ML18018B694

Rept on Operability of Pressurizer Safety Relief Valves, PORVs & Electrical Motor Operated Block Valves.

ML18018B694
Person / Time
Site:	Harris
Issue date:	04/30/1984
From:	CAROLINA POWER & LIGHT CO.
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ML18018B693	List: ML18018B692 ML18018B694
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NUDOCS 8407030108
Download: ML18018B694 (306)

Text

Report on the Operability of Pressurizer Safety Relief Valves, Power Operated Relief Valves and Electrical Motor Operated Block Valves for Carolina Power 6 Light Co.

Shearon Harris Nuclear Power Plant April, 1984 8407030i08 840628 PDR ADOCK 05000400 E PDR

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TABLE OP CONTENTS Section ~Pe e 1.0 Introduction 2.0 SRV Considerations 2.1 SRV Inlet Conditions 2 2.2 SRV Installation 4 2.3 Test Data Applicability 7 2.4 2.5 Valve Operability Evaluation Conclusions ll 13 3.0 PORV Considerations 3.1 PORV Inlet Conditions 14 3.2 PORV Installation 16 3.3 EPRI Test Data Comparisons 19 3.4 Conclusions 21 4' Electrical Motor Operated Block Valve 4.1 EMOV Installation 22 4.2 EPRI Test Information on Block Valves 24 4.3 Design Adequacy Based on Test Results 28 5.0 References 29

INTRODUCTION As a result of the Three Mile Island accident, the U. S. Nuclear Regulatory Commission required that all Pressurized Water Reactor utilities demonstrate the operability of the plant's safety and relief valves, and the structural integrity of their discharge piping and supports. These requirements were initially indicated in NUREG-0578 lj 2/

and later clarified in NUREG-0737. At the re-quest of the PWR utilities, the Electric Power Research Institute was engaged to develop and implement a test program to confirm the operability of the pressurizer PORV's and SRV's and to demonstrate the adequacy of the discharge piping and suppoxts. The PORV tests were completed in August, 1981 and the SRV tests were completed in January, 1982. The reports which summarize the test. data and studies of this program and present results applicable to the Shearon Harris Nuclear Power Plant are listed in references 3 to 6. Sections 2 and 3 of this report provide the plant specific technical eva-luations and operability determinations for the SRV's and PORV's of the Shearon Harris Nuclear Power Plant.

In addition to the SRV and PORV concerns, NUREG-0737 requires that the PWR utilities verify the operability of the block valves in;the PORV discharge line. EPRI was also engaged to develop and implement

a. test program to confirm the operability of the block valves.

The results of this test program are contained in reference 7.

Section 4 of this report contains the plant specific technical eva-luation of this test data to determine block valve operability for the Shearon Harris Nuclear Power Plant.

2. 0 SRV CONSIDERATIONS 2.1 SRV INLET CONDITIONS The Shearon Harris Nuclear Power Plant SRV inlet fluid conditions are based on considerations of FSAR transients, transients re-sulting from automatic initiation of the high pressure injection system, and cold overpressurization transients. A discussion of the derivation of the inlet conditions for these cases are con-tained in reference 5. Referring to the reference three loop Westinghouse plant, which is typical of Shearon Harris, this do-cument calculates a maximum pressurizer pressure of 2592 psia for the limiting event of a locked rotor with only SRV steam discharge.

The highest SRV pressure rise rate of 216 psi/sec. is also cal-culated to occur for this event. A maximum liquid pressure rise rate of 4 psi/sec and a peak pressurizer pressure of 2504 psia for the main feedline break was indicated in reference 5 as the li-miting SRV liquid discharge condition for the Shearon Harris plant.

The pressure range of liquid discharge expected for this condition following the initial steam discharge for the Shearon Harris plant extends from 2500 to 2504 psia. The liquid temperature range cor-responds to subcooled water between 620.1 and 623.4 F.

The limiting extended high pressure injection event resulted from the spurious activation of the safety injection syst: em at power, however, no liquid discharge from the SRV's were observed for the analysis of three loop plants. Similarly, cold overpressurization is not a design basis for the SRV's.

TABLE 2.1 Limiting Pressurizer SRV Inlet Conditions for the Shearon Harris Nuclear Power Plant Event Valve Operation Peak Pressurizer Max. Pressure Pressure (psia) Rise Rate(psi/sec)

Locked Rotor SRU Steam Discharge 2592 216 Main Feedline Break SRV Liquid Discharge 2504 4,0 (620.1 to 623.4 F)

• The maximum pressurizer pressure specified on Table 15.3,3-1 of the FSAR for the locked rotor transient is 2610 psia for three loops initially operating, and 2645 for two loops initially operating. The tabulated EPRI calculation results, however, represent a more realistic analysis than the conservative FSAR analyses. For instance, the FSAR analyses uses a conser-vative SRV setpoint pressure of 2575 psia, whereas, a more realistic value of 2500 psia was used by Vestinghouse forthe EPRI report. Additionally, an initial reactor coolant system pressure of 2280 psia was assumed in the conservative FSAR analysis, while a pressure of 2250 psia was used for the EPRI report. In any case the EPRI test pressures bound both the FSAR and EPRI report values.

2.2 SRV INSTALLATION The Shearon Harris Nuclear Power Plant is supplied with three Crosby 6M6 pressurizer safety relief valves each provided with a water loop seal upstream of the valve inlet. A listing of the important SRV information is pro-vided on Table 2.2. An isometric drawing of the piping design immediately surrounding the SRV 's is shown on Figure 2.1. The Shearon Harris SRV's are provided with a 15.28 foot length of 6" schedule 160 inlet piping which contains a 6.1 ft. insulated loop seal. Since the Crosby 6M6 SRV was specifically tested in the EPRI S/RV Test Program, a direct comparison between the test data results and the Shearon Harris Power Plant SRV conditions of operation can be made.

TABLE 2.2 Safety Relief Valve Description for the Shearon Harris Nuclear Power Plant Number of valves 3 Manufacturer Crosby Valve and Gage Type Spring Loaded Safety Valve Model 6-M-6 Design Pressure 2485 psig Design Temperature 650 F Setpoint Pressure 2485 psig Fluid Condition at Relieving Saturated Steam at Setpoint Pressure Developed Backpressure 500 psig Required Steam Capacity 380,000 ibm/hr Rated Steam Capacity 420,000 ibm/hr 5%

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2.3 TEST DATA APPLICABILITY A total of seventeen tests were performed with the Crosby 6M6 SRV with loop seal internals in the EPRI S/RV Test Program as discussed in reference 6. A comparison between the inlet piping configuration of the EPRI tests employing the Crosby 6M6 SRV and the actual Shearon Harris piping is shown on Table 2.3 .

The calculated inlet piping pressure drops for the Shearon Harris and test configurations are also specified on this table. Out of a total of seventeen tests performed, seven were cold (100 0F nominal~)loop seal-steam Casts and four were hot (350 0F nominal) loop seal-steam tests. Since the inlet piping for the Shearon Harris plant will be insulated, the calculated nominal tempe-rature is closer to the hot loop seal-steam tests; however, the Shearon Harris plant characteristics will be compared to all the loop seal-steam tests employing the Crosby 6M6 valve irregardless of loop'eal temperature. Additionally, as indicated in ref. 6, the loop seal-steam tests were performed at essentially two ring positions. The first group were performed at "lowered" ring positions and the second group at ring positions representative of typical PWR plant ring positions.

For all the loop seal-steam tests, the valve initially opened, fluttered and/or chattered at partial lift position during the loop seal discharge, except for test 1415, popped open on steam at pressures of 2580 to 2755 psia and stabilized and closed with a blowdown varying from 5.1 to 9.4/. The range of peak tank pressures of 2582 to 2760 psia measured upstream of the SRV during the tests generally exceeded the calculated peak pressurizer pressure of 2592 psia. The pressure ramp rates for the tests which varied from 3.2 to 375 psi/sec. enveloped the expected Shearon Harris peak pressure rise rate of 216 psi/sec. The range of measured peak back pressure of 227 to 710 psia also enveloped the backpressure range of 552 to 571 psia expected for the Shearon Harris plant.

Three low pressure rise rate steam to water transition tests were also performed with the Crosby 6M6 safety valve as part of the EPRI S/RV Test Program. Two of the tests included loop seals and one test contained a drained loop seal. For the tests with the "raised" typical ring positions, the valve exhibited acceptable performance for 650 F nominal flow. For the 550 0F nominal liquid flow test; however the valve chattered and the test was terminated.

The range of expected liquid temperature in the Shearon Harris plant, 620.1 to 623.4 0F, is closely represented by the test data for the transition flow tests. The pressure rise rate of 2.5 psi/sec and peak tank pressures of 2475 to 2578 psia also closely represents the Shearon Harris plant calculated values of 4.0 psi/sec and 2504 psia.

In conclusion, the EPRI S/RV Test Program conditions for testing the Crosby 6M6 SRV envelope the range of conditions expected at the Shearon Harris Nuclear Power Plant.

E' Table 2.3

.SRV Inlet Piping. Comparisons EPRI 6M6 Inlet Shearon Harris Inlet Piping Configuration Piping Configuration Total Length, in. 160.3 183.4 (including bends)

No. 180 Bends No. 90 Bends No. 45 Bends I. D. in. 6.813/4.897 5.187 Loop Seal Pater 1. 02 0.9 Volume, ft~

Loop Seal Length, ft 7.8 6.1 Transient Inlet Piping Pressure Drop for opening, psi 263 260.4 for closing, psi 181 154.3

TABLE 2.4 Crosby 6-M-6 SRV Test Summary Peak Tank Peak Opening 70 Test Valve Ring Setting Type Pres. Rise Pressure Backpres pop Time Blowdown No. Upper Lower Test Rate(psi) (psia) (psia) (sec) 906 -136 -68 Loop Seal 3.2 2582 554 NA 8.3 908 "136 -68 Loop Seal 297 2688 649 .012 8.3 910 -136 -68 Loop Seal 375 2634 227 NA 7.8 913 -44 -66 Loop Seal 375 2735 242 . 014 7.4 917 -136 .-68 Loop Seal 291 2732 245 . 016 9.0 920 -136 -68 Loop Seal 297 2725 246 .017 NA 923 -186 Loop Seal 283 2736 667 .009 7.7 929 -71 -18 Loop Seal 319 2726 710 .019 5.1 1406 -77 -18 Loop Seal 325 2703 250 NA 9.4 1415 -77 -18 Loop Seal 360 2760 255 . 012 6.2 1419 -77 -18 Loop Seal 360 2675 245 .019 NA 914e.- -44 -66 Loop Seal Trans. 1.1 2516 520 . 011 7.7 926a -186 -68 Transition 2.0 2389 445 . 013 9.4 926d -186 -68 Water(635 F) 1.5 2271 585 12. 0 931a -71 -18 Loop Seal Trans..2.5 2578 725 . 021 12.7 93 lb -71 -18 932 -71 -18 Water(635 F)

Water(550 F) '.02.5 2475 2520 700 650 NA 4.8

VALVE OPERABILITY EVALUATION The Crosby 6M6 safety relief valve tests performed in the EPRI S/RV Test Program is representative of the Crosby 6M6 valves at the Shearon Harris plant for the following reasons:

1) the tested valve and the Shearon Harris valve are identical in model, setpoint and ring position range

2) the Shearon Harris plant calculated inlet piping pressure drop due to valve opening, and pressure rise due to valve closure is represented by the corresponding test values

3) the Shearon Harris plant inlet fluid conditions are enveloped by the test conditions

4) the test valve backpressures envelope the Shearon Harris plant calculated backpressures. However, since the tests have demonstrated that the valve backpressure does not substantially affect valve operation, this item is not regarded as an im-portant consideration.

During the EPRI tests of the Crosby 6M6 valve, the valve was ob-served to oscillate through partial lift positions during loop seal discharge for the loop seal-steam tests. These 170 to 260 Hz oscillations cause water hammer acoustic pressure waves in the loop seal water in the valve inlet piping. For the 900 series tests at the "lowered" ring setting, pressures were measured between 0 and 3400-3600 psia. The 1400 series tests at the typical PWR ring setting indicated a +2450 psi oscillation about a mean pres-sure of 2650 psia with a few individual peaks exceeding 6300 psia.

For the upset and emergency conditions at 300-400 0F, the maximum

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allowable pressure for the Shearon Harris inlet piping was calculated

to be approximately 6500 psia and 9800 psia. Consequently, this piping is acceptable since all the acoustic pressure oscillations measured during the Crosby 6M6 safety valve discharge were below these maximum allowable values under both combinations. A further discussion of this phenomenon as related to the Shearon Harris Nuclear Power plant load combinations is contained in the plant specific safety and relief valve discharge piping evaluation contained in reference 8.

2.4 CONCLUSION

S This section has demonstrated that the Shearon Harris Nuclear Power Plant SRV, inlet piping configuration and fluid inlet conditions are enveloped by the EPRI S/RV Test Program data.

The tests of the Crosby 6M6 SRV demonstrate that the valve would be expected to exhibit acceptable performance during loop seal - steam discharge for the range of ring positions tested.

Valve performance during the steam-water transition and 650 F liquid tests indicated that the valve would exhibit acceptable performance for the typical PMR ring position. Unstable per-formance would be expected for steam-water transition for the "lower" ring positions and for 550 F liquid discharge at the typical PWR plant ring settings. Thus, the EPRI test data confirms the operability of the SRV's in the Shearon Harris Nuclear Power Plant for the expected operating conditions at the typical PWR ring position setting specifiedand determined by Crosby using the methodology equivalent to the method used to obtain the ring position setting of -77 and -18, used for the 6M6 valve tested in the EPRI tests.

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3.0 PORV CONSIDERATIONS 3.1 PORV INLET CONDITIONS The Shearon Harris Nuclear Power Plant PORV inlet fluid conditions are based on considerations of FSAR transients, transients re-sulting from automatic initiation of the high pressure in)ection system, and cold overpressurization transients. Reference 5 con-tains a discussion of the derivation of the inlet conditions for these transients. Referring to the three loop Westinghouse plant, which is typical of Shearon Harris, thip document calculates a maximum pressurizer pressure of 2555 psia for the limiting event of a locked rotor with both safety and relief valve steam discharge.

The highest PORV pressure rise xate of 200 psi/sec is also calcu-lated to occur for this event.

The limiting PORV liquid discharge event resulted from an extended high pressure in)ection event at power. The maximum calculated pressurizer pressure for this event was 2352 psia with a pressuri-zation rate range of 0-12 psi/sec. The valve inlet liquid tempe-rature range is 498 to 502 F for this condition. Inlet fluid conditions for the PORV's for cold overpressure protection are ap-proximately 100 to 350 0 F, and 400 to 2335 psig. The cold overpres-surization system for the Shearon Harris plant controls the PORV's by temperature and pressure instrumentation in order to relieve excess pressure when the reactor coolant system temperature falls below 350 F.

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TABLE 3.1 Limiting Pressurizer PORV Inlet Conditions for the Shearon Harris Nuclear Power Plant II Event Valve Operation Peak Pressurizer Max. Pressure Pressure Rise Rate (psia) (psi/sec)

Locked Rotor SRV and PORV 2555 200 Steam Discharge High Pressure Injection PORV Steam/Liquid 2352 0-12 at Power Discharge (498-502 F)

Cold Overpressure PORV Liquid ~415 to 2350 Protection Discharge (100-350 F) 15

3.2 PORV INSTALLATION The Shearon Harris Nuclear Power Plant is supplied with three Copes-Vulcan 3 inch globe power operated relief valves. A tabulation of the important PORV information is contained in Table 3.2. An isometric drawing showing the piping design immediately surrounding the PORV 's is shown on Figure 3.1.

The Shearon Harris PORV's are provided with 3" schedule 160 inlet piping with about a 5.8 ft. insulated loop seal. The Shearon Harris PORV model is identical to the Copes-Vulcan model tested in the EPRI S/RV Test Program; therefore, a direct comparison between the test data and the Shearon Harris Power Plant operating conditions can be made.

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TABLE 3 ~ 2 Power Operated Relief Valve Description for the Shearon Harris Nuclear Power Plant Number of Valves Manufacturer Copes <<Vulcan Type Power Operated Relief Size and Description Globe with 3" NPS, 316 with stellite plug and 17-4 PH cage Design Pressure 2485 psig Design Temperature 650 F Setpoint Pressure, 2335 psig Fluid Condition at Relieving Saturated Steam at Setpoint Pressure Maximum Steam Capacity 210,000 ibm/hr Required Opening Time 2 sec Required Closing Time 'I 2 sec Normal Air Supply Actuation Pressure 90 psig Backup Nitrogen Actuation Pressure 85 + 5 psig 17

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3.3 EPRI TEST DATA COMPARISONS As part of the EPRI/S/RV Test Program, PORV's were tested at the Marshall Steam Station and at Wyle Laboratories-6/ . Tests were specifically performed on the 3 inch Copes-Vulcan relief valve model (316 w/stellite plug and 17-4PH cage) which is the same model used at the Shearon Harris Nuclear Power Plant. The eleven steam tests performed at the Marshall Test Facility enveloped the Shearon Harris Plant's operating pressure and discharge flowrate. For all tests, the valve opened and closed on demand within the required operating time. Similarly, the nine tests performed at Wyle Laboratories enveloped the Shearon Harris Plant's operating pressure for steam, loop seal, water, and steam-liquid flow conditions. The valve flowrate was enveloped for the one steam flow test. The tested valve successfully opened and closed on demand within the required operating time for all nine tests.

The PORV tests performed at Wyle Laboratories used an actuation pressure of 86 + 1 psig. The normal and backup gas actuation pressures for the PORV's indicated on Table 3.2 are close to the test conditions.

The intent of the EPRI tests was to bound all plant operation conditions.

The small difference between the test and plant specific actuation pressures would not introduce sign'ificant differences in PORV opening and closing times. Since the long opening and closing times of the PORV's preclude the formation of any severe acoustic phenomena, the small differences in opening and closing times resulting from the small variances in actuation pressures would not be expected to affect valve operation or pipe force transients.

Since the tested PORV fully opened and fully closed on demand for each of the twenty tests and since no damage to the valve was observed that would affect future valve performance, the tests confirm the operability of the Shearon Harris Plant PORV's.

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TABLE 3.3 Copes- Vulcan (316 w/stellite and 17-4 PH cage)

" EPRI Test Summary Test No. Fluid Pressure Opening Closing Flow rate (psia) Time (sec) Time (sec) (ibm/hr)

Marshall 1 Steam 2475 1.70 1.60 236000 Marshall 2 Steam 2460 1.70 l. 50 Marshall 3 Steam 2450 1.75 1.50 Marshall 4 Steam 2455 1.65 1.55 Marshall 5 Steam 2465 1.85 1.60'.

Marshall 6 Steam 2460 1.80 50 232000 Marshall 7 Steam 2435 1.40 1.60 Marshall 8 Steam 2450 1.40 1.50 Marshall 9 Steam 2455 1.40 1.60 Marshall 10 Steam 2465 1.70 1.65 Marshall 11 Steam 2460 1.45 1.50 Wyle 71-CV-316-1S Steam 2715 0.60 1.43 255600 Wyle 72-CV-316-3W 446 F Water 2545 0.65 1.31 1008000 Wyle 73-CV-316-4W 442 F Water 675 1.01 0.60 388800 Wyle 74<<CV-316-5W 105 F Water 675 0.98 0.66 619200 Wyle 75-CV-316-6W 274 F Water 2710 0.64 1.44 1173600 Wyle 76-CV-316-2W 647 F Water 2535 0.72 1.38 540000 Wyle 77-CV-316-7S/W Transition 2532 0.70 1.37 540000 Wyle 78-CV-316-8W/W 134 F Water 2725 0.61 1.44 612000

,Wyle 79-CV-316-9N/W N2 1533 0.78 0.88 864000 20

3.4 CONCLUSION

S This section demonstrated that the Shearon Harris Nuclear Po~er Plant PORV fluid inlet conditions are enveloped by the EPRI S/RV Test Program data. Since the Shearon Harris PORV model was identical to the one tested, the ability of the Copes-Vulcan relief valve (316 w/stellite plug and 17-4PH cage) to open and close on demand within the required time for the expected Shearon Harris fluid conditions is confirmed.

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4.0 ELECTRICAL MOTOR OPERATED BLOCK VALVE (EMOV)

4. 1 EMOV INSTALLATION The block valves installed in the Shearon Harris Nuclear Power Plant are Westinghouse valves, Model MOD03003GM99FNH02G, which are motor operated gate valves with Limitorque SB-00-15 operators.

A tabulation of the important block valve information is con-tained in Table 4.1. The isometric drawing on Figure 3.1 shows the piping configuration immediately surrounding the block valves.

Since the Shearon Harris PORV's are located downstream of the block valves, the inlet conditions for the block valves are identical to those discussed in Section 3.1 for the PORV's as obtained from reference 5. The Shearon Harris block valves are similar to the two models tested in the EPRI program reference 7.

Consequently,'omparisons between the test data and the Shearon Harris Power Plant operating conditions can be made.

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TABLE 4.1 Block Valve Information for the Shearon Harris Nuclear Power Plant Valve Information:

Manufacturer Westinghouse Corporation Description Motor Operated Gate Valve Quantity Model MOD03003GM99FNH02G Design Pressure 2500 psig Design Temperature 650 F Max. 4 p for closing 2500 psi Max. 4 p for opening 2500 psi 0 erator Information:

Manufacturer Limitorq'ue Corporation Description Motorized valve operator Model SB-00-15 Nominal Motor Torque 15 ft-lb Full Stroke time to open or 10 seconds close (maximum)

Voltage 460 volts Speed 3600 rpm Wired to limit closure based on valve position 23

4.2 EPRI TEST INFORMATION ON BLOCK VALVES EPRI conducted tests on ~ Westinghouse H)D03000GM88FNBODO elec-trical motor operated block valve mounted horizontally with a Limotorque SB-00-15 operator. 7/ A power operated relief valve (PORV) was located downstream of the EMOV. Pre-evaluation tests were performed with a Rotork operator which successfully fully closed the valve for 10 cycles against full flow at a setting equivalent to a torque of 175 ft-lb. Final tests with the Westing-house supplied Limotorque operator consisted of 21 test cycles of openings and closings of the valve. The upstream pressures before opening ranged between 2345 and 2420 psig and the steam flow rate at a 2310 psig orifice pressure was estimated as 228,000 lb/hr.

The torque setting was slightly above 3.75 corresponding to a .torque of 182 ft<<lbs. The valve opened and closed satisfactorily with zero leak rate. The final evaluation test results can be found in Table 3.2-3 of reference 7, The valve was dismantled and inspected after all the tests. All components weie in good condition except for slight galling on the wedge.

Results from pre-evaluation test operations reported on Table 3.2-4 of reference 7 indicated that the valve did not open or close completely when the torque setting in the Limitorque or Rotork operators were below a setting which produced a torque lower than 175 ft-lb. Hence, the performance of the valve was found to be satisfactory only at an operator setting which produced torques of 175 ft-lb or larger.

7 Using essentially the same test configuration, tests were also con-ducted on a horizontally mounted Westinghouse EMOV model MODO3001GM99FNH02000.

The final tests indicated that the valve fully opened and closed with zero leakage on demand for 21 cycles of the evaluation tests with a SB-00-15 Limitorque operator which was wired to limit closure based on valve position. The upstream pressures before opening ranged between 24

2380 and 2445 psig. The steam flowrate at a 2315 psig orifice pressure was estimated at 234000 lb/hr. As indicated above, the successful tests were performed after Westinghouse supplied a SB-00-15 operator, redesigned the yoke, and rewired the motor operator to close on position instead of torque. The final evaluation test results can be found on Table 3.3-3 of reference 7.

The valve was disassembled and inspected after all testing was completed. There was slight galling of the wedge guides, but all other .components were in good condition.

Supplementary tests reported on Table 3.3-4 of reference 7 of the 99 Series Westinghouse EMOV were performed using a SMB-000-10 Limitorque operator which was set to close on torque. The valve failed to fully close against full flow. Consequently, the valve was shipped to Westinghouse at their request for rework. The valve was returned with a SB-00-15 operator, a redesigned yoke, and rewiring of the operator to limit closure on valve position rather than torque. Subsequent te'sting, following the final evaluation tests, were performed with the SB-00-15 operator wired to close on torque switch setting. The valve failed to close fully against full flow at the maximum torque switch setting of 2.0.

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Table 4.2 TEST VALVE DESCRIPTION EPRI TEST SERIES M-Wsl General Valve Information Manufacturer............................. Westinghouse Corporation Description.............................. Motor Operated Gate Valve Model........................ .... ...... MOD03000GM88FN80DO (88 Series)

Su lied Valve 0 erator Information Manufacturer............................. Limitorque Description.............................. Motorized Valve Operator Model t~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ SB 00 15 Nominal Motor Torque.... ~ ~

~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ 15 f't lb Voltage ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ o o ~ 575 Volts Speed.................................... 3600 rpm Limit closure based on operator torque setting

• Operator modified for a 575 volt power supply Additional Valve 0 erator Information Manufacturer............................. Rotork Description ............................. Motorized Valve Operator Model.................................... 16-NAXl Nominal Motor Torque..................... not specified Voltage ................................. 575 volts Speed.................................... not specified Limit closure based on operator torque setting 26

Table 4.3 TEST VALVE DESCRIPTION EPRI TEST SERIES M<<WS2 General Valve Information Manufacturer................................... Westinghouse Corporation Description.................................... Motor Operated Gate Valve Model.......................................... MOD03001GM99FNH02000(99Series)

Su lied Valve 0 erator Information Manufacturer. ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ Limitorque Corporation Description.. ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ Motorized Valve Operator Model o ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ SMB -000-10 ~

Nominal Motor Torque........................... 10 ft-lb V oltage.'....................................... 460 volts S peed.......................................... 3600 rpm Limit closure based on operator torque setting Modified Valve 0 erator Information Ma nufacturer................................... Limitorque Corporation D escription.................................... lgotorized Valve Operator Model e ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ o ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ o SB-00-15 Nominal Motor Torque................ ...... ~ 15 ft-lb oltage...........................

Vol ~ ......-. 460 volts Speed................,.......,............ .... ~ 3600 rpm Wired to limit closure based on valve position 27

4.3 DESIGN ADEQUACY BASED ON THE TEST RESULTS It is clear from the test results discussed in the previous section that the Shearon Harris block valves could be operated satisfactorily under predicted extreme conditions of temperature and pressure, and also withstand the dynamic loads due to the flow of steam under these conditions. However, the tests performed on both the 88 and 99 series valves do indicate the importance of employing the correct closing method by specifying the correct operator torque switch setting or changing the operator to close on position instead of torque. Subsequent to the performance of the EPRI tests, a change request was received from Westinghouse regarding the Shearon Harris EMOV requesting that the valve operator be rewired to close on position. Therefore, since the Shearon Harris EHOV is similar to the Westinghouse valves successfully tested and does possess the same operator that was successfully employed to operate both the 88 and 99 series valves and has been modified to close on position as indicated in the 99 series tests, it can be concluded that the Shearon Harris .EHOV's will operate satisfactorily.

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5.0 REFERENCES

1. NUREG-0578, TMI-2 Lessons Learned Task Force Status Report and Short Term Recommendations, Nuclear Regulatory Commission, July, 1979

2. NUREG-0737, Clarification of TMI Action Plan Requirements, Nuclear Regulatory Commission, November, 1980 3, Valve Selection/Justification Report, EPRI NP- 2292, December, 1982

4. Test Condition Justification Report, EPRI NP - 2460, January, 1983

5. Valve Inlet Fluid Conditions for Pressurizer Safety and Relief Valve in Westinghouse Plants, EPRI NP- 2296, December, 1982

6. Safety and Relief Valve Test Report, EPRI NP - 2628- SR, December, 1982

7. EPRI- Marshall Electric Motor-operated Valve (Block Valve)

'nterim Testa Data Report, EPRI NP - 2514 - LD, July, 1982

8. Analysis of Pressurizer Safety and Power Operated Relief Valve Discharge Piping for the Shearon Harris Nuclear Power Plant, Ebasco Services, Inc., April, 1984.

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CAROLINA POWER & LIGHT CO.

SHEARON HARRIS NUCLEAR POWER PLANT Analysis of Pressurizer Power Operated Relief Valve and Safety Valve Discharge Piping NRC NUREG OP7 Item 1

II. D.l April 1984

TABLE OF CONTENTS

~Pa e 1.0 Introduction 1 2.0 System Description 3.0 Description of Analysis 10 3.1 Hydraulic Model 10 3.2 Evaluation of Discharge Load on System 13 3.3 Structural Model 13 4.0 Results of Initial Analysis 27 4.1 Hydraulic Model 27 4.2 Stress Analysis 28 5.0 Alternative Discharge Piping Configurations 30 6.0 Final Analysis Results 46 6.1 Hydraulic Analysis 46 6.2 Stress Analysis 147 7.0 Conclusions 155 8.0 References 156 Appendix A - Description of CALPLOTFIII Appendix B - RELAP5 - Error Description

1. 0 INTRODUCTION This report describes the studies performed to evaluate the adequacy of the pressurizer relief piping and its support for the Shearon Harris Nuclear Power Plant as required by NUREG 0737 Item II. D.l.

(1)

The approach taken in the evaluation is consistent with the suggestions contained in the EPRI PWR Safety and Relief Valve Test Program Guide for Application of Valve Test Program Results to Plant Specific Evaluations (2)

Accordingly, the planned approach has been to develop the hydraulic loads on the piping utilizing RELAP5 (3)an EPRI verified computer code, (4) apply those transient hydraulic loads to a structural model of the piping system, compute the response of the piping system, determine the forces and moments at all points in the piping and the reaction loads on the supports/restraints, and finally verify the adequacy of the system by combining the resultant loads on the restraints and stresses in the piping to those computed for other loading conditions and comparing the results against allowable values.

The valve discharge loads on the pressurizer. relief piping system can be induced by the opening of any or all of three power operated relief valves (PORV), the opening of any or all of three spring loaded, self actuated safety valves (SRV) or a sequenced combination of PORV and SRV. Actuation of these valves allows discharge of high pressure steam from the top of the pressurizer into the discharge piping causing. pressure and momentum transients throughout the piping system. These transients create significant time varying unbalanced forces in each straight segment of the piping until steady flow is achieved.

The time histories of the discharge loading are determined throughout the system by employing the hydraulic model described in this report. This model is suitable for execution with the I'ydraulic RELAP 5 computer program.

Forces in each segment are computed by a post processor code,.CALPLOTFIII, described in Appendix A to this report ~ this postprocessor has been written by Ebasco to develop forces from the output of RELAP5 for each of the piping segments.

2. 0 SYSTEM DESCRIPTION The pressurizer relief piping system for the Shearon Harris Nuclear Plant consists of three power operated relief valves, three spring loaded self actuated valves, the interconnecting discharge piping, and the relief tank.

The isometric drawings showing the geometric configuration of the system are attached as Figure 2.1.1 through 2.1.4. Also shown on these drawings are the location of the supports and restraints.

In the event of an abnormal transient causing a sustained increase in pressurizer pressure at a rate exceeding the control capacity of the pressurizer spray, a high pressure trip level is reached which trips the reactor and opens the three power operated relief valves. If the relief provided by the opening of the three PORV is insufficient to limit the pressure rise from reaching the setpoints of the SRVs, the three safety valves will open to relieve the overpressure.

Two cases are considered to be bounding. A third realistic case has been considered but not analyzed since it would not be the design basis for discharge piping design.

The first two cases assume either actuation of all three PORVs simultaneously, or of all three SRVs simultaneously; the third case considers actuation of the three PORVs followed by simultaneous actuation of the three SRVs. For the latter the PORVs are assumed to be opened at time zero. The pressurizer pressure is

"(5) (see Table 2.1) assumed to increase linearly at a rate of 216 psi/sec,

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acPlPWQ Figure 2.1.4

TABLE 2.1 Limiting*Pressurizer SRV Inlet Conditions for the Shearon Harris Nuclear Power Plant Event Valve Operation Peak Pressurizer Max. Pressure Pressure (psia)* Rise Rate (psi/sec)

Locked Rotor SRV Steam Discharge 2592 216 Main Feedline Break SRV Liquid Discharge 2504 4.0 (620.1 to 623.4 F)

Limiting Pressurizer PORV Inlet Conditions for the Shearon Harris Nuclear Power Plant Event Valve Operation Peak Pressurizer Max. Pressure Pressure* Rise Rate (psia) (psi/sec)

Locked Rotor SRV and POR Steam Discharge

'555 200 High Pressure Injection PORV Steam/Liquid 2352 0-12 at Power Discharge (498-502 F)

Cold Overpressure PORV Liquid -415 to 2350 0

Protection Discharge (100-350 F)

• The values for the pressurizer maximum pressure in Table 15.3.3-1 of the FSAR are 2610 psia for a locked rotor transient with 3 loops initially operating and 2645 psia for a lock rotor transient with 2 loops initially operating. The FSAR transient however, has a number of very conservative assumptions which tend to raise the peak pressures. For instance, the SRVs are assumed to conservatively open at 2575 psia as compared with assumed by EPRI report. Additionally, the initial RCS pressure assumed 2500'sia to be 2280 psia for the FSAR analysis compared with 2250 for the EPRI report.

until such time that the pressure reaches 2560 psig at which time the SRVs ~

pop open. That pressure corresponds to the set pressure of the SRVs (2485 psig) plus a three percent margin which the EPRI testing determined to be an upper range of uncertainty. (6)

The three PORVs are located on three separate lines and are connected to the pressurizer nozzle through a common 6 inch line. Three power operated block valves are provided in series with the PORVs to achieve isolation if neces-sary. A 3 inch insulated loop seal is provided upstream of the PORVs. Down~

stream of the block valves the three PORV discharge lines are headered into a 6 inch line which discharges into a 12 inch header downstream of the discharge of the SRVs.

The three safety valves are each independently connected to the pressurizer safety nozzles. An uninsulated water loop seal is provided between the pressurizer and the SRVs; however, a 3 inch insulated loop seal was assumed for the initial analysis since previous experience indicated that an uninsulated water loop seal will give too high hydraulic forces. The 6 inch discharge piping from the SRVs are connected to the 12 inch common header which leads to the relief tank and to which the PORV common 6 inch discharge line is also connected.

The Copes-Vulcan power operated valves are set to open automatically at 2335 psig. Each of these valves develops a flow rate (steam) of 210,000 lb/hr.

The Crosby 6M6 safety valves are set to open at 2485 psig. Each valve has a ASME rated flow rate of 420,006 lb/hr (steam) at the set pressure. As stated above a three percent margin is added to the setpoints of the SRVs for modelling purposes.

3.0 DESCRIPTION

OF ANALYSIS 3.1 H draulic Model The pressurizer relief piping system is modelled as a network of fluid control volumes connected by )unctions. The designation of the control components is given in Figure 3.1.1, wherein each major segment is further subdivided in several subvolumes.

The thermodynamic properties in each control volume and the flow conditions at each junction are computed as a function of time following valve actuation utilizing the RELAP5 computer program. These are in turn inputted in the postprocessor program CALPLOTFIII which computes the forces at each segment of the piping. The designation of the piping segments employed in the CALPLOTFIII model is given in Figure 3.1.2. The initial steady-state temperatures of the water in the SRV and PORV loop seals are calculated using the HEATING5 computer code(')

. Three analyses have been considered corresponding to the following scenarios of valve actuation.

a. Three PORVs open simultaneously. SRV closed (Case A)

b. PORVs do not open. All three SRVs open simultaneously (Case B)

c. PORVs open. Pressurizer pressure continues to increase and SRV's open simultaneously (Case C).

10

RELAP5 HYDRAULIC MODEL OF PRESSURIZER RELIEF PIPING 22 20 15 7 8 13 19 14 18 10 17 46 12

%1 1 23 24 29 30 31 ANCHOR 44 36 38 43 42 41 40 55 52 51 47 48 49 Figure 3.1.1

CALPLOTFIII DESIGNATIONS OP LEG NUMBERS 20 28 19 15 gb 22 16 21 14 25 18 17 24 10 23 8 11 29 12 13 35 56 36 55 32 31 33 42 30 57 43 48 ANCHOR 40 45 60 67 46'7 53 50 62 51 52 66 66 Figure 3.1.2

The computer programs RELAP5 and CALPLOTFIII are described in detail in reference (3) in for the and Appendix A. Opening times PORVs and SRV were. chosen to be 480 m-sec and 9 msec for the Copes Vulcan and the 6M6 representative valves respectively. These correspond to the shortest open-ing time ascertained by the EPRI tests. (6)

The characteristics of the discharge are such that Case C is not a controlling case. Initially the system behaves as Case A until the pressure reaches the setpoint of the SRVs. At such time the SRV open against a much larger back-pressure than that of Case B.

3.2 Evaluation of Dischar e Load on the S stem The system forcing functions were evaluated at each direction change and area change of the piping. The transient forces are computed by use of the post I

processor code CALPLOTFIII, described in Appendix A, which operates on the thermodynamic and flow variables calculated in each volume and )unction of the RELAP5 hydraulic model. Figure 3.2.1 through 3.2.11 show typical force time histories in the piping segments following actuation of'the SRVs with a 3 inch MIN-K loop seal insulation. The direction of the forces is along the axis of the piping segments.

3.3 Structural Model A linear elastic structural model of the piping systems has been constructed utilizes (8) which the computer program PIPESTRESS 2010 for execution.

13

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The PIPESTRESS 2010 computer program performs a generalized response analysis of the system which is sub)ected to the simultaneous application of the

. transient hydraulic loadings in each of the piping segments.

Generalized response analysis is of course known to produce conservative results in terms of both stresses in the piping system and support/restraints loads. It is however economic to run.

A particular feature of this program is that it retains information in memory that enables it to perform selected modal superposition time history analysis of piping segments indicated to be near to or at overstressed conditions, without performing time history analysis of the entire system.

Because PIPESTRESS 2010 automatically prints out stress and stress ratios at each point in the piping system, and is reasonably economical to run, it is ideal for determining whether there is potential for problems in the existing design.

If stress in the piping and reactions on supports/restraints as conservatively predicted for the SRV actuation by the generalized response method, when combined with the previously calculated stresses and reaction loads from the appropriate loading combinations, are below or near the allowable values, then the existing system is demonstrated to be adequately designed for SRV or PARV operation.

In locations where reaction load or piping stress exceed the allowables by a significant margin a reduction in conservatism can be effected by utilizing 25

the selective time history option. If stresses or reaction loads are then computed to be in excess of allowable values, again when combined in the proper combinations with dead, live, and thermal loads, this is a positive indication that the system as presently configured may require modification.

The model utilized by the stress analysis computer code PIPESTRESS is shown on Figures 2.1.1 to 2.1.4.

26

4' RESULTS OF INITIALANALYSIS 4.1 8 draulic Model Results of the hydraulic model for the Case B Transient, i.e. SRVs open but PORVs closed, expressed as forces in selected segments of the piping system, is shown in Figures 3.2.1 through 3.2.11. Therein the leg number refers to

'the segment designation in Figure 3.1.2.

The forcing functions illustrated in these figures are typical of those which are exhibited in the corresponding segments of the other SRVs upstream and downstream piping.

The predicted hydraulic forces develo'ped by the accelerating water slug are very substantial for SRV actuation with 3 inch MIN-K loop seal insulation.

Every precaution was therefore taken to ascertain the correctness of the modelling and thus the validity of the prediction.

Me are confident that the model is correct and that it does properly predict the forces. To illustrate the degree of confidence in the prediction one may compare the forces predicted by the present model to the forces measured by the EPRI tests for a piping segment very similar to that encountered in Shearon Harris immediately downstream of the SRV.

27

4.2 Stress Anal sis Several load combinations must be considered when determining the adequacy of the present Shearon Harris pressurizer relief piping system for discharges from either power operated or safety relief valves.

Of the load combinations recommended by EPRI, (2) which are summarized in Table 4.1, the latter two combinations were not used for the following reasons:

design basis pipe rupture would either void the necessity of relieving or follow the relieving action in time so that the peak loads would not combine, and loss of coolant accidents unless originating in the system (same as pipe rupture event or quench tank relief) would not require relief.

For load combination 3 (i.e. sustained loads + simultaneous actuation of all 3 SRVs), stresses predicted by the generalized response methods excee e ceed allowable values almost everywhere.

Thus, even though the generalized response method overpredicts stresses an d reaction loads, sometimes by a considerable margin, no advantage can be gained by utilizing the selective time history of PIPESTRESS 2010. Instead, additional analyses were performed to obtain the sensitivity of loop seal conditions.

28

TASLE 4.1 Load Coabination for PORV C SRV PLANT/STSVm LOAD SERVICE STRESS LINXT OPKRATINC CONDITION COHbi NATION Class ANSI B31.1~

Upset Loads ~ Relief Valve Transient

'ustained 1.2 Sh Qpeat Sustained Loads + ObE + Relief 1 ~ 8 S~ 1.8 Sh Valve Transient Eaargency Sustained Loads + Safety Valve 2.25 S~ 1,8 Sh Discharge Transient Faulted Sustained Loads + Design basis 3.0 S~ 2.i Sh Pipe break + SSE + Nax (R.V.

DieCharg; S.V. Discharge Transients)

Faulted Sustained Loads + Lose of Coolant 3.0 S 2i4 Q Accident + SSE + Hax (R,V, Discharge; S.V. Discharge Transients) lhte Sh ~ Free TaQe I-l.i of hS% Code Sh ~ 15.9 ksi

• Seismically Designed

5.0 ALTERNATIVE DISCHARGE PIPING CONFIGURATIONS Results presented in Section 4.0 of this report indicated that the existing design of the Pressurizer Relief Valve Discharge piping is substantially overstressed following SRV actuation with a 6.14 ft loop seal insulated by 3 inch MIN-K 1301. Consequently, based on experience with other plants, an effort was begun which considers the effect of the temperature of the loop seal water upstream of the SRV on the magnitude of the downstream piping forces following actuation. Additional thermal hydraulic analyses were performed to obtain the sensitivity of the loop seal conditions for the following cases: the existing loop seal geometry with 6 inches of insulation, a shorter loop seal geometry with 3 inches of insulation, and, as a bounding cases, the existing loop seal geometry without insulation and a drained loop seal containing only steam.

The all steam actuation case is an appropriate case to look at in any case since it is valid situation for a subsequent second SRV actuation. Even if a loop seal is initially present for the first actuation, for a second actuation water in the loop seal would not be present since time is required for a steam condensation to accumulate in the loop seal piping.

The approach to eliminate or drain the loop seal has been taken by other Westinghouse plants, but in order to minimize leakage, the valve seat material should be replaced since it would be in contact with a temperature higher than the recommended value of about 300 to 350 0 F.

30

A comparison of the temperature distribution in the SRV loop seal calculated using the HEATING5 computer code is shown on Figure 5.1.1. The time histories

/

of a typical force (leg 29) for the different cases are shown in Figures 5.1.2 to 5.1.6. The effect of the average temperature in the loop seal on the peak downstream discharge line forces is illustrated in Figure 5. l. 7. This figure shows the peak positive (towards the SRV) force and the peak negative (away from the SRV) force in leg 29. Both positive and negative forces show pronounced decreases until the average temperature of the loop seal reaches 0

about 400 F, however the negative forces show a larger rate of decrease.

Additional stress analyses were performed using the same linear structural model of the SRV/PORV discharge piping system for the PIPESTRESS computer code as described previously. The stress analyses were performed for a drained loop seal, reduced loop seal and loop seal with 6 inches insulation with the existing piping and with the presence of extra strong piping downstream of the valves. The results of the above analyses show no overstressing for the drained loop seal case, one point overstressed for the reduced loop seal and 10 overstressed points for the 6 inches of loop seal insulation with the existing piping. The results for the nodal points, with the largest stresses are presented in Tables 5.1 to 5.3. The results for the loop seal with 6 inches insulation and extra strong piping reduce the number of the overstressed points downstream of the valves but create overstressing problem in the CLASS I piping upstream of the SRVs. A comparison of the values of the original support, restraint nozzle, and valve flange loads with results obtained for the 6 inches insulated loop seal case are contained in Table 5.4.

31

SRV LOOP SEAL TEMPERATURE 0

K D

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~ ~

TABLE 5.1 STRESSES IN SHEARON HARRIS PRESSURIZER RELIEF PIPING WITH SIX INCHES OF LOOP SEAL INSULATION RESULTING FROM NORMAL SUSTAINED LOADS AND SRV ACTUATION NODE PIPING TYPE OF STRESS PSI LOAD STRESS STRESS POINT CLASS PRESSURE WEIGHT SRV COMBINATION ALLOWABLE RATIO 22 ANSI B31.1 3623 973 42963 24311

.. 47559 29148 28620 28620

1. 66 1.02 3105 ANSI B31.1 3623 1214 21 ANSI B31.1 3623 757 26584 30964 28620 1.08 28 ANSI B31.1 4976 115 30775 35866 28620 1. 25 226 ANSI B31.1 3623 141 31859 35623 28620 1.25 229 ANSI B31.1 3623 93 24429 28125 28620 0.98 230 ANSI B31.1 3623 366 22535 26524 28620 0.93 1137 ANSI B31.1 2322 998 47273 50593 28620 1.77 1134 ANSI B31.1 2322 571 41392 44285 28620 1.55 1199 ANSI B31.1 2322 784 38236 31342 28620 1.10 199 ANSI B31.1 2322 421 26136 28879 28620 1.01 1100 ANSI B31.1 2322 477 26157 38956 28620 1.36 1 Assuming 700 psig 2 Using Generalized Response Method 3 Seismically Designed

0 TABLE 5.2 STRESSES IN SHEARON HARRIS PRESSURIZER RELIEF PIPING FOR REDUCED LOOP SEAL VOLUME WITH THREE INCHES OF INSULATION RESULTING FROM NORMAL SUSTAINED LOADS, AND SRV ACTUATION TYPE OF STRESS (PSI)

NODE PIPING LOAD STRESS STRESS POINT CLASS PRESSURE WEIGHT SRV COMBINATION ALLOWABLE RATIO 22 ANSI B31.1 3623 973 23746 28342 28620 0. 99 3105 ANSI B31.1 3623 1214 13325 18162 28620 0.63 21 ANSI B31.1 4976 757 5171 10904 28620 0.38 28 ANSI B31.1 4976 115 16780 21871 28620 0.76 226 ANSI B31.1 3623 141 20948 24712 28620 0. 86 229 ANSI B31.1 3623 93 17198 20914 28620 0. 73 230 ANSI B31.1 3623 366 16260 20249 28620 1137 ANSI B31.1 3623 998- 17889 22510 28620 0.79 c

1134 ANSI B31.1 2322 571 29497 32390 28620 1.13 1199 ANSI B31.1 2322 784 16605 19711 28620 0.69 199 ANSI B31.1 2322 421 7105 9848 28620 0.34 1100 ANSI B31.1 2322 477 9516 12315 28620 0.43 1 Assuming 700 psig 2 Using Generalized Response Method 3 Seismically Designed

TABLE 5.

STRESSES IN SHEARON HARRIS PRESSURIZER RELIEF PIPING WITH A DRAINED LOOP SEAL RESULTING FROM NORMAL SUSTAINED LOADS AND SRV ACTUATION NODE PIPING TYPE OF STRESS (PSI LOAD STRESS STRESS POINT CLASS V2 COMBINATION ALLOWABLE RATIO PRESSURE WEIGHT S 22 ANSI 3 B31.1 3623 18045 22641 28620 0.79 3105 ANSI 3 B31.1 3623 1214 9647 14484 28620 0.51 21 ANSI 3 B31.1 3623 757 10938 15318 28620 0.54 28 ANSI 3 B31.1 4976 115 11990 17081 28620 0.60 226 ANSI 3 B31.1 3623 141 15075 18839 28620 0.66 229 ANSI 3 B31.1 3623 93 13069 16785 28620 0.59 230 ANSI 3 B31.1 3623 366 11301 15290 28620 0.53 1137 ANSI 3 B31.1 3623 998 13259 17880 28620 0.63 1134 ANSI 3 B31.1 2322 571 15839 18732 28620 -0.65 1199 ANSI 3 B31.1 2322 784 10474 13580 28620 0.47 199 ANSI 3 B31.1 2322 421 4691 7434 28620 0;26 1100 2322 477 6357 9156 28620 0.32 1 Assuming 700 psig 2 Using Generalized Response Method 3 Seismically Designed

Table 5.4 Support and Restraint Loads for SRV Actuation with Six Inch Insulation on Loop Seal

• Load*

Original Load New FOINT FORCES IN tOUNSS -. FORCES .IN NOUNS)

NO>> FX fY f? fX fT fK RESTkAINT 2202 0% 0% 19d5% 0>> Oe 9909 RESTRAINT 2161 0% <<103$ ~ 0% 0 692 0L RESTRAINT 280 0% 0% 799e 0 0>> 11415%

RESTRAINT 2141 0% 479 0>> 0>> 1533 ~ 0>>

kESTRAINT 2213 0%

~ 0% 0>>, 2043 0 RESTRAINT 2213 0% 0% <<$ d6>> 0 0% 7702%

'00~

RESTRAINT 1113 26% 0% 0% 7381 ~ 0>> 0>>

RESTRAINT 11?4 O>> 0% <<9 0 33$ 6 ~

RESTRAINT 0% e>> 16% 0 1840 SNUBBER 1$ 0>> n. 1703% Oe 0~ 8859%

SNUBBER 1504 d631 ~ 5040% 0>> $ 91 d 449d t .. 0>>,

SNUSSER 1503 3934% 7463 ~ 0>> 4528% 8589>> 0>>

SNUBBFR 1718 0~ 9?47% 974'F e Oe 110Fde 110?d>>

SNUBBER 2510 0% 0>> 4d041 ~ 0 t ....- 0 o 71d$ 7%

82587 t SKUBBEk 230'5 0% 40412% 0% Oe Oe SNUBBER 12$ 0% 0>> $ 20% 0% 0 19355%

SNUBBER 2$ 3488 d>> 0% 0% 5'1d33 t 0% 0>>

SNUBBER 48 0% <<4?0% 1830% 0 276$ ~ 10761 ~

SNUBBER 2$ 1 Oo 845 d n. 0% 1339d>> 0>>

SNUBBER 4803 d9'1 Be <<$ 258% 0% d273 476? 0t SNUBBER 4801 3750>> 711 3% 0>> 3?20>> 70$ d>> 0~

3103 d114>> 4647% 0>> 8089% d1 48 ~ 0>>

SNUBBER SNUBBER SNUSSER 3101 2232 4229>>

0%

8022%

414%

0%

0%

4F41 0>>

~, 8993>>...

12478 ~

0>> .

0>>

SNUSSEk 2231 <<47?% 0% 0% 19032>>

SNUBBER 2222 0% 0% d4 ~

SNUBBER 813 0% 36036% 0% 0>> 36974 2201 dd06>> 0>> n. 18059 0 SNUBRER SNUBBER SNUBBER 2219 2220 16820%

0% 1027 0% 2QO37%

0>>

13456 0>>i'.dd 0%

0% 14030%

0 SNUBBER SNUSSEk SNUBBER 1123 1121 1118 n.

<<33 ~

n>>

38>>

0~

'7>>

0%

0%

0%

574500>>

0~

t 793?e 6794 ~

0%

0 0>>

SNUBBER 11?2 23% 0 0>> 6762>> 0>> 0>>

SNUBBER 199 e>> <<4>> 0>> 0 11290 0>>

SNUBBER 9161 e>> Q>> 1% 0 0 8511 ~

SNUBBER 11$ 3 0>> 4~ 0>> Oe $ 530 ~ 0%

SNUBBER 9192 21$ ~ 0>>' 0~ $ 2h8 Q h>>

SNUBBER 1161 4% ~ 0>> d 480>> 0 ~ 0~

SNUBBER 11db 0 6>> 24>> 0 0 16729 '

In Global Coordinates

Table 5.4 (Cont d)

Support and Restraint Loads Eor SRV Actuation with Six Inch Insulation on Loop Seal Po lNT Ogi gina l.

fORCES IN POUNSS Load New Load FORCES XN f11IIISI.

NO>> FX FY FZ FX FY FK

!NUBBER 1165 0>> 10>> 0>>

<<46>>

0 . SL19>> .. 0 5NUBSER 119$ 0>> 0>> 0 3d60 SNUBBER 119$ Oe 4r. 0>> 0 1140d>> 0 SIIUBBER 1198 Ce 0~ SC>> 0>> . . Oe .. R498 ~ .

SNVesER 1180 0>> -20>> 0>> 0>> $ 230>> 0>>

5NUBSER 1174 0>> 51>> 0>> 0 10391 ~ 0 SNUBBER 1123 0>> 0>> 21 ~ Oe. 0>> 4070>>

SNUBBER 1526 0>> 0>> 17>> O. 7470.

SNUBS'ER 1127 0>> 6>> 0>> 0>> 4$ 1d>> 0>>

SNUBBER 1131 Oe 114 'e 0>> Oe .. 2530 ~ Oe.

SNUBBER 220 0>> 9242>> 0 19230>>

SNUBBER 31 0>> 0>> 3173>>

0~M5~

0>> 0>> 1010$ e VAR SUI'PORT $ 01 0>> 0~ 0>> 0~

VIR SUPPORT 12$ Oe 0~ 0>> Oe 14 ~ 0~

VIR SUPPORt $ 71 0>> Oe Oe 0>> 33 VAR SUPPORT 401 0>> 0>> Oe .0 e de .....Oe.

'e

~

VIR SUPP SRt 1$ 0>> 0>> 0>>

VIR SUPPORT 45 0>> 0>> 0>> Oe 33

'VIR SUPPORT 31 G>> 0>> 0>>

VAR SUPPORT 2232 0>> 0>> 0>> 0 VAR SUPPORT 2219 Oe 0>> 0>> 14>> 0>>

VAR SUPPORT 227 0~ 0>> 0>> 0 30 VIR SUPPORT 1 127 0>> 0 0>> .Oe .. 0>>.

VIR SUPPORT 1118 0>> 0>> 0>> 0 10 ~

VAR SUPPORT 117$ 0>> 0>> Oe 0 20 0>>

1196 0>> 0>> ..0>> 'lre.. Oa VAR SUPPORt VAR SUPPORT 1191 0>>

Oe 0>> 0>> 0 0

21 27>>

' 0>>

~

VAR SUPPORT 9166 Oe Oe Oe VIR SUPPORT 11$ 0 0~ O.

'0>>

0>>

9772>>

0>>.........5 S i......O>>.

SNUBBER ross 2091 ~ 0>> 0>> Oe 5'IUBBER 70$ 1 0>> 960 Oe 0>> 9849 0>>

SNUBBER 7041 0>> <<804>>

<<333>>

0>> .. 0>>..

7 47 7>>

105lsi....

0>> 1434 4>>.

SNVBBER 7041 1736>> 0>> ~

7C31 0>> <<28 6>> 0>> 0 125 d2 ~ 0 SNUBBER

'SNUBBER VIR SUPPORT 7031 301 2430e 0>> 1$ 43 ~ 12307>> ......0 ~ ... 7556 ~

• In Global Coordinates

Table 5.4 (Cont'd)

Nozzle Loads for SRV Actuation with Six Inch Insulation on Loop Seal Original Load*

MOZSI. - .. O'NHL.lN ENNY .=.= -".. IIQNBI$.& FSi ElllSS =

FZ NX NV IIZ NOZZLE 1 962>> 1387>> <<741>> 2'3>> 1333>> <<2198 ~

NOZZLE 45 <<440>> <<198 <<814>> <<3419 ~ 933 ~ 1275 NOZZLE 62 <<485 ~ 258 68>> 161>> 22d4>> '0dd ANCIIOX <<38033>> 35>> 3580 1587 421 ~ 176 212'll NOZZLE <<0>> <<0>> <<1 ~ <<2>> 1>> <<1>>

Hew Load*

tOZIIT flII fOl)NQ f gO~OjLNOS mmmmm 9 OX~CD Z mmmmm mmmmm NgNENQ gg NO>> fX fZ IIX III Ilz NOZZLE 45 7038. 14195. 6323. 17902. 12S61. 8844.

NOZZL E 62 $ 34S 11038 3'7d7 9482>> 10943 1 54 52>>

ZII1999 212 92932~ 411 ~ 311, 2229~ZZ4 194'I NOZZLE 111 4093. 7924. Sd7$ . 1S072. SF99. 18419.

• In Global Coordinates 44

Table 5.4 (Cont'd)

Valve Flange Loads for SRV Actuation with Six Inch Insulation on Load Seal Original Load POINT PORCES IN POUNDS MOMENTS IN FOOT POUNDS NO. PX FY MK MY SRV FLANGE 13 3392 -1608 802 -345 4366 5041 SRV FLANGE 1501 -3392 1608 -802 829 -3748 -5850 SRV FLANGE 50 1408 -513 -313 47 -937 2070 SRV FLANGE 4800 -1408 513 313 -159 ~ -194 -2705 SRV FLANGE 33 1941 -1160 -564 267 2230 -1754 SRV FLANGE 3101 -1941 1160 564 -418 -2423 2304 New Loads POINT FORCES IN POUNDS MOMENTS IN POOT POUNDS NO ~ FY MK MY MZ SRV FLANGE 13 12264 34049 9161 4332 11133 12010 SRV FLANGE 1501 35566 9388 5490 5015 9557 12157 SRV FLANGE 50 6335 35338 7233 6681 9711 8350 SRV PLANGE 4800 23813 7602 5658 4567 9486 8890 SRV FLANGE 33 7520 35658 7874 4730 12572 11995 SRV FLANGE 3101 20237 8689 6035 5360 10862 15153

• In Global Cooxdinates

6. 0 FINAL ANALYSIS Following the stu d y o f a lte rnative loop seal configuration and loop seal insulation the option to maintain the existing SR V seal with 6 inches of MIN-K insulation (or its equivalent) f and modi y in g the system by changing some snubbers and restraints was chosen. This approach ach results in the least amount of changes, and hence the most economical.

6.1 H draulic 'Anal sis Results o f th e h y draulic model for the case A transi e n t , i.e. PORVs open but SRVs closed, expressed as time histories o f the h draulic forces are shown in Figures 6.1.1 through 6 . 1 . 48 Thee PORV loop seal has 3 inches of MIN-K insulation and the temperature gradient in the loo o p seal is shown in Figure 6.1.49.

The results for the case B transient, i.e.e. SRVss o p en but PORV closed expressed as hydraulic forces in segments of the piping system are shown in Figures 6.1.50 through 6.1.99. The temperature gradient in the SRV loop seal with 6 inches of MIN-K insulation as calculated by HEATING5 is shown in Figure 5.1.1.

The back pressure in the piping system which ex exists as the SRV discharge has reached steady state flow conditions are summ arized in Table 6.1.

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TABLE 6.1 SHEARON HARRIS STEADY STATE BACKPRESSURE AND FLOW RATE VALVE NUMBER DOWNSTREAM FLOW RATE CALCULATED FLOW RATS PRESSURE (PSIA) (ibm/hr) SRV RATED FLOW RATE 4 (R530) 568 499,320 1.19 14 (R529) 571 499,320 1.19 20 (R528) 552 499,320 1.19 29 (P529) 417 36 (F528) 413 40 (P527) 413

• SRV Rated Flow Rate is 420,000 ibm/hr

• + The maximum flow rate is the ASME rated flow rate times a margin of 1.15 to 1.225 (9) . EPRI actual test (6) measurements, show a maximum flow multiplier of 1.18 for Crosby 6M6 valves.

6.2 Stress Anal sis "Based on the recommendation to maintain the existing pipe geometry and to reduce the discharge loads by increasing the loop seal insulation, a stress analysis study was performed to determine the minimum number of support revisions required to meet the EPRE recommended pipe stress limits. The study considered the combined effects of sustained loads, OBE, SSE, PORV discharge and SRV discharge utilizing the stress limits of Table 4.1.

To assure minimum design changes, the generalized response analysis technique was used to economically give some guidance indicating promising solutions. However, once stress levels were reasonably reduced, the modal superposition time history technique was used to reduce stresses further at the remaining overstressed points. The study indicated that a satisfactory support arrangement would require placing additional snubbers at the following points (refer to Figures 2.1.1, 2.1.2, 2.1.3, and 2.1.4):

Point Number Snubber Function 8134 East/West Direction 2305 North/South Direction 9132 North/South Direction 2202 North/South Direction 9131 Vertical The study also indicated that the added snubbers permitted the following snubbers to be deleted:

147

Point Number Snubber Functioi'.

1718 Axial 1171 Vertical 6 Vertical 2222 Lateral Based on this revised support scheme, points of high stress in the seismically designed B31.1 downstream portion of the piping are summarized in Tables 6.2.1, 6.2.2, 6.2.3, and 6;2.4. Stresses were evaluated for Load Combinations I through,4 of Table 4.1 except that as discussed in Section 4.2.. Stresses due to Design Basis Pipe Break and Loss of Coolant Accident were not considered. Support loads for the SRV and PORV actuation cases are shown in Table 6.2.5.

Support loads were transmitted to the support/restraint design group for evaluation and all necessary redesign has been completed. Addition and deletion requirements identified above have also been included in the design modifications.

I 148

TABLE 6.2.1 High Stress Points ln Seismically Designed Downstream Portion of Piping Load Comblnatlon: Sustained Loads + Safety Valve Discharge Transient COHBINED STRESS STRESS ALLOWABLE (I.8Sh)

NODE POINT STRESS RATIO (pst) (ps I) 1505 261478 28,620 .925 1502 25,505 28,620 .871 3106 23,794 28,620 .831 4805 23,471 28,620 .820 226 26,217 28,620 .916 9126 22,980 28,620 .803 NOTE: Stresses from General Response Analysis.

Time history would indicate lower values.

TABLE 6.2.2 Rlgh Stress Points in Seismically Designed Downstream Portion of Piping Load Combination: Sustained Loads + Relief Valve Transient NODE POINT COMBINED STRESS STRESS ALLOWABLE (1.2Sh) STRESS RATIO (ps i) (psi) 21 17,140 19,080 .898 28 17,286 19,080 .906 226 18,376 19,080 .963 228 16,249 19,080 .852 9131 15,589 19,080 .817 NOTE: Stresses from general response analysis except pt. 226.

Time history analysis would indicate lower values for other points.

I 4l

TABLE 6 7 3 High Stress Points in Seismlcaily Designed Downstream Portion of Piping Load Combination: Sustained Loads + OBE + Relief Valve Transient NODE POINT COHBINED STRESS STRESS ALLOWABLE (I.8Sh) STRESS RATIO (psi) (psl) 226 26IO4 2862O ;912 9l26 25 I4 I 2862O .878 I 199 28544 2862O .997 NOTE: All stresses from Generalized Response Analysis.

Time history analysis would indicate lower values.

TABLE 6.2.4 High Stress Points in Seismically Designed Downstream Portion of Piping Load Combination: Sustained Loads + SSE + Hax {R.V. Discharge; S.V. Discharge Transients)

NODE POINT COHBINED STRESS STRESS ALLOMABLE {2'4Sh) STRESS RATIO

{psl) -

{p I )

1502 36,553 - 38))6Q .958 21 36,248 38,160 .950 47 37,953 38,160 .994 48 37,644 38,160 .986 9131 31,818 38,160 .834 9126 34,707 38, 160 .910 I I99 33,018 38,160 .865 NOTE: Stresses from General Response Analysis except pts 21, 47, 1199.

Time history analysis indicate lover values for other points.

TAg LE 6. 2.5 Support and Restraint Loads for PORV and SRV Actuation PORV Actuation Sl'V Actuation NT os C ow&pQ RESTkAlHT RROR 0~ 0~ 44IS ~ kf'STRALNT. RROR 1 - Oo 0~ tllV ~

R TRA NT V I I H kCSTkAlNT 4o RESTRAINT R lVL 0~ 10708 '

. RLVl 0~ 4ROio Oo ~

Rf STRALHT RRls Oo 0~ 170Vo RCSTRAlNT ills Stll ~ Oo 0~ RESTRAoHT 1 1 LS d't70 ~ 0~ 0~

SNWICR L4 0~ 0~ LVSt o SNQIIER 14 0~ 0~ VSSE N ICR OV 7RRR 0 0

~ NWIER Rill ~ Oo 0~ VttV~ SHUISER Rdli Oo Oo ltltl o N

SNWIER RI ltdllo Oo 0~ SNuaafk R4 SVSSRo Oo 0~

ISHWIE k NU I

VIOS lOll ~ 41VRo 0~ SHUSIE VIOS . 1V740 ~ 1 1 allo 0~

I SNUSSE R SLOS

'o ltllo SNWICR ~ SLOL VS47o IRIS ~ 5HUSIER 5101 170dL ~ ' ~

NU88 SNWICR . ~ LS 0~ ROStO ~ 0~ SNQSSER 815 S4087 ~ 0~

N VVT U

INuIICR RRRO tlSOo Oo ldllio SHU88ER RRRO 15407o 0~ liRLO~

~ NU SNWIER 1 ill Oo Rtlo sNueafk Oo Rllo Oo sNuea SHWICR t141 Oo Oo RR74o SHUSSER t141 0~ 0~ 1747o 7

HH-VNWIER llil - 177Vo 0~ 0~

NQ sHuaefk 1141 1RS7o 0~ 0~

WU SNU88fR lltd 0~ 0~ 'VRl o 5NUAlof R 1195 Oo 0~ 1771 ~

TABLE 6.2.5 (cont'd)

Support and Restraint Loads for PORV and SRV Actuation PORV Actuation SRV Actuation 5NUQQKR 1140 Oo 700o 0~

SNUQQKR SNUQQCR 1 ill 1140 Oo 0~

Oo Tito 407o 0~

II 8 SHUQQCR 1124 0~ Oo $ 75 ~ SNUQQKR 1124 0~ 0~ 514o SNUQQ SHUSQCR 1151 220 0~

0~

V4V ~ SNUQQKR SNUQQKR 1131 . Oo 444 ' ~

Oo 5'V44 ~ 220 Oo

~o 0, 1VSSS ~

Y 5 0 to Y VAR SUPPORT 124 0~ 4o VAR SUPPORT 124 0~ 0~ 15 ~ Oo Y V VAR SUPPORT VS 0~ 5~ 0~ VAR 5UPPORT 0~ 5~ 0~

YAR SUPPORT 221't Oo 5~ 0~ VAR SUPPORT 22lt 0~ 12o 0~

V Y YAR SUPPORT 1115 0~ 5~ VAR SUPPORT 1115 Oo 0~ To 0~

VAR . Y 4 VAR SUPPOPT 1 it! 0~ 10o 0~ VAR SUPPORT 1 1t 1 0~ llo 0~

YAR 5 ll 0 ~ ~ Y V

SHUSQCR 7041 14lio 0~ tlVo SHUQQKR 7041 '020 ~ 0~ Sllio SNUQQ Ov NUQ8 V V~

5NUQQCR 10vl litle Vllo 220 ~ SHUQQCR 10V1 424t ~ 157t ~ 102V ~

jN VAR SUPKNT 14 0~ 0~ VAR 5UPPORT 14 0~ 1~ 0~

NUQ 5 NU SNUQQCR tl32 504v. Oo Oo SNUQQKR t152 VOdl ~ 0~ 0~

The temperature of the loop seal water upstream of an SRV or PORV affects the magnitude of the discharge piping forces and, consequently, the stresses following actuation.

In order to meet EPRI required pipe stress limits, an approach was taken to minimize the amount of changes to the discharge piping system.

The existing pipe geometry was maintained, but the loads were reduced by increasing the SRV loop seal temperature through the addition of insulation. The minimum required average temperature of the SRV loop seal was found to be 370'F, which can be achieved by either six inches of MIN-K insulation or a metal reflective thermal insulation box or equivalent.

An extensive stress analysis study was performed to determine the minimum number of support revisions and resulted in adding five snubbers and deleting four, such that overstressing was eliminated.

Evaluation of stresses in the Class 1 portion of the piping, pressurizer nozzle loads, and valve end loads are the responsibility of Westinghouse Electric Corporation as part of their contract to perform the Class 1 analysis for this piping. However, an Ebasco preliminary stress analysis indicates a high probability that the revised scheme will be acceptable.

155

8,0 References NUREG 0737: Letter, from D G Eisenhut, NRR, USNRC, October 31, 1980

2. EPRI PWR Safe and Relief Valve Test Pro ram - Guide for A lication of Valve Test Pro ram Results to Plant S ecific Evaluations, Interim Report, Rev. 2, July, 1982.

3~ Ransome V H, Wagner R J et al, RELAPS H3d 1 Code Manual, Vol 1 & 2, EGG Idaho Inc. NUREG/CR 1826, EGG 2070 Draft, Revision 2, Sept. 1981 4~ A lication of RELAP5 MOD1 for Calculation of Safet and Relief Valve Dischar e Pi in drod mic Loads, EPRI NP- 2479, December 1982,

5. Valve inlet Fluid Conditions for Pressurizer Safe and Relief Valves

\

in Westin house- Desi ed Plants EPRI NP - 2296, December 1982.

6.

EPRI NP- 2628- SR, December 1982 7~ Turner W D, Elrod 0 C, Siman- Tov I I, HEATING5-An IBM Heat Conduction Pro ram Union Carbide Corp., Nuclear Division ORNL/CSD/TM-15, March 1977

8. User's Manual for Pi e Stress Anal sis PIPESTRESS Pro ram 2010, G Cohen, J Chesler, Ebasco Services Inc., June 1979

9. Minutes of EPRI Piping Subcommittee Meeting, May 1982.

156

40 CALPLOTFIII DESCRIPTION

Mathematical Model The CALPLOFFXXIcomputer code has been vrftten to convert the transient flov conditions calculated in a piping system by the

NKLAPQSD 1 Computer code into transient forces on the pfpiag system. Specifically, CALPLKFIITcalculates and plots the forces on straight lengths of pipe bctvecn changes in pfpe direction (bends) ~ or betvecn a change in direction and a pipe brcak. The derivatfon of the equations used ia the code are gfvcn bclov.

Strai ht Len the of Pf ee Bctvccn Directional Chan ce The force on a straight length of pfpe betveea direction changes (Figure l) fs calculated using thc moment+a equation.

7 + f/('pdv V (pV ~ d7) +~t V (pdv)

JJJ cv cs cv If the gravity term fe assumed negligfble, the follovfng equation reeults:

F s

~ V (pV ~ dA) + ~t V (pdv) ( 2) cs CV Since the force on the straight pipe length only exists ia one dimeasioa, the above equation can be vrittea ia a scalar form.

F s

~ V (pV ~ dA) + Bt Vpdv ( 3) cs CV Since thc RELhP5 M)D 1 Computer code calculates the pressures and the flovrates at different physical positions ia the piping system, it is accessary to subdivide a piping length into tvo control volume types for application of the moment+a equation. Thc first dfvfsion creates the pressure control volumes. The divfsions for the pressure control volanes are the positfons in the pipe length vherc the pressures are calculated by the computer code, and serve as the boundarfes across which thc control volume surface forces are calculated. The second control volume divisfons are duc to flov conditfons. The boundaries of the flov control volumes are located at the pipe length locations vhere flovs are calculated by the computer code. The forces in the pfpe length vhich are due to the rate of efflux of momentum across a control volume and the change of moment+a ia a control volume are calculated using the flov boundaries as flov control volume divisions.

The resultant force on the fluid across" the boundary of tbc pressure control volumes 1 ~ '2 ~ and 3y Shown in Figure lj arct FSl

> (P] P ) k~ f I FS2

~ P~ h~ PB k + P (h, A~) + I FS3 (Ps P ) B

+

The net surface force on the straight pipe length is obtained by smaning equations 4, 5, and

~sr + 'sz + 'ss 5 + 6 + Rs F ~ R S

Therefore, the force on the straight pipe length duc to surface forces is equal to the net normal and shear stresses on the pipe wall length, Thc right side of equation .3 can now be evaluated for each of the flow P

Sl control volumes h and B-P2 V2 g

2 2

h~k + ~t 8H~

bh g

+

~ hB ( 10)

Since the RELAP5 computer code calculates non thermal equlibrtum conditions for two phase flow conditions and allows the two phases to possess different velocities, the parameters of equations ( 9), ( 10) are defined as:

8p 0.5 P catering 8 + 0.5 cx t ng

$I 0.5 R + 0.5 8 gl gl 1 12 '12 ( ' 2' ,2 'g2

~ 0.5 8 + 0.5 8 (12) 2 2 p2V2 p12V12 (1-a2 + Pg2V 2 a2 (13) 2

SNnfng equations '9 and 10, and using equation '8, the net fluid force on the pfpe length can be obtained:

( zr)

If the straight length of pipe consfdered fs bounded by a directianal change and an open andy a break, the forces obtained using equation 11 aust bc modiffed ta account for the force developed at the pipe exit plane. Consequently, using the momentum equation, thc force an the straight pipe length shovn an Ffgure 2, for unchoked brcak flav, can bc vrftten as:

2 Mh unc i ( 15)

If choked break flov is dctermincd to exist by thc fluid transient computer cade, then equatfon 1S must be modiffed to account for the prcssure unbalance that occurs at the pfpe cxft plane. h rederivstion of the equation for the straight pipe length for thfs case results in the follovfng relation:

( 16.3 ar Xh K (P2 P) h ( lV) vhere:

2 2 P2 Pi + ~g 2

p2 =P h +PAL r22 5A ah 2h ('19)

( 20)

. Ncccnclature floe area body for~e of a control volume F~ surface force resultant on a controL volume I gravitational constant x force of fluid on piping control volte flovrate prcssure P pressure outside pipe control volumes R normal and shear stresses in a control volume time volume of a control volume v velocity of fluid in a control volume Greek Letters 0 density in control volume a void fraction Subscripts acceleration friction choked flov control surface

Qnc I

control volume elevation unchoked Liquid

Pisure 1 VOL 2

( + FORCE LEGEND:

~e~o ~ PRESSURE BOUNDARY

---

FLOW BOUNDARY

Pituro 2 VOL 1 VOL. 2

+ FORCE A

LEGEND:

PRESSURE BOUNDARY

~ FLOW BOUNDARY

RELAP5 CALPLOTFIII APPLICATION

SUMMARY

An error has been discovered in the RELAP5 definition of volume average velocity. This problem is discussed more fully in Appendix B. Besides being used internally in the RELAP5 code, the volume average velocity was also used by CALPLOTFIII in the pipe force calculation for the calculation of the acceleration term and to determine the pressure at a pipe break for the calculation of the unbalanced pressure term. Following receipt of the letter from Ms. E. Johnson of EG6G described in Appendix B, the CALPLOTFIII code was modified to use junction flows for the calculation of pipe forces.

Specifically, the momentum equation solution has been modified using equations 11 12 and 20 shown of pages A-2 and A-3. The modified CALPLOTFIII code ha~

been used to predict the experimental results for the subcooled experiment( (1, on pages A-9 and A-10. 'hown Since the unchoked comparison severely overpredicts the pipe forces, it is concluded that the choking option must be employed at the break location.

Even after considering the presence of choking at the break location, the code still overpredicts the break forces. Since the test geometry used two rupture disks to simulate a break, two additional runs were made. One employed a discharge coefficient of 0.8 at the break with choking, and the other assumed the flow area at the rupture disk to be 0.8 of the fu11 area and included a 0.8 discharge coefficient at the break. The latter of these two runs most closely predicted the test data leading to the conclusion that the rupture disks partially obstructed the full flow area. The additional oscillations in the measured force of the test data is probably attributable to the flexi-bility of the piping system tested. Since the RELAP5/CALPLOTFIII calculations calculate only hydraulic forces, a stress analysis code should be run using the hydraulic forces as an input in order to account for the piping and support stfffnesses. Since this additional analysis was not performed, only the initial peak force and the long term force are representative of both the hydraulic and structural loads. The forces at both these times are most closely repre-s ented by the analysis results using the partially obstructed rupture disks.

(2,3)

Similar calculational methods have been used in the EPRI S/RV test program to predict pipe forces following SRV actuation. These calculations have compared favorably to the S/RV test data. Consequently, it can be concluded that the RELAP5/CALPLOTFIII calculational method will predict conservative hydraulic forces.

A-7

(1) RELAP5/CALPLOTPIII Verification, Ebasco Applied Physics Calculation 3-E-8 No. 001, May, 1983.

(2) A lication of RELAP5/MODl for Calculation of Safet and Relief Valve Dischar e Pi in H drod amic Loads, Electric Power Research Institute, NP-2479-LD, December, 1982.

(3) A lication of RELAP5/MOD1 for Calculation of S/RV Dischar e Pi in H drod amic Load's in the CHINON Bl PWR Electric Power Research Institute, NP-2927-LD, March, 1983.

(4) REFORC A Co uter Pro ram for Calculatin Fluid Forces Based on RELAP5 Results, Electric Power Research Institute, PWR Safety and Relief Valve Test Program, EDS Report No. 01-0650-1194, Rev. 0, February, 1982.

A-8

FIGUR RELAP5 - MODEL OF SUBCOOLED BLOWDOWN EXPERIMENT COMPONENTS SINGLE VOLUME COMPONENT 1 p ~ 14.7 paia SINGLE VOlUME AIR AT70O F 24 IN. SCH 60 p ~ 923 pea auaLITV -O.o RUPTURE DISK (BREAK) COMPONENTS JUNCTION 4000000 VOL 3420000 153 pele ~ SS33 JUNC 3410000 AN AT FT.

PIPE COMPONENT 3 12' 4 IN. SCH 80 923 poh 1lI833 OLTY. FT.

VOL 3410000 ~ 0.0 40 PIPE VOLS - 1 FT. LENGTH EACH - VOLS 3010000 TO 3400000 P-923psla-auALITV 0.0 COMPONENT 2 JUNC JUNCTION 40 FT. 3400000 2000000

Figure 4 SUBCOOLED BUNDORN TEST COMPARISONS TEST 1 TEST 3 14 RELAP5 - NO CHOKING RELAP5 - WITH CHOKING 10 ir RELAP5 - CHOKlNG - .ECd ilh f I I RELAP5-.8 DISK AREA NOTE: SCALE CHANGE 0

0 10 20 30 40 50 60 70 80 90 100 200 300 400 500 600 TIME (M SEC)

APPENDIX B RELAP5 - ERROR DESCRIPTION

,During the process of implementing and verifying the RELAP5 Nod I computer code, an error has been uncovered regarding the definition of the volume average velocities. Consequently, our concerns regarding this problem were conveyed to D. Hall of EG&G in a telephone conversation on 12/13/82 and to E. Johnson of EG6G in a letter on 12/20/82. This problem is of particular importance to the industry as a whole since the volume average velocities are used to calculate transient piping forces in postprocessor computer codes such as REFORC, which was developed by EDS as part of the EPRE S/RV Program and BLAZER, release 2, version 1, which was written by EG6G.

Specifically, the problem arises from the definition of volume average ve-locity contained in equations 235 and 236 on page 122 of Vol. 1 of the RELAP5 manual . For the liquid phase calculation:

(v~)n - SlZ inlets n

A

+ 1/2 Z (ff) $ K utlets and for the gas phase calculation:

(ap) A) inlets (s y v ) A) Q A)

+ 112 (o y )

99) tlets where the subscript, K, is the volume index and j is the junction index B-1

Por the simplifying case of a constant area volume connected by one inlet and one outlet junction, the volume average velocity reduces to a simple arithmetic average af .the inlet and outlet Junction velocities. This is where the problem arises. If a volume exists across which a substantial

~

density change occurs, such as a pipe volume close to a break, or a pipe volume downstream of an opening SRV which has different junction densities due to the passage of the SRV loop seal volume, the volume average velo-city is not, correctly .calculated by the RELAP5 equations.

As an example, a simple model of a subcooled blowdown facility has been modelled in RELAP5. Figure 1 shows the RELAP5 model. The plots of the function flowrates close to and at the 'break plane are shown on Figures 2, 3 and 4. As an aid to our use of the RELAP5 computer code, Ebasco deve-loped a postprocessor plotting program. These plots for the equivalent junctions plotted on Figs. 2, 3 and 4 are shown on Figures 5, 6 and 7.

For comparison, the flowrates for the volumes bounded by the plotted three junctions are illustrated on Figures 8 and 9. These volume flowrates have been calculated using:

MFIAN ~ AVOL+(RHOP&ELPWOIDP+RHOG&ELG&OIDG)

All the variables on the right side of the above equation have been obtained from a magnetic data tape written by RELAP5. As a further comparison, the plots of the liquid density, gas density, liquid velocity, gas velocity, and gas void fraction for volume 3420000 and junction 4000000 are included on Figures 10 to 19. These plots illustrate a similarity between all the plot-ted parameters of volume 3420000 and junction 4000000 except the liquid and gas velocities. These discrepancies illustrate the cause of the differences between the flowrates of the plotted volumes and their bounding junctions shown on Figures 5 to 9. The comparison of these calculated parameters at a specific time in the transient are tabulated on Table 1.

Basically the RELAP5 calculated volume velocities are incorrect, and do not satisfy continuity. In an attempt to solve the problem with the definition of volume average velocity, a modification was made to the average volume velocity definition which occurred in subroutines IVLVEL and VOLVEL. Specifically, the parameter defined as RATIO was modified to equal RATIO ~ 1. /(AVOL(I)CARHOP(I)AVOIDP (I) )

for the liquid phase and RATIO ~ 1./(AVOL(I)+RHOG(I)&OIDG(I))

for the gas phase. These changes modified equation 235 of Vol. 1 of the RELAP5 manual to a a (rf)k +E f f) +g (ERE f 2 ( <f tf 2 j.n1et 0 (<f rf >) out lets k

B-2

for the liquid phase and resulted in similar modifications to equation 236 for the gas phase.. Unfortunately, when the sample problem being used was input into the modified version of RELAP5, the code results became unstable and "blew up" by the fourth code reduced timestep. Consequently, it is felt that either our understanding of the parameters used in the modified sub-routines were incorrect, or the relationship of the volume average velocity calculation to the full finite difference solution scheme was not fully under-stood by us.

Consequently, EG&G was asked to address this problem. It was emphasized that the solution of this problem is important to the entire PWR nuclear industry, since the RELAP5 calculated volume average velocity values are being used to calculate piping forces by all organizations currently using RELAP5 to reanalyze their Safety Relief Valve and Power Operated Relief Valve d ischarge pipe. The importance to determine the significance of this para-meter to the calculation of piping forces in order to justify the conser-vatism of the current calculation, or to recommend appropriate changes was also emphasized.

In response to our concern in a letter dated 3/8/83, E. Johnson of EG&G supplied the following explanation from Dr. V. H. Ransom, the RELAP5 section leader: "Zhe volume average velocity which is calculated in RELAP5 is &r the purpose of evaluating the momentum flux terms in the momentum equations. It is also used for evaluating the constitutive relations such as wall heat transfer, wall friction, .interphase drag, etc. The definition does not pre-serve mass continuity and should not be used where such effects are important ~

Bass continuity is only preserved at the junctions and then only under steady flow conditions. A wide variety of approximations for volume average velocity can be formulated, but none are any more accurate than a simple average.

Finer nodalization must be used to obtain increased information on the spatial velocity profile."

Dr. Ranson further advised us that the volume average velocity definition was revised in the MOD2 code to better account for sharp density and void fraction variations and that:

"For the time being you should base the momentum calculations for piping loads on the junction velocities/mass flows and retain the existing volume average velocity formulation."

Consequently, the Ebasco developed computer code, CALPLOTFIII, was modified, as described in Appendix A, to use the junction condition to calculate piping forces.

B 3

REFERENCES (1) RELAP5/CALPLOTFIII Verification, Ebasco Applied, Physics Calculation 3-E-8 No. 001, May, 1983 (2) REFORCE A Co uter Pro ram for Calculatin Fluid Forces Based on RELAPS Results Electric Power Research Institute, PWR Safety and Relief Valve Test Pxogram, EDS Report No. 01-0650-1194, Rev. 0, February, 1982 (3) Ware,A.G, BLAZER: Release 2 Version 1 Code Manual, EG6G Idaho, EGG-EA-5888, June, 1982 (4) Ransom, V. H., Wagner, R. J., RELAP5 Mod 1 Code Manual Vol. 1 6 2 EGSG Idaho, NUREG/CR 1826, EGG 2070 Draft, Revision 2, September, 1981 B -4

=

TABLE I Summary of RELAP5 Calculated Parameters at Problem Time 1.0 sec Junction Volume Junction 3410000 3420000 4000000 Liquid density 51.554 52.755 52.755 gas density 0.89170 0.67578 0.67579 gas void fraction 0.883 0.93503 0.935 liquid velocity 588.84 794.317 999.78 gas velocity 653.77 826. 77 999.78 area 0. 07986 0.07986 0.07986 flowrate 324.11 259.14 324.12

FIGURE 1 RELAP5 - INODEL OF SUBCOOLED BLOWDOWN EXPERIINENT COMPONENT 6 SINGLE VOLUME COMPONENT 1 p ~14.7 pais SINGLE VOLUME AIR AT 700 F 24 IN. SCH 60 p ~ 923 pais QUALITY~ 0.0 RUPTURE DISK IBREAK) CDMPDNENT ~

JUNCTION 4000000 ba VOL 3420000 I 153 pals ~ 6833 CPS JUNC 3410000 A18 AT FT.

120o F PIPE COMPONENT 3 4 IN. SCH 80 VOL 34'10000 r anv.

923 pais

~ 0.0 1A$33 FT.

40 PIPE VOLS - 1 FT. LENGTH EACH - VOLS 3010000 TO 3400000 P " 923 pais- QUALITY~ 0.0 COMPONENT 2 JUNCTION 40 FT.

2000000

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