ML17261A423

Forwards Addl Info Re NUREG-0737,Item II.D.1, Performance Testing of Relief & Safety Valves, in Response to 861211 Request

ML17261A423
Person / Time
Site:	Ginna
Issue date:	02/13/1987
From:	Kober R ROCHESTER GAS & ELECTRIC CORP.
To:	Lear G NRC OFFICE OF INFORMATION RESOURCES MANAGEMENT (IRM), Office of Nuclear Reactor Regulation
References
RTR-NUREG-0737, RTR-NUREG-737, TASK-2.D.1, TASK-TM NUDOCS 8702250483
Download: ML17261A423 (34)

Text

~ 6 'EGULATORY I ORNATION DISTRIBUTION:SY, (RIDS)

ACCESSION NBR: 8702250483 DOC. DATE: 87/02/13 NOTARIZED: NO DOCKET 0 FACIL: 50-244 Robert Emmet Qinna Nuclear Planti Unit i. Rochester G 05000244 AUTH. NAME AUTHOR AFFILIATION KOBERi R. W. Roche'ster Qas 5 Electric Corp.

RECIP. NANE RECIPIENT AFFILIATION LEARN Q. E. PNR ProJect Directorate 1 LEARN Q. E. Document Control Branch (Document Control Desk)

SUBJECT:

Forwards addi info re NUREG-0737I Item I I. D. ii "Perf ormance Testing of Relief h Safety Valves>" in response to 861211 re/Vest. '

DISTRIBUTION CDDE: 60460 CQPIEB RECEIVED: LTR i ENCL 1 SIZE:

TITLE: OR Submittal: TMI Action Plan Rgmt NUREQ-0737,Zc, NUREG-0660 NOTES: License Exp date in accordance with 10CFR2i 2.'109(9/19/72). 05000244 REC IP IENT COPIES RECIPIENT ID CODE/NANE LTTR ENCL ID CODE/MANE LTTR ENCL PWR-A ADTS 1 PWR-A EB 1 1 PWR-A EICSB 2 2 PWR-A FOB 1 1 PWR-A PD1 LA 1 0 PWR-A PD1 PD 04 5 5 DI IANNI. D 1 1 PWR-A PSB 1 1 PWR-A RSB ESPRIT'OPIES INTERNAL: ADN/LFNB 0 AEOD/PTB ELD/HDS4 0 IE/DEPER DIR 33.

IE/DEPER/EPB 3 3 NRR BWR ADTS NRR PAULSON, W. 1 'NRR PWR-A ADTS NRR PWR-B ADTS 1 NRR/DSRO 01 1 EXTERNAL: LPDR 03 1 1 NRC PDR 02..

NSIC 05 1 1 TOTAL NUi1BER OF COPIES REQUIRED: LTTR 30 ENCL 27

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z~dX Pizzazz / 5TAI 5 ROCHESTER GAS AND ELECTRIC CORPORATION o e9 EAST AVENUE, ROCHESTER, N.K 14649-0001 ROGER W. KOBER 'T K I. K P R O N K VICE PRKSIDKNT ELECTRIC PRODUCTION ARKA COOK TIK 546-2700 FEB j.3 1987 U.S. Nuclear Regulatory Commission Document Control Desk Attn: Mr. George E. Lear> Chief PWR Project Directorate No. 1 Washington> D.C. 20555

Subject:

NUREG-0737 Item II.D.1> Performance Testing of Relief and Safety Valves R. E. Ginna Nuclear Power Plant Docket No. 50-244

Dear Mr. Lear:

Your letter dated December ll~ 1986 requested additional information regarding NUREG-0737 Item II.D.l. The attachment to this letter provides the requested information.

Ve y ruly yours<

p.e/

Roger W. Kober Attachments

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RESPONSE TO NRC REQUEST FOR ADDITIONAL INFORMATION DATED DECEMBER 1 1 I 1986 NUREG-0737 ITEM II.D.l Question 1 The appendix of RG6E's March 4I 1983 (Reference 1) submittal showed the calculation of the inlet piping pressure drop for the Ginna safety valves. The pipe length used in the calculation of the acoustic wave pressue drop was 3.3 ft (from Table A-1 of the appendix). Figure 2-2 in the main body of the same reporti howeveri shows the safety valve inlet piping length to be at least 4.93 ft. This is not the total piping length since the length of at least one portion of the piping was not shown. The additional length could make a significant difference in the calculated acoustic wave pressure drop. Also the calculated flow0 pressure drop in the same appendix did not account for the 90 bend when calculating the L/D term. Provide the corrected pressure drop with this bend accounted for using the equation found in Rev. 2 of the EPRI Submittal Guide (Reference 2).

If the recalculated inlet pressure drop for the Ginna safety it valves exceed the pressure drop for the 3K6 test valves>

will be necessary for the ring settings of the Ginna valves to be adjusted so that the valves operate stably.

~Res ense A review of the pressure drop calculation did uncover discrepancies between the calculation moael and the as-built piping configuration. A revised pressure drop calculation has been performed in Attachment li using the procedure found in Appendix B of the "EPRI PWR Safety and Relief Valve Test Program Guide for Application of Valve Test Program Results to Plant-Specific Evaluations"I Revision 2.

Each inlet line is analyzed separatelyi as the pipe lengths and consequently the pressure drop results differ between the valves.

The results of the analysis show that Safety Valve 434 has inlet piping pressure drops below the pressure drop for the 3K6 EPRI test valve for closing and is 14.3't above the test valve pressure on opening. Safety valve 435 has inlet piping pressure drops below the pressure, drop for the 3K6 valve on closingi and is 14.6% above the 3K6 valve on opening.

The Ginna Station Safety Valves are thereforei bounded by the EPRI 3K6 valve for closing and exceed the 3K6 pressure drop for opening. Exceeding the pressure drop on opening is acceptable because opening operation of the 3K6 valve during

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the EPRI tests was stable with minor fluctuations during loop seal discharge. These fluctuations are expected and are considered normal operation for this type valve. No tests were terminated due to problems on opening. For this reasoni the Ginna 4K26 valves are expected to operate stably without adjustment and with no operational consequences resulting from the less than l5% increase in inlet piping pressure drops Question 2 The response to question 7 of our request for information (Reference 3) on the qualification of the PORV control circuitry was not sufficient to qualify the control circuitry under NUREG-0737. The response also stated the control circuitry was not included in the 10CFR50.49 review. It is the staff position that the PORV control circuitry be qualified to perform its required function for any potential environment that it may be exposed to. If any exception is taken with respect to qualification to the harsh environment for equipment important to safetyi the licensee must demonstrate that the equipment is not required to perform a safety function to mitigate the effects of any design basis acciden't when exposed to the environment caused by the accidenti and any equipment failure in any mode in the harsh e'nvironment wil'1 not adversely impact safety function or mislead the operator.

~Res onse The pressurizer PORVs are not required to perform any safety function to mitigate the effects of any design basis accident when exposed to a harsh environment. The transients that do cause a harsh environmentI such as LOCAsl steam line breaks>

feedwater line breaks or other high energy line breaks do not require operation of the PORVs to terminate the transient or mitigate the consequences of the event. Thereforei the PORVs are not required to be qualified for operation in a harsh environment.

The PORVs must also not inadvertently actuate when exposed to a harsh environment. The PORV is a spring closed valve that requires air to open. If no air is supplied to the PORV the PORV will not openi even when exposed to a harsh environment.

Air is supplied to the PORV by an ASCO solenoid valve. When the solenoid valve is energized> air is supplied to the PORV causing the PORV to open. When the solenoid valve is deenergized air is vented and the PORV closes. The ASCO valve is an environmentally qualified valve> which should not be caused to be inadvertently energized by a harsh environ-ment. The control system connected to the ASCO solenoid

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valve is a two-wire ungrounded DC system. It has been established and agreed to by the NRC (Reference 4) that there exists no credible "hot short" failure which need to be postulated for the two-wire ungrounded DC system used for Ginna control circuitry< such as the sol.noid valves.

Without a hot short the solenoid valve cannot be energized.

Since a harsh environment cannot cause the PORV itself to inadvertently operate, the ASCO solenoid which is qualified for a harsh environment will not inadvertently operate> and there is no potential for environmental qualification-related failures in the electrical system portion of the solenoid valves< it has been concluded that the PORVs will not inad-vertently open due to a harsh environment.

Question 3 In the Pressurizer Safety and Relief Line Evaluation Reporti which was included as a part of the submittal (Reference l)I the two piping discharge conditions used in the thermal hydraulic analysis were identified as the, simultaneous opening of either the two safety valves or the two PORVs.

The report did not provide specific information which is needed to define the loading cases. Please give the assumed fluid parameters used for the analysis of each piping discharge case such as the fluid state> peak pressurei temperaturei pressurization rate> and valve flow areas and justify that these loading conditions would produce the maximum loads on the piping system. In particularI provide additional information demonstrating the steam discharge case analyzed for the PORV results in piping forces that bound the liquid discharge case. Also give the analysis parameters such as the valve opening time> node spacingi and time step size used for the calculations.

~Res ense Various fluid transient analyses were performed for the pressurizer safety and relief valve piping system. Operation of the safety valves at power operation and actuation of the relief valves to mitigate cold overpressurization were cases evaluated. In generalr the two safety valves opening simultaneously and discharging without PORV flow and the two PORVs opening simultaneously without safety valve flow are the limiting design cases. Typically> the worst case valve discharge case (SOTF) is the double safety valve discharge transient for the safety valve pipingr including the inlet<

outlet and common region piping and the double relief valve discharge transient for the relief valve inlet and outlet piping. The initial conditions for the safety valve water slug discharge event included:

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P (Upstream) 2575 psia h (Steam< Upstream) = 1110 Btu/lb h (Water> Upstream) = Enthalpy based upon a temperature profile consistent with EPRI safety valve discharge test 5917> i.e.>

approximately 300 0 F at the valve inlet and saturation temperature at the steam-water interface P (Downstream) 14.7 psia The pressurizer conditions were held constant for the transient at 2575 psia and 1110 Btu/lb.

According to the results of the EPRI tests< high frequency pressure oscillations of 170-260 HZ typically occur in the piping system upstream of the safety valve while the loop seal water passes through the valve. No significant bending moment during this "simmering" phase of the transient will<

however> occur (see Rochester Gas and Electric Corporation letter from R.W. Kober to J.A. Zwolinski of the U.S. Nuclear Regulatory Commission dated Hay 24< 1985 for a discussion of this issue). The safety valve slug discharge event generates limiting loadings for the safety valve inlet> outlet and common region piping.

The initial conditions for the relief valve slug discharge case included:

P (Upstream) 2350 psia h (Steam< Upstream) = 1162.4 Btu/lb.

T (Water< Upstream) = 150 0 F P (Downstream) 14.7 psia For this case> small cold loop seals> each 1 foot long> were assumed to exist upstream of the valves. This is conservative as the piping layout is such that no or very little condensate will remain in the upstream relief valve piping.

The pressurizer conditions were held constant for the entire transient at 2350 psia and 1162.4 Btu/lb.

The initial conditions utilized for the water discharge cold overpressurization case included:

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P (Upstream) 550 psia T (Water< Upstream) = 100 0 F

'P '(Downs tr'earn) = 14.7 'psia The pressurizer conditions wege held constant for the transient at 550 psia and 100 F.

The peak forces in all discharge piping segments were at least 2 times larger for the relief valve slug discharge case when compared to the cold overpressurization case. The slug discharge event generates limiting loadings for the relief valve discharge piping.

The adequacy of the thermal hydraulic analyses can be verified by the comparison of analytical and test results for thermal-hydraulic loadings in safety valve discharge piping for EPRI Tests 908 and 917. In that evaluation< node spacing and time-step size were selected on the basis of stable solutions of the characteristic equations and matching of tegt data. The safety valve full open flow area of 0.022 ft was used in the model.

than the Crosby M-orifice This area is s/ightly smaller area of 0.025 ft for the tested valve< but results in a good analytical match of the tested fully open valve flow rate. Appropriate water temperatures were used. All pertinent data> including friction factors<

loss factors and flow areas were based upon representative calculations and the system layout. Modeling of the water was conducted with the water seal upstream of the valve prior to transient initiation. At time = 0+r the transient was initiated and the slug position was analytically calculated during and subsequent to valve opening.

The Ginna plant specific thermal-hydraulic analysis was conducted taking the same approach as was taken for the comparison to test data. Node spacing was picked consistent with the comparison and varied with pipe size and location.

Time-step sizes were utilized consistent with values utilized in the comparison> (0.004 seconds for both slug discharge cases). Valve opening times< 0.040 seconds for the safety valves and 1.00 seconds for the relief valves<2were based upon actual data. Valve flgw areas (0.0193 ft for the safety valves and 0.0174 ft for the relief valves) were selected based upon actual valve data with appropriate margins applied to account for flow rate uncertainties. All pertinent data< including friction factors> loss factors and flow areas were based upon representative calculations and the system layout. Modeling of the water slug from a temperature profile> considering initial location and movement post-transient initiation< was consistent with the comparison study. The pressurizer pressure was held constant

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through the transient at initial values. Choked flow is checked internally and automatically every time-step to ensure the proper formulation is applied at every flow path.

The highest pressure at the respective valve inlet was less than or equal to the pressurizer pressure for the two slug discharge cases and the cold overpressurization valve opening case .

The valve flow areas were adjusted prior to finalizing the thermal hydraulic analyses to account for all uncertainties and tolerances in the valve flow rate. The ASME steam flow rating for the Crosby safety valves (orifice size K ) at 2575 psia is 306 430 lb/hr. The minimum analytically determined

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steam f low through each of the safety valves in this analysis is greater than 375 000 lb/hr. This is equivalent to a f low of 122 percent of rated. The maximum expected steam flow through the Ginna Copes Vulcan PORVs is 210, 000 lb/hr.

Values greater than 275,000 lb/hr. were analytically determined. Flows greater than 130 percent of expected values were> there f ore used.

Question 4 Provide the following information on the piping and support stress analysis:

a. The submittal (Reference 1) indicated that the piping supports were analyzed in accordance with the ASME Boiler and Pressure Vessel Code~ Section III Subsection NF, but

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did not specify the allowable stresses used for the stress evaluation. Provide the allowable stress limits applicable to each load combination for the normal> upset <

emergency < and faulted conditions.

b. Provide a comparison of the worst support stresses (or loads) with the applicable allowable stresses (including those modified) to demonstrate that the piping and supports are adequate to withstand the imposed loads.

c. Give the input parameters used in the pipe dynamic analysis such as the lumped mass spacing solution time step> damping< and cut-of f frequency> etc.

limits Res onse A The pressurizer relief and safety valve piping supports were evaluated in accordance with the requirements set forth in ASME Section III Subsection NF.

< The supports were evaluated using "Design of Linear Type Supports by Elastic Analysis" described in NF-3231. 1. This subarticle refers to Appendix XVII Article XVII-2000 for the stress g .

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The stress limits as given in this article are-F = Allowable stress in tension 0.60 Sy eqn (1)

F Allowable bending stress for square and rectangular b

'qn section bars bent about their weak axis.

0.75 Sy (14)

F v Allowable shea'r stress 0.4 Sy Sy = Hinimum yield s treng th < PSI From Subarticle NF-3231.1> the allowable stress limits for Normal and Upset conditions are identical and as given in Appendix XVIX (which are given above) .

For emergency conditions< the stress limits given in XVII-2000 can be increased by one-third.

Based on 36>000 psi minimum yield strength (A-36 material) >

the allowable stresses are tabulated below.

Condition Fb/(PSX) F (PSI) F (PSI)

Normal 27g000 21 i 600 14g400 Upset Emergency 35g910 28g728 19g152 For welds< the allowable stress values given in Table NF-3292.1-1 are used.

Anchor Bolts were evaluated using allowable loads 'based on the specific size< type and embedment depth. Expansion anchor bolts governed by XE Bulletin 79-02 used a minimum factor of safety of four< while bolts subject to ductile failure used a minimum factor of safety of two.

All supports were evaluated using "Emergency" load combinations and comparing to "Normal" allowables. One support (N-628) required use of Emergency allowables with the Emergency load case< all others did not exceed Normal allowable loads.

Res onse B The worst support stresses are given in the table below.

This table identifies the support> the stress ratio (actual stress/allowable stress)< the limiting component with the support and the allowable used. Note that for supports where the limiting components are the anchor bolts; the stress ratio is based on the allowable after application of the applicable factor of safety.

S~uort Stress Ratio Limitin Item Allowable Limit N-628 0.97 Base Pla te Emergency N-625 0.95 Weld Normal N-608 0.78 Anchor Bolts Normal PS-10 0.78 Anchor Bolts Normal PS-5 0.74 Anchor Bolts Normal PS-4 0.69 Anchor Bolts Normal PS-2 0.68 Anchor Bolts Normal PS-13 0.68 Anchor Bolts Normal N-629 0.68 Clamp Normal N-601 0.67 Anchor Bolts Normal Res onse C The modeling approach used for the benchmarking effort was also used for the Ginna specific structural analysis. For a given segment< a mass point was chosen such that the hydro-dynamic force would be applied along the axial centerline of the segment. In general> the mass point nearest the center of the segment was utilized. following is a discussion of key parameters used in the structural analyses of the dynamic events:

1. D~am in A conservative system damping of 1 percent was utilized for OBE. 2 percent was utilized for SSE and the thermal, hydraulic analyses. This is much lower than the actual expected value and is below the 10 percent damping used in the structural comparison to EPRI Test 908 and 917.

2. ~Lum in Lumped mass spacing was determined to ensure that all appropriate mode shapes were accurately represented as shown on Figures 6-1 and 6-2.

3. ~Su orts The structural supports were modeled in sufficient detail to analytically represent the system.

The shock suppressors and struts were modeled by inputting a stiffness in series with the piping.

Specifically calculated stiffness values were utilized.

All supports were linear and a linear overall system analysis was conducted.

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determined within the structural program and is based upon convergence criteria that results in stable solutions. The largest time-step ever used could be 0.0001 second. The time-step is automatically adjusted such that the relative error of each modal coefficient is at least less than 10

5. Cut-off Fre uenc A cut-off frequency was used to ensure that all appropriate frequencies were included.

For the relief line analysis< a cut-off frequency greater than 700 HZ was used. A cut-off frequency greater than 1000 HZ was used for the safety valve line thermal hydraulic analysis.

The structural analysis approach is substantiated by comparisons of analytical results to test data. A discussion of the methodology utilized in performing a safety valve discharge structural analysis and comparison of analytical results to structural test results is presented in the following article:

L.C. Smith and T.M. Adams> "Comparison of Analytically Determined Structural Solutions with EPRI Safety Valve Test Results"g 4th National Con ress on Pressure Vessel and Pi in Technolo , Portland Oregon< Jun'e 19-24, 1983 PVP-Volume 74, pp. 193-199.

As noted in the response to Question 3< high frequency pressure oscillations of 170-260 HZ could occur in the piping system upstream of the safety valve< while the loop seal water passes through the valve. The safety valve slug discharge event (which occurs after the "simmering" phase)> however< generates limiting system responses for the safety valve inlet> outlet< and common region piping.

REFERENCES J.E. Maier> Rochester Gas and Electric Corp. letter to D.M.

Crutchfield> NRC> "Post-TMI Requirements< NUREG-0737'tem II.D.l> R.E. Ginna Nuclear Power Plant"i March 4< 1983 EPRI PWR Safet and Relief Valve Test Pro ram Guide for Application of Valve Test Pro ram Results to Plant-S ecific Evaluations> Revision 2i Interim Report< July 1982 R.W. Kober< Rochester Gas and Electric Corp. letter to J.A.

Zwolinski< NRC< "NUREG-0737> Item II.D.l Performance .

Testing of Relief and Safety Valves" May 24> 1985.

J.A. Zwolinksi, NRC< letter to R.W. Kober> Rochester Gas and Electric Corp., "Safety Evaluation for Appendix R to 10CFR Part 50> Items III.G.3 and III.L"

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ATTACHMENT 1 GINNA STATION PRESSURIZER SAFETY VALVE INST PIPING PRESSURE DROP CALCULATION Y-EC-248 Prepared By:

S. . Shaw, E in r ump and Valve ngineering K Electric Corp.

Reviewed By:

. L. Lace , ngineer ump and V ve Engineering g Electric Corp.

Approved By H. A. epp, ager Pump and Valv 'ngineering g Electric Corp.

January 12, 1987

ZEUL I This calculation was prepared by Westinghouse Electric Corporation at the request of Rochester Gas and Electric Corporation in support of their response to the Nuclear Regulatory Commission, regarding NUREQ-0737 Item II.D.1, Performance Testing of Relief and Safety Valves.

KQPE:

The lines analyzed herein are for the Ginna Station pressurizer safety valve inlets, valve locations 434 and 435.

PUBPEK:

The intent of this calculation is to provide plant-specific pressurizer safety valve inlet piping pressure drops, for opening and closing. These values are to be compared with the pressure drops, calculated by the same methods, found for pressurizer safety valves tested as part of the EPRI Safety and Relief Valve Test Program.

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A. Ginna Pipe Configuration Several quantities must be determined from the actual physical configuration of the inlet lines for use later in the calculation. These are: total length of pipe, length of straight pipe, and L/D. The configurations, from Ref. 1 and 6 are shown in Figure 1 for valve 434, and Figure 2 for valve 435.

1. Valve 434

a. Inlet piping consists of:

1 3" x 4" Expander 4" Schedule 160 pipe 1 45 Elbow Pressurizer Nozzle 1 90 Elbow 1 180 Elbow

b. Total Length of Pipe:

.8490 + .3333 + .3927 + .366 + .7854 + 1.250 + 1.5708 +

.427 = 5.974 ft.

c. Length of Straight Pipe:

.8490 + .3333 + .366 + 1.250 + .427 = 3.225 ft.

FIGURE 1: VALVE 434 (SHOWN ROTATED INTO PLANE)

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/ZSG FIGURE 2: VALVE 435 (SHOWN ROTATED INTO PLANE) 2

d. L/D

1. From Reference 3, Appendix A Standard 45o Elbow L/D = 16 Standard 90 Elbow L/D = 30 Standard 180 return L/D = 50 Long Radius 90 Elbow L/D = 20 All elbows used in this Ginna Station pipe are long radius, 6 inch radius for 4 inch pipe. Since the long rpiius L/D's are not directly available for the 45 elbow and the 180 returns, they will be found by using the ratio of the long radius to standard 90 values.

Long Radius 45 elbow L/D = 16 (20/30)=10.67 Long radius 180 elbow L/D = 50 (20/30)=33.33 Note: Other sources availabe for calculation of the 180 elbow, such as

, Robert P.

Benedict, and from Reference 3, "Resistance of Bends<<, show the use of this method is conservative.

2. From Reference 4, for 4>> schedule 160 pipe D = 3.438 inches = .287 ft.

3. L/D = ~~~ + 10.67 +

20

+ 33.33 75.24 287 (dimensionless)

2. Valve 435

a. Inlet piping consists of:

Pressurizer Nozzle 1 3>> x 4>> Expander 1 45 Elbow 1 90 Elbow 1 180 Elbow 4>> schedule 160 pipe

b. Total Length of Pipe:

.8490 + .3333 + .3927 + .9719 + .7854 + 1.250 + 1.5708

+ .427 = 6.580 ft.

c. Length of Straight Pipe:

.8490 + .3333 + .9719 + 1.250 + .427 = 3.831 ft.

d. L/D (Same variables used as in Valve 434 above)

L/D = ~U..287 10.67 20 33.33 77.35 (dimensionless)

B. Calculation for Pressure Drop (Reference 2)

1. For opening and again for closing for each valve a transient, and steady state pressure drop is calculated.

For each case, the larger of the transient and stealy state values is used as the controlling value for that case. Each transient pressure drop is the sun of the flow pressure difference due to pipe friction and fittings plus the pressure difference due to acoustic wave amplitude. The steady state pressure drop is due to pipe friction and fittings.

Thus there are 4 cases:

Case 1: Valve 434 Opening Case 2: Valve 435 Opening Case 3: Valve 434 Closing Case 4: Valve 435 Closing And for each case the greater of (Delta PF + Delta PA>)

or Delta P ss is used.

The equations and definitions used from Reference 2 are as follows:

The flow pressure difference due to pipe friction and fittings is given by:

If T g 2L/a, XL Delta PF +

2g ho A

If T > 2L/a, ZL Delta P = +

F 'g c Rho A

where, sunmation of expansion and contraction loss coefficients corrected if required to correspond to the inlet piping flow area. (NOTE: The contraction from the pressurizer to the inlet pipe can be assuned to be smooth and, therefore, the loss coefficient can be assumed to be zero)

(dimensionless) friction factor (dimensionless) piping equivalent length/diameter considering effects of elbows and friction (dim nsionless)

M rated valve flow rate for steam as specified in Table B-1 of Ref. 2 (lb./sec.)

gc = gravitational constant (32.2 lb-ft/lb-sec 2 )

Rho steam ensity at ncminal valve set pressure (lb/ft$ )

inlet piping flow area (ft )

a stean sonic velocity (ft/sec) 1100 ft/sec. for all calculations inlet piping length (ft) valve opening or closing time for steam inlet conditions as specified in Table B-2 of Ref. 2 (sec.)

flow rate constant for valve opening or closing as specified in Table B-2 of Reference 2.

There are two situations to consider:

If T g 2L/a, Delta PAN

= ~~

g A

~

2g Rho A 2

If T ) 2L/a, 2

Delta B~CJQ M~ a 1 P<M ~~CCM g AT 2g Rho A All param ters are defined above.

Ste -S at The steady-state flow pressure difference associated with valve opening or closing is given by:

XL ~

p Delta PF g ho All parameters are defined above. The values of the flow rate constant, C, are different for valve opening and closing and are provided in Table B-2, Ref. 2.

For this calculation, the param ters used are as follows:

a: K = K expander From Reference 4, for gradual enlargements this is:

K expander theta =

=

where Beta 2 = a1 e a2

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Beta (3.438) 23.01, derived from Reference 2

tag 5BB53 5 dimensions therefore

(.58253)

= .266 b: f= .017 (Reference 4)

C: L/D = Calculated above Valve 434 L/D = 75.24 Valve 435 L/D = 77.35 d: M = 320,000 lb/hr. (Reference 2) = 88.89 Lb/sec.

e: gc = 32.2 ft-lb/lb-sec 2

0 Rho = 7.65 lb/ft3 (Reference 4)

.0645 ft2 (Reference 4) h: 1100 ft/sec.

L = Calculated above Valve 434 L = 5.974 ft.

Valve 435 L = 6.580 ft T = Valve opening .010 sec. (Reference 2)

Valve closing .016 sec.

k: Transient:

Valve opening C = 1.11 (Reference 2)

Valve closing C = .69 Steely State C = 1.11

2. Case 1 Valve 434 Opening

a. Transient Calculation - Pressure drop due to friction 2L  ?~.'LQ. 01 1 1100 T = .010 < .011 = PJ, a

Therefore, the transient pressure difference is Delta 2 P

f (2) (32.2) (7.65) (.0645) (144)

Delta Pf - 50.97 psi Pressure drop due to acoustic wave amplitude T < PJ

~2 a

therefore, D lt P aw (32.2) (.0645) (144) (2) (32.2) (7.65)(144)

(.0645)

= 362.90 + 32.99 = 395.89 psi

The total transient pressure drop is then Delta PT = Delta Pf + Delta P aw

= 50.97 + 395.89 = 446.86 ps'.

Steady state calculation ss (2) (32.2) (7.65) (.0645) (144)

= 50.97 psi

c. Since the transient pressure drop is greater than the steady state pressure drop for Valve 434 on opening, the controlling value is the transient one, which is Delta P = 446.86 psi

3. Case 2 Valve 435 Opening

a. Transient calculation Pressure drop due to friction 2L ZSMHQ..012 1100 T = .010 ( .012 = 2J a

Therefore, the transient pressure difference is:

Delta Pf-(2) (32.2) (7.65) (144) (.0645)

= 52.15 psi Pressure drop due to acoustic wave amplitude T <?J a

therefore 2

(32.2) (.0645) (144) (2) (32.2) (7.65) (134)

(.0645)

= 362.90 + 32.99 = 395.89 psi

I The total transient pressure drop is then Delta PT = Delta Pf + Delta Paw

= 52.15 + 395.89

= 448.04 psi

b. Steady State Calculation ss (2) (32. 2) (7.65) (144) (.0645)

= 52.15 psi

c. Since the transient pressure drop is greater than the steady state pressure drop for Valve 435 on opening, the controlling value is the transient one, which is Delta P. = 448.04 psi

4. Case 3 Valve 434 Closing

a. Transient calculation, pressure drop due to friction Since T = .016 > .011 = 2L a

then Delta Pf- [.266+.017 (75.24)][(.69)(88.89)] [(~3L44. 2 (2) (32.2) (7.65) (144) (.0645)

Delta Pf - 9.08 Pressure drop due to acoustic wave amplitude T>2I a

therefore Delta P =(2) (5.974) (.69) (88.89) [(.69) (88.89) ] (~~LI]2 (32.2) (.0645) (.016) (144) (2) (32.2) (7.65) (.0645) (144)

= 153.14 + 5.87

= 159.01 psi

The total transient pressure drop is then Delta PT = Delta Pf + Delta P aw

= 9.08 + 159.01 = 168.09 psi

b. Steady State Calculation ss (2) (32.2) (7.65) (.0645) (144)

= 50.97 psi

c. Since the transient pressure drop is greater than the steady state pressure drop for Valve 434 on closing, the controlling value is the transient one, which is 168.09 psi

5. Case 4 Valve 435 Closing

a. Transient calculation, pressure drop due to friction Since T = .016 > .012 = 2I a

then Delta Pf = [.266 + (.017)(77.35)][(.69)(88.89)]

(2) (32.2) (7.65) (144) (.0645)

= 11.27 psi Pressure drop due to acoustic wave amplitude T>2l, a,

therefore

)(6.68 )(.69)(88.89) ((.69)(88.89))

(32.2) (.0645) (.016) (144) (2) (32.2) (7.65) (.0645) (144)

= 168.68 + 7.13

= 175.81 psi 10

The total transient pressure drop is then Delta PT - Delta Pf + Delta P aw

= 11.27 + 175.81

= 187.08 psi

b. Steady State Calculation Delta P ss (2) (32.2) (7.65) (.0645) (144)

= 52.15 psi

c. Since the transient pressure drop is greater than the steady state pressure drop for Valve 435 on closing, the controlling value is the transient one which is 187.08 psi.

Using the values calculated for Ginna Station, and the pressure drop information from the EPRI test from reference 2, the tabulation can be made:

Ginna 434 Ginna 435 3K6 >>F>> 6M6 >>G>>

Valve Opening 446.86 psi 448.04 psi 391 psi 263 psi Valve Closing 168.09 psi 187.08 psi 194 psi 181 psi V.

1. "Piping As-Built, Analysis Diagram Pressurizer Relief from Pressurizer to Relief Manifold", Gilbert Associates, Inc. Drawing C-381-353, Revision B, Sheets 6 and 7.

2. >>EPRI PWR Safety and Relief Valve Test Program Guide for Application of Valve Test Program Results to Plant-Specific Evaluations", Interim Report,, July 1982, Revision 2.

3. Crane Technical paper No. 410, "Flow of Fluids Through Valves, Fittings, and Piping", 1969.

4. Ibid., 1978

5. General Catalog No. 55, Ladish Co., 1954

6. "Pressurizer Upper Head Assembly", Westinghouse Electric Corp.,

Drawing 681J253.