combination Case: (1) Analysis based on allowable stress requires effective weld size 1 - Support Member F - Faulted of 0.134" ; actual effective weld size 2 - Weld E - Emergency is 0.283".
3 '- Base Plate / Anchor Ikilt U - Upset 4 - Richmond Insert / Threaded Rod
, w . -.w-. , ,
TABLE A-17 Tabulation of Worst-Case Stress vs. A!!owable Stress Representative Supports From Figure A-14 Worst Standard Component ( Worst Stressed
( Strut or Snubber ) 'j Structural Component
( Support No. Support Type Actual Lnad ' Allowable Load LC.** (o Type Actual Allowable LC." -
RC-24)97 404-C66K Snubber attached to 7.46 kips 18.29 kips F 5 2.89 kpi 4.30 kpi F frame T
~
Structural Component Types: " Load Combin;t on i Case:
.1 - Support Member F - F:ulted .
4 2 - Weld E - Emer;ency .
~ 3 - Base Plate / Anchor Bolt U - U:act 4 - Richmond Insert / Threaded Rod . ,
3 - Other: Tube Steel Local Stress n
u- - - - - _ . , - _ _ _ _ _ . - _ _ _ _ . - - _ _ - - _ _ _ _ . - - . _ _ - _ . - . _ - - _
TABLE A-18 Tabulation of Worst-Case Stress vs. Allowable Stress Representative Supports From Figure A-14 Worst Standard Component Worst St tssed
( Strut or Snubber ) Structural Component Support No. Support Type Actual lead Allowable Load LC." Type Actual Allowable L.C. "
RC-2-099-402-C86K Snubber attached to 11.12 kips 12.69 kips F 3 Interaction = 0.49 1.0 F baseplate Structural Component Tvoes: ** Load Combination Case:
h 1 - Support Member F - Faulted 2 - Weld E - Emergency .
3 - Base Plate / Anchor Bolt U - Upset 4 - Richmond InsertAhreaded Rod a-
TABLE A-19 Tabulation of Worst-Case Stress vs. Allowable Stress Representative Supports From Figure A-14 y'
4 Worst Standard Component Worst Stressed
( Strut or Snubber ) Structural Component
. Support No. Support Type Actual Load Allowable lead L.C.** Type Actual Allowable L.C. "
RC-2-101-402-C86K , Snubber attached to 4.66 kips 17.53 kips - F 5 0.75" 0.58" F(1) frame : Thick Plate Thick Plate 0
Structural Component Tyr,es: " Load Combination Case: (1) Ananis based on allowable stress requires piate th'ckness 1 - Support Member F - Faulted of 0.58" ; actual plate thickness 2 - Weld . E - Emergency is 0.75" 3 - Base Plate / Anchor Bolt' U - Up.et 4 - Richmond Insert /Threm!cd Rod -
5 - Other: Load Carrying End Plate 1
.i
TABLE A-20 Tabellion of Worst-Case Stress vs. Allowable Stress Representative Supports From Figure A-13 Worst Standard Compone . Worst Stressed
( Strut or Snubber ) Structural Component Support No. Supix>rt Type Actual Load Allowable Load L.C. " ii _ Type Actual Allowable L.C. "
r RC-2-Il0401-C86K Snubber attached to 2.63 kips- 5.36 kips F 3 Interaction = 0.77 1.0 F base plate RC-2-112-403-C86K Snubber attached to- 3.77 kips 4.76 kips 4 Interaction = 0.94 1.0 F
, baseplate j
.g-Structural Component Tyns: ** Load Combination Case:
I - Support Member F - Faulted 2 - Weld . E - Emergency .
3 - Base Plate /Anchcr Bcc I' - Upset 4 - Richmond InsertfIlireaded RW1
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. Figure A.4 COMPARISON OF THE EPRI FnRCE TfwE-HIMT9RY: FoR WE12 AMD WE33 FROM TEST 908 WITH THE-ITCHVALVE PREDICTED
. FORCE TIW -HISTnRY-A-34 g.
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! Figure A-5.. : Comparison of the EPRI Force Tiene-History for WE32 and WE33 fran Test 917 with the ITCHVALVE Prodicted Force Tilne-H1stary l
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l Figurn A-6 Camperison of- the EPRI Force Time-History. For WE32 -
t L and WE33 From Test 908 With the FIXFM3 PrMitted i
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Figure A-7_ Comparison of the EPRI Force Time History For WE32-4 -
and WE33 From Test 917 With the.FIXFM3 Predicted Force Time-History. t A-37 l
.I FIGURE A-8 SAFETY Lit 1E SECTIOf4 - LUtiPED f1 ASS LOCATIOris i Af4D APPLICATIO." P0lf1TS OF FLUID FORCES S.P. 2-053 79 Il f/
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