ML20039A727

Small & Large Break Calculations for San Onofre Unit 2, Technical Rept

ML20039A727
Person / Time
Site:	San Onofre
Issue date:	10/31/1981
From:	Chon Davis, Dobbe C, Fletcher C EG&G, INC.
To:	Lauben N, Odar F Office of Nuclear Reactor Regulation, NRC OFFICE OF NUCLEAR REGULATORY RESEARCH (RES)
References
CON-FIN-A-6047 EGG-CAAD-5630, NUDOCS 8112210208
Download: ML20039A727 (38)

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NRC FIN No. A6047 8112210208 811031 YD PDR RES

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KJ INTERIM REPORT i Accession No.

Report No. EGG-CAAD-5630 Contract Program or Project

Title:

Code Assessment and Applications Division Subject of this Document; Small and Large Break Calculations for San Onofre Unit 2 Type of Document: Technical Report Author (s): C. B. Davis C. D. Fletcher C. A. Dobbe Date of Document: October 1981 A

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v) Responsible NRC Individual and NRC Office or Division: F. Odar, NRC-RSR, N. Lauben, NRC-NRR This document was prepared primarily for preliminary or internal use. it has not received full review and approval. Since there may be substantive changes, this document should not be considered final.

EG&G Idaho, Inc.

Idaho Falls. Idaho 83415 Prepared for the U.S. Nuclear Regulatory Commission Washington, D.C.

Under DOE Contract No. DE-AC07-761D01570 NRC FIN No. A6047 I /n

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! INTERIM REPORT

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4 l ABSTRACT 1 .

Analyses were performed in support of the Nuclear Regulatory i

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Comission's licensing of the San Onofre Unit 2 Combustion Engineering pressurized water reactor for low power operation. The RELAP4/ MOD 7 computer code was used to calculate the effects of a small and a large cold j leg break on San Onofre Unit 2 while operating at 5% rated power. Results j of the calculations were used to provide boundary and initial conditions for subsequent core boiloff calculations perfonned by the Nuclear l Regulatory Commission.

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v Analyses were performed in support of the Nuclear Regulatory Comission's licensing of the San Onofre Unit 2 Combustion Engineering pressurized water reactor (PWR) for low power operation. The RELAP4/ MOD 7 computer code was used to calculate the effects of a small and a large cold leg break on San Onofre Unit 2 while operating at 5% rated power. A complete failure of all pumped Emergency Core Coolant (ECC) systems was assumed. Results of the RELAP4/M007 calculations were used to provide boundary and initial conditions for subsequent core boiloff and damage calculations performed by the Nuclear Regulatory Commission. The purpose of the core boiloff calculations was to determine the time of core damage.

The small break calculation considered a 0.0508 m diameter hole in a cold leg pipe. The calculation represented nearly 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> of the small break transient. No heatup of the reactor fuel was calculated although the core nearly uncovered prior to accumulator injection. At the time the

, calculation was terminated, the accumulators were empty and the upper plenum mixture level was 0.09 m below the bottom of the hot leg nozzles and was decreasing slowly. Hand calculations indicated that the core would not uncover for at least 2 additional hours.

The large break calculation considered a 200% double-ended offset shear of a cold leg pipe. The calculation represented 83 s of the large break transient. No significant heatup of the reactor fuel was calculated. The reactor vessel voided during blowdown and was refilled by the accumulators, which emptied at 80 s. The effect of the stored energy in the steam generators and the reactor vessel wall was significant during the transient. The energy transferred to the primary coolant from the steam generators and the reactor vessel wall was about 20 times higher than the core power at the end of the calculation. Accumulator temperature also had a significant effect upon the calculated refill of the vessel. With cold accumulator water, 80% more liquid was present in the vessel at 80 s compared to a calculation with artificially warm accumulator water.

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CONTENTS ABSTRACT ............................................................. II

SUMMARY

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1. INTRODUCTION .................................................... i 3 .

2 RELAP4 MODEL DESCRIPTION ........................................

2.1. Modifications for San Onofre Configuration ................ 3 2.2. Nodalization .............................................. 4 2.2.1 Small Break Nodalization .......................... 4 2.2.2 Large Break Nodali.*ation .......................... 4 2.3. Code Options .............................................. 5 2.3.1 Code Options for the Small Break Calculation ...... 5 2.3.2 Code Options for the large Break Calculation ...... 6 2.4. Initial Conditions ........................................ 7 2.5. Boundary Conditions ....................................... 9

3. RESULTS ......................................................... 12 3.1. Results of the Small Break Calculation .................... 12 3.2. Results of the Large Break Calculation .................... 15

4. CONCLUSIONS ..................................................... 22

5. REFEP.ENCES ...................................................... 24 FIGURES

1. Nodalization for the small break calculation .................... 25

2. Nodalization for the large break Calculation .................... 26 1

3. Calculated upper plenum and steam generator pressures--small i break ..........................................................27 l 1

Calculated secondary mixture level--small break ................. 27 -l

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5. Calculated upper plenum level--small break ......................

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6. Calculated acccmulator level--small break ....................... 28 f-s t

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7. Calculated leax quality--small break ............................ 29 8 8. Calculated upper plenum, steam generator, and accumulator pressures--large break .......................................... 29

. 9. Calculated vessel coolant inventory--large break ................ 30 l

10. Calculated core temperatures--large break ....................... 30 j 11. Calculated core and total power--large break .................... 31 l 12. The effect of ECC temperature on vessel coolant inventory--

! large break ..................................................... 31

13. The effect of ECC temperature on Upper plenum pressure--

large break ..................................................... 32 TABLES i

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1. Initial conditions for the small and large break calculations .................................................... 8

2. - Axial power factors ............................................. 11 j 3. Cnronology for the small cold leg break calculation ............. 13

4. Chronology for the large cold leg Dreak calculation ............. l4 i

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m 1. INTRODUCTION O

Two calculations were performed in support of the Nuclear Regulatory Commission's licensing of the San Onofre, Unit 2 Combustion Engineering pressurized water reactor (PWR) for low power operation.

4 Thermal-hydraulic plant conditions were simulated using the RELAP4/M007 computer code l aad a model of the San Onofre PWR which was derived from an existing model of the Combustion Engineering Calvert Cliffs, Unit 1 PWR. The plants are similar in some respects (such as primary volume and pipe sizes) and different in others (such as rated core power, and core bun'dle and steam generator configuration.) The existing Calvert Cliffs model was modified to account for these differences and a description of the changes made is included in this report. Since the primary volume to reactor coolant piping flow area ratios are the same for

two plants, results obtained using the modified Calvert Cliffs model for any size break are representative of those that would be expected using an exact model of the San Onofre plant if one were available.

Plant operating conditions at 5% of full power . sere simulated and the effects of a small and a large cold leg break were calculated using best estimate assumptions except as noted. .The small break calculation considered a 0.0508 m diameter hole in a cold leg pipe and the large break calculation considered a 200% offset shear of a cold leg pipe. In both calculations a complete failure of all pumped emergency core coolant (ECC)

(charging, high pressure injection, and low pressure injection) was assumed.

Results of the calculations were used to provide boundary and initial  ;

conditions for subsequent core boiloff and damage calculatica performed by the Nuclear Regulatory Commission using another computer code.

This report describes the model used in Section 2. Descriptions of the changes made in converting the Calvert Cliffs model to a San Onofre Model (2.1), nodalization (2.2), code options (2.3), initial conditions (2.4),

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} and boundary conditions (2.5) are included. The results of the small and i large break calculations are described respectively in Sections 3.1 and i

j 3.2. Conclusions are suninarized in Section 4 and references given in ,

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2. RELAP4 MODEL DESCRIPTIONS The RELAP4 models used to make the San Onofre small and large break calculations are described in this section. The San Onofre small and large break calculations were performed with models based on the Calvert Cliffs model described in Reference 2. The changes made to the Calvert Cliffs model to represent San Onofre are described below. Descriptions of the nodalizations, code options, and initial and boundary conditions are also presented. Input decks for the calculations are stored under configuration control number F00506 at the Idaho National Engineering Laboratory (INEL).

The computer code, version G101 of RELAP4/ MOD 7, is stored under configuration control numbers F00509 (main coding) and H0099828 (steam tables) at INEL.

2.1 Modifications for San Onofre Configuration Several changes were made to the Calvert C1iffs model to represent the 3 San Onofre configuration of the vessel, steam generators, and accumulator components.

l The model of the vessel was modified primarily to account for a different fuel geon' 2try. The original Calvert Cliffs model represented 14 x 14 fuel bundles while San Onofre contains 16 x 16 fuel bundles. The core heat slabs were modified to represent the geometry of the San Onofre 16 x 16 fuel bundles. The hydraulic geometry of the core was not altered and remained in the Calvert Cliffs configuration. In addition, the core bypass flow rates were modified to be consistent with the bypass flow rates presented in the San Onofre Final Safety Analysis Reprt3 (FSAR). The core bypass flow from the downcomer to the upper head corresponded to 0.7%

of the total vessel flow and represented the combined leakage through the upper head and the leakage due to the hot leg nozzle clearances.

g The steam generator heat removal capacity of the San Onofre PWR is greater than that of the Calvert Cliffs PWR. While the number and size of

. s tubes is the same, the average tube length is longer (18.44 vs 14.63 m)

( which results in a larger heat transfer area. The model was revised to 3

account for this difference by increasing the height of the uppermost volumes and heat slabs and the secondary mixture levels in each steam generator, a

a The San Onofre accumulators contain 60% more liquid and are operated at 2.76 MPa higher pressure than the Calvert Cliffs accumulators.

Best-estimate values for the San Oncfre accumulator liquid volume, gas volume, and pressure were used in the small and large break calculations.

2.2 Nodalization 2.2.1 Small Break Nodalization The nodalization used for the small break analysis is shown in Figure 1. This nodalization is the same as the nodalization used in the Combustion Engineering audit calculations of open power operated relief valve transients except that junctions for the power operated relief valve and charging, high pressure, and low pressure injection were removed and the pressurizer was modeled by a single volume.

The core was modeled with three volumes and six heat slabs which represented the fuel rods. Each steam generator was modeled with four primary volumes and heat slabs and a single secondary volume. One loop was modeled using a single hot leg and cold leg while the other loop was prototypically modeled using a single hot leg and two cold legs. The accumulators were connected to the downcomer upper annulus to avoid atypical primary system depressurization during injection of cold accumulator water. Additional heat slabs were located on the downcomer, core bypass, lower plenum, upper plenum, and upper head.

2.2.2 Large Break Nodalization ,,

The nodalization of the RELAP4 model used to perform the large break  ;

calculation is shown in Figure 2. The large break nodalization was eimilar

a. The word " accumulator" is used to represent the safety injection tank.

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b; c broken pump discharge leg was assumed to.f-low directly into the containment ,- 'Q

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transients. The code options used in the small break \calculat-lon were casedonexperienceobtainedatINELinapplyingRELAP4/M007to\smallbreak transients. The Code options used in the large break Calculation:were 4 Dased on the user guidelines derived during the.. assessment of,he c (.. ,~ '

RELAP4/M006 computer code.5 Tne options used ior each calculation are 4 descriDed oelow. b

- ,s 2.3.1 Code Options for the Small Break Calculation -

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Tnefollowinguser-selectedcodeinputoptionswereused'forthesmalN b break calculation. Critical flow was modeled using the -

Henry-Fauske/ Homogeneous Equilibrium Model option. A multiplier of 1.0 was.

applied to botn models. The vertical phase slip model was applied in the.

downcomer and core (Junctions 1 and 3, and 5 through 10, see Figure 1).

1 The new slip model, which uses a flow-regime dependent slip velocity correlation, was applied. The Wilson bubble rise model was used in the

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pressurizer, loop seals, and the vessel. The lower plenum, core, and upper j plenum were modeled as a vertical stack. The downcomer was also modeled as a vertical stack. A bubble gradient of 0.8 was used with the Wilson bubble

, , rise model. Tne parameters of the bubDie rise model used in the steam generator secondaries were calculated by the self-initialization feature.

, To save calculational time, the bubble rise models in the upper head and

pressurizer were tripped off and homogeneous calculations made after the respective volumes were depleted of liquid. The default and/or recommended

heat transfer options of the RELAP4/M006 computer code were used.

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Specifically, tne H S0 beat transfer surface was employed. CHF was calculated witt, the W-3 correlatio1 in the subcooled regime, Hsu and Beckoer's modified ri-3 correlation.in the saturated high flow regime, and .

_ Smitn and Griffith?s mdified Zuber correlation in the sa .urated low flow regime. Transitioilooilicq. calculated with the modified Tong-Young ,

correlation. Fjim toiling was calculated with the Condie-Bengston 111 corre la f: ion. Toe r,aturaliconvection beat transfer option was used in bcth s ,

s steam generator secocaddie;. The self-initialization feature was used to calculate initial: system pres $ure 300 temperature distributions. The enthalpy transport model was"used for initialization but was not used cpr,ing the transient. The coipress1 ole form of the momentum equation QMVMIXa U) was used at~all-junctions except that the incomp'*essible k ~

momenth equation (MVMIX = 3) was used at tne presmrizer to hot leg, 7 , JaccumuhNr to inlet annulus, and fill junctions and those junctions connect'ing the vessel with the hot or Cold legs. Tne alternate (rec 9arrisnded) water packing and. liquid mass depletion model was applied.

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Lorg power was calculated with the RELAP4/M007 reactor kinetics mMal . F1ssion power was calculated with point kinetic equations.

Reecth ity effects due to scram and feedback effects due to moderator temperature, moderator density, and fuel temperature were modeled. The reactivity curves input to the model were based on information presented in the San Onofre FSAR. Decay heat, including the effect of heavy element decay, was calculated assuming an infinite operating period.

2.3.2 Code Options for tne Large Break Calculation Tne code options used in the large break calculations were similar to those used in the small break calculation except as descrioed below.

Pnase separation in the reactor vessel was calculated with the .

vertical phase slip model whicn was applied at the same junctions as in the small oreak calculation. The Wilson bubble rise model was applied only in e tne pressurizer and upper nead.

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c Critical flow was calculated using stagnation critical flow

) properties. Tne stagnation properties option was tripped off when the upper plenum pressure dropped below 0.34 MPa O

The polytropic air expansion model was applied 'a the accumulator volumes. A polytropic exponent of 1.4 was used.

The nonequilibrium model was applied in the pump discharge legs and reactor vessel. Default values of model parameters were used except that, as recommended, the nonequilibrium model was tripped off in those volumes with a quality below l.0E-6 and the model reactivated in these volumes when the void fraction exceeded 0.25.

The first water packing model, the model which was available in RELAP4/ MOD 6, was used in the large break calculation.

2.4. Initial Conditions l The small and large break transients were assumed to start witn the plant operating at 5% rated power and 100% core flow. The basic initial conditions whicn were input to the computer model are shown in Table 1.

The initial cenditions represent a best-estimate of the actual operating status of tne plant at 5% power. The basic initial conditions were obtained, when available, from tne San Onofre FSAR. The self initialization feature of RELAP4/M007 was use to calculate the steady-state steam generator energy balance and the thermodynamic state of each control volume. Specific assumptions pertaining to the initial conditions are described below.

The initial pressurizer liquid volume was reduced to 50% of the normal volume at 100% power to approximate the effect of the pressurizer level control system on the pressurizer liquid inventory.

4 The average of the vessel inlet and outlet fluid temparatPres was Calculated oy interpolating between the average fluid temperatures presented in the FSAR for 100% power and 0% power conditions. The average 7

TABLE 1. INITIAL CONDITIONS FOR THE SMALL AND LARGE BREAK CALCULATIONS Parameter Initial Value .

Core Power, MW 169.5 8

Total Vessel Flow, kg/s 18648.

Pressurizer Pressure, MPa 15.51 Vessel Inlet Fluid Temperature, K 558.4 Vessel Outlet Fluid Temperature, K 560.1 Core Bypass Mass Flow, kg/s 261.1 Upper Head Bypass Mass Flow, kg/s 130.5 Pressurizer (Including}urgeLine)

LiquidVolume,g 12.1 Vapor Volume, m 31.1 Steam Generator Feedwater Flow, kg/s/ generator 55 Feedwater Enthalpy, J/kg 9.881 x 105 Secondary Pressure, MPa 6.62 3

Secondary Liquid Volume, g 94.38 Secondary Vapor Volume, m 132.33 Accumulator Liquid Volume m 3/ tank 49.36 Gas Volume, m3 / tank 14.36 Pressure, MPa 4.31 Temperature, K 322 Average Fuel Centerline Temperature, K 589.5 9

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4 centerline f6el temperature was calculated by interpolating between the ,

fuel-to-coolant differential temperatures presented in the FSAR for 100%  !

V and 0% power.

The axial power factors used for the core heat slabs are shown in g Table 2. The power factors represent an axial power profile obtained from the FSAR and are applicable for 35% power early in core life.

2.5. Boundary Conditions The boundary conditions applied to the small and large break calculations included steam generator feedwater and steam mass flow rates, auxiliary feedwater mass flow rates, and containment pressure. The response of most of these boundary conditions during the transients was related to the occurrence of the scram signal. The scram signal was generated on a low pressurizer pressure for these transients. Tb' low pressure set point for San Onofre was 11.03 MPa.

The occurrence of the scram signal actuated the reactor scram; time

\ / delays were built into scram reactivity vs. time curve to account for solenoid deenergization and control rod drop time. The primary coolant pumps were tripped off when the reactor scram signal occurred. The steam generator feedwater and steam lines were closed linearly between the time the scram signal occurred and 5 s af ter the scram signal occurred.

Auxiliary feedwater was initiated 30. s af ter the occurrence of the scram signal. The auxiliary feedwater flow rate was modekd as a function of secondary mixture level. The auxiliary feedwater flow. attempted to control secoadary mixture level between 13.4 m and 14.0 m above the top of the tube sheet. An auxiliary feedwater flow rate of 43.5 kg/s/ generator was applied when the secondary mixture level was below 13.4 m. No auxiliary feedwater flow was permitted when the level exceeded 14.0 m. Auxiliary feedwater

, temperature was 322 K. The accumulators were the only source of ECC during the transients.

4 "he containment was modeled as a time-dependent volume for both O transients. For the small break transient, a constant containment pressure Nl 9

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1 of 0.101 MPa was assumed. The actual magnitude of the assumed containment pressure was not important because the break remained choked in the small break calculation. For the large break transient, the cor.tainment was '

assumed to De filled with saturated steam. The containment pressure nistory was taken from the San Onofre FSAR and represented the response calculated for a 200% cold leg break from 100% power.

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TABLE 2. AXIAL POWER FACTORS

• TOP of Core 0.0946 4

' O.1655 O.1956 0.2031 0.1956 i

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3. RESULTS 3.1. Results of the Small Break Calculation O

The times at which significant events occurred in the calculation are listed in Table 3.

At time zero a 0.0508 m diameter break was opened in the cold leg.

The pressurizer emptied at 40 s and at 45.3 s the low pressurizer pressure scram setpoint was reached causing the reactor scram, reactor coolant pump trip, and termination of steam generator main feedwater and steam flows.

The primary pressure rapidly fell until, at about 270 s, it was slightly above the secondary pressure as shown in Figure 3.

The secondary mixture level, which had fallen as a result of reactor scram and secondary isolation, recovered to the normal level shortly af ter initiation of auxiliary feedwater at 75.3 s as shown in Figure 4. The secondary level was maintained at the normal level for the remainder of the transient by throttling of auxiliary feedwater.

The reactor vessel upper head drained at 390 s and the upper plenum level (Figure 5) was only 0.1 m above the core when accumulator injection began at 4158 s. Figure 6 shows the accumulators injected liquid in stages until their inventory was depleted at 7150 s. Injection in stages was caused by repressurizations of the primary system above the accumulator pressure following surges in accumulator injection. Accumulator injection caused a refilling of the vessel as noted by the upper plenum level, and a depressurization of the primary and secondary systems. Following y accumulator depletion, the system was stabilized with a primary pressure of D 862 kPa, a secondary pressure of 1930 kPa, and the upper plenum mixture level near the bottom of the hot leg nozzles. ,

The quality in the volume with the break, shown in Figure 7, increased g l to about 0.9, returned to zero during accumulator injection, then increased to nearly 1.0 following accumulator depletion.

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l TABLE 3. CHRONOLOGY FOR THE SMALL COLD LEG BREAK CALCULATION i

} Time je (s) Event

{ 0.0 Blowdown initiated a

40.0 Pressurizer emptied j 45.3 Scram signal' generated 46.0 Reactivity insertion due to scram initiated-l 75.3 Auxiliary feedwater initiated l 390 Upper head drained f

4158 Accumulator injection initiated 7150 Accumulators emptied

i. 12514 Calculation terminated i

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The calculation was terminated at 12514 s. At that time, the primary pressure was 738 kPa and decreasing at a rate of about 0.07 kPa/s. The secondary pressures were 1980 kPa in the broken loop and 1430 kPa in the intact loop. The upper plenum mixture level was 1.08 m which is 0.09 m below the bottom of the hot leg nozzles. The break mass flow rate was 2.19 kg/s and the quality of the break flow was 0.978. 12,400 kg of liquid s remained in the vessel above the core. Steam productice was being relieved primarily by two paths: through the bypass from the upper head to the upper annulus and through the intact loop which by this time had been almost completely voided. The model included a broken loop which contained two cold legs. In the two broken loop cold legs, significant amounts of liquid were still present: 3500 kg in the intact c.old leg and 3000 kg in the broken cold leg. At the problem termination time the core power was 1.786 MW, heat addition from the steam generator secondaries was approximately 10% of core power, and heat addition from other heat structures was approximately 30% of core power.

Analysis indicates that the primary pressure and break mass flow rate would continue to decrease until the energy removal rate at the break equaled the decay heat production rate. Using conditions at 689 kPa, the mass flow rate at equilibrium will be:

1.786 MW LBM Bm

= 1.90 sec = 0.862 kg/s W -c 0.293 3600 "R-- 889LBM T

1 Btu /HR 2

This corresponds to a break mass flux of 3.679 kg/s/m L equilibrium.

This break mass flux is associated with an equilibrium primary pressure of 285 kPa. Thus tne equilibrium conditions expected are a break steam flow of 0.862 kg/s at a primary pressure of 285 kPa.

Considering only the liquid in the vessel and the break mass flow rate ,

at problem termination, core uncovery could be expected at 12514 + = 18178 s O

4 This represents a lower bound estimate of core uncovery time since it i assumes none of the 6530 kg of liquid in the loops at problem termination time migrates to the vessel and the break mass flow rate does not decrease.

,o Considering only the liquid in the vessel and the equilibrium break

- I mass flow rate. The amount of liquid expected to flash in'depressurizng from 689 to 276 kPa is 12,400kgx145.17f kJ

= 871 kg 2068 g

and the core uncovery could be expected at 12514 + 12486}= 25873 s 1

1 These two estimates represent a range of the core uncovery time. Both l[ '

extremes of the range would be increased if credit were taken for any of the remaining liquid in the loops.

l 3.2. Results of the Large Break Calculation i

The times at which significant events occurred in the calculation are

listed in Table 4.

Calculated pressures for the upper plenum, intact loop steam

generator, and intact loop accumulator are shown in Figure 8. The upper plenum pressure dropped almost instantly from its initial value of' 15,500 kPa to 6,760 kPa which was slightly-lower than the saturation pressure corresponding to the initial hot leg temperature. The primary coolant temperature decreased slightly during the sudden decompression.

The calculated pressure was similar to that expected in an isothermal 4 ' blowdown because the plant was assumed to be operating at a small (2 K) hot leg to cold leg temperature differential at the start of blowdown. The upper plenum pressure then decreased at a more gradual rate and reached 15

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TABLE 4. CHRONOLOGY FOR THE 200% COLD LEG BREAK CALCULATION Time .

(s) Event 0.0 Blowdown initiated 4.6 Pressurizer emptied 5.3 Scram signal generated 6.0 Reactivity insertion due to scram initiated 13 Fuel rods dried out 13.3 Accumulator flow initiated 22 Minimum vessel coolant inventory occurred 35 Lower plenum refilled 35.3 Auxiliary feedwater initiated 52 Fuel rods quenched 80 Accumulators emptied 82.9 Calculation terminated 6

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1 containment pressure near 24 s. The pressure in the intact loop steam generator secondary, which was representative of both steam generators, decreased 30% during the calculation. Most of the pressure decrease occurred after 35 s and was due to heat transfer from the steam generator to the primary coolant, and to a lesser extent, the injection of cold

, i auxiliary feedwater. The pressure in the intact loop accumulator, which was represertative of the accumulators in both loops, began decreasing at 13 s when the pressure of the primary coolant dropped below the nitrogen cover pressure allowing ECC to flow into the primary system. The acc.umulator pressure decreased smoothly until the tank emptied at 80 s when a valve was closed to prevent air from entering the primary coolant system. RELAP4/M007 does not accurately represent air flow.

The total fluid mass in the reactor vessel, divided by the initial 4

mass , 9.689 x 10 kg, is shown in Figure 9. The vessel coolant inventory decreased rapidly as mass exited through the large cold leg break. Flow from the accumulators initiated at 13 s but most of the ECC water was O bypassed out the break until 22 s when the vessel was essentially voided

( ) with an average void fraction of 0.99. By 24 s the pressure of the primary coolant dropped below that of the containment which resulted in backflow j from the containment and the end of ECC bypass. The reflood of the core was initiated at 35 s when the lower plenum was refilled with liquid. The core was completely reflooded at 52 s. The vessel was almost full of 5

subcooled liquid at 80 s, when the accumulators emptied, although a i

two-phase fluid remained in the upper head and control element assembly regions (Volumes 12 and 11, respectively, Figure 2). The vessel contained a more mass at the end of the calculation than at the start because the fluid was cooler and denser.

The calculated fuel rod cladding temperature and corresponding fluid

$ temperature for the uppermost core heat slab are shown in Figure 10. The fluid temperature generally represented saturation temperature except after

., 52 s when the core fluid was subcooled. The calculated cladding temperature shown in Figure 10 was representative of the entire core. The

m calculated cladding temperature closely followed the fluid temperature

! V) t 17

until 13 s when the fuel rod dried out. The fuel rod temperature then departed from the fluid temperature although cooling of the fuel generally continued. The lower plenum was refilled at 35 s initiating core reflood ,

which caused a sudden decrease in cladding temperature of at least 40 K throughout the core. The core was reflooded and the fuel rods completely

(

quenched by 52 s.

l The power generated in the core due to fission plus radioactive decay had a relatively small effect upon the transient as illustrated by Figure 11 which compares the core power with the total power transferred to the primary coolant from the steam generators, fuel rods, and passive vessel heat structures. The total power transferred to the primary coolant was generally between one and two orders of magnitude higher than the core power which illustrates the importance of stored energy in the steam generators, fuel rods, and vessel heat structures for this transient. The relative unimportance of the core power is not a surprising result since the calculation was initiated at only 5% rated power and hence nearly represented an isothermal blowdown. At the end of the calculation, the steam generators and the reactor vessel wall were each transferring about )

10 times more power into the primary coolant than the decay power. Ibd calculations indicated that the vessel wall would take about 10 minutes to cool dcwn to the primary coolant temperature while the steam generators would taka about 25 minutes to cool down. The energy stored in these components was equivalent to about 350 minutes of decay power. The stored energy in these components must be modeled well to perform accurate core boiloff calculations.

MM t of the total power transferred to the primary system, see Figure 11, was from different components at different times during the calculation. For example, most of the heat transfer between 6 s and 20 s was from the steam generators. The voiding of the primary coolant system ,

thermally isolated the steam generators which reduced the total power to the primary coolant until core reflood was initiated at 35 s. The i subsequent removal of stored heat from the fuel caused a large increase in power transferred to the primary coolant. The reflooding of the core also 18

I p resulted in the entrainment of a small amount of liquid from the core to '

'( the steam generators where the liquid was evaporated resulting in -

significant heat transfer to the primary system. The steam generator heat transfer was limited by the amount of liquid entrainment. Consequently, the. calculated' liquid entrainment is a significant parameter in determining the cooldown rate of the steam' generators. The heat transfer from the fuel rods. decreased after the completion of core reflood at 52 s. At the end of the calculation, most of the heat transfer to the primary coolant was from the steam ge; 7 mtors and the reactor vessel wall.

The calculation was terminated at 82.9 s which was 2.9 s after the j accumulators emptied of' liquid. The state of the primary coolant system at

{ the end of the calculation is described below. The downcomer, lower

) plenum, core, and upper plenum were filled with subcooled liquid whereas l two-phase mixture remained in the upper head and control element assembly regions. The hot legs and pump discharge leg of each unbroken loop were also filled with liquid. The pump suction legs and steam generator U-tubes were mostly filled with steam. The fuel rods were quenched. The steam

/ generators and vessel heat slabs were contributing more power to the i primary coolant system than the reactor fuel. The calculated refilling of the vessel above the cold leg and hot leg nozzles may be unrealistic since the vessel level would probably not rise above the top of the nozzles in an accident. A more detailed nodalization in the vicinity of the nozzles and/or the use of a bubble rise model to track vessel levels would be i required to more accurately represent fluid behavior above the nozzles t

during the reflooding of the vessel.

A preliminary calculation was performed with warm water in the accumulators to determine the effect of ECC temperature on the response of

! the primary coolant system. This preliminary calculation was identical to l

., the calculation described above except that the initial ECC temperature was

372 K instead of 322 K and a different water packing model was used as 3 described later. The ECC temperature had a significant effect upon the calculated vessel fluid inventory as shown by Figure 12. The reactor p vessel refilled faster and less ECC water bypassed out the break with cold iV j

ECC water. Nearly 80% more liquid was in the vessel at the time the s

19

accumulators emptied in the calculation with cold ECC water. This additional liquid in the vessel would require more than 10 hours1.157407e-4 days <br />0.00278 hours <br />1.653439e-5 weeks <br />3.805e-6 months <br /> to be boiled away by decay heat. i e

l Figure 13 shows the effect of ECC temperature on upper plenum pressure which explains the effect of ECC temperature on vessel inventory shown in Figure 12. The injection of cold ECC caused more condensation which resulted in a lower primary coolant pressure compared to the calculation with warm ECC. When the primary coolant pressure was below containment pressure, such as between 25 s and 35 s, steam flowed from the containment to the downcomer which prevented ECC bypass and promoted vessel refill.

The initiation of core reflood at 35 s and subsequent removal of stored energy from the fuel rods and steam generators produced steam which pressurized the primary coolant. In the calculation with cold ECC water, the primary coolant pressure again decreased below containment pressure shortly af ter the core was quenched at 52 s which prevented any further ECC bypass and promoted vessel refill. In the calculation with warm ECC, less condensation occurred which allowed the heat transfer from the fuel rods and steam generators to maintain the primary coolant pressure above containment pressure which promoted ECC bypass and even resulted in a reduction of vessel inventory. The major difference between calculations was that less condensation occurred with warm ECC which allowed the primary coolant pressure to generally exceed containment pressure thus promoting ECL bypass.

RELAP4/MG97 has two water packing models. One model, called the first water packing model, was available in earlier versions of the code such as RELAP4/M006. An alternate water packing model was developed an i recomended for use in RELAP4/ MOD 7. The computer run time was found to be sensitive to the selection of the water packing model although the calculated results were not sensitive to model selection. The calculation ,

with the cold (322 K) ECC water was run with the first water packing model and averaged 180 cp s per real s. The calculation with the warm (372 K)  :

ECC water was run with the alternate water packing model before 59 s and with the first model af ter 59 s. The warm ECC water calculation 20

<x averaged 1000 cp s per real s between 45 and 58 s and 60 cp s per real s

/ I

/ after 59 s. Changing from the alternate to the first water packing model improved the computer run time by nearly a factor of 20. The calculation with the alternate water packing model ran slowly because the time step control algorithim selected the input minimum time step of 2.5 x 10-5 3 to satisfy the saturation line crossing criterion as a result of liquid filling control volumes. Discussions with code development personnel indicate that the use of the alternate water packing model may be inconsistent with the code's time step control algorithim for some trans ien ts . The alternate water packing model probably could have been used economically with a large minimum time step.

O)

v i

4 21

l

4. CONCLUSIONS The following conclusions were derived as a result of the calculations. ,

1. For the small break, core uncovery time estimates, assuming no e

remaining loop liquid migrates to the vessel, were 18178 s using the break flow rate at problem termination time, and 25873 s using the expected equilibrium break flow rate.

The small break calculation was terminated at 12,514 seconds for economic reasons and estimates of core uncovery time were made by hand calculation. Both of the above estimates would be increased if some of the remaining loop liquid was assumed to migrate to the vessel.

2. The stored energy in the heat structures and steam generator secondaries was important in the large and small break calculations because of the low initial core power.

At the end of the small break calculation heat addition from the heat structures was approximately 30% of the core power and heat addition from the secondaries was approximately 10% of core power. At the end cf the large break calculation, the power deposited in the primary coolant from the steam generators and the reactor vessel wall was about 20 times larger than t!e core decay power (see Figure 11.). The remaining stored energy in the steem generators and reactor vessel was equivalent to about 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> of decay power. The stored energy in these components must be modeled well to perform accurate core boiloff calculations.

3. ECC temperature had a significant effect upon the ECC bypass and vessel refill results for the large break.

Calculations performed with cold (322 K) and warm (372 K) ECC water showed that the condensation caused by cold ECC water affected the t differential pressure between the primary coolant and the containment (see Figure 13) which reduced ECC bypass resulting in substantially more liquid 1

22

s in the vessel when the accumulators emptied. The additional liquid in the

~

vessel with the cold ECC would reauire more thaa 10 hours1.157407e-4 days <br />0.00278 hours <br />1.653439e-5 weeks <br />3.805e-6 months <br /> to be boiled away Dy decay heat.

4. A detailed nodalization and/or the use of a bubble rise model may be necessary to accurately track vessel mixture levels once the levels approach tne cold leg and hot leg nozzles during the refill of the vessel.

The dowocomer and upper plenum refilled completely in the large break calculation altnough in reality the vessel would probably not refill above the nozzles.

The following conclusion is applicable to the usage of RELAP4/M007.

1. Tne large break calculation computer run time was sensitive to the selection of water packing model.

As described in Section III.3, computer run time for the large break calculation was improved by a f actor of 20 wheh the first water packing model was used instead of the alternate water packing model. The alternate water packing model, which was developed for use with RELAP4/M007, may not be consistent with the code's time step Control algorithim for some transients. Tne economical use of the alternate water packing model may require the use of large minimum time step sizes.

P y

l 23 l

5. REFERENCES

1. S. R. Behling et al., RELAP4/M007--A Best Estimate Computer Program to Calculate Thermal and Hydraulic Pnenomena in a Nuclear Reactor or Meiatea dystem, tusu laano inc, fiURLU/UK-199d, August 196I. *

2. C. B. Davis and C. D. Fletcher, Audit Calculations of Open Power Operated Relief Valves in a Combustion Engineering PWR, EGG-CAAP-5054, y Novemoer 1979.

3. Final Safety Analysis Report, San Onofre Nuclear Generating Station Units 263, Southern California Edison Company and San Diego Gas and Electric Company.

4. Code Assessment Branch, Assessment of the RELAP4/M006 Thermal-Hydraulic Transient Code for PWR Experimental Applications, EGG Idano Inc., CAAP-TR-78-035, December 1978.

5. RELAP4/M006 A Computer Program for Transient Thermal-Hydraulic Analysis of Nuclear Reactors and Related Systems User's Manual, EG&G Idaho, Inc., COAP TR 003, January 1978.

O i.

O 24

i

) b )  !. )

~.- ._.- -

RELIEF 53 MAIN STEAM 3 51 g PRESSURIZE R STEAM __

GENERATOR '2D 9

5 8

09) HAIN FEED 62 -

' k I35 52 3 RELIEF MAIN 39 _ $ $

FEED h h_ L2) 46 ~ 134 STEAM g yg VESSEL 22@ g 12 g 9 13 GENERATOR b

42 23 11 Id 19g il 3

63 FEED

_  ? '

j To MAIN 1 R 131 ,

(x # _J-50 FEED g @21 @ @ T2[-

~

14T 3,

3 4@

pggp 40 s I h 5 8 51

- 2k\

W M [125 '

_6 h CONTAINMENT 45 @ PUMP 4

10 1 2 h

'g__..sl6 3

e 515 9 gm a s

W s

[ 2 BREAK h Volume ACCUMULATORS 15 N

@ ileat Slabs

@'jk _

94 I -b d 29 PUMP 93 3_

X '

Junction 41 4

W Denotes Wilson Bubble Rise Model Ck 5

Denotes Vertical Slip Model '

Figure 1. Nodalization for the small break calculation.

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R 4 3 N ^

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-. . .. . .- ~_ . - . _ _ . . - . . .

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l' e 20000 i i j{ . UPPER PLENUM

) g. --- STEAW GENERATOR SECONDARY 2500

'5000 -

1 I

^-

Initial upper plenum pressure -

m ,

g O

-2000 0 2 Q. -

M d -

O.

• -500 '

10000 --

• 3 u 4 = 3 I H .

• g l 4 Initial S. G. secondary pressure 1000 m l 0.

a.

4 5000 -

" .i , ..

. -500 7 ----....................

i 0 O 3

0 5000 10000 15000 l Time (s) l Figure 3. Calculated upper plenum and steam generef or

pressures -- small break.

1

14.5 , ,

( Normal level -47 34 -

.{ TW --4s

^

n E -

v -

-45 O C

j C 1 -

0 13.5ar -

0

?

44 =

? I o i e '

u q -43 C 4 13 <- -

l-  : Auxiliary feedwater initiated i -

I

-42 i

, 12.5 -

! O 5000 10000 15000 l

Time (s)

Figure 4 Calculated secondary mixture level -- small breck.

i i~ 27 i

~ 1

O A . 4

- 12 g Top of upper plenum 1 _-10 3-O E 8 O v

c --

e g ot leg elevations

. 2 H

-s $0 0 -

l -4 g .

-- \ Top of core 2

'If ' *

' 0 0

0 5000 10000 15000 Time (s)

Figure 5. Calculated upper plenum mixture level - smcil breck.

' 8 i

Normal i

-25 O

level s- , _-20 v

b v

-15 C C 0 4 -

0

- , =

0 1 o

$ -10 g W W 2 -

Accumulator -5 emptied 0 O f, 0 5000 10000 15000 Time (s)

Figure 6. Calculated cccumulator level - small break. l O

28

.=- _. - ._

i

. I i

i i

+

1 4,

4

e a

I i

,} V 4

I '

a j 0.8 -

4 5 ' .

14

I s L 0.6 -

j >

t

=

c 0 f 1

3 I j O 0.4 -

[, -

4 I ,

4  : j 1

i 0.2 -

[ p -

i O

3 0 5000 10000 15000 i Time (s)

! Figure 7. Calculated leak quality -- small breck.

i 1

20000 , , ,

3 UPPER PLD4UM j . --- STEAM GENERATOR SECONDARY 2500 j -- ACCUWULATOR l 15000 - -

g -2000 7 Q.

.x 7o.

j -

v

! ' 10000-- - 1500 e 4 u m O y 1 m .

i " a

• e 1

.t000

[ ~~~~- ------..... .....,___________---~~-~~ --- g 5000 - -

---

* - - ECC injection 3o0 i -s. initiated Accumulator i

_ ~. . ,_,---- empti ed ..~. .-._

0 f)

, 0 25 50 75 10 0 0

j Time (s)

{ Figure 8. Calculated upper plenum, steam generator, and j } accumulator pressure -- large break.

I i

f i V i

i i 29 i

S f.5 , , , ,

Accumulator 4 E emptied O

e N -

._ 1

=

C

+--- Core reflooded C

O 0.5 - -

o 0

L u.

g End of bypass ' ' '

i I

0 O 20 40 60 80 20 Time (s)

Figure 9. Calculated vessel coolent inventory - iorge breck.

600 , , ,

O FUEL CLADDING

--- FLUID n 550 -

Dryout

~

4

' OU

, Reflood initiated *

$ ',M a 500 - .

2 8 '., 400 2 O s y/$y e

\,

[e 4m [

, , Core reflooded e

-300

/

400 -

i l6 *,

200 350 O 25 50 75 10 0 4 Time (s)

Figu r e 10. Calculated core temperatures - targe break.

(

O 30

f L]

a 2000 i i 4 i I

CORE

--- TOTAL Reflood 1500 -

initiated L

^ $

3 -

3, 1000 -

,! Core reflooded -

u ,- .

e .. .

.l 3:

,- l ',i, ..

l.

...e'.

$ 500 E j '. ' A .' ',' -

l

, s

'j, f,, , ' l , .s ,,,

. , . . .', l 4 -

0 - -

-500 O 20 40 60 80 10 0 Time (s)

F i gu r e 11. Calculated core end total power -- large breck.

5 V

1.5 , , , ,

COLD ECC 2 --- WARM ECC D

E 0

g 1 e Accumulator

\ -

emptied

~

O

• c ,,...

• ' '~, ,

o , ,

- 0.5 - ' ' ~

,l ' ' , , -

u o

.,~---- '

o u

L., ,-

l

'- 1 o

5 0 20 40 60 80 l

10 0 Time (s) l Figure 12. The eff er:t of ECC terrporoture on vessel coolant g inventory - large breck.

K 31

1 I

I 9

4 L

900  ; , , ,

-  ! UPPER PLENUW (CCLD ECC) -12 0 l --- UPPER PLENUM (WARM ECC)

'. - - CONTAINWENT

! -100 m i O o 600 -

! Flow to Containment _  ;;

-so S C

• ; u 5 ., so S i '

o E .\

u 3g i

>3 l s'l- w p j

/ MReflood initiated 20 Flow from Containment 0 o 0 25 50 75 10 0 Time (s)

Figu r e 13. The e f f ec t of ECC temperature on upper plenum pressure -- large breck.

l r

l t

i I

1 O

l 32 1