PLA-3630, Responds to NRC Request for Addl Info Re Requested Rev to Tech Specs for Facility Ses Related to Rwcu/Hpci/Rcic Temp Based Steam Leak Detection Isolation Setpoints

Responds to NRC Request for Addl Info Re Requested Rev to Tech Specs for Facility Ses Related to Rwcu/Hpci/Rcic Temp Based Steam Leak Detection Isolation Setpoints

ML18026A410
Person / Time
Site:	Susquehanna
Issue date:	08/19/1991
From:	Keiser H PENNSYLVANIA POWER & LIGHT CO.
To:	Butler W Office of Nuclear Reactor Regulation
Shared Package
ML17157A805	List: ML17157A804 ML17157A806 ML17157A807 ML17157A808 ML17157A809 ML17157A810 PLA-3630, Responds to NRC Request for Addl Info Re Requested Rev to Tech Specs for Facility Ses Related to Rwcu/Hpci/Rcic Temp Based Steam Leak Detection Isolation Setpoints
References
PLA-3630, NUDOCS 9108260071
Download: ML18026A410 (35)

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ACCELERATED DISTRIBUTION DEMONSTRATION SYSTEM REGULATORY INFORMATION DISTRIBUTION SYSTEM (RIDS)

ACCESSION NBR:9108260071 DOC.DATE: 91/08/19 NOTARIZED: NO DOCKET FACIL:50-387 Susquehanna Steam Electric Station, Unit 1, Pennsylva 05000387 50-388 Susquehanna Steam Electric Station, Unit 2, Pennsylva 05000388 AUTH.NAME AUTHOR AFFILIATION KEISER,H.W.

Pennsylvania Power 6 Light Co.

RECIP.NAME RECIPIENT AFFILIATION BUTLER,W.R.

Project Directorate I-2 SUEGECT:

Responds to NRC request for addi info re requested rev to tech specs for face.lity SES related to RWCU/HPCI/RCIC temp based steam leak detection isolation setpoints.

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0, Pennsylvania Power 8 Light Company Two North Ninth Street~Altentowri, PA 18101-1179~215/774-5151 Harold W. Kelser Senior Vice President-Nuclear 215/774%194 Director of Nuclear Reactor Regulation Attention: Dr. W. R. Butler, Project Director Project Directorate I-2 Division of Reactor Projects U.S. Nuclear Regulatory Commission Washington, D.C.

20555 SUSQUEIIANNA STEAM ELECTRIC STATION RESPONSE TO REQUEST FOR ADDITIONAL INFORMATIONON PROPOSED AMENDMENTS 138 TO LICENSE NO. NPF-14 AND 92 TO LICENSE NO. NPF-22: REVISIONS TO TEMPERATURE LEAKDETECTION RWCU/HPCI/

RCIC SETPOINTS PLA-3 30'ocket Nos. 50-387 and 50-388

Dear Dr. Butler:

This letter is in response to the NRC Staff's request for additional information regarding our requested revisions to the Technical Specifications for Susquehanna SES related to RWCU/HPCI/RCIC temperature based steam leak detection isolation setpoints.

The following material provides additional background explanation for our requests in PLA-3487 dated January 9, 1991. Later material provides specific responses to the Staff questions contained in an NRC letter dated June 13, 1991.

The actions requested by PP&L in its Proposed Amendments 138 (NPF-14) and 92 (NPF-22) resulted from efforts to reconstitute the design bases for the temperature based steam leak detection and isolation circuitry in rooms within secondary containment which interface with the reactor coolant system as a result of the response to NRC Violation 88-15-01.

They were also based on operating experience considering the normal seasonal variations of reactor building temperatures and design bases temperatures listed in FSAR Chapter 3 for maximum room temperatures under postulated accident conditions.

Reconstitution of design bases started with modeling each room in which steam leak detection circuitry is installed, and calculating room temperature response to postulated leak rates.

Initial analyses used leak rates of 5 gpm, and subsequently, 25 gpm.

We used our Compartment Transient Temperature Analysis Program (COTTAP) computer program, which is discussed

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V FILES R41-2, A17-2 PLA-3630 Dr. W. R. Butler below in our response to Staff Request 2, to generate room thermal response curves under leak conditions.

The response also provides summary results of the calculations.

Our response (below) to Staff Request 1 provides a functional basis for considering design bases leakage rates of 25 gpm for automatic isolation, and our response to Staff Request 3 provides other precedents for selection of a 25 gpm leak rate as a design bases.

RE ET 01 It is not clear why a change to the existing temperature setpoints is necessary.

The licensee stated that the existing setpoints willnot allowfor timely detection ofa 5 gpm leak, but the licensee did not dePne what leak rates the existing setpoints willdetect.

The current capability must be dered and the acceptability ofthe existing condition must be addressed.

R~NE The majority of the Technical Specification temperature based isolation setpoints in the RWCU area, HPCI area, and RCIC area remain unchanged by our request.

The changes that were requested include high ambient and high differential temperature setpoints in the RWCU penetration room, and high ambient setpoints in the HPCI/RCIC room cooler air inlets.

In the RWCU area, the proposed change would eliminate an excessively small margin between isolation setpoints in the RWCU penetration room and peak summer temperatures.

(The setpoints for the RWCU Pump rooms and RWCU Heat Exchanger rooms will remain at their present values.)

At Susquehanna SES, several inadvertent isolations of RWCU have occurred, with resultant impact on plant chemistry and pump seal performance.

The most recent inadvertent isolation occurred in June 1991 on Unit 2 and was reported to the NRC.

Since Susquehanna SES began operation, several isolations of RWCU have occurred because of inadequate margin between isolation setpoint and peak temperature in the RWCU penetration room. Even without occurrence of an actual isolation, summer temperatures in the RWCU penetration room without adequate margin invokes unnecessary challenges to plant operations personnel.

In the HPCI area the proposed change to the two high temperature isolation circuits would eliminate an inconsistency between those circuits with setpoints of 147'F (mounted on air inlets to the room coolers), and two other circuits with setpoints of 167'F (wall mounted).

Room cooler outlet air does not impact directly on any of the temperature monitoring elements used for isolation actuation.

Both the local and wall mounted temperature circuits would be subject to similar temperature environments in the event of a room steam leak and should have the same setpoint.

~a FILES R41-2, A17-2 PLA-3630 Dr. W. R. Butler The HPCI room temperature in the event of an accident is described in FSAR Table 3.11-6 as 130'F maximum.

The smaller margin between setpoint and maximum room temperature under these conditions (when an inadvertent isolation would be unacceptable) argues that the two circuits in question should have their setpoints raised to match the two wall mounted temperature circuit setpoints.

All of the remaining high ambient temperature and high differential temperature circuit isolation setpoints in the HPCI room and in the HPCI pipe routing area would remain at their present value.

The RCIC area steam leak detection circuity is functionally identical to the HPCI area circuity, and is covered by the discussion above.

The requested changes to selected Technical Specification setpoints have a different motivation from our requests to redefine the design bases for the bulk of the temperature based isolation setpoints as described in our FSAR. The requested setpoint changes are intended to eliminate anomalies and inconsistencies for a limited number of setpoints.

Redefinition of the design bases for all temperature based isolation setpoints represents the culmination of a two year effort to establish an analytically consistent and uniform design bases for the bulk of the existing temperature based isolation circuit setpoints without changing those setpoints.

Our analyses defined room temperature response for defined leak rates under both summer and winter conditions.

The specific response in each room was different and is discussed for the HPCI, RCIC, and RWCU rooms below.

For all three rooms, the temperature rise resulting from a postulated 5 gpm leak was distinct but not significantly different from temperature variations due to seasonal differences or to loss of HVAC functions. Actual room temperatures uhder non leak conditions vary on a seasonal basis by as much as 25 to 30 degrees.

This variation is a substantial fraction of the temperature rise calculated for a 5 gpm leak.

Thus, establishing high ambient isolation setpoints based on 5 gpm leaks under winter conditions would result.,in an unacceptably small margin between the isolation setpoint and the high room temperatures expected in summer.

Conversely, establishing the setpoint based upon the initial room temperature under summer conditions would'produce an unacceptably long time to reach the trip point under a postulated 5 gpm leak in initial winter conditions.

The opposite relationships exist for the high differential temperature circuits.

These conclusions, which raised question on the acceptability of a postulated 5 gpm leak rate as the design bases, were reported to the NRC as an emerging design issue under the provisions of 10CFR50.9 in PP&L letters PLA-3214 dated 7/24/89, PLA-3315 dated 1/16/90, and PLA-3443 dated 9/21/90.

These letters also discussed concurrent analyses at higher projected leak rates which demonstrated the adequacy of the existing Technical Specification setpoints to isolate the leaking line within a reasonable timeframe, and to protect the plant and the public.

FILES R41-2, A17-2 PLA-3630 Dr. W. R. Butler The process for establishing safe and effective setpoints for the leak detection circuitry followed the following strategy.

We first established the status of using 25 gpm leak rates in the industry by contacting GE Company, and other licensees.

(See response to Question ¹3.)

We performed offsite dose calculations to show the acceptability of the leak size (see response to Question ¹5). We then calculated the room transient thermal response with 25 gpm leaks, and determined the temperature reached at the end of a 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> period.

We considered that temperature as an Analytic'Limitand performed a setpoint calculation (see response to Question ¹4.)

Where the resultant setpoint was greater than the present setpoint listed in the Technical Specifications, and where adequate margin to unnecessary isolations currently exists, we reselected the Analytic Limitat a time shorter than the 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> point to retain the present setpoints.

For most of the rooms, this time was four hours or less.

For the RWCU penetration room, thermal response required setting of the Analytic Limitat the 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> point.

Specific data for the thermal response capabilities of the three areas is discussed below:

HPCI Area:

The current technical specification high ambient trip points are 167'F.

Calculations indicate that the temperature rise resulting from a 5 gpm steam leak at rated process conditions would take in excess of 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> to reach this setpoint.

Calculations were reperformed for HPCI with a presumed leakage rate of 25 gpm at rated process conditions. Temperature response shows the existing setpoint (167'F) is reached at about 18 minutes and the analytic limitis reached in 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.

We estimate that a smaller leak rate of 12-15 gpm (rated temperature and pressure conditions) would cause the setpoint to be reached in 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />; The high differential temperature response is similar to the high ambient temperature response.

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It The HPCI room cooler inlet high ambient temperature setpoints are inconsistent with the HPCI wall mounted high ambient temperature setpoints as discussed above.

Although

'oth sets of circuitry monitor the same environmental conditions, the air cooler inlet temperature circuit is set 20'F lower than the wall mounted circuit.

The proposed change increases the margin above maximum design room temperature, and therefore reduces the possibility of an inadvertent isolation particularly during an accident when room temperatures are elevated.

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'I FILES R41-2, A17-2 PLA-3630 Dr. W. R. Butler RCIC Area:

The current technical specification high ambient trip points are 167'F.

Calculations indicate that the temperature rise resulting from a 5 gpm steam leak at rated process conditions would require up to eight hours to reach this setpoint.

Calculations were reperformed for RCIC with a presumed leakage rate of 25 gpm at rated process conditions.

The temperature response shows the existing setpoint (167'F) is reached in about 10 minutes and the analytic limitis reached in under 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.

We estimate that a smaller leak rate of 10 gpm (rated temperature and pressure conditions) would cause the setpoint to be reached in.4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />. The high differential temperature response is similar to the high ambient temperature response.

The discussion for the RCIC high ambient temperature on the room cooler inlet is the same as that discussed for HPCI above.

RWCU Area:

The current technical specification high ambient trip setpoints are 147'F in the pump room circuits and heat exchanger room circuits.

They are 118.3'F in the penetration room circuits from which most of the spurious isolations have occurred.

Calculations indicate that the temperature rise resulting from a 5 gpm leak at process conditions, starting from initial winter room thermal conditions, would be insufficient to reach all existing setpoints within a 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> period.

The analyses of a 25 gpm leak at rated pressure and temperature conditions, starting from a winter condition showed that the proposed setpoint of 131'F would be reached in less'than 1.5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> and the analytic limit would be reached in 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />.

The time required to reach the isolation setpoint (147'F) for the RWCU pump room and the heat exchanger room would be 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> and 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> respectively.

The analysis for the differential temperature setpoint is similar to the high ambient except that the longer time response is associated with initialsummer conditions (low delta T).

RE ET 2

The licensee credits certain temperature calculations in its safety analysis, but the details ofthe temperature calculations were not provided for stag review. It is not clear to the staff what assumptions were made and whether those assumptions are acceptable or not.

Also, specij7c details regarding the application ofthe computer model COTTAP were not provided for staff review.

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FILES R41-2, A17-2 PLA-3630 Dr. W. R. Butler R

Attachment A contains a user's manual for the COTI'AP computer code and copy of a paper recently published in Nuclear Technology which describes the methodology used in the COTlAP program and presents some of the verification calculations which have been performed.

The user's manual presents some of the calculations which were performed against problems that have exact analytical solutions.

The referred paper presents the methodology along with calculations which have been benchmarked against calculations performed with the CONTEMPT computer program.

In addition, the program and computation package have been independently reviewed by Gilbert Associates.

PP&L also maintains a Quality Assurance file/package for the COTTAP computer code.

Attachment B contains a summary of the calculations which were performed for each room and upon which the revised temperature setpoints were based.

Calculations were performed for each room under a variety ofconditions (for example, summer and winter initial conditions and various break sizes) and have been independently reviewed.

The attachment presents a summary which includes the methodology and assumptions for each calculation along with the representative results which were used to calculate the revised setpoints.

RE ET The licensee arbitrarily selected 25 gpm as the design basis leak rate forall areas, stating that the 25 gpm leak basis is consistent with GE design specifications, with the basis used at other BWRs and with the Technical Speci(Ication Improvement Program (TSIP).

The TSIP cannot be credited since the program currently has not been approved; and details relative to the GE design spectftcations and other BWRs was not provided for stagconsideration.

The effects of the design basis leak on equipment, emergency operating procedures and personnel were not discussed (assuming prolonged operation near the design basis'eak rate).

Additionally, no discussion relative to ASME Code requirements was pr'ovided.

A i'election of25 gpm as the design bases leak rate was considered when calculations using the FSAR values of 5 gpm leakage were unable to support a consistent methodology for setpoint selection in all rooms with steam leak detection functions.

These issues were reported to the NRC (see response to Request I), and investigation into a 25 gpm leakage rate as design bases was started.

Decision to select 25 gpm as the design bases leakage value for HPCI/RCIC/RWCU area rooms was based on the following:

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pt FILES R41-2, A17-2 PLA-3630 Dr. W. R. Butler A limitingcriteria was that the resultant high ambient temperature setpoints for system isolation had to include significant margin above maximum room temperatures under all non leak conditions to preclude inadvertent isolations. The computer modelling capabilities demonstrated that the previously assumed leakage value of5 gpm could not meet this criteria for most of the rooms examined.

This conclusion applied equally to differential temperature setpoints.

An assumed 25 gpm leakage rate allowed a uniform approach to setpoint calculation, retention of most of the existing setpoints which incorporate adequate margin, and allowed raising those few setpoints (as requested) where insufficient margin existed.

Retention of the existing setpoints where analysis indicated some increase could be justified by a 25 gpm design bases leak also avoided possible conflicts with fire suppression initiation setpoints in those areas which included such systems.

2.

Leak detection can be considered as a safety function with the purpose of minimizing or precluding the potential for a high energy line break (for which independent and diverse detection and isolation systems exist). FSAR Table 5.2-10 correlates leak rates to crack size up to cracks associated with unstable piping rupture for different pipe sizes and stresses.

A leak rate of 25 gpm can be seen from that Figure to be less than those leak rates associated with the onset of unstable pipe rupture.

3.

An assumed leakage of 25 gpm for calculating isolation setpoints was consistent with recommendations provided in GE document EDE-17-0689.

GE advised that they were using 25 gpm as a design bases on all recent design activity.

Other existing BWR designs have been accepted using a leakage rate of 25 gpm as a design bases for leak detection.

These include Perry, Grand Gulf, Clinton, and River Bend.

5.

Off site dose calculations using leak rates of 25 gpm demonstrated acceptable safety consequences.

See the response to Request 5.

After notifying NRC of our findings in 10CFR50.9 reports, room thermal analysis and setpoint calculations continued with the 25 gpm leak rate value.

The effects of the new design bases leak rate on equipment, procedures, and personnel were assessed and found to be minimal.

The operating procedures require operator rounds into the HPCI and RCIC areas once per day.

All areas with steam leak detection circuitry have their temperatures (and differential temperatures) available in the main control room for monitoring.

I 1,'i FILES R41-2, A17-2 PLA-3630 Dr. W. R. Butler The alarm response procedures identify specific action required including observation, confirmation, isolation, and repair ofleaks.

Visual observation ofa steam leak, or rising room temperatures, or the occurrence of a pre-isolation temperature alarm in the main control room would invoke operator action without attempting to quantify the leak rate, or waiting for the temperature to reach the isolation setpoint.

Prolonged operation with any significant leak is not anticipated.

The emergency procedures are symptom based, and not event based. Ifa steam leak occurred, it would be isolated and repaired.

Therefore, an increase in the analytic design bases leak rate would have no effect on emergency, alarm response, or operating procedures.

An increase in the defined design bases for leak detection should have no effect on personnel.

Leaks willbe detectable by the operational considerations mentioned above, and by diverse alarm systems such as area radiation monitors at levels far below design bases leakage rates.

Operators would not enter areas with direct evidence of leakage, except if required as part of a planned evolution under controlled conditions with appropriate protective equipment. Therefore, the proposed changes should have no effect on station personnel.

Allequipment required to function within the environmental zone of the leak is included in our equipment qualification program.

The equipment is qualified for the effects of a high energy line break - HELB.

Ifa leak were to occur, the system would be isolated and ifappropriate an LCO (for HPCI and RCIC) would be entered.

The faulty pipe willbe repaired in accordance with the Susquehanna SES Welding and Non-Destruction Examination Manual. This manual is based on the requirements of ASME Section 3 (NC-2500, ND-2500) and Section 9.

The repair willbe inspected in accordance with ASME Section ll. Prior to declaring the system operational, an evaluation of the leaks effect on other area equipment willbe conducted and appropriate action taken.

E T 4

The methodology used in establishing the temperature setpoints was not described in detail, including consideration for instrument errors; the licensee did not describe to what extent industry standards were being used in establishing the temperature setpoints; and the logic used in selecting the system process conditions was not explained.

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I'i FILES R41-2, A17-2 PLA-3630 Dr. W. R. Butler R~H The setpoints are calculated using the method in PP&L Design Guide, "Instrumentation and Control Setpoint Calculation Methodology".

This safety related setpoint process utilizes the General Electric methodology and was used by General Electric in the calculation of the Susquehanna SES Technical Specification values.

Generic setpoint calculations based upon the PP&L Design Guide were prepared to establish the process for calculation of isolation setpoints for both high ambient temperatures and high differential temperatures.

An analytical limitwas defined based upon the thermal response curves calculated using the COTI'AP room models.

The Allowable Value, Trip Setpoint (Tech Spec Setpoint), and Process Setpoint (as-installed setpoint) were defined starting from the Analytic Limit.

The margin behveen the Analytic Limit and the Allowable Value accounts for instrument and calibration inaccuracy.

The margin between Allowable Value and Trip Setpoint accounts for instrument drift. Driftvalues are derived from manufacturer's specified drift accuracy or from historical plant data if appropriate.

The margin between Trip Setpoint and Process (as-installed) setpoint, usually based on the drift value, provides additional assurance that actual setpoints would not drift above Technical Specification Allowable Values.

Setpoint calculations were then completed for individual rooms, after selecting an Analytic Limit from the room thermal response curves calculated with the COTI'AP code.

For the thermal calculations, leaking fluid was assumed to be at the process pressure and temperature conditions that would exist during normal power operation of the system that was presumed to be leaking.

Because each individual room temperature response has a unique time dependent function, different times had to be selected for each room to determine the Analytic Limit.

Thermal response curves were calculated for both summer and winter conditions permitting the High Ambient Analytic Limit to be selected from the winter response curve, and the High Differential Temperature Analytic Limit to be selected from the summer response curve since these were the most conservative selections for the respective functions.

The individual room setpoint calculations were then completed as described for the generic calculations above.

Iterations between setpoint calculations and room thermal response calculations were required when initiallyconsidering 5 gpm and subsequently 25 gpm as the design bases leakage.

Where this process produced an Allowable Value and Trip Setpoint greater than prescribed by the Technical Specifications, and where no other anomalous factors existed (such as inadequate margin to worst case room temperatures),

the existing setpoints were left unchanged.

This essentially established an effective design bases leakage rate less than the 25 gpm used for the COTTAP room thermal response calculation, and would FILES R41-2, A17-2 PLA-3630 Dr. W. R. Butler apply to'the HPCI and RCIC room setpoints.

Our response to Request 1 provides estimates of the leakage rates which would cause isolation actuation after a four hour peflod.

RE T

The licensee stated that the radiological consequences of a coolant leak outside primary containment was analyzed, but the details of the analysis was not provided for staff review.

R The radiological consequences of a coolant leak outside primary containment was analyzed in PP&L calculation SE-B-NA-078.

Fifty gpm of reactor grade water was assumed to leak into secondary containment at a concentration of 4.0 uCi/gm Dose Equivalent Iodine-131. This'is the maximum allowable coolant concentration of iodine for Susquehanna SES operation. No credit for removal, holdup or decay was taken. The period of the leak was assumed to be 48 hours5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br /> after which environmental leakage was terminated.

The analysis concludes that the resultant offsite and control room doses fall far below 10 CFR 100 offsite dose limits and 10 CFR 50, Appendix A, GDC-19 control room dose limits.

An analysis ofa reactor steam leak was also conducted in PP&Lcalculation FX-C-DAM-010. A 50 gpm water equivalent steam leak was assumed to occur for a 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> period.

No credit for removal, holdup or decay was taken.

This analysis also concluded that the resultant offsite and control room doses fall far below 10CFR100 offsite dose limits and 10CFR50, Appendix A, GDC-19 control room dose limits.

Calculations SE-B-NA-078 and FX-X-DAM-010 which document the radiological analysis are included in Attachment C.

Ifyou have any questions, please contact Mr. C.T. Coddington at (215) 774-7915.

Very truly yours,

. W. Keiser Attachments FILES R41-2, A17-2 PLA-3630 Dr. W. R. Butler cc:

NRC Document Control Desk (original)

NRC Region I Mr.

G.

S.

Barber, NRC Sr. Resident Inspector Mr.

J.

J.

Raleigh, NRC Project Manager

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ATTACHMENT A

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COTTAP: A COMPUTER CODE FOR SIMULATIONOF THERMAL TRANSIENTS IN SECONDARY CONTAINMENTS OF BOILING WATER REACTORS

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'ARK A. CHAIKOand MICHAELJ. MURPHY Pennsylvania Power dc Light Company, Allentown, Pennsylvania 18101 Received December 1, 1989 Accepted for Publication September 12, 1990 The Compartment Transient Temperature Analysis Program (COTTAP) was developed by the Pennsylva-nia Power & Light Company forpostaccident boiling water reactor (BWR) secondary containment thermal analysis. The code makes use ofpreviously developed implicittemporal integration methods and sparse ma-trixinversion techniques to allow modeling ofan en-tire BIVR secondary containment. Investigations were made with a model consisting of 121 compartments and 767 heat-conducting slabs.

The simulation pre-sented involves the numerical integration of20 101 or-dinary differential equations over a 30-h simulation period. Two hours ofCPU time were required to carry out the calculation on an IBM3090 computer.

The COTTAP code considers natural convection and radi-ation heat transfer between compartment air and walls through a detailed finitedifference solution ofthe slab conduction equations. Heat addition from hot piping and operating equipment, and cooling effects associated with ventilation flows and compartment heat removal units are also included. Additional capabilities of COTTAPinclude modeling ofcompartment heatup re-sulting from steamline breaks and simulation ofnat-ural circulation cooling in compartments with flow paths at differing elevations.

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~:eM48&KL4CF5FZ~~'"-""">V~>ZMSÃ I~ INTRODUCTION Under postaccident conditions, boiling water reac-tor (BWR) secondary containment ventilation systems typically isolate to prevent fission product release to the environment. Since cooled air is no longer circu-lated through the secondary containment, increased compartment temperatures result. Predictions of post-accident compartment temperatures are necessary to determine whether safety-related equipment is sub-jected to temperatures that'exceed its maximum design values. Safety-related equipment must be operable un-der postaccident conditions in order to effect the safe shutdown of the reactor.

After an accident, the secondary containment ventilation system operates in a recirculation mode to promote air mixing between compartments and to dilute locally concentrated radioactive isotopes.

Original design calculations for Pennsylvania Power

& Light Company's (PP&L) Susquehanna Steam Electric Station (SSES) assumed that air recircula-tion provided enough mixing to produce a fairly uniform temperature distribution throughout all sec-ondary containment compartments.

For this reason, a single-compartment transient model was used in the simulation of postaccident conditions. Recent investi-gations based on steady-state calculations have shown, however, that significant temperature variations can exist between compartments.

These temperature variations were large enough to prompt a detailed NUCLEAR TECHNOLOGY VOL. 94 APR. 1991

Chalko and Murphy POSTACCIDENT BWR SECONDARY CONTAINMENTTHERMALANAL/$1$

multicompartment transient analysis of the secondary containment.

To reanalyze the postaccident transient behavior of the SSES secondary containment, PP&L developed the Compartment Transient Temperature Analysis Pro-gram (COTTAP). Development of this program began after an evaluation of available codes revealed that none were capable of performing a sufficiently detailed simulation owing to the large number of heat-conduct-ing structures found in the SSES secondary contain-ment. For example, the CONTEMPT code,'hiCh is probably the most widely used containment analysis program, can model as many as 999 compartments but is limited to 99 heat-conducting slabs.

In contrast, COTTAP can model up to 1200 heat-conducting slabs and 300 compartments. It also contains models that describe heat dissipation from operating electrical equipment and process piping. A COTTAP model of the SSES-1 and -2 secondary containment structures consists of -120 compartments and 800 heat-conduct-ing slabs.

The CONTAINcodex's a more recently developed containment simulation program with complex mod-eling capabilities. It is, however, designed specifically for primary containment simulation and is not well suited for secondary containment modeling because it has no provisions for energy input to compartments from heat loads such as electrical panels, lighting, mo-tors, and hot piping.

A description of the COTTAP code, including as-sumptions, governing equations, numerical solution methods, and code limitations is given in Sec. II. Rep-resentative results of the SSES-1 and -2 secondary con-tainment analysis are presented in Sec. III, and code verification is discussed in Sec. IV.

II. OESCRIPTION OF THE COTTAP CODE II.A. Compartment Mass and Energy Balances The COTTAP code allows for air and water vapor mass transfer between compartments by means of forced ventilation, leakage, and natural circulation flows. A forced ventilation flow model describes heat-ing/ventilating/air conditioning systems, and a leakage model simulates intercompartment flows that are gen-erated by pressure differentials. In addition, a natural circulation model simulates gravity4riven flows between compartments connected by flow paths at differing elevations.

Steam can also be added to a compart-ment as a result of pipe breaks or removed through condensation and rain-out. Airand water vapor mass conservation equations for a compartment with N ventilation paths, NI leakage paths, and N, natural cir-culation paths are given by dp Ivu W

V= Z WuJYuj+ Q WgYIJ+ Z Wcj(YcJ Y) dt JSR 1

and dpw V

di where Nu NI

= g WuJ(1 YuJ) + g Wij(I YIJ) j~i J=l JJc

+ Z Wcj(Y Ycj) + Wbs Wcond Wro j~l (2)

V= compartment volume (m3) l = time (s) pp= compartment air and water vapor densities, respectively (kg/m3)

WJ, Wlj, H~ = mass flow rates associated with j'th ventilation, leakage, and cir-culation paths, respectively (kg/s)

Y= mass fraction of air within com-partment Yj, YIJ air mass fractions in donor com-partments for ventilation path j and leakage path j, respectively Yj mass fraction of air in adjoining compartment associated with cir-culation path j Wb, = rate of steam addition due to pipe breaks (kg/s)

Wd= steam condensation rate (kg/s)

W, = rain-out rate (kg/s).

The values Wj and W~ are positive for flow into the compartmerit and negative for flow out of the com-partment, whereas the circulation rate Wj is always a positive quantity. Ventilation paths are described by their associated mass flow rates and identification numbers of source and receiving compartments.

Ven-tilation flows can be tripped offor on at any time dur-ing a transient by supplying appropriate trip-logic data.

Leakage, circulation, and pipe break models are dis-cussed in Sec. II.C along with other special purpose models.

In formulating the compartment energy balance, it is assumed that air behaves as an ideal gas. Moreover, for the transients of interest, partial pressures of wa-ter vapor are typically(I atm. Therefore, it is assumed that the steam speciflic enthalpy depends only on tem-perature, i.e., the vapor enthalpy is equal to the en-thalpy of saturated steam at the temperature of the gas mixture. The partial pressure of water vapor within a compartment is computed from the ideal gas equation of state, and the total compartment pressure is calcu-lated as the sum of the air and water vapor partial pressures.

With these assumptions, the compartment energy balance becomes NUCLEARTECHNOLOGY VOL. 94 APR. 1991 45

Chaiko and Murphy POSTACCIDENT BKVR SECONDARY CONTAINMENTTHERMALANALYSIS paT paCp,(T) dCpa(T) dhg(T) dT

+p p a pR dT lu Iv a

a

= VTCp,(T)'

Vlt (T) "

dt 't

+ VT R

'"+R dt dt

+ Qh'ghs + Qpanei + Qmoror + Qcooier + Qpiping

+ Qmisc + Qslab + Qbreal' IVbsltg(Pbreak)

IVroJtf(T)

IVcondJ>f(T)

>u

+ Z Vvj[YvjTvJCpa(Tvj) + (

vj)llg(Tvf)]

J=l NI

+ g ~ij[YiiTtjCpa(Tij) + ( I Ytj)ltg(Tji)]

Jmi

+ Z IVcj[ YjTcj Cp (Tcj )

YTCpa (T) j~l where

+ (I Ycj)ling(Tcj)

( I Y)ltg (T)],

(3)

T = compartment gas temperature (K)

Cp,(T) = specific heat of air at temperature T (J/kg K) ltg(T) = specific enthalpy of saturated water vapor at temperature T (J/kg)

R, = ideal gas constant for air (288.7 J/

kg K)

R= ideal gas constant for water (461.4 J/kg K)

Qlighs, Qpanelt Qmosor~ Qcoolers Qpiplngs Qmisc

= compartment heat loads due to light-ing, electrical panels, motors, air coolers, hot piping, and miscellane-ous equipment (J/s)

Q,i,b rate of heat transfer to compartment air/water vapor mixture from sur-rounding slabs (J/s)

II.B. Slab Model In the secondary containment of a BWR, compart-ment walls, ceilings, and floors are generally concrete slabs that range in thickness from -0.3 to -2 m. To determine the heat transfer rate between a compart-ment atmosphere and the bounding concrete slabs, the one-dimensional heat conduction equation

IJT, cl2'T, at 'x'4) is solved for each slab. Here, T, (K) is the slab temper-ature, and x (m) is the spatial coordinate. Since the thermal diffusivitye, (m /s) is supplied as input for each slab, materials other than concrete can be mod-eled provided that slabs are of uniform material com-position. This one-dimensional description assumes that slab edge effects do not significantly affect the overall rate of heat transfer.

Boundary conditions on slab temperature are given by Pb,k total compartment pressure if pipe contains saturated liquid (Pa)

Pb,z pipe fluid pressure if pipe contains saturated steam (Pa) hg(Pb,h) = specific enthalpy of saturated water vaPor at Pressure Pbreah (J/kg) ltf (T) = specific enthalpy of saturated liquid water at temperature T (J/kg)

TJ, Tij donor compartment temperatures for ventilation pathjand leakage path j, respectively (K)

Tj = temperature in adjoining compart-ment associated with circulation path j (K)

Compartment heat loads from lighting, electrical pan-els, motors, and miscellaneous equipment are main-tained constant unless they are tripped on, off, or exponentially decayed during the transient. Hot piping and room cooler loads vary with compartment temper-ature and can also be tripped on or off. In addition, hot piping heat loads can be exponentially decayed using the heat load decay model discussed in Sec.

II.C.7.

46 Qb,h = heat transfer rate to air/water vapor mixture from liquid exiting break as it cools to compartment temperature (J/s) 8'b, = mass flow rate of steam exiting break (kg/s) and

= [T, (t) Ts(0, t)]

cJT, hl Bxo k,

= [Ts(Ls t) T2(t)],

(6)

IJT, hg

BxL, k,

NUCLEAR TECHNOLOGY VOL. 94 APR. I99t

Chaiko and Murphy POSTACCIDENT BIVR SECONDARY CONTAINi4IENTTHERMALANALYSIS where Ti (t), Tz(t) = temperatures of compartments ad-jacent to the slab k, = slab conductivity (J/m s K)

L, = slab thickness (m) hi llz heat transfer coefficients (J/

m-s K).

The solution of Eq. (4) subject to Eqs. (5) and (6) gives the rates of energy transfer from the slab surfaces to the adjacent gas mixtures.

The coefficients hi and h. account for natural convection, radiation, and condensation heat transfer.

In the absence of condensation, the coefficient hi can be expressed as hi = hi+ hi,,

(7) where hiand hi, arc the natural convection and ra-diation components, respectively.

Natural convection coefficients are expressed in terms of the Nusselt number, which in turn is a func-tion of the Rayleigh and Prandtl numbers. For the co-efficient hi, the appropriate relation is free convection from a vertical plate. For horizontal slabs, free-convection coefficients depend on whether the surface is being heated or cooled by the surround-ing gas mixture. As recommended by Holman," the correlation of Fujii and Imura is used with the mod-ified characteristic length proposed by Goldstein et al 6 to compute the coefficient for an arbitrarily shaped slab with heated surface facing upward or cooled sur-face facing downward. In cases where the upper sur-face is cooled or the lower surface is heated, the correlations of Lloyd and (vloran are used.

Diatomic gases such as nitrogen and oxygen are es-sentially transparent to thermal radiation; however, the emissivity of water vapor with respect to therinal radi-ation is significant.a In COTTAP, radiant energy ex-change between a slab surface and water vapor contained within the surrounding gas mixture is modeled through the use of an effective radiation heat transfer coeffi-cient [see Eq. (7)). For the applications of interest, tem-perature differences between a slab surface and the surrounding gas mixture are relatively small (typically

<5 K). Therefore, the following approximate relation proposed by Hottel and Sarofirn for small tempera-ture differences is used to compute thc radiation coef-ficient:

where Nu = "

=f(Ra,Pr),

hiCt.

(8)

(<s+ I) hi, = '4+ a+ b c)e,<<,aT, (10) 2 Ct = slab characteristic length k = gas thermal conductivity and the Rayleigh and Prandtl numbers ture are, respectively, defined by for the gas mix-Pr,

=p

k where a = Stefan-Boltzmann constant (5.669 x 10 '/

mz s K4) e, = slab emissivity T,= average temperature, which is defined by T~= f(T" + T,((,g)/2)',

(I I)

(9) where where g = acceleration due to gravity (9.8 m/sz)

P = coefficient of thermal expansion (K ')

v = kinematic viscosity (mz/s) a = thermal diffusivity(mz/s) t( = dynamic viscosity (kg/m s)

Cp = specific heat of the air/water vapor mixture (J/kg K).

T = gas temperature (K)

Ts,/ slab surface temperature (K) c,= emissivity of water vapor evaluated at T,.

The Cess-Lian'quations, which give an analytical approximation to the emissivity charts of Hottel and Egbert," are used to compute the water vapor emis-sivity. In Eq. (10), c has the value 0.45, and a and b are obtained through differentiation of the Cess-Lian emis-sivity equations Gas mixture properties used in the calculation of free convection coefficients arc evaluated at the thermal boundary layer temperature, which is taken as the av-erage of the slab surface temperature and the bulk gas temperature.

For vertical slabs, coefficients are calculated from the correlation proposed by Churchill and Chu'or and 8 ln (e,( T,P, P., P,L,))

a 8 ln(PL,)

Bin[a,(TP~ P P L( ))

8 In(T)

(12)

(13)

NUCLEAR TECHNOLOGY VOL. 94 APR. I99i 47

Chaiko and Murphy POSTACCIDENT BIVR SECONDARY CONTAINMENTTHERMALANALYSIS where P, = air partial pressure (Pa)'=

water vapor partial pressure (Pa)

L

= average mean beam length (m).

Condensation on a slab surface occurs when the surface temperature drops below the dew point (the saturation temperature of water evaluated at the par-tial pressure of water vapor in the compartment) of the air/water vapor mixture. Heat transfer coefficients for condensation conditions are calculated using the exper-imentally determined Uchida" correlation, which in-cludes the diffusional resistance effect of noncondensible gases on steam condensation rates In COTTAP, initial compartment temperatures, pressures, and relative humidities are specified as in-put data. An initial slab temperature profile is deter-mined by computing the steady solution to Eqs. (4),

(5), and (6) corresponding to the initial compartment conditions. This implies that compartments have been maintained at their initial conditions long enough for slabs to attain steady-state temperature profiles.

II.C. Special Purpose Models The COTTAP code includes specialized models to simulate the effects of pipe breaks, hot piping, and compartment air coolers. Leakage and natural circu-lation models are also included to describe intercom-partment mass transfer. In addition, the code includes a simplified slab model, a heat load decay model, and a compartment model in which temperature,

pressure, and relative humidity are specified as a function of time.

II.C.I. Pipe Break Model Within the scope of the present model, pipes may contain steam or saturated liquid water. Input data de-fine the total mass flow through the break Wt(kg/s) along with the time at which the break develops and the length of time over which fluid loss occurs. For pipes containing saturated liquid, the steam flow rate Wtexiting the pipe (kg/s) is calculated from the en-ergy balance W>IAf(Pp)

Wbsirg(P) + ( Wpl W5$)irf(P) s (I4) which describes the isenthalpic expansion of fiuid from pipe pressure P~ to compartment pressure P. The liq-uid fraction, which does not flash as it leaves the pipe, is assumed to cool to compartment temperature, and the dissipated sensible heat is transferred directly to the compartment air/water vapor mixture. For the case where a pipe contains steam, all of the mass and energy exiting the break is deposited directly into the compart-ment gas mixture.

Rain-out phenomena can be important in compart-ments containing pipe breaks. For example, following 48 where W, = 0,0 ifRH( 0.99, (16)

RH = relative humidity W, = total steam flow rate into the compartment (kg/s)

C,) = constant that is supplied as part of the input data (kg/s).

II.C.2. Hot Piping Model In many secondary containment compartments, the major heat source consists of piping that contains reactor steam or coolant. The heat addition rate to a compartment air/water vapor mixture from a hot pipe is calculated from Q~l g=

t t, t[ f T(t)]

where (17)

U~ = overall heat transfer coefficient (J/m s K)

L~ = pipe length (m)

D~ = outside diameter of the pipe (or insulation if the pipe is insulated) (m)

Tf = pipe fluid temperature (K)

T = compartment temperature.

The overall heat transfer coefficient is calculated by the code based on initial compartment conditions; the co-efficient is then maintained constant throughout the transient.

II.C.3. AirCooler Model Cooling units are used in a number of secondary containment compartments to remove heat generated by equipment such as emergency core cooling systems (ECCS) injection pumps and high-voltage buses and transformers. Heat removal rates of cooling units are calculated from Q,,(t) = C, [T(t) T,oot(t)j (Ig)

NUCLEAR TECHNOLOGY VOL. 94 APR. I99I isolation of a pipe break (due to valve closure, for in-stance) a compartment begins to cool and condensa-tion continues to occur on surrounding walls. For a sufficiently fast cooldown rate, condensation alone does not prevent compartment air from becoming sat-

urated, and thus moisture droplets (rain-out) form within the gas mixture. To maintain compartment rel-ative humidity less than or equal to unity, the rainout rate W(kg/s) is calculated from the following empir-ical model:

Wp = 200 (RH 0.99)max( WC,~ )

ifRH) 0.99 (15) and

C"a+o an" Murphy POSTACCIDENT BWR SECONDARY CONTAINMENTTHERMALANALYSIS where T<<t(t) = average of the inlet and outlet cooling water temperatures C,<<1 = constant that is computed from spec-ified initial values of the cooling load Q

p/ the inlet cooling water tempera-ture, the cooling water flow rate, and thc compartment temperature T.

An energy balance on the cooling water yields the out-let cooling water temperature.

P1, Pz

~ik Ktk Alk Plk pressures of the compartments associated with the leakage path (Pa) leakage rate (kg/s) irreversible pressure loss coefficient leakage area (m2) gas density within the compartment sup-plying the. leakage flow (kg/m3).

It is assumed that inertial effects do not significantly affect leakage rates.

II.C.5. Natural Circulation Model A natural circulation model simulates gravity-driven mixing in compartments connected by flow paths at differing elevations. The circulation rate W, (kg/s) is obtained from II.C.4. Leakage Models The COTTAP leakage model simulates pressure-induced intercompartmental mass transfer through openings such as doorways and ventilation ducts. In-tercompartment leakage is calculated by balancing the pressure differential between the compartments with an irreversible pressure loss. Thus, the leakage rate sat-isfies KtkIVg.(t ) J IVrk(t )

where ll.C.7. Heat-Load Decay Model Cooling of a component such as a pipe filled with hot stagnant fluid or a pump that has ceased operat-ing is simulated through the use of a lumped-param-eter heat transfer model. Most compartments in the secondary containment have a large thermal capacity because of the bounding concrete slabs. It is therefore assumed that the component temperature changes on a faster time scale than the compartment air temper-ature; i.e., the air temperature is assumed to remain fairly constant during the cooldown of the component.

With this assumption, the component heat dissipation rate Q,(t) is governed by dQc(t)

Q ( )

dt (22)

This model also describes intercompartment, gravity-driven circulation flows that can develop at open door-ways (see the analysis of Brown and Solvason'.

II.C.6. Thin Slab Model The detailed slab model discussed in Sec. II.B is not required to describe heat transfer through thin slabs that have little thermal capacitance.

Slabs of this type, e.g., refueling floor walls, have nearly linear tem-perature profiles, and thus the heat flow through a thin slab can be calculated by the use of an overall heat transfer coefficient U>>. The rate of heat transfer through a thin slab is obtained from q>>(t) = U>>A>>[T1(t) T,(t)]

where A>> = thin slab heat transfer area (m")

T1 Tp = temperatures of the compartments sepa-rated by the slab (K).

Values of U>> (J/m s K) are supplied as part of the code input data (one value for each vertical slab and two values for each horizontal slab). For horizontal slabs, two values of U>> are required because free-convection film coefficients depend on the direction, upward or downward, of heat flow through the slab.

2g(pz(t) Pt(t)] (~ <)

t. Kt/IAtpz(t)]+ E<</(Ap,(t)]j where Qc(tO) = Qco (23) where pt, p2 = densities of the air/water vapor mixtures within the two adjacent compartments (kg/m ) (here it is assumed that p2 is the gas density for the cooler compartment)

E, Ft = elevations of the upper and lower flow paths (m)

A,At upper and lower flow path areas (mz).

where McCm Yc UcAc (24)

M, = mass of the component (kg)

C~ = specific heat of the component (J/kg K) and 7, (s '), the thermal time constant of the compo-nent, is given by NUCLEARTECHNOLOGY VOL. 94 APR. 1991 49

Chaiko and Murphy POSTACCIDENT BIVR SECONDARY CONTAINMENTTHERMALANALYSIS U, = overall heat transfer coefficient (J/m2 s K)

A, = component heat transfer area (m~).

In Eq. (23), to (s) is the time at which the cooldown process begins, and Q, which is supplied as input data, is the heat dissipation rate prior to cooldown. So-lution of Eqs. (22) and (23) gives the exponential-decay approximation used in COTTAP to model heat dissi-pation of cooling components.

The component time constant 7, is specified as input data except in the case of hot piping, where it is calculated by the code from the piping description data.

II.C.8. Tt'tne-Dependent Cotnparttnent Model With the time-dependent compartment (TDC) model, environmental conditions within a compart-inent are specified as a function of time; i.e., temper-ature, pressure, and relative humidity versus time are supplied as tabular input data. This model is particu-larly useful for representing outside air conditions, in-cluding solar and thermal radiation effects.

The influence of solar and long-wave atmospheric radiation on exterior buildup surfaces can be described by spec-ifying the effective Sol-Air temperature'4 in the TDC instead of the actual outside air temperature.

In sec-ondary containment analysis, the TDC model is also useful for describing transient conditions within the primary reactor containment, which are generally known from the results of detailed licensing basis cal-culations.

dTsi = oisTsxxi i (2S) where i = 1,2,3,...,N, the number of equally spaced grid points Ti = slab temperature at grid point i T i = finite difference approximation to the second-order spatial derivative at grid pointi.

Following the approach used by Pirkle and Schiesser'n the MOL solution of parabolic equa-50 II.D. Numerical Solution Methods An energy balance and two mass balances are solved for each compartment to determine gas temperature, air mass, and water vapor mass. In addition, the one-dimensional heat conduction equation is solved for each slab. Before computing the numerical solution of the governing equations, partial differential equations describing heat flow through slabs are approximated by sets of ordinary differential equations (ODEs). This is accomplished through application of the method of lines (MOL). In the MOL, a finite difference approx-imation is applied only to the spatial derivative in Eq. (4), giving tions, fourth-order central difference formulas are used to compute T,; at interior grid points:

A six-point sloping difference formula is used to ap-proximate T,; at i = 2 and i = N I:

I Tsxx2 =

2 (10Tsi I 5TsZ 4Ts3 + 14Ts4 6Tss+

Ts6) + O(h )

(27) and 1

TsxxN-i =

2 ( IOTsiv 15 Tsiv-i 4Tstv-3 1262

+ 14Tsiv-3 6Tstv-4 + Tsiv-s)

+O(~4)

(28)

For the end points, where the normal derivatives are specified through convective boundary conditions, the following finite difference approximations, recom-mended by Pirkle and Schiesser,'s are used to com-pute T~i.

I 415 32 T

=T i + 96T2 36T3+ T4 sxxi 12' s

s s

3 s

Tss 50/3.'Tsxi) + O(h )

(29) 3 4

and 1

415 Ts iv =

Tiv + 96Ttv i 36Tiv z 1282 6

s s

32 3

+ Tsh!

3 TsÃ-4 + SOATsxW 2

+O(~4)

(30)

In Eqs. (29) and (30), thc normal derivatives Tsxi and T~> are evaluated in accordance with Eqs. (5) and (6),

the convective boundary conditions; i.e.,

and hi Ts i = (Ti Tsi) ks h2 Tsx3 = (Tsiv Tg) ks (31)

NUCLEAR TECHNOLOGY VOL. 94 APR. I99i I

Tsxxi =

12 (Tsi-2 + 16Tsi-i 30Tsi+ 16Tsi+i

128, Tst+z) + O(~")

(26) where i = 3,4,...,N 2 6 = spacing between grid points.

Chaiko and Murphy POSTACCIDENT BWR SECONDARY CONTAINMENTTHERMALANALYSIS Allgoverning equations are now expressed in terms of ODEs of the form

= F(y, t) with y(0) = yo (32)

II.E. Code Limitations in Modeling Accident Scenarios The followingmodeling limitations have been iden-tified in the current version of the COTTAP code:

1. Fission product transport among compartments is not modeled.

NUCLEAR TECHNOLOGY VOL. 94 APR. I99t Solutions of Eq. (32) exhibit rapid initial adjust-ments in compartment air temperature caused by the relatively small thermal capacitance ofthe air contained within the compartment. Moreover, slab temperatures undergo rapid initial changes in narrow regions near the boundaries, resulting in the formation of spatial thermal boundary layers. In the numerical integration of Eq. (32), small time steps are required to simulate these initial transients. As the initial transient response decays, however, it is desirable to increase step sizes in order to reduce the computation time required to fol-low the slowly varying part of the solution. Equations, such as Eq. (32), which exhibit initial temporal bound-ary layer structures are termed stiffdifferential systems (see the discussion in Ref. 16), and because of stabil-ity limitations, they cannot be solved efficiently with explicit integration schemes.

For this reason,.an im-plicit scheme was selected for COTTAP.

Numerical integration of the governing Eq. (32) is carried out with the LSODES code,'hich uses the implicit backward differentiation methods proposed by Gear for the solution of stiff systems. The LSODES code also employs sparse matrix inversion techniques in solving the implicit finite difference equations. With these numerical integration features, it is feasible to carry out the integration of the large differential sys-tems that arise in the simulation of secondary contain-ment transients.

As an illustration of the problem dimension, simulation of the SSES-I and -2 secondary

'ontainments under postaccident conditions required the solution of 20101 coupled ODEs.

For these large-scale

problems, reevaluation of code-calculated slab heat transfer coefficients at every time step leads to unacceptably long computation times. To alleviate this difficulty,the frequency of re-evaluation (number of steps between reevaluation of coefficients) is a parameter supplied as input to the code. Sensitivity calculations on small-scale problems representative of postaccident secondary containment transients indicate that coefficients can be reevaluated as infrequently as once per ten steps without introducing significant errors in the results. The CPU time require-ments were reduced by a factor of 4 when coefficients were reevaluated at every tenth time step.

2. Cooler modeling does not describe moisture re-moval under conditions where the cooling coil temper-ature is below the dew point of the inlet gas mixture.

3. Pipe break modeling is valid only for lines con-taining steam or saturated liquid; breaks involving the release of subcooled liquid cannot be described.

/

4. Compartment flooding events cannot be simu-lated because all liquid is assumed to exit through com-partment floor drains.

III. RESULTS OF SSES SECONDARY CONTAINMENT ANAIYSIS FOR POSTACCIDENT CONDITIONS This section gives representative results for a COT-TAP simulation of the combined SSES-I and -2 sec-ondary containments under postaccident conditions.

The thermal responses of the Units I and 2 secondary containments are coupled by heat transfer through common walls that separate the two structures. The SSES model consists of 105 compartments, 16 time-dependent compartments, 767 slabs, 38 thin slabs, and 505 heat loads. The simulation was carried out for 30 h and required 124 min of CPU time on an IBM 3090 computer. Note that most of the CPU time is required to simulate the rapidly varying part of the transient that occurs within the first few hours of the event.

Thus, substantially longer simulation times do not sig-nificantly increase CPU time requirements.

For this analysis, it is assumed that a loss-of-coolant accident (LOCA) occurs in SSES-I and a false LOCA signal (a spurious signal that indicates loss of reactor coolant and leads to ventilation system'sola-tion and operation of ECCS injection pumps) is gen-erated on SSES-2.

Under postaccident conditions, ECCS injection pumps comprise the key equipment within the secondary containment structure. The ECCS consists of the residual heat removal (RHR), core spray, and high-pressure coolant injection (HPCI) sys-tems. These systems receive electrical power from high-voltage buses contained within emergency switch gear and load center rooms. Figure I shows the calculated temperature response within a SSES-I RHR pump room (each unit contains two RHR pump rooms and two core spray pump rooms). Initially,the air temper-ature increases rapidly because of the small thermal ca-pacitance of the air within the'compartment.

As air temperature increases, a balance between compartment heat sources and losses to compartment air coolers and slabs begins to develop. At this time, air temperature starts to increase on the slow time scale governed by the slab thermal capacity and transport properties. An initial rapid temperature rise followed by a much slower temperature increase is characteristic of all com-partment heatup transients. After I h of operation, this particular RHR pump switches from the injection mode of operation to the suppression pool cooling 51

Chaiko and Murphy POSTACCI DENT BiVR SECONDARY CONTAINMENTTHERMALANALYSIS 322 320

~ 31S.

I-E o 316.

cL 314 312 0

5 10 15 20 25 30 Time (h)

Fig. 1. Simulation of postaccident temperature response within SSES-I RHR pump room for LOCA on SSES-I and false LOCA on SSES-2.

317 316

~~ 315.

I-314 E

o 313.

312 31 1 z 310 0

5 10 15 20 25 30 Time (h)

Fig. 3. Simulation of postaccident temperature response within SSES-I HPCI pump room for LOCA in SSES-I and false LOCA in SSES-2.

mode. As a result of increased compartment heat loads associated with the change in operating mode, the tem-perature again increases rapidly until a ncw balance between the heat-generation and heat-loss rates is at-tained.

The temperature response within a SSES-I core spray pump room is shown in Fig. 2. Core spray op-eration begins at the start of the event and ceases I h later. Temperature decreases rapidly at this point be-cause, once pump operation is terminated, no signif-icant heat loads remain in the compartment. Figure 3 illustrates the temperature response of the SSES-I HPCI system, which also begins operation at the start of the accident. In this case, however, compartment temperature continues to increase when the system ceases operation at I h into the transient. This occurs because piping heat loads within this compartment are substantial. When HPCI pump operation stops, an as-sociated room cooling unit also ceases operation. Upon shutdown of the cooling unit, slowly decaying piping heat loads rapidly increase compartment temperature until a balance between heat generation and heat losses to compartment slabs is approached.

Figure 4 gives the temperature within a SSES-I load center room that

~

317 L

E c

316 I-E 315

~ 314-313 O

0 5

10 15 20 Time (h) 25 30 Fig. 2. Simulation of postaccident temperature response within SSES-I core spray pump room for LOCA in SSES-I and false LOCA in SSES-2.

309 hC P 308 Eoo

~ 307.

O 0

306 0

5 10 15 20 25 30 Time (h)

Fig. 4. Simulation of postaccident temperature response withinSSES-I load center room for LOCAin SSES-I and false LOCA in SSES-2.

52 NUCLEAR TECHNOLOGY VOL. 94 APR. 1991

Chaiko and Murphy POSTACCIDENT BIVR SECONDARY CONTAINMENTTHERMALANALYSIS supplies electrical power to emergency equipment. In this compartment, heat loads remain essentially con-stant throughout the transient.

From the results of this analysis, it is determined that under postaccident conditions, some of the equip-ment within the secondary containment would be ex-posed to temperatures that exceed their qualification values. Consequently, components were reassessed for operation at higher temperatures, and in some in-stances equipment was relocated to compartments with less severe environmental conditions. Furthermore, a

procedure was developed to instruct plant operators to shed nonessential electrical loads within 24 h after an accident in order to moderate the temperature re-sponses within secondary containment compartments.

IV. EVALUATION OF CODE ACCURACY 315

~ 310 L

~~

305 300 COTTAP CONTAIN 0

2 4

6 8

10 Time (h)

As part of the verification process for the COT-TAP code, calculational results were compared with those obtained with the CONTAIN(Ref. 2) program, which has been verified through comparison with ex-perimental data.'

Although the CONTAIN code does not accommodate a direct heat input (such as from operating mechanical or electrical equipment) to a compartment, useful problems can nevertheless be formulated in order to investigate the modeling and computational accuracy of COTTAP. Two such prob-lems were formulated for code verification. The first problem tests the COTTAP compartment mass and en-ergy balance calculations and the slab heat transfer simulation. This problem consists of a single compart-ment that has a 1000-m3 volume and contains air at 300 K and 101325-Pa initial temperature and pressure.

Concrete slabs, which range in thickness from 0.1 to I m, form the walls of the compartment. Allslabs have a uniform, initial temperature of 300 K. To add heat to the compartment, the air in contact with the outer surface of one slab (the slab that is 0.1 m thick) is sud-denly increased to 400 K at t = 0. In addition, at 50 s into the transient, air with a temperature of 500 K is in-jected into the compartment at a 0.26 kg/s flow rate.

Outer surface temperature rise and air injection con-ditions were selected to effect significant, but not ex-cessive, temperature and pressure response.

Figures 5 and 6 present a comparison of the COT-TAP and CONTAIN calculation results for the first test problem. The temperature and pressure simula-tions both show excellent agreement; note that the pressure response curves given in Fig. 6 completely overlap. In Fig. 5, the initial temperature

increase, which is due to injection of hot air into the compart-ment, begins to level offat -0.5 h. Heat addition by means ofconduction through the externally heated slab then begins to occur, causing a further but less rapid increase in temperature.

The second test problem considered for code ver-Fig. 5. Comparison of COTTAP and CONTAINcompart-ment temperature simulations for test problem I, 0.20 0.18 0.16 I 0.14 0.12 COTTAP CONTAIN 0.10 0

2 4

6 8

10 Time (h)

Fig. 6. Comparison of COTTAP and CONTAIN.compart-ment pressure simulations for test problem I.

ification involves modeling of compartment tempera-ture and pressure behavior under conditions where high-energy steam is injected into the compartment. In this problem, condensation effects strongly influence the rate of temperature and pressure increase. Com-partment physical description data are the same as that for test problem l. In this case, however, the only heat source is the steam entering the compartment at a 0.20 kg/s flow rate and a 2.7756 x 10'/kg enthalpy.

This flow rate and enthalpy are characteristic of a small steam leak within a secondary containment com-partment. Figures 7 and 8 show a comparison of the NUCLEARTECHNOLOGY VOL. 94 APR. t99t 53

Chaiko and Murphy POSTACCIDENT BiVR SECONDARY CONTAINMENTTHERMALANALYSIS 450 ACKNOWLEDGMENTS

" 4oo n,

E 350 300 COTTAP CONTAIN Thc authors thank Jack G. Refling, James E. Agnew, Mark R. Mjaatvedt, and Leonard J. West for their many helpful suggestions during the course of this work. We also thank Lisa Walsh for typing the manuscript.

REFERENCES

1. C. C. LIN, C. ECONOMOS, J. R. LEHNER, G.

MAISE, and K. K. NG, "CONTEMPT4/MOD4: A Multi-compartment Containment System Analysis Program,"

BNL-NUREG-51754, Brookhaven National Laboratory (1984).

0.6 0.5

~

=

g 0.4 N 03-C7 o-0.2 COTTAP CONTAIN r

0.1 0

5 10 15 20 Time (h)

Fig. 8. Comparison of COTTAP and CONTAINcompart-ment pressure simulations for test problem 2.

0 5

10 15 20 Time Ih)

Fig. 7. Comparison of COTTAP and CONTAINcompart-ment temperature simulations for test problem 2.

2. K. K. MURATAet al., "User's lvfanual for CONTAIN I.l:A Computer Code for Severe Nuclear Reactor Accident Containment Analysis," NUREG/CR-5026, Sandia Na-tional Laboratories (1989).

3. S. W. CHURCHILLand H. H. S. CHU, "Correlating Equations for Laminar and Turbulent Free Convection from a Vertical Plate," Int. J. Heat Mass Transfer, 18, 1323 (1975).

4. J.

P. HOLMAN, Heal Transfer, 4th ed., p. 250, McGraw-Hill Book Company, New York (1976).

5. T. FUJII and H. IMURA,"Natural Convection Heat Transfer from a Plate with Arbitrary Inclination," Inr. J.

Heat Mass Transfer, 15, 755 (1972).

6. R. J. GOLDSTEIN, E. M. SPARROW, and D. C.

JONES, "Natural Convection Mass Transfer Adjacent to Horizontal Plates," Ini. J. Hear Mass Transfer, 16, 1025 (1973).

7. J. R. LLOYDand W. R. MORAN, "Natural Convec-tion Adjacent to Horizontal Surface of Various Planforms,"

ASME 74-WA/HT-66,.American Society of Mechanical Engineers (1974).

8. D. Q. KERN, Process Hear Transfer, p. 690, McGraw-Hill Book Company, New York (1950).

9. H. C. HOTTEL and A. F. SAROFIM, Radiative Transfer, McGraw-HillBook Company, New York (1967).

COTTAP and CONTAINsimulation results. The re-sults show good agreement even though the codes em-ploy considerably different approaches in the calculation of condensation rates on slab surfaces. The COTTAP code uses the experimentally determined Uchida'ondensation coefficient, while CONTAIN carries out a detailed computation of the thermal re-sistances associated with the gas boundary layer and the condensate film.

10.

R. D. CESS and M. S. LIAN,"ASimple Parameteriza-tion for the Water Vapor Emissivity," Inr. J. Hear Transfer, 98, 676 (1976).

11. H. C. HOTTEL and R. B. EGBERT, "Radiant Heat Transmission from Water Vapor," Am. Insl. Chem. Eng.,

38, 531 (1942).

12. H. UCHIDA, A. OYAMA,and Y. TOGO, "Evalua-tion of Post-Incident Cooling Systems of Light-Water Power Reactors," Proc. 3rd Inr. Conf. Peaceful Uses of Atomic Energy, Geneva, Switzerland, 1964, Vol. 13, p. 93, United Nations (1965).

54 NUCLEAR TECHNOLOGY VOL. 94 APR. 1991

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Chaiko and Murphy POSTACCIDENT BWR SECONDARY CONTAINMENTTHERMALANALYSIS

13. W. G. BROWN and K. R. SOLVASON, "Natural Con-vection Through Rectangular Openings in Partitions-I Ver-tical Partitions," Int. J. Heat Mass Transfer, 5, 859 (1962).

14. ASHRAE Handbook 1985 Fundamentals, American Society of Heating, Refrigerating and Air-Conditioning En-gineers, Atlanta, Georgia.

15. J. C. PIRKLE, Jr. and W. E. SCHIESSER, "DSS/2: A Transportable FORTRAN 77 Code for Systems of Ordinary and One, Two and Three-Dimensional Partial Differential Equations," presented at 1987 Summer Computer Simula-tion Conference, Montreal, Canada, 1987.

16. C. W. GEAR, Munerical Initial Value Problems in Or-dinary Differential Equations, Chap.

11, Prentice-Hall, En-glewood Cliffs, New Jersey (1971).

17. A. C. HINDMARSH, "ODEPACK, A Systematized Collection of ODE Solvers," Scientific Computing, Vol. I,

p. 55, R. S. STEPLEMAN et al., Eds., IMACS Transac-tions on Scientific Computation, North-Holland Publishing Company, Amsterdam (1983).

18. K. K. MURATAand K. D. BERGERON, "Experimen-tal Validation of the CONTAIN Code," Proc. IltltLIVR Safety Information Mtg., Gaithersburg, Maryland, October 24-28, 1983, SAND-83-1911C, Sandia National Laborato-ries (1983).

19. K. K. MURATAet al., "CONTAIN: Recent Highlights in Code Testing and Validation," Proc. Int. Mtg. Light IVater Reactor Severe Accident Evaluation, Cambridge, Massa-chusetts, August 28-September I, 1983, American Nuclear Society (1983).

Mark A. Chaiko [BS, 1980, and MS, 1983, chemical engineering, Penn-sylvania State University (PSU); PhD, applied mathematics, Lehigh Univer-sity, 1989] is a project engineer-nuclear systems at the Pennsylvania Power &

Light Company. His current technical interests include boiling water reactor stability analysis and thermal-hydraulic modeling of reactor systems.

Michael J. Murphy (BS, mechanical engineering, 1982, and MS, nuclear engineering, 1986, PSU) is a project engineer-nuclear systems with the Penn-sylvania Power & Light Company. He is currently involved in simulation of anticipated transient without scram and severe accident analysis.

NUCLEAR TECHNOLOGY VOL. 94 APR. 1991 55