ML20133D226

Safety Analysis for Upper Head Injection Elimination

ML20133D226
Person / Time
Site:	McGuire, Mcguire
Issue date:	09/30/1985
From:	DUKE POWER CO.
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ML20133D223	List: ML20133D219 ML20133D226
References
NUDOCS 8510080364
Download: ML20133D226 (111)

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MCGU1RL NUCLEAR STATION SAFETY ANALYSIS FOR UH1 ELIMINATION l

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15.6.4.1 Identification of Causes and Accident Description

, Acceptance Criteria and Frequency Classification A loss-of-coolant accident (LOCA) is the result of a pipe rupture of the reactor coolant system (RCS) pressure boundary. For the analyses presented here, a major pipe break (large break) is defined as a rupture with a total cross-sectional area equal to or greater than 1.0 f t .2 This event is considered an ANS Condition IV event, a limiting fault. See Section 15.0.1 for a discussion of Condition IV events.

A minor pipe break (small break), as considered in this section, is defined as a rupture of the reactor coolant presgure boundary (Section 5.2) with a total

cross-sectional area less than 1.0 f t in which the normally operating l charging system flow is not sufficient to sustain pressurizer level and pressure. This is considered a Condition III event, an infrequent fault. See Section 15.0.1 for a discussion of Condition !!! events.

The Acceptance Criteria for the loss-of-coolant accident is described in 10 CFR 50.46 as follows:

a. The calculated peak fuel element cladding temperature is below the

requirement of 2200*F.

b. The cladding temperature transient is terminated at a time when the core geometry is still amenable to cooling. The localized cladding oxidation limits of 17% are not exceeded during or after quenching,

c. The amount of hydrogen generated by fuel element cladding that reacts chemically with water or steam does not exceed an amount corresponding to interaction of 1% of the total amount of Zircaloy in the reactor.

i d. The core remains amenable to cooling during and after the break,

e. The core temperature is reduced and decay heat is removed for an j extended period of time, as required by the long lived radioactivity remaining in the core.

These criteria were established to provide significant margin in ECCS performance following a LOCA.

In all cases, small breaks (less than 1.0 ft ) yield results with more 2

margin to the Acceptance Criteria limits than large breaks.

Description of a Laroe Break LOCA Transient Should a major break occur, depressurization of the RCS results in a pressure decrease in the pressurizer. The reactor trip signal subsequently occurs when the pressurizer low pressure trip setpoint is reached. A safety injection

$1gnal (515) is generated when the appropriate setpoint is reached. The countermeasures will limit the consequences of the accident in two ways 1

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[ a.

i Reactor trip and b: rated water injection complement void formation in causing rapid reduction of power to a residual level corresponding to

. fission product decay heat. However, no credit is taken in the LOCA analysis for baron content of the injection water aiding in shutdown.

In addition, the insertion of control rods to shut down the reactor is neglected in the large break analysis,

b. Injection of borated water provides for heat transfer from the core and prevents excessive clad temperatures.

The sequence of events following a large break LOCA are presented in Figure 15.6.4-1.

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 fission product decay, hot internals and the vessel continues to be transferred to the reactor coolant. At the beginning of the blowdown phase, the entire RCS contains subcooled liquid which transfers heat from the core by forced convection with some fully developed nucleate boiling. Thereafter, the core heat transfer is based on local conditions with transition boiling and forced convection to steam as the major heat transfer mechanisms.

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The heat transfer between the Reactor Coolant System and the secondary system may be in either direction depending on the relative temperatures. In the .

case of continued heat addition to the secondary, secondary system pressure increases and the main steam safety valves may actuate to limit the pressure.

Makeup water to the secondary side is automatically provided by the auxiliary feedwater system. The SIS actuates a feedwater isolation signal which ,

j isolates normal feedwater flow by closing the main feedwater isolation valves .

and also initiates emergency feedwater flow by starting the auxiliary feedwater pumps. The secondary flow aids in the reduction of Reactor Coolant System pressure.

When the Reactor Coolant system depressurires to approximately 600 psia, the accumulators begin to inject borated water into the reactor coolant loops.

l Since the loss of offsite power is assumed, the reactor coolant pumps are assumed to trip at the beginning of the accident. The effects of pump j coastdown are included in the blowdown analysis.

The blowdown phase of the transient ends when the RCS pressure (initially  !

assumed at 2280 psta) falls to a value approaching that of the containment atmosphere. Prior to or at the end of the blowdown, the mechanisms that are responsible for the bypassing of emergency core cooling injection water into '.

the RCS are calculated not to be effective. At this time (called end of bypass) refill of the reactor vessel lower plenum begins. Refill is complete when emergency core cooling water has filled the lower plenum of the reactor vessel, which is bounded by the bottom of the fuel rods (called bottom of core recovery time).

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The r; flood phase cf the transicnt is defined as the time period lasting from the end extent of the that refillcore until the reactorrise temperature vessel has been has been filled with water to the terminated.

From the later stage of blowdown and then the beginning of reflood, the safety injection accumulator tanks rapidly discharge borated cooling water into the RCS, contributing to the filling of the reactor vessel downcomer. The downcomer water elevation head provides the driving force required for the reflooding of the reactor core. The low head and high head safety injection pumps aid in the filling of the downcomer and subsequently supply water to maintain a full downcomer and complete the reflooding process. The safety injection pumped flow as a function of pressure is given in Table 15.6.4-6 for the large break cases.

Continued operation of the ECCS pumps supplies water during long-term cooling. Core temperatures have been reduced to long-term steady-state levels associated with dissipation of residual heat. Af ter the water level in the refueling water storage tank (RWST) reaches a minimum allowable value, coolant for long-term cooling of the core is obtained by switching to the cold leg recirculation phase of operation in which spilled borated water is drawn from the containment sump by the low head safety injection (RHR) pumps and returned to the RCS cold legs. The Containment Spray System continues to operate to further reduce containment pressure. Approximately 15 hours1.736111e-4 days <br />0.00417 hours <br />2.480159e-5 weeks <br />5.7075e-6 months <br /> after initiation of the LOCA, the ECCS is realigned to supply water to the RCS hot legs in order to control the boric acid concentration in the reactor vessel.

Descr10 tion of Small Break LOCA Transient Ruptures of small cross section will cause expulsion of the coolant at a rate which can be acconinodated by the charging pumps. These pumps would maintain an operational water level in the pressurizer permitting the operator to execute an orderly shutdown. The coolant which would be released to the containment contains the fission products existing at equilibrium.

The maximum break size for which the normal makeup system can maintain the pressurizer level is obtained by comparing the calculated flow from the Reactor Coolant System through the postulated break against the charging pump makeup flow at normal Reactor Coolant System pressure, i.e., 2250 psia. A makeup flow rate from one centrifugal charging pump is typically adequate to sustain pressurizer level at 2250 psia for a break through a 0.375 inch diameter hole. This break results in a loss of approximately 17.25 lb/sec.

Should a larger break occur, depressurization of the Reactor Coolant System causes fluid to flow into the loops from the pressurizer resulting in a pressure and level decrease in the pressurizer. Reactor trip occurs when the low pressurtzer pressure trip setpoint is reached. During the earlier part of the small break transient, the effect of the break flow is not strong enough to overcome the flow maintained by the reactor coolant pumps through the core as they are coasting down following reactor trip. Therefore, upward flow through the core is maintained. The Safety injection System is actuated when the appropriate setpoint is reached. The consequences of the accident are limited in two ways:

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1. Reactor trip and berated water injection complement void formation in the i .

core and cause a rapid reduction of nuclear power to a residual level

, corresponding to the delayed fission and fission product decay.

2. Injection of borated water ensures sufficient flooding of the. core to

prevent excessive clad temperatures.

t Before the break occurs the plant 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 continues to be transferred to the Reactor Coolant System. The heat transfer between the Reactor Coolant System and the secondary system may be in either direction i depending on the relative temperatures. In the case of continued heat addition to the secondary, system pressure increases and steam dump may )

occur. Makeup to the secondary side is automatically provided by the '

{ auxiliary feedwater pumps. The safety injection signal stops normal feedwater

, flow by closing the main feedwater line isolation valves and initiates auxiliary feedwater flow by starting auxiliary feedwater pumps. The secondary I flow aids in the reduction of Reactor Coolant System pressures.

i l When the RCS depressurizes to 600 psia, the cold leg accumulators begin to i inject water into the reactor coolant loops. Due to the loss of offsite power  :

1 assumption, the reactor coolant pumps are assumed to be tripped at the time of reactor trip during the accident and the offects of pump coastdown are

included in the blowdown analyses.

i 15.6.4.2 Analysis of Effects and Conseauences i

Methods of Analysis l

The requirements of an acceptable ECCS Evaluation Model are presented in Appendix K of 10 CFR 50 (Reference 3). The requirements of Appendix K J regarding specific model features were met by selecting models which provide 4 1

significant overall conservatism in the analysis. The assumptions made 1 pertain to the conditions of the reactor and associated safety system 1 4

equipment at the time that the LOCA occurs and include such items as the core i peaking factors, the containment pressure, and the performance of the ECCS

! system. Decay heat generated throughout the transient is also conservatively 1 calculated as required by Appendix K of 10 CFR 50. The thermal-hydraulic l j analyses reported in this section were performed with an upper head fluid  ;

temperature of Tcold* I i i l Larae Break Evaluation Model i

j The analysis of a large break LOCA transient is divided into three phases: l

(1) blowdown, (2) refill, and (3) reflood. There are three distinct transients analyzed in each phase: (1) the thermal-hydraulic transient in the l RCS, (2) the pressure and temperature transient within the containment, (3) i and the fuel and cladding 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 of the LOCA. l i

The description of the various aspects of the LOCA analysis methodology is given in References 4, 10, 13 and 14. These documents describe tho major

phenomena modeled, the interfaces among the computer codes, and the features l

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cf the codes which ensure compliance with the Acceptance Criteria. The SATAN-VI (Reference 5), WREFLOOD (Reference 6), LOTIC (Reference 7), BART (Reference 13), BASH (Reference 14) and LOCTA-IV (Reference 8) codes are used to assess the core heat transfer geometry and to determine if the core remains amenable to cooling throughout and subsequent to the blowdown, refill, and reflood phases of the LOCA. The SATAN-VI computer code analyzes the thermal-hydraulic transient in the RCS during blowdown. The WREFLOOD and BASH computer codes are used to calculate the thermal-hydraulic transient during the reflood phase of the accident. The BART computer code is used to calculate the fluid and heat transfer conditions in the core during reflood.

The LOTIC computer code is used to calculate the containment pressure transient during all three phases of the LOCA analysis. Similarly, the ,

LOCTA-IV computer code is used to compute the thermal transient of the hottest '

fuel rod during the three phases. Fuel parameters input to the LOCTA-IV code were taken from a new version of the PAD code (Reference 9).

SATAN-VI is used to calculate the RCS pressure, enthalpy, density and mass and energy flow rates, as well as steam generator heat transfer between the primary and secondary systems, as a function of time during the blowdown phase of the LOCA. SATAN-VI also calculates the accumulator water mass ard internal -

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 and refill phases, these data are transferred to the WREFLOOD code. The mass and energy release rates during blowdown and reflood are transferred to the LOTIC code for use in the determination of the cor.tainment pressure response during these phases of the LOCA. Additional SATAN-VI output data from the end of blowdown and refill, including the core pressure and the core power decay transient, are input to the LOCTA-IV code.

BASH is an integral part of the ECCS evaluation model which provides a more realistic thernal-hydraulic simulation of the reactor core and RCS during the reflood phase of a LOCA. Instantaneous values of accumulator conditions and safety injection flow at the time of completion of lower plenum refill are provided to BASH by WREFLOOD. Figure 15.6.4-2 illustrates how BASH has been substituted for WREFLOOD in calculating transient values of core inlet flow, enthalpy, and pressure for the detailed fuel rod model, LOCTA. A more detailed description of the BASH code is available in Reference 14. The BASH code provides a much more sophisticated treatment of steam / water flow phenomena in the reactor coolant system during core reflood. A more dynamic interaction between the core thermal-hydraulics and system behavior i:

expected, and recent experiments have borne this out. In the BASH code reflood model, BART provides the entrainment rate for a given flooding rate, then a system model determines loop flows and pressure drops in response t.

the calculated core exit flow. An updated inlet flow is used to calculate a new entrainment rate. This system will produce a more dynamic flooding transient, which reflects the close coupling between core thermal-hydraulics and loop behavior.

The LOTIC code is a nuthematical model of the ice condenser containment.

LOTIC is described in detail in Reference 7. LOTIC is run using output from SAlAN and WREFLOOD, which provide the necessary mass and energy releases to the containment. In this analysis the WREFL000/LOTIC system is used only to provide containment boundary conditions required by BASH.

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i The LOCTA cede is a computer program that evaluates fuel, cladding and coolant temperatures during a LOCA. A more complete description than is presented here can be found in Reference 8. During refill and reflood, the ECCS model uses a code consisting of LOCTA coupled with BART to yield a significant improvement in fuel rod behavior prediction. In the BART/LOCTA detailed fuel rod model, for the calculation of local heat transfer coefficients, the empirical FLECHT correlation is replaced by the BART code. BART employs rigorous mechanistic models to generate heat transfer coefficients appropriate to the actual flow and heat transfer regimes experienced by the LOCTA fuel rods. This is considered a more dynamic realistic approach than relying on a static empirical correlation.

Small Break LOCA Evaluation Model The NOTRUMP computer code is used in the analysis of loss-of-coolant accidents due to small breaks in the reactor coolant system. The NOTRUMP computer code is a state-of-the-art one-dimensional general network code consisting of a number of advanced features. Among these features are the calculation of thermal non-equilibrium in all fluid volumes, flow regime-dependent drif t flux calculations with counter-current flooding limitations, mixture level tracking logic in multiple-stacked fluid nodes, and regime-dependent heat transfer correlations. The NOTRUMP small break LOCA emergency core cooling system (ECCS) evaluation model was developed to determine the RCS response to design basis small break LOCAs and to address the NRC concerns expressed in NUREG-0611. " Generic Evaluation of Feedwater Transients and Small Break Loss-of-Coolant Accidents in Westinghouse Designed Operating Plants.'

In NOIRUMP, the RCS is nodalized into volumes interconnected by flowpaths.

The broken loop is modeled explicitly with the intact loops lumped into a

, second loop. The transient behavior of the system is determined from the l l governing conservation equations of mass, energy and momentum applied throughout the system. A detailed description of NOTRUMP is given in i References 11 and 15.

The use of NOTRUMP in the analysis involves, among other things, the representation of the reactor core as heated control volumes with an associated bubble rise model to permit a transient mixture height -

calculation. The multinode capability of the program enables an explicit and detailed spatial representation of various system components. In particular, it enables a proper calculation of the behavior of the loop seal during a loss-of-coolant transient.

Cladding thermal analyses are performed with the LOCTA-IV (Reference 8) code which uses the RCS pressure, fuel rod power history, steam flow past the uncovered part of the core, and mixture height history from the NOTRUMP hydraulic calculations, as input.

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A schematic representation of the computer code interfaces is given in Figure l

l 15.6.4-3.

. The small break analysis was performed with the approved Westinghouse ECCS Small Break Evaluation Model (References 8,11 and 15).

( Large Break Input Parameters and Initial Conditions Table 15.6.4-1 lists important input parameters and initial conditions used in the large break analyses.

Small Break Input Parameters and Initial Conditions Table 15.6.4-1 lists important input parameters and initial conditions used in the small break analyses.

The axial power distribution and core decay power assumed for the small break analyses are shown in Figures 15.6.4-60 and 15.6.4-61.

Safety injection flow rate to the Reactor Coolant System as a function of the system pressure is used as part of the input. The Safety Injection (SI) system was assumed to be delivering to the RCS 25 seconds after the generation of a safety injection signal.

For these analyses, the SI delivery considers pumped injection flow which is depicted in Figure 15.6.4-62 as a function of RCS pressure. This figure represents injection flow from the SI pumps based on performance curves degraded 5 percent from the design head. The 25 second delay includes time required for diesel startup and loading of the safety injection pumps onto the emergency buses. The effect of flow from the RHR pumps is not considered here since their shutoff head is lower than RCS pressure during the time portion of the transient considered here. Also, minimum safeguards Emergency Core Cooling System capability and operability has been assumed in this analysis.

The hydraulic analyses are performed with the NOTRUMP code using 102% of the licensed NSSS core power. The core thermal transient analyses are performed with the LOCTA-IV code using 102% of licensed NSSS core power.

Larae Break Results Based on the results of the LOCA sensitivity studies (Reference 12), the limiting large break was found to be double-ended cold leg guillotine (DECLG). Therefore, only the DECLG break is considered in the large break ECCS performance analysis. Calculations were performed for a range of Moody break discharge coefficients (CD ). Consistent with the methodology described in Reference 16 the break size which resulted in the worst case for minimum safety injection was used in a calculation in which no failures of the ECCS were assumed (Maximum safeguards). The results of these calculations are summarized in Tables 15.6.4-2 through 15.6.4-5.

Figures 15.6.4-4 through 15.6.4-44 present the parameters of principal interest from the large break ECCS analyses. Transients of the following parameters are presented for each discharge coef ficient analyzed, and where appropriate for the worst break maximum safeguards case.

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Figure 15.6.4-4 The follcwing quantities are presented at the hot spot through (location of maximum clad temperature) on the hottest fuel Figure 15.6.4-15 rod (hot rod):

1. fluid quality

2. mass velocity

3. heat transfer coefficient The heat transfer coefficient shown is calculated by the LOCTA-IV code.

Figure 15.6.4-16 The system pressure shown is the calculated pressure in through the core. Core flowrates are also presented.

Figure 15.6.4-21 Figure 15.6.4-22 These figures show the hot spot clad temperature transient through and the clad temperature transient at the burst location.

Figure 15.6.4-29 The fluid temperature shown is also for the hot spot location.

Figure 15.6.4-30 These figures show the core reflood transient.

through Figure 15.6.4-37 Figure 15.6.4-38 These figures show the cold leg accumulator delivery through during blowdown.

Figure 15.6.4-40 Figure 15.6.4-41 The pumped safety injection during reflood and the through calculated containment pressure are presented for the Figure 15.6.4-44 CD = 0.6 DECLG maximum and minimum safeguards cases.

The maximum cladding temperature calculated for a large break is 1962*F which is less than the Acceptance Criteria limit of 2200*F of 10 CFR 50.46. The maximum local metal water reaction is 6.2 percent, which is well below the embrittlement limit of 17 percent as required by 10 CFR 50.46. The total core metal water reaction is less than 0.3 percent for all breaks, as compared with the 1 percent criterion of 10 CFR 50.46, and the cladding temperature transient is terminated at a time when the core geometry is still amenable to cooling. As a result, the core temperature will continue to drop and the ability to remove decay heat generated in the fuel for an extended period of time will be provided.

Small Break Results As noted previously, the calculated peak cladding temperature resulting from a i small break LOCA is less than that calculated for a large break. A range of l small break analyses are presented which establishes the limiting break size.

The results of these analyses are summarized in Tables 15.6.4-7 and 15.6.4-8.

Figures 15.6.4-63a through 15.6.4-71 present the principal parameters of interest for the small break ECCS analyses. For all cases analyzed, the following transient parameters are included:

a. RCS pressure

b. core mixture height

c. hot spot clad temperature 89920:10/092685 15.6-14 s

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• 4 For the limiting break analyzed (3 inch), the following additional transient

. parameters are presented (Figures 15.6.4-72 through 15.6.4-74):

a. core steam flow rate l

b. core heat transfer coefficient '

I 1

c. hot spot fluid temperature l l

The maximum calculated peak cladding temperature for the small breaks analyzed is 1488'F. These results are well below all Acceptance Criteria limits of 10 CFR 50.46 and no case is limiting when compared to the results presented for large breaks.

Transition Core Impact The large break loss-of-coolant accident (LOCA) analysis presented herein for McGuire Units 1 and 2 considered a fuel core of optimized fuel. This is consistent with the methodology employed in the Reference Core Report 17 x 17 Optimized Fuel Assembly (OFA) for 17 x 17 0FA Transition (WCAP-9500).

When assessing the impact of transition cores on large break LOCA analysis, it must be determined whether the transition core can have a greater calculated peak clad temperature (PCT) than either a complete core of the reference design or a complete core of the new fuel design. For a given peaking factor, the only mechanism available to cause a transition core to have a greater calculated PCT than a full core of either fuel is the possibility of flow redistribution due to fuel assembly hydraulic resistence mismatch. This hydraulic resistance mismatch may exist only for transition cores and is the only unique difference between a complete core of either fuel type and the transition core.

The difference in fuel assembly resistance (K/A2) for the two assembly designs [17 x 17 Standard /17 x 17 0FA] may impact two portions of the large break LOCA analysis model. One is the reactor coolant system (RCS) blowdown portion of the transient analyzed with the SATAN-VI computer code, where the higher resistance 17 x 17 0FA assembly has less cooling flow than the 17 x 17 standard fuel assembly. While the SATAN-VI code models the crossflows between the average core flow channel (N-1 fuel assemblies) and a hot assembly flow channel (one fuel assembly), experience has shown that SATAN-VI results are

not significantly affected by small differences in the hydraulic resistance between these two channels.

To better understand the transition core large break LOCA blowdown transient phenomena, conservative blowdown fuel clad heatup calculations have been

, performed to determine the clad temperature effect on the new fuel design for mixed core configurations. The effect was determined by reducing the axial flow in the hot assembly at the appropriate elevations to simulate the effects of the transition core hydraulic resistance mismatch. In addition, the W blowdown evaluation model was modified to account for grid heat transfer enhancement during blowdown for this evaluation. The results of this analysis have shown that no peak clad temperature penalty is observed during blowdown.

Therefore, it is not necessary to perform a new blowdown calculation for transition core configurations because the Evaluation Model blowdown calculation' performed for the full 17 x 17 0FA core is conservative.and 9

bounding.

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i 89920:10/092685 15.6-15

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The other portion of the LOCA calculatien impacted by hydraulic resistance mismatch is the core reflood transient. Fuel assembly design specific

^ analyses have been performed with a version of the BART computer code which accurately models mixed core cases during reflood. Westinghouse transition core designs including specific 14 x 14,15 x 15 and 17 x 17 standard to 0FA transition core cases were analyzed. For each of these cases, BART modelled both fuel assembly types and predicted the reduction in axial flow at the appropriate elevations. As expected, the increase in hydraulic resistance mismatch for the 17 x 17 0FA assembly was shown to produce a reduction in reflood steam flow rate for the 17 x 17 0FA assembly at the mixing vane grid elevations during the transition core period. This reduction in steam flow rate is offset by the fuel grid heat transfer enhancement predicted during reflood. The various fuel assembly specific transition core analyses performed resulted in peak clad temperature increases of up to 10*F for core axial elevations where PCTs can possibly occur. Therefore, the maximum PCT penalty possible for 17 x 17 0FA during transition cores is 10'. Once a full core of the 17 x 17 0FA fuel is achieved, the large break LOCA analysis with UHI removed will apply without the crossflow penalty.

15.6.4.3 Environmental Consecuences The postulated consequences of a LOCA are calculated for 1) offsite and 2) control room operators.

Offsite Dose Consequences The offsite radiological consequences of a LOCA are calculated based on the following assumptions and parameters.

1. 100 percent of the core noble gases and 25 percent of the core iodines are released to the containment atmosphere.

2. 50 percent of the core iodines are released to the containment.

3. Annulus activity which is exhausted prior to the time at which the annulus reaches a negative pressure of -0.25 in, w.g. is unfiltered.

4. ECCS leakage begins at the earliest possible time sump recirculation can begin.

5. ECCS leakage occurs at twice the maximum operational leakage.

6. Bypass leakage is 7 percent of total containment leakage.

7. The effective annulus volume is 50 percent of the actual volume.

8. The annulus filters become faulted at 900 seconds resulting in a 15 percent reduction in flow.

9. Elemental iodine removal by the ice condenser begins at 600 seconds and continues for 2540 seconds with.a removal efficiency of 30 percent.

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10. One of the containment air return fans is assumed to fail.

11. The containment leak rate is fifty percent of the Technical Specification limit after 1 day.

12. Iodine partition factor for ECCS leakage is 0.1 for the course of the l accident.

13. No credit is taken for the auxiliary building filters for ECCS leakage.

14. The redundant hydrogen recombiners and igniters fail. Therefore, purges are required for hydrogen control.

15. The annulus reaches equilibrium after 200,000 seconds such that the only discharge is due to inleakage.

16. Water density at 160'F is used to calculate the sump water mass.

17. Other assumptions are listed in Table 15.6.4-11.

Based on the model in Appendix 15A, the thyroid and whole body doses are calculated at the exclusion area boundary and the low population zone. The doses are presented in Table 15.6.4 -11 and are within the limits of 10 CFR 100.

Control Room Operator D6se .

The maximum postulated dose to a control room operator is determined based on the releases of a Design Basis Accident. In addition to the parameters and assumptions listed above, the following apply:

1. The control room pressurization rate is 1,000 cfm; the filtered recirculation rate is 1,000 cfm.

2. The unfiltered inleakage into the control room is 10 cfm.

3. Other assumptions are listed in Table 15.6.4-12.

15.6.5 A NUMBER OF BWR TRANSIENTS Not applicable to McGuire.

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REFERENCES FOR SECTION 15.6

1. Burnett, T. W. T., et. al. , "LOFTRAN Code Description", WCAP-7907, June 1972.

2. Chelemer, H., Boman, L. H. , Sharp, D. R., " Improved Thermal D' esign Procedures", WCAP-8587, July 1975.

3. " Acceptance Criteria for Emergency Core Cooling System for Light Water Cooled Nuclear Power Reactors",10 CFR 50.46 and Appendix K of 10 CFR

50. Federal Register, Volume 39, Number 3, January 4,1974.

4. Bordelon, F. M., Massie, H. W. and Borden, T. A., " Westinghouse ECCS Evaluation Model-Summary", WCAP-8339, (Non-Proprietary), July 1974.

5. Bordelon, F. M. , et. al. , " SATAN-VI Program: Comprehensive Space Time Dependent Analysis of Loss of Coolant", WCAP-8302, (Proprietary) June 1974, and WCAP-8303, (Non-Proprietary), June 1974.

6. Kelly, R. D., et. al., " Calculated Model for Core Reflooding After a loss of Coolant Accident (WREFLOOD Code", WCAP-8170 (Proprietary) and WCAP-8171 (Non-Proprietary), June 1974

7. Hsieh, T. and Raymond, M., "Long-Term Ice Condenser Containment LOTIC Code Supplement 1", WCAP-8355 Supplement 1, May 1975, WCAP-8354 (Proprietary), July 1974.

8. Bordelon, F. M. , et. al . , "LOCTA-IV Program: Loss of Coolant Transient Analysis", WCAP-8301 (Proprietary) and WCAP-8305, (Non-Proprietary),

June 1974.

9. Rahe, E. P., Westinghouse letter to Thomas, C. O., U.S.N.R.C., Letter Number NS-EPR-2673, October 27,1982,

Subject:

" Westinghouse Revised PAD Code Thermal Safety Model", WCAP-8720, Addendum 2 (Proprietary).

10. Eicheldinger, C., " Westinghouse ECCS Evaluation Model, February,1978 Version", WCAP-9220 (Proprietary) February,1979, and WCAP-9221 (Non-Proprietary) February, 1978.

11. Lee, H., Tauche, W. D., Schwarz, W. R., " Westinghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code", WCAP-10081-A, August 1985.

12. Salvatori, R., " Westinghouse Emergency Core Cooling System - Plant Sensitivity Studies", WCAP-8340, (Proprietary) July 1974.

13. Young, M., et. al., "BART-1A: A Computer Code for the Best Estimate Analyzed Reflood Transients", WCAP-9561-P-A,1984 (Westinghouse Proprietary).

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l 89920:10/092685 15.6-18 s

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1

. I

14. Kabadi, J. N., et. al., " BASH: An Integrated Core and RCS Reflood Code :

. for Analysis of PWR Loss-of-Coolant Accidents", WCAP-10266,1984 (Westinghouse Proprietary). )

15. Meyer, P. E. and Kornfilt, J., "NOTRUMP, A Nodal Transient Small Break and General Network Code", WCAP-10080-A, August 1985.

16. Rahe, E. P. (Westinghouse), letter to Tedesco, R. L. (USNRC),

No. NS-EPR-2538, December 1981.

89920:10/092685 15.6-19

TABLE 15.6.4-1

, Input Parameters Used in the ECCS Analyses Parameter Laroe Break hmallBreak Peak Linear Power (kw/f t) 12.88 12.21 (includes 102% factor)

Total Peaking Factor, Fg 2.32 2.32 Power Shape Chopped See Figure Cosine 15.6.4-60 Fuel Assembly Array 17 X 17 17 X 17 Optimized Optimized Nominal Cold Leg Accumulator 950 950 Water Volume (ft 3/ accumulator)

Nominal Cold Leg Accumulator 1350 1350 Tank Volume (ft 3/ accumulator)

Minimum Cold Leg Accumulator . 600 600 Gas Pressure (psia)

Pumped Safety Injection Flow See Table See Figure 15.6.4-6 15.6.4-62

• Steam Generator Initial Pressure (psia) 987.0 987.0 Steam Generator Tube 5 5 Plugging Level (%)

I i

8992Q:10/091285 15.6-20

J TABLE 15.6.4-2 Larae Break LOCA Time Secuence of Events CD = 0.8 CD = 0.6 CD = 0.4 DECLG DECLG DECLG 4

(sec) (sec) ,__( sec)

Start 0.0 0.0 0.0 Reactor Trip Signal 0.46 0.46 0.47 Safety Injection Signal 2.6 2.7 2.9 Cold Leg Accumulator Injection 12.7 15.3 21.0 Pump Injection 27.6 27.7 27.9 End of Bypass 28.42 33.1 44.5 End of Blowdown 28.6 34.8 44.6 Bottom of Core Recovery 45.3 50.9 64.1 Cold Leg Accumulator Empty 67.2 71.0 79.1 89920:10/092685 15.6-21 s

- -.-____-_z_. __

TABLE 15.6.4-3 Larae Break LOCA Time Seauence of Events Maximum Safeguards CD = 0.6 DECLG (sec)

Start 0.0 Reactor Trip Signal 0.46 Safety Injection Signal 2.7 Cold Leg Accumulator Injection 15.3 Pump Injection 27.7 End of Bypass 33.1 '

End of Blowdown 34.8 i Bottom of Core Recovery 49.7 Cold Leg Accumulator Empty 75.6 i

9 89920:10/092685 15.6-22

- - - , - -- - ~ , , , - - - . - . . , . . . , . - - . , , - , - - -

TABLE 15.6.4-4 Larae Break LOCA Results Fuel Claddina Data CD = 0.8 CD = 0.6 CD = 0.4 DECLG DECLG DECLG RESULTS Peak Clad Temperature (*F) 1944 1962 1915 Peak Clad Temperature Location (ft) 6.25 6.25 6.75 Local Zr/H 2O Reaction (max), (%) 5.88 6.16 4.52 Local Zr/H2 O Location (ft) 6.00 6.00 6.75 Total Zr/H 2O Reaction (%) <0.3 <0.3 <0.3 Hot Rod Burst Time, (sec) 63.3 56.7 92.7 Hot Rod Burst Location, (ft) 6.00 6.00 5.50 l

, 8992Q:10/092685 15.6-23

--- ,, -.... ,, -..,- -.--- __, . _ . ., , , . . - . , _ - _ . , - - - - , - , , - , .- - - . ~ - - - - - , - _ , - - -

TABLE 15.6.4-5

, Larae Break LOCA Results Fuel Cladding Data ,

Maximum Safeguards CD = 0.6 DECLG RESULTS i

Peak Clad Temperature (*F) 1933 Peak Clad Temperature Location (ft) 6.50 Local Zr/H 2O Reaction (max), (%) 4.71 Local Zr/H2 O Location (ft) 6.00 Total Zr/H 2O Reaction, (%) <0.3 Hot Rod Burst Time, (sec) 56.7 Hot Rod Burst Location, (ft) 6.00 8992Q:lD/092685 15.6-24 s

TABLE 15.6.4-6 Safety Injection Pumped Flow e

Assumed for Breaks Greater Than or Equal to 10 Inches MIMIMUM SAFEGUARDS:

Pressure SI Flow (psia) (1b/sec) 14.7 493.2 34.7 437.2 54.7 378.6 74.7 315.5 114.7 198.4 214.7 100.2 614.7 81.2 1014.7 58.5 3014.7 0.0 MAXIMUM SAFEGUARDS:

Pressure SI Flow (Dsia) (lb/sec) 14.7 1078 34.7 997 54.7 914 74.7 826 94.7 731 114.7 621 214.7 171 314.7 166

\

1 8992Q:lD/092685 15.6-25 b

TABLE 15.6.4-7

, Small Break LOCA Time Secuence of Events 2 in 3 in 4 in 6 in (Sec) (Sec) (Sec) (Sec)

Start 0.0 0.0 0.0 0.0 Reactor Trip 23.58 19.24 9.09 6.35 Top of Core Uncovered N/A 1026.5 681.9 289.2 Cold Leg Accumulator Injection N/A N/A 891.9 393.2 Peak Clad Temperature Occurs N/A 1737.6 982.0 453.5 Top of Core Covered N/A >2400 1642.1 503.2

)

8992Q:10/092885 15.6-26

L.

. TABLE 15.6.4-8 Small Break LOCA Results Fuel Claddina Data 2 in 3 in 4 in 6 in RESULTS Peak Clad Temperature (*F) N/A 1488 1348 1189 Peak Clad Location (ft) N/A 12 12 12 Local Zr/H 2O Reaction (max), (%) N/A 0.81 0.18 0.078 Local Zr/H2O Reaction Location Ft. N/A 12 12 12 Total Zr/H 2O Reaction, (%) N/A <0.3 <0.3 <0.3 Hot Rod Burst Time, (sec) N/A N/A N/A N/A Hot Rod Burst Location, (ft) N/A N/A N/A N/A 1

e l

8992Q:10/092685 15.6-27  !

, l 1

No Change From

. 1984 FSAR Update Table 15.6.4-11 (Page 1 of 2) f Parameters for LOCA Offsite Dose Analysis  !

4

1. Data and assumptions used to estimate i radioactive source from postulated

• j accidents

.1 1 a. Power Level (MWt) 3565

b. Failed fuel 100% of fuel rods in core 1 c. Activity released to containment atmo-sphere from failed fuel and available for 2

release (percent of core activity)

, i i Noble gases 100 Iodines 25

d. Iodine activity released to containment ' 50 sump and available for release via ECCS leakage outside contaiment (percent of core activity)

e. Iodine fractions (organic, elemental, Regulatory Guide 1.4 and particulate)

2. .0ata and assumptions used to estimate act- '

ivity released l a. Containment free volume 3

Upper containment volume (ft3) 6.70?*05 4

Lower containment voh.me (ft3) 3.68E+05 l Total containment free volume (ft3) 1.038E+06 l i b. Containment leak rate (percent of j containment volume per day) i 4

0< t < 24 hrs 0.2 r

t > 24 hrs 0.1 r

c. Bypass leakage fraction 0.07  !

d. Annulus ventilation iodine filter i l

t efficiency (percent) i 95 i .

15.6-28 L

i lif84 Upda:e 0

No Change From 1984 FSAR Update Table 15.6.4-11 (Page 2 of 2) ' .

Parameters fnr LOCA Offsite Dose Analysis

e. Total annulus ventilation flow rate (cfm) 0-900 sec 8000 900+ sec 6800

f. Time at which annulus ventilation system is operable (sec) 23

g. Annulus volume (ft3) 422,361

3. Dispersion data .

a. Distance to exclusion area boundary (m) 762

b. Distance to low population zone (m) 8850

c. X/Q at exclusion area boundary (sec/m3) .

0-2 hrs 9.0E-04

d. X/Q at low population zone (sec/m3) 0-8 hrs 8-24 hrs 8.0E-05 (

5.2E-06 1-4 days 1.7E-06 4+ days 3.7E-07

4. Dose data

a. Method of dose calculations Appendix 15A b .- Dose conversion assumptions Appendix 15A

c. Doses (Rem)

Case 1 (With ECCS leakage)

Exclusion Area Boundary Whole Body 3.0 Thyroid 2.0E+02 Low Population Zone Whole Body 6.2E-01 Thyroid 6.5E+01 Case 2 (Without ECCS leakage)

Exclusion Area Boundary Whole Body 2.9 Thyroid 1.8E+02 Low Population Zone Whole Body 6.1E-01

') /

Thyroid 5.7E+01 15.6-29 1984 Update

No Change From 1984 FSAR Update Table 15.6.4-12 (Page 1 of 2)

Parameters for LOCA Control Room Dose Analysis

1. Data and assumptions used to estimate radioactive source from postulated accidents

a. Power Level (MWt) 3565

b. Failed fuel 100% of fuel rods in core

c. Activity released to containment atmos-phere from failed fuel and available for release (percent of core activity)

Noble gases 100 Iodines 50

d. Iodine activity released to contaiment e 50 sump and available for release via ECCS leakage outside containment (percent of core activity)

e. Iodine fractions (organic, elemental, C and particulate)

Regulatory Guide 1.4

2. Data and assumptions used to estimate act-ivity released

a. Containment free volume (ft3) 1.038E+06

b. Containment leak rate (percent of containment volume per day) 0<t<24 hrs tI24 hrs 0.2 0.1

c. Bypassleakagefraction 0.07

d. Control room pressurization rate (cfm) 1000 e.

Control room filtered recirculation rate (cfm) 1000

f. Control room unfiltered:in-leakage (cfm) 10

g. Control room volume (ft3) 116,000

h. Control room pressurization and re- .

99 circulation removal efficiencies for

. iodine (percent) 15.6-30 1984 Update t '

No Change From 1984 FSAR Update Table 15.6.4-12 (Page 2 of 2) -

Parameters for LOCA Control Room Dose Analysis

3. Dispersiondata-ControlroomintAkex/Q(sec/m3) 0-8 hrs 1.0E-03 8-24 hrs 7.0E-04 1-4 days 4.5E-04 4+ days 2.4E-04

4. Oose data  ;

a. Method of dose calculations Appendix 15A

b. Dose conversion assumptions Appendix 15A

c. Doses (Rem)

Whole body Thyroid 2.2E-01 2.6E+01 Skin 4.1 l

1 15.6-31 s

1984 Update

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APPENDIX B NON-LOCA TRANSIENTS 9

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15.1.4 INADVERTENT OPENING OF A STEAM GENERATOR RELIEF OR SAFET 15.1.4.1 Identification of Causes and Accident Description The most severe core conditions resulting from an accidental depressurization of the Main Steam System are associated with an inadvertent opening of a single steam dump, relief, or safety valve. The analyses performed assuming a rupture of a main steam line are given in Section 15.1.5.

The steam release as a consequence of this accident results in an initial in-crease falls.

in steam flow which decreases during the accident as the steam pressure ture and The energy removal from the RCS causes a reduction of coolant tempera-pressure. In the presence of a negative moderator temperature coefficient, the cooldown results in an insertion of positive reactivity.

The analysis is performed to demonstrate that the following criterion is satisfied:

Assuming a stuck rod cluster control assembly, with offsite power avail-able, and assuming a single failure in the Engineered Safety Features System, there will be no consequential damage to the core or Reactor Coolant System after reactor trip for a stream release equivalent to the (

spurious opening, with failure to close, of the largest of any single steam dump, relief, or safety valve.

The following systems provide the necessary protection against an accidental depressurization of the Main Steam System. -

1.

Safety Injection System actuation from any of the following:

a.

Two-out of-three low steamline pressure signals in any one loop

b. Two-out-of-four low pressurizer pressure signals.

c. Two-out-of-three high containment pressure signals.

2.

The overpower reactor trips (neutron flux and AT) and the reactor trip occurring in conjunction with receipt of the safety injection signal.

3.

Redundant isolation of the main feedwater lines.

Sustained high feedwater flow would cause additional cooldown. Therefore, in addition to the normal control action which will close the main feed- -

water valves following reactor trip, a safety injection signal would trip the main feedwater pumps and will generate a feedwater isolation signal which will rapidly close all main feedwater control valves, isolation valves, and pump discharge valves.

15.1-8 19Bi Update 4

4. Trip of the fast-acting main steam isolation valves (designed to close in less than 5 seconds) en:

a. Two-out-of-three low steamline pressure signals in any one loop.

b. Two-out-of-four high-high containment pressure signals,

c. Two-out-of-three high negative steamline pressure rate signals in any one loop (used only during cooldown and heatup operations).

Accidental depressurization of the secondary system is classified as an ANS Con-dition 11 event, a fault of moderate frequency. See Section 15.0.1 for a discussion of Condition II events.

15.1.4.2 Analysis of Effects and Consecuences Method of Analysis The following analyses of a secondary system steam release are performed for this section:

1. A full plant digital computer simulation using the LOFTRAN Code (Reference

1) to determine RCS temperature and pressure during cooldown, and the effect of safety injection.

2. Analyses to determine that the DNB design basis is met.

The following conditions are assumed to exist at the time of a secondary steam system release:

1. End-of-life shutdown margin at no-load, equilibrium xenon conditions, and with the most reactive rod cluster control assembly stuck in its fully withdrawn position. Operation of rod cluster control assembly banks during core burnup is restricted in such a way that addition of positive reacti-vity in a secondary system steam release accident will not lead to a more-adverse condition than the case analyzed.

2. A positive moderator density coefficient corresponding to the end-of-life rodded core with the most reactive rod cluster control assembly in the fully withdrawn position. The variation of the coefficient with temperature and pressure is included. The K,ff versus temperature at 1000 psia corresponding to the positive moderator density coe'ficient used is shown in Figure 15.1.4-1.

3. Minimum capability for injection of high concentration boric acid solu-tion corresponding to the most restrictive single failure in the Safety Injection System. This corresponds to the flow delivered by one charging pump delivering its full contents to the cold leg header. No credit is taken for the' low concentration boric acid which must be swept from the safety injection lines downstream of the refueling water storage tank prior to the delivery of concentrated boric acid ( " ppm from the refueling water storage tank) to the reactor coolant loops.

1900 15.1-9 1984 Update

4. The case studied is a steam flow of 248 pounds per second at 1100.+4.a PN.  ;

with offsite power available. This is the maximum capacity of any single steam dump, relief, or safety valve. Initial hot shutdown conditions at time zero are assumed since this represents the most conservative initial e

condition. Should the reactor be just critical or operating at power at the time of a steam release, the reactor will be tripped by the normal overpower protection when power level reaches a trip point. Following a trip at power, the RCS contains more stored energy than at no-load, the average coolant temperature is higher than at nc-load and there is appreciable energy stored in the fuel. Thus, the additional stored energy is removed via the cooldown caused by the steam release before the no-load conditions of RCS temperature and shutdown margin assumed in the analyses are reached. After the additional stored energy has been removed, the cooldown and reactivity insertions pr'oceed in the same manner as in the analysis which' assumes no-load condition at time zero. However, since the initial steam generator water inventory is greatest at no-load, the magnitude and duration of the RCS cooldown are less for a steam line release occurring at power.

E. In computing the steam flow, the Moody Curve (Reference 4) for f(L/D) = 0 i used.

6. Perfect mo hture separation in the steam generator is assumed.

Res ul t,s, The calculated time sequence of events for this accident is listed in Table 15.1.2-1. b The results presented are a conservative indication of the events which would occur assuming a secondary system steam release since it is postulated that all of the conditions described above occur simultaneously.

Figures 15.1.4-2 and 15.1.4-3 show the transient results for a steam flow of 248 lb/sec at 1100 psia. -

The assumed steam release is typical of the capacity of any single steam dump, relief, or safety valve. Safety injection is initiated automatically by low pressurizer pressure. Operation of one centrifugal charging pump is assumed.

Baron solution at.hMS pp enters the RCS from the refueling storage water tank (RWST) providing suffr: tent negative reactivity to prevent core damage.

1900 The transient is quite conservative with respect to cooldown, since no credit is taken for the energy stored in th+ system metal, other than that of the fuel elements, or the energy stored ir the other steam generators. Since the tran-sient occurs over a period of abcat ~ minutes, the neglected stored energy is likely to have a significant effe:t ir. slowing the cooldown.

15.1.4:3 Environmental Consec uenc! s The inadvertent opening of a singla ste2m dump relief or safety valve can re-suit in steam release from the sectndary system. If steam generator leakage exists coincident with the failed fael c)nditions, some activity will be re-leased.

15.1-10 1984 Update

15.1.4.4 Conclusions The analysis shows that the criteria stated earlier in this section are satis-fied. For an accidental depressurization of the main steam system, the DNB l

design limits are not exceeded. This case is less limiting than the steam line rupture case described in section 15.1.5.

15.'1.5 STEAM SYSTEM PIPING FAILURE 15.1.5.1 Identification of Causes and Accident Description The steam release arising from a rupture of a main steam line would result in l an initial increase in steam flow which decreases during the accident as the '

steam pressure falls. The energy removal from the RCS causes a reduction of coolant temperature and pressure. In the presence of a negative moderator tem-peratur.e coefficient, the cooldown results in an insertion of positive reactivity.

If the most reactive rod cluster control assembly (RCCA) is assumed stuck in its fully withdrawn position after reactor trip, there is an increased possibility that the core will become critical and return to power. A return to power foi-lowing a steam line rupture is a potential problem mainly because of the high power peaking factors which exist assuming the most reactive RCCA to be stuck

( in its fully withdrawn position. The core is ultimately shut down by the boric acid injection delivered by the Safety Injection System.

The analysis of a main steamline rupture is performed to demonstrate that the following criteria are satisfied:

Assuming a stuck RCCA with or without offsite power, and assuming a single failure in the engineered safety features, the core remains in place and in-tact. Radiation doses do not exceed the guidelines of 10 CFR 100.

~

Although DNB and possible clad perforation following a steam pipe. rupture j

s are not necessarily unacceptable, the following analysis, in fact, shows that no DNB occurs for any rupture assuming the most reactive assembly st'uck in its fully withdrawn position. The DNBR design basis is discussed in Sec-

' tion 4.4.

The major rupture of a steamline is the most limiting cooldown transient and is analyzed at zero power with no decay heat. Decay heat would retard the cool-down thereby reducing the return to power. A detailed analysis of this transient with the most limiting break size, a double-ended rupture, is presented here.

The following functions provide the protection for a steamline rupture:

1.

Safety Injection System actuation from any of the following:

a.

Two-out-of-three low steamline pressure signals in any one loop.

b. Two-out-of-four low pressurizer pressure signals.

c. Two-out-of-four high containment pressure signals.

15.1-11 1981. Update s

r 2.

The overpower reactor trips (neutron flux and AT) and the reactor trip occurring in conjunction with receipt of the safety injection signal.

3.

Redundant isolation of the main feedwater lines.

Sustained high feedwater flow would cause additional cooldown.

in addition to the normal control action which will close the main feed-Therefore, water valves, the safety injection signal will trip the main feedwater pumps and will generate a feedwater isolation signal which will rapidly close all main discharge feedwater control valves, isolation valves, and pump valves.

4.

Trip of the fast-acting main steam isolation valves (designed to close in less than 5 seconds) on:

a.

Two-out-of-three low steam line pressure s'ignals in any one loop.

b.

Two-out of-three high-high containment pressure signals.

! c.

Two out of-three high negative steam line pressure rate signals in any one loop (used only during cooldown and heatup operations.

l The main break in the steam steamisolation line. valves will-be. fully close within 10 seconds of a large of all valves would completely terminate the blowdown.For breaks downsteam of th location, no more than one steam generator would experience an uncontrolledFor anyi brea blowdown even if one of the isolation valves fails to close. A description of s, '

steam line isolation is included in Section 10.3.2.

• . Steam throat of flow theissteam measured by monitoring dynamic head in nozzles located in the generator.

The effective throat area of the nozzles is l 1.4 square feet, which is considerably less than the main steam pipe area.

Thus, any the nozzles also serve to limit the maximum steam flow for a break at -

location. '

A major steamline rupture is classified as an ANS Condition IV event. See Section 15.0.1 for a discussion of Condition IV events.

Effects sented in of this minor secondary system pipe breaks are bounded by the analysis pre-section.

Condition III events, as described in Section 15.0.1.3. Minor secondary system pip 15.1.5.2 Analysis of Effects and Consecuences Method of Analysis The analysis of thle steam pipe rupture has been performed to determine:

1.

The core heat flux and RCS temperature and pressure resulting from the cooldown following the steam line break.

has been used. . The LOFTRAN 4444L (Reference 1)

C ,j ,

15.1-12 1984 tpdate

2. The thermal and hydraulic behavior of the core following a steam line break. A detailed thermal and hydraulic digital-computer code, THINC, has been used to determine if DNB occurs for the core conditions com-puted in item 1 above.

Studies have been performed to determine the' sensitivity of steam line break results to various assumptions (Reference 5). Based upon this study, the fol-lowing conditions were assumed to exist at the time of the main steam line break accident:

1. End-of-life shutdown margin at no-load, equilibrium xenon conditions, and the most reactive RCCA stuck in its fully withdrawn position. Operation of the control rod banks during core burnup is restricted in such a way that addition of positive reactivity in a steamline break accident will not lead to a more adverse condition than the case analyzed.

2. A positive moderator density coefficient corresponding to the end-of-life rodded core with the most reactive RCCA in the fully withdrawn position.

The variation of the coefficient with temperature and pressure has been in-cluded. The K,ff versus temperature at 1000 psia corresponding to the positive moderator density coefficient used is shown in Figure 15.1.4-1.

The effect of power generation in the core.on overall reactivity is shown in Figure 15.1.5-1.

The core properties associated with the sector nearest the affected steam generator and those associated with the remaining sector were conservatively combined to obtain average core properties for reactivity feedback calcul-ations. Further, it was conservatively assumed that the core power dis-tribution was uniform. These two conditions cause underprediction of the reactivity feedback in the high power region near the stuck rod. To verify the conservatism of this method, the reactivity as well as the power dis-tribution was checked for the limiting statepoints for the cases analyzed.

This core analysis considered the Doppler reactivity from the high fuel ,

temperature near the stuck RCCA, moderator feedback from the high water enthalpy near the stuck RCCA, power redistribution and nonuniform core inlet temperature effects. For cases in which steam generation occurs in the high flux regions of the core, the effect of void formation was also included. It was determined that the reactivity employed in the kinetics analysis was always larger than the reactivity calculated in-cluding the above local effects for the statepoints. These results verify conservatism; i.e., underprediction of negative reactivity feed-back from power generation.

1900

3. Minimum capability for injection of boric acid (4rn e ppm from the RWST) solution corresponding to the most restrictive single failure in the Safety Injection System. The Emergency Core Cooling System consists of three systems: 1) the passive accumulators, 2) the Residual Heat Removal System, and 3) the Safety Injection System. Only the Safety Injection System and the accumulators are modeled for the steam line break accident analysis.

15.1-13 1984 update

The actual modeling of the Safety Injection System in LOFTRAN is described in Reference 1. The flow corresponds to that delivered by one charging pump delivering its full ' flow to the cold leg header. No credit has been taken for the low concentration borated water which must be swept from the lines downsteam of the RWST prior to the delivery of high concentration boric acid to the reactor coolant loops.

For the cases where offsite power is assumed, the sequence of events in the Safety Injection System is the following. After the generation of the safety injection signal (appropriate delays for instrumentation, logic, and signal transport included), the appropriate valves begin to operate and the charging pump starts. In 10 seconds, the valves are assum ed to be in their final position and the pump is assumed to be at full speed. The volume containing the low concentration borated water is swept before

~~~

the ppm water reaches the core. This delay, described above, is in-herenti included in the modeling.

1900 In cases where offsite power is not available, an additional 10 second delay is assumed to start the diesels and to load the necessary safety injection equipment onto them.

4.

Design value of the steam generator heat transfer coefficient including allowance for fouling factor.

5. Since the steam generators are provided with integral flow restrictors with a 1.4 square foot throat area, any rupture with a break area greater than 1.4 square feet, regardless of location, would have the same effect \

on the NSSS as the 1.4 square foot break. The following cases have been considered in determining the core power and RCS transients:

. a. Complete severance of a pipe, with the plant initially at no load conditions with full reactor coolant flow and offsite power available.

b. Case (a) with loss of offsite power simultaneous with the steam line break and initiation of the safety injection signal. Loss of offsite power results in reactor coolant pump coastdown.

6. Power peaking factors, corresponding to one stuck RCCA and n:n uniform core inlet coolant temperatures, are determined for end of core life. The cold-est core inlet temperatures are assumed to occur in the sector with the stuck rod. The power peaking factors account for the effect of the local void in the region of the stuck control assembly during the return to power phase following the steam line break. This void in conjunction with the large positive moderator density coefficient partially offsets the effect of the stuck assembly. The power peaking factors depend upon the core power, temperature, pressure, and flow, and thus are different for each case studied, The core parameters used for each of the two cases correspond to values determined from the respective transient analysis. .

Both cases above assume ~ initial hot standby conditions at time zero since this represents the most pessimistic initial condition. Should the reactor be just critical or operating at power at the time of a steam line break,  ;

i 15.1-14 193,cpe,e, 1

the reactor will be tripped by the normal overpower protection system when power level reaches a trip point. Following a trip at power, the RCS con-tains more stored energy than at no-load, the average coolant temperature is higher than at no-load, and there is appreciable energy stored in the fuel.

Thus, the additional stored energy is removed via the cooldown caused by the steam line break before the no-load conditions of RCS temperature and shut-down margin assumed in the analyses are reached. After the additio.nal stored energy has been removed, the cooldown and reactivity insertions proceed in the same manner as in the analysis which assumes no-load condition at time zero. A spectrum of steam line breaks at various power levels has been analyzed in Reference 5.

7. In computing the steam flow during a steam line break, the Moody Curve (Reference 4) for f(L/D) = 0 is used.

C. The Upper ";;d !nf;;tien (UMI) i: :f u?:ted. The ::tu tier pre :u ':-

ihu UHI .. eLv.e the seturetion pic;;ur; for th: 'n;; tic: :: l:nt #- t:

upper h;;d. 'h: ' :urg: Of ccid UH! ::ter E::p: thf: ir :t*ve ::r' -*

'r;; 'l;;hing nd ' rem r t rding the prc::ur; d ;r :::. 'h Off :t Of U"!

i: : f::ter pre::ure decre :: '+4:5 ia tu"* pe--4+e -a*a e"#a+y #afa*+4+-

flGo int; the CGie. The:6 cff;;t ;r i;ry ;m;I' :nd r: Cult: :" ~::

ignicantly ff::ted.

These assump'tions are discussed more fully in Reference 5.

Results The calculated sequence of' events for both cases analyzed is shown on Table 15.1.2-1.

The results presented are a conservative indication of the events which would occur assuming a steam line rupture since it is postulated that all of the conditions described above occur simultaneously.

Core Power and' Reactor Coolant System Transient Figures 15.1.5-2 through 15.1.5-4 show the RCS transient and core heat flux following a main steam line rupture (complete severance of a pipe) at initial no-load condition (case a).

Offsite power is assumed available so that full reactor coolant flow exists.

The transient shown assumes an uncontrolled steam release from only one steam generator. Should the core be critical at near zero power when the rupture occurs the initiation of safety injection by low steam line pressure will shut down the reactor. Steam release from more than one steam generator will be prevented by automatic trip of the fast acting isolation valves in the steam lines by low steam line pressure signals, high-high containment pressure signals, or high negative steam line pressure rate signals. Even with the failure of one valve, release is limited to no more than 10 seconds for the other steam generators while ti.e one generator blows down. The isolation valves are designed to be fully closed in less than 5 seconds from receipt of a closure signal, and the closure signal is expected to be generated within 5 secones.

15.1-15 198 Upda:e I

As shown in Figure 15.1.5-3 the core attains criticality with the RCCAs inserted (with the design shutdown assuming one stuck RCCA) shortly af ter boron solution at 2,"CC ppm enters the RCS. The continued addition of boron results in a peak corep\wersignificantlylowerthanthenominalfullpowervalue.

1900 The calculation assumes the boric acid is mixed with, and diluted by,' the water flowing in the RCS prior to entering the reactor core. The concentration after mixing depends upon the relative flow rates in the RCSx fr;- th: U"!, and in the Safety Injection System. The variation of mass flow rate in the RCS due to water density changes is included in the calculation as is the variation of flow rate in the Safety Injection System due to changes in the RCS pressure.

The Safety Injection System flow calculation includes the line losses in the system as well as the pump head curve.

Figures 15.1.5-5 through 15.1.5-7 show the salient parameters for case b, which -

corresponds to the case discussed above with additional loss of offsite power at the time the safety injection signal is generated. The Safety Injection System delay time includes 10 seconds to start the diesel in addition to 10 seconds to start the charging pump and open the valves. . Criticality is achieved later and the core power increase is slower than in the similar case with offsite power available. The ability of the emptying steam generator to extract heat from the RCS is reduced by the decreased flow in the RCS. The peak power remains well below the nominal full power value.

It should be noted that following a steam line break only one steam generator blows down completely. Thus, the remaining steam generators are still avail-able for dissipa. tion of decay heat after the initial transient is over. In s_

the case of loss of offsite power this heat is removed to the atmosphere via the steam line safety valves.

Marcin to Critical Heat Flux A DNB analy;is was performed for both of these cases. It was found that both cases has a min,imum DNBR greater than the limit value. '

Affect of Continued Auxiliary Feedwater Addition A supplementary analysis was performed to determine the potential for unacceptable worsening of the reactor return-to power as a result of continued addition of auxiliary feedwater following a main steam line break. The main steam line break transient as analyzed here should be insensitive to continued auxiliary feedwater addition since the limiting core conditions, as described above, occur within the first minute due to the initial high cooldown rate.

During this time the primary to secondary heat transfer rate from the blowdown of the initial steam generator water inventory is several orders of magnitude greater than the rate due to the additional auxiliary feedwater, even when runout flow is assumed. The supplementary analysis, assuming auxiliary feedwater at runout flow conditions as described in Section 6.2.1.3.12, was evaluated in Reference 6. This evaluation found that the transient was insensitive to continued auxiliary feedwater addition, and, therefore, that the main analysis above remained bounding.

15.1-16 1984 update

15.1.5.3 Env4 ronmental Consecuences The postulated acciden.ts involvinj release of steam from the secondary sys-tem do not result in a release of radioactivity unless there is leakage from the RCS to the secondary system in the steam generators. A conservative analysis of the potential offsite doses resulting from this accident is pre-sented considering equilibrium operation based upon 1 percent defective fuel and a 1 gpm steam generator leak rate prior to the postu!ated accident.

The primary and secondary coolant activities correspond to limits set by Tech-nical Specifications prior to the accident.

The following assumptions and parameters are used to calculate the activity release and offsite dose for a postulated steam line break:

1. Prior to the accident, an equilibrium activity of fission products ex-ists in the primary and secondary systems caused by primary-to-secondary leakage in steam generators.

2. The total primary-to-secondary l'eak rate is 1.0 gpm, with 0.25 gpm in each steam generator during the first 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. After 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />, no leakage is assumed in the faulted steam generator a.,d 0.333 gpm per generator in the non-faulted steam generators.

3. Offsite power is lost when the steam line break occurs.

{. 4. The steam release for the defective steam generators terminates in 2 hours2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br />. The steam release from the non-faulted steam generators con-tinues for eight hours.

5. All noble gases which leak to the secondary side are released via the steam release.

6. The iodine partition factor during the accident is 0.1.

7. Other assumptions are listed in Table 15.1.5-2. -

• Based on the model described in Appendix 15A, the thyroid and whole body doses are calculated at the exclusion area boundary and the low population zone. The results are presented in Table 15.1.5-2. The doses are within the limits of 10 CFR 100.

15.1-17 1984 UPd 8t*

4 REFERENCES FOR SECTION 15.1

1. L . . . . . , T . ',, . T . , m e '

o..,. "LC.PT.O.:, C.. C ....,... ..", ,, CAT 7:-:7 , .'... ;;::.

2. Chelemer, H., Boman, L. H., Sharp D. R., " Improved Thermal Design Procedures," WCA?-8567, July 1975.

3. " Westinghouse Anticipated Transients Without Trip Analysis", WCAP-8330, August 1974.

4. Moody, F. S., " Transactions of the ASME, Journal of Heat Transfer",

Figure 3, page 134, February 1965.

5. Hollingsworth, S. D. and Wood, D. C., " Reactor Core Response To Excessive Secondary Steam Releases", WCAP-9226 (Proprietary) and WCAP-9227 (Non-Proprietary), January 1978.

6. Letter, T. M. Novak (NRC) to H. B. Tucker, October 14, 1982.

Enclosure:

Safety Evaluation Report of Duke Power Co. response to I. E.Bulletin 80-04. .

Bu.enett , T.W.T. , et a.l ., " LoFTRAh! Cocle l)escriphon," g.

Wc AP - 7907 - P- A (Proprietory), WCAP 1907-A k (Non- Proprietary) , Aprii 1984.

i I

l 15.1-18 198t* UPd ate i

Table 15.1.2-1 (Page 2 of 2) ,

Time Secuence of Events for Incidents Which Cause an Increase In Heat Removal By The Secondary System Accident Event Time (sec.)

Inadvertent Opening of Inadvertent opening of 0 a Steam Generator Relief one main steam safety valve or Safety Valve Pressurizer empties Le' kk 19 1 2,400. ppm boron reaches core  ::: 22.0 Steam System Piping Failure

1. With offsite power Steam line ruptures 0 available Pressurizer empty -44. ll Criticality attained iH9 13

-ib4GG ppm boron reaches core -se 20 1900 , ,

2. Without offsite power Steam line ruptures 0

. Pressurizer empty Wbs 12.

Criticality attained War 16 ,

2,000 ppm boron reaches core 4 28 lC)00 l

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