ML20137K404
| ML20137K404 | |
| Person / Time | |
|---|---|
| Site: | North Anna  |
| Issue date: | 01/17/1986 |
| From: | VIRGINIA POWER (VIRGINIA ELECTRIC & POWER CO.) |
| To: | |
| Shared Package | |
| ML20137K386 | List: |
| References | |
| NUDOCS 8601240053 | |
| Download: ML20137K404 (69) | |
Text
r ATTACHMENT 1 SAFETY EVALUATION FOR 75 STEAM GENERATOR PLUGGING LEVEL IN NORTH ANNA UNIT NO. 1 8601240053 860117 PDR P ADOCK 05000338 PDR
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r SAFETY EVALUATION FOR 7% STEAM GENERATOR PLUGGING LEVEL
1.1 INTRODUCTION
Recent eddy current inspections of North Anna Unit No. I stean generator tubes indicated a need to plug additional tubes beyond the Cycle 5 level of 3.17%.
The plugging level was increased to 3.89% during the North Anna Unit 1 Cycle 5 to Cycle 6 outage. Since it is possible that future eddy current inspections will require that the plugging level be increased to a level beyond the current 5% limit approved in Reference 1, an assessment of the impact on LOCA and non-LOCA analyses at higher levels of tube plugging (up to 7%) is being provided here.
1.2 NON-LOCA ANALYSES Since sufficient tube plugging can reduce primary system flow, result in more severe pump coastdown characteristics, and reduce the primary system volume, the impact of tube plugging on non-LOCA transient analyses was evaluated (Reference 2).
The Reference 2 evaluation provided a figure with a conservative estimate of actual RCS flowrate versus steam generator tube plugging level for the North Anna Power Station. This figure shows that at 7?. tube plugging the primary systen flowrate would still be more than 2.5% greater than the current North Anna thermal design flow. Thus the current licensing analyses, in which flowrate is an important concern (DNB limited events), remain valid at 7f.
plugging.
Seven percent steam generator tube plugging results in a three percent change in the total loop resistance. Reference 2 indicated that this change would have a minimal impact on the pump coastdown curve.
Reference 2 determined that 7% steam generator tube plugging would reduce the RCS volume by 2 percent and this would reduce the dilution times by 2 percent.
However, during dilution events it was determined that sufficient time (approximately 60 minutes) remained to criticality to allow operator action.
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7 1.3 LOCA ANALYSIS The current docketed LOCA analysis (Reference 1) is based on an assumed plugging level of 5%. A new LOCA analysis, which was recently subnitted by Reference 3 in support of the North Anna core uprating, assumes a plugging level of 7%. As discussed below, this analysis, which is provided as Attachment 2, bounds the cu: rent licensed design conditions and is being submitted to justify operation at the current licensed conditions with 7%
stean generator tube plugging, Table 1 provides a summary of~ differences in plant conditions between the current licensed analysis and the Attachment 2 analysis. The primary conservatisms in the new LOCA analysis are the higher reactor power and the lower reactor coolant flow, both of which tend to increase the peak clad temperatere during a LOCA. The 50 psia reduction in steam pressure and the one degree difference in vessel average coolant temperature will have a minor impact on peak clad temperature. Previous experience shows that the effects of the power and RCS flow assumed will dominate the other differences to produce a conservative peak clad temperature. Therefore, operation at the current licensed reactor thermal power of 2775 MWt will be bounded by the results of the uprated analysis when operating at the FQ limit (2.15) 1etermined from the new analysis. The approach of using a LOCA-ECCS analysis performed at higher reactor power to support lower power operation has previously been used (See Reference 4).
Three aspects must be considered when evaluating the effect of steam generator tube plugging level on small break LOCA transients. They are reduced heat transfer area, the increased initial temperature difference between the primary and secondary side and the countercurrent flow limit (CCFL). Since only a small portion of the steam generator heat transfer area is required to provide an effective heat sink during a small break transient, plugging some steam generator tubes will not affect small break LOCA transients in view of the available heat transfer area. The increased temperature difference between the prima ry and secondary side disappears Wediately after the break when the secondary side pressure reaches the steam generator safety valve setpoint and therefore, has no impact on the 84R629NPW165 3 o
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i transients. The CCFL has been shown to have insignificant impact on the peak clad temperature for small levels of steam generator tube plugging (up to approximately 20%). Since the plugging levels will not exceed 7%, there will be no impact on small break LOCA and the currently approved small break LOCA analysis (Section 15.3.1 of the North Anna UFSAR) remains bounding.
1.4 TECHNICAL SPECIFICATIONS CHANGES The necessary Technical Specification changes for North Anna Unit No. I associated with reducing the F0 limit from the currently licensed value of 2.20 to the value of 2.15 are provided in Attachment 3. The affected sections of the Technical Specifications for Unit No. I are Sections 3.2.2 and 3.2.6 and Figure 3.2-2.
In conjunction with the lower FO limit, Attachment 4 provides a new Core Surveillance Report for North Anna 1 Cycle 6.
1.5 CONCLUSION
S The proposed increase in stean generator tube plugging and corresponding reduction in the FQ limit has been reviewed against the criteria of 10 CFR 50.59 and does not involve any unreviewed safety questions. The specific bases for this detennination are as follows:
- 1. The revised specifications will not increase the probability of occurrence or consequences of any malfunction or accident previously addressed. The re-analyzed large break LOCA analysis, which is attached, shows that operation under the revised specifications would not result in any increase in accident consequences. The analysis assumptions for the remainder of the UFSAR Chapter 15 transient analyses have not changed and they remain bounding.
- 2. No new accident types or equipment malfunction scenarios will be introduced as a result of operating in accordance with the revised specifications.
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- 3. The margin of safety, as defined in the basis for the affected Technical Specifications, is not reduced. Operation at the lower F0 limit will not reduce the margin to the LOCA acceptance limits.
The proposed changes do not pose a significant hazards consideration as defined in 10 CFR 50.92. This conclusion is based on Example vi of those types of license amendments that are considered unlikely to invnive significant hazards considerations. Example vi, which was published in the Federal Register, Vol. 48, No. 67, April 6,1983, p.14870, " Standards for Determining Whether License Amendments Involve No Significant Haza rds Considerations, Interim Final Report" partially states, "A change which either may result in some increase to the probability or consequences of a previously analyzed accident or may reduce in some way a safety margin, but where the results of the change are clearly within all acceptable criteria with respect to the system or component specified in the Standard Review Plan". This change merely adjusts the FQ limit to reflect the results of the LOCA analysis performed with an increased steam generator plugging level. The results of the new LOCA analysis are within all acceptable criteria of 10 CFR 50.46 and the appropriate safety margins are maintained. As discussed above, the proposed Technical Specification amendment would not:
- 1. Involve a significant increase in the probability or consequences of an accident previously evaluated, or 2 Create the possibility of a new or different kind of accident previously evaluated, or
- 3. Involve a significant reduction in a margin of safety.
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1.6 REFERENCES
- 1. W. L. Stewart (Virginia Power) to H. R. Denton (NRC) Serial No. 726 December 30, 1982.
- 2. R. H. Leasburg (Virginia Power) to H. R. Denton (NRC) Serial No. 080, February 12, 1982.
Serial No. AEP:NRC:0745M, Docket No. 50-315, License No. DPR-58, dated August 23, 1984.
i 1
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TABLE 1-1 COWARISON OF REACTOR COOLANT SYSTEM PARANTERS Design Conditions Current Uprated License Power -
Analysis Analysis (Reference 1) (Reference 3)
Reactor Power, MWt 2775 2898 Total Reactor Flow, GPM 285,000 278,400 Reactor Coolant Temperature. *F 587.8 586.8 (Vessel Average)
Stean Generator Steam Pressure, psia 900 850 84R629NPW165 7
ATTACHMENT 2 D
LOCA-ECCS SAFETY EVALUATION FOR NORTH ANNA UNIT NO. 1 l
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2.1 INTRODUCTION
A re-analysis of the Emergency Core Cooling System (ECCS) performance for the postulated large-break LOCA has been performed in compliance with Appendix K to 10 CFR 50. The results of this re-analysis are presented here, and are in compliance with 10 CFR 50.46, " Acceptance Criteria for Emergency Core Cooling Systems for Light Water Reactors." This analysis was performed with the NRC-approved 1981 model with BART version of the Westinghouse LOCA-ECCS evaluation model (Ref. I and 12). The analytical techniques used are in full compliance with 10 CFR 50, Appendix K.
As required by Appendix K of 10 CFR 50, certain conservative assumptions were made for the LOCA-ECCS analysis. The assumptions pertain to the conditions of the reactor and associated safety system equipment at the time that the LOCA is assumed to occur, and inc,lude such items as the core peaking factors, the containment pressure, and the performance of the Emergency Core Cooling System. All assumptions and initial operating conditions used in this reanalysis were the same as those used in the previous LOCA-ECCS analysis (Ref. 2), with the following exceptions: ,
- 1. The core power level has been increased to 2898 MWt.
- 2. Seven percent steam generator tube plugging was assumed.
- 3. A thermal design flow of 92,800 gpm per loop was used.
- 4. The 1981 LOCA-ECCS evaluation model with BART (Ref.10 and 11) was used to perfo m this analysis.
With the above changes incorporated into the analysis, it was found that the assumed heat flux hot-channel factor could be 2.15 and still ensure compliance with the 10 CFR 50.46 acceptance criteria.
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2.2 ACCIDENT DESCRIPTION A LOCA is the result of a rupture of the reactor coolant system (RCS) piping or of any line connected to the system. The system baundaries considered in the LOCA analysis are defined in Section 3.6 of the North Anna FSAR.
Sensitivity studies (Ref. 3) have indicated that a double-ended cold-leg guillotine (DECLG) pipe break is limiting. Should a DECLG break occur, rapid
~
depressurization of the reactor coolant system occurs. The reactor trip signal subsequently occurs when the pressurizer low-pressure trip setpoint is reached. A safety injection system (SIS) si,gnal is actuated when the appropriate setpoint is reached and the high-head safety injection pumps are activated. The actuation and subsequent activation of the Emergency Core Cooling System, which occurs with the SIS signal, assumes the most limiting single-failure event. These countermeasures will limit the consequences of the accident in two ways:
(
- 1. Reactor trip and borated water injection complement void formation in causing rapid reduction of power to a residual level corresponding to fission product decay heat. Nc credit is taken in the analysis for the insertion of control rods to shut down the reactor.
- 2. Injection of borated water provides heat transfer from the core and prevents excessive clad temperature.
Before the break occurs, the unit is in an equilibrium condition, i.e., the heat generated in the core is being removed via the secondary system. During blowdown, heat from decay, hot internals, and the vessel continue to be transferred to the reactor coolant system. At the beginning of the blowdown phase, the entire reactor coolant system contains subcooled liquid that
} transfers heat from the core by forced convection with some fully developed 8
nucleate boiling. After the break develops, the time to DNB is calculated, consistent with Appendix K of 10 CFR 50. Therefore, the core heat transfer is based on local conditions, with transition boiling and forced convection to steam as the major heat transfer mechanisms. During the refill period, it is assumed that rod-to-rod radiation is the only core heat transfer mechanism.
The heat transfer between the rea.: tor coolant system and the secondary system i
84R629NPW165 10 s
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f may be in either direction, depending on the relative temperatures. For the case of continued heat addition to the secondary side, secondary-side pressure increases and the main safety valves may actuate to reduce the pressure.
Makeup to the secondary side is a'utomatically provided by the auxiliary feedwater system. Coincident with the safety injection signal, normal feedwater flow is stopped by closing the main feedwater control valves and tripping the main feedwat h pumps. Emergency feedwater flow is initiated by starting the auxiliary feedwater pumps. The secondary-side flow aids in the reduction of RC9. pressure. When the reactor coolant system depressurizes to 600 psia, the accumulators begin to inject borated water into the reactor coolant loops. The conservative assumption is then made that injected accumulator water bypasses the core and goes out through the break until the termination of bypass. This conservatism is again consistent with Appendix K ;
of 10 CFR 50. In addition, the reactor coolant pumps are assumed to be tripped at the initiation of the accident, and effects of pumps coastdown are included in the~ blowdown analysis.
The water injected by the accumulators cools the core, and subsequent operation of the low-head safety injection pumps supplies water for long-term cooling. When the _ reactor water storage tank. (RWST) is nearly empty, long-term cooling of the core is accomplished by' switching'to the recirculating mode of core cooling, in which the spilled borated water is drawn from the containment sump by the low-head safety injection pumps and "
returned to the reactor vessel.
The containment spray system and the recirculation spray system operate to return the containment environment to subatmospheric pressure.
2.3 ANALYSIS -
The large-break LOCA transient;is divided.,for analytical purposes, into three phases: blowdown, refill, and reflood. :There are three distinct' transients analyzed in each phase, including the thermal-hydraulic transient in the reactor coolant rystem, the pressure and temperature transient within the containment, the fuel clad temperature transient of the_ hottest fuel rod in the core. . Based on these considerations, a system of interrelated computer codes has been developed for the analysis.
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The description of the various aspects of the LOCA analysis methodology is given in WCAP-8339 (Ref. 4). This document describes the major phenomena modeled, the interfaces among the computer codes, and the features of the codes that ensure compliance with 10 CFR 50, Appendix K. The SATAN-VI, C0CO, WREFLOOD, BART, and LOCTA-IV codes, which are used in the LOCA analysis, are described in detail in WCAP-8306 (Ref. 5), WCAP-8326 (Ref. 6), WCAP-8171 (Ref.
7), WCAP-9695 (Ref. 10), WCAP-10062 (Ref. 11), and WCAP-8305 (Ref. 8),
respectively. These codes assess whether sufficient heat transfer geometry and core amenability to cooling are preserved during the time spans applicable .
to the blowdown, refill, and reflood phases of the LOCA. The SATAN-VI computer code analyzes the thermal-hydraulic transient in the reactor coolant system during blowdown, and the COC0 computer code calculates the containment pressure transient during all three phases of the LOCA analysis. Similarly, i the LOCTA-IV computer code is used to compute the thermal transient of the i hottest fuel rod during the three phases.
SATAN-VI is used to determine the RCS pressure, enthalpy, and density, as well as the mass and energy flow rates in the reactor coolant system and steam-generator secondary, as a function of time during the blowdown phase of the LOCA. SATAN-VI also calculates the accumulator mass and pressure and the pipe break mass and energy flow rates that are assumed to be vented to the containment during blowdown. At the end of the blowdown, the mass and energy release rates during blowdown are transferred to the C0C0 code for use in the determination of the containment pressure response during this first phase of the LOCA. Additional SATAN-VI output data from the end of the blowdown, including the core inlet flowrate and enthalpy, the core pressure, and the core power decay transient, are input to the LOCTA-IV code.
With input from the SATAN-VI code, WREFLOOD uses a system thermal-hydraulic model to determine the core flooding rate (i.e., the rate at which coolant enters the bottom of the core), the coolant pressure and temperature, and the quench front height during the refill and reflood phases of the LOCA.
WREFLOOD also calculates the mass and energy flow rates that are assumed to be vented to the containment. Since the mass flowrate to the containment depends upon the core pressure, which is a function of the containment backpressure, the WREFLOOD and C0C0 codes are interactively linked. With the input and 84R629NPW165 12
boundary conditions fron WREFLOOD, tne mechanistic core heat transfer model in BART calculates the fluid and heat transfer conditions in the core during reflood.
LOCTA-IV is used throughout the analysis of the LOCA transient to calculate the fuel and clad temperature of the hottest rod in the core. The input to LOCTA-IV consists of appropriate thermal-hydraulic outputs from SATAN-VI, WREFLOOD and BART, and conservatively selected initial RCS operating conditions. These initial conditions are summarized in Table 2-1 and Figure 2-1.
(The axial power shape of Figure 2-1 assumed for LOCTA-IV is a cosine curve that has been previously verified (Ref. 9) to be the shape that produces the maximum peak clad temperature.)
The C0C0 code, which is also used throughout the LOCA analysis, calculates the containment pressure. Input to C0C0 is obtained from the mass and energy flowrates assumed to be vented to the containment, as calculated by the SATAN-VI and WREFLOOD codes. In addition, conservatively chosen initial containment conditions and an assumed mode of operation for the containment cooling systen are input to C0CO. These initial containment conditions and assumed modes of operation are provided in Table 2-2.
2.4 RESULTS Tables 2-1 and 2-2, and Figure 2-1 present the initial conditions and modes of operation that were assumed in the analysis. Table 2-3 presents the time sequence of events, and Table 2-4 presents the results for the double-ended cold-leg guillotine break for the C = 0.4, 0.6, and 0.8 discharge D
coefficients. The double-ended cold-leg guillotine break has been determined to be the liniting break size and location based on the sensitivity studies reported in Reference 3. The analysis resulted in a limiting peak clad temperature of 2160.6*F for the C = 0.4 case, a maximum local cladding D
oxidation level of 6.69%, and a total core metal-water reaction of less than 0.3%. The detailed results of the LOCA re-analysis are provided in Tables 2-3 through 2-6 and Figures 2-2a through 2-18c. The figures show the following:
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- 1. Peaking Factor vs. Core Height - Figure 2-1 shows the cosine power shape used in the analysis.
- 2. Mass Velocity - Figures 2-2a through 2-2c show the mass velocity at the clad burst and hot-spot locations on the hottest fuel rod for the discharge coefficient used.
- 3. Heat Transfer Coefficient - Figures 2-3a through 2-3c show the heat transfer coefficient at the clad burst and hot-spot locations on the hottest rod for the discharge coefficient used. The values of heat transfer coefficient that are shown were calculated by the LOCTA-IV code based on equations for heat transfer in the nucleate boiling, transition boiling, film boiling, and steam cooling regimes.
- 4. Core Pressure - Figures 2-4a through 2-4c show the calculated pressure in the core for the discharge coefficient used.
- 5. Break Flowrate - Figures 2-Sa through 2-Sc show the calculated flowrate out of the break for the discharge coefficient used. The flowrate out of the break is plotted as the sum of flow at both the pressure vessel end and the reactor coolant pump end of the guillotine reak.
- 6. Core Pressure Drop - Figures 2-6a through 2-6c show the calculated core pressure drop for the discharge coefficient used. The core pressure drop is interpreted as the pressure immediately before entering the core inlet to the pressure just outside the core outlet.
- 7. Peak Clad Temperature - Figures 2-7a through 2-7c show the calculated hot-spot clad temperature transient and the clad temperature transient at the burst location for the discharge coefficient used. The peak clad temperature for the limiting discharge coefficient of 0.4 is 2160.6 F at the 6.75 ft elevation in the core.
- 8. Fluid Temperature - Figures 2-8a through 2-8c show the calculated fluid temperature for the hot spot and burst locations for the discharge coefficient used.
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- 9. Core Flow - Figures 2-9a through 2-9c show the calculated core flow, both top and bottom, for the discharge coefficient used.
- 10. Reflood Transient - Figures 2-10: through 2-10c show the reactor pressure vessel downcomer and core water levels for the discharge coefficient used. Figures 2-11a through 2-11c show the core inlet velocity for the discharge coefficient used.
- 11. Accumulator Flow - Figures 2-12a through 2-12c show the calculated flow for the dischargt coefficient used. The accumulator delivery during blowdown is discarded until the end of bypass is calculated.
Accumulator flow, however, is established in the refill-reflood calculations. The accumulator flow assumed is the sum of that injected in the intact cold legs.
- 12. Pumped ECCS Flow (Reflood) - Figures 2-13a through 2-13c show the calculated flow of the emergency core cooling system for the discharge coefficient used.
- 13. Containment Pressure - Figures 2-14a through 2-14c show the calculated pressure transients for the discharge coefficient used. The analysis of this pressure transient is based on the data given in Tables 2-2, 2-5, and 2-6.
- 14. Core Power Transient - Figures 2-15a through 2-15c show the core power transient calculated by the SATAN-VI code for the discharge coefficient used.
- 15. Break Energy Release - Figure 2-16 shows the break energy released to the containment for the limiting discharge coefficient of 0.4.
- 16. Containment Wall Heat Transfer - Figure 2-17 shows the containment wall heat transfer coefficient for the limiting discharge coefficient of 0.4.
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- 17. Fluid Quality - Figures 2-18a through 2-18c show the fluid quality at the clad burst and hot-spot locations (location of maximum clad temperature) on the hottest fuel rod (hot rod) for the limiting breaks.
2.5 CONCLUSION
S For breaks up to and including the double-ended rupture of a reactor coolant pipe, and for the operating conditions specified in Tables 2-1 and 2-2, the emergency core cooling system will meet the acceptance criteria as presented in 10 CFR 50.46, as follows:
- 1. The calculated peak fuel rod clad temperature is below the requirement of 2200 F.
- 2. The amount of fuel element cladding that reacts chemically with water or steam does not exceed 1% of the total amount of Zircaloy in the reactor.
- 3. The clad temperature transient is terminated at a time when the core geometry is still ame.iable to cooling. The localfzed cladding oxidation limits of 17% are not exceeded during or arte, quenching.
- 4. The core remains amenable to cooling during and after the break.
- 5. The core temperature is reduced and the long-term decay heat is removed for an extended period of time.
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2.6 REFERENCES
- 1. Letter from J. R. Miller, NRC, to E. P. Rate, Westinghouse, dated December 1, 1981.
- 2. Letter from W. L. Stewart, Vepco, to H. R. Denton, NRC, dated December J0, 1982 (Serial No. 726). .
- 3. R. Salvatori, Westinghouse ECCS Sensitivity Studies, WCAP-8356, July 1974.
- 4. F. M. Bordelon et al., Westinghouse ECCS Evaluation Model - Summary, WCAP-8339, July 1974.
- 5. F. M. Bordelon et al . , SATAN-VI Progran.: Comprehensive Space-Time Dependent Analysis of Loss-of-Coolant, WCAP-8306, June 1974.
- 6. F. M. Bordelon and E. T. Murphy, Containment Pressure Analysis Code (C0CO), WCAP-8326, June 1974.
- 7. R. D. Kelly et al., Calculational Model for Core Reflooding After a Loss-of-Coolant Accident (WREFLOOD Code), WCAP-8171, June 1974.
- 8. F. M. Bordelon et al., LOCTA-IV Program: Loss-of-Coolant Transient Analysis, WCAP-8305, June 1974.
- 9. Letter from C. M. Stallings, Vepco, to E.G. Case, NRC, Serial No. 092, dated February 17, 1978.
- 10. Young, M. Y. et al., BART-A1: A Computer Code for the Best Estimate Analysis of Reflood Transients, WCAP-9695, January.1980.
- 11. Chiot., J. S. et al., Models for PWR Reflood Calculations using the BART Code, WCAP-10062, December 1981.
- 12. Letter from C. O. Thomas, NRC to E. P. Rahe, Westinghouse dated December 21, 1983.
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TABLE 2-1 INITIAL CORE CONDITIONS ASSUMED FOR THE DOUBLE-ENDED COLD-LEG GUILLOTINE BREAK (DECLG)
Calculational Input Core Power (fiWt) 102% of 2898 Peak linear power (kW/ft) 102% of 12.225 Heat flux hot-channel factor (Fq ) 2.15 N
Enthalpy rise hot-channel factor (F 3H) 1.55 3
Accumulator water volume (ft , each) 1025 Reactor vessel upper head temperature equal to T hot Limiting Fuel Region and Cycle Cycle Region Unit 1 All All regions Unit ? All All regions i
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TABLE 2-2 CONTAI MENT DATA 6
Net free volume 1.916 x 10 ft 3 Initial conditions" Pressure 9.6 psia Temperature 90 F RWST temperature 35 F Outside temperature -10*F Spray system a 2
Number of pumps Runout flowrate (peroperating) pump 2000 gpm Time in which spray is effective 59 see Structural heat sinks a Thickness (in.) Area (ft 2), with allowance for uncertainties 6 concrete 8,393 12 cor. crete 62,271 18 concrete 55,365 24 concrete 11,591 27 concrete 9,404 36 concrete 3,636
.375 steel, 54 concrete 22,039
.375 steel, 54 concrete 28,933
.500 steel, 30 concrete 25,673 26.4 concrete, .25 steel, 120 concrete 12,110
.407 stainless steel 10,527
.371 steel 160,328
.882 steel 9,894
.059 steel 60,875 a
See Section 6.3.3.12 of the FSAR for a detailed breakdown of the containment heat sinks and for justification of the other input parameters used to calculate containment pressure.
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TABLE 2-3 TIE SEQUENCE OF EVENTS DECLG DECLG DECLG CD = 0.4 C = 0.6 D CD = 0.8 (sec) (sec) (sec)
Start 0.0 0.0 0.0 Reactor trip 0.64 0.62 0.61 Safety injection signal 2.12 1.69 1.46 Accumulator injection 15.70 11.90 9.76 Pump injection 27.12 26.69 26.46 End of bypass 31.72 26.65 23.82 End of Blowdown 31.72 26.65 23.82 Bottom of core recovery 45.25 39.74 37.03 Accumulator empty 55.94 51.06 48.27 TABLE 2-4 RESULTS FOR DECLG C = 0.4 CD = 0.6 D CD = 0.8 Peak clad temperature, F 2160.6 2013.7 1829.7 Peak clad location, ft 6.75 7.25 7.25 Local-Zr/H2O reaction (max),% 6.69 3.77 1.92 Local Zr/H 'O location, ft 6.0 6.75 7.25 2
Total Zr/H O reaction, % <0.3 <0.3 <0.3 2
Hot-rod burst time, sec 40.70 66.00 73.00 Hot-rod burst location, ft 6.0 6.75 6.775 84R629NPW165 20
TABLE 2-5 REFLOOD MASS AND ENERGY RELEASES DECLGD (C = 0.4)
Total Mass TotalEgergy Time (sec) Flow Rate (lb/sec) Flow Rate (10 Btu /sec) 45.252 0.0 0.0 46.427 0.013 0.000171 56.412 87.30 1.059 72.137 209.96 1.382 91.587 261.33 1.457 113.137 270.53 1.413 137.687 294.51 1.406 200.937 317.58 1.360 TABLE 2-6 BROKEN LOOP ACCUMULATOR FLOW TO CONTAINMENT DECLG (CD = 0.4) a Time (sec) Mass Flow Rate (lbm/sec) 0.00 4010.1 1.01 3622.2 3.01 3105.1 5.01 2762.6 7.01 2510.3 10.01 2226.3 15.01 1898.9 20.01 1674.6 25.01 1516.9 b
29.01 1575.3 a
For energy flowrate, multiply mass flow rate by a constant of 59.60 Btu /lbm.
b For energy flowrate at this time, multiply mass flowrate by 54.09 Btu /lbm.
i 84R629NPW165 21
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