ML20133Q106

Review of Trac Calculations for Calvert Cliffs PTS Study

ML20133Q106
Person / Time
Site:	Calvert Cliffs
Issue date:	04/30/1985
From:	Jo J, Rohatgi U BROOKHAVEN NATIONAL LABORATORY
To:	NRC OFFICE OF NUCLEAR REGULATORY RESEARCH (RES)
References
CON-FIN-A-3266, REF-GTECI-A-49, REF-GTECI-RV, TASK-A-49, TASK-OR BNL-NUREG-51887, NUREG-CR-4253, NUDOCS 8511010391
Download: ML20133Q106 (114)

Text

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NUREG/CR-4253 BNL-NUREG-51887 10c55507ES77 i 1^N1R4 US NRC ADM-DIV 3: TI3C POLICY

e. DJB 4GT BR PDR NJREG h5 IN3 TON DC 00555 REVIEW 0F TRAC CALCULATIONS FOR CALVERT CLIFFS PTS STUDY J.H. Jo and U.S. Rohalgi Date Published - April 1985 i

LWR CO[E ASSESSMENT AND APPLICATION GROUP DEPARTMENT OF NU', LEAR ENERGY, BROOKHAVEN NATIONAL LABORATORY UP(ON, LONG ISLAND, NEW YORK 11973 I*

l Prepared for the U S Nuclear Regulatory Commission l

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Office of Nuclear Regulatory Research Contract No. DE-ACO2-76CH00016

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NUREG/CR-4253 BNL-NUREG-51887 AN. R-4 REVIEW 0F TRAC CALCULATIONS FOR CALVERT CLIFFS PTS STUDY J.H. Jo and U.S. Rohalgi Manuscript Completed - March 1985 Date Published - April 1985 LWR CODE ASSESSMENT AND APPLICATION GROUP DEPARTMENT 0F NUCLEAR ENERGY BROOKHAVEN NATIONAL LABORATORY UPTON, LONG ISLAND, NEW YORK 11973 PREPARED FOR UNITED STATES NUCLEAR REGULATORY COMMISSION OFFICE OF NUCLEAR REGULATORY RESEARCH WASHINGTON DC 20555 NRC FIN A-3266

ABSTRACT Six selected transient calculations out of thirteen perf ormed by LANL using the TRAC-PF1 code f or the llSNRC PTS study of the Calvert Cliff s Nuclear Power Plant have been reviewed in depth at BNL.

Simple hand calculations based on the mass and enerqv balances have been perf ormed to predict the ten-perature and pressure of the reactor system, and the results have been compar-ed with those of TRAC. Comparison was also made between the TRAC and RETRAN calculations for two of these transients, which were perf orned by ENSA.

In general, the esults calculated by TRAC appear to be reasonable based on the comparison with RETRAN and hand calculations.

l

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EXECUTIVE SUt*ARY Brookhaven National Laboratory was requested by the USNRC to review the thermalhydraulic calculations perf arned by Los Alamos National Labo rat o ry (LANL) and Idaho National Engineering Laboratory (INEL) as a part of an NRC progran to study the Pressurized Thermal Shock saf ety issue.

The thernal-hydraulic calculations f or this study were perf ormed at LANL and INEL using the latest versions of the TRAC-PWR and RELAP5 codes, respectively.

This report presents the results of the BNL review of the selected transients calculated by LANL usinq the TRAC PF1 code for the Calvert Clif f s Nuclear Power Plant.

LANL perf ormed TRAC calculations for thirteen transients.

TRAC input decks and steady-state results for these calculations were reviewed at BNL and a quick p rel imi na ry review of all thirteen calculations was also perf ormed.

No najor discrepancies were found in the input deck or the steady-state calculations.

BNL selected six transient calculations out of thirteen perf orned by LANL for further in-depth review.

In order to provide a quantitative review of these calculations, a simple method was developed at BNL to predict the p ri ma ry systen temperature on the basis of the overall ener gy balance.

In this approach, the whole reactor systen, including the secondary sides of the steam generators and the netal structures, is lumped into a single volume and represented by a single average temperature.

It was shown that the primary temperatures calculated by TRAC were indeed in close agreenent with those obtained by simple hand calculations for most transients.

Since the prinary and secondary pressures were more dif f icult to predict by simple analysis due to signif icant nonequilibrium ef f ects involving condensation and evaporation, two liniting pressures based on adiabatic and equilibrium assumptions were calculated and compared with the TRAC-calculated pressure.

The cctual pressure is somewhere hetween these two extrenes.

It was found that the TRAC pressure was usually closer to the adiabatic than to the equilibrium pressure.

The sane approach was used to extrapolate the calculations and to predict the ultinate state of the systen.

In ceneral, the +emperatures and pressures of the p rima ry systen calculated by TRAC have been found to be quite reasonable.

The seconda ry pressures calculated by TRAC indicate that the TRAC code may have some dif ficulty with the condensation model and f urther work is needed to assess the code calculation of the U-tube steam generator pressure when the cold auxiliary feedwater is introduced into the steam generator.

However, this uncertainty is expected to have no s ignif icant ef f ect on the transient calculations.

-v-l l

ACKNOWLEDGMENT The cuthors wish to express their appreciation to Dr. P. Saha of BNL f or his valuablo comments and suqqestians.

They also wish to acknowledge its. J.

E. Koenig of ' ANL for oroviding inf ormation about the LANL calculations.

Ack-nowledgment is oko given to Ms. Ann C.

Fort and Laura Zaharatos for typing this report.

3

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TABLE OF CONTENTS Page No.

ABSTRACT.

jjj EXECUTIVE

SUMMARY

y l

ACKNOWLEDGMFNT..

yj ABRREVIATIONS.......

x 1.

Introduction.

1 2

2.

Transient 1: 1-f t Steam line Break in HZP Condition.

6 3.

Transient 11: Full Double-Ended Guillotine Steam line Break with MSIV Failure to Close Durina HZP Operation.

15 2

4.

Transient 3: HFP l-f t Stean line Break in HFP Condition.

20 5.

Transient 6: AFW Overfeed From HFP.

26 6.

Transient 7A: Small-Break LOCA in a Hot Leq.

32 7.

Transient 9: t'FW Overf eed to One Stean Generator.

38 8.

Sunnary and Conclusions.....

42 REFERENCES....

43 APPENDIX A - BNL Review of PTS Input and Steady-State Calculations Perf ormed by LANL f or Calvert Clif f s Nuclear P1 ant.

47 APPENDIX B - LANL Response to the BNL Review of the Calvert Clif f s.

Input Decks and Steady-State Calculations.......

49 APPENDIX C - A Preliminary Review of TRAC Calculations for Calvert Clif f s PTS Stu1y.

53 APPENDIX D - Extrapolation of Existing PTS Calculations With or Without Changes in Bourdary Conditions.........

93

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LIST OF FIGURES Ficure No.

Page No.

2.1 Transient 1: Liquid Temperature.........

8 2.2 Transient 1: Liquid Temperature........

8 2.3 Transient 1: System Pressure.......

9 2.4 Transient 1: System Pressure and Pressurizer Level.....

9 2.5 Transient 1: Steam Generator Pressure.....

10 2.6 Transient 1: Liquid Temperature in the Downconer.

10 2.7 Transient 1: Pressure in the Vessel Downcomer..

11 2.8 Transient 1: Pressurizer Water Level.......

11 2.9 Transient 1: Pressure in the Steam Generators..

12 3.1 TrcnSient 11: Liquid Temperature.........

17 3.2 Trarsient 11: Pressurizer Pressure........

17 3.3 Tra isient 11: Steam line Pressure..

18 3.4 Tra isient 11: Temperature in the Vessel Downtoner.

18 3.5 Tra isient 11: Pressurizer Pressure....

19 3.6 Tra isient 11: Pressure in the Steam Generators..

19 4.1 Transient 3: Total Integrated Steam Flow f rom Steam Generators..........

24 4.2 Transient 3: Downconer Liquid Temperature...

24 4.3 Transient 3: Downconer Pressure..

25 4.4 Transient 3: Pressure in the Steam Generators..

25

(

5.1 Transient 6: Liquid Temperature..

29 5.2 Transient 6: Total Steam Flow From Steam Generator.

29 5.3 Transient 6: Primary Pressure...

30 5.4 Transient 6: Seconda ry Pressure:.........

30 5.5 Transient 6: Comparison of the Liquid Tenperature Calculated by TRAC Code and BNL Method Using Two Dif f erent Auxiliary Feedwater Fl ow R a t es.............

31 6.1 Transient 7A: Liquid Tenperature..

35 6.2 Transient 7A: P rima ry Pressure.

35 6.3 Transient 7A: Total Steam Discharge....

36 6.4 Transient 7A: Secondary Pressures..

36 6.5 Transient 7A: Vapor Space Volume.....

37 7.1 Transient 9: Liquid Temperature..

40 7.2 Transient 9: SG Liquid Temperatures.......

40 7.3 Transient 9:

System Pressure..

41 7.4 Transient 9: Stean Generator Secondary Side Pressure.

41

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J LIST OF TABLES Table No.

Page No 1.1 Cal vert Clif f s PTS Transients...........

4 2.1 St;nario Description - Transient No. 1.......

13 2.2 Comparison Between TRAC and Design / Plant at Hot Zero Power Conditions...............

14 3.1 Scenario Description - Transient No.11......

16 4.1 Scenario Description - Transient No. 3.......

21 4.2 Comparison Between TRAC and Design / Plant at Full Power Conditions..................

22 5.1 Scenario Description - Transient No. 6.

28 6.1 Scenario nescription - Transient No. 7A 34 7.1 Scenario Description - Transient No. 9.......

39 l

1 i

I

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ABBREVIATIONS ADV Atmospheric Dump Valve AFAS Auxiliary Feedwater Actuation Signal AFW Auxiliary Feedwater BG&E Baltimore Gas and Electric Company BNL Brookhaven National Laboratory HFP Hot Full Power HPI High Pressure Injection HZP Hot Zero Power LANL Los Alamos National Laboratory LOCA Loss of Coolant Accident MFRV Main Feedwater Regulating Valve MFW Main Feedwater MSIV Main Steam Line Isolation Valve ORNL Oak Ridge National Laboratory PORV Power Operated Relief Valve PTS Pressurized Thermal Shock SG Steam Generator SGIS Steam Generator Isolation Signal SIAS Safety Injection Actuation Signal TBV Turbine Bypass Valve TSV Turbine Stop Valve USNRC U.S. Nuclear Regulatory Commission

-X

1.

INTRODUCTION Rapid cooling of the reactor pressure vessel during a transient or acci-dent accompanied by high coolant pressure is ref erred to as pressurized ther-nal shock (PTS).

In late 1981 the U. S. Nuclear Regulatory Commission (NRC) designated PTS as an unresolved saf ety issue and developed a task action plan (TAP A-49) to resolve the issue.

The saf ety issue exists because rapid cooling at the reactor vessel wall inner surf ace produces thermal stresses within the wall. As long as the f rac-ture toughness of the reactor vessel is high, overcooling transients will not cause vessel failure. However, NRC staf f analyses (SECY-82-465) showed that certain older plants with copper impurities in vessel weldments may become sensitive to PTS in a f ew years as the nil-ductility transition tenperature of the weld material gradually increases. The purpose of the thermal-hydraulic analyses presented in this report is to better understand the behavior of a plant during various kinds of postulated severe overcooling transients with l

multiple f ailures of equipment and without operator corrective action.

The understan jing gained f rom these detailed calculations will be used to interpo-late coalint temperature and pressure responses in the downconer f or other postulated transients using a simplified mass-and-energy balance approach.

For each of these postulated transients, Oak Ridge National Laboratory (ORNL) will then calculate the reactor vessel temperature distribution and stresses during the transient and the conditional probability of vessel f ailure if the transient should occur. ORNL will publish a report that integrates these re-suits to estimate the likelihood of PTS driving a crack through the reactor vessel wall and to identify important event sequences, operator and control actions, and uncertainties.

This series of analyses is intended to provide inf ormation to help the NRC staff confirm the bases f or the screening criteria in the proposed PTS rule (proposed *.0CFR 50.61) and determine the content required f or licensees' plant-specific saf ety analysis reports and the acceptance criteria f or correc-tive neasures.

The Nuclear Regulatory Commission (NRC) has selected three plants repre-senting FWRs supplied by three vendors in the United States f or detailed PTS study. These are: Oconee-1 (Babcock and Wilcox), Calvert Clif f s (Conbustion Engineering), and H. B. Ro)inson (Westinghouse Electric). Oak Ridge National Laboratory, which is the lead contractor f or the entire PTS study has identi-fied several groups of transients with nultiple equipment f ailure and with no corrective operator action which could lead to severe overcooling in these plants. The thermal-hydraulic calculations f or these transients were per-f orned at the Los Alamos National Laboratory (LANL) and the Idaho National Engineering Laboratory (INEL) using the latest versions of the TRAC-PWR and RELAPS codes, respectively.

The Oconee-1 transients were divided between LANL and INEL, with some transients common to both. The Calvert Clif f s and Robin-son transients were assigned to LANL and INEL, respectively.

-1_

Brookhaven National Laboratory (BNL) was requested by the NRC to review and compare the plant input decks developed at LANL and INEL, and to review the calculation results. This report presents the results of the BNL review of the selected Calvert Cliff s calculations perf ormed at LANL.

LANL perf ormed TRAC calculations of thirteen transients, as shown in Table 1.1 [1].*

Four of these, Transient Nos.1, 2,10, and 11, were steam line break accidents initiated f rom the hot zero power (HZP) condition. The remainder were various transients initiated f rom the hot f ull power (HFP) con-dition. TRAC input decks and steady-state results f or these calculations were reviewed at BNL, and comments were transmitted to NRC, LANL, and ORNL study nembers in January 1984. Also, a quick preliminary review of all thirteen calculations was perf ormed at BNL and a letter report was sent to the study members in February 1984 For the sake of completeness, copies of these com-munications are presented in Appendices A and C, respectively. Also included in Appendix B is the LANL response to the BNL review of the input decks and the steady state calculations. Six of the thirteen transients have been se-lected f or detailed review at BNL: Transients 1 and 11 f or the HZP condition, and Transients 3, 6, 7A, and 9 f or the HFP conditions. The reasons f or their selection are as f ollows:

1.

All f our HZP transients are steam line break accidents. Two are 2

1-f t steam line breaks and the other two are full double-ended guillotine steam line breaks. Transients 1 and 11, representing one each f rom the two dif f erent break sizes, have been selected f or this detailed review. The same transients were also calculated using the RETRAN code by ENSA f or BGSE. Thus, a comparison between the TRAC and RETRAN calculations can also be made.

2.

Basically, three dif f erent categories of transients were initiated from the HFP condition: (a) steam line break or valve f ailure- (Transients 3, 4, and 4A), (b) small-break LOCA (Transients 7 and 7A), and (c) steam genera-tar overf eed (Transients 6, 8, and 9).

Transient 5 is a combination of pri-mary and secondary f ailures (PORV and ADV stuck-open).

Transient 3 has been selected to represent the stean line break / valve 2

failure transients.

The break size of this transient (1 f t ) is the same as that of Transient 1 of the HZP case, and this allows comparison of the same transients initiated at twn extreme power levels. Transient 7A has been se-lected as representative of the small break LOCA transients, since Transient 7 involved artificial and thus unrealistic flow blockage of the pri-mary loop. Transients 6 and 9 have been selected f or the steam generator overf eed cases, representing the AFW and MFW overf eed, respectively.

In order to provide a quantitative review of the above TRAC calculations, a simple method has been developed to predict the primary system temperature based on the mass and energy balances.

In this approach, the whole reactor system, including the secondary sides of the stean generators and the netal structures (unless otherwise mentioned), is lunped into a single volume and the energy balance is applied to that volume. However, separate mass balance

• It should be noted that the transient calculations described in Tabit 1.1 were purely hypothetical and not necessarily probable. The transients were chosen to give as much insight as possible in a minimun set of calculations to the ef f ect of certain operator and equipment f ailures, even when the proba-bility of the combination of these f ailures was extremly low.

.- 2 --

equations are used f or the primary system and the secondary side of each SG.

Details of this method can be f ound in Section 2 of Appendix D (Simple Ap-i l

proach). This approach assumes that the temperature dif f erences between the cold and hot leqs of the primary loops and between the primary and secondary sides of SGS are relatively small. It will be shown in this report that the primary temperatures calculated by TRAC were indeed in close agreement with those obtained by simple hand calculations for most transients.

The primary and secondary pressures have been more dif f icult to analyze with this sinple approach, especially when the cold water is entering into the pressurizer or the secondary sides of the SGs. Due to the signif icant non-J equilibrium ef f ect, the pressure prediction depends largely on the condensa-tion or evaporation rate, which is dif ficult to estimate by simple analysis.

Many f actors af f ect the condensation and evaporation rates, such as tempera-ture of the liquid and vapor, mass flow rate, mixing of the incoming water with the bulk water, and the mode of heat transf er between the liquid, vapor and wall. Theref ore, in sone transient calculations, attenpts have been nade h

to conoare the oressurizer water levels obtained by the TRAC and BNL single q

calculations instead of the pressures.

It has been observed that the trend of the pressurizer pressure calculated by TRAC closely approximates the trend of the water level in the pressurizer in many transients. Whenever possible and applicable, calculation f or the pressurizer pressure has been based on the adiabatic and/or equilibrium assumptions (Section 4 of Appendix D, Pressurizer Model). The adiabatic approach assunes no mass and energy transf er between the liquid and vapor phases (no condensation or evaporation). The prossure thus calculated is expected to be the lower bound of the actual pressure when the pressurizer is beinq enutied and the upper bound when the pressurizer is being f illed. Dn the other hand, the equilibrium approach assumes that the phases are in complete equilibrium, and it is expected to provide the upper bound pressure when emptying and the lower bound when f illing. The actual pressure is expected to be somewhere between these two extreme pressures, it has been f ound that the pressure calculated by TRAC is usually closer to the adiabatic than to the equilibrium pressure.

A similar nonequilibrium ef f ect has also been observed in the secondary side pressure of SG calculated by TPf/, especially when the SG is being f illed with the cold auxiliary f eedwater (AF a).

In several transients, the secundary pressure remains high while the tenperature declines. This indicates high nonequilibriun ef f ect.

It appears that f urther code assessnent work is needed to verif y the code calculation of the U-tube steam generator pressure when the cold auxiliary feedwater is introduced into it.

However, this uncertainty is not expected to af f ect the transient calculations signif icantly.

A similar approach may be used f or the extrapolation of the calculations if necessary.

In fact, attempts have been made, whenever possible, to predict the ultinate state of the systen beyond the calculated tine.

Abbreviations such as TRAC temperature or BNL pressure are f requently I

used in the followinq discussion f or convenience. They mean the temperature l

calculated by the TRAC code or the pressure obtained by the simple hand calcu-lations by RNL staf f, respectively. /

s t

~-

Table 1.1 Calvert Cliffs PTS Tr an s i en t s initial Equipment Transient Plant initiating Fallures on Operator No.

Descriptive Title State Event Demand Actions 2

2 1.

1-tt steam line break at Hot 01 1.0-ft hole in None None j

standby Power steam 1Ine A 2.

Full double-ended Hot 05 Full steam line Auxillary feedwater None guillotine steam line Power break (AFw) is

r. o t isolated l

broak 2

2 3.

1-ft steam line break at 1005 1.0-ft hole in None None tu11 power Power steam line i

4 Turbine-trip with turbine-1005 Turbine trip T8v sticks wide None o

bypass valve (TBV) stuck Po-er open I

open 4a.

Turbine trip with one TBV 1005 Turbine trip TBv & MSiv stuck open and one MStV stuck open Power 5.

Primary power-operated &

1005 Popv transfers I ADi opens on demand N

e atmospheric-dump valve Power to elde open and sticks open tADV) stuck open 6

AFw overfeed after AFa 1005 MFw system trips AFw delay for AFn valves response failure 20 min.

opened fully at 20 min.

2 7.

Small break toss-of-coolant 1005 An 0.02-ft no,,

y0,,

y,n, accident with blocked Power appears in the hot natural circulation leg

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2 2.

TRANSIENT 1: 1-FT STEAM LINE BREAK IN HZP CONDITION 2

This transient was initiated by a 1-f t break at the main steam line dur-ing the HZP operation. No other equipment f ailure or operator action was as-suned. The transient scenario as specif ied by ORNL is shown in Table 2.1.

The initial steady state f or the HZP calculated by TRAC is shown in Table 2.2 Figure 2.1 shows the downcomer liquid temperature calculated by TRAC with the system average temperature obtained by BNL hand calculation, as discussed in the previous section. Two BNL-calculated temperatures are shown in the figure. One is calculated with the assumption that heat transf er between the j

wall of the reactor (and other structures) and liquid is instantaneous and, thus, the metal temperature changes with the liquid temperature. The other calculation assumes that the heat transf er is so slow that the metal tempera-ture does not chance. The real temperature should be between these two ex-trenes. The TRAC downconer temperature initially agrees well with the tempera-ture calculated without the metal mass accounted f or, and then it eventually approaches that calculated with the metal mass accounted f or, as expected, l

This indicates that the metal takes a considerably longer time to cool. The liquid temperatures calculated by TRAC at the various locations are shown in Fiaure 2.2, along with the BNL system average tenperatures, with and without the netal heat transf er during the initial 1500 seconds. The figure shows that the downconer temperature may be representative of the systen average tempera-ture and, again, both TRAC and BNL calculations agree very well.

5 Fioure 2.3 shows the systen pressures as calculated by TRAC and the RNL staff. The BNL pressure is calculated on the assumption of adiabatic compres-sion during the filling stage. As discussed in the preceding section, the adiabatic assumption yields the highest rate of pressure increase during com-pression. The actual pressure is expected to be lower than this, as is the Case in this caiculation. The figure also snows, f or conparison, the water level in the pressurizer as calculated by BNL. As expected, the pressure and the RNL water level behave similarly. However, the TRAC pressurizer level de-creases while the TRAC pressure increases between 170 and 550 seconds, as shown in Figure 2.4 This annears to be contradictory.

i Figure 2.5 shows the TRAC pressure of the secondary sides of both steam generators. The saturation pressures corresponding to the BNL average temper-ature and the TRAC intact steam generator temperature are also shown in the f igu re.

These would be the expected pressures of the stean generators if the equilibrium condition prevails. The broken steam generator pressure stays at the atmospheric pressure as it hetones empty, as expected. However, the in-tact stean generator pressure remains much higher than the saturation pressure and also shows several sharp turns. A similar steam generator pressure re-spnnse is observed in several other transients when the steam generator is being f illed with cold AFW. This is apparently related to the severe non-equilibrium ef f ect caused by the TRAC condensation nodel.

It appears that the TRAC condensation model underpredicts the condensation rate and, thus, overes-tinates the n^n-equilibrium ef f ect. However, this is not expected to alter the course of the rest of the transient significantly, since the SG pressure is not involved in the control of the system at ter the initial 100 seconds in this transient.

l The TRAC calculation was terminated at 7200 seconds. After 7200 seconds, the system temperature is expect.ed to continue to decrease until it eventually reaches 3S7 K, where the decay heat balances with the cooling by the charging and the AFW.

There is a corresponding RETRAN calculation perf ormed by ENSA f or BGAE available f or this transient f or the initial 1000 seconds. Figure 2.6 shows good agreement between the downcomer temperature calculated by RETRAN and those obtained by TRAC and BNL calculations. Figure 2.7 shows that the RETRAN pressure is virtually identical to the TRAC pressure, while the BNL pressure based on the adiabatic assumption is higher than these, as expected. Fiqure 2.8 corpares the pressurizer water level f or all three calculations. The re-sults f rom the BNL and RETRAN calculations agree closely, while the TRAC level shows a somewhat contradictory trend between 170 and 550 seconds, as mentioned earlier. Figure 2.9 shows the pressure in the steam generators f rom both RETRAN and TRAC calculations. The saturation pressure corresponding to the system average temperature calculated by BNL is also shown in the figure. The BNL saturation pressure matches the broken SG pressures for both TRAC and RETRAN calculatinns very closely. However, the intact SG pressure f or TRAC increases while the RETRAN pressure continues to decrease. As discussed ear-lier, f urther work is needed to clarify this uncertainty.

In sunnary, both TRAC and RETRAN codes present reasonable results except f or the TRAC intact SG pressure, which may have an insignificant ef f ect on the final results.

, 1 1

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CAUTION: The scenario simulated contains significant conservatisms ia. operator actions and equipment failures. For details, see the scenario descriptions.

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560 DOWNCOMER (THAC) 550 SYSTEM AVERAGE 520 TEMPERATURE LTH E

g-525 METAL L ATENT HEAT-4 80 w L

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Figure 2.1 Transient 1: Liquid Temperature CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipnent failures. For details, see the scenario descriptions 575 T-T

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200 400 600 800 1000 1200 1400 1600 TIM E (s)

Figure 2.2 Transient 1: Liquid Temperature _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

CAUTION: The scenario simulated contains significant conservatisms in Operator actions and equipment failures For details, see the scenario descriptions.

PSI M Pa IO I

l l

1 I

I I

2500 -

p g

- 10 16 2000 - 14

~

/

,/

8 a

12 F

-[

E

/

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/

6Q#

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SYSTEM PRESSURE (TRAC)

I I

--- SYSTEM PRESSURE (BNL; - 4N

-f AD BATIL MODEL)

T 8

-- PRESSURIZER WATER LEVEL h

1000-

/

(BNL) pg

+, /

6 J'

ct 1

I I

I I

I I

I

-4 O

O 1000 2000 3000 4000 5000 6000 7000 0000 TIME (s)

Figure 2.3 Transient 1: System Pressure CAUTION: Tre scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions.

Psi MPa

]I8 i

l I

i i

i I

- g SYSTEM PRESSURE (TRAC) 16-

. - PRESSURIZER LEVEL (TRAC)

--- PRESSURIZER WATER LEVEL (BNL)-

7 pe 2000-14

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W g

--12 9 5 W" w

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_ 4w m 1500

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- 38 8*

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i j

i i

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200 400 600 800 1000 1200 1400 1600 TIME (s)

Figure 2.4 Transient 1: System Pressure and Pressurizer Level

CAUTION: The scenar10 simulated contains significant conservatisms in operator actions and equipment failures. for details, See the scenario descriptions.

1000 7

T T~ ~T T- ~~ T 1~

1 INTACT LOOP (TRAC)

--- BROKEN LOOP (TRAC) 6 SATURATION PRESSURE 900 CORRESPONDING TO THE INTACT SG LIQUID TEMP (TRAC) - 800

^ SATURATION PRESSURE ES CORRESPONDING TO THE S

SYSTEM AVERAGE TEMP.(BNL) - 700_"

.5 t

600S g4

'/

to a

500 $

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300 I.

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200

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0 1- =1- - 4 r-- =i-0 2000 4000 6000 8000 TIME (s)

Fiqure 2.5 Transient 1: Steam Generator Pressure CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For detdiis, see the scenario descriptions.

600 p- -~ T --

T I

T 550

- RETRAN BROKEN LOOP _

-- RETRAN INTACT LOOP i~'

\\

(3 500 o TRAC

-w'\\

I

^ DNL (WITH METAL)

Sl 450 (1 BNL (WITHOUT METAL) _ ay400 .q\\x 'N, [j350 N,\\ t-it' 'A o 3 R W.........: e 250 L---- I-1 1 - -- T 0 200 400 600 800 1000 1200 TIME (s) Fiqure 2.6 Transient 1: Liquid Temperature in the Downcomer _ _ _ _ _ _ _

CAUTIUN: Ine scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 2500 i i i i - RETRAN 2000_. TRAC a BNL a ct ^ a E 1500-O en e ^ Q-1000 500O 200 400 600 800 1000 1200 TIME (s) Figure 2.7 Transient 1: Pressure in the Vessel Downconer CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 8 1 I I I I I TRAC y

---

R E TR AN

] --- BNL y6 "i ct U ?, A ./* ~ / 5 .// N / E f,' S // $2 5 b -l Y / Y 0 \\m~ ct/ i I I I I 1 0 200 400 600 800 1000 1200 1400 1600 TIME (s) Figure 2.8 Transient 1: Pressurizer Water Level CAUTION: The scenario simulated contains significant conservatisms in operator { actions and equipment failures. For details, see the scenario descriptions. L I I I I I - BROKEN LOOP -- INTACT LOOP (RETRAN) -- BROKEN LOOP - INTACT LOOP (TRAC) o SATURATION PRESSURE CORRESPONDING 1000 TO SYSTEM AVERAGE TEMPERATURE (BNL) =j800 g 600 }, 3,, a 's N E .\\, ' x ~ ~ ~. a a. 400 \\ 200 \\ N ^ O I I 7 ~ ~ " ~ --- f 0 200 400 600 800 1000 1200 TIME (s) Figure 2.9 Transient 1: Pressure in the Steam Generators l TABLE 2.1 SCENARIO DESCRIPTION TRANSIENT NO. 1 Plant Initial State - Just prinr to transient initiator General

Description:

Hot standby, 0% Power af ter 100 hr of shutdown Systen Status Turbine: Not latched Turbine Bypass Valves (TBV): Automatic control Atmospheric Dunp Valves (ADVs): Automatic control Charging Systen: Automatic control Pressurizer: Automatic control Engineering Saf ety Features: Automatic control Power Operated Relief Valves (PORVs): Automatic control Reactor Control: Manual Main Feedwater: In bypass mode, manual control to provide zero level in SGs; I condensate pump, 1 booster pump, I t1F WP operating on steam supplied by unit 2. Aux Feedwater: Automatic control Main Strean Isolation Valves (MSIVs): Open, autonatic control Main Feedwater Isolation Valves (MFIVs): Open, automatic control 2 Transient Initiator - A 1.0-f t hole appears in stean line A outside containment upstrean of the MSIV and downstream of the flow restrictor. Equipment Failures which occur during the accident transient if the equipnent is denanced. None Operator actions /inacticns a. Operator will turn off all RCPs 30 seconds af ter SI AS based on low pressurizer pressure, b. Operator f ails to turn of f Charging pumps prior to f ull repressurization. c. Operator f ails to control repressurization, d. Operator f ails to maintain level in intact SG. e. Operator f ails to respond to high SG alarm at 30". j f. Operator f ails to respond to high SG alarm at 50"..

TABLE 2.2 Comparison Between TRAC and Design / Plant at Hot Zero Power Conditions Design / Plant TRAC Data Predictions PRIMARY SIDE Power 100 hr after shutdown 9.38 MW Decay Heat 17.38 MW Pump Power Pressure 15.52 MPa 15.52 MPa (2250 psia) (2250 psia) Mass Flow 19300 kg/s 19700 kg/s (42,549lb/s) (43,431 lb/s) Average Temperature 550.9 K 551.8'K (532 F) (533.6 F) Pressurizer Level 3.68 m 3.68 m (144.0 in) (144.0in) SECONDARY SIDE Feedwater flow per SG 10.1 kg/s 11.8 kg/s (22.3 lb/s) (26.0lb/s) SG Dome Pressure SG 11 (TRAC 6.20 MPa 6.17 MPa component 22) (900 psia) (895.5 psia) SG 12 6.20 MPa 6.17 MPa (900 psia) (895.5 psia) Feedwater 299.8*F 299.8 F Temperature (80.0 F) (80.0'F) TBV % Open 5.0 SG Liquid Mass 102,058 kg 102,058 kg (225,000lb) (225,000 lb) 3. TRANSIENT 11: FULL DOUBLE-ENDED GUILLOTINE STEAM LINE BREAK WITH MSIV FAILURE TO CLOSE DURING HZP OPERATION This transient is initiated by a full double-ended guillotine break in a steam line during the HZP operation. In addition, it is assumed that both MSIVs fail to close on SGIS and that the operator turns off the AFW system at eight minutes after the beginning of the transient. Transient scenario as spe-cified by ORNL is shown in Table 3.1. The major differences between this transient and Transient 1 are the ad-ditional failure of the MSIVs to close and the operator action to turn off the AFW. The break size is not considered to be a major difference, since the lo-cation of the break is downstream of the flow restrictor which is located at the exit of the SG. Figure 3.1 shows the TRAC downcomer temperature and the BNL system aver-i age temperature. They match very closely. The reactor rapidly cools down and 1 depressurizes due to blowdown of both SGs. The system temperature converges l to the saturation temperature corresponding to the atmospheric pressure (373'K or 100 C) as the SGs dry out and depressurize to the atmospheric pressure. 2 The blowdown continues and the system remains at this temperature until both 3 SGs are finally empty. We estimate this time to be approximately 7800 seconds based on the energy and mass balance. Once SGs dry out, the temperature will slowly rise, since charging flow is not sufficient to balance the decay heat. It will eventually reach the steady-state temperature where the decay heat balances with cooling due to charging. We estimate this temperature to be ap-proximately 545*K. A substantial amount of water is entrained through the break in the be-ginning of the transient. This entrainment slows the rate of the temperature and pressure decreases. This may mean that the initial cooling could be some-what faster or slower depending on the adequacy of the TRAC entrainment mod-l el. The entrainment will also affect the timing of the dryo'ut of the steam generators. However, the effect of entrainment on the final outcome of this transient is not expected to be significant. i Figure 3.2 shows the TRAC primary pressure and the pressurizer water lev-el from the TRAC and BNL calculations. They all show a consistent trend. As expected, the secondary side pressure of both SGs decreases approximately a-long the saturation pressure corresponding to the BNL system average tempera-ture, as shown on Figure 3.3. A corresponding RETRAN calculation performed by ENSA and BG&E is also available for the first 600 seconds of this transient. Figure 3.4 compares the downcomer temperature of the RETRAN, TRAC and BNL calculations. They all agree very well. Figure 3.5 shows significant deviation between the pressuri-zer pressures calculated by the TRAC and RETRAN calculations. The RETRAN pressure continues to decrease while the TRAC pressure increases after 200 seconds. This indicates that the RETRAN pressurizer model is closer to equi-librium than that of TRAC. Fiqure 3.6 shows that the RETRAN SG pressure matches that of a TRAC as well as the saturation pressure corresponding to the BNL average system temperature, as expected. In summary, the calculated results of this transient by both codes appear j to be very reasonable. l l__ m _ _.. _ _,.. - ~, _, -.,,,,, ~.

TABLE 3.1 SCENARIO DESCRIPTION TRANSIENT NO. 11 Plant Initial State - Just prior to transient initiator General

Description:

Hot standby, 0% Power af ter 100 hr of shutdown l System Status Turbine: Not latched Turbine Bypass Valves (TBV): Automatic control Atmospheric Dump Valves (ADVs): Automatic control Charging System: Automatic control Pressurizer: Automatic control Engineering Saf ety Features: Automatic control i Power Oper3ted Relief Valves (PORVs): Automatic control Reactor Control: Manual Main Feedwater: In bypass mode, manual control to provide zero level in SGs; I condensate pump, I booster pump, 1 MFWP operating on steam supplied by unit 2. Aux Feedwater: Automatic control Main Stream Isolation Valves (MSIVs): Open, automatic control Main Feedwater Isolation Valves (MFIVs): Open, automatic control Transient Initiator - A full double-ended guillotine pipe break in steam line A upstream of the MSIV and downstream of the flow restrictor. Equipment Failures which occur during the accident transient if the equipment is demanded. Both MSIVs f ail to close. Operator actions / inactions i t a. Operator will turn off all RCPs 30 seconds af ter SIAS based I on low pressurizer pressure, b. Operator fails to turn off charging pumps prior to full I repressurization. l c. Operator f ails to control repressurization. d. Operator f ails to maintain level in intact SG. e. Operator f ails to respond to high SG alarm at 30". f. Operator f ails to respond to high SG alarm at 50"... l g. Operator turns off the AFW at 8 minutes. h. Operator f ails to manually close the stuck-open MSIVs. l l l

CAUTION: Tne scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 590 600 i i i i 560 -TRAC DOWNCOMER TEMPERATURE ^ 530h BNL SYSTEM AVERAGE - 500^ 5 TEMPERATURE b $500 E a o N470 - 400Q e m E440 E 3 - 3003 s 410 H 9 o 9 a 380 o n-o 9 - 200 9 >350 320 100 290O 500 1000 1500 2000 2500 3000 3500 TIME (s) Figure 3.1 Transient 11: Liquid Temperature LAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 7 xlO I'0 I I I I I i 1.6 M*L 12 l4 ~ 10 y m A'. - a$ 3 1.2 6-e. / y 1.0 h-g' 6y {O.8 ~/- 3 3 4 $ .6 {'

/

(TRAC) 2 - PRESSURIZER PRESSURE O a PRESSURIZER LEVEL (TRAC) [

['

O O' 4 - PRESSURIZER LEVEL (BNL) y k" "/ -- -2 ' 0.2 I I O.O O 500 1000 1500 2000 2500 3000 3500 TIME (s) Figure 3.2 Transient 11: Pressurizer Pressure CtuTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. x t 06 7 i i i i i i 900 6 STEAMLINE PRESSURE (TRAC)-

---

SATURATION PRESSURE CORRESPONDING TO SYSTEM - 750 2 '\\

^5 AVERAGE TEMPERATURE (BNL) g [4 ' 'N - 600g h3 - \\ - 450h h2 N - 300 a N w tc i ~~._,- 150 ct 0 - 0 I I I I I I I I I O 50 10 0 150 200 250 TIME (s) Figure 3.3 Transient 11: Steam Line Pressure CAUTION: The scenario simulated contains significant conservatisms in operator i actions and equipment failures. For details, see the scenario descriptions. 600 - i i i i 5501 RETRAN o TRAC 500 a BNL g p450 a D 400 5 g350 w

• 300 o

~ 250 0 I I I 200 O 10 0 200 300 400 500 600 TIME (s) Figure 3.4 Transient 11: Temperature in the Vessel Downcomer CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 2600 i g i i 2400 2200 RETRAN s 2000

---

T R AC

1800 4

$1600 -I W1400 -1 cr 31200 '} 01000 ( E 800 - ( - - - - - - ' ' - - - ~ ~ ~ _ _ 600 400 200 ~ I I I I I 0 O 10 0 200 300 400 500 600 TIME (s) Figure 3.5 Transient 11: Pressurizer Pressure CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 1000 i i i i i RETRAN 800 o TRAC a BNL rn 3 600 6 w C U$ $400 e CL 200 ^ O a ^' O O 10 0 200 300 400 500 600 TIME (s) Figure 3.6 Transient 11: Pressure in the Steam Generators 1 I l 2 4. TRANSIENT 3: HFP 1-FT STEAM LlhE BREAK IN HFP CONDITION 2 Transient 3, which is initiated by a 1-f t break at the main steam line, is similar to Transient 1 except that it started during the HFP operation. In general, the course of the transient is also similar to that of Transient 1. Table 4.1 shows the transient scenario as specified by ORNL. The initial steady state f or HFP operation is presented in Table 4.2. Figure 4.1 compares the total steam flow f rom the steam generators in-cluding break flow for both transients. The steam flow rate is higher f or Transient 3 due to the higher initial energy level (higher average tempera-ture) and higher decay heat production f or the HFP condition. However, the initial water inventory is less for Transient 3. These combined effects cause the broken SG to become empty much f aster than f or Transient 1. Figure 4.2 shows the TRAC downcomer temperature and the BNL system average temperature. As in Transient 1, the BNL average temperature is initially higher than the TRAC temperature due to the delayed cooling of metal walls, but they eventual-ly converge. As expected, the minimum temperature is much higher in Transient i 3 than in Transient 1 owing to higher initial temperature, higher decay heat, and less initial water inventory in the SGs. Once the SG dries out, the tem-perature starts to slowly increase, since the decay heat production is still higher than the cooling by HPI and AFW. The temperature is expected to con-tinue to increase at an estimated rate of 1.5'C for every 1000 seconds until it reaches approximately 535'K around 6000 seconds, and then start to decrease slowly owing to decreased decay heat production. Figure 4.3 shows the TRAC and BNL primary pressures and BNL pressurizer water level. The TRAC pressure agrees very closely with the BNL pressure which represents the highest rate of pressure rise due to the adiabatic compression assumption. The pressure in this transient is expected to be closer to adiabatic conditions than in Tran-sient 1, since the temperature of the water surging into the pressurizer is higher and is increasing. Figure 4.4 shows the TRAC pressure in the SG secondary sides and the sa-turation pressure corresponding to the BNL average temperature. As expected, the broken SG pressure stays at the atmospheric pressure as it becomes empty. The intact steam generator pressure also agrees well with the saturation pres-sure, contrary to Transient 1, where a severe nonequilibrium eff ect is exhi-bited in the intact steam generator. The intact SG is estimated to be com-pletely full at about 550 seconds. In summary, the TRAC calculation of this transient appears to De reason-able. l ' l I

TABLE 4.1 SCENARIO DESCRIPTION TRANSIENT NO. 3 Plant Initial State - Just prior to transient initiator General

Description:

100% Power steady state Systen Status Turbine: Automatic control Turbine Bypass Valves (TBV): Automatic control Atmospheric Dump Valves (ADVs): Automatic control Charging System: Automatic control Pressurizer: Automatic control Engineering Safety Features: Automatic control Power Operated Relief Valves (PORVs): Automatic control Reactor Control: Automatic Main Feedwater: Autonatic Aux Feedwater: Automatic control Main Stream Isolation Valves (MSIVs): Open, automatic control Main Feedwater isolation Valves (MFIVs): Open, automatic control 2 Transient Initiator - A 1.0-f t hole appears in steam line A outside containment upstream of the f151V and downstream of the flow restrictor. Equipment Failures which occur during the accident transient if the equipment is demanded. None Operator actions / inactions a. Operator will turn of f all RCPs 30 seconds af ter SIAS based on low pressurizer pressure. b. Operator fails to turn off charging pumps prior,to f ull repressurization. c. Operator f ails to control repressurization, d. Operator f ails to maintain level in intact SG. e. Operator f ails to respond to high SG alarm at 30". f. Operator f ails to respond to high SG alarm at 50"... -

TABLE 4.2 Comparison Between TRAC and Design / Plant at Full Power Conditions. Design / Plant TRAC Data Predictions PRIMARY SIDE Core power 2694 MW 2700 MW 3 3 Vessel flow 25.27 m /s 25.28 m /s (401,121 gpm) (401,324 gpm) APvessel 0.28 MPa (40.65 psid) AP 0.19 MPa 0.24 MPa sg (28.15 psid) (34.60 psid) APloop 0.54 MPa 0.538 MPa (78.73 psid) (76.28 psid) Thot 585.7 K 585.6*K (594.6*F) (595.1 F) Tcold 559.3 K 559.6 K (547.0*F) (547.6*F) ATvessel 26.4*K 26.4 K (47.6*F) (47.5*F) SECONDARY SIDE Feedwater flow per SG 749 kg/s 737 kg/s (5.95 Mlb/hr) (5.85 Mlb/hr) SG Dome Pressure SG 11 5.90 MPa 5.90 MPa (856 psia) (852.9 psia) SG 12 5.86 MPa 5.89 MPa (850 psia) (853.7 psia) MFW Pump Discharge 1 Pressure MFW 11 7.8 MPa 7.67 MPa (1130.7 psia) (1112.6 psia) MFW 12 7.63 MPa 7.57 MPa (1106.7 psia) (1097.4 psia)

Table 4.2 (cont) Design / Plant TRAC Data Predictions Feedwater 494.8 K 496.2*K Temperature (431.0*F) (433.5*F) MFRV % open -90 88.9 SG liquid mass 62,350 kg 64,600 kg (137,458 lb) (142,419lb) l i i l l

CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 120,000 i i i . i i i i i i ~ ' 5100,000 3o J ' 80,000 m m< 2 60,000 TRANSIENT I 2 <w TRANSIENT 3 y 40,000 J<l5 20,000 F 0 O 500 1000 1500 TIME (s) Figure 4.1 Transient 3: Total Integrated Steam Flow From Steam Generators CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. t 600 590 i i i i i i i 570 - 560 g550 - 520E t w w g530 480$ ~~~~~~~~~~~ D Q 510 - ' 44o$ 's _. $490 g l - TRAC - 400W 470 BNL ~ 9 9 8450 8 - 320 430 410 - 280 I I I I I I I 390 O 500 1000 1500 2000 2500 TIME (s) Figure 4.2 Transient 3: Downcomer Liquid Temperature,

CAUTION: The scenario simulated contat.is significant conservatisms in operator actions and equipment failures. For details, see trie scenario descriptions, i 1.8 107 i i 1.4 107 ~/'/. , 10 _ 1.6 107 _ '.. ~,_ g 3 n /* , C'. - 6]w 7 S 1.2 10 / _ 43 p- ~~~~ 2" $1.0107 g w 8.0 106:- ow a u' n .3 -TRAC PRECURE iE g 6.0 106- - BNL PHESSURE 3 4.0 106

---

BNL PRESSURIZER WATER d

LEVEL m 2.0 106 a 0.0 ' 500 1000 1500 2000 2500 3000 3500 TIME (s) Figure 4.3 Transient 3: Downcomer Pressure CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions, al06 0 1 I I I I I I I I I I l 7 BROKEN (TRAC) INTACT (TRAC) O SATURATION PRESSURE 6 [ CORRESPONDING TO THE ~ i AVERAGE TEMPEPATURE(BNL) -S5 q u a4 x v) o m O *b..- 3 c. 0 o 2 i 0 O 400 800 1200 1600 2000 2400 REACTOR TIME (s) Figure 4.4 Transient 3: Pressure in the Steam Generators 4 - _ _. _.

t 5. TRANSIENT 6: AFW OVERFEED FROM HFP This transient is initiated by the loss of MFW due to MFW pump trip. Ad-ditionally, the AFW does not start on AFAS until AFW pumps to both steam gen-erators are started by the operator at 20 minutes into the transient. The transient scenario is shown in Table 5.1 4 Upon loss of the MFW, the SG level starts decreasing, and the reactor and turbine are tripped on low SG level. The ADVs and TBVs also open at the same time. The average primary temperature and pressure start decreasing rapidly. I However, since the ADVs and TBVs are programmed to open and close on the pri-mary temperature of 552 K, the primary temperature is maintained at this set ) point until both SGs are empty. Figure 5.1 shows that both TRAC and BNL cal-culations maintain this temperature until about 800 seconds. Figure 5.2 shows the total steam flow through the ADVs and TBVs necessary to maintain this con-dition during this period f or both TRAC and BNL calculations. It shows good agreement. This also holds the secondary temperature at this temperature (or slightly below 550*K) and the SG pressure at the saturation pressure ( 60 bar) corresponding to this temperature (Figure 5.4). In the meantime, the primary pressure slowly starts increasing due to the pressurizer heaters (Fig-l ure 5.3). Once both SGs dry out, there is no cooling of the primary system. 1 The primary temperature starts increasing due to the decay heat, and the pri-mary pressure increases f aster until it reaches the PORV set point. The secondary pressure starts decreasing since the ADV/TBVs are still open due to high primary temperature. This decreasing pressure causes the SGIS to isolate the TBVs. However, since the ADVs still remain open, the secondary pressure continues to decrease. The results obtained by both the TRAC and BNL calcula-tions generally match the expected trend as described above up to this point j (1200 seconds). In the TRAC calculation, the opening of ADV was programmed to depend on the primary temperature only. However, there was a question if it should also j depend on the secondary pressure. If ADVs are programmed this way, the i secondary side pressure would be kept at the set pressure of the ADVs (proba-bly 60 bar) and the secondary side temperature at the corresponding saturation temperature (552"K). This, in turn, would keep the average primary tempera-ture near this temperature, which would be equivalent to controlling the pri-mary temperature. The only time the secondary side behavior deviates f rom this course significantly would be when the secondary side water level is very low and, therefore, the steam generator as a heat sink is lost anyway. Therefore, even this change of programming on the control of the ADVs would not make any significant dif f erence for the rest of the transient. I At 1200 seconds af ter of the transient began, the operator activates the AFW pumps to both SGs. The specified AFW flow rate for this transient is much higher than the normal AFW flow rate (about 5 times). Figure 5.1 compares the TRAC downcomer temperature with the BNL system average temperatures calculated i by use of two dif f erent AFW flow rates. BNL 1 represents the result obtained by use of the same AFW flow rate as in the TRAC calculation, and BNL 2 repre-sents the expected result if normal AFW flow rate is used. For either case, initiation of AFW causes rapid cooling of the primary system. The BNL 1 tem-perature matches the TRAC temperature very closely. The BNL 2 temperature de-creases much more slowly, as expected. Figure 5.3 shows the TRAC primary pressure and the BNL pressurizer water level by using both AFW flow rates. I .--.---w, -,,m -4 c--,-

---

y---

---e--.= mm--

J The BNL 2 level is shown to remain much higher than that of the BNL 1. This may indicate that the primary pressure may never reach the SIAS set point (121 1 bar) and/or the HPI pump shutof f head (88.7 bar)or it may take a much longer 1 time. This would f urther slow down the temperature drop and also lessen the rate of pressure increase by injecting less HPI/ charging water into the pri-mary system i Figure 5.4 shows that the SG pressure f or both the TRAC and BNL 1 calcul-l ations rapidly decreases due to the temperature drop f ollowing the AFW initia-tion. It also indicates that the SGs are completely filled at about 3000 seconds f or both calculations. On the other hand, the BNL 2 results show that j the secondary side pressure stays high and then starts to decrease slowly. The SGs are estimated to be full at about 9500 seconds according to the BNL 2 calculation. For both the BNL 1 and BNL 2 calculations, the temperature is expected to continue to decrease beyond 6000 seconds, until it eventually reaches the steady-state temperature, where the decay heat balances with cool-ing due to charging and AFW flow, if the AFW flow persists. We estimate this I temperature to be 300*K and 380'K f or the BNL 1 and BNL 2 cases, respec-tively. However, the feedwater supply system may run out of water long before I that time, especially for the higher AFW flow case (BNL 1). ] The TRAC calculation shows severe asymmetric pressure and pressure oscil-1 lation between the two steam generators (Figure 5.4) when the AFW is intro-duced. This appears to be due to condensation caused by introduction of a j large amount of cold AFW. However, this phenomenon may not aff ect the pri-mary pressure and temperature significantly, since the overall cooling may de-pend on the total amount of AFW rather than its distribution. Also, this j phenomenon is not expected to happen f or the normal AFW flow rate case. In summary, the TRAC calculated results appear to be acceptable, given the high AFW flow rate. I ) i

• This transient was recalculated by LANL using the normal AFW flow rate, which corresponds to the BNL2 case. Figure 5.5 compares the temperatures cal-culated by TRAC and BNL simple method for both AFW flow rates. They agree very closely for each case.

3 I i ! j 4 i i I

TABLE 5.1 SCENARIO DESCRIPTION TRANSIENT NO. 6 Plant Initial State - Just prior to transient initiator General

Description:

100% Power steady state System Status Turbine: Automatic control Turbine Bypass Valves (TBV): Automatic control Atmospheric Dump Valves (ADVs): Automatic control Charging System: Automatic control Pressurizer: Automatic control Engineerina Saf ety Features: Automatic control Power Operated Relief Valves (PORVs): Automatic control Reactor Control: Automatic Main Feedwater: Automatic Aux Feedwater: Automatic control Main Stream Isolation Valves (ftSIVs): Open, automatic control Main Feedwater Isolation Valves (MFIVs): Open, automatic control Transient Initiator - Both Main Feedwater pumps trip simultaneously. Equipment Failures which occur during the accident transient if the equipment is demanded. Aux Feedwater pumps f ail to start. Operator actions / inactions a. Operator will turn off all RCPs 30 seconds af ter SIAS based on low pressurizer pressure. b. Operator fails to turn off charging pumps prior to full .i repressurization. c. Operator f ails to control repressurization. d. Operator f ails to maintain level in intact SG. e. Operator f ails to respond to high SG alarm at 30". f. Operator f ails to respond to high SG alarm at 50"... I g. Operator initiates actions to correct Aux flow problem and overrides l Aux flow control to provide max' flow at 20 minutes. Aux flow control valves turned wide open, and all AFW pumps started. 1 Note: Terminate computer run if and when auxiliary feed tanks are empty. l I ' l l

CAUTION: The scenario simuleted contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 600 i i i i i i i - 600 2 550 1

• N s ~~ -.

-. soo-[ g ~ y .%.s~~~. w w x x D500 \\ a s N r 4 \\ 400 4 \\ C E 450 \\ E s s s s w s - 300 W H 's H e400 's s a 5 5 9 TRAC ( DC) - 200 g J 350

---

BN L I ( AVER AGE )

J - - BNL 2 ( AVERAGE) 100 300O 2000 4000 6000 8000 TIME ( s ) Figure 5.1 Transient 6: Liquid Temperature CAUTION: Tne scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. xlO3 80 i i i i i 70 3 xlO 25 3 5 60 o 3 125 3o 3 50 d ] t 00 m < 40 m j y s / TRAC 75 o o 30 -/ BNL W 50$ e 20 e 0 Y $_ 10 -{ 25 z O l O I -10O 200 400 600 800 1000 1200 1400 1600 TIME (s) Figure 5.2 Transient 6: Total Steam Flow From Steam Generators CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 20 i i i i i 18 E 10~ ,. C - j 16 p ~ 014 '~ ui /,, - 5 _l2 r, f '. g $IO t___ln-N o \\ - 03 $8 N e TRAC PRESSURE d 6 1 4 PRESSURIZER LEVEL- -52

---

B N L I Dm

--- B N L 2 m 2 PRESSURIZER LEVEL E O I I O 2000 4000 6000 8000 TIME (s) Figure 5.3 Transient 6: Primary Pressure CAUTION: The scenario simulated contains significant conservatisms in operator actions ard equipment failures. For details, see the scenario descriptions. 10 i i i I400 TRAC STEAM GEN-A - TRAC STEAM GEN-B 8 ---BNLi - 1200 -- BNL 2 l r--- H b  ? )$ lO----- - 1000 9,, 6 I O ~ l \\- I - 800 @ m l l m [0 4 \\ - 600 $ x x \\ .t I ~.s n. o. L :: I N 400 \\i. I 2 l- - 200 0 O O 2000 4000 6000 8000 TIME (s) l Figure 5.4 Transient 6: Secondary Pressures J i

CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 600 I I - 600 2 550 1 p.5 s : --. s, ' - 500-C g' W w --.s, m g 3500 - s -V. 5 N - 400 4 e N C E 450 's E 2 r 's - 300b \\ w H e400 's a s 5 TR AC i (DC) 5 o -- TR AC 2 (DC) - 200 g i350 - ----- BN L I ( AVER AGE ) J - - BNL 2 ( AVERAGE) 100 3000 2000 4000 6000 8000 TIME ( s ) Figure 5.5 Transient 6: Comparison of the Liquid Temperature Calculated by TRAC Code and BNL Method Using Two Dif f erent Auxiliary Feedwater Flow Rates ' i

i i t 6. TRANSIENT 7A: SMALL BREAK LOCA IN A HOT LEG 2 l This transient is initiated by a small break (0.02 ft ) in a hot leg dur-ing HFP operation. The transient scenario is presented in Table 6.1. Figure 6.1 shows the temperatures at various locations calculated by TRAC l and the BNL system average temperature. The BN! temperature agrees ressonably well with the average of the hot and cold leg temperatures. Figure 6.2 shows l the TRAC primary pressure and the saturation pressure corresponding to the BNL l system average tenperature. The figure indicates that the primary pressure is very close to the saturation pressure, as expected for LOCA, but still the primary system is generally subcooled. At the initiation of the break, the primary pressure starts decreasing j rapidly (Figure 6.2). At 2100 psia (144.8 bar), the low primary pressure trips the reactor and turbine. ADVs and TBV open at the same time. The pri-mary pressure continues to decrease due to continued loss of mass through break, triggering the $1A3 at 1740 psia (121 bar) and reactor coolant pump trip at 30 seconds af ter that. Introduction of HPI makes up some of the water l lost through the break and reduces the rate of depressurization. l The water lost through the break is mostly in the liquid form and, there-f are, does not contribute significantly to the energy loss f rom, or cooling of, the system. (This would be a major difference between this transient and a similar small break LOCA transient due to f ailure of the PORV to close. The i primary temperature would decrease f aster in the PORV LOCA transient.) Cool-ing in this transient is achieved mainly by the opening of the ADV/TBVs and the release of some steam f rom the SGs. Opening these valves initially causes I a sharp drop of the primary temperature. However, since the ADV/TBVs are con-i trolled to open or close at the average primary temperature of 552*K, the average primary temperature is maintained at 552 K once it reaches this tem-perature (Figure 6.1). Figure 6.3 shows the total steam flow necessary to maintain this condition during this period f or both TRAC and BNL calculations, which shows very good agreement. This state continues until either the AFW starts due to low SG 1evel (which does not appear to happen in this transient) I or the continued cooling by the HPI finally exceeds the decay heat. This brings the primary temperature down below 552'K and closes the ADV/TBVs. The primary temperature continues to f all due to cooling by HPI/ charging. Since the break flow is slightly higher than the HPI, the system pressure also con-tinues to decline slowly until the becak flow finally balances with the HPI flow and then levels off. In the TRAC calculaticn, the TBVs are closed a lit-tie earlier than expected because of SGIS at 502 seconds, which is caused by high containment pressure. (Information on how the containment pressure was l calculated was not available.) However, this does not make any significant i diff erence since the ADVs remain open until the primary temperature f alls be-low 552 K. The secondary temperature also slowly decreases at the same rate as the primary temperature starts to decrease. The TRAC calculation shows that secondary pressure stays constant for a long tine (Figure 6.4), while it is expected to decline corresponding to the saturation pressure of the declin-ing primary tenperature. Also, one of the SGs started to empty and asymmetric pressure starts to develop at about 5000 seconds, while there appears to be no particular event to cause this. The cause of this anomaly appears to be the inadequate condensation model of the TRAC code. d l l ., =

During the presentation of this transient calculation by LANL on December 13, 1983, there was some discussion concerning the stagnation of the primary loops due to voiding at the steam generator U-tubes. Subsequently, LANL is-sued a report [2] contending that voiding at the U-tube did not happen be-cause 1) the primary system remained subcooled during the entire transient due to the nonequilibrium eff ect built into the TRAC code and, thus, the reverse heat transf er f rom the SG was not suff icient enough to cause boiling of pri-mary fluid at the U-tubes; 2) the liqu' d level in the upper plenum did not de-crease below the hot leg penetration level. BNL generally agrees with these arguments. As discussed earlier, Figure 6.2 shows that the primary pressure remains above the saturation pressure curing the entire transient. This is partly due to the non-equilibrium effect, as mentioned by LANL, and partly to the fluid in the upper done remaining hotter so that its saturation pressure is higher than the rest of the primary system. Figure 6.5 shows the vapor space volume in the primary systec di ring the transient calculated by BNL. It confirms that the upper dome is never completely empty in this particular transient. In summary, it appears that the sequence of events of the calculations generally follows the expected trend, and the TRAC-calculated results are reasonable except for the SG secondary side pressure. i . i 1 J

TABLE 6.1 SCENARIO DESCRIPTION TRANSIENT NO. 7A Plant Initial State - Just prior to transient initiator General Psscription: 100% Power steady state ) System S'.atus j Turbine: Automatic control i Turbine Bypass Valves (TBV): Automatic control Atmospheric Dump Valves (ADVs): Operating / Automatic control Charging System: Automatic control Pressurizer: Automatic control Engineering Safety Features: Automatic control Power Operated Relief Valves (PORVs): Automatic control Reactor Control: Automatic Main Feedwater: Automatic control Aux Feedwater: Automatic control Main Stream Isolation Valves (MSIVs): Open, automatic control Main Feedwater Isolation Valves (MFIVs): Open, automatic contro.* 2 Transient Initiator - A 0.02 f t hole appears in the hot leg of loop A. Equipment Failures which occur during the accident transient if the equipment is demanded. None Operator actions / inactions a. Operator will turn off all RCPs 30 seconds af ter SIAS based on low pressurizer pressure. b. Operator f ails to turn off charginq pumps prior to f ull repressurization. c. Operator f ails to control repressurization. d. Operator f ails to maintain level in intact SG. e. Operator f ails to respond to high SG alarm at 30". f. Operator f ails to respond to high SG alarm at 50"... n - _

CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 600 i i i i i i i 600 575 7 [ 560 C h550 6% M' 520 0 a Q 525 -q s'N 480D ' se s 's N s 440a 's 5 N a.500 s Nb s w w 475 N 400F H o o 5 -- HOT LEG (TRAC) g 450 --- COLD LEG (TRAC) 360 5g 8 o SYSTEM AVERAGE (BNL) 320 ' 425 400 l I O 2000 4000 6000 8000 TIME (s) Figure 6.1 Transient 7A: Liquid Temperature CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 20 i I l l - 2800 'O ~ - TRAC I6 >. BNL (SATURATION PRESSURE-2400 CORRESPONDING TO SYSTEM } I4 AVERAGE TEMPERATURE) - 2000 3 I2 E $a 10 1600y a m m O8 l200 m E 6 o o 800 4 o 400 2 I I 0 O O 2000 4000 6000 8000 TIM E (s) Figure 6.2 Transient 7A: Primary Pressure I

CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. i i i i i i 30,000 -P--------- '.B. x l TRAC 6 ,I --- B N L ' 20,000 ff F 1 I m O I W IQOOO [- E I I I I O' O 1000 2000 3000 4000 5000 6000 7000 REACTOR TIME (s) Figure 6.3 Transient 7A: Total Steam Discharge CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 10 I I I I I I I 1400 STEAM GEN.lt (TRAC)

---

STEAM GEN.12(TRAC) j 6

~ o SATURATION PRESSURE - I200 CORRESPONDING TO THE AVERAGE SYSTEM TEMPERATURE (BNL) ] - 1000 g 36 W 0 's - 800[ l w o l 'S m a m $4 o 600$ m w 1 E 400' 2 200 I I I I I I 0 O O 2000 4000 6000 8000 TIME (s) Figure 6.4 Transient 7A: Secondary Pressures i l

CAUTION: The scenario simulated contains significant c)nservatisms in operator actions and equipment failures. For details, see the scr,1ario descriptions. O i i i i i i g-da E 20 ~ S li w i a: 2 40 1-3 ' ' ~, ' ',, ' gg ws - $8 g 60 a L_ 9 80 I I I I I 100 O 1000 2000 3000 4000 5000 6000 7000 TIM E (s) Figure 6.5 Transient 7A: Vapor Space Volume 1 i, --'1

l i l 7. TRANSIENT 9: MFW OVERFEED TO ONE STEAM GENERATOR This transient is initiated when a MFRV fails to close when the turbine trips during the HFP operation. This allows full f eedwater flow to the af-fected SG until the MFW pumps are finally tripped due to low condenser / hot well inventory. The transient scenario as specified by ORNL is shown in Table 7.1. 1 i Figure 7.1 shows the liquid temperature at various locations calculated i by TRAC as well as the BNI. system average temperature. During the initial 300 seconds while the MFW continues, the primary liquid temperature declines sharply. A large temperature diff erence between the hot and cold legs also persists due to the continued cooling by MFW. The affected SG completely l fills at about 130 seconds and the excess f eedwater is released through the l TBV/ADVs. The HPI begins at about 250 seconds and contributes to further cooling. At about 300 seconds when the MFW finally stops due to low conden-l ser/ hot well inventory, the primary temperature starts climbing rapidly, since the decay heat f ar exceeds cooling by the HPI, which is the only cooling mech-anism for the entire system at the moment. When the primary temperature fi-nally reaches 552 K, the TBV/ADVs open and start releasing steam. BNL esti-mates this time to be about 4200 seconds based on the energy balance. Opening of the TBV/ADVs maintains the primary temperature at 552'K. During this peri-1 i od, the SG continues to lose steam and the AFW starts due to low liquid inven-l tory. Once the AFW starts, the system temperature is estimated to drop at the j rate of about 7 C for every 1000 seconds until it levels off at about 380 K. As shown in Figure 7.2, the secondary temperature generally f ollows a i similar trend to that of the primary temperature. The intact SG temperature generally remains much higher than other parts of the system during the ini-i tial 1000 seconds. This may indicate some stagnation of the intact loop. The behavior of the primary pressure is similar to that o' the primary temperature, as shown in Figure 7.3. The system pressure and pressurizer water level calculated by BNL are also shown in the figure for comparison. As discussed in the introduction, the BNL pressure is obtained on the basis of j the adiabatic assumption, which is the maximum pressure attainable during com-pression. The TRAC pressure is lower than this, as expected. Figure 7.4 shows the TRAC pressure in the secondary sides of the SGs. l The pressure of the aff ected SG remains near TBV set point (61.1 bar) since it is completely full at 130 seconds. However, the intact SG pressure is also l shown to stay above 60 bars even during the initial 300 seconds when the tem-i perature declines steeply. This appears to indicate a severe nonequilibrium effect, which is observed in several other transients. The actual pressure is expected to be somewhere between this pressure and the saturation pressure corresponding to its liquid temperature. Both steam generator pressures start I to decline at 4800 seconds when AFW is initiated, as they should. In summary, the results calculated by TRAC for this transient appear to J be reasonable except f or some parts of the intact SG pressure. . l

TABLE 7.1 SCENARIO DESCRIPTION TRANSIENT NO. 9 Plant Initial State - Just prior to transient initiator General

Description:

100% Power steady state System Status Turbine: Automatic control Turbine Bypass Valves (TBV): Automatic control Atmospheric Dump Valves (ADVs): Automatic control Charging System: Automatic control Pressurizer: Automatic control Engineering Saf ety Features: Automatic control Power Operated Relief Valves (PORVs): Automatic control Reactor Control: Automatic Main Feedwater: Automatic Aux Feedwater: Automatic control Main Stream Isolation Valves (MSIVs): Open, automatic control Main Feedwater Isolation Valves (MFIVs): Open, automatic control Transient Initiator - Turbine trip. Equipment Failures which occur during the accident transient if the equipment is demanded. Main feedwater to SGA fails to run back (remains at 100% power characteristics). Operator actions /ir. actions a. Operator will turn off all RCPs 30 seconds af ter SIAS based on low pressurizer pressure. b. Operator f ails to turn off charging pumps prior to full repressurization. c. Operator f ails to control repressurization. d. Operator f ails to maintain level in intact SG. e. Operator f ails to respond to high SG alarm at 30". f. Operator f ails to respond to high SG alarm at 50"... 1 I i.- -

CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 580 i i i i i 575 560 - 550 ,.. 7=:w.3.E;bq ~....- 525 b E \\ i w w . h, Y $ 540 t N \\g 9;S' - 500 Q \\ ,/ m E olI -lk '/ - 475 E E 520 /

-/ /

s 2 g . ll)' ll 450 W w w H I e-500 - 9 t[ l TRAC HOT LEG (Af fected 8 Intact) o j [! I - - TRAC COLD LEG ( Affected) pf TRAC COLD LEG (Intact 1) - 425 o j,: g J J - y,it 1 lll ---TRAC COLD LEG (Intact 2) 400 480 il,lt7 i O BNL SYSTEM AVERAGE 375 460 O 1000 2000 3000 4000 5000 6000 TIME (s) Figure 7.1 Transient 9: Liquid Temperature CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. 600 i i i i i i i i i i i - 600 580 7 - 570g [560 - 540 w n m i s540 4,- ,-o - SIOH q 4 C g - 480y E520 2 2 - 450 W w F 500 -p / INTACT SG (TRAC) a U ij'

---

AFFECTED SG (TRAC) - 420 ;

c580 -p o BNL AVERAGE o J 390 1 460 - 360 440 l O 1000 2000 3000 4000 5000 6000 TIME (s) l Figure 7.2 Transient 9: SG Liquid Temperatures -

i 1 CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions, j 18 i i 4 , -%%gglg,r 10} 16 j--- I / ~ d d 14 e' 5 ~l ,' a 6m a 0-e j i 5 I / _ 4D 12 w i 3 e i t p t $ 10 $ i TRAC PRESSURE 26 E l l' --- BNL PRESSURE (ADI ABATIC MODEL) Oa ,//

---

BNL PRESSURIZER WATER c.

i 8 i- / j / LEVEL - -2 E k/ - -4 6 s/ 9 l l I I I 4 4 i O 1000 2000 3000 4000 5000 6000 TIM E (s) y. i ~ _ Figure 7.3 Transient 9: System Pressure CAUTION: The scenario simulated contains significant conservatisms in operator actions and equipment failures. For details, see the scenario descriptions. d 70 i 60 ........ I " - / o/ i 50 7 O / oS40 w g 30 w i j */ INTACT SG (TRAC) m' I AFFECTED SG (TRAC) -f 20 --e SATURATION PRESSURE i CORRESPONDING TO THE SYSTEM AVERAGE TEMP.(BNL) 10 } I I I I I O O 1000 2000 3000 4000 5000 6000 i TIME (s) j ~- Figure.7.4 Transient 9: Steam Generator Secondary Side Pressure j + 4 g# # r

8.

SUMMARY

AND CONCLUSIONS Several selected transient calculations performed by LANL using the TRAC-PF1 code for the USNRC PTS study of the Calvert Cliffs Nuclear Power Plant have been reviewed at BNL. Six of the thirteen calculations consisting of two HZP and f our HFP transients have been selected f or the in-depth review. Sim-pie hand calculations, based on the mass and energy balances of the entire reactor system, have been performed to predict the temperature and pressure of the reactor system, and the results have been compared with those obtained by TRAC. In general, the temperatures and pressures of the primary system calcula-ted by TRAC have been very reasonable. The secondary pressures calculated by TRAC appear to indicate that the TRAC code hss some difficulty with the con-densation model-and further work is needed to assess the code calculation of the U-tube steam generator pressure when the cold auxiliary feedwater is in-troduced to the steam generator. However, it is not expected that this uncer-tainty would affect the transient calculations significantly. ) 1 i i 'I f, i 1

REFERENCES 1. J. E. Koenig et al., " TRAC Analyses of Potential Overcooling Transients at Calvert Cliffs - 1 f or PTS Risk Assessment," Los Alamos National Laboratory, December 1983. 2. G. D. Spriggs and R. Smith," Review of TRAC Results for small Break LOCA," Communication with J. N. Reyes, J r., NRC, Los Alamos National Laboratory, February 1,1984. i

1 i APPENDIX A e i J l j as - e 1 4 4.

_=. BROOKHAVEN NATIONAL LABORATORY MEMORANDUM DATE: January 12, 1984 M: Pradip Saha FROM: J. Jo i W BJECT: BNL REVIEW 0F PTS INPUT AND STEADY-STATE CALCULATIONS PERFORMED BY LANL FOR CALVERT CLIFFS NUCLEAR PLANT This is an interim report on the progress of the review of PTS input decks and steady state calculations performed by Los Alamos National Laboratory (LANL) for the Calvert Cliffs nuclear plant which is owned by Baltimore Gas & Electric Co. (BG8E). l As you know, we were not involved in the initial stage of the PTS study on the Calvert Cliffs plant when the geometric and plant operation data were collected from Combustion Engineering and BG&E. Most of the information used 1 for this review was based on the handouts obtained from LANL at the two meet-ings I attended for the study group (one at LANL on June 27, 1983 and the other at BG8E on September 20,1983) and the Calvert Cliffs FSAR. There fo re, as was in the Oconee review, the main emphasis was placed on a general review j of the input deck and steady state output rather than confirmation of specific numbers used in the input deck. Since LANL developed two separate input decks, one for Hot Zero Power (HZP) and the other for Hot Full Power (HFP) conditions, these two de..ks were compared for consistency. The versions of input listings reviewed here were those obtained from LANL at the September review meeting. It is my understanding that some of these input, especially that for HFP, have been modified since then. The TRAC codes used for LANL calculations were some intermediate versions of the TRAC-PFl/ MODI code. It appeared that dif ferent versions of the code l were used for the HZP and HFP inputs. BNL recently obtained a draft copy of i the TRAC-PF1/ MODI manual. However, the input listings were not consistent with some sections of the manual. Most of the findings listed here were comfrunicated to Ms. J. Koenig of LANL over the telephone on November 10, 1983. Specific details of the review and its findings are summarized as follows: I

]

A. Review of Input 1. Connections of the components were checked to ensure that they were consistent with noding diagram including the number of nodes. Several minor discrepancies were found, but they were insignificant. 2. Symmetry of the loops was checked. Several minor non-symmetry were found but they were either real differences in the plant or 4 insignificant. 3. Safety injection tanks and low pressure injection systems were not model ed. It appeared that these components were judged not to be needed for the specified transients. 4. Total volume of the primary side agreed very closely with that given in the FSAR. Individual pipe diameters and volumes including core, primary side of steam generator and pressurizer matched those given in the FSAR. In the secondary side of the steam generator, several differences in volume were found between the HZP and HFP inputs. a. The steam don.e volume of HFP deck = 20.33 m3 HZP deck = 45.40 m3 These differences were already identified by the LANL staff at the September meeting and the steam dome volume of the HFP input deck was corrected. b. The downcomer volume of HFP deck = 28.52 m3 HZP deck = 39.2 m3 1 This difference appears to be significant. Wrong downcomer volume would result in wrong water mass or in the wrong water level if the water mass was matched. Either of these differences may have some effect on the course of the overcooling transient. c. There were significant differences in the hydraulic diameters of the tube regian and the heat transfer area between the primary and the secondary sides of the two input decks. HFP HZP Hydraulic Diameter 0.073 m 0.005 m Heat Transfer Area 9088/10353 m2 7338/8361 m2 (Inner Wall /0 uter Wall) i i - 47 _ 4 -r---- -,c.3-v-3-,r -,,4 %e, -r--

J It appears that the hydraulic diameter was adjusted in the HZP deck while the heat transfer area was increased in the HFP input to match the heat transfer. I do not know if this adjustment is required to match the plant condition, or which is the better way to obtain proper heat transfer if necessary. This would indicate deficiency of the code and should be ex-pl ained. j 5. Different friction factor options were used for different components and sometimes even in the different cells of the same component. Also, no frictions were used in many places. These appear to be questionable and may af fect the natural circulation flow rate which is important in the PTS study. 6. Friction factors of very large magnitude (1021 and 1010) were used at the exit of the steam generator to the steamline and auxil-iary feedwater line, respectively, in the HZP input. However, I did not notice a very high pressure drop across these junctions in the steady state output which was expected from-the high friction factors used. l 7. The single phase homologous curve option was used (not fully degraded two-phase homologous) for the Reactor Coolant Pump (RCP) modeling. i This may cause slightly faster coastdown of the pump' if some vapor 4 I exists when it is tripped. But this effect is considered to be very j minor. 8. In the pressurizer heater model, a certain amount of heat was taken out of the pressurizer when the heater was off. This was explained to me as a heat loss from the pressurizer to the atmosphere. How-ever, since this heat loss would be balanced with addition of heat by the heater no net heat should be removed from the pressurizer during the steady state. 9. Most of the trips and controllers were incorporated correctly as j specified in the LANL handouts except: 4 a. Extensive time smoothing was used in the estimation of the f steamline pressure which was used for Turbine Bypass Valve (TBV) control in the HZP input. This may reduce the sensitivity of those controls, b. The following equation was used to calculate the pressurizer l evel : i L= Ap - 9445 m, 4792./2

where a p was the pressure difference across the pressurizer. i I could not find how the number 4792.72 was obtained here. 3 l (If it is p 611 kg/m, g = 9.8m/sec2.) g, it should be 5987.8 since p = c. Auxiliary feedwater actuation signal ( AFAS) logic was different between the HZP and HFP input decks. Auxiliary feedwater ( AFW) to both steam generators was actuated by the same trip so that i the AFM would be delivered to both steam generators once it was tripped for the HZP input deck, while AFW would be delivered only to the steam generator with higher pressure on asymmetric steam generator detection for the HFP deck as it should be. d. The atmospheric dump valve (ADV) was programmed to remain closed for the HZP case. However, this may be needed to be opened later when the secondary side pressure increases. B. Steady State 1. Generally very good steady state was obtained for the HFP case. 2. Acceptable steady state was obtained for the HZP case. The secondary side energy balance was off by 10%. However, for a very low power case, this may not be important (it takes an extremely long computer time to obtain good steady state for a low power case). 3. The liquid volume in the pressurizer was less than in the FSAR for the HFP case: FSAR 800 ft3 HFP 678 ft3 However, this may be due to operational differences from the FSAR. 4. The liquid in the pressurizer of the HFP case was highly subcooled. I do not believe this is real and could cause a delay of pressure rise even if the heater was activated. 5. For both cases, i.e., HFP and HZP, the void fraction suddenly de-creased just above the tube region of the secondary side (separator region). This does not appear to be real but may not be important because no heat transfer was involved in this region. -

6. There was very little recirculation from the heated tube region to the downcomer (about 10%). This is contrary to the expectation. (I understand the recirculation ratio is very high for low power oper-ation.) The main feedwater was heated mainly by the wall heat trans-fer in the downcomer for the tube region, not by mixing with the re-circulation flow. This needs to be checked. However, this may not have any significant effect on the transient. In summary, the input decks and the steady states were acceptable. However, some of the items mentioned above, particularly items 4.b, 4.c, 6 and 9 of the input review and Items 4 and 5 of the steady state review, should be further addressed or explained by the LANL staff. A m.>_-,_-4 m a em-u.m 2 e au-a a k I f i t 1 4 1 a l T j APPENDIX B 4 ). l. v s .a l 1 l l ,f I f ( l e i l i i i I 2 1 'I t 51 - i .I ,_,-,,--.,_,,...--,. -., ~, _ - -,,,,.. -,,.,, _, -, _, -.., - -. _,.

1 I l APPENDIX B: LANL Response to the BNL Review of the Calvert Cliffs Input Decks and Steady State Calculations input deck: 3 A1. Refer to 4b - The correct downcomer volume is 39.2 m. This was the vol-ume used in the full power input deck as well as the hot zero power input deck. In the full power deck, however, some of the downcomer volume was placed in a TRAC component above the one that was labeled as the down-Comer. A2. Refer to 4c - In order to match steady state conditions with TRAC, it is necessary to adjust the heat transfer area or the hydraulic diameters. Here, two different analysts took separate approaches. Adjusting the heat-transfer area is probably preferable because it will not affect the pressure drops. A3. Refer to 5 - Some erroneous friction factors were found and corrected for later calculations. Transient 1 and 2 were thought to have unrealistic flow oscillations (although small) when the loops should have been almost stagnant. A4. Refer to 6 - These high friction factors are flags to invoke the phase-separation option at that cell face. AS. Refer to 8 - Heat los; from the pressurizer would occur when the heaters had tripped on low level. A6. Refer to 9b - The equation that we finally used was: Level = ap - 9517 4949 This equation gave the correct level indication at steady state. A7. Refer to 9c - The AFW logic was the same for all transients. In earlier decks, however, the AFW flow was zeroed out in a restart deck on an asyn-metric-SG-pressure signal instead of being automatically zerced out. Steady State: 3 A8. Refer to 4 - For the particular steady state that you studied, this was an error. Calculations initiated from this steady state were rerun if the PTS-review group thought it was necessary. A9. Refer to 6 - An exact steady state was not obtained for the hot-zero power case. However, it was thought to be adequate for the purpose of initiating a transient. 4.. - -

,an a=>. --..t. .-a-a - - a i I ( i a APPENDIX C f I l 1 i 1 1 3 i t ) 4 l -l f .t, I i 1 I i. i f Ii i j 4 ... __. _,- _... _.. _.. -,.~ . _,. _,. -. _ _ -. -.. _, _. - -. _,.. ~,., --m ..-m-,-- -,, - --

A PRELIMINARY REVIEW OF TRAC CALCULATIONS FOR CALVERT CLIFFS PTS STUDY

• J. H. Jo LWR Code Assessment and Application Group Brookhaven National Laboratory Department of Nuclear Energy Upton, New York 11973 January 1984
• Prepared under the auspices of the U. S. Nuclear Regulatory Commission.

A Preliminary Review of TRAC Calculations For Calvert Cliffs PTS Study This report documents the preliminary BNL comments on the TRAC calcula-tions performed by LANL for the USNRC PTS study of Calvert Cliffs nuclear plant. A quick review of all the TRAC calculations was performed at BNL based on the LANL draft report [1] and the handouts obtained at the two Calvert Cliffs PTS study group meetings (September 20, 1983 at Baltimore Gas and Electric Co. and December 13, 1983 at Los Alamos National Laboratory). A more detailed review of a few selected transients will follow. LANL performed TRAC calculations of 13 transients as shown in Table 1 (re-produced from the LANL draf t report [1]). Four of these, i.e., Transient Nos. 1, 2,10 and 11 are the hot zero power (HZP) steam line break accidents and the rest of them are various transients for the hot full power (HFP) condi-tion. For the sake of convenience, the order of transients discussed below f has been changed from Table 1. The HZP transients will be first discussed, followed by the HFP transients.

1. HZP Steamline Break Accidents 1.1 Transient 1: 1-ft* Steamline Break This transient was initiated by a 1-ft2 break at the main steam line dur-ing the HZP operation. There was a misstatement in Table 4.1 (Sequence of Events of Transient 1) of the LANL draft report that the minimum pressure was a reached at 700 seconds. Figure 1 (Figure 4.2 of the LANL report) shows that the minimum pressure was reached at about 170 seconds. Figure 2 showed the pressurizer level was decreasing while the pressure of the primary system was increasing between 170 seconds and 550 seconds. This appeared to be contra-dictory. The downcomer liquid temperature (Figure 3) increased substantially from 400 K to 425*K between 1300 seconds and 3000 seconds. However, the decay heat did not appear to be large enough to sustain this temperature rise. A simple calculation based on the energy balance indicated that the temperature should level off during this period. Part of this temperature rise may be due to the gradual mixing of the hotter i r head liquid with the bulk of the liquid in the system. This point will be further checked in the detailed review. The steam line pressure of the intact loop (Figure 4) showed a sharp re-versal when the intact steam generator was isolated at 20 seconds and continued to increase. This appeared to be contradictory since the primary side temperature of the intact loop continued to decrease during this period. The steam line pressure of the intact steam generator showed a sudden drop at about 1600 seconds. However, there was not any particular event to cause this drop at this time. This may have been caused by some sudden condensation computed by the TRAC code which may not be realistic. (Similar behavior was observed in many other transients.) 1 In summary, the TRAC calculation of this transient was acceptable. How-ever, the cause of the temperature rise between 1300 and 3000 seconds and the intact steam generator pressure need further investigation. 1.2 Transient 10: 1-ft* Steam Line Break with Two RCPs Left Running This was a similar transient as Transient 1 except that two of the four RCPs were lef t running when the RCP trips were on. The sequence of events should have been the same as Transient I until the pump tripped.

However, there were some differences in the timing of the events. There appeared to be substantially more feedwater delivered into each steam generator in the be-ginning of the transient than Transient 1.

This may be partly responsible for the above discrepancy. There was no charging flow in this transient after the HPI was terminated at 950 seconds (Figure 5) while charging continued in all other transients. The downcomer liquid temperature (Figure 6) at the end of the transient calculation (at 5300 seconds) was about 30 K higher than that of Transient I at the same time (425'K for the Transient I and decreasing; 455*K for Transient 10 and stabilized). This difference can be accounted for by the extra energy added by the pumps and appeared to be in the right range. There was no HPI/ charging after 950 seconds; yet the rate of pressurization was about the same as Transient I where the charging continued. Absence of charging should lessen the pressurization rate. In summary, the major difference between this transient and Transient 1 would be the extra energy added by the pumps, which resulted in the higher downcomer liquid temperature. The calculated results, with the exception of the primary pressure, appeared to be reasonable. The pressure increased faster than Transient 1 despite no charging. This point needs further clari-fication from the LANL staff. !

1.3 Transient 2: Full Double-Ended Guillotine Steam Line Break This transient was initiated by a full double-ended guillotine break in a steam line. Additionally, the Auxiliary Feedwater (AFW) to the broken steam generator was not manually isolated. Although the size of the break of this transient was very different from that of Transient 1, the total break flow and the pressure history of the secondary side of both steam generators were not that different from Transient 1. This was expected since the location of the break was after the flow re-strictor and, in any case, the amount of integrated mass loss was eventually equal to the initial liquid mass in the broken steam generator. Based on the overall energy balance, the major difference between this transient and Trans-ient I would be the failure of the operator to isolate the AFW rather than the break size. This additional AFW to the broken steam generator would act as a continuous heat sink and eventually bring the entire system to the saturation temperature of the broken steam generator, i.e., 100 C (373 K), which the calculation did (Figure 7). This condition would continue until the broken steam generator was com-pletely fi,lled with water. Then the entire system temperature would slowly decrease again toward about 350 K (the expected steady state liquid temperature to balance the decay heat with HPI/AFW flow) if the AFW to the broken steam generator persisted. We estimate this time to be approximately 12000 seconds. In summary, the TRAC calculation of this transient appeared to be very reasonable. _

1 1.4 Transient 11: Full Double-Ended Guillotine Steam Line Break with Failure of MSIV to Close i This transient should be the same as Transient 2 except that both MSIVs i i failed to close on the Steam Generator Isolation Signal (SGIS) so that both i SGs continued to blowdown. However, the timing of SGG was different between these two transients. Also the timing of the Auxiliary Feedwater Actuation Signal (AFAS) was very different. These differences should be clarified by the LANL staff. However, these details in the beginning of the transient would not affect the course of the transient significantly. This appears to be a relatively simple transient to analyze. The system would rapidly cool down due to the blowdown of both SGs until the system l reached the saturation temperature corresponding to the atmospheric pressure (i.e., 373*K or 100*C). The system would remain at this temperature until l j both SGs were finally empty. The calculation confirmed this trend (Figure 8). Based on the TRAC calculated break flow and assuming that the break flow was l all vapor, the system should reach 100*C at about 90 seconds. However, con-I siderably longer time was taken to reach this temperatur? for the TRAC calcu-i i lation. This may indicate that a substantial amount of liquid was entrained through the break. This entrainment would also explain the slower depres-l surization rate than expected from the simple energy balance. The secondary side pressure of the SGs decreased approximately along the eaturation pressure of the system temperature (Figure 9). The system reached the PORV se't point earlier than Transient 2 despite similar temperature be-f l havior. This' may be due to the longer HPI period. Both steam generators were l expected to be empty at 7800 seconds based on the liquid mass in the SGs at 3000 seconds. i In summary, the TRAC calculated results of this transient appeared to be i l very reasonable. 60 - I ++-,---m--e. .wm,,,,- e,w--- ,,--m,--%e-w w er-nvn----mm-evm m n mwe -fm'- op w --i wv wr--w*-t-r--** v 7 v-ww +g e=,--ee,ur-w-

• --e+-*+

t **- - - - - *- +--

l l l 2. HFP Transients 2 2.1 Transient 3: 1-ft Steam Line Break 9 Transient 3, which was 1.0 ft2 steam line break from full power, was similar to Transient 1 except it started from HFP. This transient calculation was repeated because there was an error in the steady state temperature in the pressurizer, which might have caused faster depressurization than expected. This could result in longer periods of safety injection and eventually lower system temperature. Comparison between the original and corrected calcula-tions confirmed this. The primary temperature of the corrected run started to deviate from that of the original run at about 300 seconds, consistent with l the HPI flow rate which showed a substantial difference between the two runs after 300 seconds. The broken steam generator dried out substantially earlier l in the corrected calculation than in the original calculation. This appeared l somewhat puzzling since the timing of the SGIS was about the same between both runs. The magnitudes of the primary temperature drop before, and rise after the SG dry-out were in the right range for both runs based on the total break 1 flow and decay power by ANS curve (ANS decay power curve was used throughout this review of the HFP transients. Information on actual power generated was not available at the time of review; the power was calculated by the point kinetics with the reactivity table in all TRAC HFP calculations). The broken steam generator dried out earlier (less initial water mass) and the primary pressure reached at the PORV set point earlier than the HZP case (Transient 1) as expected. The calculated results with the corrected steady state pressurizer temperature appeared to be reasonable. 2.2 Transient 4: Turbine Trip with TBV Stuck Open 1 Transient 4 was not reviewed here since there was an error in the input which caused a significant difference in timing for the major events such as.

t the SI AS. The initial portion of this calculation (0 - 570 seconds, up to SGIS) would be the same as that of Transient 4A as discussed below. 2.3 Transient 4A: Turhire Trip With One TBV and One MSIV Stuck Open This transient is basically similar to Transient 3. The major difference I was the rate of mass loss at the break. The rate of mass loss through the 4 break (i.e., stuck-open valves) was much smaller in this transient a nd, 1 accordingly, this was a much milder transient. The rate of temperature drop (Figure 10) was lower, the minimum temperature was higher and the pressure l (Figure 11) changed much slower than Transient 3. The temperature change matched approximately those obtained by simple hand calculations based on the a mass and energy balances. The primary system pressure leveled off at 600 seconds and started increasing at 1000 seconds while the primary system temperature continued to decrease. There was no HPI or charging flow during this period. This appears to be contradictory and needs further explanation. i The timing of the broken steam generator dryout was within the expected range. The pressure of the intact steam generator changed approximately along the i saturation pressure of the calculated temperature, which was expected. j In summary, the calculated results appear to be reasonable except for some portion of the primary pressure response. 4 2.4 Transient 7: Small Break LOCA With Artifically Blocked Natural Circula-tion l Transient 7 was not reviewed since this calculation involved artificial and unrealistic blockage of the primary loop. i 2.5 Transient 7A: Small Break LOCA 2 i This transient was initiated by a small break (0.02 ft ) in a hot leg. l Figures 12 and 13 show the calculated primary temperature and pressure, re-i spectively. At the initiation of the break, the primary pressure would start l l I ,,,_,__,---,n.n.n,-

decreasing sharply. At some point (2100 psia), the low primary pressure would trip the reactor and turbine. ADVs and TBV would open at the same time. The prima ry pressure would continue decreasing due to continued loss of mass through break, triggering the HPI and pump trip at 1275 psia. Introduction of HPI would reduce the rate of depressurization and pressure would increase much slowly. Meantime, opening and closing of ADVs and TBVs would maintain the average primary and secondary temperature at 552 K (set temperature of the ADVs and TBVs) or near it. This state would continue until either the AFW started due to low SG level (which did not appear to happen in the calcula-tion) or the continued cooling by HPI eventually brought down the primary tem-perature below 552*K. This would close the ADVs/TBVs. The primary tempera-ture would continue to fall due to the cooling by HPl. Since the break flow j was slightly higher than the HPI, the system pressure would continue to drop 1 j slowly until the break flow finally balanced with the HPI flow. The secondary pressure and temperature would also slowly decrease along the primary temper-ature and its saturation pressure. In the calculation, the TBVs were closed ) because of SGIS, which was caused by high containment pressure. (Information was not available on how the containment r,res sure was calculated.) This closed the TBVs a little earlier than expected. However, this may not have made any significant dif ference since the ADVs were still open. The calcu-lation showed a i onstant secondary pressure (Figure 14) maintained for a long time af ter the ADVs were closed despite the declining primary temperature, and some asymmetric pressure between the two SGs at around 5000 seconds (Figure i 14). Also, one of the SGs started to empty or boil while there appeared to be no event to cause this. l ! i

In summary, the sequence of events of the calculations generally appeared to follow the expected trend except the-SG pressure. There was insufficient information in the handout to check the timing of the events or the magnitude of the temperature and pressure change. 2.6 Transient 5: Primary PORV and Secondary ADV Stuck Open The transient was similar to Transient 7A until one of the ADVs failed to close. Also, it appeared that the flow through the PORV was smaller than the break fl ow of Transient 7A. After the reactor trip, the stuck-open ADV, instead of maintaining the primary temperature by opening and closing, allowed the continuation of steam generator blowdown. This caused the prima ry temperature and pressure to continue to decrease (Figures 15 and 16;. The HPI increased as the system pressure decreased and, at some point (around 70 bar), the HPI flow matched the PORV flow. This stabilized the pressure. If this condition persisted, the primary temperature would eventually drop near the HPI water temperature. Again, there was not enough information available such as PORV and ADV fl ow rate, in the handout to check the magnitude of the tenperature drop or timing of the events. 2.7 Transient 6: AFW Overfeed This transient was initiated by the loss of MFW due to MFW pump trip. Ad-ditionally, the AFW did not start on AFAS until 20 minutes into the transient when AFW pumps were started to both steam generators by the operator. Upon the loss of the MFW, the SG level would start decreasing, and reactor and turbine would be tripped on low SG level. The ADVs and TBVs would also open at this time. The average primary temperature and pressure would start decreasii,g rapidly. However, since the ADVs and TBVs were progranned to open and close on the primary temperature, this would maintain the primary tem-perature at the set point, i.e., 552 K. This also would hold the secondary temperature at this temperature (or slightly lower ~ 550"A) and the SG pres-sure at the saturation pressure (~60 bar) corresponding to this temperature. Meantime, the primary pressure would slowly start increasing due to the pres-surizer heaters. This state would continue until both SGs became empty or nearly empty. At this point, there would be no cooling of the primary system. The primary temperature would start increasing due to the decay heat and the primary pressure would increase faster until it reached the PORV set point. The secondary pressure would start decreasing since the ADV/TBVs were still open due to high primary temperature while SGs became empty. This decreasing pressure would cause the SGIS which would isolate the TBVs. However, since the ADVs would still remain open, the secondary pressure would continue de-creasing. The calculated results generally matched the expected trend as described above (Figures 17 and 18). As programmed in the calculation, the opening of ADV depended on the primary temperature only. However, there was a question if it should also depend on the secondary pressure. If ADVs were programmed this way, it would have kept the secondary side pressure at the set pressure of the ADVs (pro-bably 60 bar) and its temperature at the saturation temperature (552*K) of this pressure. This, in turn, would have kept the average primary temperature near this temperature. This would be equivalent to controlling the primary temperature. The only time when the secondary side behavior deviated from this course significantly would have been when the secondary side water level was very low and, therefore, the steam generator as a heat sink was lost any-way., Therefore, this change of programming on the control of the ADVs may not have made any significant difference for the rest of the transient. - -.

At 20 minutes after the beginning of the transient, the operator activated the AFW pumps to both SGs. This caused rapid cooling and depressurization of the primary system. Low primary pressure activated the SIAS. Meantime, the low primary temperatve caused the ADVs to close. The secondary pressure started to increase and the secondary temperature rapidly increased to the prima ry temperature. The primary pressure a nd temperature continued de-creasing until HPI flow began. The pressure started increasing while the tem-perature continued its downward trend. At this point, the calculation showed severe asymmetric pressure and temperature between the two steam generators (Figure 19). This was explained by '..le LANL staff as instability caused by condensation which is not considered realistic. However, this may not affect the primary pressure and temperature significantly, since the overall cooling may depend on the total amount of AFW rather than its distribution. In summary, the calculated results appeared to be reasonable except for the asymnetric SG behavior after the AFW was introduced. 2.8 Transient 8: MFW Overfeed to Both SGs This transient was not reviewed, since not enough information was avail-able in the handout. It appears that the temperature drop before, and rise after the MFW was tripped off, were of correct magnitude. 2.9 Transient 9: MFW Ovarfeed to One SG This transient was initiated when a Main Feedwater Regulating Valve (MFRV) failed to close when the turbine triaped. This allowed full feedwater flow to the affected SG until the MFW pumps were tripped. This was a relatively dif-ficult transient to analyze. Considerable heat transfer continued in the af-fected steam generator and, consequently, there was a large temperature l i

I dif ference between the primary and second sides of the system, and the cold and hot legs of the prima ry loops for a substantially long period of time (until the MFW pump tripped at 300 seconds). Initial steep drop of temperature and pressure (Figures 20 and 21) of the 1 System due to the loss of the steam in both SGs through the TBVs and ADVs were within the expected range. The continued drop of system pressure and i temperature due to continued MFW to the affected SG and the sharp increase of i the temperature and pressure af ter the MFW was discontinued at 300 seconds, were also as expected. The timing of the affected SG fill-up (about 130 j second) also matched that of the simple hand calculation based on the MFW flow rate and the initial SG water inventory. However, the continued high pressure l of tre intact SG until 300 seconds despite decreasing temperature was not expected (Figure 22). There was high mass flow in the steam line of the J affected SG between 200 and 300 seconds. However, neither the primary temperature e the secondary pressure appeared to be high enough to open TBV/ADVs during this period. j There was a sharp leveling of the system temperature at around 3200 j seconds, yet it appeared that no major event occLrred at this time. This leveling of temperature appeared to be a little too sudden. Decrease of the l temperature around 4800 seconds was expected since AFW started at this time. However, the cause of the AFAS at this time was not clear because there was no I major mass loss from the steam generator, and the temperature remained high. As mentioned earlier, this was a difficult transient to assess. There are 4 I still several points which need further investigation. I i i I ! a j

1 3. Plan for the Detailed Review Among the thirteen transient calculations performed by LANL, six trans-ients have been se'ected for the detailed review. They are Transient Nos. I and 11 for the HZP transients and 3, 6, 7a and 9 for the HFP transients. All four HZP transients are steam line break accidents. Two of them were i 1-ft2 steam line break and the other two are rull double-ended guillotine l steam line breaks. Transients 1 and 11, representing two dif ferent break sizes, are selected for further review. There are also corresponding RETRAN calculations done by ENSA for BG&E available for these two transients. Re-suits of the TRAC and RETRAN calculations will be compared for these two transients. There are basically three dif ferent categories of HFP transients. These are: (a) steam lini break / valve failure (Transients 3, 4 and 4a), (b) small break LOCA (Transients 7 and 7a) and (c) runaway-feedwater (Transients 6, 8 and 9) transients. Transient 5 was a combination of primary and secondary failures (PORV and ADVs stuck-open). Transient 3 has been selected to represent the steam line break / valve i 2 failure transient. The break size of this transient (1-f t ) is the same as ^ that of Transient 1 of the HZP case and this will allow comparison of these tro transients initiated at two extreme power levels. Transient 7a is selected for the small break LOCA transient, since Trans-ient 7 involved artificial and thus unrealistic blockage of the primary loop. Transients 6 and 9 are selected for the runaway feedwater cases, repre-senting the AFW and MFW overfeed, respectively. It should be noted from Table I that the transients selected for detailed 4 \\ review (Transient Nos. 1, 3, 6, 7a, 9 and 11) do include most of the more severe overcooling transients calculated by LANL for the NRC Calvert Clif fs PTS study. j ' i I

REFERENCES 1. J. G. Koenig, et al., " TRAC Analyses of Potential Overcooling Transients at Calvert Cliffs-1 for PTS Risk Assessment," Los Alamos National Labora-tory, December 1983. I I i 1 I f t - - -

1 l i i 1 TABLE 1 CALVERT CLIFFS PTS TRANSIENTS Inst tal Iquilm nt Transient Plant Initiattnq lailures on Operation Minimum Minimum No. Descriptive Title State twent Deman1 Actions I(K ) P(MPa) Repressurization 2 2 { l. 1-ft steam line break at flot 0% l.0-ft hole in Nne hone 3% 4,8 ygg i standby Power steam line A ?. Full double-ended Hnt 01 Full steam line Auxiliary fec h ter hone 377 3.7 yes quillotine steam line l' owe r break ( AF W) 19

• st isolatect break 2

2 3. I ft steam line break at 100% l.0-ft hole in None None 468 6.0 res full power Power steam line 4 Turbine-trip with turbine. 1001 Turbine trip IBV sticks wide None 530 10,8 y, g bypass valve (IBV) stuck Power open ,i open I 44 Turbine trip with one IBV 100% Turbine trip 1BV & MSIV stick open 500 11.4 y,, s y and one MSIV stuck open Power 8 5. Primary power-operated and 100% PORV trarsfers I ADV orens on demand hone 407 6.0 fe j atmnspheric-dur'p val ve Power to wide cren and sticks ope *t ( ADV) stuck open i ) 6 AFW overfeed af ter AFW 1001 MfW system trips AfW delay for 20 AFW valves 375 6.5 res l response failure Power off m in, opened fully I at 20 min. L 2 7. Small break loss-of-conlaet 100 An 0.07 ft bole Noe None 342 2.6 t;o I accident with blocked Power appears in the hot natural circulation le9 i 7a. Small break 1rCA with no 100; Nr ce hone 440 3.8 No i artificial flow blockage Power I H. Main fee.twater overfeed 1001 Turbine trip 2 MFRVs stf(6 None 4PO 7.0 Ves Power open i 9 Main feedwater overfeed to 100% Turbine trip 1 MFRV sticks Non* 490 6.4 yes nn, Sr. Twr open 2 10 I ft steam line break with Hot O! 1.0 f t hole in hone None 446 3.9 res i ? R(Ps left operattnq Powr steam line 11. Full cineshle-ended guillotine Hot 01 Full steam line M5ivs fail to AFW turned 376 4.5 res d steam line break Power break close off at 8 min, i 1 I

l l s CC-LO FT**2 MSLB FROM HZP SYSTEM PRESSURE l 14 i i i j -2500 b J -2250 14 - s' -2000 2 I' T u-1 - 950 5 3 a 'e I a -=00 m- ~ g - 125 0 I s- ~ -1000 6- ~ -750 \\. t 0 M 2d00 3dOO 4d00 5d00 4800 m 8000 r,m.(s) i Figtre 1 Transient 1 v e f - 71,- i r

PRESSURIZER LEVEL FRAME 120 COMrnNENT 1 ROD 0 NODE 1

1.

• i i

i i i i i i i 8. 7 L 6. oo T ~ l o 5. xUo _.) m 4. _ao 5z O

3. '.-

U 2.

1. :.

O' u. 2$0. dT. dJ. $1. Is!E. ~T7l3. Ida. IThT. Bb. 21u 0. 1 REACTOR TIME (SECONDS) Figure 2 Transient 1 .s; 1 l l, l z

I CC-1.0 F P'2 MSLR FROM HZP DOWNCOMER UOUID TEMPERATURE Livtt ? 575 -500 35o. - 520 y $33 .......... THUA 2 8 - THDA 3 { Q -440 500-T3 2 -400 l Aps. e2 l 3 -340 Y 450 - I -= /sN, ~~ _,,0 400-j - 240 375 0 1000 2000 3000 4000 5000 0000 7000 0000 Tirm ($ i Figure 3 Transient 1 l l l.

CC-10 FP"2 MSLB FROM HZP STEAMUNE PRESSURE INTACT tOOP 7 -1000 -900 8- -800 S- -700 ~ 'E' ?u -600 2 e 4-3 / s U f N [ -500 g* 3-400 300 2- -100 I 0 9000 2000 3000 4000 5000 6000 7000 8000 Time (s) Figure 4 Transient 1 i 74 -

l 70- - 1$0 60- - 12 0 50- -10 0 k. 40- ^ O v -80 o 3 O 30-o 60 C m m a o 5 1 20- -40 to- - -20 0- -0 ~ --20 0 1'11 400 GOO E00 1000 1200 1400 1G02 Time (s) l. Figure 5 Transient 10: Total HPI Flow .I.

e (tvtt 5 SG" -540 l 540-4 ~ - 510 \\ THETA 1

---

-- - - - THETA 2 520-

, -480 2 ( - THETA 3 h E -450 E k 3 500-o '. h L E -J -420 Q- { \\ y l80-0 2 3 . h 390 a tT 3 { U- '3 \\, 460- -360 \\ 440-330 -300 Sb0 1000 15b0 2000 25'00 3N00 3500 4N00 4500 5000 5500 O Time (s) Figure 6 Transient 10: 00wncomer Liquid Temperature g.

Ss 0 - 600 360- ~ SCO 53c-l' !\\ 500-c E i L T - 40c w 470-I V' 4% E 440- \\ g g -3D0 g d 2 4to g 9 h a b 380-r- -200 350-- 320-10 0 290 0 1000 2000 3000 4000 5000 6000 7000 8000 TiWE (s) Figure 7 Downcomer Liquid Temperature During Transient 2 i l

590 - -600 i i i i 560-530- } ~ 500-I C S L w -400 to 470-Q h 5 E m Q-440-2 7 d -300 g 410-g o a 3 0 380-c _x._....j = = - : 1 = ww. - - - =w_-. ; - m ] a -200 350-320- -10 0 290-i i i 0 500 1000 1500 2000 7500 3000 3500 TIME (s) Figure 8 Transient 11: Full Steam Line Break with Stuck Open MSIVS Downcomer Liquid Temperature _ _ - _ _ - _ _ _ _ _ _ - _ - _ _ _ _ _ _ _ -

7000000-e i -900 6000000-l i -750 5000000-4000000 LOOPB - -600 9 i t ** * * * +

• LOOP A k

6 v w 450 l*J cr @3000000 ~ g 0 M l W 2000000- . -300 1000000 -. -15 0 0- - -0 -1000000-0 25 50 75 10 0 12 5 150 175 200 225 250 TIME (s) Figure 9 Transient 11: Full Steam Line Break with Stuck Open MSIVS Steamline Pressure..

LEvtl 5 570-4 i -555 560-THETA 1 -540 .......... THETA 2 550-8 - THETA 3 -525 F v e i E y 540 - 510 e CL E E 530 . -495 d2 .T y .'5 s* E

• J

-480 J 520-k,' - d. k&h l .sss 510 - -450 500. 0 250 500 750 1000 1250 1500 1750 2000 2250 2500 Time (s) Figure 10 Transient 4A: Downcomer Liquid Temperature l

16 i -2300 15.5 - -2200 15-14.5 - - 210 0 0 14 - M[ -2000 E 13.5 - E 8 8 E -1900 i 13-12.5 - -1800 l 12 - -1700 11.5 - l 11- -1600 i i i i i 0 250 500 750 1000 1250 1500 1750 2000 2250 2500 Time (s) Figure 11 Transient 4A: System Pressure..

(LEVEL 7) E00-THETA CELL 1 -600 575- ~~ ~"- THETA h 2 THETA CELL 3 -560 550- - - - - - - THETA CLLL 4 e. . D THETA CELL 5 p g N 3 v 525-s.N -480 IEa s h. Db E 500-440 b \\o e =5 475- '#00 9 8 m s J -360 },% 450 I' -320 425- / -280 400- \\ 0 1000 2000 3000 1000 5000 0000 7000 8000 TIME (s) Figure 12 Calvert Cliffs Unit 1 - PTS Transient 7A: Reactor Vessel D0wncomer

20 i -2800 18- -2400 16-14 -2000 9 9 R 12 - m } -1600 O u u 10-01 m W 120c d 8-tr a 6' ~ 800 \\ 'N 4._ 400 2-l 1 0 -0 0 1000 2000 3000 4000 5000 6000 7000 8000 TIME (s) Figure 13 Calvert Cliffs Unit 1 - PTS Transient 7A: Primary Pressure.

10-i -1400 STEAM GEN.11 -noo 3 - ----- - STEAM GEN.12 -1000 Q 6-2,- - ~ - "_ .s00 S W W a a 't S g 600 ji 4 Q. Q. 400 2- -200 o-0 0 1000 2000 3003 4000 5000 6000 7000 80C0 TIME (s) Figure 14 Calvert Cliffs Unit 1 - PTS Transient 7A: Secondary Pressures s g 575- -560 l l 550- -520 THETA 1

---

THETA 2

- _480 525-g m - THETA 3 C E E 500- - -440 E 3 2 e ..i o u ,h% _ -400 D W 475 - a E E 32 %g'- } -360 ~ -320 425-l - v, -280 T 400- - 24 0 l l 375-0 1000 2000 3000 4000 5000 0000 7000 8000 Time (s) l l Figure 15 Transient 5: Downcomer Liquid Temperature l l 1 l l l _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

f 18 - i -2500 16 - -2250 14 - - -2000 o Rj g2_ - -1750 m v v 2 i 2 0 0 -1500 E 10- -1250 B- -1000 6-I -750 4-0 1000 2000 3000 4000 5000 GC00 7000 8000 Time (s) Figure 16 Transient 5: System Pressure. _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

(LEVEL 7) 600-600 THETA CELL 1 - - +-- THETA CELL 2 550~ -+--- THETA CELL 3 -500 - 1HETA CELL 4 g k -- THETA CELL 5 .00-p a w -400 g e Q" 450-V w W sw g ,9 a o ~ O D D 400-- o Og* -200 350 10 0 300 0 1000 2000 3000 4000 $000 6000 /000 8000 TIME (s) Figure 17 Calvert Cliffs Unit 1 - PTS Transient 6: Reactor Vessel Downcomer.

20 i 6 2800 18 - -2400 16-I i g I 14 - 2000 9 9 ct 12 --

• G e

-1600 O u a 10 -- g In M . -1200 g. 6- -800 4- -400 2-0 -0 0 1000 2000 3000 4000 6000 6000 7000 8000 TIME (s) Figure 18 Calvert Cliffs Unit 1 - PTS Transient 6: Primary Pressure i

7 -1000 ( STEAM GEN.11

---

STEAM GEN.12

-900 6 -- -800 9 y-5- T re, -700 S v l '. w a .l '. Ta v3 m u) w l -600 4-l m a f -500 8 3- \\ -400 \\ l.;. ! 2- -300 6 i 0 200 400 BOO 800 1000 1200 1400 1600 TIME (s) Figure 19 Calvert Cliffs Unit 1 - PTS Transient 6: Secondary Pressures 1 l l-

M i i - 560 540- -540 F THETA 1 550 - -520 m - - - - THETA 3 C 6 540 - 2 j -500 550-g g I -480 E e g_ 4 u ','5 Sto - -460 g s 500- ... uo k 40- - 420 400 O S00 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 6000 Tme(s) Figure 20 Transient 9: Downcomer Liquid Temperature

'8 i i i i i i i -2500 hgk gh - 16 - i 14 - - -2000 r T.Pi -\\ - -1750 i z"- 12 - a ) e e ~ ~" io _ E -1250 g. -m 4- -750 4 0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 6000 Tme(s') Figure 21 Transient 9: System Pressure ( l-

IN1ACI LOOP 6.35 -970 l 6.30 -910 6.25-9 T I'. Q-6.20-1 a ...g. 2 2 8 8 f 6.15-i -890 i 6.10 - -880 6.05-i 6, a 200 400 con aco 1000 1200 1400 1600 l Time (s) Figure 22 Transient 9: Steamline Pressure l I , l

I 1 APPEND 1x 0 i

l r BROOKHAVEN NATIONAL LABORATORY 4 MEMORANDUM DATE: January 12, 1984 To: P. Saha ) FROM : U. S. Rohatgi and J. Jo i j Extrapolation of Existing PTS Calculations With or Without

SUBJECT:

Changes in Boundary Conditions l NRC has requested LANL and INEL to compute primary side response to var-ious hypothetical accident scenerios using the advanced codes such as RELAP5 and TRAC-PF1. However, there are many probable event sequences and only a few l of them could be considered for detail calculations. Tnese calculations will provide downcomer liquid temperatures and wall heat transfer coefficients 1 which will be used in stress analysis code. However, the other possible tran-sients will not be calculated by advanced codes, but the downcomer fluid tem-perature, wall heat transfer coefficient and primary pressure will be estimat-ed using the results of other transients with similar features and simpli-fied balance equations. This memorandum describes some approaches to system-atically extrapolate the calculation from any time in the transient. i 1. Multi Volume Approach: In this approach primary and secondary sides are modeled as separate vol-umes with heat transfer in the steam generator. The heat transfer calcula-i 4 tion in the steam generator takes into account the liquid level. Primary Side Mass balance: dM (I) dt HPI

• N ~ BR c

Energy balance: h(Mh + M,h ) = Qd*O+Omis

• N pp p

p HPI HPI + h -W h h cc BR BR (2) n ~ OSGI l._

where the M, M, h are primary side total mass, system metal equiva-g g p lent fluid bass and average enthalpy, and W HPI. W e WB R Od

• 6 c

and Q gt are HPI, charging and break mass flow rates, 6;cSy* heat, Omis 3 ptrip power, energy input through spray, and heat transfer to secondary side of the steam generator, respectively. Secondary side Each steam generator will have the following set of balance equations: Mass balance: -W Msi " Wfwi sti (3) t Energy balance: hsi) " Nfwi fwj - Wsti gh j + Q Gi (4) (Msi h S Here, M, h3 are secondary side total mass and average enthalpy, and s W7w, h a nd Wst, are feed water flow rate, feed water enthalpy and steamNo,wrate,respectively. Ms1 = Al pgj + A (Lt-l)P j i gj si si = ALjp h j + A(Lt - l )Dgj gi (6) M h jg i h QSGi = P(L H j + (Lt - L )Hgj)(Tp-Tsi) (7) jg i Tp=Tp (h ) (3) p hj=hg (T j) (9) g s hj=hg (Tsi) (10) g gj g (Tsi) (11) P "p gj g (Tsi) (12) o "o Rare A, L, L, P, T, Tsi, Hgjand Hgj are flow area, liquid level, total j t p height, perimeter, primary side temperature, secondary side temperature, i liquid and vapor region heat transfer coefficients, respectively. In general there are two stram generators and one may be blowing down. The three volumes in this situation are; primary side and two secondary side for two steam gen-erato rs. So there are twenty-one variables; M, h, T for the primary p p o side and M. h, h, h, Q3g, L, T P t* Ag s for each steam generator. There 3 g g s are also twenty-one equations, i.e., (1), (2) and (8), and two sets consisting of Equations (3) to (7) and (9) to (12). l This system of equations can be further simplified, l ! 1

OSGi = Hoj (Tp - Tsi)Li (13) where 0 'I 'U H = Li,o(T -Tsi)o of p Here H, QSG0' l and (T -Tsi) are heat transfer coefficient, steam generator g o p heat transer rate, liquid level and tenperature difference in steam generator at the time from where the extrapolation will begin. It has been asstsned that most of the heat transfer will be in the liquid phase and the net heat trans-fer will be proportional to the liquid level. In this model, H represents o an average heat transfer coefficient and assumed to be constant during the time period of the extrapolation. This approach implies that as the steam generator secondary side fills up the heat transfer will improve. The other important variable is the break flow rate and it will be estimated from the primary pressure which will be computed separately from the pressurizer analysis. 2. Simple Approach Previous approach will require detailed analysis of each steam generator and in many instances, that is not required either for extrapolation or check-ing the detail calculations. This simple approach will apply to situations where the heat transfer between the primary and secondary side is small and variation in fluid temperature throughout the system is small. The system, consisting of primary side and all the secondary sides can be modeled as single volume. In computing system energy the contribution due to secondary side steam energy can be neglected. The metal part of the system stores a significant amount of energy and it is accounted for by estimating equivalent liquid mass and adding it to the system fluid nass. The balance equations are: h(M)=WHPI + Wc-WBR (14) p (Msi) " Wfwi ~ Wsti (15) h{(M+M+IM s1)h) = Q +0 +0 mis +W h HPI HPI d p p m +Whce~W h h BR BR + IWfwi fwj h - IWsti sti where h is the average fluid enthalpy for the system. The other variables are the same as described in the previous approach. For the system with two steam generators, there are 4 equations in four unknowns which are, M, Ms1' p Ms2 and h.. _ _ _ _ _ - _ _ _ _ _ _

3. Boundary Conditions Both approaches described so far require high pressure injection (HPI), charging and feed water conditions. These are generally known as they are input to the system and in some instances are function of the pressure of the volume in which they are introduced. The flow through the breaks and valves are also functions of the conditions such as pressure and void fraction of the volume in which they are located. This will require modeling pressurizer and stean generator secondary side separately to estimate the pressures. The steam generator secondary side model has been described in this fomulation. The secondry side pressure is assumed to be saturation pressure corresponding to its temperature. In the cases where saturation pressure exceeds the TBV pressure setting the valve will open and steam will be released. This steam flow will also depend upon the secondary side pressure. 4. Pressurizer Model This coaponent controls the pressure in the primary side through sprays and heaters. However, during the transient there is flow through the surge line which will affect the pressurizer pressure. The model described here will predict primary side pressure for cases where the primary side has no vapor except in the pressurizer. The surge line flow will depend upon the contraction or expansion of the liquid in the remaining primary side. The balance equations are as follows: hML*Wsr + Wsp - (17) r d-M =-WBR + r (18) y dt d - Lt It (M h ) = Qh+W h h sr sr+Wsp sp - rh (19) y d-- ( M h ) = - W h BR y + rh (20) y y v dt WBR = f(P) (21) s t = U + Sv = pressurizer volume (22) V t 1 v p =p (P,h ) (23) y y y hr=hL or h, depending upon surge line flow direction. (24) s p h

The unknowns of the model are M, h, M, h, WBR, r, Wsr, hsr, p and p L L y y which are liquid inventory and enthalpy, vapor inventory and enthalpy, break flow, vapor generation rate, surge line flow and enthalpy, pressure and vapor density. Howaver, there are only eight equations for ten unknowns and two more equations are needed. The primary side liquid has much larger volume compared to the liquid volume in the pressurizer and small expansion or con-traction of primary loop liquid will have significant change in the liquid inventory of the pressurizer. Wsr " Y dp /dt (25) p g (P, T ) o =o p g are primary side liquid volume, liquid density and temper 3turb, and T Here V, p p This still leaves this formulation shor; of one equation which will come fran assunption on the processes of vapor generation. In the first limiting case it can be assumed that there is no vapor generation and it is a frozen case, which implies that liquid could become superheated and vapor could become subcooled depending upon the direction of surge line flow. In this case r=0 (26) This completes the fomulation. However, it can be further simplified if the vapor expansion or contraction can be represented by some polytropic processes P/pk = constant (27) This simplication will replace Equations (20) and (23) by Equation (27) and also vapor enthalpy h will not be needed. The variable k is 1.0 for iso-themal process and is 1.33 for isentropic process. The second limiting case is where the liquid and vapor are both saturated and any addition of mass or energy will change the pressure along the satura-tion conditions. The system of equations is as follows: L=Wsr + Wsp - r (28) gv = _ wBR + r (29)

\\ dh h sr (h g h ) + Wsp (h -h ) - M s f sp f L dp r= (30) hfg WBR " WBR(P) (31) of = pf(p) (32) g " o (p) (33) o g hr = h (P) (34) f hfg = hfg(P) (35) Vt" = pressurizer volume (36) V f g hsr = hf or h, depending upon the direction of flow. (37) p d.2f Wsr _ Vp dt (38) So there are eleven equations and eleven unknowns which are, M, M, r, L y WBR Pfpg, h, hfg, hsr* Wsr f and P. These two cases will provide limiting pressure histories for the primary side. In case of flo.v into the pressurizer, the frozen case will predict higher pressure than the equilibrium case, while for flow out of the pressuri-zer the frozen case will predict lower pressure as flashing of pressurizer liquid in equilibrium case will try to maintain the pressure. Equations (26) and (38) restrict this model to transients which have intact primary side with only vapor region in the pressurizer. 5. Solution Pro:edure All the differential equations are linear and so simple Euler types of integration can be used. Most of the equations are essentially equation of state and can be replaced by steam table. As most of the changes during PTS transients occur over the long tenn, these equations can be finite differenc-ed. This will make it possible to use hand calculations to estimate i temperature and pressure history on the prinary side, af cc: C. Yuelys-Miksis R. J. Cerbone _ _ _ _ _ _ _ _ _ _ _ _.

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U.S. NUCLE AR REGUL ATORY COMMISSION NUREG/CR-4251 BIBLIOGRAPHIC DATA SHEET BNL-NUREG-51P/ TITLE AND SuuTITLE (Aan Vo/ume No,1aooroor, ares 2.(ten,ef,iej f Review of TRAC Calculations for Calvert Cliffs PTS Study / J RE CIPIE N T'S CESSION NO Au r HO H <Si $. D AT E HEPC[T COYPLE TE D "oN'" l'E^" J. H. Jo and U. Rohatgi fsarch 1985 Pt H F OHMINO OHG ANi/ A Ti. N AME AND MAILING ADONESS IInclude I0 Code / DATE H[pOHT ISSUE D Department of Nuclea Energy

• N'[

l^" Brookhaven National L ratory 6 "[ * *" * ' Upton, Long Island, New ork 11973 .,p. ~., 1 SPONSOHING OHG ant /ATION NAYL, ) M AILING ADOHESS Itacr oc I,o Cooel u Division of Accident Evalua n / Office of Nuclear Regulatory search [1 FIN NO. ) U.S. Nuclear Regulatory Connis n Washington, DC 20555 f A3266 lYvt CF Htv0HT PtnoooCO nt o tincru <ve caress Technical Report 3 bVPPLE VE N T AHY NO f t S _[ 1,3 tt ra, gspra f w I k 6 AB5TH ACT (200

• n<ns ne seis; Six selected transient calculations out 6. thirtee performed by LANL using the TRAC-PF1 code for the USNRC PTS study of the Calve Cliff Nuclear Power Plant have been re-viewed in depth at BNL. Simple hand calcui ions ased on the mass arid energy balances have been performed to predict the temperatu a

pressure of the reactor system, and the results have been compared with those of Comparison was also made between transients, which were performed by the TRAC and RETRAN calculations for two of th, appear to be reasonable based on the ENSA. In general, the results calculated by T " comparisoit with RETRAN and hand calculations. t k ! / AL Y WOHUS AND DOCUint NT AN ALYSIS 17a CE SC HIP T OHS Pressurized Thermal Shock Calvert Cliffs Review TRAC l i /s ice N Ts F iE HS OPE N E N DE D TE RMS 18 AV AIL ABILITY ST A TE ME NT 19 SE CURI TY C(ASS (Ta.s reporrt 21 NO OF PAGES n assihed Unlimited 10 SECUHi T Y CL A% f fho pe# // I' H ' C t 5 3 i MC 6 Onu 33% ses su}}