ML20198H362
| ML20198H362 | |
| Person / Time | |
|---|---|
| Site: | Callaway  |
| Issue date: | 01/28/1986 |
| From: | Schnell D UNION ELECTRIC CO. |
| To: | Harold Denton Office of Nuclear Reactor Regulation |
| References | |
| ULNRC-1247, NUDOCS 8601310011 | |
| Download: ML20198H362 (99) | |
Text
_
f seu s
UNION ELECTRIC COMPANY 1901 Crctiot street. St. Louis January 28, 1986 Donald F. Schnell Vice Preddent Mr. Harold R. Denton, Director Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Washington, D.C.
20555
Dear Mr. Denton:
ULNRC-1247 DOCKET NUMBER 50-483 CALLAWAY PLANT LARGE BREAK LOCA ANALYSIS
References:
- 1) ULNRC-1207 dated 11/15/85
- 2) NRC letter from B. J. Youngblood to D.
F.
Schnell dated 12/30/85
- 3) NRC letter from B. J. Youngblood to D.
F.
Schnell dated 1/10/86 Reference 1 transmitted an application for a reload license amendment for Callaway Cycle 2, which included a large break Loss of Coolant Accident (LOCA) analysis based on the Westinghouse BASH model.
In response to Reference 2 which notes that NRC has not yet completed its generic review of BASH, a new large break LOCA analysis has been completed using the Westinghouse BART model.
The BART model is the most recent Westinghouse, large break LOCA model which has received NRC approval.
The BART analysis is attached herewith and is identified as Attachment G to the reload license amendment for Cycle 2 (Reference 1).
Attachment G should be added to your copics of the reload submittal.
The BART analysis used the same input assumptions (i.e., uprated power and 10% steam generator tubes plugged) as the previous BASH analysis and should be used for review and approval of the Cycle 2 amendment request.
In addition, it is requested that NRC review and approve the BASH analysis as soon as practical following generic approval of WCAP-10266.
With submittal of the attached BART analysis, the review and approval schedule for the BASH analysis can proceed independent of the Cycle 2 amendment request.
The BASH analysis is contained in Reference 1 as Attachment D and will provide additional operating margin.
Also attached herewith and identified as Attachment E, Revision 1 is a revised Significant Hazard Evaluation.
The revised evaluation is in response to Reference 3 and updates the
\\
evaluation based on the BART analysis versus the BASH analysis.
8601310011 860120 g
PDR ADOCK 05000403 P
PDR McAng Address-RQ Box 149, St. Louis MO 63166
e %
No additional Technical Specification changes are necessary, l
other than those previously identified in Reference 1.
This submittal closes Callaway License Condition 2.C.13 (NPF-30).
Continued NRC attention to the Callaway Cycle 2 reload amendment request is appreciated.
As you are aware, our 6-week refueling outage commences on March 1, and Cycle 2 is expected to commence in early April.
If there are any questions please contact us.
Very truly yours, Donald F.
Schnell DS/bjk Attachments: Attachment G - LOCA Accident Analysis (BART)
Attachment E, Rev. 1 - Significant Hazard Evaluation
u o
s STATE OF MISSOURI )
)
Donald F.
Schnell, of lawful age, being first duly sworn upon oath says that he is Vice President-Nuclear and an officer of Union Electric Company; that he has read the foregoing document and knows the content thereof; that he has executed the same for and on behalf of said company with full power and authority to do so; and that the facts therein stated are true and correct to the best of his knowledge, information and belief.
By jf Donald F.
Schnell Vice President Nuclear SUBSCRIBED and sworn to before me this d[
day of m
198 d4h@
BARBARA (PFAFF[
NOTARY PUBt10, STATE CF MISSOURI MY COMMISSION EXPlRES APRIL 22, 1989 ST. LOUIS COUNTY s
e s
cc:
Gerald Charnoff, Esq.
Shaw, Pittman, Potts & Trowbridge 1800 M.
Street, N.W.
Washington, D.C.
20036 Nicholas A.
Petrick Executive Director SNUPPS 5 Choke Cherry Road Rockville, Maryland 20850 G. C. Wright Division of Projects and Resident Programs, Chief, Section lA U.S. Nuclear Regulatory Commission Region III 799 Roosevelt Road Glen Ellyn, Illinois 60137 Bruce Little Callaway Resident Office U.S. Nuclear Regulatory Commission RR#1 Steedman, Missouri 65077 Paul O'Connor Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Mail Stop 316 7920 Norfolk Avenue Bethesda, MD 20014 Ron Kucera, Deputy Director Department of Natural Resources P.O. Box 176 Jefferson City, Missouri 65102
b bcc:
34 56-0031. 2.1 (103 6) w/a.
3456-0021.6 3456-0756.1 Nuclear Date DFS/ Chrono D.
F.
Schnell J.
F. McLaughlin J.
E.
Birk W. II. Weber F.
D.
Field R. J. Schukai M. A. Stiller U/a.
S. E. Miltenberger J/*
II. Wuertenbacher D. W. Capone A. C. Passwater M/*
T.
II. McFarland P. Wendling/a.
R.
E.
Shafer w D.
D. J. Walker w/
(Bechtel)/
G. P. Rathbun (KG&E) w J. II. Smith G56.37 (CA-460) u/a.
Compliance (J.
E. Davis)w/a.
l NSRB (Sandra Auston) 6 6. f /A f
h o
r 3
ULNRC-1247 Attachment G LOCA ACCIDENT ANALYSIS USING THE BART CODE
15.6.5 LOSS-0F-COOLANT ACCIDENTS RESULTING FROM A SPECTRUM 0F POSTULATED PIPING BREAKS WITHIN THE REACTOR COOLANT PRESSURE BOUNDARY 15.6.5.1 Identification of Causes and Freauency Classification A Loss-of-Coolant Accident (LOCA) is the result of a pipe rupture of *he RCS pressure boundary.
For the analyses reported here, a major pipe break (large break) is defined as a rupture with a total cross-sectional area equal to or 2
greater than 1.0 square foot (ft ).
This event is considered an ANS Condition IV event, a limiting fault, in that it is not expected to occur during the lifetime of the plant but is postulated as a conservative design basis. See Section 15.0.1 for a discussion of Condition IV events.
i J
l 8874Q:lD/012486 15.6-4a
A cin:r pipe break (small break), as considered in this section, is defined as a rupture of the reactor coolant pressure b:undary with a total 2
cross-sectional area less than 1.0 ft in which the normally operating charging system flow is not sufficient to sustain pressurizer level and pressure. This is considered a Condition III event, in that it is an inf requent f ault which ney occur during the life of the plant. See Section 15.0.1 for a discussion of Condition III events.
The Acceptance Criteria for the LOCA are described in 10CFR50.46 (Reference 4) as follows:
1.
The calculated peak fuel element clad temperature is below the requirement of 2200*F.
2.
The am6udt of fuel element gladding that reacts chemically with water or steam does not exceed 1 percent 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 amenable to cooltag.
The localized cladding oxidation limit of 17 percent is not exceeded during or af ter quenching.
4.
The core remains amenable to cooling during and af ter the break.
5.
The core temperature is reduced and decay heat is removed for an extended period of time. This is required to remove the heat from the long lived radioactivity remaining in the core.
These criteria were established to provide significant margin in Emergency Core Cooling System (ECCS) performance following a LOCA.
Reference 3 includes a study of the probability of the occurrence of RCS pipe failures.
2 In all cases, small breaks (less than 1.0 ft ) yield results with more margin to the Acceptance Criteria limits than large breaks.
88740:10/012486 15.6-5
B 15.6.5.2 Seauence of Events and Systems Operations
)
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 signal is generated when the appropriate setpoint is reached. These countermeasures will limit the consequences of the accident in two ways; Reac' or trip and borated water injection complement void formation in the t
a.
core and cause a rapid reduction of nuclear power to a residual level corresponding to the delayed fission and fission product decay heat.
However, no credit is taken during the LOCA blowdown for negative reactivity due to boron content of the injection water.
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 the fluid medium for heat transfer from the core and prevents excessive clad temperatures.
Description of a large Break LOCA Transient The sequence of events following a large break LOCA is presented in Figure 15.6-3.
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, film boiling, and forced convection to steam as the major heat transfer mechanisms.
The heat transfer between the RCS 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 8874Q:10/012486 15.6-6
atmosph;ric relief and/or 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 safety injection signal actuates a feedwater isolation signal which isolates nornal feedwater flow by closing the main feedwater isolation valves and also initiates auxiliary feedwater flow by starting the auxilia'.y feedwater pumps. The secondary flow aids in the reduction of RCS pressure.
When the RCS depressurizes to approximately 600 psia, the cold leg accumulators begin to inject borated water into the reactor coolant loops.
Since the loss of of f site power is assumed, the reactor coolant pumps are assumed to trip at the time of reactor trip during the accident. The effects of pump coastdown are includcd in the blowdown analyses.
The blowdown phase of the transient ends after the RCS pressure (initially assumed at a nominal 2280 psia) 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 water injected 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 (culled bottom of core recovery time).
The reflood phase of the transient is defined as the time period lasting from the end-of-refill until the reactor vessel has been filled with water to the extent that the core temperature rise has been 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 centrifugal charging, safety injection, and RHR 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 15.6-3.
8874Q:10/012486 15.6-7
Ccntinued 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 generation. After the water level of the refueling water storage tank 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 residual heat removal pumps and returned to the RCS cold legs. The Containment Spray System continues to operate to further reduce containment pressure. Approximately 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-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.
Description of Small Break LOCA Transient Ruptures of small cross section will cause expulsion of the coolant at a rate which can be accommodated 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 charging pump is 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 pressurizer 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 8874Q:10/012486 15.6-8 1
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:
1.
Reactor trip and borated water injection complement void formation in the core and cause a rapid reduction of nuclear power to a residual level corresponding to the delayed fission and fission product decay heat.
2.
Injection of borated water ensures sufficient flooding of the core to prevent excessive clad temperatures.
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 fission product 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 depending on the relative temperatures.
In the case of continued heat addition to the secondary, secondary system pressure increases and steam relief via the atmospheric relief and/or safety valves may occur. Makeup to the secondary side is automatically provided by the auxiliary feedwater pumps. The safety injection signal isolates normal feedwater flow by closing the main feedwater isolation valves and initiates auxiliary feedwater flow by starting the auxiliary feedwater pumps.
The secondary flow aids in the reduction of Reactor Coolant System pressure.
When the RCS depressurizes to approximately 600 psia, the cold leg accumulators begin to inject borated water into the reactor coolant loops.
The vessel mixture level starts to move up to cover the fuel before the accumulator injection for most breaks.
For all breaks, the accumulator injection provides enough water supply to bring the mixture level up to the upper plenum region where it is maintained.
Due to the loss of offsite power assumption, the reactor coolant pumps are assumed to be tripped at the time of reactor trip during the accident and the effects of pump coastdown are included in the blowdown analyses.
88740:10/012486 15.6-9
15.6.5.3 Analysis of Effects and Crnseauences 15.6.5.3.1 Method of Analysis The requirements of an acceptable ECCS evaluation model are presented in Appendix K of 10CFR50 (Reference 4).
The requirements of Appendix K regarding specific model features were met by selecting models which provide a significant overall conservatism in the analysis.
The assumptions made pertain to the conditions of the reactor and associated safety system equipment at the time that the LOCA occurs and include such items as the core peaking factors, the containment pressure, and the performance of the ECCS.
Decay heat generated throughout the transient is also conservatively calculated as required by Appendix K of 10 CFR 50.
Large Break LOCA Evauation Model The analysis of a large break LOCA transient is divided into three phases:
1) blowdown, 2) refill, and 3) reflood. There are three distinct transients analyzed in each phase:
- 1) the thermal-hydraulic transient in the RCS, 2) the pressure and temperature transient within the containment, and 3) the fuel and 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 of the LOCA.
The description of the various aspects of the LOCA analysis methodology is given in References 5,10,11,16 and 17. These documents describe the major phenomena modeled, the interfaces among the computer codes, and the features of the codes which ensure compliance with the Acceptance Criteria. The SATAN-VI, WREFLOOD, COCO, and LOCTA-IV codes which are used in the LOCA analysis, are described in detail in References 6 through 9.
Modifications to these codes are specified in References 10 through 12.
The BART code is described in References 16 and 17.
These 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 and the WREFLOOD computer code is used to 8874Q:10/012486 15.6-10
calculate this transient during the refill and reflood phases of the accident. The BART computer code is used to calculate the fluid and heat transfer conditions in the core during reflood. The C0C0 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.
The large break analysis was performed with the approved December,1981 version of the Evaluation Model (Reference 11), with the approved 1984 version of BART (Reference 16).
SATAN-VI is used to calculate the RCS pressure, enthalpy, density, and the mass and energy flow rates in the RCS, as well as steam generator energy 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 and 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. Also at the end-of-blowdown, the mass and energy release rates durin; olowdown are transferred to the C0C0 code for use in the determination of the containment pressure response during these phases of the LOCA. Additional SATAN-VI output data from the end-of-blowdown, including the core pressure, and the core power decay transient, are input to the LOCTA-IV code.
With input f rom 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 addition to the containment through the break. Since the mass flow rate to the containment depends upon the core flooding rate and the local core pressure, which is a function of the containment backpressure, the WREf LD00 and COC0 codes are interactively linked. With input and boundary conditions from WREFLOOD, the mechanistic core heat transfer model in BART calculates the hydraulic and heat transfer conditions in the core during reflood.
LOCTA-IV is used throughout the 8874Q:10/012486 15.6-11
analysis of the LOCA transient to calculate the fuel clad temperature and metal-water reacticn of the hottest rod in the core. A schematic representation of the computer code interfaces for large break calculations is i
shown in Figure 15.6-4.
The C0C0 code is a mathematical model of the containment.
C0C0 is run using mass and energy releases to the containment provided by SATAN and WREFLOOD.
COCO is described in detail in Reference 8.
Calculated pressures from C0C0 are presented in Table 15.6-4.
The LOCTA-IV code 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 9.
In the 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."
8874Q:lD/012486 15.6-12 1
~
In NOTRUMP, the RCS is nodalized into volumes intercennected by ficwpaths.
Th2 brok n loop is modeled explicitly with the intact loops lumped into a second loop. The transient behavior of the system is determined f rom the governing conservation equations of mass, energy and momentum applied throughout the system. A detailed description of NOTRUMP is given in References 14 and 18.
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 9) 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.
Schematic representations of the computer code interfaces are given in Figure 15.6-5.
The small-break analysis was performed with the approved Westinghouse ECCS Small Break Evaluation Model (References 9, 14, and 18).
15.6.5.3.2 Input Parameters and Initial Conditions Larae Break Input Parameters and Initial Conditions Table 15.6-2 lists important input parameters and initial conditions used in the large break analyses. Both maximum and minimum safeguards ECCS capability and operability has been assumed in this analysis to determine the worst case.
8874Q:10/012486 15.6-13
~
Small Break Input Parameters and Initial Conditions Table 15.6-2 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 d
analyses are shown in Figures 15.6-45 and 15.6-46.
Note that the power shape shown in Figure 15.6-45 is the most limiting for small break transients rather than a chopped cosine power shape with a peak at the center (Reference 14).
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.
This includes the assumption of a 12 second delay in the startup of the diesel generator.
For these analyses, the SI delivery considers pumped injection flow which is depicted in Table 15.6-9 as a function of RCS pressure. This table represents injection flow from the centrifugal charging (CCP) and safety injection (SI) 1 pumps. The 25 second delay includes time required for diesel startup and loading of the CCP and SI 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 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 NSSS design thermal power. The core thermal transient analyses are performed with the LOCTA-IV code using 102% of reactor core design thermal power.
15.6.5.3.3 Results Large Break Results Based on the results of the LOCA sensitivity studies (Reference 15), the limiting large break was found to be the double ended cold leg guillotine S
8874Q:10/012486 15.6-14
(DECLG). Therefere, enly the DECLG break is censidered in the large break ECCS perfermance analysis.
Calculaticns were perfcrmed for a range of Moody break discharge coefficients under minimum safeguards conditions.
The results of these calculations are summarized in Tables 15.6-5 and 15.6-6.
Figures 15.6-6 through 15.6-44 present the parameters of principal interest from the large break LOCA analyses. Transients of the following parameters are presented for each discharge coefficient analyzed, and where appropriate for the limiting discharge coefficient (C = 0.4 DECLG) with maximum D
safeguards assumed.
Figure 15.6-6 The following quantities are presented at both the clad through burst location and at the hot spot (location of maximum Figure 15.6-17 clad temperature) on the hottest fuel 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 15.6-18 The system pressure shown is the calculated pressure in
~
through the core. Core flow rates are also presented.
Figure 15.6-23 Figure 15.6-24 These figures show the hot spot clad temperature transient through and the clad temperature transient at the burst location.
Figure 15.6-31 The fluid temperature shown is also for the hot spot and burst location.
Figure 15.6-32 These figures show the core reflood transient.
through Figure 15.6-39 8874Q:lD/012486 15.6-15
Figure 15.6-40 Thesa figures show the cold leg accumulator delivery through during blowdown and the pumped safety injection during Figure 15.6-44 reflood.
The maximum clad temperature calculated for a large break is 2153*F, which is less than the Acceptance Criteria limit of 2200*F of 10CFR50.46. The maximum local metal-water reaction is 7.68 percent, which is well below the embrittlement limit of 17 percent as required by 10CFR50.46. The total core metal-water reaction is less than 0.3 percent for all breaks, as compared with the 1 percent criterion of 10CFR50.46, and the clad 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. A small impact of cross flow for transition core cycles is conservatively evaluated as at most a 10'F effect, which can be accommodated in the margin to 10 CFR 50.46 limits.
Small-Break Results As noted previously, the calculated peak clad temperature resulting from a small-break LOCA is less than that calculated for a large break. A range of small-break analyses is presented which establishes the limiting break size.
It has been shown that a 2" break does not show any core uncovery (Reference 14) therefore this analysis has not been performed for this report. The results of these analyses are summarized in Tables 15.6-7 and 15.6-8.
Figures 15.6-47 through 15.6-55 present the principal parameters of interest for the small-break ECCS analyses.
For all cases analyzed, the following transient parameters are presented:
a) RCS presssure b) Core mixture height c) Hot spot clad temperature 8874Q:10/012486 15.6-16
Far the li;iting break analyzsd (3 inch), the following additicnal transient parameters are presented (Figures 15.6 --56 through 15.6 --58):
a.
core steam flow rate b.
core heat transfer coef ficient c.
hot spot fluid temperature The maximum calculated peak cladding temperature for the small breaks analyzed is 1299'F. The maximum local metal-water reaction is 0.23% which is well below the value observed in large breaks.
Further, the total core metal-water reaction is less than 0.3% for all cases analyzed and the clad temperature transients turn around 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.
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.
15.6.5.4 Radiolooical Consequences See Attachment F of Reference 13.
15.6.6 A NUMBER OF BWR TRANSIENTS This section is not applicable to Callaway.
88740:10/012486 15.6-17
15.
6.7 REFERENCES
1.
Burnett, T. W.
T., et. al., "LOFTRAN Code Description", WCAP-7907-P-A (Proprietary), WCAP-7907-A (Non-Proprietary) April 1984.
2.
- Chelemer, H., Boman, L.
H., Sharp, D. R., " Improved Thermal Design Procedures", WCAP-8587, July 1975.
3.
" Reactor Safety Study - An Assessment of Accident-Risk in U.S. Commercial Nuclear Power Plants," WASH-1400, NUREG-75/014, October 1975.
4.
" 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.
5.
Bordelon, F. M., Massie, H. W. and Borden, T. A.,
" Westinghouse ECCS Evaluation Model-Summary", WCAP-8339, (Non-Proprietary), July 1974.
6.
Bordelon, F.
M., et al., " SATAN-VI Program: Comprehensive Space Time Dependent Analysis of Loss of Coolant" WCAP-8302, (Proprietary) June 1974, and WCAP-8306, (Non-Proprietary), June 1974.
^
7.
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 8.
Bordelon, F. M. and Murphy, E.
T., " Containment Pressure Analysis Code (C0CO)," WCAP-8327 (Proprietary) and WCAP-8326 (Non-Proprietary), June 1974.
i 9.
Bordelon, F. M., et, al., "LOCTA-IV Program: Loss of Coolant Transient Analysis", WCAP-8301, (Proprietary) and WCAP-8305, (Non-Proprietary), June 1974.
l 8874Q:10/012486 15.6-18 j
- 10. 8;rdelon, F. M., et al., "Westingh*use ECCS Evaluatien Model -
Supplementary Infcrmatien," WCAP-8471-P-A, April,1975 (Proprietary) and WCAP-8472-A, April, 1975 (Non-Proprietary).
- 11. " Westinghouse ECCS Evaluation Model, 1981 Version," WCAP-9220-P-A, Rev. 1 (Proprietary), WCAP-9221-A, Rev.1 (Non-Proprietary), February,1982.
- 12. Letter from C. Eicheldinger (Westinghouse) to D. B. Vassallo (NRC), Letter Number NS-CE-924 dated January 23, 1976.
- 13. Letter ULNRC-1207, dated November 15, 1985.
- 14. Lee, N., Rupprecht, S. D., Schwarz, W. R., Tauche, W. O., " Westinghouse Small Break ECCS Evaluation Model Using the NOTRUMP Code " WCAP-10054-P-A (Proprietary) and WCAP-10081-A (Non-Proprietary) August 1985.
- 15. Salvatori, R., "Westinghoase Emergency Core Cooling System - Plant Sensitivity Studies", WCAP-8340, (Proprietary) July 1974.
- 16. Young, M., et. al., "BART-1A: A Computer Code for the Best Estimate Analyzed Reflood Transients", WCAP-9561-P-A, 1984 (Westinghouse Proprietary).
- 17. Chiou, J. S., et al., "Models for PWR Reflood Calculations Using the BART Code," WCAP-10062.
- 18. Meyer, P. E.,
"NOTRUMP, A Nodal Transient Small Break and General Network Code", WCAP-10079-P-A (Proprietary) and WCAP-10080-A (Non-Proprietary),
August 1985.
8874Q: 10/012486 15.6-19
TABLE 15.6-2 Input Parameters Used in the LOCA Analyses Parameter Larae Break Small Break Reactor Core Design Thermal Power * (Mwt) 3565 3579 Peak Linear Power (kW/f t)*
13.47 12.77 at 6.0 ft at 10.0 f t Total Peaking factor (Fgy) at peak 2.32 2.20 Power Shape Chopped See Figure
- Cosine, 15.6-45 Fz " I 497 Fuel Assembly Array 17 X 17 17 X 17 Optimized Optimized Nominal Cold Leg Accumulator 850 850 3
Water Volume (ft / accumulator)
Nominal Cold Leg Accumulator 1364 1364 3
Tank Volume (ft / accumulator)
Minimum Cold leg Accumulator 600 600 Gas Pressure (psia)
Pumped Safety Injection Flow See Table See Table 15.6-3 15.6-9 Steam Generator Initial Pressure (psia) 935.0 935.0 Steam Generator Tube 10 10 Plugging Level (%)
Containment Parameters (See FSAR Section 6.2)
Initial Flow In Each Loop (lb/sec) 9710.7 Vessel Inlet Temperature (*F) 554.93 Vessel Outlet Temperature (*F) 619.47 Reactor Coolant Pressure (psia) 2280.
Two percent is added to this power to account for calorimetric error.
Reactor coolant pump heat is not modeled in the large break LOCA snalyses.
8874Q:10/012486
TABLE 15.6-3 Safety In.ioction Pumped Flowo Assumed for Breaks Greater Than or Equal to 10 Inches MIMIMUM SAFEGUARDS:
Pressure SI Flow (osia)
(lb/sec) 14.7 505.03 34.7 429.24 54.7 350.76 74.7 274.05 114.7 159.01 214.7 102.01 614.7 79.73 1014.7 53.06 2014.7 0.0 MAXIMUM SAFEGUARDS:
Pressure SI Flow (psia)
(1b/sec) 14.7 1237.26 34.7 1149.26 J
54.7 1056.86 74.7 953.89 94.7 851.89 114.7 732.36 214.7 171.0 314.7 165.9 1014.7 124.4 2014.7 47.16 j
Sum of the Centrifugal Charging, Safety Injection and Residual Heat Removal pump flows, 8874Q:10/012486
TA8LE 15.6-4 Containment Pressure (Psia)
Minimum Safeauards C
= 0.4 C ) = 0.6 C ) = 0.8 Time ECLG
)ECLG
)ECLG 0.0 14.7*
14.7*
14.7*
10.0 31.42 33.90 35.91 20.0 35.94 37.65 38.26 30.0 36.45 35.85 35.51 40.0 34.79 33.93 33.80 59.0 32.12 31.90 31.92 79.0 31.03 31.17 31.22 99.0 29.84 30.17 30.23 149.0 27.63 27.97 28.03 199.0 25.77 26.03 26.08 Maximum Safeauards CD = 0.4 Time DECLG 0.0 14.7*
10.0 31.42 20.0 35.94 30.0 36.44 40.0 34.74 59.0 32.02 79.0 30.76 99.0 29.24 149.0 25.99 199.0 23.29
- Assumed initial conditions for the containment are 14.7 psia and 90*F.
88740:10/012486
TABLE 15.6-5 (ShGet 1)
Lara9 Break LOCA Results Fuel Claddina Data Minimum Safeguards CD = 0,8 CD = 0.6 CD = 0,4 DECLG DECtG DECLG RESULTS Peak Clad Temperature (*F) 1986 2079 2153 Peak Clad Temperature Location (ft) 7.00 7.00 7.00 local Zr/H O Reaction (maximum %)
2.77 3,68 7.68 2
Local Zr/H O Reaction 7.00 7.00 6.50 neximum reaction (ft), Location for 2
Total Zr/H O Reaction, (%)
<0.30
<0.30
<0.30 2
Hot Rod Burst Location, (ft) 6.00 6.00 6.50 8874Q:10/012486
TABLE 15.6-5 (Sheet 2)
Laroe Break LOCA Results Fuel Claddina Data Maximum Safeauards CD = 0.4 DECLG RESULTS Peak Clad Temperature (*F) 1920 Peak Clad Temperature Location (ft) 7.00 Local Zr/H O Reaction (maximum %)
3.04 2
Local Zr/H O Reaction, Location for maximum 6.50 reaction (2ft)
Total Zr/H O Reaction, (%)
<0.30 2
Hot Rod Burst Location, (ft) 6.50 8874Q:1D/012486
TABLE 15.6-6 (Sheet 1)
Larae Break LOCA Time Sequence of Events Minimum Safeguards CD = 0.4 CD = 0.6 CD = 0.8 DECLG DECLG DECLG Time (sec)
(sec)
(sec)
Break 0
0 0
Reactor Trip Signal 0.765 0.741 0.727 SI-Signal 1.67 1.36 1.18 Intact Loop Accumulator Injection 20.4 15.5 12.9 Pump Injection 28.67 28.36 28.18 End of Bypass 39.753 30.328 27.59 End of Blowdown 39.753 30.353 27.59 BOC Time 55.34 44.267 41.627 Intact Loop Acc Empty 61.54 54.801 51.58 Hot Rod Burst Time 69.60 48.00 51.20 Peak Clad Temperature (PCT) Time 117.20 106.20 100.40 l
8874Q:10/012486
v-,
TABLE 15.6-6 (Sheet 2)
Laroe Break LOCA Time Seauence of Events Maximum Safeguards C = 0.4 0
DECLG (sec)
Break 0.0 Reactor Trip Signal 0.741 S-1 Signal 1.36 Intact Loop Accumulator Injection 15.5 Pump Injection 28.36 End of Bypass 30.32 End of Blowdown 30.351 BOC Time 54.464 Intact Loop Accumulator Empty 62.670 Hot Rod Burst Time 69.80 Peak Clad Temperature (PCT) Time 121.60 8874Q:10/012486 4
TABLE 15.6-7 Small Break LOCA Time Seauence of Events 3 in 4 in 6 in (Sec)
(Sec)
(Sec)
Break 0.0 0.0 0.0 Reactor Trip Signal 8.42 5.13 3.01 Safety Injection Signal 17.39 13.55 9.81 Safety Injection Begins 42.39 38.55 34.81 Top of Core Uncovered 1059 673 338 Cold Leg Accumulator Injection N/A 902 385 Peak Clad Temperature Occurs 1642 947 450 Top of Core Covered 2100 1178 466 l
t I
i 1
i 8874Q:10/012486 l
1
TABLE 15.6-8 Small Break LOCA Results Fuel Claddina Data 3 in 4 in 6 in RESULTS Peak Clad Temperature (*F) 1299 1184 969 Peak Clad Temperature Location (ft) 12 12 11.75 Local Zr/H O Reaction (maximum %)
0.23 0.084 0.071 2
Local Zr/H O Reaction, Location for 12 12 11.25 2
maximum reaction (ft)
Total Zr/H O Reaction, (%)
<0.3
<0.3
<0.3 2
Hot Rod Burst Time, (sec)
N/A N/A N/A Hot Rod Burst Location, (ft)
N/A N/A N/A 4
1 8874Q:10/012486
TABLE 15.6-9 Safety Injection Flow For Small Break Analysis
- Pressure (Psia)
SI Flow (1b/sec) Per Loop 14.7 111.6 114.7 107.47 214.7 103.2 314.7 98.7 414.7 94.07 514.7 89.3 614.7 84.36 714.7 79.26 814.7 73.94 914.7 68.36 1014.7 62.41 1114.7 55.94 1214.7 48.58 1314.7 39.24 1414.7 22.99 1514.7 15.31 1614.7 12.61 1714.7 9.73 1814.7 7.43 1914.7 4.91 2014.7 1.98 l
2114.7 0.0 Sum of the Centrifugal Charging and Safety Injection pump flows.
8874Q:10/012486
A BREAM OCCURS 15366.1 REACTOR TRIP SIGNAL (COWAENSATED PRESSURIZER PRESSURE)
PUMPEO SAFETY INJECTION SIGNAL (HI-I CONT. PRESS OR LO PRESSURIZEA P)
L PUWPED SAFETY INJECTION BEGINS (ASSUMING OFFSITE POWER AVAILABLE)
O W
COLD LEG ACCUWULATOR INJECTION W
N
~
CONT.> EAT REu3 VAL SYSTEM INITIATION (ASSUMIN3 OFFSITE POWER AVAILABLE)
PUWREO SAFETY INJECTION BEGINS ( ASSUMING LOSS OF OFFSITE POWER) yr END OF BLOr.jown REFILL 3r BOTTOu OF CORE RECOVERY R
COLD LEG ACCUuULATORS EupTY L
0 0
0 CORE QUENC&EO h
O SWITCH TO COLO LEG RECIRCULATION ON RwST LOW-LOW-1 LEVEL ALARM N
(AUTOMATIC AND MANUAL ACTIONS)
G T
SWITCH TO LONG-TERH RECIRCULATION (MANUAL ACTION)
E l
R M
C t
O i
O L
LONG-TERM HOT LEG RECIRCULATION (MANUAL ACTION) g..
N j
G Y
Figure 15.6-3 Sequence of Events for Large Break Loss of Coolant Analysis
I i-BLOWDOWN A END OF BIDWDOWN REFILL,REFIDOD t
4 4
(EOB)
MMA UA
=
g dl FUEL ROD THERNAL,
ROD GEOMETRY g HEAT TRANSFER COEFFICIENT NECHANICAL CONDITIONS Y
DURING BIDWDOWN BART HEAT TRANSFER CONDITIONS DURING REFIDOD JL l
l SATAN I
I CORE INLET FIDW,
I INLET ENTHALPY MASS, ENERGY RELEASE l
RCS, CORE THERNONYDRAULIC g
CONDITIONS DURING BIDWDOWN RCS I
CONDITIONS l
AT EOB l
(_---__-----l------------_q I
WREFIDOD/ COCO 8
8 8
I I
WREFIDOD g
CAIEUIATES BREAK MASS, ENERGY RELEASE 1
JL I
l I
g I '
g i
i I
COCO e
CAI4ULATES CONTAINMENT PRESSURE I
s E
s L-------______--------____-s Figure 15.6-4 APPROVED WESTINGHOUSE APPENDIX K IDCA EVALUATION MODEL WITH BART 1
w 15366 3 1
N L
0 0
T C
R U
CORE PRESSURE, CORE T
M FLOW, MIXTURE LEVEL A
3 P
AND FUEL ROD POWER HISTORY j
i t
O< TIME < CORE COVERED V
m, 1
1 l
Figure 15.6 5 Code interface Description for Small Break Model
I e
LEGEPO:
PEAK-PEAK CLAD TEW'ERATURE LOCATION BURST-BURST LOCATION AS SHOUP4 IN TABLE 15.6-5 QUALITY OF FLUID
- BURST, 6.00 FT( )
PEAK, 7.00 FT(*)
- 3. 6 3.4 I.2 1.
y q
/
.e "r
9
/
b
.e
/
b
)
JI I
1 5 g
4 r
1 i
d R
0 2
e see se' se2 ies TIME 15CC)
Figure 15.6-6 Fluid Quality, CD = 0.8 DECLG, MINSI
l s
Y QUALITY OF FLUID BURST.
6.50 FT( )
PEAK, 7.00 FT(*)
1.s s.a 1.2 l-1 S
k-t
,g 3
Y W
V 2
k 6
.6 r
I b
/
I, g
i 5
0
.2 e
e see se i,2 ges
?!NC 15EC3 Figure 15.6-7 Fluid Quality, CD = 0.4 DECLG, MINSI tr
i e
~
QUALITY OF FLUID
- BURST, 6.00 FT( )
PEAK, 7.00 FT(*)
1.s 34 1.2
/
\\)J W
/
v
.O l
f 9
/
1 3u
.s g
D
]
I i
m ll r
8..
e
,,e ier ses TIME ISECl Figure 15.6-8 Fluid Quality, CD = 0.6 DECLG, MINSI
.o t
QUALITY OF FLUID BURST.
6.50 FT( )
PEAK. 7.00 Fil*)
s,s 1.e I.2 1.
f:1.
i
.s y
y v
2 o
6
.6 i
s E*
l 1 l 1
2 e
ies
,,a t,,
1es TINC 8SCCI Figure 15.6-9 Fluid Quality,CD=0.4DECLG,MAXSI i
9 me H
tr)
Z H
2:
E o
O O.
.)
U J
P-q>
w O
e m
x 4
o w
s g
II O
O N
~b u
_e o
w C.
/
- e. $
to e
y 7
2 r
~
c 7
(
o N
in so
]
w I
-t~,..
'E
~
~m
%N.
N
_N-
>=
w UO
.J w>
3 2
R 3
R R
R 7
E 1335 E AJ/81)
All3013A EStM
o e
i MASS VELOCITY
- BURST, 6.50 FT( )
PEAK, 7.00 FT(*)
198 l
t
.S M U
A k 68 f
fU
=
^
d
\\
\\\\ l as
\\
h L
US e
- /#
j W
E l
1 -re
-de 188 18' Ig2 gg5 TIME ISECl t
MINSI Figure 15.611 Mass Veiocity, CD = 0.4 DECLG l
8 i
MASS VELOCITY BURST.
6.00 FT( )
PEAK, 7.00 FT(*)
G8 A
t 3..
o T\\
M) d6
\\
w f
3 c%a-1 ge s
I l
/
r--
n ed$
I 38' IB 332 gg5 T]NC (SEC)
Figure 15.6-12 Mass Velocity, CD = 0.6 DECLG, MINSI
i MASS VELOCITY
- BURST, 6.50 FTl )
PEAK, 7.00 FT(*)
les L
e-e,i i
h fl e
d
\\
\\hl
\\
43 s
3 N
5' t A i
1 /
6
=1 as 4
8 les 15 192 105 TIME ISECI i
Figure 15.6-13 Mass Velocity,CD=0.4DECLG,MAXSI e
n 1
I HEAT TRANS. COEFFICIENT
- BURST, 6.00 FT( )
PEAK, 7.00 FT(*)
te5 Y
I
_i, t
L
$ te? Y i
^
s,-
' 1-y sw
\\ /
s.,
t a
I
,v Asw~
G L
r >~
^-
[
k Ll N V
8u p sol G=
E Y
seee 2e se se se ice 12e las ise see 2ee 22e TIT ISECD Figure 15.614 Heat Transfer Coefficient. CD = 0.8 DECLG. MINSI
is HEAT TRANS. COEFFICIENT BURST, 6.50 FT( )
PEAK, 7.00 FT(*)
te5 Y}
ll
{f
~
C a
i g is?
_-a L
5
's l 1 G
I
\\
E l
\\
A A
fj
- ~~Y g
I
=='
W yv 9 sol G
E Y
lee 8
2e de se se see 32, i.e ese i,,
i TIME ISECl 9
Figure 15.6-15 Heat Transfer Coefficient. CD = 0.4 DECLG, MINSI
i t
HEAT TRANS. COEFFICIENT BURST, 6.00 FT( )
PEAK, 7.00 FT(*)
i,s Y
r 1
6 a
\\
I t3 192 I
^
l
~
y s
t i
n m a~~
3~~
s i.,
n
!TI rV.
G L
.C
%m Am
^~-
m E
WiV
'V oe ISI N
5 a
Y l
ISO 1
9 29 49 69 99 199 129 149 ISO 199 299 229 TIMI' ISECl Figure 15.6-16 Heat Transfer Coefficient CD = 0.6 DECLG, MINSI e
8i HEAT TRANS. COEFFICIENT BURST.
6.50 FT( )
PEAK. 7.00 Fil*)
le5 v
y
.f A
5 l[
m si l
- is?
A w
\\
\\
a s
i 1
n a
E
/
\\
M Tal^#'dW C~-
g
, f oV
't v
4 u
Isl i
I Y
se*
e 2e se se se see 12e see ise see 2ee 22e flME' ISEC8 i
Figure 15.6-17 Heat Transfer Coefficient,CD=0.4DECI4,MAXSI i
i
15366 16 0
LD exi>
0
~
DibI O
ci
~
to n o
S w
o m
a O
O 6-4 I
l O
F N
g-E i
0 e
~
e t!I e
O
.6 m
I I
I a
O O
o O
O O
O O
O O
O O
O O
O m
N N
(VISd) 380SS3Bd
,,_,--,--,---,-----,,----,,-y,
,.,-,,,.,,,,--,--p
15366 18 Oin o
i
- H mX 2
z Hx o
d n-a o
8 to a
m w
C.
11 O
O 0
>-4 O
F N
a.
eCU S
E e
o
.a e
}
I 1
I I
o o
o o
o o
o O
O O
O O
m O
m o
m N
N 1
(VISd) 38nSS38d
,-e-
,,,--.,_p---,---m
,,,-.,,,,---.,,-,-_.-_-n---
1536617 Oln i
4 d
+
5 5=
O F) d O
w S
tn a
ed b
cu 1
O F
e g
~
N a.
e 0
R 6
2 O
t g
is I
t i
I I
O O
O O
O O
O O
O O
O O
LD O
tn O
in N
N (Visa) aanssssa j
5$E 05 0,
4 ISN I
M 0
G 3
L
)
C C
E E
D S
8
(
0
=
E DC M
,e I
ta M
0 T R
O 2
w T
lo T
F kO e
r P
B o
\\
O C
l T
1 2
-6 5
1 0
e s
r 1
ug i
F f
Jl O
0 0
0 0
0 0
0 0
0 0
0 0
0 1
0 0
5 5
0 0
7 5
2 2
5 7
oW* oJ ~ WH<E>b'tN i
2,
1
)i
a W
15366-20 o
e O
6 o
+
$x Y
i t
E o
er On y
w a
C W
v.
m
?
w 8
s e
s oF 1
r s
N o
o b
F u-F.
E 4 '- O 3
m R
o 5
m h.
Jm o
8 8
8 o
o a
o a
8 o
o o
m o
o m
m
?
(03s/e1) 3,t y g g g ),,_ 7 c_
-.--- a
A a
15366-21 O
LD
~
H to H
O x
~
PO
(
g O
s LLI u
m E
ii e,
d I
(,1.)
Il 2
o m
o y
O H i
Lk
~
O N
bh[
t O
E O
$u b
4 O
te E
m l
2 l
l O
O O
O O
O O
O O
O O
O O
O O
O LD LD O
O b
LD N
N LD b
I i
(oss/97) 31v8Mola-z l
l I
1 l
, - - - _ _, - _., - -. - _. - -.. _,. _ _, - - - - -.. -,.,.. - -. -, - - - - - - - - - - -. - - - - - - - - - - - -, - - - - - ~ - -
~~ -
n.
i 1
4 l
{
CLAD AVG. TEMP. HOT ROD
- BURST, 6.00 FT( )
PEAK, 7.00 FT(*)
i i
NM 1
i C
W 8
N N
N i
ises - a
~
s w
=
E x
i 1"
N o
i g ses g
I d
\\
1 I
's as as se os see see see ise les aos 22e I
isnc istc I
1, Figure 15.6-24 Peak Clad Temperature, CD = 0.8 DECLG, MINSI
I SN I
M 82 2
G L
)
C
(
6 E
9 D
T 2
F 4
0 0
0 8
=
9 xN D
1 7
C e
K 8
r A
6 u
x \\
1 t
E a
P rep 6
m
)
4 w
lC e
1 T
(
T d
F a
9I 2C 0
1 E 5
S k
I ae 6
E P
6M y
1 T 5
8I 2
T
-6 S
R 5
6 U
1
/
ig 9
B e
ru D
6 F
6 O
3 R
TO e
H.
d P
M E
8 T.
}
2 V
G VA f
g t
D s
e
=
8 A
e 8
e L
s s
5 a
1 I
C
_ meh-O te ! gd
CLAD AVG. TEMP. HOT ROD
- BURST, 6.00 FT( )
PEAK, 7.00 FT(*)
2508 C
A 0 2ees N
1598 6'/
N e
x
,i-s g see d
5 29 48 66 SS 196 126 149 166 ISO 296 226 TINC ISECI Figure 15.6-26 Peak Clad Temperature, CD = 0.6 DECLG, MINSI l
CLAD AVG. TEMP. HOT ROD
- BURST, 6.50 FT( )
PEAK, 7.00 FT(*)
2see C
0 2ees 4
~
x N
N m
isee 8
\\
N E
Y J
t'M V
e 9a g see d
8 e 2e de se se les las see Ise see 2ee 22e TINC ISECI t
Figure 15.6-27 Peak Clad Temperature,CD=0.4DECLG,MAXSI
is FLUID TEMPERATURE
- BURST, 6.00 FT( )
PEAK, 7.00 FT(*)
2ees
- 1758 0
E 1599 1258 I
sees l'
\\
i i f
N N
e-2 s
x o
U 2se 5
25 48 66 96 186 128 146 168 ISO 283 229 TlHE (SCCI t
Figure 15.6-28 Fluid Temperature, CD = 0.8 DECLG, MINSI
i FLUID TEMPERATURE
- BURST, 6.50 FT( )
PEAK, 7.00 FT(*)
2ees
[ 175e m
15ee
/
125e N
tese s
N
^
5 76e h
w see d
25e
'e 2e de se se see les 14e tse see 2ee 22e TIME ESEC1 Figure 15.6-29 Fluid Temperature, CD = 0.4 DECLG, MINSI
{
l
ii l
FLUID TEMPERATURE
- BURST, 6.00 FT( )
PEAK, 7.00 FT(*)
rees C 17se M\\^
^
isee s
'/ %C K
~
sese h
\\
\\
\\
$ sees
/
\\
N a
t e
s U
ih
, f
\\
s
\\
3,,
9 N
S T
2se
~
'e 2e de se se see i2e las Ise les Pee are TIME ISEC1 Figure 15.6-30 Fluid Temper $ture, CD = 0.6 DECLG, MINSI
i i
e l
t l
FLUID TEMPERATURE
- BURST, 6.50 FT( )
PEAK, 7.00 FT(*)
me
- - 17se 5g 1598 bb. _
/
1258 7y7 I
k,%
i.
oc 758 e
v 3'
ase 5
29 49 68 99 198 129 id8 166 ISO 296 229 TIME ISECl Figure 15.6-31 Fluid Temperature,CD=0.4DECIA,MAXSI
't
,ll11 l
ess I
S e
N es I
M 2.
G 8
L 5
C 6
d E
/v D
t s
8 C
e 0
d C
=
N D
E W
e C
R s
s O
O ls D
C ve fI ei Le ec y
st e
r u
t C
x eN i
I 51 M
21 do e
lo f
e e
R 23 es
/
I 5
6 1
e e
r
)
e u
T F
ig s
F
(
L E
e V
1 s
E c
c L
e o
7_
E R
e E
5 s
T e
s.
s e
s.
7 2
2 7
i 2
s A
1 8
W
- $0" 5E'
l.
i WATER LEVEL (FT) 2e.
17.s
/~
sf i s' f
DOWA/COMEK E
/
g 12 5 _pc 0a se.
G"
- 7. s C04 Kn r
W 2.s f e
se see ise 2ee 2se see sse des ase see sse TIME isECl Figure 15.6-33 Reflood Mixture Levels, CD = 0.4 DECLG, MINSI
I WATER LEVEL (FT) 28 17 5 15-A hf MCctME R
12.5 a
d Is.
erW G'
75
$O ?E m
/
/
2.5 Pf>t i f
8 55 196 156 298 2b5 598 558 des 458 See 550 TIME ISECO Figure 15.6-34 Reflood Mixture Levels, CD = 0.6 DECLG, MINSI
l d.
WATER LEVELIFTI 28 17.5 DOW NCo Ms g 15.
- 12.5 d
- roc.
6 o
, t e.
W 3
- 7. s (ORE 3
6.
e
- 2. s f i
Goc.
g Se IN ise 2ee ese see 5s,
,3, sse TINE ISECl t
Figure 15.6-35 Reflood Mixture Levels,CD=0.4DECLG,MAXSI
i FLOOD RATE (IN/SEC) 2,
- 1. 5 w
N x
M E
i N
1.
s S
6
.5 8
58 ISO 155 209 256 5ee 556 496 458 568 558 TINC ISECB t
Figure 15.6-36 Core inlet Velocity, CD = 0.8 DECLG, MINSI
ll t
FLOOD RATE (IN/SEC) 2.
- ,1.5 na N
I C
E ac 1.
, ~ -
e 1
i
.5 i
8 56 196 ISO 208 255 500 556 408 458 500 558 TIME ISEC) e Figure 15.6-37 Core inlet Velocity, CD = 0.4 DECLG, MINSI f
-w--
g a
6 0
9 O
En W
Z
-2 E
d aOw e
O E
8o n
E, oo 3
-S 5>
g-2 8
e en 1
3Y I
u O
I e
~
d 9
g
/
,x' 4
a.
/
.E O
Z E
at!
ElC co e
e O
of J
e d
1335/ Nil 3pa 0007J 1
f li l
I t
l f
i FLOOD RATE (IN/SEC) 2 3 1.s bi i
=
N g-w g
S w
N
.s a
4 i
ese see ise see ese see 5se see ese see sse TINC ISEC3 f
t Figure 15.6-39 Core Inlet Velocity,CD=0.4DECLG,MAXSI l
t 4
h
,r,"g o
S i
0 I
4 I
SN I
M G
L 0
C E
I 3
)
D C
8 E
0 S
=
(
D C
E no x
M itc I
e 0 T n
j i
I 2
ro ta lu mu ccA 04
-6 i
0 5
1 i
1 e
ru ig F
o 0
o 0
0 g
0 b
0 0
0 o
0 0
8 S
4 2
_ owws$ _ sod e
i 1
l
I t
t 8000 l
6000
,oW (n
N 3
4000 Wo J
h_
2000 i
2 I
I I
I O
O IO 20 30 40 sO TIME (SEC) i Figure 15.6 41 Accumulator injection, CD = 0.4 DECLG, MIN &MAXSI 8 :
U
1536640 O
LD O
~
+
H CS Z.
HE O
ci f
g au O
w W
o
{i m
m o
il w
8 2
8 H
O F
N s
ow 3
E B
o<
~
O 9
e, y
E 1
e l
O O
O O
O O
O O
O O
O O
O O
m D
N (035/GT MO7:1
s 00 2
I r
I S
0 l
l I
6 I
1 M,
GL C
ED I
)
S 4
(
0 D
=
O D
O C
0 L 2 F w
I 1
E o
R l
F FO no T
i I
R t
A c
T e
S jn R
I 0 E 8 T d
I e
FA p
m E
uP M
I 3
T 4
I I
6 5
1 e
I ru 0
g I
4 iF I
F 0
0 0
0 0
0 0 O f
6 6
7 1
4 I
4 3
2 1
~ o0 sod 5I_h
{
i J
i
W.,'
((.
1536642
'i:
o 1,
o i
N
)
t i-I!. '
1 H
cox O
to
- x S
U W
Q n
m
-o it o
C o
U o
o a 3
O N g s
- g C
C L6 0
w HC o
0 2
O H
E 3
Oo r a
k E
S U3 v
e H
iw sn H
O Cu po ww l
1 I
I O
o o
o o
o oO p
n v
4
+
6 N
(035/G7) Mold
i 14 1--
12 tt-N3g 10 a
[
8 u)3O O_
6 OOI 4
V O
2 I
O i
O 2
4 6
8 10 12 CORE HEIGHT (FT)
Figure 15.6-45 Small Break Power Dis'tribution Assumed for LOCA Analysis b
i 100
__=
5 2 TOTAL DECAY HEAT (WITH 4% SHUTDOWN MARGIN) y 2
o 16' a-s
=
5 Z g
i 7
3 2
Nm
-2 H
lo 4
5 3
g z 2
3 I I I Illi I
I i 11111 I
IIlilli l I ll Illi i
l ll 11ll ig I
0 l
2 3
IG 2
5 10 2
5 fo 2
5 10 2
5 10 2
5 104 TIME AFTER SHUTDOWN (SECONDS)
Figure 15.6-46 Core Power (including Residual Fission) After Reactor Trip (Applies to all Small Breaks) t
a 15336-45 j
O o
LDN l
l OO ON e
.!!r d
O O
E ID R
O E
Lij o-W U
x x
b 2
E O
=
0 3
O 2
8 n
o '
sy D
W e
O Ei O
.=
LD,
f I
l i
i I
I O
O O
O O
O O
O O
O O
O m
O m
O W
N N
(VISd) 38nSS38d 508
D 15366-46 oo e
i m
I I
N I
I I
I I
i t
i oo i
l
- o kl W I N
o 01
[I
.E o
oI ol L
0li kl E
ol
.e o
i el 21 O
o
=
ri oI m-g F
O q
l Fl W
i o
m I
ol 3
l l
O (1}
a 1
2 o-m i
I o w g
l l
o a
I I
3a 1
I
- n I
g i
I 4
I I
o S
I j
o e
LD I
I E
I I
I I
I I
l
[
,'I O
o o
o o
o o
O
+
M N
(la) 73A37 38n1XIW 3803 i
)
1
4 7
1536647 O
~
O LD N
O O
ON
.E-e4>
2 3
m O
b' Oin r
U E
w W
Oa C/J W2 o
I e-.
O F g
O e
O 5
~
a>
O O
- M CDT O
C m
O in N
.!?
u.
I I
I I
I O
O O
O O
O O
O O
O O
O O
O O
O O
O O
O to N
N (3-s338930) dW31 '9AV GV73 0024 LoH l
l
d a
1536648 OOON O
LD b
O O
LD e
.Er O
a LD N
2
-o c
til m
O m
o O
E O
ttl e
2 w
F F
O 2
LD 6
b 8e O
ui O
~
LD
.9 u.
O
- LDN I
I I
o O
O O
O O
O O
O O
O O
W O
W O
W J
i N
N (visa) 38nss38a s3s
A 1536649 oo i
O I
l N
I I
I om wl l
b af I
~
oUI g
o L I wl o
o 01 mg m
.g C
o
~
a; UI ol el g
F o
l o;
D
~*1 I
Ml N
~
e oI u
m F
o 5
HI
@ ~w o
m 3
of 8
~
o I
-w I
t 1
em
~
l I
H F
o I
J m
f I
n y
I g
V I
l
?
o l
l o
=
=
l I
m e
l
.5 i
~
I i
o l
l m
m I
l I
l i
,l il o
o a
N (14 13^3138nixIn geog
15366-50 l
oOO N,
O in
)
n O
O E
m p
a>
eo O
16 5
-o R
w c
O m O
g O
v, w
- E s
~
u H
O 5
to gi n
a v
I 20 O
O g
to e
ui
~
M e
a O
E tD N
I I
I I
I o
O O
O O
O O
o O
O O
O O
O O
LD O
LD O
LD r0 N
N (J-S338930) 008 10H 'dW31 '9AV OV70
O 1536641 OO f
LO OO
+
a i~
vi O
t O
B to
^O u
W a-W 0
x
.x W
8 2
.s e,
O O
N 3O Ins
/
O O
2 u.
i
/
I I
O O
O O
O O
O O
O O
O O
m O
m O
O N
N (VISd) 380SS38d S08
- $h oo s
0 I
0 4
E-e R
m o-iT C-x_
F-sV t
h O-ig e
0 M-I 0
H e
\\_
O
)
r T-3 u
C tx T-E i
o S
M B-e
(
ro C
E k
E_
M ae I
r o_
0 T R
B C_
I 0
ge 2
L d
F_
lo C
i O_
"6 P_
4 o
-5 T__
6 0
5 1
i 0
e r
1 u
ig F
g o
0 0
0 0
o g
5 4
3 2
l p u_m J_w $
E agx.2 Wmoo
15363 53 O
Ob O
O O
o
.Es Y
eo O
a O
e LD f
n U
m W
2 U)
U 5
c.
W2 E
I O V O
e 4
m F
a 2
g io O
e.
O m
.9 u.
l l
I O
O O
O O
O O
O O N
~
O O
O O
O O
O LD O
LD O
LD to N
N (J-S338930) 008 10H 'dW31 '9AV OVlO
O e
1536664 OO ON i
l OO (O
/
a -
3
_e e
LL O
E O O g
N W g
U1 B
O W
-s 2
H E
F F
o a
O 2
U 0
6 to 9o C
O B
O
.e
+
l I
~
O O
O O
o o
m W
O W
(cas/91)3Mo1,1 Wysts 3800 3.tvy
HEAT TRANS. COEFFICIENT. BTU /FT2-HR *F N
O N
U1 O
O O
O l
N U1 O
O O
O O
O O
I I 1 1 1lll l l l lllll l
l l l1111 m
E a
.m O
O e-9e F$
m 8
O o
O 5
d P
r M
1
-4 U) a m
5 3-G o
O g
O
=
5'5' NO OO j
~
N U1 OO i
l l
SS-99CS L
a E
15366-56 O
O LD N
j OO ON
.5s Y
ez5 O
b O
E tD Er O
'E LD 2
in u.
8 c.
ld M
2 8
z O
l-y O
g O
m osa o
O b)
?
O e
O ui LD 2
a i-1 I
I I
I O
O O
O O
O O
O O
O O
O O
O O
m O
m O
m 70 N
N (a-s338930) asnivsaanal oInla
e ULNRC-124 7 T
Attachment E, Rev. 1 SIGNIFICANT SAFETY HAZARD EVALUATION FOR CALLAWAY PLANT TRANSITION TO OFA FUEL I.
INTRODUCTION As discussed in Reference 1, Union Electric Company has made the decision to replace Westinghouse 17 x 17 Low-Parasitic (LOPAR) fuel, currently being utilized at the Callaway Plant for Cycle 1, with Westinghouse 17 x 17 Optimized Fuel Assemblies (OFAs).
To implement this decision, Union Electric requested an amendment to the Callaway Operating License to incorporate the Technical Specification changes as specified in Attachment A of Reference 1, which are necessary to support a transition from a Westinghouse 17 x 17 LOPAR fueled core to a Westinghouse 17 x 17 OFA fueled core.
These Technical Specification changes are the result of several changes to design and analytical methodology as compared to those employed for the Cycle 1, LOPAR core.
The discussions contained in Sections III and IV of this evaluation address each of the parts of the amendment request and provide an analysis using the standards of 10CFR50.92 (i.e.,
the three factor test) :
II.
SUMMARY
OF DESIGN AND ANALYTICAL CHANGES The changes in design from Cycle I which are the subject of the proposed amendment are as follows:
a)
Installation of Westinghouse 17 x 17 Optimized Fuel Assemblies in place of Westinghouse 17 x 17 Low Parasitic fuel assemblies; and b)
Installation of Westinghouse Wet Annular Burnable Absorber (WABA) rods, instead of borosilicate glass burnable absorber rods.
The changes in analytical methodology employed in the safety evaluation of the proposed amendment are as follows:
a)
Use of the Improved Thermal Design Procedure (ITDP) to meet the DNB design basis, and the WRB-1 DNB Correlation, where applicable, instead of the W-3 R-Grid DNB Correlation; b)
Use of the revised PAD code thermal safety model to provide more realistic fuel temperature data; c)
Use of the NOTRUMP small break LOCA evaluation model in lieu of the WFLASH model which was used originally; and
$ s d) -Use of both the BART and the BASH large break LOCA evaluation models in lieu of the February 1978 model which was used originally.
Reanalysis was performed at an increased maximum core thermal power of 3565 MWt for all safety analyses which were required for the change to OFA.
III.
DESIGN CHANGES Union Electric has evaluated each of the changes in design against the Significant Hazards criteria of
On the basis of the results of these evaluations, Union Electric has concluded that the changes do not involve any significant hazards.
The following are summaries for each design change and the conclusions reached.
III.1 OPTIMIZED FUEL. ASSEMBLY (OFA)
As was stated, this amendment incorporates the Technical Specification changes which are necessary for the replacement of Westinghouse 17 x 17 LOPAR fuel assemblies with Westinghouse 17 x 17 OFAs.
This replacement will be performed over a number of cycles.
In Cycle 2, the core will include one reload region of OFA.
In subsequent cycles, additional regions of OFAs will be used until the core is completely OFA.
Detailed descriptions of the OFA fuel and the Westinghouse reload methodology can be found in WCAP-9500-A, " Reference Core Report 17 x 17 Optimized Fuel Assembly", and WCAP-9272-P-A, " Westinghouse Reload Safety Evaluation Methodology", respectively.
Both of these reports have received generic 4 approval from the NRC and have been referenced in connection with the approval of similar license amendments for other facilities.
The use of the OFA fuel increases the efficiency of the core by reducing the amount of parasitic material and reduces fuel cycle costs by optimizing the water to uranium ratio.
a)
This change does not involve a significant increase in the probability or consequences of an accident previously evaluated.
The OFA is mechanically compatible with the LOPAR fuel assembly, reactor internal interfaces, fuel handling and refueling equipment, and spent fuel storage racks.
This facilitates the transition.
The assembly envelope and grid centerline elevations for both designs are identical, thereby minimizing mechanical and hydraulic interaction.
However, there are several design differences between the OFA and LOPAR assemblies.
The OFA has reduced fuel rod, guide thimble, and instrumentation tube diameters, and has
A e six intermediate Zircaloy (mixing vane) grids instead of six intermediate Inconel (mixing vane) grids in the LOPAR assembly.
The reduced OFA guide thimble tube inside diameter provides sufficient diametric clearance for burnable absorber rods, source rods, and thimble plugs.
The Zircaloy grids have thicker and wider straps than the Inconel grids which compensate for differences in material strength properties.
The overall 17 x 17 OFA pressure drop is essentially the same as that for the 17 x 17 LOPAR fuel assembly.
b)
This change does not create the possibility of a new or different kind of accident from any accident previously evaluated.
This is based on the fact that the method and manner of plant operation is unchanged.
c)
This change does not involve a significant reduction in a margin of safety.
In a mixed core of OFA and LOPAR assemblies, the different grid structures and rod diameters result in a localized flow redistribution between an adjacent OFA and LOPAR assembly.
The effect of this localized flow redistribution is bounded by applying to the OFAs transition penalties of 2% DNBR and 10 deg-F LOCA Peak Clad Temperature (PCT) (see Section IV.3) which are used in the design and safety analyses.
As a rasult of the smaller OFA fue rod diameter, for the me power level, the OFA ro..
's a higher volumetric a
heat generation and surface Fea. flux than the LOPAR assembly rods.
In addition, *:
reduction in the OFA guide thimble tube diameter r<sults in an increase in the control rod scram time to the dashpot from 2.2 to 2.4 seconds.
This increase, as well as the other effects of the changes in design, have been incorporated in the non-LOCA and LOCA transient analyses.
The results of the new non-LOCA analyses, presented in Attachment C, indicate that the ANS Condition II acceptance criteria, as endorsed by NRC via NUREG-0800, are still met.
The results of the new LOCA analyses, presented in Attachments D and G, indicate that the acceptance criteria of 10CFR50.46 are still met.
In summary, the physics characteristics for the OFA fuel are only slightly different from those of the LOPAR fuel.
The differences are within the normal range of variations seen from cycle to cycle and are well within the envelope of conditions analyzed.
Therefore, there is no significant reduction in the safety margin.
. 5 III.2 WET ANNULAR BURNABLE ABSORBERS (WABA)
Wet Annular Burnable Absorber rods will be used in the Callaway reload cores which utilize 17 x 17 OFA fuel.
The WABA design has annular aluminum oxide - boron carbide (A1 0 B C) absorber pellets contained within 7 9 g
two concentric Zirdaloy tubes with water flowing through the center tube as well as outside the outer tube.
Cycle 1 used an annular borosilicate glass burnable absorber with a gas-filled central tube and outer stainless steel clad.
There are improved fuel cycle benefits associated with the WABA design.
These result from lower parasitic neutron absorption by Zircaloy as compared to stainless steel tubes, the increase in water fraction in the burnable absorber cell, and a reduction in the boron penalty at the end of each cycle.
The design of the WABA is discussed in detail in WCAP-10021-P-A, Rev.
1,
" Westinghouse Wet Annular Burnable Absorber Evaluation Report", which has received generic NRC approval and has been referenced in connection with the approval of similar license amendments for other facilities.
For the WABA rods, there is a larger core bypass flow than that for the glass absorber rod.
In order to assure that the total core bypass flow remains within its design basis limit, i
the number of WABA rods used for Cycle 2 and subsequent reloads will be limited to less than the allowable number of WABA rods approved in WCAP-10021-P-A, Rev.
1.
a)
This change does not involve a significant increase in the probability or consequences of an accident previously evaluated.
The referenced WCAP evaluates the material properties and mechanical performance of the WABA design.
The WABA design requirements are satisfied and the cladding integrity is maintained throughout the design lifetime.
b)
This change does not create the possibility of a new or different kind of accident from any accident previously evaluated.
This is based on the fact that the method and manner of plant operation is unchanged.
c)
This change does not involve a significant reduction in a margin of safety.
The referenced WCAP evaluates the nuclear and thermal-hydraulic design and concludes that the WABA rods satisfy all performance and design requirements for their design life.
III.3 DESIGN CHANGES SIGNIFICANT HAZARDS
SUMMARY
The Commission has provided guidance concerning the
?
application of the standards in 10CFR50.92 by providing certain examples (48FR14870).
This amendment request is i
1
similar to example (vi), 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 as specified in the Standard Review Plan, NUREG-0800.
The design changes specified do not involve a significant increase in the probability or consequences of an accident or other adverse condition over previous evaluations; nor create the possibility of a new or different kind of accident or condition over previous evaluations; nor involve a significant reduction in the margin of safety.
Based on the above, the design changes do not present a significant hazard.
IV.
ANALYTICAL CHANGES Union Electric has evaluated each of the changes in analytical methodology against the Significant Hazards criteria of 10CFR50.92.
On the basis of the results of these evaluations, Union Electric has concluded that the changes do not involve any significant hazards.
The following are summaries for each change and the conclusions reached.
IV.1 IMPROVED THERMAL DESIGN PROCEDURE (ITDP) AND WRB-1 DNB CORRELATION The analytical method employed with OFA fuel to satisfy the DNB design basis is the Improved Thermal Design Procedure (ITDP).
This methodology is described in WCAP-8567, " Improved Thermal Design Procedure," which has received generic approval from the NRC and has been referenced in connection with the approval of similar license amendments for other facilities.
This methodology is also discussed in WCAP-9500-A.
The ITDP treats plant operating and f ael parameters in a statistical manner to assure that DNBR limits are met.
For Cycle 1 with LOPAR fuel, a standard thermal-hydraulic methodology was used which considered plant and fuel best-estimate parameters and their uncertainties in a non-statistical manner to assure that DNBR limits were met.
The new analyses also utilize the WRB-1 DNB Correlation instead of the W-3 R-Grid Correlation which was utilized for Cycle 1.
The WRB-1 Correlation provides a more accurate critical heat flux prediction than the W-3 R-Grid Correlation.
This methodology is described in WCAP-8762-P-A, "New Westinghouse Correlation WRB-1 for Predicting Critical Heat Flux in Rod Bundles with Mixing Vane Grids", which has received generic NRC approval and has been referenced in connection with the approval of similar license amendments for other facilities.
This i
S methodology is also discussed in WCAP-9500-A, referenced above in connection with the discussion of OFA.
Use of the ITDP model and WRB-1 Correlation in the non-LOCA analyses makes available additional DNBR margin without encroaching on required safety margins.
A portion of the additional margin, assuming a maximum core thermal power of 3565 MWt instead of the licensed rated thermal power of 3411 MWt, is used in the safety evaluations provided in Attachment B, and the safety analyses required for the change to OFA and provided in Attachment C.
These evaluations and analyses of the applicable DNB-limiting transients were carried out at the assumed higher and, therefore, conservative maximum core thermal power level of 3565 MWt.
The W-3 R-Grid DNB Correlation used in the analysis of the Callaway Cycle 1 design was developed from experimental DNB studies conducted with fluid flowing inside heated tubes, modified so as to be applicable to the test results for the LOPAR fuel design.
Following development of the W-3 R-Grid DNB Correlation, Westinghouse developed the WRB-1 Correlation.
The WRB-1 Correlation is based exclusively on rod bundle tests with fluid flowing outside the heated tubes, a condition which is much more representative of the actual core configuration than the single tube tests on which the W-3 R-Grid Correlation is based.
The WRB-1 Correlation is a better predictor of DNB for the actual nuclear reactor geometries.
The DNBR limit for the WRB-1 Correlation is therefore lower than for the W-3 R-Grid correlation, since the data scatter is smaller.
Using the same statistical methods used to calculate the W-3 R-Grid DNBR limit of 1.30, the WRB-1 Correlation DNBR limit is 1.17.
The 1.17 DNBR limit has been accepted by the NRC as satisfying the DNB design basis when utilizing the WRB-1 Correlation.
The ITDP methodology is employed to show that the DNB design basis is satisfied.
Uncertainties in plant operating parameters, nuclear and thermal parameters, and fuel fabrication parameters are considered statistically such that there is at least a 95% probability with a 95%
confidence level that the minimum DNBR will be greater than or equal to 1.17 for the limiting power rod under normal operation and anticipated operational occurrences.
Plant parameter uncertainties are used to determine the plant DNBR uncertainty.
This DNBR uncertainty, combined with the DNBR limit, establishes a DNBR value which must be met in plant safety analyses.
Since the parameter uncertainties are considered in determining the design DNBR limit, the plant safety analyses are performed using i
nominal values of input parameters without uncertainties.
- _ _.. _ _ ~ _. _. _.,
- T In addition to the above considerations, a plant-specific DNBR margin has been considered in these analyses.
In particular, the DNBR limits of 1.42 and 1.45, for thimble and typical cells, respectively, were employed in the safety analyses.
The DNBR margin between the DNBRs used in the safety analyses (1.42 for thimble cells and 1.45 for typical cells) and the design DNBR limits (1.32 for thimble cells and 1.34 for typical cells) is more than sufficient to compensate for the DNBR penalties associated with transition core and rod bow effects and allows for design flexibility.
a)
This change does not involve a significant increase in the probability or consequences of an accident previously evaluated.
This is based on the fact that the DNB design basis and correlation do not involve any design change but only analytical changes.
b)
This change does not create the probability of a new or different kind of accident from any accident previously evaluated.
This is based on the fact that the method and manner of plant operation is I
unchanged.
c)
This change does not involve a significant reduction in a margin of safety described in the discussion provided above and the WCAPs referenced therein.
The change in the DNBR limit associated with the change in the DNB correlations in no way implies a reduction in the safety margin for Callaway.
This is true because the DNB design basis (i.e., that there is a 95% probability with a 95% confidence level that the hottest rod does not experience DNB) remains unchanged.
IV.2 PAD The revised PAD code thermal safety model is used to perform the fuel temperature calculations for the Callaway reload cores.
This model features the following changes:
1)
A modified annular gap conductance model 2)
Isotropic burnup-dependent fuel densification 3)
A 0.7 multiplier on the creep rate 4)
Best-estimate basis modeling techniques 5)
Uncertainties obtained from a statistical combination of model uncertainty and uncertainties due to i
manufacturing variations are added to the calculated temperatures.
The revised PAD code thermal safety model, based in part on more test data, is described in WCAP-8720, Addendum 2,
" Westinghouse Revised PAD Code Thermal Safety Model",
which has received generic approval from the NRC and has been referenced in connection with the approval of similar license amendments for other facilities.
Use of the revised PAD code thermal safety model provides more realistic initial fuel temperatures for the safety analyses.
The use of these more realistic temperatures has also been approved by the NRC in their SER for WCAP-8720, Addendum 2.
The fuel temperatures calculated with the revised PAD code thermal safety model are lower than those calculated with the model used for the Cycle 1 analyses.
a)
This change does not involve a significant increase in the probability or consequences of an accident previously evaluated.
This is based on the fact that no design change is involved, but only analytical modeling techniques are changed.
b)
This change does not create the possibility of a new or different kind of accident from any accident previously evaluated.
This is based on the fact that the method and manner of plant operation is unchanged.
c)
This change does not involve a significant reduction in a margin of safety.
The revised PAD code thermal safety model is an NRC-approved methodology which results in more realistic temperature predictions, lower than those calculated for the Cycle 1 analyses.
IV.3 SMALL BREAK LOCA ANALYSIS The change to OFA from LOPAR requires reanalysis of the small break loss of coolant accident (LOCA) to demonstrate compliance with the requirements of 10CFR50.46 and to develop OFA-specific peaking factor limits.
Because Union Electric desired that the large and small break LOCA analyses be applicable to Callaway over the long term, they were performed at an assumed upper bound design thermal power level of 3565 MWt and with the assumption that 10% of the tubes in the steam generators have been plugged.
This also involved the use of fuel data based upon the revised PAD code thermal safety model (WCAP-8720, Addendum 2) as described above.
In Cycle 2, Callaway will operate at a core power level of 3411 MWt.
Only a few of the tubes in the steam generators have been plugged and the likelihood that as many as 10% will have
. Q to be plugged is considered to be very low.
These assumptions lead to higher peak clad temperatures (PCT) than would otherwise be predicted and are therefore very conservative.
The small break LOCA analysis was performed using the Westinghouse NOTRUMP model described in WCAP-10054-P-A, which has been approved generically by the NRC and has been referenced in connection with the approval of similar amendments for other facilities, a)
This change does not involve a significant increase in the probability or consequences of an accident previously evaluated.
This is based on the fact that no design change is involved, but analytical models have been updated to the most recently approved versions.
b)
This change does not create the possibility of a new or different kind of accident from any accident previously evaluated.
This is based on the fact that the method and manner of plant operation is unchanged, c)
This change does not involve a significant reduction in a margin of safety in that the results of this analysis show the PCT is well below the 2200 degree F limit of 10CFR50.46.
IV.4 LARGE BREAK LOCA ANALYSIS The change to OFA from LOPAR requires reanalysis of the large break loss of coolant accident (LOCA) to demonstrate compliance with the requirements of 10CFR50.46 and to develop OFA-specific peaking factors.
Union Electric has elected to perform this analysis at conditions which will be applicable over the long term as described in Section IV.3.
The large break LOCA analysis was performed twice, once using the Westinghouse BART model described in WCAP 9561-P-A and a second time using the Westinghouse BASH model described in WCAP-10266.
The BART model is Westinghouse's most recently NRC-approved, while the BASH model is currently under NRC review with approval anticipated in the near future.
The results of the analyses show the peak clad temperature (PCT) from the BART analysis to be below the 2200 degree-F limit of 10CFR50.46.
The BASH model embodies the most up to date Westinghouse technology and leads to considerable improvement in operating margin.
Using BASH, the PCT is well below the regulatory limit.
O During the transition from LOPAR to OFA, because the hydraulic resistances of these assemblies differ slightly, the possibility of flow redistribution during the reflood phase of a large break LOCA exists.
This can lead to a PCT for the transition core that is somewhat higher than that for a full core of LOPAR or OFA.
Westinghouse studies performed using the BART model show that, in a transition core, the effect of the difference in hydraulic resistance between a LOPAR assembly and an OFA assembly is to reduce the steam flow velocity in the OFA at the elevation of the mixing vane grid during the reflood phase of the LOCA transient.
This effect is offset somewhat by the improvement in heat transfer attributable to the effects of the mixing vane grids.
The net effect on calculated PCT is an increase of not more than 10 deg-F.
This penalty is applicable only durir.g the transition core period and is only applied to the results of the large break LOCA analyses, which are performed assuming a full core of either LOPAR or OFA fuel.
The OFA-specific peaking factor limits were developed in a manner consistent with the methodology described in the NRC approved " Reference Core Report 17 x 17 Optimi ed Fuel Acsembly", WCAP-9500-A.
a)
This change does not involve a significant increase in the probability or consequences of an accident previously evaluated.
This is based on the fact that no design changes are involved with the exception of a bounding assumption that up to 10% of the tubes in the steam generators have been plugged.
This assumption has been used in the revised LOCA analyses, using the NRC approved BART model, and the PCT remains below the regulatory limit and the radiological consequences are not significantly affected as reported in Attachment F to Reference 1.
b)
This change does not create the possibility of a new or different kind of accident from any accident previously evaluated.
This is based on the fact that the method and manner of plant operation is unchanged, c)
This change does not involve a significant reduction in a margin of safety in that the analyses, using the NRC-approved BART model, show the PCT is below the regulatory limit of 2200 deg-F.
IV.5 ANALYTICAL CHANGES SIGNIFICANT HAZARDS
SUMMARY
The Commission has provided guidance concerning the application of the standarris in 10CFR50.92 by providing certain examples (48FR14870).
This amendment request is
" O similar to example (vi), 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 as specified in the Standard Review Plan, NUREG-0800.
The analysis changes specified do not involve a significant increase in the probability or consequences of an accident or other adverse condition over previous evaluations; nor create the possibility of a new or different kind of accident or condition over previous evaluations; nor involve a significant reduction in the margin of safety.
In addition, the analysis changes, with the exception of the use of the BASH large break LOCA evaluation model, have all received generic approval from the NRC and have been referenced in connection.with the approval of similar license amendments for other facilities.
The large break LOCA analysis was also performed using the BART model which has received NRC approval.
Based on the above, the analysis changes do not present a significant safety hazard.