ML20214S039
| ML20214S039 | |
| Person / Time | |
|---|---|
| Site: | Summer  |
| Issue date: | 03/31/1987 |
| From: | Case G, Harris B, Larson J EG&G IDAHO, INC. |
| To: | NRC |
| Shared Package | |
| ML20214S016 | List: |
| References | |
| CON-FIN-A-6492, RTR-NUREG-0737, RTR-NUREG-737, TASK-2.D.1, TASK-TM EGG-NTA-7639, NUDOCS 8706090104 | |
| Download: ML20214S039 (93) | |
Text
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EGG-NTA-7639 March 1987 i
INFORMAL REPORT 4
Idaho':
National ANALYSIS OF PRESSURIZER SAFETY VALVE DISCHARGE Engineering --
PIPING OF THE V.C. SUMMER NUCLEAR POWER STATION, l.aboratory UNIT 1
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JManaged by the U.S.
Department G. S. Case ofEnergy B. L. Harris J. R. Larson i
G. K. Miller l
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Prepared far the work pertonned und,,
um ocJfl#l78 U.S. NUCLEAR REGULATORY COMMISSION 8706090104 870325 '
PDR ADOCK 05000395 P
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i DISCLAIMER This book was prepared as an account of work sponsored by an agency of the United States Government. Neither the United States Government nor any agency thereof, nor any of their employees, makes any warranty, express or implied, or assumes any legal liabelity or responsibihty for the accuracy, completeness, or usefulness of any information, apparatus, product or process disclosed, or represents that its use would not infnnge pnvately owned rights. Referer<es herein to any specific commercial product, process, or service by trade name, trademark, manufacturer, or otherwise, does not necessanly constitute or imply its endorsement, recommendation, or favonng by the United States Government or any agency thereof. The views and opinions of authors expressed herein do not necessarily state or reflect those of the United States Government or any agency thereof.
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ANALYSIS OF PRESSURIZER SAFETY VALVE DISCHARGE PIPING OF THE V. C. SUMMER NUCLEAR POWER STATION, UNIT 1 G. S. Case B. L. Harris J. R. Larson G. K. Miller a
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ABSTRACT Based on the requirements of NUREG 0737,Section II.O.1, " Performance Testing of BWR and PWR Relief and Safety Valves," all operating plant licensees are required to demonstrate the structural integrity of their as-built pressurizer safety valve / power operated relief valve piping systems. Thus, the licensee for the V. C. Summer Nuclear Power Station, Unit 1, among other licensees, performed such an analysis and submitted the results to the Nuclear Regulatory Commission.
This report describes an analysis by EG&G Idaho, Inc. of the V. C. Summer pressurizer safety valve inlet and discharge piping for fluid loads caused by safety valve actuation.
The analysis was intended to confirm results obtained by the licensee concerning adequacy of the piping system for fluid discharge loads. A thermal-hydraulic analysis was first performed to generate time histories for the fluid forces acting on the system. A structural analysis was then performed on the safety valve piping to calculate its response to these fluid forces.
Results from the analysis performed supported the licensee's conclusion that the system design is adequate for hot loop seal and steam discharge through the safety valves.
1 An analysis of piping loads caused by a cold icop seal and steam discharge was also performed.
Results of the cold loop seal discharge analysis showed several pipe elements were significantly overstressed.
The fluid forces developed during this analysis, in conjunction with the structural model, were also used to evaluate the accuracy of several simplified approaches used by various licensees in their analyses of safety valve piping systems.
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SUMMARY
An analysis was performed to assess the adequacy of the pressurizer safety valve discharge piping located at V. C. Summer Nuclear Power Station, Unit 1, for actuations of the safety valves.
The piping analyzed consisted of the inlet piping to three safety valves, the piping from each safety valve to a common header, and the piping from the header to the pressurizer relief tank.
Stress results from the analyses were compared to limits specified by the American Society of Mechanical Engineers (ASME)
Boiler and Pressure Vessel Code,1983 Edition.
Results from the analysis were also used to evaluate conclusions reached by the licensee regarding l
the adequacy of this piping system.
The analysis was performed by EG&G Idaho, Inc., for the U.S. Nuclear Regulatory Commission under FIN No. A6492.
The analysis described herein was focused exclusively on safety valve discharge loading, and the resulting stresses were compared to ASME Code Section III Level C Service Limits. As such, this does not constitute a complete safety evaluation of the piping system since that would also require consideration for normal sustained loads, operating basis earthquake, safe shutdown earthquake, and discharges through the power operated relief valves.
Because of the significance of safety valve discharge loading, however, the magnitude of stresses calculated in this I
analysis serve as a useful indicator as to whether the system is properly designed.
4 A three-step approach was used to evaluate the adequacy of the system for a safety valve discharge event.
First, a time history thermal-hydraulic calculation of the fluid mass flow and hydrodynamic state of the system during safety valve discharge was performed. The thermal-hydraulic results were then used to calculate the fluid forces exerted on the piping.
Finally, a structural analysis program was used to iii r
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determine the dynamic response of the piping.
The calculated stresses were then compared to allowable values to assess the adequacy of the system.
The RELAPS/M002 program was used to perform two thermal-hydraulic simulation of the loop seal discharge through the safety valves.
Discharges of both a hot and a cold loop seal liquid were simulated.
Fluid forces were calculated using a program named BLAZER, which transformed RELAP5/M002 output into forces that could be input into the structural analysis program. The SAP IV program was then used to perform a time history solution for the structural response to the fluid forces.
The hot loop seal analysis included explicit modeling of the heated loop seal liquid immediately upstream of the safety valves. A valve model which simulated the simmer and pop behavior of full-scale safety valves was used. The peak stresses in the system were caused by the loop seal liquid being forced through the discharge piping at a high velocity.
Results of the analysis, which is considered a best estimate calculation cf the piping i
response to a valve actuation event in the V. C. Summer plant, showed pipe stresses to be well within allowable values.
The analysis performed by the licensee on this system showed the system to be adequate for safety valve actuations. The analysis presented
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in this report supports the licensee's conclusion, but does not necessarily verify the licensee's analysis methodology.
Values of stresses calculated by the licensee for the system were not available for comparison with the results presented herein. Also, the discussion of modeling assumptions included in the licensee's submittal was incomplete.
Confirmation of the licensee's analysis method requires the availability of details of the licensee's assumptions and results.
Because the V. C. Summer plant contains hot loop seals, the analysis of a cold loop seal discharge is not necessary for determining the adequacy 4
of the V. C. Summer plant's piping.
Since some power stations do employ iv
cold loop seals, an analysis of a cold loop seal discharge was performed to determine the expected severity of piping stresses in such plants.
The cold loop seal liquid was modeled in accordance with available guidelines.
These guidelines call for placing the loop seal liquid downstream of the valve after isenthalpic expansion to atmospheric pressure. The loop seal liquid was placed in a few volumes immediately downstream of the valves at a void fraction near one-half. The analysis of safety valve discharge with a cold loop seal calculated overstresses in several pipe elements, partitilarly in pipe bends where stress indices are high.
The structural model and fluid forces from this analysis of the V. C. Summer safety valve piping system were also used to evaluate certain simplified approaches used by various utilities in their analyses of the safety valve piping system.
The simplified approaches evaluated were: (a) neglecting the effect on the piping system response of axial stretching forces in each pipe segment, (b) modeling the snubbers as rigid supports rather than modeling the flexibility of the snubbers, and (c) approximating dynamic results by performing a static analysis in which the peak net force on each straight pipe segment was multiplied by a dynamic load factor of 2.
Results of the evaluation indicate that (a) neglecting axial extension in the pipe segments caused by opposing forces in the segments had a negligible effect on the calculated piping response, (b) modeling the snubbers as rigid supports rather than incorporating the flexibility of the snubbers in the analysis yielded significantly nonconservative results in portions of the piping system, and (c) performing a static analysis in which the dynamic forces were amplified by a dynamic load factor of 2 also proved to be nonconservative for a portion of the calculated responses.
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i CONTENTS ABSTRACT..............................................................
11 h i
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SUMMARY
iii 1.
INTRODUCTION.....................................................
1 t
2.
DESCRIPTION OF PIPING CONFIGURATION..............................
4 3.
DESCRIPTION OF TYPICAL LOOP SEAL DISCHARGE.......................
7 4.
THERMAL-HYORAULIC CALCULATIONS...................................
8 3
i 4.2 Compute r Code De sc ri p ti on...................................
8 l
4.2 Noda11zation................................................
8 4.3 Boundary Conditions.........................................
15 4.4 Modeling Options............................................
16 i
4.4.1 Requested Time Step...............................
16 4.4.2 Choking Model.....................................
16 j
4.4.3 Flow Resistances..................................
17 4.5 Hot Loop Seal Simulation....................................
17 j
4.6 Cold Loop Seal Simulation...................................
18 1
1 5.
ANALYTICAL MODELING METHOD.......................................
21 E
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5.1 Structural Model............................................
21
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5.2 Piping Supports.............................................
22 i
j 6.
PRESENTATION OF ANALYTICAL RESULTS..............................
23 i
7.
THERMAL-HYDRAULIC EXPLANATION FOR DIFFERENCES IN l
STRESS RESULTS..................................................
33 1
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9.
OTHER ANALYSIS TOPICS............................................
37 8.1 Including Axial Stretching Forces...........................
37 i
8.2 Modeling Snubbers as Rigid Supports.........................
39 i
i 8.3 Approximating a Dynamic Analysis with a Static Analysis
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and a Dynamic Load Factor...................................
43 9.
CONCLUSIONS AND RECOMMENDATIONS..................................
45 I
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REFERENCES.......................................................
48 APPENDIX A--METHOD OF DYNAMIC ANALYSIS................................ A-i APPENDIX B--EFFECT OF N00ALIZATION ON RELAP5/ MOD 2 CALCULATIONS OF VOID FRACTION DURIN9. SAFETY VALVE DISCHARGE............ B-1 FIGUF,ES 1.
Schematic of V. C. Summer overpressure protection system.........
5 2.
Isometric drawing of safety valve discharge piping...............
6 3.
RELAPS model of piping upstream of the safety valves.............
9 4.
RELAPS model of piping between valve "A" and header..............
10 5.
RELAPS model of piping between valve "B" and header..............
11 l
6.
RELAP5 model of piping between valve "C" and header..............
12 7.
RELAP5 model of header...........................................
13 8.
RELAPS model of piping between header and relief tank............
14 9.
RELAPS calculation of volume vapor void fraction downstream of safety valve..................................................
34 10.
Fl uid wave passi ng through two elbows............................
38 TABLES 1.
B and 8 stress indices for elbows..............................
24 1
2 j
2.
Stresses in selected straight pipe segments for hot loop seal analysis....................................................
26 1
3.
Stresses in selected pipe elbows for hot loop seal analysis......
27 i
4.
Forces in selected snubbers for hot loop seal analysis...........
28 5.
Stresses in selected straight pipe segments for cold loop seal analysis....................................................
30
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6.
Stresses in selected pipe elbows for cold loop seal analysis.....
31 7.
Forces in selected snubbers for cold loop seal analysis..........
32 vit
10 8.
Stresses in selected straight pipe segments for 10 lb/in.
snubber stiffness...............................................
40 10 9.
Stresses in selected pipe elbows for 10 lb/in. snubber stiffness........................................................
41 10 10.
Forces in selected snubbers for 10 lb/in. snubber stiffness........................................................
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ANALYSIS OF PRESSURIZER SAFETY VALVE DISCHARGE PIPING OF THE V. C. SUMMER NUCLEAR POWER STATION, UNIT 1 1.
INTRODUCTION An analysis of the pressurizer safety valve piping in the V. C. Summer Nuclear Power Station, Unit 1, for hydrodynamic loads caused by safety valve actuation was performed to assess the adequacy of the piping for such loading.
Results from the analysis were, to the extent possible, used to evaluate conclusions reached by South Carolina Electric and Gas Company regarding the adequacy of this piping system. The V. C. Summer Plant is a Westinghouse three-loop pressurized water reactor (PWR).
Information from the study will be used by the U.S. Nuclear Regulatory Commission (NRC) in evaluating safety valve piping of similar PWR plants. The work was performed by EG&G Idaho, Inc., for the NRC under FIN No. A6492.
According to NUREG 0737,Section II.D.1, " Performance Testing of BWR and PWR Relief and Safety Valves," all operating plant licensees are required to demonstrate the structural integrity of as-built safety and relief valve piping systems on a plant specific basis.
Thus, the licensee for the V. C. Summer plant, among other licensees, performed such an analysis and submitted the results to the Nuclear Regulatory Commission.1 The analysis described in this report was intended to evaluate results obtained by the licensee concerning adequacy of the piping system for fluid discharge loads.
The present analysis was focused exclusively on safety valve discharge loading. As such, this analysis alone does not constitute a complete safety evaluation of the piping system per the NUREG 0737,Section II.D.1, program. That would require consideration of load combinations involving
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normal sustained loads, operating basis earthquake, safe shutdown earthquake, and discharges through the power operated relief valves.
Because of the significance of safety valve discharge loads, however, results from this analysis serve as a useful indicator as to design adequacy of the system.
1
A three-step approach was used to evaluate the adequacy of the piping system. First, a time history thermal-hydraulic calculation of the fluid mass flow and hydrodynamic state during safety valve discharge was performed.
Results from the thermal-hydraulic calculation were then used to calculate the fluid forces exerted on the piping.
Finally, a structural analysis program was used to calculate the response of the piping to the fluid forces by solving the governing equations of motion for the system.
The analysis included calculation of stresses in the piping that were compared to allowable stresses so as to determine whether the piping is adequate from a safety standpoint.
The thermal-hydraulic calculation was performed with the RELAP5/M002 computer code.2 The code calculated parameters of fluid flow, such as temperature, pressure, void fraction, density, velocity, etc., versus time.
Two distinct simulations of fluid flow were performed using RELAP5/M002. One simulation modeled valve discharge with heated loop seals upstream of the safety valves; the other modeled cold loop seals. The hot loop seal simulation is representative of a safety valve actuation event in the V. C. Summer plant.
The cold loop seal simulation is expected to be characteristic of plants employing cold loop seals upstream of the valves.
Fluid forces exerted on the safety valve piping during valve discharge were determined from the thermal-hydraulic calculations using a computer program called BLAZER.3 The program read fluid data from a RELAPS/ MOD 2 output tape, computed force-time histories at specified points on the piping, and generated the force histories in a format that could be input directly into the <tructural analysis code. The forces were then applied to the finite element model of the piping.
Response of the safety valve piping system to the induced fluid 4
loading was calculated using the SAP IV structural analysis program.
The system was represented by a finite element model consisting of lumped masses connected by massless elements.
Fluid loads were applied to the model in the form of force-time histories and the program generated the time history response of the model by performing a direct step-by-step integration on the set of linear equations describing the motion of the j
system.
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The safety valve piping downstream of the valves was assessed according to requirements of the 1983 American Society of Mechanical Engineers (ASME) Code Section III for Class 2 components.5 Piping upstream of the safety valves was evaluated according to the ASME Code Section III for Class 1 components.
The calculated stresses were compared j
to Level C Service Limits, which correspond to an emergency condition.
Consideration of a faulted condition would require that safety valve discharge loads be combined with safe shutdown earthquake loading. The stresses for this condition, were it considered in this analysis, would be compared to higher Level D Service Limits.
Certain approximations have been made by various utilities in the i
I analyses of the safety valve piping systems in their plants. The use of these simplified approaches raised questions as to the adequacy of these analyses. The structural model and fluid forces developed for the V. C. Summer analysis were also used in separate analyses to evaluate the use of these approximations. They were (a) neglecting the effect on the dynamic analysis of axial stretching forces in each pipe segment, (b) modeling the snubbers as rigid rather than flexible supports, and (c) approximating dynamic results by performing a static analysis using the peak loads multiplied by a dynamic load factor of 2.
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2.
DESCRIPTION OF PIPING CONFIGURATION The overpressure protection system in the V. C. Summer Nuclear Power Station, Unit 1, consists of three safety valves, three power operated relief valves (PORVs), three PORV block valves, and all interconnecting piping extending from the pressurizer tank to a relief tank downstream. A schematic showing the overpressure protection system is presented in Figure 1.
The analyses described in this report were focused on actuations of the three safety valves.
The safety valves and their associated discharge piping are shown in an isometric in Figure 2.
The safety valves are Crosby model HP-BP-86 6M6 valves. The inlet lines to the safety valves begin at connections to the pressurizer tank and have loop seals containing liquid that is maintained at an average temperature of 380 F.
The discharge lines from the three valves meet at a common header downstream, which then leads to the pressurizer relief tank.
The safety valve discharge piping is supported by a total of 36 snubbers.
The piping is 6-in. Sched ile 160 between the pressurizer and safety valves, 6-in. Schedule 40 pipe between the safety valves and header, and is 12-in.
Schedule 40 S between the header junction and relief tank.
Connected to the safety valve discharge header is the discharge line from the three PORVs. A portion of this line, which contains sections of 6-in.
Schedule 40 and 80X, was included in these analyses to account for its effect on the safety valve piping response.
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N Block valve Poweroperated relief valve 6" x 3" reducer Pressurizer 7 3010 Figure 1.
Schematic of V. C. Summer overpressure protection system.
Safety valve "C" i
To pressurizer Safety valve "B" To pressurizer N
N PORV' attachm\\
line ent Safety valve "A" To pressurizer 7
Pressurizer relief tank 7 3005 Figure 2.
Isometric d. awing of safety valve discharge piping.
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3.
DESCRIPTION OF TYPICAL LOOP SEAL DISCHARGE This section describes safety valve behavior during loop seal discharge. A loop seal discharge event was chosen for this analysis because, from a safety standpoint, loop seal discharge is more important than steam discharge. A loop seal discharge will exert much higher forces on the piping than a steam discharge.
Oscillatory behavior is observed when a safety valve with an upstream liquid loop seal initially opens.
The valve typically oscillates between 0 and 20% of its full cpen position for about a second. During this second, much of the liquid initially contained in the loop seal is transported downstream of the valve. This first second of loop seal discharge is termed " valve simmer." At the end of the valve simmer period, the valve pops to the full open position. Valve pop takes a few milliseconds. Because the loop seal liquid causes the greatest piping stresses as it is propelled through the discharge piping, accurate modeling of the liquid behavior during valve simmer and pop is necessary.
The loop seals upstream of the safety valves in the V. C. Summer plant contain liquid water maintained at an average temperature of 330 F.
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THERMAL-HYDRAULIC CALCULATIONS i
Two separate thermal-hydraulic simulations of relief valve discharge 1
were performed. One simulation, a best estimate calculation of safety valve ' discharge in the V. C. Summer plant, ' calculated discharge with hot loop seals placed upstream of the valves. The other simulation used the V. C. Summer piping configuration to characterize valve discharge events in plants that employ cold loop seals upstream of the safety valves.
This section discusses the analytic techniques used to calculate the behavior of the fluid within the safety valve discharge piping system. The computer code, system nodalization, boundary conditions, and modeling options used in both simulations are described.
The safety valve modeling and initial conditions are presented for each simulation.
4.1 Computer Code Description 1
The RELAPS/ MOD 2 computer code was used to calculate fluid behavior during safety valve discharge.
RELAP5/M002 is an advanced, one-dimensionai computer program for the analysis of reactor systems.
It is based on a nonhomogeneous, nonequilibrium hydrodynamic fluid model.
The fluid is described by a continuity, momentum, and energy equation for each phase.
Choked flow is calculated by a correlation.
Correlations for wall friction, form losses, interphase friction and mass transfer, and wall heat transfer to the mixture are also included in the code. Air in the piping is treated thermodynamically.
The numeric solution technique is semi-implicit.
Cycle 36.04 of RELAp5/M002 was used for the hot loop seal simulation; cycle 20 was used for the cold loop seal simulation.
4.2 Nodalization This section presents the nodalization used with both RELAP5/M002 simulations.
The nodalization of the portion of the overpressure protection system connected to the safety valves is found in Figures 3 through 8.
The 8
V Number of equally spaced volume cells j-Length (f t)
(6) f afety valve (3) 0.7334 r__
2.76f1 (8)
(8) 2.2971 Pressurizer (
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model (6)
Time dependent 1.441
, Vertical volume q__
r Horizontal 7 3016 Figure 3.
RELAP5 model of piping upstream of the safety valves.
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Safety valve A 1.777 Horizontal l
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Numberof equally j
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&(6) 45' (10) 90*
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b 7-3015 Figure 4.
RELAP5 model of piping between valve "A" and header.
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Safety (7) valve B 1.777 Horizontal 7.839 (10)
Numberof equally spaced volume cells r(8) 45' (8)
Length (ft) 1.256 (1)
- 6" x 12" reducer 0.818 (2) 73014 Figure 5.
RELAPS model of piping between valve "B" and header.
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Safety a Vertical valve C 1.771 Horizontal 12.351 (10)
Numberof equally spaced volume cells r (8)
Length (ft)
(8) 3.4665 45*
7 3013 Figure 6.
RELAP5 model of piping between valve "C" and header.
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geometric configuration was derived from drawings furnished to the NRC.
The drawings are contained in Reference 6.
As the power operated relief valves were not opened for the calculations presented herein, the valves and associated piping were excluded from the codel.
To include the additional piping would have exceeded the storage capability of the CDC 176 computer. Also, the effect of the omitted piping on the resultant hydraulic behavior of the remaining system was felt to be negligible.
The figures show the orientation of the model, the location of elbows and piping' junctions, the length of piping segments, and the number of volume cells comprising the segments.
The flow paths from the pressurizer to the three safety valves are identical and thus shown only once in Figure 3.
As the flow paths from each valve to the common discharge header are different, they are individually shown in Figures 4, 5, and 6.
Figure 7 shows the common header junctions and Figure 8 shows the exit piping connecting the common header to the pressurizer relief tank.
The last two piping segments shown in Figure 8 are approximate representations of exit piping located inside the pressurizer relief tank as design information was not available.
4.3 Boundary Conditions For both simulations, the pressurizer contained saturated steam only.
The pressurizer pressure was linearly increased from 2376 psia to 2592 psia in the first second of the transient and held constant at 2592 psia thereafter. These pressurizer boundary conditions are consistent with those listed in Reference 7.
The valves opened when the pressurizer pressure was 2500 psia.
The dischirge piping contained water vapor at atmospheric pressure.
The segment of,9scharge piping inside the pressurizer relief tank was liquid filled to represent the actual pipe condition. The relief tank was modeled with a time dependent volume in which atmospheric conditions were imposed.
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Heat transfer from the fluid to the piping wall was not modeled.
This assumption simplified the model considerably.
Ignoring heat transfer to the wall results in calculated hydraulic forces that are somewhat conservative.8 4.4 Modeling Options Modeling choices made for both simulations which deserve special mention involve the maximum time step requested, the use of the choking model, and the representation of flow resistances at pipe fittings.
4.4.1' Requested Time Step The maximum time step is set by the criteria:
at < ax/cn where ax is representative of the smallest volume length in piping legs for which the calculation of forces is desired, e is the expected sonic velocity, and n is an integer.
It is anticipated that the shockwave downstream of the valve could travel at twice the sonic speed. To limit a pressure or mass gradient front (shockwave) from passing through a volume cell in one time step, it is recommended that n = 2.
A representative value of the sonic velocity for vapor at atmospheric pressure in the downstream piping is about 1400 ft/s.
For two phase flow (after a liquid slog has broken up) the sonic velocity in the two phase fluid is about
-4 500'ft/s. A maximum time step of 10 seconds was specified for both simulations.
4.4.2 Choking Model The choking model was applied upstream of the valve, at the orifice area representing the valve, and at the exit junction.
The model was not applied in the downstream piping.
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4.4.3 Flow Resistances 9
Flow resistance coefficients were taken from Crane for elbows.
For 10 the angled tees, values were estimated from the RPL handbook assuming
. flow in all branches. The resistance coefficients were converted to pipe roughnesses and applied over an arbitrary number of volume cells at each elbow and tee junction. The conversion was required for BLAZER's use of RELAP5/M002 output to calculate fluid forces.
4.5 Hot Loop Seal Simulation The hot loop seal simulation was performed as the first step in the analysis to determine the adequacy of the V. C. Summer plant safety valve pipin'g system during valve discharge. Modeling of the loop seal liquid was consistent with a method shown to produce good agreement with measured forces. A valve model which simulated valve simmer and pop behavior was employed.
In this simulation, the hot loop seal liquid was placed upstream of the valve.
Placing the loop seal liquid upstream of the valve accurately represents the fluid's actual initial condition. A linear temperature profile rising from 220 F at the valve to 540 F (near saturation) at the other end of the loop seal was used for the initial liquid condition. This temperature profile results in an average fluid temperature of 380 F.
Modeling of the cold loop seal discharge necessitated placing some vertical pipe sections in a horizontal position (see Section 4.6).
To maintain consistency between the two simulations, the affected sections of pipe wc-also horizontal in the hot loop seal simulation. The omission of gravitational effects in the affected sections was expected to have negligible effect on the calculated hydraulic behavior and resultant hydraulic forces.
The RELAP5/M002 servo valve model was used to mimic safety valve simmer and pop behavior in this simulation. The valve was cracked open at.
the typi ci valve set pressure (2500 psia).
The valve area was held i
17
constant during the simmer period; no attempt was made to mimic the oscillating behavior of a real valve. After most of the loop seal liquid passed through the valve, the valve was popped fully open.
The safety valve full flow area was determined using test 11 information indicating a steam flow rate of 441,100 lbm/hr with an upstream tank pressure of 2575 psia. The model of the overpressure protection system was exercised at steady state with the pressurizer pressure fixed at 2575.
The valve flow area was varied until the correct steam flow rate was achieved. This flow area was then used in the thermal-hydraulic analysis for the V. C. Summer plant.
The servo valve controls were adjusted to best mimic the measured valve mass flow, upstream pressure, and downstream pressure recorded in Electrical Power Research Institute (EPRI) valve discharge experiments.12 Several RELAPS/ MOD 2 runs using a model of the EPRI nest facility were performed in adjusting the servo valve controls. The valve was cracked open at a specified valve inlet pressure.
The valve was popped open on a value of upstream void fraction that produced the best agreement with the measured parameters. The valve crack area was then adjusted so that the time from crack to pop equaled the measured simmer time.
The servo valve model was able to match the data quite well.
Once the servo valve was adjusted to produce good agreement with the data, the model and its controls were used for each valve in the RELAP5/M002 input deck of the V. C. Summer plant. All three safety valves were set to open at 2500 psia in this simulation.
However, separate controls regulated the behavior of each valve.
4.6 Cold Loop Seal Simulation 3
The cold loop seal simulation indicates the severity of a valve actuation event in a safety valve piping system which employs cold loop seals. A cold loop seal discharge exerts much greater forces on the piping system than a hot loop seal discharge because less of the liquid flashes to steam in the cold loop seal case.
18
During the simmer period, much of the liquid initially contained.in j
the loop seal is transported downstream of the valve.
For this simulation, valve simmar was modeled following guidelines provided in Reference 8 for modeling cold loop seal discharges. The guidelines prescribe modeling the fluid conditions that exist downstream of the valve at the instant of valve pop, and employing a valve model that simply pops open from a fully closed position.
In Reference 8, this modeling approach provided a thermal-hydraulic calculation that was used to produce a reasonable prediction of measured piping loads in a full-scale experimental facility.
To model the actual fluid condition at valve pop, the loop seal liquid was distributed uniformly (after expansion to atmospheric pressure in a constant enthalpy process) in the first two piping legs downstream of the valve in each flow path. This process resulted in a void fraction near one-half for the piping legs.
The remaining piping was assumed to be filled with water vapor at atmospheric pressure.
l To keep the liquid distributed downstream in position prior to valve pop, the vertical piping sections from the valves to the common discharge header were placed in a horizontal position.
The omission of gravitational
{
effects was expected to have negligible effect on the hydraulic-behavior f
and resultant hydraulic forces.
Valve pop was modeled as a linear increase from zero to full flow area I
in 7 ms. The safety valve full flow area was the same area used in the hot loop seal simulation.
To maximize the hydraulic forces on the downstream piping, the opening times of Valves B and C were slightly delayed to cause an approximate superposition of the mass flow peak from each path as the flow through the individual pipes rea:hed the junction points in the common downstream header.
The delay time coupled with the pressure ramp rate resulted in Valve C opening at about a 6 psi higher pressure than Valve A, a value well i
within the uncertainty of the opening pressure.
l 19
T The maximum forces calculated by BLAZER occurred in the common header when the three mass flow peaks converged.
However, the SAP IV structural analysis calculated maximum stresses for-the piping system upstream of the common header before any discharge fluid reached the header. Thus, it was not necessary to adjust valve opening times to identify piping that is most
' likely to fail during safety valve discharge.
)
i 1
i 3
1 i
1 5
e a
4 20
5.
ANALYTICAL MODELING METHOD The safety valve piping was analyzed by developing a finite element model for use in the SAP IV structural analysis code.
The following description of the method used for modeling the piping system includes discussion of the structural model and treatment of the piping supports.
The discussion in this section is limited to the piping model used to analyze the system. A detailed explanation of the methods used to perform the structural dynamic analysis, which includes the calculation of fluid forcing functions and the time history solution for piping response to the forcing functions, is contained in Appendix A.
5.1 Structural Model The finite element model represents the actual system as an assemblage of massless clastic members connected at discrete nodal points with the system mass concentrated at the nodal points. A sufficient number of nodal points were used so that the discretized model would respond to the hydrodynamic loading in the same manner as the actual continuous system.
This was accomplished by placing nodes at intervals close enough that all natural frequencies of vibration in the lateral directions that can be excited by the hydrodynamic loading on the piping (approximately 100 Hz or less) were included in the structural response calculations.
The model consisted of 164 pipe and 36 boundary elements.
Pipe elements were used to represent the safety valve piping, and boundary elements were used to represent snubbers.
The model extended from attachment points to the pressurizer tank at the end of each safety valve inlet pipe to an attachment point at the pressurizer relief tank. A portion of the PORV piping was included in the model to include its effect on the safety valve piping.
The attachment points to the pressurizer tank and relief tank were assumed to be fixed in this analysis since the details on the attachments were unavailable. The three safety valves were treated as concentrated 21
weights of 950 lb each with the center of gravity located 30 in above the pipe centerline. The actuator was modeled as a rigid beam.
The stresses calculated using the 30-in, distance to the center of gravity of the valve were considered conservative. This distance was not known at the time the model was developed.
Reducing this distance would reduce stresses in the piping near the valve but would have little effect at more distant locations, such as near the header.
5.2 Piping Supports The safety valve piping is supported at several locations along its length.
In addition to the anchor points mentioned in the previous section, there are 36 one-directional snubbers. The type of snubbers actually used in the V. C. Summer plant were not known by EG&G at the time this analysis was performed.
Therefore, the stiffness value used to model the snubbers in the analysis was based on stiffness data for typical varieties of Pacific Scientific snubbers, which are representative for this type of installation'. A stiffness of 100,000 lb/in, was chosen because three common size Pacific Scientific snubbers have stiffnesses on this order.
They are the PSA-1, PSA-3, and PSA-10 with stiffnesses of 60,000 lb/in., 75,000 lb/in. and 240,000 lb/in., respectively. The next larger snubber, the PSA-35, is not as commonly used as those already mentioned.
It has a stiffness of 700,000 lb/in.
e 22
6.
PRESENTATION OF ANALYTICAL RESULTS The analytical results include calculated stresses at several critical locations on the piping and loads on the snubbers.
Stresses in the piping were calculated from SAP IV output according to Equation (9) of Paragraph NC-3653.1,Section III, of the 1983 ASME Boiler 5
and Pressure Vessel Code as follows:
P 0
M max, + B B
SQL = B1 2t 2 Z n
where S
stress due to hydrodynamic loads
=
QL P
peak pressure, psi
=
max D
outside diameter of pipe, in.
=
g n minal wall thickness, in.
t
=
n M
resultant moment on cross-section due to relief valve
=
B loads section modulus of pipe, in.3 Z
=
B,B2=
primary stress indices for product under investigation y
(0.5 an 1.0 for straight pipe and various values for d
elbows per Table 1).
23
TABLE 1.
B AND 8 STRESS INDICES FOR ELBOWS 1
2 Bend Radius 0
0 Pipe Size (in.)
1 2
6-in. Schedule 160 9
0.196 1.588 6-in. Schedule 40 9
0.0 3.275
~
12-in. Schedule 40S 18 0.0 4.123 6-in. Schedule 80X 30 0.441 1.064
(
24
The values of B shown in Table 1 for 6-in. Schedule 40 and 12-in.
2 Schedule 405 pipe are quite large in comparison to the stress index on the moment term used in the 1980 version of ASME Section III, Class 2.
If the 1980 Code were used, the stress indices would be reduced from 3.275 to 2.46 i
for 6-in. Schedule 40 pipe and from 4.123 to 2.32 for 12-in. Schedule 40S pipe. However, the Level C Service Limit would also be reduced from 2.2S to 1.2S. Therefore, it is not readily apparent as to which Code h
h version gives the most conservative results.
The pressure that causes safety valve opening, approximately 2400 psi, was used for P in the inlet piping.
P was neglected for the max max piping downstream of the safety valves since the downstream pressure was much smaller. The conclusions reached later in this report concerning piping adequacy are not affected by addition of the P term in the max stress equation.
I The maximum values for bending moments were taken from the SAP IV computer output to calculate M f r use in Equation (9). The M term, B
B which is the resultant of the two bending moments and the torsional moment on the cross-section, was found from the maximum values of each of the three moment components at a particular location without regard to the time when each of the moments achieved its maximum value. This procedure was conservative, but not overly conservative, because one moment component was l
generally much larger than the other two and, therefore, dominated the
(
resultant.
The hot loop seal analysis performed is a best estimate calculation of piping stresses expected during safety valve discharge.In the V. C. Summer plant.
For this case, the stresses in the most highly stressed straight segments and pipe elbows are shown in Tables 2 and 3.
The stresses for all elements in the piping system were well within the allowable stresses.
The maximum forces in the snubbers for the hot loop seal analysis are shown in Table 4.
These forces are well within the allowable force of 54,600 lb for a PSA-35 snubber.
25 7,
e e,
=.
TABLE 2.
STRESSES IN SELECTED STRAIGHT PIPE SEGMENTS FOR HOT LOOP SEAL ANALYSIS 8
P,,x o Total Level C D
1 Equation (9)
Allowable 2t" B2 "B Stress Stress Element (See Note)
Z (psi)
(psi) 1 5535 1067 6602 28,050 44 0
468 468 36,450 45 0
466 466 36,450 78 0
1805 1805 36,450 115 5535 1240 6775 28,050 116 0
1587 1587 36,450 145 5535 1344 6879 28,050 146 0
1449 1449 36,450 l
NOTE: The pressure, Pmax, f r the inlet piping was taken as 2400 psi.(or
, /
I approximately the pressure required to open the valve). This pressure reduces at some time after valve opening. The pressure downstream of tha safety valve was assumed to be small.
Neglecting this term downstream'does not affect conclusions reached from consideration of stress results.
I 1
l:
i i
I l
a i
4 1
26 1
I
TABLE 3.
STRESSES IN SELECTED PIPE ELB0WS FOR HOT LOOP SEAL ANALYSIS B P 0
y max o Total Level C Equation (9)
Allowable 2t" B M 2 B Stress Stress Element (See Note)
Z (psi)
(psi) 2B 2170 1496 3666 28,050 i
l 8B 2170 639 2809 28,050 3350 3350 36,450 278 1718 1718 36,450 518 5.156 5156 36,450 81B 108B 2170 2242 4412 28,050 114B 2170 486 2656 28,050 627 627 36,450 1198 142B 2170 1043 3213 28,050 1079 1079 36,450 1498 1267 1267 36,450 1618 NOTE: The pressure, pmax, for the inlet piping was taken as 2400 psi (or approximately the pressure required to open the valve).
This pressure reduces at some time after valve opening.
The pressure downstream of the safety valve was assumed to be small.
Neglecting this term downstream does not affect conclusions reached from consideration of the stress results.
e 27 i
y;, -
~
bl 2, t, 3
4, -, -
4 f
- ~
3
~
. -,, r TAB [.b4.9 FORCES IN SELECTED SNUBBERS F0F HOT-LOOP' SEAL ANALYSIS Force Location (1b)
- (' ;
)'
Header '
5327.
,s.
c-First Vert'ical Run Below' Header-2761
'I
. g J
- Next Horizontal Run 1560
~
Next Vertical Run 3683 b
1 i' k i.
t y
e i
i n
4'.
f t
i I
b i
f M
4 28 t-t
/
nr.
.m..n
,-n,
,e
. -- -~-.
y-r,-.
--m-.
The cold loop seal analysis was performed to determine the severity of-stresses that might occur in a safety valve piping system that employs cold loop seals.
For this case, the stresses in the most highly stressed straight segments and pipe elbows are shown in Tables 5 and 6.
The stresses in several straight elements, particularly in the header region where the safety valve lines and PORV line meet (elements 78, 116, and 146), exceeded the allowable stress. Many elbow element stresses were above the allowable stress, with the highest stress (129,146 psi) occurring in element 27, the third pipe bend downstream from Valve "B".
The allowable stress for the Class 1 inlet piping was 28,050 psi.
The Class 2 allowable stress for the discharge piping was 36,450 psi.
The maximum forces in the snubbers are shown in Table 7.
Since the sizes of the snubbers used in the V. C. Summer Plant were not available at the time this report was prepared, the snubbers were assumed to be equivalent to a Pacific Scientific model PSA-35, which is one of the largest size snubbers available. The allowable force for the PSA-35 is 54,600 lb.
Table 7 shows that four snubbers have a force greater than 54,600 lb.
Snubber manufacturers produce another type of support, a 0
compensating strut, which can carry a load of up to 3.7 x 10 lb.
Therefore, equipment is available to support the loads shown in Table 7.
All of the stresses presented above correspond to safety valve discharge loading alone and were compared to Level C Service Limits.
Consideration of a faulted conolcion would require that safety valve discharge loads be combined with safe shatdown earthquake loading. This would obviously increase the stresses presented above, but Level D Service Limits would increase the allowable stress from 2.2 S to 3.0 S -
h h
Because of the large safety margin in the stresses calculated for the hot loop seal simulation, those stresses combined with seismic stresses would almost certainly meet the faulted condition allowables. The presented results indicate that the V, C. Summer plant piping is adequate, but that a similarly designed system employing cold loop seals would probably be j
overstressed.
29
TABLE 5.
STRESSES IN SELECTED STRAIGHT PIPE SEGMENTS FOR COLD LOOP SEAL ANALYSIS 0
P,,x o Total D
1 2t Equation (9)
Allowable n
8 M 2 B Stress Stress Element (See Note)
Z (psi)
(psi) location 1
5535 24,374 29,909 28050 Inlet to valve "B"
-44 0
29,510 29,510 36450 Between junction 45 0
28,435 28,435 36450 area and relief tank 78 0
65,802 65,802 36450 Between valve "A" and header 115 5535 22,690 28,225 28050 Inlet to valve "A" 116 0
47,321 47,321 36450 Between valve "C" and header 145 5535 20,443 25,978 28050 Inlet to valve "C"
146 0
37,904 37,904 36450 PORV line NOTE: The pressure, pmax, f r the inlet piping was taken as 2400 psi (or approximately the pressure required to open the valve).
This pressure reduces at some time after valve opening.
The pressure downstream of the safety valve was assumed to te small.
Neglecting this term downstream does not affect conclusions reached from consideration of the stress results.
J 1
e 4
o 30 7
n.
TABLE 6.
STRESSES IN SELECTED PIPE ELBOWS FOR COLD LOOP SEAL ANALYSIS B P 0
y max o Total Equation (9)
Allowable 2t" B M 2 B Stress Stress Element (See Note)
Z (psi)
(psi)
Location 28 2170 35,381 37,551 28,050 Inlet to valve "B" 88 2170 32,910 35,080 28,050 Inlet to valve "B" 278 0
129,146 129,146 36,450
-Between valve "B" and header 51B 0
104,823 104,823 36,450 Between junction area and relief tank 818 0
114,526 114,526 36,450 Between valve "A" and header 1088 2170 29,041 31,211 28,050 Inlet to valve "A" 1148 2170 32,108 34,278 28,050 Inlet to valve "A"-
1198 0
128,547 128,547 36,450 Between valve "C" and header 142B 2170 27,202 29,370 28,050 Inlet to valve "C"
1498 0
78,646 78,646 36,450 PORV line 1618 0
128,547 128,547 36,450 PORV line NOTE: The pressure, pmax, f r the inlet piping was taken as 2400 psi (or
^
approximately the pressure required to open the valve). This pressure reduces at some time after valve opening.
The pressure downstream of the safety valve was assumed to be small. Neglecting this term downstream does not affect conclusions reached from consideration of the stress results.
e 31 a-
TABLE 7.
FORCES IN SELECTED SNUBBERS FOR COLD LOOP SEAL ANALYSIS Force Location (lb)
Header 133,350 First Vertical Run Below Header 189,760 Next Horizontal Run 74,689 Next Vertical Run 152,230 e
32
7.
THERMAL-HYDRAULIC EXPLANATION FOR DIFFERENCES IN STRESS RESULTS Two calculations of piping stress, based on two different thermal-hydraulic simulations, were performed in this work. As documented in Section 6, the stress results for the two cases are significantly different. This section discusses details of the two thermal-hydraulic simulations that explain the differences between the two estimates of piping stresses.
The difference between the calculations of safety valve piping stresses is due to different void fractions calculated in the system by the two thermal-hydraulic simulations.
In the cold loop seal simulation, volume vapor void fractions as low as zero (the volume was full of liquid water) were calculated in the piping between the safety valves and the header. Void fractions of about one-half were calculated in the header and much of the relief tank piping as the loop seal liquid passed through the system.
In contrast, significant flashing of the loop seal liquid as it passed across the valve junction in the hot loop seal thermal-hydraulic simulation resulted in a vapor void fraction that never dropped lower than 0.6 in the discharge piping.
In all but the first ten feet of the piping immediately downstream of each valve, the void fraction was greater than i
0.93.
As the density of liquid water is much greater than the density of steam, the momentum forces exerted by the discharge fluid on the piping will be greater for the case of lower vapor void fraction.
In the cold loop seal simulation, void fractions as low as zero are calculated even though the liquid was initially distributed with a void fraction near one-half.
The loop seal liquid initially placed downstream of the valve was not moving at valve actuation.
Modeling the liquid as stationary allows the acceleration of liquid from an upstream volume (nearer the valve) into a downstream volume as the valve opens, to increase 1
the amount of liquid in the downstream volume.
Figure 9 shows the decrease in vapor void fraction (increase in liquid fraction) occurring in a volume downstream of the valve that results from the initial motionless placement 33
~. _
E t
F i
1 i
i i
0.8 C
O 5
Valve 8
0.6 pens k
f 1
.V-o>
0.4 o-a.
v 0>
0.2 0
i.
0.5 0.6 0.7 0.8 0.9 Time (s) i f
i i.
7 Figure 9.
RELAPS calculation of volume vapor void fraction.
downstream of safety valve.
1 i
i i
-34 j-L
of the loop seal liquid downstream of the valve. Once this water spike is calculated, it is carried down the piping and causes the maximum piping loads throughout the entire system.
The liquid that gets distribued downstream of the-valve will, in reality, have a significant velocity. The velocity of the liquid will determine'to what extent the void fraction in an actual discharge will exhibit'the behavior shown in Figure 9.
It is expected that the liquid I
velocity at valve pop is small-enough that the void fraction history in an actual cold loop seal discharge is-quite like that shown in Figure 9.
The guidelines followed in placing the cold loop seal liquid downstream of the safety valves produced good agreement with forces l
measured in'the EPRI valve test facility.8 However, the V. C. Summer plant loop seal volume is 50% larger than the loop seal volume in the EPRI facility. The larger loop seal resulted in more liquid being placed I
downstream of the valves in the V. C. Summer plant simulation than in the simulation of the EPRI facility. Thus, it is likely that RELAPS calculated
[
a larger water spike for the V. C. Summer plant than for the EPRI facility. The larger water spike could cause a conservatively high l
calculation of forces and stresses in the piping system. At this time, there is no information that shows applying the guidelines to a slightly larger scale will still result in an accurate calculation of piping stress.
Still, it is believed that, for this cold loop _ seal simulation, the guidelines result in realistic stress calculations.
Modeling of valve simmer in the hot loop seal thermal-hydraulic
~
simulation is believed to result in a slightly high calculation of vapor
[
void fraction (too little liquid) in the discharge piping. The oscillatory I
behavior of the valve was not modeled in the hot loop seal simulation; a small constant valve flow area was instead us'ed to simulate valve simmer.
A relatively large plug of liquid probably passes through the valve each j
time the valve flow area reaches a relative maximum during the oscillations of actual valve discharge. Almost no liquid flows through the valve at the minimum flow area of the oscillations.
By contrast, the thermal-hydraulic simulation models valve flow during simmer as a steady stream.
It is 35 i
expected that the plugs of liquid characteristic of actual discharge will.
not flash to steam as quickly as the modeled stream-like discharge. Hence, void fraction just downstream of the valve may be overpredicted in the hot loop seal thermal-hydraulic simulation. The actual discharge, consisting of more liquid than calculated, could result in piping loads slightly greater than those calculated. Also, oscillations of the valve may increase stresses experienced by piping in the vicinity of the valve.
However, the actual stresses are still most likely to fall within allowable values.
The stress calculations are quite sensitive to the void fractions calculated by the thermal-hydraulic-simulations.
In addition to the modeling technique used to represent the condition of the loop seal liquid, nodalization of the discharge piping can have a significant impact on.the calculation of void fraction. Appendix B presents results from parametric RELAP5/M002 calculations of valve discharge in tne EPRI facility which show f
the effect of nadalization on void fraction predictions. Data on void fraction during valve. discharge is not available in.the EPRI test results.
Extreme care must be exercised in performing the thermal-hydraulic simulation of valve discharge. Guidelines provided from analyses of-other facilities may not be applicable to a specific piping system.
In all events, benchmarking of the techniques used in performing ch an analysis against actual data should be performed.
i 1
i I
i
\\
i i
i I
l 36 i
f 8.
OTHER ANALYSIS TOPICS Some approximate methods have been used by various. utilities in their analyses of pressurizer safety valve piping systems. The review of these analyses per NUREG 0737 Section II.D.1 requirements raised questions as to the adequacy of the approximations. A-secondary purpose of the investigations _ conducted herein was to utilize the structural model of the V. C. Summer piping systems to evaluate -three of these approximate ~
approaches. They were:
(a) neglecting the effect on the dynamic analysis of axial stretching forces in each pipe segment, (b) modeling the snubbers l.
as rigid rather than flexible supports, and (c) approximating dynamic results by performing a static analysis _using the peak loads multiplied by'.
i a dynamic load factor of 2 at all loading locations in the piping system.
All of the evaluations discussed below were performed for the cold loop seal case.
8.1 Including Axial Stretching Forces The axial forces which are calculated by BLAZER for a fluid wave passing through two elbows are shown in Figure 10. The two forces on a i
pipe segment, F and F, are called subforces.
BLAZER has the option A
g 1
of creating a combined force on the pipe segment, i.e., FA+F' A
B common approximation used in the dynamic analysis of safety valve piping in pressurized water reactor systems is to use only the combined force'on the segment and thereby neglect the effect of axial stretching produced by the subforces. Computer runs were made with and without the subforces applied so as to measure the significance of this effect in the V. C. Summer safety valve piping.
It was found, in this particular analysis, that the moments in the piping system were essentially the same for both computer runs, i.e., the maximum difference was about 1 to 2% Thus, the action of subforces could have been neglected when combined forces were applied to each straight pipe segment.
~
37 i
,. ~., _. - - -. -. _ _, - ~
_, ~ -, - -
, - -.... - -.... - -. - - - _ - - - _. - _ ~ ~. -,. - - -
r 3
s 0 Flow In
- a. Time 1 FB=0 FA A
////////
B
~
9/
h&
a Flow in
- b. Time 2 > Time 1 h
FA FB_
///////////////
+/
lL
/
/
C u Flow in
- c. Time 3 > Time 2 7 3009 1
Figure 10.
Fluid wave passing through two elbows.
38
i I
l Results from this investigation indicate that consideration of axial stretching effects was not necessary for this piping system and quite possibly unnecessary for overpressure protection systems in other PWRs.
It should not be concluded, however, that consideration of axial stretching of the pipe is not necessary in any piping system. An example of a system where axial stretching caused by subforces changed the resulting bending moments in the piping is provided in Reference 13.
8.2 Modeling Snubbers as Rigid Suoports At least one utility has modeled a portion of the restraints in its safety valve piping system as rigid when, in reality, these restraints likely have some flexibility that was not included in the analysis. The effect of modeling the snubbers in the V. C. Summer system as rigid was evaluated by performing an analysis in which all snubber stiffnesses were 10 increased from 100,000 lb/in. to 10 lb/in.
The stresses in the straight segments and pipe elbows from this analysis are shown in Tables 8 and 9, respectively. A comparison with Tables 5 and 6 shows that, as would 10 be expected, stresses in the piping were smaller for the 10 lb/in.
case. This comparison shows that some of the stresses decreased rather significantly.
The snubber loads are shown in Table 10.
Seven snubber loads exceeded the allowable load for a PSA-35 snubber of 54,600 lb. The maximum load in any snubber was 135,250 lb in the first vertical run below the header.
This is smaller than the maximum load from the 100,000 lb/in. snubber stiffness case (189,760 lb).
However, as expected, most snubber loads 10 increased for the 10 lb/in. case. A comparison with Table 7 shows that some snubber loads changed substantially.
The difference in support stiffness caused significant differences in stresses because rigid supports do not allow deflection of the system at the support points, whereas snubbers having a more realistic stiffness allow some deflection. As their stiffnesses increase, snubbers generally carry a greater load. A redistribution of stresses occurs in which the piping stresses generally decrease. However, the natural frequency of a 39
10 TABLE 8.
STRESSES IN SELECTED STRAIGHT PIPE SEGMENTS FOR 10 LB/IN.
SNUBBER STIFFNESS 0 P D
1 max o Total Level C Equation (9)
Allowable 2t" 0 N 2 B Stress Stress Element (See Note)
Z (psi)
(psi) 1 5535 17,129 22,664 28,050 21 0
21,875 21,875 36,450 38 0
14,355 14,355 36,450 77 0
18,007 18,007 36,450 78 0
16,362 16,362 36,450 107 5535 10,727 16,262 28,050 115 5535 16,680 22,215 28,050 116 0
29,428 29,428 36,450 137 5535 13,535 19,070 28,050 145 5535 18,646 24,181 28,050 146 0
15,538 15,538 36,450 155 0
10,535 10,535 36,450 158 0
6,790 6,790 36,450 NOTE:
The pressure, pmax, f r the inlet piping was taken as 2400 psi (or approximately the pressure required to open the valve).
This pressure reduces at some time after valve opening. The pressure downstream of the safety valve was assumed to be small. Neglecting this term downstream does not affect conclusions reached from consideration of the stress results.
e 40
10 TABLE 9.
STRESSES IN SELECTED PIPE ELBOWS FOR 10 LB/IN. SNUBBER STIFFNESS 0 P D
1 max o Total Level C 2t Equation (9)
Allowable B2 "B Stress Stress Element (See Note)
Z (psi)
(psi) 2B 2170 25,061 27,231 28,050 88 2170 19,174 21,344 28,050 328 0
51,659 51,659 36,450 41B 0
19,666 19,666 36,450 76B 0
49,777 49,777 36,450 818 0
40,865 40,865 36,450 85B 0
43,949 43,949 36,450 918 0
38,551 38,551 36,450 938 0
47,417 47,417 36,450 1008 0
58,213 58,213 36,450 1198 0
55,514 55,514 36,450 1208 0
58,983 58,983 36,450 1308 0
69,779 69,779 36,450 144B 2170 25,507 27,677 28,050 1618 0
6,067 6,067 36,450 NOTE: The pressure, pmax, f r the inlet piping was taken as 2400 psi (or approximately the pressure required to open the valve). This pressure reduces at some time after valve opening.
The pressure downstream of the safety valve was assumed to be small. Neglecting this term downstream does not affect conclusions reached from consideration of the stress results.
i i
41
10 TABLE 10.
FORCES IN SELECTED SNUBBERS FOR 10 LB/IN.' SNUBBER STIFFNESS Force Location (lb)
Header 133,460 First Horizontal Run Below Header 104,590 First Vertical Run Below Header 135,250 Next Horizontal Run 103,270 Next Vertical Run 99,711 Two Snubbers Nearest Header 72,639 on Line from Valve A-58,854 i
42
flexible snubber could be excited by the vibratory motion of the piping, leading to significant stresses in the support and nearby piping.
Based on these results, the expedient practice of treating all supports as rigid elements in an overpressure protection system is not recommended. The supports should be modeled according to their actual stiffness values.
8.3 Approximating a Dynamic Analysis with a Static Analysis and a Dynamic Load Factor At least one utility performed a static analysis on its safety valve piping employing a dynamic load factor of 2 in lieu of performing a dynamic analysis. To evaluate whether this procedure yields suitably conservative results, two static analyses were performed on the V. C. Summer system in which the dynamic forces were amplified by a factor of 2.
In the first static analysis the peak net force from the force-time history for each pipe segment in the V. C. Summer model was multiplied by 2 and applied to that pipe segment regardless of when the peak force may have occurred. The results showed that some pipe bending moments from the static analysis were very conservative and others were very nonconservative relative to the dynamic analysis.
For example, the static analysis showed a resultant moment for element 51 of only 665,000 in.-lb compared to 6
1.12 x 10 in.-lb frem the dynamic analysis.
In the second static analysis, the net dynamic force acting on each pipe segment at a selected point in time was multiplied by a factor of 2 i
and applied to the model as a static force.
The static analysis was performed at a time of 0.181 seconds.
This time was chosen because the forces in the 12-in. piping in the vertical section nearest the pressurizer relief tank were the greatest at that time. Again, certain bending moments were nonconservative.
The largest moment for this run was only 6
770,000 in.-lb compared to 1.12 x 10 in.-lb for the dynamic analysis.
It is possible that selection of another time may produce more conservative 43
results, but it is not likely that the static analysis results for any selected time would exceed the dynamic analysis results everywhere in the piping system. Both static analysis techniques investigated gave nonconservative results in portions of the piping system.
Evidently the frequency content of the fluid forcing functions excited system natural frequencies to produce dynamic amplification greater than 2.
A further complication is that the forcing functions are applied to the piping system in a sequence as the fluid moves down the pipe, resulting in a complex interaction between force application and structural response.
Because of the complexity of the piping system and the load histories applied to the system, a time-history dynamic analysis is generally required to accurately determine the dynamic response of the piping system.
e 44
9.
CONCLUSIONS AND RECOMMENDATIONS The pressurizer safety valve piping system for the V. C. Summer Nuclear Station, Unit 1, was analyzed for stresses caused by hydrodynamic loads due to a safety valve actuation.
Stresses were calculated for each of two independent thermal-hydraulic simulations of loop seal discharge through the safety valves.
The stress results from the two calculations were significantly different.
One simulation, considered a best estimate calculation of safety valve discharge in the V. C. Summer plant, modeled hot loop seals upstream of the safety valves.
The other simulation was a conservative estimate of a cold loop seal discharge, and indicates the severity of safety valve discharge events in plants employing cold loop seals.
The hot loop seal calculation showed all stresses to be well within allowable values according to criteria of the ASME Code Section III,1983 Edition.
This calculation supports the licensee's conclusion that the safety valve piping system is adequate / An analysis of seismic loads and PORV discharges would also have to be performed to completely validate the licensee's conclusions. The results from the cold loop seal calculation showed that the stresses in the straight sections of piping were generally acceptable but that several pipe elbows were overstressed.
This calculation also showed several snubbers of assumed stiffness to be overloaded. The snubber stiffnesses were estimated because the size of snubbers installed in the Summer plant were not known at the time this analysis was performed.
The magnitude of the stresses is highly dependent on the thermal-hydraulic predictions of void fraction in the piping downstream of the safety valves. As the density of liquid water is much greater than the density of steam, the momentum forces exerted by the discharge fluid on the piping will be greater for the case of lower void fraction.
RELAPS/M002 predictions of void fraction during safety valve discharge were found to be quite sensitive to how the initial state of the loop seal liquid is modeled and to the nodalization of the discharge piping.
The use 45
of RELAP5/M002, or a similar computer code, does not in itself ensure that the resultant thermal-hydraulic prediction will be suitable for accurately determining stresses in a safety valve piping system.
Guidelines for modeling loop seal discharge behavior in one facility may not be appropriate for another facility, particularly if empiricism exists in the guidelines.
For instance, guidelines developed using data from the EPRI valve test facility prescribe simulating valve simmer by distributing a portion of the cold loop seal liquid downstream of the valve.
This distribution of loop seal liquid produced a thermal-hydraulic simulation that was used to accurately predict the measured piping forces.
As discussed in Section 7, although it is believed the available guidelines give reasonable results, applying this guideline to a larger loop seal may result in an overly conservative calculation of void fraction, and therefore of stresses, in the discharge piping.
Because of sensitivities in nodalizatien and empirical guidelines to the calculation of void fraction, extreme care must be taken in performing a thermal-hydraulic calculation of safety valve discharge.
The analysis performed by the licensee showed the design of the overpressure protection system to be adequate for safety valve actuations.
The analysis presented herein supports the licensee's conclusion.
Since few of the details of the licensee's analysis were available, the methodology used therein could not be evaluated directly.
Therefore, results of the analysis described herein do not necessarily verify the licensee's methodology.
In any case, the licensee's results would be expected to differ to some degree from results of the analysis contained in this report. As previously mentioned, the calculated fluid forces are sensitive to techniques used in thermal-hydraulic modeling of loop seal discharge behavior.
Different techniques for calculating the fluid forces from fluid data generated in the thermal-hydraulic analysis can also lead to differences in the calculated forces.
Differing approaches to the structural dynamic analysis could affect results obtained.
For example, use of a modal superposition time history solution rather than direct integration could lead to different results depending on how contributions l
46
i from some of the higher frequency modes are treated. What these differences in analytical methodology indicate is that any two organizations analyzing a loop seal discharge through the safety valve in~
an overpressure protection system will likely produce differences in solutions to the same problem. Whether conclusions reached regarding
~
safety of the system are appropriate depends on the accuracy of assumptions made in the analysis.
Conclusions reached from the consideration of other topics, as discussed in Section 8, are summarized below:
1.
Including axial stretching forces on the pipe segments in the system, in addition to the combined (or net) forces, caused no appreciable change in resulting bending moments in the piping system for this particular analysis. This result supports the practice of neglecting these balancing axial forces in othar 2
analyses of pressurizer safety valve piping systems. However, it cannot be concluded from this study that exclusion of this axial stretching effect is justified for all piping systems.
l 2.
Modeling snubbers as rigid restraints rather than using the actual snubber stiffness caused significant changes in the pipe stresses and support loads. As would be expected, the support loading generally increased while the pipe stresses diminished.
l The practice of modeling all supports as rigid restraints is not recommended since inaccurate results may well be obtained.
3.
Performing a static analysis in which the peak dynamic forces I
were amplified by a dynamic load factor of 2 'resulted in some stresses that were conservative and others that were l
significantly nonconservative.
Thus, a dynamic load factor of 2 is not high enough to ensure a conservative analysis. Because
~
of the complexities involved in the dynamic analyses of fluid discharge transients, the use of static representations is not recommended.
I i
47
10.
REFERENCES 1.
Gilbert Associates, Inc., Pressurizer Relief System Piping and Support Evaluation Report for Virgil C. Summer Nuclear Station, July 30, 1982.
2.
V. H. Ransom et al., RELAP5/M002 Code Manual, NUREG/CR-4312, August 1985.
3.
J. R. Olsen, BLAZER:
Release 3 Version 1 Code Manual, EG&G Internal Technical Report RE-A-82-045, June 1982.
4.
R. C. Guenzier, IBM 360/75 and CDC7600 Version of SAP IV, A Structural Analysis Program for Static and Dyna,mic Analysis of Linear Systems, Aerojet Nuclear Co., Report TR-775, Jar.uary 1976.
5.
American Society of Mechanical Engineers, ASME Boiler and Pressure Vessel Code,Section III, Division 1, Subsection NC, July 1983, Winter 1983 Addenda.
6.
O. W. Dixon, Jr., Vice President Nuclear Operations, South Carolina Electric and Gas Company ltr. to Harold R. Denton, Director, Office of Nuclear Reactor Regulation, U.S. Nuclear Regulatory Commission,
" Virgil C. Summer Nuclear Station, Docket No. 50/395, Safety and Relief Valve Report, NUREG-0737, Item II.O.1," July 30, 1982.
7.
A. Meliksetian et al., Valve Inlet Fluid Conditions for Pressurizer Safety and Relief Valves in Westinohause-Designed Plants, Westinghouse Electric Corporation Interim Report, February 1982.
8.
R. K. House, Application of RELAP5/ MOD 1 for Calculation of Safety and Relief Valve Discharge Piping Hydrodynamic Loads, Intermountain Technologies Inc., Interim Report, March 1982.
9.
Flow of Fluids Through Valves, Fittings, and Pipe, Crane Company, Technical Paper No. 410, 1979.
- 10. Aerospace Fluid Component Designer Handbook, RPL-TDR-64-25, February 1970.
11.
Safety and Relief Valve Test Report, Electric Power Research Institute, Interim Report, April 1982.
12.
EPRI Safety Valve Test Report Vol. 6, Test Results for Crosby Safety Valve Model HB-BP-86, 6M6, Combustion Engineering Inc., EPRI Research Project V102-2, Interim Report, July 1982.
13.
A. G. Ware, " Moment Loads Induced by Pressure and Momentum Forces in Piping," ASME Journal of Pressure Vessel Technology,104, November 1982, pp. 268-271.
i l
1 48
GD 9
APPENDIX A METHOD OF OYNAMIC ANALYSIS i
9 e
A-f
i CONTENTS 1.
INTRODUCTION..................................................... A-1 2.
CALCULATION OF FLUID FORCES...................................... A-1 3.
APPLICATION OF FLUID FORCES...................................... A-6 4
4.
SAP IV TIME HISTORY SOLUTION..................................... A-6 5.
TIME HISTORY SOLUTION TECHNIQUE.................................. A-8 6.
TREATMENT OF INITIAL CONDITIONS.................................. A-9 7.
DAMPING IN PIPING SYSTEM........................................ A-9 FIGURES A-1. RELAP5 repfesentation of two pipe bends.......................... A-2 A-2. Generalized fluid container...................................... A-3 A-3. Typical fluid force-time history at the header................... A-5 A-4. Application o' fluid forces to a pipe segment.................... A-7 a
A-il
1.
INTRODUCTION The analysis performed to predict the dynamic response of the safety valve piping is accomplished in a series of operations.
Fluid forces exerted on the piping during discharge are calculated from fluid pressures and momenta obtained from RELAP5/M002 output and are applied to the finite element model. A time history solution for the dynamic response of the piping is generated using the SAP IV structural analysis program.
Performance of the analysis requires special consideration for initial conditions and damping in the piping system.
The operations are discussed in detail in the sections that follow.
2.
CALCULATION OF FLUID FORCES The fluid forces exerted on the safety valve piping during safety valve discharge were calculated from RELAP5/M002 predicted fluid pressures and momenta using a computer program called BLAZER. This program reads the fluid data from a tape, computes force-time histories at specified points on the piping system, and generates force-time histories in a format that can be input directly into the structural analysis code.
The forces were calculated using the R5 FORCE subroutine in BLAZER.
In RELAP5/M002 each pipe leg is subdivided into several control volumes as shown in Figure A-1.
The forces shown in Figure A-2 were calculated by R5 FORCE for each control volume.
Based on the cross sectional view of the arbitrarily shaped container and the notation specified in Figure A-2, the net force on a control volume can be expressed as:
2 F = - (Py1 + pg u)Ay3 + (PI2
- PI")AI2 + 'El El -P E2 E2 + TA A
s I
A-1
J 1
4 4
i l
1 i
Pipe with two elbows I
i f
SF1 SF2
+ ---
+
CV5 CV6 CV7 CV8 CV9 CV10CVII a
i CV4 CV12 1
CV3 CV13 I
l CV2 CV14 Cv1 Cvis RELAPS representation l
Note: CV = controlvolume SF = sub force 7 3006 i
1 Figure A-1..RELAPS representation of two pipe bends.
A-2
l l
i l
.-u.
i i
e Pl1, Ajj PE1, AE1 p
,4
'Pl,U @Aj
'PE2eAE2 i
r P, U;
@A12 4
rsoor Figure A-2.
Generalized fluid container, A-3
where appropriate volume surface area (as shown in Figure 10)
A
=
fluid pressure P
=
fluid velocity u
=
fluid density
=
p shear force per unit area.
t
=
The control volume forces were then combined to produce subforces acting on the pipe as shown in Figure A-1.
For this case, subforce 1 may be defined as the sum of control volume forces 5, 6, and 7, and subforce 2 as the sum of control volume forces 8, 9, 10, and 11.
The BLAZER program contains an option of adding all subforces on a pipe segment into a single combined force, which amounts to the net force c \\l on the segment. Application of only a combined force on a pipe segment in the stress analysis of a system, however, neglects the effect ofs axial stretching caused by opposing subforces acting at the ends of the segment.
This is more thoroughly discussed in Section 8.1 of this report.
/
A further feature of the BLAZER program is that, once a force-time history is produced, the number of force-time pairs in the history can be reduced while maintaining the peaks and important characteristics of the history. This is accomplished by concentrating the greatest number of force-time pairs retained into regions where the history changes most rapidly.
In this analysis, the original force histories were reduced in size to 390 force-time pairs to stay within computer core allocations.
A typical fluid force-time history, corresponding to the header section of the model, is shown in Figure A-3.
Both the original and reduced forcing functions are plotted, showing that the reduced fur,ction appropriately matches the original.
A-4 ki
n ORIGINAL FORCING FUNCT;0N i
AT NOCE G3 0
=
l
. m.
-200 j
j E
-400
- s
-600 i
3
-800 d
/
w -1000 II v
5 u.
-1200 B
-1400
-1600
-1800 g,,.
0.0 0.5 1.0 1.5 2.0 TIME (SECl REOUCED FORCING FUNCTION AT NODE G3 200 1
-4co
"=
-600 E
a
-e00 d
/
w -1000 j
o 3
^
=
u.
-1200
- s
-1400 l
-1600 l
ll 1
H
=
i
-1800 5
g.
c.0 c.s t.o 1.s 2.0 f
^
T:ME tsEci i
Figure A-3.
Typicai fluid force-time history at the header.
)
A-5 4
r 3.
APPLICATION OF FLUID FORCES In the cold loop seal analysis the combined (or net) force acting on
}
each pipe segment was applied at some point along the length of the segment in the structural model. As noted in Section 2 above, though, application cf just the combined force does not include the offects of pipe stretching caused by opposing subforces acting at the ends of the segment.
- Thus, equal and opposite subforces were also applied at the ends of all segments to account for this axial stretching effect.
In this manner, the total of the-three forces applied to each segment eqJaled the net force on the segment as required, and any axial extension ir the segments was included in the piping response. This loading configuration is depicted in Figure A-4.
In the hot loop seal analysis performed en the system, the combined force acting on each segment was again applied at some point along the I
length of the segment.
In this analysis, however, subforces were applied only at the ends of segments having length greater than 5 ft.
Because of results reported in Section 8.1 of this report, the omission of subforces on shorter segments was expected to have a very minimal effect on the calculated piping response.
4.
SAP IV TIME HISTORY SOLUTION The SAP IV structural analysis program was used to perform the analysis for the dynamic response of the safety valve piping to fluid transient loading. The program employs the finite element method and numerical integration techniques to solve for static and dynamic response of linear structural systems.
l The SAP IV program was especially useful because of its ability to solve the equations of motion for the system by the direct integration method.
Use of direct integration analysis eliminates the need to perform an eigenvalue-eigenvector solution and assures that only very high A-6
u I
i f
/g I **
Flow direction Subforce A Subforce-A
-+-
l 73008 i
Figure A-4.
Application of fluid forces to a pipe segment.
i l
A-7 m
c
frequency response is filtered from the solution, provided that a small enough time step is used. Use of the direct integration solution technique is discussed in greater detail in Section 5 below.
The SAP IV computer program outputs internal forces and bending moments at the ends of the finite element:.
In this analysis, the program output a force and bending moment for each coordinate axis at the ends of all pipe elements.
It also output forces in the boundary elements.
5.
TIME HISTORY SOLUTION TECHNIQUE Determination of the dynamic response of the structural model to the fluid loads using the finite element method required solving the following set of linear equations:
M'u' + CO + Ku = R where M, C, and K are the mass, damping, and stiffness matrices of the element assemblage; and u, 0, U, and R are displacement, velocity, acceleration and generalized fluid vectors, respectively.
The mat ices M and K are determined from program input corresponding to the characteristics of the structural model. The matrix C is calculated from the relation C = aM + SK, since Rayleigh damping is assumed in a SAP IV analysis.
(The damping constants a and S are discussed in Section 7 below.) The vector R is determined from the force-time histories that are applied to the structural model.
The vectors u, u, and U represent the response of the system and are generated from the solution to the set of equations above.
From the vector u, the internal forces and moments of the system are calculated.
The solution for the equations was obtained in this analysis by direct step-by-step integration using a time step of 0.001 s.
This time step was chosen to include all response up to a frequency of at least 100 Hz in the analysis.
A-8
The result of the time-history solution was a set of values for vectors u, u, and 'u' at each time step.
This system of vectors constitutes the time history response of the piping system to the fluid loads applied.
From vector u, the SAP IV program computed the internal forces and bending moments in the system.
6.
TREATMENT OF INITIAL CONDITIONS The piping system is subjected to internal pressures even before discharge through the safety valves begins.
These pressures create nonzero fluid forces on the system at the beginning of the time-history solution period. The SAP IV program, however, only computes a time-history solution when all applied forces are equal to zero at time zero.
For this reason, the initial force for each force-time history was subtracted from its corresponding history prior to performing the time history dynamic solution.
In this way all applied force time histories did indeed begin with a zero value at time zero.
In these particular analyses the largest initial force was found to be 120 lb, so the stress due to initial forces was neglected.
7.
DAMPING IN PIPING SYSTEM As stated in Section 5 above, the SAP IV program assumes Rayleigh damping when a structure is analyzed by the direct step-by-step integration technique.
Rayleigh damping relates the percentage of critical damping
(() to the circular natural freauency (w) of the structure as follows:
w
& = ya + 6 y e
A-9
where a and S are damping constants for the system.
This relationship between damping and natural frequencies of the structure results in a damping matrix C that can be calculated from the I
l structure's mass and stiffness matrices (M and K, respectively) as follows:
C = aM + SK.
-4 In this analysis, values ~of 0.19218 for a and 2.27 x 10 for 6 were established to givs an average damping'of 1% up to a frequency of 100 Hz.
4 A-10 l
i.
__ ]
9 e
APPENDIX B EFFECT OF N00ALIZATION AND TIME STEP SIZE ON RELAP5/M002 CALCULATIONS OF VOID FRACTION DURING SAFETY VALVE DISCHARGE e
B-1
.m a
L CONTENTS i
1.
INTRODUCTION..................................................... B-1 2.
FACILITY AND TEST MODELE0........................................
B-2 i
-3.
IMPORTANCE OF VOID FRACTION CALCULATION..........................
B-4 4.
CALCULATIONS PERFORMED...........................................
B-5 5.
NODALIZATION EFFECTS.............................................
B-6 5.1 Heated Loop Seal...........................................
B-6 t
5.2 Cold Loop Seal.............................................
B-6 6.
' TIME STEP SIZE EFFECTS........................................... B-14 4
7.
CONCLUSIONS...................................................... B-18 8.
REFERENCES.....................................................
14.
FIGURES B-1.
Representation of the EPRI valve test facility for test 917.....
B-3 B-2.
Effect of nodalization on RELAP5 calculations of void fraction (position 1) during hot loop seal discharge............
B-7 B-3.
Effect of nodalization on RELAP5 calculations of void fraction (position 2) during hot loop seal discharge............ B-8 B-4.
Effect-of nodalization on RELAPS calculations of void fraction (position 3) during hot l loop' seal discharge............
B-9 B-5.
Effect of nadalization on RELAPS calculations of void fraction (position 1) during cold loop seal discharge........... B-10 B-6.
Effect of nodalization on RELAPS calculations of void fraction (position 2) during cold loop seal discharge........... B-11 i
B-7.
Effect~of nodalization on RELAPS calculations of void fraction (position 3) during cold loop seal discharge........... B-12 B-8.
Effect of time step size on RELAPS calculations of void fraction (position 1) during hot loop seal discharge............ B-15 i
B-11 l
m
(
B-9.
Effect of time step size on RELAPS calculations of void fraction (position 2) during hot loop seal discharge............ B-16 B-10. Effect of time step size on RELAPS calculations of void fraction (position 3) during hot loop seal-discharge............ B-17 O
i i
O^
. e B-iii
1.
INTRODUCTION Based on requirements of NUREG 0737,Section II.D.1, " Performance Testing of BWR,and PWR Relief and Safety Valves," all operating plant licensees are required to demonstrate the structural integrity of their as-built pressurizer safety valve / power operated relief valve piping
~
systems.
In fulfilling this requirement, many licensees have used
~
I thermal-liydraulic analyses of valve discharge to calculate the forces imposed on the piping system. Thus, the determination of the adequacy of the system is strongly dependent on the thermal-hydraulic analysis.
l The RELAP5 computer code has been used by some licensees to calculate the thermal-hydraulic behavior of valve discharge events. This appendix presents RELAPS calculations of safety valve discharge that show how-sensitive the RELAPS results are to changes in nodalization and time step size.
1 e
B-1
2.
FACILITY AND TEST MODELED 2
A RELAPS model of the EPRI safety valve test facility was used for the calculations presented below. The EPRI safety valve testing program was performed to obtain full scale operability data for safety valves over the full range of fluid condition under which they may be expected to opercte.
EPRI test 917 was modeled in the calculations. presented.
Test 917 is a test of model HB-BP-86, 6M6 Crosby valve with heated liquid in the loop seal just upstream of the valve. A representation of the test configuration used in test 917 is shown in Figure B-1.
A loop seal test was selected because the loop seal discharge subjects the piping system to higher stresse.s than a steam or single phase liquid discharge. Thus, the loop seal case is of greater concern for evaluating the adequacy of a piping system.
1 4
B-2 l
.. =. _ - _
1
]
Position 1 Safety _ /
Vdve n' 6.0 2.4 2.8 2.0 7.7 Position 2 2A 6.2
[] 12" x 8" Reducer v
Pressurizer Verticd Length (ft) 17.3 Atmosphere -
Position 3 3.2 Horizontal f
43.5 Figure B-1.
Representation of the EPRI vdve test facility.
for Test 917.
4 NSLOO499 2
(
3.
IMPORTANCE OF VOID FRACTION CALCULATION The greatest forces caused by a valve discharge event occur in the piping downstream of the valve as the loop seal liquid is propelled through the system. The density of the loop seal liquid inflicts large momentum forces on the piping system.
Thus, an' estimation of forces experienced by the system is very sensitive to the thermal-hydraulic calculation of void fraction.
Because the calculated void fraction controls the piping forces 4
predicted for valve discharge events, a comparison of void fractions can best illustrate the sensitivity of RELAP5 calculations to nodalization and time step size. Positions at which comparisons between void fraction calculations were made are indicated in Figure B-1.
l t
I S
B-4
.----...,.--1,-,-
--,,r,.-
e
,m m--.
4.
CALCULATIONS PERFORMED Three RELAPS calculations of EPRI test 917 were performed, including a base case, a finely noded calculation, and a calculation with a reduced time step.
The base case nodalization consists of two PIPE components joined by a servo valve with time dependent volumes at each end of the system. The loop seal portion of the pipe contains node lengths ranging between 0.5 and 1.2 ft.
Node lengths of 0.5 ft are used in the first 6.0 ft downstream of the safety valve.
Thereafter, node lengths of approximately 2.0 ft are used. The finely noded calculation was made with node lengths in the downstream piping one-fourth as long as the base case nodes.
The reduced time step calculation used the base case nodalization but a time step roughly one-tenth of the courant limit (which limited the base case run) was specified.
Two additional RELAPS calculations were performed which substituted a cold loop seal for the heated loop seal of test 917. One cold loop seal calculation was performed with the base case nodalization; t'.e other used the finely noded model.
In all five calculations, an identical set of controls regulated the servo valve behavior to approximate the opening action of an actual safety valve. Simmer and pop behavior was approximated.
The loop seai liquid was placed upstream of the valve in all cases. An identical pressurizer pressure history, resulting in the valve opening at 0.6 s into the calculation, was used in all cases.
Cycle 36.04 of RELAPS/ MOD 2 performed all five calculations.
\\
l B-5
5.
NODALIZATION EFFECTS 5.1 Heated Loop Seal Figures B-2, B-3, and B-4 show comparisons of the RELAP5 calculated void fractions at each of three positions in the downstream piping.
Figure B-2 shows that immediately downstream of the valve, the finely noded calculation predicts significantly more liquid content (lower vapor void fraction) than the base case.
Figure B-3 shows that, at position 2, most j
of the loop seal liquid has flashed to steam in both cases, but the finely 4
noded calculation still predicts a lower vapor void fraction in the loop seal discharge. As shown in Figure B-4, both calculations show that essentially all the liquid has flashed by the time the loop seal fluid has reached position 3.
The figures indicate a modest sensitivity of the RELAP5 calculations to nodalization for the hot loop seal discharge.
Only in the piping immediately downstream of the valve would estimated piping forces be expected to differ significantly between the two calculations.
The finely noded calculation would lead to the calculation of higher piping forces in the few feet immediately downstream of the valve.
5.2 Cold Loop Seal Figure B-5 shows that, for the cold loop seal discharge, both the base case and finely noded calculation predict a large liquid content immediately downstream of the valve shortly after the valve opens. As some of the liquid in the cold loop seal is below 212 F, not all of the liquid will flash to steam.
In general, then, the cold loop seal runs will calculate a greater liquid content (less vapor void fraction) than the hot loop seal runs.
Figures B-6 and B-7 show that, at positions 2 and 3, the finely noded case calculates a minimum vapor void fraction much lower than the base case does. The finely noded case would be expected to result in a significantly higher estimated peak force throughout the piping system.
B-6
= -,
N O
1 i
.\\
,1 i
8 i
g i
- %I g
s i
l
/
h
, A' g
p g
o i
f C
i '_r r
.3 0.9
'q o
I Il gl 5
I l BASE CASE g
li
-- FINE NODING h
0.8 4
O 0.5 1
1.5 -
2 Time (s) -
Figure B-2. ~ Effect of nodalization on RELAP5 calculations of void froction (position 1) during hot Icop seal discharge.
O B-7
k I
1
,{
8 i
i~
1 ll a-
)
0.99 It
~
I lI]
A ll o
3 0.98 ll o
E BASE CASE la
~ ~ FINE NODING i,
.r-0.97 g;
0 11 3
& 0.96 l
L.
g O
I 1
0.95 1
1 I
0.94 O
0.5 1
1.5 2-Time (s)
Fi gu r e B-3.
Effect of nodalization on RELAP5 calculations of void fraction (position 2) during hot loop seal discharge.
l 1
l B-8
4 I
1 i
i a
i I
I
/
la g /
'\\
l y
\\
I V
i i c
I gI O
ill Z
BASE CASE i
f 8
-- FI NE NOD I NG
/
t
/
it I'
.3 0.99 ti I
o>
L O
0.98 0
0.5 1
1.5 2
Time (s) i Figure B-4.
Effect of nodalization on RELAP5 calculations of void fraction (position 3) during hot loop sea! discharge.
1 1
B-9
i 1
i i
f 4
-i n
t i
I l
C 1
O l
o r j -0.75 8,
l 'l BASE CASE y
I, FINE NODING i
i 3
I i
g i
1 I
c.
i i
- 8. 0.50
(
/
o 0.25 O
0.5 1
1.5 2
Time (s)
Figure B-5.
Effect of nodalization on RELAP5 calculations of void fraction (position 1) during cold loop' seal discharge.
9 2
4 B-10
4 4
1 i,
i s' l
l 1
$\\\\ll 8
2 it 8
BASE CASE si
-- FINE NODING 1l lt u 0.8
'5 is II g
I 8.
l 8
f 0.6 O-0.5 1
- 1.5 2
Time (s)
Figure B-6.
Effect of nodalization on RELAP5 calculations of void fraction (position 2) during cold loop seal discharge.
4 B-11 4
,n r-
1 1
s e
n i
V i
I l 1 1 1 1 I
lI1l 8
l1il -
z
)
8 1,1
)
I! '
30.9 BASE CASE ll S
-- FINE NODING t
ll e
ll ii Il lg i
l I
I 0.8 l
0 0.5 1
1.5 2
l l
Time - (s)
Fi gu r e 8-7.
Effect of nodalization on RELAP5 calculations of void -
fraction (position 3).during cold loop seal discharge.
o B-12
Figures 8-5, B-6, and B-7 indicate a significant sensit'vity of the RELAPS calculation to nadalization for the cold loop seal discharge.
Throughout the discharge piping, except immediately downstream of the valve, forces estimated from the finely noded case are likely to be significantly larger than those estimated from the base case.
f t
i 9
h 1
B-13
6.
TIME STEP SIZE EFFECTS-Figures B-8, B-9, and B-10 show only small differences between RELAPS -
calculations of void fraction between the base case and the reduced time step case at the three positions investigated.
Reducing the time step did not change the calculation of thermal-hydraulic conditions downstream of
(
the safety valve during a hot loop seal discharge. Consequently, little difference between piping forces calculated from the base case and reduced time step run is expected.
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Figure B-8.
Effect of time step size on RELAPS-calculations of void - fraction (position 1) during hot loop seal discharge.
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Figure B-9.
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discharge.
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Effect of time step size on RELAP5 calculations of void fraction (position 3) during hot loop seal discharge.
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CONCLUSIONS The RELAP5 calculations performed support the following conclusions:
o RELAP5 simulations of safety valve loop seal discharge show that calculations of void fraction are sensitive to the nodalization used, particularly in the case of a cold loop seal discharge.
Because of the impact of void fraction on estimates of forces in a piping system, extreme care should be used if RELAPS is to provide a calculation which is appropriate for determining the adequacy of a safety valve piping system.
o Reduction of the requested time step below the code imposed courant limit does not significantly alter RELAP5 thermal-hydraulic calculations of safety valve discharge when a hot loop seal is employed upstream of the valve.
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REFERENCES 1.
V. H. Ransom et al., RELAP5/M002 Code Manual, NUREG/CR-4312, August 1985.
c 2.
EPRI Safety Valve Test Report Vol. 6, Test Results for Crosby Safety i
Valve Model HB-BP-86, 6M6, Combustion Engineering Inc., EPRI Research-Project V102-2, Interim Report, July 1982.
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BIBUOGRAPHIC DATA SHEET EGG-NTA-7639 Sgt insTavCTiohs og Tut stytast 3 f atti ago st.,e rirLE J leave.LANs ANALYSIS OF PRESSURIZER SAFETY VALVE DISCHARGE PIPING OF THE'V.C. SUMMER NUCLEAR POWER STD. TION, UNIT 1 A Q Af t stront COMPLETED MONin vtAA March 1987
. Auv ca's>
G. S. Case B. L. Harris
. oart ago.r issuto J. R. Larson G. K. Miller
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March 1987 7 Pte 70AMI4G CmGANs2Af eOm Mt 90 MAsu%G AOQatss t,aca aste Cee#
4 pmOJECT4TA4KMpOME uS11T NoMetA e
v EG&G Idaho, Inc.
.,0 o Ga.,a u t.
P. O. Box 1625 Idaho Falls, ID 83415 A6492 s
@ SPQq $QAING QAGANig Af eQN.sawg AmeO Maill4G ACOntS3 #saes ee te ce.o#
lie.Tvrt08 REPORT w
Division of PWR Licensing - B Informal Office of Nuclear Regulatory Commission Washington D.C 20555
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N/A 12 SUP*LivtNT AMT 40TES 4
1 l
l Based on the requirements of NUREG 0737,Section II.D.1, Performance Testing of BWR rid PWR Relief and Safety Valves," all operating plant licensees are required.to
, sdemonstrate the structural integrity of their as-built pressurizer safety valve / power operated relief valve piping systems. Thus, the licensee for the V. C. Summer Nuclear Power Station, Unit 1, among other licensees, performed such an analysis and submitted, the results to the Nuclear Regulatory Commission.
This report describes an analysis by EG&G'Idano, Inc. of the V. C. Summer pressurizer safety valve inlet and discharge piping for fluid loads caused by safety
{
i valve actuation.
Results from the analysis performed supported the licensee's conclusion that the system design is adequate for hot loop seal and steam discharge
'through the safety valves. The fluid forces developed during this analysis, in conjunc-I tion with the structural model, were also used to evaluate the accuracy of several simplified approacties used by various licensees in tfiefr analyses of safety valve piping systems.
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