ML20114A991
| ML20114A991 | |
| Person / Time | |
|---|---|
| Site: | Mcguire, Catawba, McGuire  |
| Issue date: | 11/30/1991 |
| From: | DUKE POWER CO. |
| To: | |
| Shared Package | |
| ML19303F031 | List: |
| References | |
| DPC-NE-3001-A, NUDOCS 9208240229 | |
| Download: ML20114A991 (276) | |
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l DUKE POWER COMPANY McGUIRE NUCLEAR STATION CATAWBA NUCLEAR STATION MULTIDIMENSIONAL REACTOR TRANSIENTS AND SAFETY ANALYSIS PHYSICS PARAMETERS METHODOLOGY DPC-NE-3001-A - November 1991 4 Nuclear Engineering Group Nuclear Services Division Nuclear Generation Department Duke Power Company 9200240229 920817 PDR ADOCK O$000369
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,(paasegjo, UNITED STATES f w j NUCLE AR REGULATORY COMMISSION WASHINGTON, D. C. 20555
! Novemoer 15, 1991 j
Docket Nos. 50 369, 50-370 50 413 and 50-414 l Mr. H. B. Tucker, Senior Vice President Nuclear Generation Duke Power Company P. O. Box 1007 Charlotte, North Carolina 28201-1007
Dear Mr. Tucker:
SUBJECT:
SAFETY EVALUATION ON TOPICAL REPORT DPC-NE-3001, " MULTIDIMENSIONAL REACTOR TRANSIENTS AND SAFETY ANALYSIS PHYSICS PARAMETERS" (TAC N0s. 75954/75955/75956/75957) The NRC staff has reviewed Duke Power Company Topical Report DPC-NE-3001,
" Multidimensional Reactor Transients and Safety Analysis Physics Parameters,"
dated Janaury 29, 1990, as supplemented by letters dated February 13 and June 3,1991. The staff has found the topical report to be acceptable for referencing it, licensing analyses for the McGuire and Catawba Nuclear Stations
-subject to the conditions in section 4.0 of the attached Safety Evaluation.
This concludes our review activities in response to your submittals regarding Topical Report DPC-NE-3001 addressed by TAC numbers 75954/75955/75956/75957. Si ely, imothy . Reed, Project Manager Project Directorate 11-3 Division of Reactor Projects - I/II Office of Nuclear Reactor Regulation E. 'osure: Safety Evaluation cc: See next page i +
. a %. Catawba Nuclear Station Duke Power Company McGuire Nuclear Station cc: Mr. R. C. Futrell Mr. Alan R. Herdt, Chief Regulatory Compliance Manager Project Branch #3 Duke Power Company U.S. Nuclear Regulatory Commission Catawba Huclear Site 101 Marietta Street, NW, Suite 2900 Clover, South Carolina 29710 Atlanta, Georgia 30323 Mr. A.V. Carr. Esq. North Carolina Electric Merbership Duke Power Company Corp. 422 South Church Street P.O. Box 27306 - Charlotte, North Carolina 28242-0001 Raleigh, North Carolina 27611 J. Micnael McGarry, III, Esq. Saluda River Electric Cooperative, Winston and Strawn Inc. 1400 L Street, N.W. P.O. Box 929 Washingtcr., DC 20005 Laurens. South Carolina 29360 North Carolina HPA-1 Senior Resident inspector Suite 600 Route 2, Box 179N P.O. Cox 29513 York, South Carolina 29745 Raleigh, North Carolina 27626-513 Regional Administrator, Region 11 fir. Frank Modrak U.S. Nuclear Regulatory Conmission Project flanager, Mid-South Area 101 Marietta Street, NW, Suite 7900 ESSD Projects Atlanta, Georgia 30323 Westinghouse Electric Corporation MNC West Tower - Bay Pal Pr. Heyward G. Shealy, Chief P.O. Box 355 Bureau of Radiological Health Pittsburgh, Pennsylvania 15230 South Carolina Dept. of Health . and Environmental Control County Manager of York County 2600 Lull Street York County Courthouse Columbia, South Carolina '0201 York, South Carolina 29745 Ms. Karen E. Long Richard P. Wilson, Esc. Assistant Attorney General Assistant Attorney General North Carolina Dept. of Justice S.C. Attorney General's Office P.O. Box 629 P.O. Box 11549 Raleigh, North Carolina 27602 Columbia, South Carolina 29?ll Mr. R. L. Gill, Jr. Piedmont Municipal Power Agency Licensing 121 Village Drive Duke Power Company Greer, Sauth Carolina 29651 P.O. Box 1007 Charlotte, North Carolina 28201-1007
Catawba Nuclear Station . Duke Power Company McGuire Nuclear Station County Manager of Mecklenburg County Dr. John M. Barry 720 East Fourth Street Department of Environmental Health Charlotte, North Carolina 28202 Mecklenburg County j 1200 Blythe Boulevard i Charlotte, North Carolina 28203 l l Mr. R. O. Sharpe Mr. Dayne H. Brown, Director Compliance DepartrF'rt of Environmental Health Duke Power Company an( lMttral Resources McGuire Nuclear Site Division of Radiation Protection 12700 Hagers Ferry Road P. O. Box 27687 Huntersville,-North Carolina 28078-8985 Raleigh, North Carolina 27611-7687 Senior Resident inspector Mr. M. t ' .ckman c/o U.S. Nuclear Regulatory Commission Vice Pres u.<nt , Catawba Site 12700 Hagers Ferry Road l Duke Power Company Puntersville, North Carolina 28078 P. 0.. Box 256 Clover, South Carolina 29710 Mr. T. C. Mctiee kin Vice President, McGuire Site Duke Power Company 12700 Hagers Ferry Road Huntersville, North Carolina 28078-8985
/,[pm noy), UNITED STATES y NUCLEAR REGULATORY COMMISSION
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..... SAFETY EVALUATION FOR DPC-flE-3001-P "MULT101 MENS 10NAL REACTOP TRANSIENTS AND SAFETY ANALYSIS PHYSICS PARAMETERS METHODOLOGY"
1.0 INTRODUCTION
By letter dated January 29,1990 (Ref.1), Duke Power Company (DPC) submitted topical report DPC-NE-3001-P, " Multidimensional Rnctor Transients and Safety - Analysis Physics Parameters Methodology." The methodology described in this topical report expands on the currently approved reload design analyses of Reference 2 and is intended for application to the Catawba and licGuire Nuclear Fower Stations. The report includes the overall methodology for using bounding reference analyses together with key safety parameters for analyzing the required Final Safety Analysis Report (FSAR) Chapter 15 events as well as the DPC reference analyses for selected transients involving multidimensional neutronics. The bounding analysis methodology used by DPC to ensure that the accident analysi for the reference core conservatively bounds the reload core is describc in the topical report. The important key safety parameters for each Chapter 15 event are identified, and the methods for calculating these parameters are described. New DPC bounding reference analyses are given for (1) the rod ejection accident (REA), (2) the steam line break accident (SLBA), and (3) the dropped rod accident (DRA). The new reference analysis for the REA is performed with three-dimensional spatial neutronics, and the analyses for SLBA and DRA are performed with a point kinetics model. The new reference ; analyses are snalyzed in detail and shown to satisfy the appropriate 10 CFR _ Part 100 cose limits, the departure f rom nucleate boilin ratio (DNBR) safety limit, the fuel enthalpy limit, and the American Society of itechanical Engineers reactor coolant system pressure limit. In reload applications, DPC will show that the reference analysis is bounding by demonstrating that the event-specific key safety parameters of the reload core are within the conservative envelope of the reference analysis. The topical report is reviewed in Section 2, and the safety evaluation of the DPC methodology is summarized in Section 3. The limitations imposed , concerning the licensing application of the DPC methoo.1 are given in t Section 4 , The following summary and technical evaluation include the contribution of Brookhaven National Laboratory as staff consultant under FIN No. A-36B6. 2.0
SUMMARY
OF THE TOPICAL REPORT DPC's topical report DPC-NE-3001-P (1) identifies the key safety parameters, (2) describes the methods for calculating these parameters, and (3) gives the new reference analyses for the rod ejection accident, steam line break accident and dropped rod accident. The DPC nethods associated with these analyses are summarizec in the following sections.
i: 4
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2.1 Key Safety Parameters The key safety. parameters play a crucial role in the DPC reload methodology.
. By comparing these parameters ;or a reload core to the values deterinined for the reference analysis, any nonconservatism in the reference analysis may be !
identified and the need for a new safety analysis may be established. The OPC methodology defires both generic and event-specific safety carameters. l The initial core power distribution, scram reactivity, effective delevad neutror fraction and decay constants, .and prompt neutron lifetime are import for many transients and are considered to be generic. A conservative st.. i reactivity is determined using the minimum shutdown marpin allowed by the technical specifications and the rod insertion limits, together with a minimum rate of reactivity insertion. This rate is detennined using the measured roo speeds and a conservative correlation between rod insertion and fractional inserted "eactivity. The core power distribution is assumed to provide the maximum peaking allowed (including perturbations) by the-F and F technical specification limits. Forrapidtransients3minimumdelahedneuNonfraction bete is used; for slow transients, which are insensitive to beta, a maximum value of beta is recommended. For conservatism the initial fuel temperature is taken to be the maximum. The FSAR Chapter 15 accident analyses are reviewed in Chapter 2 of the topical report, and event-specific key safety parameters are identified for each event. The events' analyzed include feedwater malfunctions, increased steam flow,- turbine trip, loss of nonemergency ac power, loss of coolant flow, rod ejection, inadvertent actuation of_the emergency core cooling system, and loss-of-coolant accidents. On the basis of the dynamics of the transient, the conservative direction for each key safety parameter is given for the events analyzed. The event-specific key safety paraceters include the Doppler temperature coefficient (DTC), moderator temperature coefficient (FTC), shutdown margin-(SDil), accident reactivity, and critical beron concentration. 2.2 Determination-of the Key Safety Parameters The key safety parameters are determined using physics codes and methods that the NRC staff has approved (Ref. 2) or is reviewing.(Ref. 3). The three-di- m ional static power distribution and depletion calculations are performed wit ,10E-P or SIltVLATE-3P. The models used are based on accumulated operating hist y .of previous reload cycles. The static physics parameters include control rod worth, shutdown margin and trip reactivity. The calculated parameters depend on the three-dimensional power. shape ano, consequently, core loading, rod insertion, and time in life. The shutdown margin and rod worth calculations are performed at beginning-of-cycle (BOC) and at end-of-cycle (EOC). The shutdown margin calculations assume the highest worth rod is stuck in its fully withdrawn positior and account for the power defect, rod insertion, and calculational uncertainties.
2 .;, w The transient parameters inc'ude the MTC, DTC, delayed neutron parameters, and baron worths. The temperature coefficients are calculated using a static model and account for powcr level and cycle exposure. The boron worths are determined using a set of perturbed s+atic calculations, and the dependence on power level, moderator temperature, fut:1 exposure, and control rod insertion is included. 2.3 New Reference Analyses for the Rod E.iection, m ,m Line Break and DrocDed Roc Accments The topical report includes new reference analyses for the rod ejection, steam line break, and dropos rod accidents. The detailed methods for analyzing the events are pretented together with the resulting consequences, margin to limits, and acceptance criteria. The cases analyzed are extreme and should bound most reload cores. In pr ctice, whether these reference analyses actually bound the postulateo events for the reload core will be determined by comparing the key safety parameters, which are given for each of the DPC reference analyses. 2.3.1 Rod Ejection Accident Reference Analysis The rod ejection accident (REA) reference analysis consists of three distinct and coupled analyses: (1) a core neutronics analysis (2} a core thermal-hydraulics analysis, and (3) a systems thermal-hydraulics analysis. The core neutronics response to the rod ejection reactivity insertion is calculated with the Electric Power Research Institute (EPRI) ARROTTA code (Ref. 4). ARROTTA calculates the three-dinensional power / flux solution in (x,y,z) geometry both for the static analyses requiri:d for determining the key parameters and for the core response during the transient. The ARROTTA core model uses one radial noce per assembly and twelve axial nodes. The fuel and reflector two-group cross sections and nuclear input parameters were oetermir.ed with CAS!10-3 (Ref. 3). The local cross sections are a function of fuel and moderator temperature and relative water density. ARROTTA uses assembly discontinuity factors, calculated by CASMO-3, to account for the local heterogeneities within the fuel assembly. ARROTTA includes a core thermal-hydraulics model that is identical to the one included in the EPRI BEAGL program (Ref. 5). The ARROTTA time-depend:mt core power / flux solution is used as input to the subchannel core thermal-hydraulics analysis performed with VIPRE-01 (Ref. 6). The NRC has approved VIPRE-01 for referencing in licensing analyses. VIPRE-01 calculates the core flow distribution and coolant conditions, fuel rod temperatures, and DNBR during the REA. VIPRE-01 uses a single-channel fuel conduction model together with the ARROTTA time-dependent peak pin power to calculate the peak fuel enthalpy. The transient pressure is calculated using a multichannel VIPRE-01 model and the ARROTTA time-dependent power / flux solution. A RETRAN-02 (Ref. 7) system model is used to determine the reactor coolant system (RCS) response during the REA. The RETRAN model which is based on the McGuire/ Catawba model (Ref. 8), uses the VIPRE-01 core coolant expansion rate to determine the limiting RCS transient pressure.
The CEA reference analysis model is based on a Cycle 2 Catawba-1 core and is-performed at hot-full-power (HFP) and hot-zero-power (HZP) conditions at BOC and EOC. The. nuclear cross sections have been adjusted to increase the power peaking and ejected rod worth (at off-center location D-12) so that the reference analysis will bound expected core reloads. The D-12 rod is ejected in-(a conservative) 0.1 second. The core inlet flow is reduced by P.2 percent to account for measurement uncertainty in the HFP case. The core inlet flow is reduced an additional 54 percent for the (two-pump) HZP case. The reference analysis indicated that the REA results in a maximum fuel enthalpy of 133 cal /gm and a peak system pressure of 2609 psig which are lower than the corresponding limits. The number of rods in departure from nucleate boiling (DNB) was less than 37 percent, and the resulting offsite doses were well within the 10 CFR Part 100 limits. The REA key safety parameters-(which vary from cycle to cycle) are the moderator and Doppler temperature coefficients, the delayed neutron fraction, and the ejected rod worth. 2.3.2 Steam Line Break Accident Reference Analysis The steam line break accident reference analysis consists of (1) a RETRAN-07 systems analysis of the RCS response to the steam line break, (2) a NODE-P (Ref. 2) or SIMULATE-3P (Ref. 3) neutronics calculation of the core power distribution at the time of minimum DNBR, and (3) a VIPRE-01 core thermal-hydraulics analysis of the minimum DNBR. 'The systems analysis model is based on the ficGuire/ Catawba model (Ref. 8). To model the thmnal mixing, the RETRAN-02 model was modified to include parallel flowpaus, with one path
-connected to the faulted loop and the other path representing the intact loops.
Special mixing junctions are included to allow for thermal mixing. The RETRAM-02 neutronics feedback is included, using_ a precalculated k noceratortemperaturefunctionandaDopplertemperaturecoefficieN7versus . A range of break sizes was evaluated to determine the limiting break size. The three-dimensional power distribution is calculated for the asymmetric core conditions determined by RETRAN-02 at the time of minimum DNBR (MDNBR). This power distribution is then used in the VIPRE-01 multichannel steady-state calculation to determine the MDNBR. The VIPRE-01 calculation uses the calculated asymetric core inlet temperature and flow as boundary conditions. - The initial conditions and boundary conditions used in the reference analysis are generally conservative. These include a low pressurizer level and RC5
'1cr and. a high RCS temperature arid steam generator water inventory. The core is initially at hot-zero-power conditions to maximize the cooldown. The reference analysis is performed with and without offsite power for a limiting break size of 1.4 f tr. In both cases, the MONBR is greater than 1.45, which-
.is greater than the limiting MDNBR value.
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2.3.3 Droceed Rod Accident Reference Analysis
-The dropped roo transient is analyzed with a RETRAN-02 plant systems model.
The input dropped rod worth and core moderator and fuel reactivity coefficients are determined with NODE-P. The RETRAN-02 model calculates the initial reduc-tion in power, bank withdrawal and moderator cooldown, and the final minimum DNBR statepoint. The DNBR analysis is performed with a multichannel VIPRE-01 model using the RETRAN-02 systems input together with a detailed three-dimensional power distribution determined with either NODE-P or SIMULATE-3P. For the reference analysis, a single-loop Catawba 1 RETRAN-02 model is used. The uncertainty in plant operating variables is accounted for by using either the statistical core design (SCD) methodology or conservative upper-limit input values. _ This includes a low pressurizer level and a max 1 mum average fuel temperature. The reference analysis is performed at BOC, MOC and EOC for a range of droceed rod worths. The reference analysis calculations indicate that the MDNBR does not reach the SCD DNBR limit for the cases analyzed. The key physics parameters for the dropped rod analysis that will be used to ' evaluate each DPC reload core are the initial radial peakino F axial flux l shape,moderatoranddopplertemperaturecoefficients,droppedho,dworth,and bank withdrawal worth. l 3.0 EVALUATION DPC's topical report DPC-NE-3001-P provides the physics methods that will be used to evaluate reload cores against precalculated reference analyses. The focus of this review was on the identification of the key safety parameters and on the reference analyses provided for the rod ejection, dropped rod, and steam line break accident. The initial review of the topical report resulted in a series of questions. This evaluation included the review of DPC-NE-3001-P and i DPC's responses to these questions in References 9 and 10. The evaluation of the major issues raiseo during this review is summarized in the following sections. 3.1 Key Safety parameters Both generic and event-specific key safety parameters are used in the DPC methocology. The identified parameters are based on the dynamics of the transient, the sensitivity with respect to the safety parameter,'and the approach to safety limits. The list of key safety para'.1eters in Table 2-1 of the topical report does not include all the modeling input data used in_the reference analysis that may change with a new reload core design. However, DPC intends to exami_ne all of the thermal-hydraulic and mechanical parameters, as well as the physics key parameters, in validating the reference analysis for application to a reload core, l l
In Chapter 3 of the topical report and in Response 4 of Reference 9, DPC has indicated that the key safety parameters will be determined using approved codes and methods. However the SIMULATE-P methods described in DPC topics 1 report DPC-f;E-1004 (Ref. 3),are being reviewed by the NRC staff and should not be used until they have been approved. 'DPC intends to use the key safety parameters identified in Table 2-1 to evaluate the licensing events. The only issue raised was the prompt neutron lifetime which DPC does not consider as a key safety parameter for the uncontrolled withdrawal of the rori cluster control assembly. In Resrmse 27 of Reference 10, DPC indicated that for the bank withdrawal at power, ' : prompt neutron lifetime affects the bank reactivity insertion rate. Howeve DPC has performed sensitivity calculations that indicate that the itDNBR is essentially unaffected by changes in the prompt neutron lifetime. For the withdrawal of a single rod, the limiting statepoint occurs well af ter the rod is completely withdrawn and there is only minimum sensitivity to the prompt neutron lifetime. OPC has also performed sensitivity calculations for the bank withdrawal from subcritical. These calculations indicate only a very weak sensitivity of the transient te the neutron lifetime. On the basis of the above, the staff concludes that the determination and application of the key safety parameters in the DPC reload methodology are acceptable. 3.2 Rod E,iection Accident Analysis The RETRAN-0? systems model and VIPRE-01 core thermal-hydraulics model used in the DPC rod eiection accident (REA) analysis are based on the Babcock f Milcox liark-BW fuel and the Westinghouse optimized fuel designs. The introduction of new fuel designs (involving enanges in loss coefficients, dimensions, etc.) may invalidate tne applicability of the reference analysis to the reload cere. This is also a concern for the steam line break and dropped rod accident analyses. OPC has' indicated in Responses 1,18, and 24 of Reference 9 that when a new fuel. design is included in a cycle reload core, the impact of the design changes on all analyses will be evaluated and a reanalysis will be performed if necessary. The ARROTTA analysis neglects the change in the assembly-wise flow distribution and assently crossflow during the REA. This is considered to be a good approximation, since nu significant heating is transferred to the moderator until after the power transient has been reversed in both the. bot-full-power (liFP) and hot-zero-power (H2P) cases, and no bulk boiling occurs until af ter local departure from nucleate boiling (DNB) occurs. . l l
o.n u - The nuclear cross sections in ARROTTA are represented as functions of the local moderator density, control rod insertion, and fuel and moderator temperature. The ARROTTA REA model includes the local fuel exposure dependence by defining a set of about 75 distinct fuel compositions. However, a typical reload core consists of a continuous three-dimensional fuel exposure distribution and about 1000 unique fuel compositions. DPC collapses this set of fuel compositions from about 1000 down to about 75 using the SIGTRAN code, in response to question 9 Reference 9, DPC states that the ARROTTA sensitivity calculations in which the number of compositions was increased by about 50 percent indicated no significant change in the ARROTTA predictions. In the DPC methodology, the REA safety parameters (F moderator temperature coefficient,dopplertemperaturecoefficient, beta,hc,dejectionworth,andthe pin power census) are calculated for the reference core in which the cross sections in the neighborhood of the ejected rod have been adjusted to increase the rod worth and local peaking. This is an appropriate definition of the key saf ety parameters, since it ensures consistency between the REA reference results ano the cycle reloao core safety parar.eters. The temperature coefficient and rod reactivity safety parameters are calculated by standard
'tatic eigenvalue differencing. The temperature reactivity feedback is calculated using an isnthermal analysis for the HZP case (Response 13, Ref. 9).
At HFP the feedbacks are determinea by an increase in uniform inlet temperature and an increase in core thermal power. The ejected rod worth is calculated without feedback (Response 13 Ref. 9). The ARROTTA code uses the analytic nodal method of Reference 11 and the thereal-hydraulics model of Feference 5. The flux solution is calculated in two groups, and the use of discontinuity factors allows an accurate reconstruction of the local pin-wise power distribution. In Reference 12, DPC has indicated that ARROTTA is only used for the REA and, because of the rapid nature of this event, the neutronics solution rather than the moderator feedback ef fects are most important for this application. As qualification of the ARR0TTA neutronics solution, Combustion Engineering (CE) under contract to Electric Power Research Institute (EPRI) compared the ARROTTA coce Ref.13) to the NRC-approved CE FERfl1TE code (Ref. It). Comparisons were made for steady-state conditions and for an off-center REA in half-core geometry from HZP conditions. Good agreement was obtained for the REA transient core power, peak assembly power, core average fuel temperature ard peak f uel temperature, and steacy-state power distributions. In Reference 15, l additional ARROTTA comparisons are given for four static calculations and . two transient calculations, two of which included thermal-hydraulics feedback. These comparisons indicate good agreement in the static eigenvalues, transient core power, and both static and transient power distributions. In Response 5 of Reference 9 DPC stated that Version 1.02 of ARROTTA was used in the HERMITE benchmarking calculations, which is the same version as that used in the Reference 15 analyses and the DPC REA reference analysis. The EPRl/CE ARROTTA-HERMITE comparisons in Reference 13 and the ARROTTA comparisons of Reference 15 indicate that ARROTTA provides an accurate calculation of the rod ejectiun transient. l l
8_ l On the basis of the above and the responses in Reference 10, the staff concluoes that the DPC analysis of the REA is acceptable. ' 3.3 Stean Line Break Accident Analysis In the DFC steam line break accident (SLBA) analysis the stuck rod is located in the sector of the core associated with the faulted loop. This results in i maxir.um core peaking, but also results in minimum inlet temperature. DFC has I shown in Response 37 of Reference 10 that the location uf the rod in the I faulted loop sector results in an MDNBR and is conservative. Two of the key factors affecting the DPC steam line break response are the steam generator inventory and the auxiliary feedwater flow to the faulted steam generator. The Catawba-units have higher feedwater flow to the faulted steam generator than McGuire. In addition, of the two Catawba units, Catawba 2 has the highest initial steam generator inventory. In its Response 30 of Reference 10, DPC has indicated that Catawba 2 also has a higher steam generator inventory than both McGuire units at the SLBA initial conditions. DPC's selection of Catawba 2 as the bounding unit for the SLBA is therefore conservative. On the basis of the above and the responses in Reference 10, the staff concludes that the DPC analysis of the SLBA is acceptable. 3.4 Dropped Rod Accident Analysis The measured core power is a primary factor in determining the power overshoot in the response to the dropped rod in the dropped rod accident (DRA). The location of the dropped rod can produce a core power tilt and adversely affect the measured core thermal power. As indicated in Responses 42 and 43 of Reference 10, DPC assumes a control rod system failure that results in the limiting power tilt and a minimum measured core thermal power. This assumption maximizes the DRA power overshoot and minimizes the margin to DNB. In addition, DPC assumes the control withdrawal stops are inoperative allowing the power overshoot to proceed above 103 percent power (Response 43, Ref.10). In the dropped rod event the least negative temperature coefficient provides a conservative minimum feedback to the power transient,.but also results in a ntnconservative minimum positive reactivity insertion resulting from the cooldown. DPC has performed sensitivity calculations which indicated that the feedback reactivity dominates the core response and a least negative temperature coefficient provides the bounding conservative DRA analysis. On the basis of the above and the responses in Reference 10, the staff concludes that the DPC analysis of the DRA is acceptable.
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-g-3.5 Applications of Codes and Methodolocy ine RETRAN-02 and VIPRE-01 codes are used in the DPC rod ejection, steam line break, and dropped rod accident analyses. In Response 29 of Reference 10, DPC
,has indicated that the coplication of these codes is outside the limitations of their present NRC approval in two instances: (1) PETRAN-02 MOD 005 is used to determine the boron transport in the steamline break accident analysis and (2) the VIPRE-01 heat transfer correlations are used for post critical heat flux ICHF) analyses in the rod ejection accident (REA) analysis.
In response to the VIPRE-01 concerns, DPC has indicated that the application of the post-CHF heat transfer correlations in the REA analysis affects both the peak fuel enthalpy and rtactor coolant syszem (RCS) peak pressure calculations. DPC has determinec through sensitivity analysis that the available post-CHF correlations result in conservative fuel temperatures or have a negligible effect on the peak fuel temperatures or beth. In view of the large (factor of about 2) margin between the calculated REA peak fuel enthalpy and the fuel enthalpy limit, this is acceptable. DPC has also evaluated the effect of the post-CHF heat transfer correlations on the peak RCS pressure and found they have less than a 14-psig effect, which is within the available margin to the RCS pressure limit. Consequently, DPC's application of the VIPRE-01 post-CHF heat transfer correlations in the P.EA analysis is acceptable. Since no additional information is provided on RETRAN-02, the approval of the DPC transient analysis methods is contingent on the approval of M00005 of RETRAN-02 for boron tr:nsport calculations. The thermal-hydraulics methodology described in the DPC tcpical report DPC-NE-3000 (Ref. 8) has been used in the DPC transient analyses. The limitations of the flRC approval of the DPC-NE-3000 thermal-hycraulics methodology will, therefore, also apply ~ to the transient analysis methodology of DPC-NE-3001-P. 4.0 LlHTTATIONS The staff has reviewed in detail the DPC reactor transients and safety analysis physics parameters methodology topical report and the supporting dccumentation in References 9 and 10. The topical report documents the DPC reload key safety parameter methods and the reference analyses for the rod e,iection, l croppec roc, and steer line break accidents. On the basis of this review,-the staff concludes that the DPC methodology is acceptable for performing licensing ai alyses under-the conditions stated in Section _3 of this evaluation and summarizec as follows:
-(1) The licensing application of the Sill!JLATE-3P static methods for deter-mining the key safety parameters requires NRC approvhl of the reference topical report, DPC-NE-1004 (Section 3.1).
(2) The licensing application of the DPC-NE-3001-P transient analysis methods requires NRC approval of MOD 005 of RETRAN-02 for boren transport calculations (Section 3.5). (3) The licensing application of the DPC-NE-3001-P transient analysis methods { requires NRC approval of the thermal-hydraulics topical report OPC-NE-3000 ! (Section 3.5).
5.0 REFERENCES
- 1. Letter, H. B. Tucker (DPC) to U.S. Nuclear Regulatory Commissic " Safety Analysis Physics Parameters and Multidimensional Reactor Transieas Methodology," DPC-NE-3001-P, January 29, 1990.
- 2. D.ske Power Company, " Nuclear Physics Methodology for Reload Design,"
DPC-NF-2010A, June 1985.
- 3. Duke Power Company, " Nuclear Design tiethodolugy Using CAS!1C-3/SIMt' LATE-3P,"
DPC-NE-1004, January 1990. 4 L. D. Eisenhart, "ARR0TTA An Advanced Rapid Reactor Operationc . Transient Analysis Computer Code," Computer Code Documentation Package, EPRI NP-7375-CCML, Volume 1, Electric Power Research Institute Aegust 1991.
- 5. D. J. Diamond, H. S. Cheng, and L. D. Eisenhart, "BEAGL-01, A Cuputer Code for Calculating Rapid Core Transients," EPRI NP-3243-CCM, Volume I, Electric Power Research Institute, 1983.
- 6. Letter from C. E. Rossi (NRC) to J. A. Blaisdell (UGRA), " Acceptance for Referencing of Licensing Topical Report, VIPRE-01: A Therinal-Hydraulic Analysis Code for Reactor Cc,res," EPRI NP 7511-CCH, Vol. .-5, liay 1986.
! 7. Electric Power Research Institute, "RETRAN-02: A Program for Transient Thermal-Hydraulic Annalysis of Complex Fluid Flow Systems," EPRI NP-1850-CCHA, Revision a, November 1988.
- 8. Letter, H._ B. Tucker (DFC) to NRC, " Thermal-Hydraulic Transient Analysis
- - Methodology," DPC-NE-3000, September 29, 1987.
l 9. Letter, H. B. Tucker (DPC) to U.S. Nuclear Regulatory Commission," Response L to-Request for Additional Informatier Relative to Topical Report OPC-FE-3001-P," February 13, 1991.
- 10. Letter, H. B. Tucker (DPC) to U.S. Nuclear Regulatory Commission, " Response to Request for Additional Information Relative to Topical Report DPC-NE-3001-P,"
- i. June.3, 1991.
- 11. K. S. Smith, "An Analytic f:odal Method for Solving the P-Group, Multi-Dimensional, Static and Transient Neutron Diffusion Equations," Nuclear Ergineering Thesis, Department of Nuclear Engineering, Massachusetts Institute of Technology, Cambridge, Massachusetts, February 1979.
l L
.aw
- 12. Letter, H. B. Tucker (DPC) to U.S. Nuclear Regulatory Comission, " Topical Report OPC-NE-3001-P," September 14, 1990.
- 13. Electric Power Research Instutua, "ARROTTA-HERMITE Code Comparison,"
NP-6614 EPRI, December 1989.
- 14. P. E. Rohan, S. G. Wagner, S. E. Ritterbusch, "HERMITE: A Multidimensional Space-time Kinetics Code for PWR Transients," CENPD-188-A, Corrioustion Engir.eering, Inc., July 1976.
- 15. Electric Fower Research Institute, "AROCTTn Validation and Verification - -
Standard Benenmarks Set," Project 1936-6, July 1989. 06te: November 15. 1991
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1 T DUKE PQWER COMP 6NY McGUIRE NUCLEAR STATION CATAWBA NUCLEAR STATION PAULTIDIMENSIONAL REACTOR TRANSIENTS AND SAFETY ANALYSIS PHYSICS PARAMETERS METHODOLOGY DPC NE-3001-A November 1991 Nuclear Engineering Group Nuclect Services Division N. wear Generation Department Duke Power Company o i ,' , , , . - + - - - - - , , ,-
Abstrac_t This report describes the Duke Power Comr.any methodologies for: 1) simulating
-the FSAR Chapter 15 events characterized by multidimensional reactor transients, and ii) systematically confirming thet reload physics parameters important to Chapter 15 transients ard accidents are bounded by values assumed in the licensing analyses. The multidimensional reactor transients described are the rod ejection accident, the main steam line break, and the dropped rod ,
transient. The analytical approaches combine neutronics calculations with system and core hermal-hydraulics simulations. It is concluded that appli-cations of the methodologies and physics parameters checks will result in conservative predictions of the consequences, and that the applicable acceptance criteria are met. This report is applicable to the McGuire and Catawba Nuclear Stations. 1
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MULTIDIMENSIONAL REACTOR TRANSIENTS AND SAFETY ANALYSIS PHYSICS PARAMETERS Table of Contents
1.0 INTRODUCTION
2.0 DETERMINATION OF SAFETY ANALYSIS PHYSICS PARAMETERS 2.1 Overview 2.2 Generic Parameters _ 2.3 Discussion of FSAR Chapter 15 Transients and Accidents 2.3.1 Feedwater System Malfunctions That Result in a Reduction in Feedwater Temperature (15.1.1) 2.3.2 Feedwater System Malfunction Causing an increase in Feedwater Flow (15.1.2) 2.3.3 Excessive Increase in Secondary Steam Flow (15.1.3) 2.3.4 Inadvertent Opening of a Stear Gene ator Relief or Safety Valve (15.1.4) 2.3.5 Steam System Piping Failure (15.1.5) C.3.6 Loss of External Load (15.2.2) 2.3.7 Turbine Trip (15.2.3) 2.3.8 Inadvertent Closure of Main Steam Isolation Valves (15.2.4) 2.3.9 Loss of Condenter Vacuum (15.2.5) ; 2.3.10 Loss of Non-Emergency AC Power (15.2.6) 2.3.11 Loss of Normal Feedwater Flow (15.2.7) 2.3.12 Feedwater System Pipe Break (15.2.8) 2.3.13 Partial loss of Forced Reactor Coolant Flow (15.3.1) 2.3.14 Complete loss of Forced Reactor Coolant Flow (15.3.2) 2.3.15 Reactor Coolant Pump Locked Rotor (15.3.3) 2.3.16 Reactor Coolant Pump Shaft Break (15.3.4) 2.3.17 Uncontrolled Bank Withdrawal From Suberitical (15.4.1) 2.3.18 Uncontrolled Bank Withdrawal at Power (15.4.2) 2.3.19 Dropped Rod (s) and Dropped Bank (15.4.3) 2.3.20 Statically Misaligned Control Rod (15.4.3) 2.3.21 Single Control Rod Withdrawal (15.4.3) il l __ _ ___ _ ____ _ ____ .._ _______ __ _ _____________ _ ___ _
i Table of Contents [ cont'dl 2.3.22 Improper Startup of the Fourth Reactor Coolant Pump (15.4.4) 2.3,23 Moderator Dilution Accident (15.4.6) 2.3.24 Rod Ejection Accident (15.4.8) 2.3.25 Inadvertent ECCS Actuation (15.5.1) 2.3.26 CVCS Halfunction Resulting in Increase in Primary Inventory-(15.5.2) 2.3.27 Inadvertent Opening of a Pressurize- Relief or Safety Valve (15.6.1) 2.3.28 Instrument Line Rupture (15.6.2) 2.3.29 Steam Generator Tube Rupture (15.6.3) 2.3.-30 Lors of Coolant Accidents (15.6.5) 2.4 Reload Cycle Evaluation 3.0 CALCULATION OF KEY SAFETY ANALYSIS PHYSICS PARAMETERS 3.1 - Control Rod Worth Calculations 3.2 Reactivity Coefficients and Kinetics Parameters 4.0 ROD EJECTION ANALYSIS 4.1 Overview 4'1.1 Description of Rod Ejection Accident 4.1.2 Acceptance Criteria 4.1.3 Analytical Approach 4.2 Simulation Codes and Models
.4.2.1 Nuclear Analysis 4.2.2 Core _ Thermal-Hydraulic Analysis L 4.2.2.1 VIPRE-01 Code Description 4.2.2.2 Fuel Temperature and.Enthalpy Calculation j 4.2.2.3 ~ Coolant-Expansion Rate Calculation-L -4.2.2.4 DNBR Evaluation 4.2.3- System Thermal-Hydraulic Analysis
[ 4.3 AR20TTA Analysis
.d.3.1 ' Initial. Conditions ~4.3.2 Boundary Conditions 4.3.3 :Results iii l
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Table of Contents (cont'd) 4.4- VIPRE Analysis 4,4.1 Elnitial Con'ditions . 4.4.2- Fuel Temperature and Enthalpy L4.4.3 Coolant Expansion Rate 4;4.4- DNBR and Fuel Pin Census 4.5; RETRAN- Analysis 4.5.1 Init'ial Conditions 4.5.2 Boundary Conditions
- - 4. 5. 3 Results
- 14. 6 Dose Con 2seguences '
4.7 Cy_cle Specific Evaluation. 15.0 STEAM.LINE BREAK ANALYSIS 5.1 Overview 5.1.1 Description of Ste:m Line Break Accident 5.1.2 Acceptance Criteria 5.1.3 Analytical = Approach 5.2: Simulation Codes and Models 5.2.1 System Thermal-Hydraulic Analysis 5.2.1.1' Selection of a Bounding Unit 5.2.1.2 Modifications ~to Base Plant Model . 5.2;-1.3 Break Modeling 5.2.2. Nuclear Analysis 5.2.2.1 Core Physics' Parameters
- 5.2.2.2 Power: Distributions
- 5. 2. 3 - ' Core Thermal-Hydraulic Analysis 5.2.3.1 -VIPRE Code Description 5.2.3.2 Analysis Methodology ,
5.3 Tran s i en t - Analysi s-- 5.3.1- Initial Conditions 1 5.3.2 Boundary Conditions
.b.5.2.1 - Availability of Systems and Components 5.3.2.2- Response Times 5.~3.2.3: Flow From Interfacing Systems iv 4
.v.- ,p _-.% , , s. .- . _ . . . _ . . , - - . . _ .._ -e-_,- _--_ .
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Table of Contents Leont'd) 5.3.2.4 Engineered Safety Features Actuation Setpoints 5.3.2.5 Boron injection Modeling 5.3.2.6 Core Kinetics Modeling 5.4 _Results and Conclusions. s 5.4.1 Primary and Secondary System Response , 5.4.2 . Core. Response 5.4.2.1 -Axial and Radial Power Distributions 5.4-2.2. Minimum DNBR Results 5.5 Cycle specific Evaluation 6.0 DROPPED ROD ANALYe'" 6.I' Overview 6.1.1 Description of Dropped Rod Accident . 6.1.2 Acceptance Criteria 6.1.3 Analytical Approach 6.2 -Simulation Codes and Models
-6.2.1 System Thermal-Hydraulic Analysis 6.2.2 Nuclear Analysis 6.2.3 : Core Thermal-Hydraulic Analysis 6.3 - -Transient Analysis-6.3.1 Initial-Conditions 6.3.2: Boundary Conditions-
- 6. 3. 2. ~1. Physics Parameters.
.- '6.3.2.2 Reactor-Protection-System
-6.3.2.3 Power Range Nuclear Instrumentation 6.3.2.'4 Rod Control System-6.3.2.5 -Pressurizer Pressure and Level Control-4 6.3.2.6 Main Feedwater and Turbine Control 6.4 Results and Conclusions 6.4.1 System Transient Results.
6, 4 .1.'l - Typical Beginning-of-Cycle Response [ 6.4.1.2- Typical End-of-Cycle Response g 6.4.1.3 Limiting Statepoint' Selection i: 6.4!2 Core Response ! v l
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Table of Contents (. cont'd) 6.4.2.1 Statepoint Conditions-6.4.2.2 Thermal-Hydraulic Results for DNBR
- 6. 5 - Cycle Specific Evaluation Appendix A - REAC10R VESSEL THERMAL MIXING EVALUATION f.
l-I l I r I o l. i. Vi , __ _ - _ , _ , . , . . . . , , .,.~,.,--,.. .,.m .-r , . . , . -
List of figures F_iqur_e Tit 1e 1-1 Reload Core Safety Analysis Verification Process 4-1 , Catawba 1 Cycle 2 Assembly Enrichments and Fuel Exposures 4-2
.. a VIPRE Model for Fuel Temperature and DNBR Calculations 4-3
[ [VIPRE Model,r Reactor Coolant Expansion Rate Calculation 4-4 Rod Ejection Accident Control Rod Locations 4-5 BOC, HFP ARO Power Distributions 4-6 EOC, HFP ARO Power Distributions 4-7 Trip Reactivity Curve 4-8 BOC, HFP Core Power vs. Time ' 4-9 BOC, HFP Power Distribution at 0.0 seconds 4-10 BOC, HFP Power Distribution at 0.09 seconds 4-11 BOC, HZP Core Power vs. Time 4-12 BOC, HZP Power Distribution at 0.0 seconds 4-13. BOC, HZP Power Distribution at 0.2 seconds 4-14 BOC, HZP Power Distribution at 0.34 seconds 4-15 -E00,-_HFP Core Power vs. Time 4-16 EOC- HFP Power Distribution at 0.0 seconds 4-17 EOC, HFP Power Distribution at 0.09 seconds 4-18 E00, HZP Core Power vs. Time 4-19 EOC, HZP Power Distribution at 0.0 seconds 4-20' EOC, HZP Power Distribution at 0.13 seconds 4-21 'E00, HZP Power Distribution at 0.17 seconds 4-22 BOC, HFP MARP-Curves 4-23 BOC, HZP MARP' Curves 24 EOC, HFP MARP Curves 25 EOC, HZP MARP Curves 4-26 80t. HFP Number of Pins in DNB by Assembly 4 BOC, HZP Number of Pins in DNB by Assembly 4-28 EOC, HFP Number.of Pins in DNB by Assembly 4-29 EOC, HZP Number of " ins in DNB by Assembly vii
List of Figures (cont'd) Fiqure Title 4-30 HFP BOC Case Core Coolant Volume Expansion Rate 4-31 HFP BOC Case RCS Pressure Response 5-1 RETRAN Reactor Vessel Model 5-2 K-effective vs Moderator Temperature 5-3 VIPRE Model 5-4 Offsite Power Maintained Steam Line Pressure 5-5 Offsite Power Maintained Cold Leg Temperature 5-6 Offsite Power Maintained Hot Leg Temperature 5-7 Offsite Power Maintained Core Baron Concentration 5-8 Offsite Power Maintained Core Reactivity 5-9 Offsite Power Maintained Neutron Power 5-10 Offsite Power Maintained Core Heat Flux 5-11 Offsite Power Maintained Pressurizer level 5 Offsite Power Maintained Pressurizer Pressure 5-13 Off site Power Maintained Break Mass Flow Rate - 5-14 Offsite Power Maintained Core Mass Flux 5-15 Offsite Power Lost Steam Line Pressure - 5-16 Of fsite Power lost Cold Leg Teinperature 5-17 Offsite Power Lost Hot Leg Temperature 5-18 Of fsite Power Lost Core Boron Contentration 5-19 Offsite Power Lost Core Reactivity 5-20 Offsite Power Lost Neutron Power 5-21 Of f site Power Lost Core Heat Flur 5-22 Offsite Power lost Pressurizer Level 5-23 Offsite Power Lost Pressurizer Pressure 5-24 Offsite Power Lost Break Mass Flow Rate 5-25 Offsite Power Lost Core Mass Flux 5-26 Offsite Power Maintained Typical Assembly Radial Power Distribution 5-27 Of f site Power Lost Typical Assembly Radial Power Distribution ' 6-1 VIPRE[ ]Model 6-2 BOC F-Delta-H vs. Dropped Rod Worth 6-3 MOC F-Delta-H vs. Dropped Rod Worth viii l
List of Figures (cont'd) i f_isun Title 6-4 EOC F-Delta-H vs. Dropped Rod Worth 6-5 Minimum Tilt vs. Dropped Rod Worts 6-6 100 pcm BOC Case Actual and Indicated Reactor Power 6-7 100 pcm BOC Case Control Bank D Position 6-8 100 pcc SOC Case Reactor Coolant System T-ave 6-9 100 pcm BOC Case Pressurizer level . 6-10 100 pcm BOC Case Pressurizer Pressure 6-11 400 pcm EOC Case Actual and Indicated Reactor Power 6-12 490 pcm EOC Case _ Control Bank D Position 6-13 400 pcm E0C Case Reactor Coolant System T-ave 6-14 400 pcm EOC Case Pressurizer Level 6-15 , 400 pcm EOC Case Pressurizer Pressure 6-16 pcm BOC Cases Reactor Power 6-17 pcm BOC Cases Reactor Coolant System T-ave 6 pcm BOC Cases Pressurizer Pressure F-19 pcm MOC Cases Reactor Power 6-20 pcm MOC Cases Reactor Coolant System T-ave 6-21 pcm MOC Cases Pressurizer Pressure 2? pcm EOC Cases Reactor Power 6-23 pcm EOC Cases Reactor Coolant System T-ave 6-24 pcm EOC Cases Pressurizer Pressure 6-25 pcm Limiting Cases Reactor Power - 6-26 pcm Limiting Cases Reactor Coolant System T-ave 6-27_ pcm Limiting Cases Pressurizer Pressure 6-28 - Axial Power Shapes A-1 Mixing Test The mocouple Locations -
-A-2 -A-3 A-4 A-5
, A-6 A-7 ix
l List of Tables 1 Sur-mary of Safety Analysis Physica Parameters 4-1 Rod Ejection Transient Kinetics Parameters - J 4-2 Rod Ejection Transient Initial Conditions 3 Rod Ejection-ARROTTA Resuits 4-4 Rod Ejection Reload Checklist 5-1 Sequence of Events for 1.4 ft' Split Break with Offsite Power Maintained - 5-2 Sequence of Events for 1.4 ft' Split Break with Offsite Power Lost at 51-l Actustion L b E X 1 . g-:rm wp - *mmtm3*gN1-Sw"g-+WyvP % Mem--w~4'"m'T yg+p-v= M w TY f-'"
_. _ . ~ . . .. _- _ . _. _ . _ _ _ _ . . = List of Abbreviations ARD All Rods Out BOC Beginning of Cycle CETC Core Exit Thermocouples " CHF Critical Heat Flux CVCS Chemical and Volume Control System DEG Double-ended Guillotine DNBR Depar'.ure from Nucleate Boiling Ratio DTC Doppler lemperature Coefficient ' ECCS Emergency Core Cooling System EFPD Effective Full Power Days EOC End of Cycle Fah Ratio of the integral of linear power along the rod to the average rod power Fq Maximum local heat flux / average fuel rod heat flux FSAR Final Safety Analysis Report ilFP Hot Full Power
' HIP Hot Zero Power KW/FT Kilowatts per foot LM Limiting Mixing LOCA Loss of Coolant Accident MARP Maximum Allowab'e Radial Power MTC Moderator. Temperature Coefficient NI Nuclear Instrumentation DAC Operatcr Aid Computer PORV. Power-Operated Relief Valve PZR Pressurizer RCS Reactor Coolant Systei RIL Rod Insertion Limits
.RPS Reactor Protection System RTD Resistance Temperature Detector SCD' Statistical Core Design SG Steam Generator SI Safety Injection swd Steps withdrawn TM Thermal Mixing l x1
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l l 1.0 INIR000CT10N This report describes the methodologies to be used by Duke Power Company to simulate multidimensional reactor transients and to verify that the key physics parameters calculated for a reload core are bounded by values assumed in the licensing Chapter 15 analyses. This report is applicable to the McGuire and Catawba Nuclear Stations, which are 4-loop 3411 MWt Westinghouse units. These methodologies expand on the NRC-approved reload design methods of DPC-NF-2010A (Reference 1-1) and on the system and core tran tient thermal-hydraulic simulation methods of DPC-NE-3000 (Reference 1-2). Chapter 15 accident analyses show that the design of a reactor and its asso-tiated systems will mitigate the events of various postulated accidents and ensure that the consequences of these accidents are acceptable. These analy-ses, hereafter referred to as the " reference safety analyses", along with the facility Technical Specifications, establish the bases and conditions for safe operation of the plant. Important parts of t..e reference analyses include values of parameters assumed in the analyses, performance characteristics of the mitigating systems, and the analytical models used. Values of parameters selected in the reference analyses are chosen to bound values expected during the life of the plant. Performance character'.stics of the mitigating systems are modeled to give conservative performance characteristics and produce bounding consequences for each of the accidents. For each fuel cycle design, each reference analysis is validated by examining all of the key physics, thermal-hydraulic, and mechanical parameters which are assumed in the reference analysis and whicn might be affected by a reload design. These values are compared to those calculated for the particular cycle. If all parameters are within the envelope of values assumed in the reference analysis, then the analysis is valid and no reanalysis is necessary, if, however, one or more of the plant parameters assumed in the reference analysis are found to be nonconservative for the reload cycle, those accident analyses which are affected by the nonconservative parameters must be reeval-uated or the loading pattern must be revised. This validation process is shown schematicalij in Figure 1-1. 1-1
input to these checks will come from various groups associated with the reload design and safety analysis of reload cores including nuclear design, safety analysis, mechanical, and thermal-hydraulic groups. This checklist concept applies only to changes in important p;rameters resulting from refueling of the reactor core. Other changes in plant systems or setpoints might necessi-tate the reanalysis of certain transients independently of the reload design process, in some cases, the effects of plant modifications might be incorpo-rated into the reload analysis. The first two chapters of this report con-centrate on the generation of key physics parameters and the methods for using these parameters to validate existing safety analyses. Chapter 2 specifies those physics parameters deterr.ined to be important for each Chapter 15 event. The appropriateness of selecting a maximum, minimum or nominal value for each parameter is justified. Future reanalyses of Chapter 15 transients would use the specifications of Chapter 2 to determine physics data required to perform a conservative analysis. Chapter 3 describes the nuclear design methods employed to calculate values of the important safety analysis physics parameters. These parameters can be influenced by core composition, boron concentration, control rod position. power level, xenon distribution, and other considerations. The approach taken to determine a conservative value of a parameter is to utilize the results of Chapter 2, and then investigate combinations of the above factors as permitted by Technical Specifications. Three FSAR Chapter 15 events involve significant asymmetric core power peaking and require evaluation of the core response from a multidimensional simulation perspective. These events are the steam line break (15.1.5), the dropped rod transient (15.4.3), and the rod ejection accident (15.4.8). In ceder to conservatively predict the transient response, a combination of neutronic, system thermal-hydraulic, and core thermal-hydraulic simulation codes is employed. Depending on the dynamics of the particular analysis, it is possi-ble in some situations to adequately and conservatively model aspects of the transient with static methods. Otherwise, an explicit transient evaluation is performed. The analyses presented are intended to bound future reload core designs. As such, a cycle-specific check of important safety analysis physics 1-2
4 parameters and/or a ilmited scope analysis will be all that is necessary to confirm that the existing analysis results remain bounding. The rod ejection accident analysis methodology is presented in Chapter 4. The rapid core transient response is simulated with the three-dimensional space-time transient neutronics nodal code ARROTTA (Reference 1-3). The rod elec-tion accident is analyzed at full power and zero power at both beginning and ! end-of-cycle. The core thermal response is modeled with the VIPRE-01 (Refer- ! ence 1-4) code. Peat fuel enthalpy, a core-wide DNBR evaluation and transient core coolant expansion are calculated, The Reactor Coolant System pressure response is simulated with the RETRAN-02 (Reference 1-5) code. The results of the analysis are shown to meet all acceptance criteria, lhe stean line break accident analysis methodology is presented in Chapter 5. The system thermal-hydraulic analysis is performed with RETRAN-02. The worst case scenario, which occurs at zero power at end-of-cycle, is presented. Cases both with and without -of f site power are analyzed. The core power ' peaking at the return-to power statepoint condition and including the worst stuck rod is determined. The approach to DNBR is then predicted with VIPRE-01. The results show that the DNBR limit is not exceeded for the limiting cases. The dropped rod transient analysis methodology is presented in Chapter 6. The transient response to single and multiple dropped rods from within the same group are evaluated. A complete range of dropped rod worths at beginning, middle, and end-of-cycle are analyzed. The system thermal-hydraulic analysis is performed with RETRAN-02. The core power peaking at the limiting. statepoints'is evaluated with VIPRE-01 to demonstrate that the DNBR limit is
-not exceeded.
-The analyses presented in Chapters 4-6 of this report are intended to replace the existing FSAR Chapter 45 analyses. Reanalysis of these accidents in the L future by Duke Power Company will use the methods described in this report.
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r , _ References l l 1-1 Nuclear Physics Methodology for Reload Design, DPC-NF-2010A, Duke Power Co., June 1985. 1-2 Thermal-Hydraulic Transient Analysis Methodology, OPC-NE-3000, Duke Power Co., July 1987. 1-3 'ARP.0TTA: Advanced Rapid Reactor Operational Transient Analysis Computer Code, Computer Code Documentation Package, EPRI. 1-4 VIPRE-01: A Thermal-Hydraulic Code for Reactoi 0 es; EPRI NP-2511-CCMA Revision 2. EPRI, April, 1987. I 1-5 RETRAN A Program for Transient Thermal-Hydraulic Analysis of Complex Floid Flow Systems, EPRI NP-1850-CCMA Revision 4. EPRI, November 1988, i l l' l l 1-4 I
Figure 1-1 Reload Core Safety Isnalysis Verification Process Reload Cycle
- Design h
1 reload cycle design parameters ! Y Reload Parameter all values conservative
)
Comparison ( l some values Current S a fe tI' non-conservative y Analyses Valid Detailed Review of Analyses j i effects acceptable u Revisri Licensing Basis c / effects uncertain or Analyses
\ unacceptable F
Reanalysis of consequences Event (s ) acceptable consequences unacceptable 1-5 l
2.0 DETERMINATION OF SAFETY ANALYSIS PflYSICS PARAMETERS 2.1 Overview FSAR Chapter 15 transients and accidents must be conservatively analyzed to ensure that the applicable fuel design limits, system overpressure design limits, and dose consequences are not exceeded. Each transient and accident , analysis incorporates a set of assumptions, which when combined in a consis-tent.or conservative manner, produce conservative analysis results. These analyses bound the licensed operating conditions and modes for the current plant design and fuel cycle. An important subset of the analysis assumptions includes the core physics parameters necessary to characterize the initial conditions and transient response of the core. The relative importance of various physics parameters and the sensitivity to variations in the values of
'the parameters varies between transients. However, it is possible to identify, for each event, a set of physics parameters which are significant and directly affect the results of the analysis. Once these key parameters have been determined, then the impact of variation in the range of values due to a_ change in the core loading pattern and eperating history can be assessed.
A conservative or consistent value can then be selected for analysis, or several combinations can be analyzed to ensure the transient response is-bounded. The purpose of this chapter is to review and identify the key physics parame-ters for each FSAR Chapter 15 event. The conservative direction for each parameter (e.g., minimum / maximum) is identified where important. Table 2-1 summarizes the key parameters identified in this chapter. The actual analysis values are:not provided but can be obtained by referencing the current valid licensing analysis for each event and plant. 2.2. Generic parameters
- Some of the important safety analysis physics parameters can be considered generic in that the value of the parameter is-important for many transient analyses. The descriptions of the following generic parameters are not repeated for each specific transient in Section 2.3.
2-1 , s - .,_,--r.-- ., ,.-,~.,.-m,-m+. S.,r. .,. . , , , ,-. ~ _ e , _ ,m... . m.
_m__. _ _ _ . _ _ . _ _ _ _ _ _ - . _ . _ _ . _ . _ . . . _ . . - _ . _ _ _ . _ _. - - _ _ . __ Reactivity _ Insertion followinq_.R uctor Trip The reactivity insertion following reactor trip is a combination of a minimum available tripped rod worth and a normalized reactivity insertion rate. The minimum available tripped rod worth assumed in safety analyses must ensure, as i a minimum, that the shutdown margin in Technical Specifications is preserved. , This shutdown margin assumes that the most reactive rod remains in the fully withdrawn position and that the other control rods drop from their power dependent insertion limits. The normalized reactivity insertion rate is determined by bounding control rod drop times as determined by plant testing, i and by developing a. conservative relationship between rod position (*. inserted) and normalized reactivity worth. Initial Core Power Distribution
- Technical Specifications require that the core power distribution remains within prescribed. limits during power operation. These power peaking limits are typically expressed as limits on total peak (F9) and radial peak (Fg).
Fg limits are typically a function of elevation in the core and might also vary as a function of burnup and power level. The transient and accident analyses assume that any core power distribution permitted within normal operating limits is a valid initial condition. For those transients *in which the initial power distribution has a significant impact on the course of the event, perturbed power distributions allowed by operating limits are considered. These events are discussed individually in Section 2.3. Effective Delayed Neutron Fractions and Decay Constants The delayed neutron parameters are mainly important during rapid reactivity excursion transients. For an event such as the rod ejection accident, the minimum value of the effective delayed neutron fraction (beta-effective) is conservative, If the transient is not characterized by a rapid change in reactivity, then the value of-beta-effective'is not significant. The. values of the fractions and decay constants for each delayed neutron precursor group are not key parameters, and typical values are sufficisat. 2-2 g g --wy
- 9.. )-M.i cyyg -ri i ye yy .y g. y s mvvp-is---r w y yiweim-pyg , 9ga-y-.- 'r4m-m.-.-w--m-u-,,w -g m r-wyw ~ p-- - .
i l i prompt Neutron lifetime The prompt neutron lifetime is mainly important during rapid reactivity excursion transients. This parameter is not a key parameter, and so typically beginning t.d end of-cycle values are used consistent with the limiting core condition for the transient. Initial fuel Temp _eratures Both the initial core average fuel temperature and the initial hot spot temperature are important to a conservative evaluation of the transient core response. These temperatures are determined using approved methods (Reference 2-1). The initial hot spot temperature is determined in a manner consistent with the initial power distribution and appropriate hot channel factors. Fuel temperatures are assumed to be conservative when taken at the maximum values. 2.3 Dis _cussion_of FSAR Chapter 15 Transients and Accidents 2.3.1 Feedwater System Malfunctions That Result in a Reduction in ' Feedwater Temperature (15.1.1) P' This transient is bounded by 15.1.2 and 15.1.3, which are discussed in Sec-tions 2.3.2 and 2.3.3. 2.3.2 Feedwater System Malfunction Causing an increase in Feedwater Flow (15.1.2) < This transient is initiated by a failed-open main feedwater control valve which results in an increase in main feedwater flow. Due to the increase in the secondary heat sink, the primary coolant temperature decreases. The transient response is most conservative in the presence of a most negative moderator temperature coef ficient (MTC) which will result in the maximum increase in reactor power. Similarly, a least negative Doppler temperature coefficient (DiC) is conservative since this maximizes the core power response. The HTC and DTC are the only key physics parameters for this event. 2-3
l l 2.3.3 Excessive increase in Secondary Steam riow (15.1.3) This transient is initiated by an increase in secondary steam flow, which can result from the turbine governor valves opening or from a spurious steam dump to condenser event. Due to the increase in the secondary heat sink, the primary coolant temperature decreases. Similar to the discussion in Section 2.3.2, a most negative MTC and a least negative DTC are conservative modeling assumptions. Additional analyses are typically performed with the Rod Control System in automatic for both the above MTC/DTC case and a case with minimum reactivity feedback. These cases require conservative modeling of the Rod _ Control System and associated physics parameters. 2.3.4 Inadvertent Opening of a Steam Generator Relief or Safety Valve ( 15.1. 4 ) This transient is it.'tiated by uncon**o led secondary depressurization result-ing from a failure of a secondary steam oimp valve, safety valve, or PORV. The worst case scenario begins from a no-load condition. The resulting primary system overcooling can cause a loss of core shutdown and a return-to-power can occur prior to boron injection from the actuation of safety injec-tion. Since the steam system piping failure (15.1.5) is analyzed to the acceptance c.riteria that are applicable to the 15.1.4 transient, and since it bounds 15.1.4 in all aspects, there is no basis for analyzing 15.1.4. The 15.1.5 analysis can sirnply be referenced. 2.3.5 Steam System Piping Failure (15.1.5) This transient is initiated by a rupture of a main steam line. The worst case scenario begins from a no-load condition. The resulting primary system over-cooling causes a loss of core shutdown and a return-to power conditinn occurs. This trar, stent is analyzed by assuming a conservatively large reactivity insertion as the core cools down. The power increase is exacerbated by assuming a least negative Doppler coefficient. The boron concentration in the safety injection flowpath and the baron worth are both minimized. Due to the assumption of a stuck rod, the core power distribution at the limiting statepoint will be highly peaked. Consequently, the core power distribution 2-4 l ________-_ _ _ ___ __ _____ _ _ -___ ___--__--- - - -. - A
I must be_ evaluated to quantify the number of fuel pins exceeding the DNBR limit. 2.3.6 Loss of External Load (15.2.2) i This transient is bounded by 15.2.3, which is discussed in Section 2.3.7.
-2,3.7 Turbine Trip (15.2.3)
This transient is initiated by a rapid closure of the turbine stop valves, resulting.in a decrease in the secondary heat sink. As a result, primary coolant temperatures increase. The transient response is most conservative at beginning-of-cycle where the MTC and DTC are least negative. The least aegative HTC and DTC maximize the pre-trip core power response. With this approach, the mismatch between heat source and heat sink is conservatively maximized. 2.3.8 Inadvertent Closure of Main Steam Isolation Valves (15.2.4) This transient is bounded by 15.2.3, which is discussed in Section 2.3.7.
- 2. 3. 9. -
Loss of Condenser Vacuum (15.2.5) This transient is bounded by 15.2.3, which_is discussed in Section 2.3.7. 2.3.10 Loss of Non-Emergency AC Power (15.2.6) This transient is initiated by a loss of non-emergency AC power. Similar to the turbine trip transient in Section 2.3.7, this transient is basically a loss-of-heat sink event. -Therefore, the conservative physics parameters are least negative KIC and DTC, 2.3.'11 _ Loss of Normal feedwater Flow (15.2.7) This transient is_ initiated by a loss of main feedwater flow. Similar to the
-_ turbine trip transient in Section 2.3.7, this transient'is basically a loss of l
2-5
~_ _ - _ _ _ _ _ _ _ _ _ _ _ ___ _ - , . ,s., __ _ _
l i
- bes . sink event. Therefnre, the conservative physics parameters are least
, negative Mit and DTC. 2.?.12 feedwater 5), tem pipe Great (15.2.8) This transient is initiated by a rupture of a main feedwater line. Similar to the turbine trip transient in Section 2.3.7, this transient is basically a loss of heat sink event, lherefore, the conservative physics parameters are least negative MTC and DTC. 2.3.13 Partial loss of forced Reactor Coolant flow (15.3.1) This transient is initiated by a trip of one reactor coolant pump. Due to the decrease in core flow, the core average moderator temperature increases. In ordet to maximize the pre-trip core power response and conservatively predict the minimum DNBR, least negative MTC and DTC are appropriate. 2.3.14 Complete loss of forced Reactor Coolant flow (15.3.2) This transient is initiated by a simultaneous trip of all four reactor coolant pumps. Similar to the single reactor coolant pump trip transient in Section 2.3.13, the conservative physics parameters are least negative MTC and DTC.
~
2.3.15 Reactor Coolant pump Locked Rotor (15.3.3) This transient is initiated by an instantaneous seizi're of a reactor coolant pump which results in a rapid decrease in loop and core flow. Similar to the single reactor coolant pump trip transient in Section 2.3.13, the conservative physics parameters are least negative MTC and DTC. 2.3.16 Reactor Coolant Pump Shaft Break (15.3.4) This transient is bounded by 15.3.3, which is discussed in Section 2.3.15. 2-6
2.3.17 Uncontrolled Bank Withdrawal From Subcritical (15.4.1) This transient is initiated by a malfunction in the Rod Control System which results in the withdrawal of two sequentici control banks from a subcritical condition. In order to maximize the pre-trip core power response, a most positive MTC and a least negative DTC are conservative assumptions. The reactivity addition rate resulting from the rod withdrawal is taken to be the maximum credible value. This value is a combination of two sequential control banks moving at maximum speed over the span of rod positions resulting in the maximum differential summed rod worth. 2,3.18 Uncontrolled Bank Withdrawal at power (15.4.2) f This transient is initiated by a malfunction in the Rod Control System which results in the withdrawal of two sequential control-banks at power. Unlike most other transients, the uncontrolled rod withdrawal at power analysis typically requires that a soectrum of cases be analyzed in order to confirm that the minimum DNBR limit'is not exceeded. Due to the increase in core power peaking at reduced power levels, the analyses must consider all power levels at viable worst case initial conditions. The limiting reactivity addition rate is also not obvious and so all rates up to the maximum credible value (refer to Section 2.3.17) must be considered. A most positive MTC is combined with a least negative DTC, The impact of the rod withdrawal on the core power distribution is another parameter requiring evaluation. 2.3.19 DroppedRod(s)andDroppedBank(15.4.3) .This transient is initiated by a malfunction in the Rod Control System which results in one or more rods from.the same group dropping into the core, Key physics parameters include the dropped rod worth, the total worth of the controlling rod groups which are available for withdrawal, the flux incident on the excore power range flux detectors, and the post ~ drop core power distri-button. In order to bound the effect-of thermal feedback, bounding values for HTC and DTC as a function of core burnup must be analyzed. The dropped bank transient generally results in a reactor trip on low pressurizer pressure. For those_ events which do not result in a reactor trip, the core 2-7
power peaking is bounded by the dropped rod (s) transient due to the symmetric nature of a dropped rod bank. Therefore, the dropped rod bank is not ana-lyzed. ' i 2.3.20 Statically Misaligned Control Rod (15.4.3) i 1 The statically misaligned control rod evaluation considers the situation where i one Bank D control rod is mispositioned relative to the remaining Bank D rods. l The single rod can be at any position while the bank is within its normal operating limits. The important physics parameter is the core power distri-bution resulting from the asymmetric condition. l l 2.1.21 $1ngle Control Rod Withdrawal (15.4.3) l This transient is initiated.from full power by a spurious withdrawal of one ' Bank D rod. Key physics parameters include the integral worth of the Bank D rod beginning from the full power insertion limit, the flux incident on the excore power range flux detectors, and the core power distribution resulting from the asymmetric condition. In order to maximize the core power response, a least negative MTC and a least negative DTC are selected. 2.3.22 Improper Startup of the Fourth Reactor Coolant Pump (15.4.4) This transient is initiated by an improper matJal restart of the fourth reactor coolant pump while at power. Due to the resulting increase in core flow, core average temperature decreases. Cold water originally in the idle loop is also transported towards the core by restarting the pump. In order to conservatively maximize the core power response, a most negative MTC and a least negative DTC are selected. 2.3.23 Moderator Dilution Accident (15.4,6) Moderator dilution events can result from malfunctions or misoperation of the 1 makeup and letdown systems. These events can occur in various operating modes as detai' led in the FSARs. The important physics parameters in each mode are the seme. These are the critical boron concentration and the initial boron 2-8
f concentration. The initial boron concentration is determined by a prescribed initial value or by adding to the critical boron concentration the mode-specific shutdown margin converted to ppmb. In order to be conservative, the boron concent..tions should be large, since the effect of a dilution will be greater. The baron worth used to convert the shutdown margin to ppmb should be conservatively large. 2.3.24 Rod Ejection Accident (15.4.8) The rod ejection accident is initiated by a mechanical failure of the control rod drive mechanism pressure hooing. The event is evaluated at both hot full l power and hot zero power conditions at both beginning and end-of-cycle. For each condition, the physics parameters are selected in a consistent manner to p conservatively bound the transient response. A conservatively high ejected rod worth is evaluated. The MTC is specified as least negative to minimize negative reactivity addition via thermal feedback. The DTC is specified as least negative to minimize thermal feedback and maximize core power response. Beta is also minimized to maximize the core power response. The resulting core power distribution including the maximum total peak are key parameters. 2,3.25 Inadvertent ECCS Actuation (15.5.1) This transient is initiated by a spurious actuation of the Emergency Core
}
Cooling System, which results in boron injection into the primary system. Reactor power decreases slowly until a reactor trip occurs. All thermal margins increase during this transient, and there are no important physics parameters. 2.3.26 CVCS Malfunction Resulting in Increase in Primary Inventory (15.5.2) 1his transient is bounded by 15.5.1.
_ _ _ _ - - _ _ _ . . _ . _ . . _ _ _ . .. __.__ ___ _ ____ - _ - . _ ____.__.m+ 2.3.27 Inadvertent Opening of a pressurizer Relief or Safety Valve (15.6.1) This transient is initiated by a spurious lifting of a pressurizer relief or safety valve and a failure to close. A loss of primary coolant results and i primary pressure decreases until reaching the reactor trip setpoint. There are no important physics parameters associated with this event. 2.3.28 Instrument Line Rupture (15.6.2) lhis transient, similar to 15.6.1, does not involve any important physics parameters. > t 2.3.29 Steam Generator Tube Rupture (15.6.3) ; This transient, similar to 15.6.1, does not involve any important physics parameters. 2.3.30 Loss of Coolant Accidents (15.6.5) L The only important physics parametcr for LOCA is the initial power distri-bution. Linear heat flux (kw/ft) limits are established as a function of core elevation. These-limits may also account for differences in fuel assembly design and burnup. L - i- 2.4 Reload Cycle Evaluation The important physics parameters in Table 2-1 are evaluated each reload cycle to ensure that values assumed in the current licensing analyses bound the reload core. Accidents for which the physics parameters are not bounced would be reevaluated to ensure acceptable accident consequences or the core would be redesigned so the physics parameters fall within the limits assumed in the Lreference analysis. ! 1 2-10 l l l l
- ,, M. - > - , ra - . . .#. - - . r-r -. , . - - . , - -ev--...r.m,- . ,m,_-,w . - , + . - - - *. - - . , - - - . e,.. e ,-,---+,.w.-me v--v.
I 1
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l 1 Re f e rence s..
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1 2-1 Y. H. Hsil, et al., TACO 2: Fuel Pin Performance Analysis, Revision 1, 1 DAW-10141PA, June 1983, l 1 l j l B 11 L
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Table 2-1 5 mary of_ Safety Analysis Physics Paraaeters FSAR Cor.servative Report Key Paramete-s Direction Section Teansient Or Accident Section 3
- Raactivity insertion - Minimum wo-th Generic N/A 2.2 following reactor trip - Slowest inseetion
- Initial core power - Maximum power distribution ceaking per Tech 5pec
- Effective delayed - Minimum for rapid neutron fraction and -eactivity tesn-decay constant: sients
- Ma*imu- for all other transients va - Ncminal precursor L
group fractions and decay constants
- Initial fuel tamperatures - Maximum
- MTC/DTC Case 2 - Most negativt I'.3.2 Feedwater flow increase 15.1.2 - Least negative 2.3.3 Increase in steam flow 15.1.3
- MTC/DTC Case 2 - Zero MTC (Increase in stean flow only) - Least negative (DTC)
- MTC - Mest negative 2.3.4 SG safety valve failure 15.1.4 - teast cegative 15.1.5
- DTC 2.5.5 Steam line break
- SI boron concentration - Minieum
- Baron worth - Minimum
- Core power distribution - Maximu peaking with stuck rod I
- Least negative Turbine trip 15.2.3
- MTC 2.3.7 15.2.6
- DTC - least r.ecative 2.3.10 Loss of AC power 2.3.11 Loss of feedwater flow 15.2.7
t 4, , Table 2-1 (coi .) Report FSAP. Section Conservative Transient Or Accident Section Key Parameters ___ Direct f or. 2.3.12 Feedwater line break 15.2.8 2.3.13 Partial loss of flow 15.3.1
- MTC - Least negative 2.3.14 Complete loss of flow 15.3.2
- DTC - Least negative
- 2.3.15 Locked rotor 15.3.3
- Core power distribut*ur - Maximize number of (locked rotor only) pins in DN3 2.3.17 Uncontrolled rod withdrawal 15.4.1
- MTC - Most positive from subcritical '
- DTC - Leas'. negative i
- Reactivity addition rate - Maximum 2.3.18 Uncontrolled rod withdrawal 15.4.2
- MTC/DTC - Most positive MTC
- Least negative DTC
- Excore detector signal - Minimin indicated vower i
- Reactivity addition rate
- Small to maximum 2.3.19 Dropped rod (s) 15.4.3
- MTC - Bounding vs burnup Dropped rod bank - DTC - Bounding vs burnup
- Dropped red worth - Small to maximum
- Available red worth - Maximum for withdrawal
- Excore detector tilt - Minimum indicated .
power I
- Core power distribution - Maximum peaking l with dropped rod j
. 2.3.20 Statically misaligned rod 15.4.3
- Core power distribution - Haximum peakina i wit.' misaligned rod -
i I l
i Table 2-1 (cont'd) l FSAR Conservative l Report Direction Transient Or Accident Section Key Paramete;_s Section 15.4.3
- MTC - Least negative j 2.3.21 Single rod withdrawal - Least negative
- DTC
- Worth of sinol? rod - Maximan
- Core power dis -ibution - Maximize number with rod wittidrawn of pins in Drib
- Excore detector tilt - Minierum indicated power Fourth RCP startup 15.4.4
- MTC - Most negative 2.3.22
- DTC - Least negative 2.3.23 Moderator dilution 15.4.6
- Critical boron concentration - Highest
- Initial baron concentration - Closest to m critical concen-1 tration 15.C.8
- MTC - Most positive 2.3.24 Rod ejection
- DTC - Least negative
- Ejected rod worth - Maximum I
- Beta-effective - Minimum
- Core power distribution - Maximum total with ejected rod peak j
- Maximize number t.
of pins in DNB 15.6.5
- Initial core power - Maximum kw/ft 2.3.30 Loss of coolant accident vs. core e'evation distribution
; a
, s
3.0 CALCULA110N OF KEY SAFETY ANALYSIS PilYSICS PARAMETERS 1hree-dimensional core models such as EPRI-NGDE-p (Reference 3-1) and SIMU-LATE-3P (Reference 3-2) are used to calculate core physics parameters and power distributions, in some cases, simpler two-dimensional calculations may be performed with P0Q (Reference 3-1) or SIMULATE-3p. In these cases, appro-priate corrections for flux redistribution effects are made. Core physics parameters are calculited as part of the safety analysis for each reload core using NRC-approved methodology to systematically confirm the physics parameters for a reload core are bounded by the licensing Chapter 15 analyses. The models used to perform these calculations are based on the available operating history of the previous reload cycle to assure best estimate calculations, Determination of whether a nuclear-related physics parameter is within the bounding value assumed in the reference safety analy-sis must be made by performing explicit calculations of the parameter, or by comparison to values generated in previous relead core designs, Comparison to previously calculated physics parameters (to determine if the physics parame-ter is bounding) is only performed if the reload core being analyzed is , similar to previously analyzed reload cores. These comparisons can be per-
-formed to determine the bounding nature of a physics parameter because of the predictable behavior of most physics parameters a* i function of reactor power, rnoderator temperature, burnup, ard solube Nron 'oncentration. The calculation of control rod worths, reactivity coefficients, and kinetics parameters are described below.
3.1- Control Rod Worth Calculations The primary purpose of control rods is to provide adequate shutdown capability during normal plant cperation and accident conditions. Control rods are'also used to maintain criticality during rapid reactivity changes such as those that would occur during typical load follow maneuvers. They can also be used to offset reactivity changes produced from fuel depletion and changes in boron concentration, xenon concentration, and moderator temperature. However, control-rods are maintained at or near their all rod out (ARO) position during 3-1
nominal power operation and a,e normally only used to compensate for rapid reactivity changes. Control rod integral and differential rod worths are sensitive to local and global power distribution changes. Since the placement of fresh and depleted fuel assemblies produe.es unique power distributions, it is necessary to analyze control rod worths for each reload core. Rod worth related calcula-tions that are evaluated for each reload core are:
- Shutdown margin
- Trip reactivity
- Control rod insertion limits
- Maximum differential rod withdrawal at power Harimum differential rod withdrawal from subtritical ,
- Dropped rod worth
- Ejected rod worth Shutdown Margin Shutdown margin calculations are typically performed for each reload core at i
beginning of cycle (00C) and end of cycle (EOC) at various power levels including hot full power (HFP) and hot zero power (HZP) conditions. These calculations are typically performed in three dimensions, taking into account ,
-the power defect, stuck rod worth, allowance for rods being at their power dependent insertion limits, and rod worth uncertainty.
Trip Reactivity The minimum trip reactivity and the trip reactivity shape are evaluated for each reload core. Trip reactivity is defined as the amount of negative reac- , tivity inserted into the reactor core following a reactor trip. Allowances for the highest worth stuck rod and for the control banks at the rod insertion limits are taken into account. If the results from this calculation are not bounded by the trip reactivity assumed in the safety analysis, reanalyses of the affected accidents are performed with a new minimum trip reactivity. 3-2
The minimum normalized trip reactivity shape is also analyred for each reload core. This calculation is performed from HfP, and is structured to conserva-tively delay the amount of negative reactivity inserted into the reactor core versus rod position. The highest worth stuck rod is assumed stuck in itf fully withdrawn position after trip. This conservatism is achieved by allowing for a bottom peaked power distribution. Control Rod Insertion limits Control rod insertion limits serve several functions and are dependent upon the acceptable results of power peaking analyses, shutdown margin calcula-tions, ejected rod worth calculations, and inserted reactivity assumptions for safety analysesi Verification of the rod insertion ilmits from a peaking standpoint is performed in the operating limits and RPS setpoint analysis performed for each reload core design. The methodology used to perform this analysis is discussed in detail in Reference 3-3. Rod insertion limits also impact the available shutdown margin by influencing the magnitude of the rod insertion allowance. The rod insertion allowance is calculated at various burnups and includes allowances for top peaked power distributions. Rod insertion limits also impact the ejected rod worth and the amount of worth available for withdrawal for accidents sensitive to this parameter, Maximum Differential Rod Withdrawal from Power The maximum differential rod worth at power is calculated for each reload core at BOC and EOC. This calculation is performed to assure that inputs to the uncontrolled bank withdrawal at power accident are bounded. The maximum differential rod worth of any two control banks is calculated assuming normal overlap and adverse axial power distributions, while adhering to the power dependent rod insertion limits. Maximum Differential Rod Withdrawal from Suberitical The maximum differential rod worth from subcritical is calculated for each reload core at BOC and EOC. This calculation is performed to assure that inputs to the uncontrolled bank withdrawal from suberitical or low power 3-3
accideni are bounded. The calculation of this parameter assumes that control banks move in 100% overlap with the reactor at HZP. 1he impact of adverse axial-power distributions is also considered in the calculation of the maximum differential rod worth. Dropped Rod Worth The maximum allowed dropped rod worth is calculated at both BOC and EOC. Limiting combinations of dropped rods are evaluated to determine the maximum dropped rod worth. This value is compared against the reference analysis value to ensure that the safety analysis remains bounding. Dropped rod worths are calculated by evaluating the reactivity difference produced from a control rod or rods dropped from the HFp ARO condition. Ejected Jod Worth Ejected rod worths are calculated at BOC and EOC for both HFP and HZP condi-tions. Initial conditions for the ejected rod worth calculation are established by assuming that the control rods are at their rod insertion limit and by.1mposing a positively skewed power distribution. The rod worth calculation is performed by ejecting the control rod from the rod insertion limit to the ARO condition and calculating the reactivity difference. All possible rods are analyzed to determine the highest worth ejected rod. Conservatisms in the calculation of this rod worth and the resulting peaking factors produced from the rod ejection are retained by holding both the moderator and fuel temperature distributions constant. 3.2 Re_acti ity Coefficients and Kinetics Parameters The dynamic behavior of a reactor core during load follow maneuvers, tran-sients, and accident conditions can be described in terns of reactivity co-efficients. The magnitude and sign of these coefficients affect the reactor stability during transient and accident conditions. Reactivity coefficients are defined as the change in reactivity produced from a change in reactor power, moderator density, fuel temperature or boron concentration. The 3-4
l moderator density effects are often expressed in terms of moderator tempera-ture. Since these coefficients are a strong function of exposurt. they are calculated at several exposure statepoints during core life. ..ea . .ity coefficients are also influenced by changes in moderator temperature, reactor power, and soluble boron concentration. The statepoints at which reactivity coefficients are evaluated are chosen to ensure that the assumptions made in the specific accident analyses remain bounded.. For example, the moderator dilution accident at power is sensitive to the most positive moderator temperature coefficient and the steam line break accident is sensitive to the most negative (or least positive) isother- , mal temperature coefficient. The calculation of the moderator temperature - coef ficient, and fuel teraperature coef ficients and the statepoints at which these coefficients are evaluated are discussed below. The calculation of critical boron concentrations, boron worths and kinetics parameters follow. Moderator Temperature Coefficient The moderator temperature coefficient (MTC) is defined as the change in core reactivity resulting from a change in moderator temperature. Bounding coef- .ficients '(least and most negative) are calculated for each reload core. The following parameters are considered in the evaluation of the moderator tem-perature coefficient to ensure that conservative results are obtained.
- Soluble boron .
- Cycle exposure
- Control rods
- Moderator temperature The calculation of the HTC is typically performed using a three-dimensional core model. The moderator temperature coefficient '.s calculated by inducing a change in moderstor temperature (and, therefore, density) about the average temperature of-interest and dividing the resulting reactivity change by the change in moderator temperature.
3-5
l i Dop_pler Temperature Coefficient The Doppler (or fuel) temperature coef ficient (DTC) is defined as the change in tore reactivity resulting from a change in fuel temperature. The most and ! least negative DTCs are calculated for each reload core considering the core burnup and power level. The DTC is calculated by performing a set of two cases which vary the fuel temperature about a mean fuel temperature. The reactivity difference between the two fuel temperatures divided by the change in fuel temperature is the definition of the DTC. 01Cs are of ten quoted at various power levels by equating changes in reactor power to changes in mean fuel temperature. Kinetics parameters The dynamic behavior of the reactor core is determined to a large degree by the presence of delayed neutrons. Delayed neutron fractions and decay con-stants are calculated for six effec +ive delayed neutron groups. The total beta effective is the sum of the six group effective fractions and is, along with prompt neutron lifetime, calculated at BOC and EOC conditions. Critical _ Baron Concentrations and Baron Worths Critical and shutdown boron concentrations are calculated as a function of reactor power, exposure, temperature, and control rod positions as allowed by. the power dependent rod insertion limits. Differential boron worths are also calculated as a function of various combinations of the above variables. The results of these calculations are compared to inputs for several accident analyses, References 3-1 Nuclear Physics Methodology for Relond Design, DPC-NF-2010A, Duke power Company, June 1985. 3-2 Nuclear Design Methodology Using CASMO-3/ SIMULATE-3P, DPC-NE-1004, Duke Power r ompany. 3-6 u
3-3 Nuclear Design Methodology for Core Operating limits of Westinghouse Reactors, DPC-NE-20llP, Duke Power Company, January 1988. N mm 3-7
.. - . - - , . .-_ - -~.
i 4,0 R00 EJECTION ANALYSIS
- 4.1 Overvfew 4.1.1 Descriptica of Rod Ejection Accident 7he rod ejection accident is dest- bed in esAR Sectier. 15,4.8 (Reference 4-1).
The acc16ent is initiated by a failure of the control rod drive mechanism housing, which allows a control rod to be rapidly ejected from the reactor by the Reactor Coolant System pressure. If the reactivity worth of the ejected
.sntrol rod is large enough, the reactor will become prompt crittcal. The resulting power excursion will be limited by the fuel temperature feedback and the accident will be terminated when the Reactor Protection System trips the reactor on high neutron flux and the remai.'ing control rods fall into the core. The mechanical design and testing of the control rod drive mechanisms and housings make this event unlikely, If a control rod election should occur, the nuclear design ef the reactor core and limits on control rod insertion will limit any potential fuel damage to acceptable levels.
4.1.2 Accepttace Criteria The rod ejection accident is classified as an ANS Condition IV event. Three
- acceptance criteria are applicable as required per NUREG-0800 Section 15.4.8 .(Reference 4-2). The radially averaged fuel pellet enthalpy shall not exceed 280 cal /gm at any axial location. This criterion-ensures that a coolable core-geometry is maintained. Acceptable offsite dose consequences must be shown by being "well_within" the 10CFR100 dose limits of E5 rem whole-body and 300 rem to the thyroid. "Well within" is to be' interpreted as less th'an 25% of the above values. The radionuclide source term is determined by conservativaly predicting the number of fuel pins exceeding the ONB limit and the percentage of melted fuel. The peak Reactor Coolant System pressure must be within Service Limit C as defined by'the ASME Code (Reference 4-3), which is 3000
- psia (120% of the 2500 psia design pressure).
4-1
_ . ~ __. . .-_ _ _ _ _ . _ . _ . __ . ._ _ _ 1
'4.1.3 Analytical Approach-The complexity of the core'and system response to a rod ejection event re- -quires the application.of a- sequence of computer codes. The rapid core power excurston is simulated with a thrBe-dimensional transient neutronic and ,
thermal-hydraulic model _using the ARROTTA code (Reference 4-4). The resulting transient core power distribution results are then input to VIPRE-01 (Refer-ence 4-5) core thermal-hydraulic models. The VIPRE models calculate the peak fuel pellet enthalpy, tie allowable power peaking to avoid exceeding the DNBR limit, and the core coolant expansion. rate. The allowable power peaking is then used along with a post-ejected condition fuel _ pin census to determine the percent of pins in DNB. The coolant _ expansion rate is input to a RETRAN-02 (Reference 4-6) model of the Reactor Coolant System to determine the peak pressure resulting from the core power excursion.
-4.2 Siinulation Codes and Models 4.2.1 Nuclear Analysis The response of the reactor core to the rapid reactivity insertion from the control rod ejection is simulated with the ARROTTA code. -ARROTTA computes a three-dimensional power distribution (in rectangular coordinates) and reac-tivity or power level for both static and transient applications. The neutronics-solution in ARROTTA is based on the Analytic Nodalization Method as 1 developed-.for QUANDRY (Reference 4-7).
The neutronics method generates _an exact solution to the neutron diffusion equations if the shape of the transverse leakage function is assumed to be of - a'known_ quadratic form. In the limit of small node sizes, the equations revert to the same limit as the standard flux-centered finite difference , equations. ARROTTA uses a full two group representation of the diffusion equations and up to six delayed neutron groups. A complete description of the 1 theory and equations solved in ARROTTA can be found in Reference 4-4 l l 4-2
The ARR0TTA.model for the rod ejection analysis is based on a best estimate-model of Catawba 1 Cycle 2-that is adjusted as described in Section 4.3.1 to ' produce conservative results. .The assembly enrichments, burnable poison loading and assembly exposures for Catawba 1 Cycle 2 are shown in Figure 4-1. The neutronics model is based on the Westinghouse optimized fuel loaded in the reactor for that cycle. The ARROTTA model has one node per fuel assembly in the radia1 direction and twelve equal length fuel. nodes in the axial direction. In addition to the fuel, 'there are two rows of reflector in the radial direction on the outside of the core and one plane of reflector nodes on both the top and bottom of the core. The-reflector row next to the fuel consists of homogenized steel baffle and reactor coolant while the outer ow contains jue+ coolant. The axial reflector planes consist of homogenized reactor coolant, assembly structure, and some vessel structure, All-fuel and reflector cross sections and assembly discontinuity factors (ADFs) were taken from CASMO-3 (Reference 4-8) assembly lattice calculations. The two group conventional cross sec+. ions are processed by a series of auxil-iary programs and input to ARROTTA in the following form: I = R - ( A + B X + C X2) + (1-R) - (D + E.X + F X8) + (T, -%) + hM - g)f f where R=0 for no control rod,- R=1 for a control rod fully inserted in the node, A, B,_ C, D, E, F, dI/dT, and dI/dT are 7 determined from the CASMO-3 cross sections for both energy groups for all cross sections including Itr,
-Ia, vI and KI g. The dI/dT term is only used for the fast group cross f f sections. ~For a PWR, the X term is defined as the change in relative water 3
density from a reference value (density / reference density - 1). Also, there are microscopic coniributions to absorption and removal for soluble boron and to just absorption for xenon, iodine, samarium, and promethium. Because of , limitations in the auxiliary programs, the variations of all cross-sections-i. against X are the same either with or without a control' rod present (A/D, but E=B and F=C). 4-3
, - - ~ --
l l The cross sections are not functionalized again,t fuel exposure because each set of cross sections, called a composition, is only valid for a unique fuel exposure and enrichment combination. The auxiliary program that function-alizes the cross sections finds fuel nodes of similar exposure and identical enrichment 6'id assigns a single composition to those nodes at thei. avera a exposure, q ADFs are used to account for heterogeneities within assemblies for which the homogeneous fiux solution cannot account. ARROTTA allows a different ADF for each face of a node, but CASMO-3, because it is an infinite lattice program, , provides only a single radial ADF. These ADFs are input to ARROTTA as a singie rac ial value for each fuel node and reflector-fuel interface. Since ADFs are not strong functions of fuel exposure, they are input as single values for each assembly instead of a value for each composition. ADFs are not used in the axial direction. The ARR0TTA thermal-hydraulic model is comprised of a fluid dynamics model and a fuel pin heat transfer model. The fluid dynamics model is an inhomogeneous, non-equilibrium, two phase, closed channel model that uses separate energy equations for each phase and accounts for six possible flow regimes. The heat conduction model is based on spatially averaged, time dependent equations for the average pellet temperature. The thermal-hydraulic parameters calculated by ARROTTA are used to update the cross section model for the nuclear calculations; they are not used to determine fuel performance during the transient. The ARR0TTA code has been benchmarked against numerical steady-state and transient standard benchmark problems. The results of these benchmarks are documented in Reference 4-9 and show that ARROTTA agrees very well with the reference solutions. ARROTTA has also been benchmarked to a separate rod ejection transient simulation for a four-loop Westinghouse reactor (Reference 4-10). The benchmark case compares ARROTTA to HERMITE, a code which has received NRC review and approval for use in control rod ejection analyses. The ARROTTA model for this benchmark problem is very similar to the model used in this control rod ejection analysis. The results from this benchmark 4-4
problem also show excellent agreement. These benchmark problems clearly demonstrate that ARR0TTA is an acceptable code to use in analyzing the rod
-ejection accident.
ARROITA is used to calculate the core power level versus time during the rod
-ejection transient. Also, the radial, axial, and total peaking by assembly is calculated at each timt step during the transient. This information is used by VIPRE to determine the fuel temperature, enthalpy and the amount of fuel failure due to DNB, 4.2.2 Core Thermal-Hydraulic Analysis 4.2.2.1 VIPRE-01 Code Description 'The VIPRE-01 code is used for the rod ejection analysis thermal evaluations.
VIPRE-01 is a subchannel thermal-hydraulic computer code developed for EPRI by Battelle Pacific Northwest Laboratories (BPNL). The VIPRE-01 code has been reviewed by the NRC and was found to.be acceptable for refa encir in licens-ing applications (Reference 4-11). With the subchannel analysis approach, the nuclear fuel element is divided ( into a number of quasi one-dimensional channels that communicate laterally by ! diversion crossflow and turbulent mixing. However, VIPRE-01 is also capable of simulating single subchannel geometry. Given the geometry of the reactor
. core and coolant channel, and_the boundary conditions or forcing functions, .VIPRE-01 calculates core flow distributions, coolant conditions, fuel-rod temperatures and the departure from nucleate boliing ratio (DNBR) for steady-state and transient conditions. VIPRE-01 accepts all necessary bound-ary conditions that-originate either from a system transient simulation code -such as RETRAN, or a transient core neutronics simulation code such as
- - ARROTTA. Included is the. capability to impose different boundary conditions on.different segments of the core model. For example, different transient inlet temperatures, flow rates,_ heat flux transients, and_even different transient assembly and pin radial powers or axial flux shapes can be modeled.
l 4-5
4.2.2.2 Fuel Temperature and Enthalpy Calculation In order to show that the peak fuel enthalpy acceptance criteria described in Section 4.1.2 is met, al VIPRE model with fuel conduction is utilized to calculate the maximum hot spot fuel temperature and enthalpy during the transient. Given the
. VIPRE calculates the transient maximum hot spot average fuel temperature and the ,
maximum radial average fuel enthalpy. Detatis regarding the " s-VIPRE model and initial and boundary cenditions follow. Model Description i l l l l [ Axial- Power Distributions l l- During the transient the hot assembly axial power distributions change mainly due to the motions of the ejected control rod and due to the insertion of control rods as the reactor trips'. VIPRE is able to accept different axial power distributions during the transient. For each transient case, L m Gum b 4-6
M eul N 9M I4 . WAE N 1 NG tuu i W 4-7
Operating Gas Gap Conductivity Condition- { Btu /hr-ft 8-F.) Heat Transfer Correlations Sensitivity studies have been performea to justify the use of the heat trans-fer correlations. for_ the four major segments of the boilirig curve as shown below. For single phase forced convection - Forsaturatednucleateboilingregime-[
}
)
Fortransitionboilingregime-[ For film boiling regime - ' } } I The critical heat flux correlation used to define the peak of the boiling curve is the W-35 correlation. The minimum DNBR value for which transition boiling occurs is set to be 1,30. Flow Correlations For the rod ejection analysis, the subcooled void, the bulk void, and the twophasefrictionmultiplieraremodeledbyusingthe[. correlations, respectively, The jus _tification of using these models is based on the results of the sensi-tivity analysis of different void models to the transient fuel temperature calculation, I i l 4-8 l
I i l Other Thermal-Hydraulic ' Correlations l Pressure losses due to frictional drag are calcolated in VIPRE for axial flow. The friction factor for the pressure loss in the axial direction is determined frcm at empirical correlation as: f = A x Re 0 where Re is the Reynolds number _. The code evaluates both a turbulent and laminar set of coefficients and selects the maximum. The values selected for parametersAandBarebasedon[ s . .
-Turbulent flow: A= l B=
Laminar flow: A= 8= I ., . . The local hydraulic form loss coefficient is set as a constant to model the irrecoverable axial pressure loss as shown below. AP = KG2/2pg c where: K = spacer grid form loss coefficient G = mass flux, lbm/sec-ft' p = density, Ibm /ft' g = 32.174 lb-ft/sec t-lb c f Cpnservative Factors Flow area reduction - the subchannel flow area is reduced by 2% to account for variations in as-built subchannel flow area. Hot channel flow rate reduction - the hot assembly inlet flow is conserva-tively reduced by 5% from the nominal assembly flow. An appropriate engineering hot channel factor is applied to account for variations in the fabrication variables which affect the heat generation rate along the flow charinel. 4-9
a .
' Direct Coolant Heating The amount of heat generated in the coolant is 2.6% of the total power.
Fuel Enthalpy Calculation VIPRE-01 does not perform a fuel enthalpy calculation. Thus, the fuel enthalpy for a given fuel temperature during the transient is calculated separately f rom VIPRE based on the equation obtained f rom MATPRO (Reference 4-13). _, i FENTHL = FENTHL(T) - FENTHL(Tref) with FENTHL = gp j 0
. E' - + K,ex (-E g/RT)
Where: FENTHL = fuel enthalpy (J/kg) T = temperature (K) Y = oxygen to metal ratio = 2.0 R = 8.3143 (J/mol
- K) 0 = the Einstein temperature (K) _
= 535.285 for UO, K = 296.7 (J/kg - K) 3 K,= 2.42 x 10 8 (J/kg - K2)
K,= 8.745 x 10' (J/kg) ED = 1.577 x 10' (J/mol) FENTHL(T 7 ,f) = fuel enthalpy at any desired reference temperature . The above fuel enthalpy correlation is only valid for a fuel temperature greater than about 300K (80.3 F) (Peference 4-13). The reference temperature is 300K, 4-10
. -_ _ - . . - - . -. -. =
4.2.2;3 Coolant Expansion Rate Calculation-If the peak fuel enthalpy criterion is met, there is littir chance of fuel dispersal into the coolant. Therefore, the Reactor Coolant System expansion rate may be calculated using conventional heat transfer from the fuel and prompt heat generation;in the coolant. This rate must be calculated with the consideration cf the spatial power distribution befo-e and daring the transient since this rate, at any lecation in the reactor core, depends on the iritial smount of subcooling and the rate of change of the heat added into the coolant channels. A }VIPREmodel.isconstructedforthispurpose. Using the - VIPRE calculates the flow rate in each channel during the transien+. Using the VIPRE channel flow rates, the total coolant expansion rate can be calculated. This total coolant expansion rate is input to a RF~RAN plant transient model for simulating the resulting pressure-response. Model Description 4 5
='
m. 4-11
Axial Power Distributions , . , ~. Radial Power Distributions Fuel Conduction Model umB 4-12
He_at Transfer Correlations Heat transfer correlations used for the four majnr segments of the boiling curve are as shown below, For single phase forced convection I ' Forsaturatednucleateboilingregime-[ ] For transition boiling regime -[ ] Forfilmboilingregime-{. ] The critical heat flux correlation used to defina the peak of the boiling curve is the W-35 correlation. The minimum DNBR value for which transition boiling occurs is set to be 1.30. Flow Correlations For the coola:.t volume expansion calculations, the subcooled void, the bulk void, and the two phase friction multiplier are modeled by using the - correlations, respectively. Other Thermal-Hydraulic Correlations Refer to Section 4.2.2.2. _ Calculation of the Reactor Coolant Expansion Rate From VIPRE Flow Rates From the model results, the inlet and axit mass flow rates and densities for each channel can be obtained. The instantaneous volume expar-sion rate at time t for each channel, Q$ (ft'/sec), is first calculated as shown below. inlet Qi (f t'/sec) = Mi, pi, exit exit _ Mi, pi, inlet where: i = channel index M = mass flow rate, I b.n/sec p = density, lbm/ft 2 4-13
Then the instantaneous core volume expansion rate at time t is: I]Q'g Q (it'/sec) = i=1 The above calculations are repeated for dif ferent times to obtain Q(t) during the transient. 4.2.2.4 DNBR Evaluation To' determine the offsite dose consequences, an analysis is performed using the VlpRE code to determine the percentage of the core experiencing DNB. Those fuel pins which exceed the DNBR limit are assumed to fail.
]TheCHF
-correlation used is the BWCMV correlation (Reference 4-14) and the DNBR limit is 1.331 (1.331 = 1.10 x 1.21 where 1.210 is the correlation design limit and the 1.10 factor adds 10% margin). Analyses have shown that to prevent DNB,
{. .
)Afuelpincensusisthenperformedtodeterminethenumberoffuel pins.in the core experiencing DNB. .g l- 4-14 l
1 I Last, the pin power for every pin is compared to the appropriate MARP value. If the pin power is higher than the MARP, then that pin is in DNB. - Model Description A is used fcr the rod ejection
- }modelofa 3.. '
DNBR calculations. The. u sed for the DNBR calculation is identical to the one used in Section 4.2.2.? (Figure 4-2), l Axial Power Distributions , Radial Power Distributions 4-15
Fuel Conduction Model Heat Transfer Correlations For the DNBR calculations, only the single-phase forced convection and nucle-ate boiling heat transfer modes are applicable. The[ 3correla-tion is used for the single phase forced convection mode. The[,
]correlationisusedforthenucleateboilingregime. Justifica-tion for using these correlations is based on [
] The critical heat flux ;
correlation used to define the peak of the boiling curve is the W-35 correlation. The minimum DNBR value for which transition boiling occurs is set to 1.30. Flow Correlations The[ 3correlationforthetwophasefrictionmultiplier. Justification for using these correlations is based on [ b Other Thermal-Hydraulic Correlations ; l i Refer to section 4.2.2.2. ) l 4-16 l
4.2.3 System Thermal-Hydraulic Analysis The Reactor Coolant System resp'nse to a rod ejection accident is primarily a rapid'pressurl7ation due to the increase in heat transfer associated with the power excursion. - lhe VIPRE analysis of the coolant expansion rate described in Section 4.2.2,3 produres an expansion rate which conservatively models
~
}TheVIPREresult is input to a RETRAN-02 model of McGuire/ Catawba as The RF1RAN-02 model is the ase model described in detail in Reference 4-15.
-4.3 ARROTIA Analysi.s 4.3.1 Initial Conditions The control rod ejection transient is analyzed at four statepoiats for Catawba Unit 1, Cycle 2: beginning-of-cycle (80C) at hot zero power (HZp) and hot full power (HFP) and end--of-cycle (E0C) at HZP and ilFP, Because of the modifications to the ARROTTA model that are described below, analysis of this core is expected to bound any expected future reload cycle. The ejected control rod is located at core location D-12. Figure 4-4 shows this location in the reactor. The control rod in location D-12 is part of Control Bank D (hereaf ter referred to as Bank D). At the HZP rod insertion limit D is the only bank-fully inserted. At the HFP insertion limit, it is the only bank in the core. The central control rod (location H-8, also a member of Gank D) is not chosen as the ejected rod because sensitivity studies showed that c higher f gwould be achieved by ejecting a given worth from D-12 than from H-8 due to
-the asymmetric power distributions produced when 0-12 is ejectad. Since the higher Fg is more conservative, D-12 is chosen as the ejected rod.
For the HZP statepoints, the reactor is initially critical at a very low power
-. level with control rods at the insertion limit: Bank D at 0 steps withdrawn (swd), Bank C at 47 swd, and Bank B at 162 swd. The rod is fully withdrawn at approximately_226 swd and at 0 swd the control rod tip is approximately 2.65 inches above the bottom of the core. No allowance is mole for a bank of rods being mispositioned lower than indicated (higher worth) because the ejected 4-17
rod is initially f ully . inserted. If either Bank C or B were mispositioned, the worst effect would be to increase the ejected rod worth, However, the ejected rod worth is alreaay assumed to be conservatively high, so this ef f ect is accounted for. Tne control rod is ejected in 0.1 seconds at constant velocity for the HZP cases. This is significantly faster than physically possible, even when friction is ignored in the ejection time calculation. Sensitivity studies show that the peak' power level attained during the transient is slightly' higher (more conservative) for a faster ejection time. Thus, the control rod ejection results are conservative with respect to control rod ejection time. For the HFP statepoints, the reactor is initially at 102% of rated power with Bank D at 149 swd. This is 12 steps beyond the insertion limit to make allowance for the bank being mispositioned. The control rod at D-12 is ejected in 0.058 seconds at constant acceleration. This acceleration is consistent with the ejection time of 0.1 seconds used in the the HZP cases. me m 4-18
1
= w The moderator temperature coef ficient (MTC) is adjusttd to the values 3bown in s Table-4-1. Although the MTC has very little ef fect on the the transier:t' results, it was adjusted to be Greater than_the tachnical specification Iimits for the BCC statepoints. For the EOC statepoints, the MTC was adjusted to be greater (less negative) than any expected for future reload cycles. The MTC wasadjustedintheARROTTAmodelby(
The Doppier (or fuel) temperature coefficient (UTC) is important to this transient because the negative reactivity from the increased fuel temperature is the only effect that limits the power excursion and starts to shut down the reactor. The DTC is adjusted to the values shown in Table 4-1. These values are greater (less negative) than any expected for future reload cycles. The DTC is-adjusted by . The effective delayed neutron fraction _(B) and the ejected rod worth both determine the transient response of_the reactor, P'e peak power level at-tained-during- the transient will: increase for smaller values of S and-1,arger values of the ejected rod worth. The ejected rod worth is adjusted by ,, W 4-19
S is input to the model by six delayed groups for each composition. Since 6 is dependent on enrichment and burnup, all the cnmpositions are different from each other. Bisadjustedby( P' The total effect of all of the changes to the Catawba Unit 1 Cycle 2 model is to create an ARROTTA model that will bound all expected reload cycles for both McGuire and Catawba Nuclear Stations. The varlons limiting parameters are listed in Table 4-1. 4.3.2 Boundary Conditions The fuel and core thermal-hydraulic boundary conditions are listea on Table 4-2 The thermal-hyoraulic description of the foal used in ARROTTA represents B&W Mark-BW fuel. The rod ejection transient is very nearly adiabatic through the time that the peak power level is limited by the fuel temperature feedback. Since the-Mark-BW fuel contains a higher mass of fuel than the Westinghouse OFA fuel, it will heat up more slowly during adiabatic events. The slow heatup decreases the fuel temperature- feedback, resulting in a higher, more conservative, transient core power response. The reactor trip signal is generated when the third highest excore detector reaches either 37% for_the HZP cases or 118% for the HFp cases. 'This modeling is based on a single failure of the highest detector and a two-out-of-the-remaining-three trip coincidence logic. 'The excore signals are synthesized from the power densities of several assemblies that are near the excore locations. The remaining control rods fall into t.he reactor starting, at 0.5 seconds after the trip signal is generated. During the'rea: tor trip, the ejected rod and the remaining rod with the highest worth ar< assumed not to fall into the raactor. To conservatively model the reactor trip, not all of the control rod banks are allowed to drop, and some of the banks that are dropped-have their worth reduced by cross section-adjustments. The net shutdown margin in all cases is less than 250 .pcm. Also, the' negative reactivity inserted due to the reactor trip is not 4-20
allowed to exceed the conservative trip reactivity curve that is shown on Figure 4-7. The integral worth of the falling control rods is computed for several different axial positions Of therodsattheinitialconditions.[ e 4.3.3 Results Core power versus time, as calculated by ARR0TTA, is shown in Figures 4-8, 4-11, 4-15, and 4-18 for the four cases. Table 4-3 summarizes the results of the four cases. The ARROTTA analysis used time steps of 0.001 seconds through the time of peak power. After the peak power occurred, the time step size was relaxed to 0.010 seconds. Sensitivity studies showed that for the initial power excursion the peak core power level is reduced for time steps shorter than 0.001 seconds. Thus, the time step selection of 0.001 seconds through the time of the peak power level is conservative. After the peak, sensitivity studies showed very little change in the results for time steps up to 0.010 seconds. For the HFP statepoints, the core power increas s rapidly as the control rod is ejected (Figares 4-8 and 4-15). The power continues to increase until the Doppler feedback, caused by the increasing fuel temperature, becomes large enough to turri the excursion around. The power level then continues to decrease.as the fuel-temperiture approaches an equilibrium value. Due to the . rapid initial power increase, the reactor trip on high flux oreurs very early. Rod insertion completes shutdown of the reactor. Insertion-begins after the peak power due to the trip delay. Since the trip reactivity for each transient is normalized to~the conservative trip reactivity curve of Figure 4-7,_ rod motion has a minimal effect until the rods approach the bottom of the Core. The HZP statepoints differ fron the HFP statepoints in that there is no initial thermal-hydraulic feedback and the reactor becomes prompt critical. The initial power increase continues long after the control rod is ejected (Figures' 4-11 and 4-18). Since the reactor is prompt critical, it quickly reaches a high power level before the fuel heats up enough for the Doppler 4-21
feedback to turn the power excursion around. The power level then decreases almost as fast as it increased, until near-equilibrium is reached. The reactor is thenLshut down by control rod insertion resulting from the high flux trip. The ARROTTA initial radial power distribution for the BOC HFP case is shown on Figure 4-9. The power distribution at the time of the pea, power, which is concurrent with the highest, radial and nodal peaks, is shown in Figure 4-10. For the EOC HFP statepoint, these power distributions are shos - in Figures 4-16 and 4-17. The initial power distribution for BOC HZP statepoint is shown in-Figure 4-12. The power distribution at the time of the highest radial and nodal peaks is shown on Figure 4-13 and the power distribution at the-time of the highest pcwer is shown on Figure 4-14 For the E0C HZP statepoint, these power distributions are shown on Figures 4-19, 4-20, and 4-21. The core power level versus time is a key input from the neutronics calcula-tion to the thermal-hydraulic evaluation discussed in the next section. In addition to total core power versus time, the as discussed in Section 4.2.2.4. The . to evaluate the number of pin failures due to DNB as described in Section 4.4.4. 4.4- VIPRE Analysis 4.4.1 initial Conditions During the-rod ejection transient, the reactor core coolant pressure increases due to the coolant expansion as a result of the reactor power excursion, i~ 4-22
The core inlet flow rate for the HFP case, 339,972 gpm, is derived by reducing the assumed technical specification minimum measured flow, 382,000 gpm, by 9% for assumed bypass flow and by 2.2% for measurement uncertainty. The two pump core inlet flow rate for HZP is 46% of the 4 pump HFP flow. The initial core inlet flowrates are 3",972 gpm and 156,387 gpm for HFP and HZP respectively. . B 4 The initial core coolant inlet temperatures include an allowance of +4 F for control deadband and measurement error. Core inlet troperatures of 561.4 F and 561.0 F are used throughout the transient analyses for HFP and HZP, respectively. The initial core power for the HZP and HFP cases are 0.0% and 102.0% of 3411 MWt. However, for the.HZP case an assumed 2% of 3411 MWt of decay heat is added into the average power generated by ARROTTA. The transient core average power for the four operating conditions are shown in Figures 4-8, 11, 15, and 18. 4.4.2 Fuel Temperature and Enthalpy - The fuel temperatures and enthalpies are calculated for the four transient cases. The maximum fuel temperature and enthalpy during the transient are shown below. Maximum Maximum Maximum Maximum Centerline Fuel Average Clad Surface Fuel Average Fuel Temp. Temp. Temp. Enthalpy Case (F) (F) (F) (cal /am) HZP, BOL 4090 3220 1410 133 HFP, BOL 3890 2812 1101 113 HZP, EOL 3190 2872 1172 116 HFP, EOL 3066 1956 940 75 4-23
The'above results show that during the transient the maximum centerline fuel temperature is well below the fuel melting temperature of 4700 F (Reference 4-1), and that the maximum fuel average enthalpy is well below the acceptance criterion _ radially averaged fuel enthalpy of.280 cal /gm. Since the fuel pellet does no. melt-during the accident, the activity due to the fuel pellet will not contribute to the dose calculation results. 4.4.3 Coolant Expansion Rate The Bf HFP rod ejection transient results in the highest coolant expansion rate. Figure 4-31 shows the instantaneous core coolant expansion -ate (ft /sec) as a function of transient time. 2 The initial expansion rate corresponds to the full power initial condition and the resulting decrease _in coolant density due to sensible heating in the core. The result shows that a peakexpansionrateofi 4.4.4 DNBR'and Fuel Pin Census I The DNBR calculations are performed for the four operating conditions. The DNBR results are expressed as a family of curves of maximum allowed radial power (MARP) versus assembly axial peak and location (Figures 4-22 through 4-25). When the radial power peak of the fuel pin exceeds-the MARP during the transient, DNB is assumed to occur and the cladding fails. The fuel pin census is performed to determine the number of. failed fuel pins during the accident. Results are shown in Figures 4-26 through 4-29 and are summarized below. 1 Operating % of Fuel Pins Conditions Experiencina DNB HZP, BOL. 10.7 HFP, BOL 36.9 HZP, EOL 19.6 l HFP, EOL 14.4 l l 4-24
The above results show that the HFP, BOC case has the largest number of nins experiencing DNB. The of f site <iose consequences are analyzed based on 50% of the fuel pins experiencing DNB to conservatively bound the above results. 4.5 R_ETRAN Analysis 4.5.1 Initial Conditions The RETRAN model pressure response to the rod ejection transient is primarily a function of the coolant expansion rate and the pressurizer code safety valve relief capacity, which are input as boundary conditions as discussed in the following section. Most parameters such as initial primary temperature have little impact on the pressure response due to the very short duration of the simulation (3.5 seconds) which results in minimizing tenperature transport effects. Sensitivity studies were performed which demonstrated that thermal effects were not significant. Therefore, the transient was evaluated with nominal hot full power initial conditions with the exception of pressurizer pressure and level. These two parameters clearly impact the system pressure response to the coolant expansion. Pressurizer Pressure The peak pressure response is cor.servatively bounded by using a m;ximum error- ; edjusted value of 2295 psig. This is the nominal hot full power pressure of 2235 psig with a 60 psi uncertainty allowance for elevated pressurizer pressure Pressurizer level A high initial pressurizer level decreases the volume of the steam bubble -thereby increasing the compressibility effect. The-limiting hot full power programmed pressurizer level is 61.5%. The initial condition uncertainty allowance for reduced level is 9%. The initial level is therefore 70.5%. 4-25
- -- -_ .- ~ _ . .
4.5.2 Boundary-Conditions Primary system boundary conditions which significantly effect the pressure
~
response include the coolant volume expansion rate, reactor power, and pressurizer safety valve modeling. These boundary conditions are discussed separately below. In order to conservatively bound the pressure response, the pressurizer PORVs and spray are defeated. Full primary system flow is maintained-to maximize the reactor vessel pressure drop and hence maximize the lower plenum pressure. The maximum system pressure occurs at the bottom of the lower plenum. . Secondary side boundary conditions were determined to have minimal impact on the transient due to the short duration of the simulation. Nevertheless, conservative assumptions were made to conservatively bound the pressure response. The turbine is assumed to trip immediately on reactor trip. Main , feedwater isolation is assumed to be initiated at time 0.0, and the isolation valves are ramped closed over a bounding 2.5 second interval. The condenser is assumed not to be available, and the steam generator PORVs are assumed to be inoperable. The steam generator safety valves are available for secondary steam relief, however, the simv~e' ion is terminated before these valves are challenged. Reactor Coolant Volume Expansion Rate m Mut - 4-26
w. Reactor' power
~
1
~~
Pressurizer Safety Valve Modeling
]
l 4 The pressurizer code safety valves function as overpressure mitigation equipment. The nominal lift setpoint is increased by 3% to account for calibration-allowance. The valves are assumed to open linearly until they are fully open at a pressure 3% above_the adjusted 'ift setpoint. The valve modeling then includes a hysteresis.effect that keeps the valves fully open until the_ pressure decreases to 5% below the adjusted lift setpoint. 4.5.3 Results The Reactor Coolant System pressure response to the rod ejection is shown in Figure 4-31. The pressure plotted represents the pressure at the bottom of-the reactor vessel lower plenum where the highest system pressure occurs. Figure 4-31 shows that a peak system pressure of 2728 psig is reached in 1.9 seconds. The-peak pressure is within the acceptance criterion of 3000 psig
= discussed-in Section 4.1.2, 4.6- Dose Consequences A conservative evaluation of-the rod ejection a'ccident is performed to determine the resulting radiological consequences. Methods used to perform this evaluation are identical to those used in the present licensing evaluation for Catawba Nuclear Station. No fuel melting occurs for either Catawba or McGuire Nuclear Stations. The value for the number of pins assumed toLenter DNB is conservatively selected to be 50% to bound the results given in Section 4.4.4. Dose results for McGuire and Catawba Nuclear Stations are as follows:
4-27
Exclusion Area Boundary Acceptance McGuire Catawba Criterion Whole Body 0.598 0.480 6.25 Thyroid 50.29 30.55 75 Low Population Zone Acceptance McGuire Catawba Criterion Whole BnJy 0.080 0.057 6.25 Thyroio 9.283 3.165 75 These results show that the offsite dose consequences from a conservative rod ejection analysis are well within the dose limits stated in 10CFR100. 4.7 Cycle Specific Evaluation Due to the conservative assumptions and modeling used in the ARROTTA model, it is anticipated that for reload cores, no new ARROTTA cases will be necessary. LThe determination as to whether the existing ARROTTA cases remain bounding
.will be made_by performing a cycle-specific reload check of the key physics input parameters listed in Table 4-4, These parameters will be calculated using standard steady-state neutronics codes approved by the NRC for reload
-design. If the key parameters remain bounded then rio new ARROTTA analyses are necessary; otherwise, an evaluation, reanalysis, or redesign of the reload core will be performed. A DNB pin census will be performed for the reload. cycle, as described in Section 4.3.3. The. ejected rod worth shall be calculated with the fuel and moderator temperatures frozen in the pre-ejected condition or uniform .throughout the core (either method will generate conservative results). Also, the xenon distribution will be skewed to force a top peaked power dist-ibution to make the ejected rod worth higher (for tb cases) and to make the DNB pin census more conservative. The power dis . ion with the ejected rod out 4-28
will be used for the DNB pin census. The calculated percent fuel failure dut ' to DNB will be compared for each cycle to the fuel failure limit assumed in the dose calculation. If the cycle specific value is less than the limit, ' then the existing ssfety analysis is still valid. Otherwise, an evaluation, a new dose calculation, reanalysis, or new reload design will be performed as appropriate.
.RgqLegej 4-1 McGuire Nuclear Station, Final $afety Analysis Report, 1988 Update.
42 USNRC Standard Review Plan, Section 15.4.8, NUREG-0800, Rev. 2, July 1981. 4-3 ASME Boiler and Pressure Vessel Code, Section Ill, " Nuclear Power Plant Components", ASME. 4-4 ARROTTA: Advanced Rapid Reactor Operational Transient Analysis Computer Code, Computer Code Documentation Package, EPRI. 4-5 VIPRE-01: A Thermal-Hydraulic Code for R. actor Cores; EPRI NP-2511-CCMA Revision 2, EPRI, April, 1987. 4-6 RETRAN-02: A Program fo. Transient Thermat-Rydr.ulic Acalysis of Complex Fluid Flow Systems EPRI NP-1850-CCMA devision 4. EPRI, " * - 1988. 4-) K, S, Smith, "An Analytic Nodal Method for Solving tbc ,?-Group, Multidimensional, Static and Transient Neutron Diffutton Equations," Nuc. Eng. Thesis, Dept. of Nuc, Eng. , MIT, Cambridge, MA. , (February, 1979). 4-8. Nuclear Design Methodology Using CASMD-3/ SIMULATE-3P, DPC-NE 104, Duke Power Company. 4-29
l l 4-9 "ARR0i1A Validation and Verification - Standard Benchmarts Set" EPR1 Research Project 1936-6, July 1989 Prepared by $. Levy, Inc, 4-10 ARROTTA-HERMITE Code Comparison, EPRI NP-6'14, December 1989. 4-11 Letter from C. E. Rossi (NRC) to J. A. Blaisdell (UGRA), " Acceptance for Referencing of Licensing Topical Report, V1 PRE-01: A Thermal-Hydraulic Analysis Code for Reactor Cores," EPRI-NP-2511-CCM, Vol. 1-5, May 1986. 4-11: Y. H. Hsii, et al., TACO 2: fuel Pin Performance Analysis, Revision 1, BAW-10141PA, June 1983. 4-13 Donald L. Hagrmann, et al., MATPRO-Version 11 (Revision 2): A Handbook of Materials Properties for Use in the Analysis of light Water Reactor fuel Rod Behavior, NUREG/CR-0479, August 1981. 4-14 D. A. Farnsworth and G. A. Meyer, BWCMV Correlation of Critical Heat flux in Mixing Vane Grid Fuel Assemblies, BAW-10159P, Babcock and Wilcox, May 1986. 4-15 Thermal-Hydraulic Transient Analysis Methodology, DPC-NE-3000, Revision 0, Duke Power Company, July 1987. 4-16 Nuclear Physics Methodology for Reload Design, DPC-NF-2010A, Duke Power Company, June 1985. l
}
4-30 l.
l l Table 4-1 Rod Ejection Transient Kinetics Param<eters Parameier BOC HIP 90C HFP EOC H7P LOC HFP Ejected Rod Worth (pcm) 76? 201 907 200 MTC (ptm/F) 7.1 0.5 -9.3 DiC (pcm/F) -9.9
-0.9 -0.9 -1.2 -1.2 Delayed Neutron Fraction 0 00551 0.0055 0.004 0.004 r
Table 4 c Rod Ejection Transient Initial Conditions Parameter BOC HZP BOC HTP EOC HZP EOC HFP Initial Power (MWt) 3411E-9 3479.22 3411E-9 3479.22 y Initial Power (%) 1.00E-7 102.0 1.00E-7 102.0 C Core riow (gpm) 156387.3 339972.4 156387.3 339972.4 Inlet Temperature (F) 561.4 561.0 561.4 561.0 Reactor Pressure (psia) 2305.0 2305.0 2305.0 2305.0 Fission Power Fra: tion in Coolant 0.026 0.026 0.026 0.026 Table 4-3 Rod Ejection ARR0TTA Results i BOC BOC EOC EOC Parameter HIP HFP HZP HFP Time of peak power, sec. 0.339 0.097 0.172 0.097 Ps k power level, % of full power 1d76 142 5555 167 Peat nodal powe- relative 12.79 3.17 18.11 3.76 to core average Peak assembly power relative 6.16 1.92 7.90 2.18 to core average Time that trip setpoint 0.296 0.064 0.156 0.061 reached, sec. Time of the beginning of 0.796 0.564 0.656 0.561 the tripped rod motion 4-31
T a b'; p 4-4 Rod Ejection Reload Chect. list i BOC DOC EOC EOC Parameter HFP HZP HFP HZP Ejected Rod Worth (pcm) less than 200 720 200 900 6 greater than 0.0055 .00551 .004 .004 DTC (pcmi'F) less than -0.9 -0.9 -1.1 -1.1 MTC (pcm/*F) less than 0.0 +7.0 -10.0 -10.0 DNB Census less than 50% 50% 50*; 50's F g less than 4.12 16.62 4.88 23.55 i I l { 1-l l 1 4-32 1 1 -- -- . . - - - - . - . - - - . _ . . - _ . . .
I Figure 4-I Catawba Unit 1. Cycle 2 Assembly Enrichments and Fuel Exposures I 1 H 0 F , E D C B A 1.6 00 -J . '
- 06 31-20 3 l*05 24*15 31*06 31 15 34/00 6 12 43- ! 72 -15.39 10 84 14'91 If d4 14 01 3 00 21 79 2459 27 73 23 60 2502 23 35 25 04 10.74 _
3 : 20 3l=00 24a16 31*00 2 4* Ib 3 "00 3 2/00 9 35 39 11.98 14 50 8 86 15 26 13 00 0 00 2765 24 67 25 0B 2199_ 26 23 25 04 10 27 2 +1" I b 3 2/06 24"I2 3 2/06 31*Ib 3 4/00 10 14 59. 0 00 15 53 0.00 14 10 0 00 26 25 14 31 27 12 14 17 - 26 16 , 53 2 4'00 32/I2 24ai6 3 2/04 24 12 11 16.24 0.00 14 44 0.00 15 74 27 72 13 63 25 91 12 55 20 4S 2.4= 12 31*00 3.2/00 12 15 76' 9 47 0 00
-26.63 22.14 .10 00 3.1/00 2 4*12 Enrichment /" of bps 13 0.00 15.76 SOC Exposure (L JD/t1TU) ;
12.18 20.29 EOC Exposure
' Note. 2 4"12 means a 12 BP cluster was Dulled from the 2 4% enriched assembly at the end of the last cycle.
3.2/12 means the 3 25 enriched assembly currently has a 12 SP ClU$ter. P 4-33 y- -+ r.,J ---, ------.m- . ,,--4.r-.4, ---.2.,----.*,uy.
+,-ys,y muw.-.-.-g. .,w.,. ,,--,w-,
---y ,y-.* y,rm---w----y =, r---,,.-,, . wre,
Figure 4-2
,VIPRE Model for Fuel Temperature and DNBR Calculations I
- e ,
4-34
d Figure 4-3 , VIPRE Model for R5 actor Coolant Expansion Rate Calculation
+
b yy rod number xx channel number 4-35 l l
l l I l rigure 4-4 ! PSAR Section 15.4.8 - Rod Ejection Accident l Control Rod Locations p r J N M L
- e- .
r [ .. e , i, l 1 l 7 D C S SA j D CA en L; 1 4 SA D SE D SA I 5 L: LD i l 6 0 C A C B 7 SD SD 8 C SE A D A SC C 9 SD SB 10 S C A C D ! l 11 SD SC 12 SA D SE D SA 13 SC SD SB SD 14 SA B C D SA 15 Control NumDer Shut d>wn Number Bank Of Rods Bav Of Roos A 4 SA S 0 0 SS B C 8 SC 4 0 5 SD 4 Totai 25 SE 4 Total 28 i mem mm E;ecteo Rea Location 4-36
~.
Figure 4-5 ' FSAR Section 15.4.8 - Control Rod Ejection SOC, HFF, ARO Power Distributions i BeforeandAfter[ ! 3 H 3 F E D C B A 0 00 1. t 4 t 12 1.15 0 95 1.13 1 12 1 04 0 065 0 95 0 96 1 01 0 84 1 01 0 90 09: 0 67 1 03 1.06 1 09 0 93 1 07 1 04 1.19 1.14 1 ;I i 15 0 96 1 17 0 96 1 11 0 99 9 0 95 0 94 1 01 0 as 1.00 0 88 1 00 0 00 1 03 1 03 1 09 0 97 I lo 0 96 1 10 1.16 1.12 1 lb 0 90 1 13 0 95 1.23 1 07 0 06 to 0 96 1.01 0.93 1.26 0 97 1 26 1 04 0 02 1.06 1 09 1.03 1 36 1.09 1.34 1,17 1,16 1.14 0 96 1.23 0 94 1.13 0 95 1 to 0 40 1,00 0.09 1.26 1.32 1.25 ti 1.25 1.18 0 40 1.08 0 97 1.36 1.39 1 40 1 40 1.37 0 71 0 95 1 17 0.95 1.13 0.91 1.10 0.90 12 0.84 1.00 0 97 1.25 1.35 1.32 0 98 0.93 1.10 1 09 1.40 1.50 1 40 1.35
~ .141 0 96 1.23 0.95 1 10 1 09 0 30 13 1.01 0 BB 1.26 1.25 1.32 1.13 1.40 1 07 0 96' 1.34 1 40 1.40 1 41 0.76 1.1 ) I ii 1.06 1.10 0.89 0.30 14 0 90 1 01 1.04 1.18 0.98 0 40 1.05 1.10 1.17 1.37 1.35 0.76 1.05 0 99 0 68 0 40 AssemDly Power Before Changes 15 0 91 0 88 0 82 0 40 Assembly Power Af ter Changes 1.19 1 16 1.15 0.71 Peak Pin Af ter Changes 1
4-37
.s
- - .. e .- - - - . . - , - , , . - . - - - . . - . . ~ . .._,,s,, . _ _ .
Figure 4-6 FSAR Section 15.4.8 - Control Rod Ejection EOC, HFP, ARO Power Distributions BeforeandAfter[ m l
.i H G F E D C B A 0 76 1 00 0 90 1 03 0 90 1 05 1 04 i O!
6 0 62 0 03 0 84 0 90 0.79 0 93 0 91 0 57 0 63 0 00 0 00 0 93 0 94 0 97 0 95 1 09 1 00 0 97 1 03 0 92 1 10 0 94 1 06 0 90 9 0 63 0 02 0 90 0 84 1 01 0 06 0 95 0 06 0 07 0 07 0 99 0 90 1.07 0 92 1 05 1 DS 0 96 1 03 0 95 1 27 0 99 1 26 1 05 0 93 10 0 64 0 91 0 90 1 29 1.00 1 29 1 02 O S6 0 80 0 99 0 98 1 33 1 00 1 33 1 12 1'3 1.03 0 92 1.27 1 01 1 30 1 00 1 14 0 50 ii 0.90 0 84 1 29 1.30 1.36 1.30 1 22 0 50 0 93 0 91 1.33 1 40 1.40 1 40 1 37 0 02 t 0 90 1.10 0.99 1.30 0.99 1 12 0 96 i; O 79 1.01 1.00 1.36 1 38 1.36 1.06 0 84 1.07 1 08 1.40 1.50 1 40 1.37 1.05 0,94 1.26 1 00 i.;2 1,10 0 47 13 0 93 0 86 1 28 1 30 1.36 1.19 r ' '2 0 97 0 92 1 33 1.40 1 40 1.40 0 o9 1.0 1 1 06 1.05 1.I4 0.96 0 47 ja #o' 0 95 1.02 1 21 1.06 0 52 0.95 1.06 1.12 1.37 1.37 0.89 1.01 0.96 0.92 0 49 AssemDly Power Before Changes 15 0.88 0.06 0 85 0.50 Assembly Power Af ter Changes 1.09 1.00 1.13 0.82 Peak Pin Af ter Changes 4-38
\
Figure 4-7 FSAR Section 15.4.8 - Control Rod Ejection Trip Reactivity Curve 1.0 . 0.9 - 41 0.8 .- t O ' U 0.7 . O o - 0.6 '
- ~ -
a ' 5- - 0.5 = a - u < m - O -
- a. 0,4 -
O - a 0.3 - m -
~ - @ e CC )
0.2 . 0,1 , 7 0.0 . . . 0 10 20 30 40 50 60 70 80 90 100
% Control Rod Inserted 4-39
i rigure 4 8 FSAR Sectiert 15. 4. 6 - Contr:1 Tu>d Eject. ion BOC HFP Core Power vs. Time 160 140- A. t 120 - - . re
- e. -
u y 100 O a G- . 9 - 8 80 M 60 , N e 40 s 20 - W d W W W T W W W g W W W W W W W W W W W W V i 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 Time (Seconds) 4-40 l
Figure 4-9 FSAR Section 15.48 - Centrol Red Ejectton BOC HFP Assomely Power Distribution at 0.0 Seconds A P N M L K J H 0 T [ 0 C 0 A 1 0 e0 0 85 0 92 0 96 0 92 0 85 0 40 2 0 38 0 95 1.18 1.07 f.05 i.03 105 1 07 1 18 0 95 0 38 3 0.38 1 05 1.22 1 22 1.28 0 91 1.05 0 91 1.28 1 22 1 22 1 05 0 38 4 0 95 1 22 1 06 19 0 99 1.12 0 68 1.12 0 99 1 19 I !9 1.22 0 95 5 0 40 1.18 1 22 1.19 1.30 1.29 0.92 1,04 0 92 1 29 1 30 1 19 f.22 1.19 0 40 6 0 85 1 07 1 28 0 99 1.29 0.96 1.04 0 97 1 03 0 96 29 0 99 1.28 1.07 0 85 7 0 92 1 05 0 92 1.12 0 92 1.03 0 93 0 91 0 93 1.04 0 92 1 12 0 91 1 05 0 92 8 0.96 1.03 1.06 0.88 1.04 0 97 0.92 0.51 0 92 0.97 1 04 0 88 1 06 1.03 0 96 9 0.92 1.05 0.91 1.12 0.92 1 04 0.93 0.92 0.93 1.03 0 92 1.12 0.92 1.05 0.92 to 0 85 1.07 1.28 0.99 1.29 0.96 1.03 0.97 1.04 0.96 1.29 0 99 1.28 1.07 0.85 r 1 11 0 40 1.19 1.22 1.19 1.30 1.29 0.92 1.04 0.92 1.29 1.30 1.19 1.22 1.18 0.40 L J .. 12 0.95 1.22 1.08 1.19 0.99 1.12 0.88 1.12 0.99 1 19 1.08 1.22 0 95 13 0.38 1 05 1.22 1.22 1.28 0.91 1.06 0.91 1.28 1.22 1.22 1.05 0.38 14 0.38 0.95 1.18 1.07 1.05 1.03 1.05 1.07 1. i B 0 95 0.38 15 0.40 0.85 0.92 0.95 0 92 0.85 0 40 P 1
- Peak Assembly Power k A 4-41
rigure 4-10 TSAR Section 15.4.8 - Control Rod Ejection BOC HFP Assembly Power Distribution at 0.09 Seconds A P tJ M L K J H 5 f C 0 C D A 1 0 33 0 70 0 76 0 60 0.77 0 72 0 34 2 0.31 0 77 0 97 088 0 87 0 86 0 89 0 91 1 00 0 82 C 33 3 03I O B5 0.99 1.00 f .00 076 0.90 0.79 1.1 I l.00 1 07 0 93 0 34 4 0 77 0 99 0 07 0 97 0 82 0 94 0.75 0 97 0 87 1 05 0 94 1 10 0 87 5 0 33 0 97 1.00 0.97 108 1 08 0.79 0 92 0 82 1.16 1.19 109 1 14 1.12 0 39 6 0.70 0 88 1 06 0 82 1.05 0 82 0 91 0.08 0 95 090 1 22 0 96 125 1 00 0 85 7 0.76 087 0.77 094 0.79 0.91 0 83 0 85 0.90 1 03 0 93 l 15 0 95 f I1 0 97 0 000 0.07 0.90 0.75 0.92 0 88 0 85 0 48 0.94 104 1.14 0 98 1.19 1.16 1.07 9 0.78 089 0.78 0.98 0.82 0.96 0 90 0.94 1 01 1.18 1.09 f .3 6 1. I 2 1.27 1.09 to 0,72 0.91 1.11 0.87 1.16 0.90 1.03 1.04 1.18 t .17 1.65 1.31 169 f .3 9 1 07 11 0.35 l.02 1.06 1.05 4.19 1.22 0.93 1.14 109 1.65 1.82 1.76 1.78 1.64 0 55 l 12 0.83 1.07 0.94 1.09 0.95 1.15 0.98 1.36 1.31 1.76 1.88 1.92 1.40 r 1 13 0.33 0,93 1.10 1.14 1.25 0.95 l.19 1.12 1,69 1.77 1.92 164 059
- L A 14 0.34 0.87 1.12 1.06 1.10 1.16 1.27 1.39 1.64 1.40 0.59 15 0.39 085 0.95 1.07 1.09 1.07 0.55
- Peak Assembly Power k J 4-42 1
l l
Figure 4-11 FSAR Section 15.4.8 - Control Rod Ejection BOC HZP Core Power vs. Time 4 10 , i 10 3 I 1 102 .i I
) %
! % m 1 m -
10
! N j
5 10 0; u S' 10'l-o I a, e 10 - 2,: u l l o . u 19 3. . 10'4, ! 10' 0 ! 10 6. 10 ' 7 ,' '
!)
10-8' ,, ,,, , , , ,, , , , , , , , 0.0 0.5 1.0 1.5 2.0 2.5 3.P 3.5 4.0 Time (Seconds) 4-43 _ . . . - _ _ _ ~ , _ , . . _ .
1 l Figure 4-12 ; FSAR Section ISA 8 - Control Rod Ejection 90C HZP Assembly Power DistrtDutton at 0.0 Seconds n c t. M t. t a n . r t o c e a 1 0 30 0 68 0 63 0.50 0 63 0 66 0 30 2 05 1.19 1.33 0 84 0.76 0 45 0.76 0e4 1 33 1.19 0!O 3 0 50 14! 1.59 157 1 39 0 E2 0 et 0 e2 1 39 I57 1 59 I 4! O!O 4 1 20 159 0 09 1.51 1 19 1 16 0 84 1 15 1.19 iSI 0 09 I 59 1.20 5 0 40 l 35 150 1.51 f .67 1 35 0 09 102 0 09 1.35 67 1.51 150 t ?5 O 40 6 0 70 0 05 14 1.20 1.35 054 0 88 0 89 0 00 0 54 1.35 120 1 40 0 85 0 70 7 064 0 77 0 03 1.16 0.89 0 00 0 78 0.76 0.78 D e8 0.89 1.16 0.82 0 77 004 8 0.58 0 45 0 e7 0 85 1.03 0 89 0.76 0.27 0 76 0 89 I 03 0 85 0 07 0.45 0 Se 9 064 0 77 O e2 1.16 0 89 0 88 0.78 0 76 0 78 0 88 0 89 1.16 0 83 0 77 064 10 0.70 0 05 1.40 1.20 1.35 054 0 88 0 09 0.88 0.54 l.35 1.20 1.40 0.e5 0 70 F 7 11 0 40 1.35 1.58 1.51 1.67 1.35 0.89 1.02 0.89 l.35 1 67 151 1 58 1.35 0 40 L .)
'2 1.20 1.59 0.89 1.51 1.19 1.16 0 84 1.16 1 20 1 51 0.89 1.59 1.20 13 0.53 i el 1,59 1.57 1.39 0.82 0 86 0 82 1.39 1.57 1.59 1.41 0.50 14 050 1.19 f.33 0.84 0.76 0.45 0.76 0.04 1.33 1.19 0.50 15 0.38 0.68 0.63 0 58 0.63 0 68 0.38 P 1
- Pea < Assembly Power k J t
l 4-44
1 l Figure 4-13 FSAR Sectton 15A 8 - Control R0d Enctl0n DOC b2P Assembly Power DIStrIbut10n at 0 2 Seconds ri P t, M L K J H 6 i [ ; C h A 1 0.06 0.12 0.12 0.13 0.17 0 21 0.12 2 0 00 0 19 0 21 0.14 0.15 0 11 0 22 0.21 0 46 0 43 0 19 3 0 00 0 ?2 0 25 027 0 20 0.10 0 23 0 20 0 49 050 0 62 O!e 0 22 4 0.19 0 25 0 15 0 27 0 25 0 29 0 25 0 41 0 47 0 03 0.39 0 73 O!7 5 0 00 0 22 0 07 0 27 0 33 0 30 025 0 36 0 30 002 Of3 0 79 0 84 0 73 0 23 6 0 12 0.14 0 26 0 25 0 30 0.14 0 31 0 39 0 46 0 32 0 06 0 79 0 92 C'7 0 49 7 0 12 0 15 0 18 28 0 25 0 31 0 34 0.43 0 57 0 77 0 82 1 09 0.77 0 72 0 61 0 0.13 0.11 0 23 0 25 0 30 0 39 0 42 0 21 0.78 1 07 1 38 1 21 1 27 Ot5 0 83 i 0.17 0 22 0 26 0 41 0 36 0 46 0.57 0 78 1.02 1 40 1.68 2 40 1 75 160 1 27 10 0.21 0 27 0 48 0 47 0 61 0.31 0.77 1.07 1 40 1,12 3 41 3 23 3 79 2.17 1.70 11 0.13 0 46 0.58 0 63 0 83 0 85 0.82 I.37 1 68 3 40 5 08 5 24 5 22 4 12 1.14 7 7 12 0 43 06l 039 0.79 0.79 1.09 1.20 2.39 3 22 5.23 5 23 6.16 4 09 h J 13 0 19 0.57 0 72 0.83 0 91 0.77 1.26 1.74 3.78 5.20 6.15 5 35 i e2 I4 0.21 0 56 0.71 0.56 0.71 0.64 1.58 2. I 5 4 07 4 07 1.8 l 15 0.22 0.48 0 60 0 82 1,25 1.66 1.08 Y M
- Peak Assembly Power k M 4-45
Figure 4-14 FSAR Section 15A 8 - Control Rod Eject 10n BOC HZP Assembly Power DistrtDu! ton at 0 34 Seconds it r t. M L L J H 0 t ( 0 c t, A I 0 09 0 10 0 10 0 10 0 20 0 26 0 15 2 0 11 0 26 0 ?O 0 00 0 20 0 14 0 20 0 33 0 56 C !3 0 23 1 3 0 11 0.30 0 35 0.36 0 35 0 23 0 29 0.32 0 59 0 70 C 73 009 05 4 0 20 0 35 0 20 0 37 0 32 0 30 0 31 0 49 0 55 0 74 O c6 0 05 Of6 5 009 0 30 0 36 0 37 0 43 0 39 0 31 C 42 0 42 0 71 0 94 O e9 0 95 0 03 0 26 0 0 16 00 0 35 0 32 0 39 0 10 0 36 0 44 0 51 0 35 0 93 0 e5 1 01 004 055 7 0.16 00 0 23 0 36 0.31 0 36 0 38 0 46 0 60 0 79 0 64 1 13 0 01 0 70 0 65 8 0 to 0.14 0 29 0.31 0 42 0 44 0 46 0.22 0.79 1 06 1 35 I le 1 26 006 0 05 9 0 22 0,20 0 32 0 49 0 42 0.51 0 60 0.78 1.00 1 35 1 60 2 20 1 67 1.56 1 26 to 0 26 0 33 0 59 0.55 0.70 0 35 0.79 1.05 1 35 1.07 3.16 2 96 3.54 2.11 1.65 11 0,16 0.56 0.69 0.73 0.93 0.92 0 83 1.34 f .59 3.15 4 61 4 77 4 76 3 02 1.07 r 7 12 0.52 0.73 0 45 0 00 0 05 1,12 1.10 2.25 2 95 4 76 4 72 555 3 74 L A 13 0 23 0 68 0.84 0 94 1.00 0 80 1.25 1 66 3 53 4 74 5 54 4 83 164 14 0 25 0 65 0 82 0 63 0.75 0 66 1.55 2 09 3.79 3 '2 164 15 0.25 0 54 0 64 0 84 1 24 I 61 1 02 P 7
- Peak Assembly Power k ..
l 4-46 l l
l Figure 4-15 i FSAR Section 15.4.8 - Control Rod Ejection i EOC HFP Core Power vs. Time ] 1 180 4 160 4 e 140 -- -
.- 120 -_
s 2 n 100; A ~ - 3 80 N '
\
60
\
N.
\
40 20 x % T T 3 Y Y 5 3 V W W W W 1 W T T T T T T T l 0,0 0.5 - 1.0 1,5 2.0 2,5 3.0 3.5 4.0 Time (Seconds) 4-47
.. - _ . - -- - - - . _ . . . - . - - . - . - - , - _ - - - - - . _ . , , . _ _ . . ~ . . . . -.
rigure 4-16 TSAR Section 15.4,8 - Control Rod Ejection EOC !!TP Assembly Pcaer Distribution at 0.0 Seconds R P N M L K J H 0 I ( 0 C D A 1 0 50 0 08 0 09 0 90 0 09 0 00 0!O 2 0 50 1 05 1.23 ' 04 0 98 0 94 0 90 1 04 1 23 1 05 0 50 3 0 50 1.15 131 1 29 1.31 0 80 0 96 0 00 131 129 13? 1 15 O*0 4 1(5 i 31 1 20 1.33 1 01 1 03 ODI i 03 I 01 1 33 120 1 31 1 05 5 O!O I 23 1 29 I33 1 30 1.38 0 05 0 91 0 05 1 71 1 38 I 33 1 29 123 0 50 6 0 85 1 05 1.31 1 01 1.31 0.91 0 90 0 02 0 90 0 91 1 31 1 01 111 4 05 0 08 7 0 00 0.90 0 00 1 03 0 05 0 90 0.70 0 77 0 90
- 0. 7 e 0 05 1 03 0 00 0 98 0 89 8 0 90 0 94 0 96 0 81 0.91 0 02 0.77 0 49 0.77 0 82 0 91 O el 0 96 0 94 0 90 4 0.09 0.98 0.80 1.03 0 65 0.90 0.78 0.77 0.78 0.90 0.05 0 08 1.03 0.96 0 09 10 0.08 1 05 1.31 1.0 ! 1.31 0.91 0.90 0.e2 0.90 0.9 i i .3 i i .0 i i .3 i i.05 0 ee
~
r 7 11 0 50 1.23 1.29 1 33 1.38 1.31 0 65 0.91 0.05 1.31 1.30 1.33 1.29 123 0.50 L A 12 1.05 1.31 120 1.33 1 01 1.03 0 01 1.03 1.01 1.33 1.20 1.31 1 05 13 0.50 1.15 1.31 1.29 l.31 O b8 0 96 0 00 1.31 1.29 1.31 1.15 0.50 14 0 50 1 05 1.23 1.04 0.98 0.94 0.98 1.04 1.23 1.05 0 to 15 0 50 0.88 0.09 0.90 0 09 0 68 0.50 YM
- Peak Assembly Power k J 4-48
rigure 4-17 TSAR Section 15.4.8 - Control Rod Ejection EOC HrP Assert.bly Power Distributicn at 0.09 Seccnds k r it '1 L k J n 0 t C C- C b - 1 0 39 0 f'8 0 70 0 72 0 72 l 0 72 0 42 2 0.38 0 00 0 94 0 BI 0 77 0 75 0 60 0 6C ! .03 0 00 0 43 3 0 38 0 07 0 99 0 99 1 02 0 70 0 77 0 73 1 10 1 09 I 11 099 0 44 4 0 00 C 99 0 91 1.03 0 79 0 02 0 66 0 07 0 97 1 15 102 1 15 0 95 5 0 39 0 94 0 99 1 03 1 07 104 0 69 0 76 0 74 I I5 1.23 12C 1 19 1 15 0 20 6 000 0 01 102 0 79 1 04 0.74 0 75 0 71 0 01 D e4 1 22 0 97 127 1 03 0 00 7 0 70 0 77 0.70 0 02 0 69 0.75 0 60 6 69 0 75 0 69 0 07 107 0 93 1 05 0 95 0 0 72 0 75 0.77 0 66 0 77 0 71 0 69 0 46 0 79 0 90 1 03 0 94 1 12 l.10 1 05 9 0 72 0 00 0.73 0 07 0 74 0.81 0 75 0,79 0 00 1 00 1 07 1 33 1,15 1 26 1 12 10 0.72 0 86 1.10 0.87 1.15 0 04 0 09 0.90 1.08 1. l ? I 78 1 43 164 1 45 1.19 11 0 42 1.03 1 09 1.15 1 23 1.22 0.87 1.03 1.07 1.78 / 03 2 05 1.97 1 00 0 72 P 1 12 0 09 1,11 1 02 1 20 0.97 1.07 0 94 1.33 1 43 2 05 2.18 2. l; l 62 L A 13 0 43 0 99 1.15 1.19 1.27 0.93 1.12 1.15 1.84 1 96 2 13 l e6 0.90 14 0 44 0.94 1.15 l 03 1.05 1.10 1.26 1 45 1.80 162 0 80 15 0 48 0.08 0.95 1.05 1.12 1.19 0 72 I 9
- Peak Assembly Power WA 4-49
Figure 4-18 FSAR Section 15.4.8 - Control Rod Ejection EOC llZP Core Power vs. Time 10^ , i 103[ ; ; 102[' !
~
s. 10 1
\m ..--
a . N s~_
# m a 9 b
y 10 0 , [ O :. a . 9 10*1 , o ! o : 10-2l' , 10-3, : i .i 10'4, _
- i 10-5_
I 10-6[l 10-7 ) , , - - . . , ,,,, ,, - - u
' . 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4*0 Time (Seconds) 1 ,,
4-50 s
--..h
Tigure 4-19 FSAR Section 15.4. 8 - Cont rol Rod Ejection EOC H"P Assembly Power Distribution at 0.0 seconds D P 14 M L K J H 0 r [ t ,_ r: A 1 0 52 0 77 064 0 56 0 64 0.77 0 52 2 OfD 1 44 150 0 92 0 72 0 44 0 72 0 92 1 50 1 44 0 68 3 Ofe 1.71 1 02 174 I 44 0 76 0.72 0 76 1 44 1 74 I 82 1 71 Ofe 4 1 44 1 02 1.11 1 60 1.11 0 96 0 69 0 96 I il 1 60 1.11 i e2 1 44 5 O !! 150 1 74 1 60 1 69 I 26 0 ?! 0 74 0.71 1 20 1 69 I 60 174 50 O!2 6 0 70 0 93 1 45 1.11 1 26 0 49 0 60 057 0 60 0 49 1 26 1.11 1 45 0 93 0 78 7 Of5 0 72 0 70 0 96 0.71 0 00 0 49 0.44 0 40 0 60 0.71 0 96 0 76 0 72 0 65 8 O!f 0 44 0 72 0 09 0 74 0 50 0 44 0.18 0 44 0.58 0 74 0 69 0 72 0 44 0 56 9 0 65 0 72 0 76 0 96 0 71 0 60 0 48 0.44 0 40 0 60 0.71 0.96 0 76 0 72 0 65 10 0 7e 0 93 i 45 i.i i i .2e 0 49 0 60 0.57 O eo 0 49 i .2e i.ii ix5 0 93 0 7e 11 OTO 1.50 1.74 1.60 1.69 1.26 0.71 0.74 0.71 1.26 169 1 68 1.74 1.50 052 r 7 12 1.44 1.02 1.11 1.68 1.11 0.96 0.69 0.96 1.11 1.60 1.11 1.e2 1 44 L A 13 0 68 1.71 1.02 1.74 1.44 0 76 0 72 0,76 1.44 1.74 1.82 ! .71 068 14 0 68 1.44 1.50 0.92 0.72 0 44 0 72 0.92 1 50 1.44 068 15 0.52 0.77 0 64 0.56 0 64 0 77 0 52 Y M
- Peak AssemDly Power N A 4-51
Figure 4-20 FS AR Sect 10n 15.4 9 - Control Rod Ejection EOC HZP Assembly Power Distributton at 013 Seconds n n u M L r. s o c r t 0 t 0 - 1 0 03 0 05 0 05 0 06 0 09 0.13 0.10 2 004 0 00 0 09 0 06 0 06 0 00 0 12 0 16 0 31 0 32 0 16 3 0 04 0 09 0 10 0 11 0 10 0 07 0 10 0 14 0 30 0 39 0 44 0 44 O tf 4 0 0e 0 07 O iO O ii 0 09 O ii O ii 0 20 0.27 e 44 03i 0:3 e 43 5 OC3 0 09 0 11 O il 0 13 Ott 0 10 0.15 0 86 0 39 O!6 0 50 0 61 0"3 O IC 6 0 0$ 0 00 0 10 0 09 0.11 0 06 0 12 0 17 0 23 0 20 0 50 0 53 0 67 0 42 0 3f 7 0 01 OC0 0 07 0 11 0 to 0.12 0 14 0 20 0 32 0 40 0 50 0.79 0 00 0 54 0 47 0 0 00 00* 0.10 0.11 0 15 0 17 0 20 0,14 O SO 0 77 1 09 1.05 1 08 0.58 0 75 9 0 09 0.12 0.14 0 20 0.18 0.23 0 32 0 50 0 74 1.16 1.62 2.31 1.00 1.57 1.31 10 0.13 0.16 0.30 0.27 0 30 0.20 0 40 0.70 1 16 I 20 3.03 3.52 4 30 2 40 1.96 11 0.10 0,31 0 39 0 44 0 56 0.58 0.50 1 09 161 3.03 6 20 6 75 6 47 4 98 ' 56 F 1 12 0 32 0 44 0.31 0 50 0.53 0.70 1.05 2.31 3.52 6.75 7.19 7 90 5.39 L A 13 0 16 0.44 0.53 0 61 0 67 0.59 1.07 1.00 4 37 6 46 7.90 7 17 2 60 14 0.10 0 43 0 53 0 41 0.53 0.57 1.56 2 39 4 97 5 30 2 60 15 O. t 9 0.30 0 49 0 75 1.30 1.95 1.56 P 1 l
- Peak AS5emDly Power <
k J 4-52 i i I
1 FAgure 4-21 PSAR Section 15.4.8 - Control Rod Ejection EOC H2P Assembly Power Distribution at 0.17 Seconds D 0 1. M L A J H G f E D C D ^ l 0 06 009 0 09 0.11 0 16 0 21 0 16 2 0 07 0 15 0 16 0 11 0 11 OC9 0 19 0 26 0 48 0 40 0 24 3 0 07 0 10 0 19 0 20 0 18 0 12 0 16 0 22 0 46 059 C CS 064 0 27 4 0 15 0 19 0 12 0 20 0 15 0 17 0 17 C: 30 0 40 0 64 0 45 0 76 0(2 5 0 06 0 16 0 20 0 00 0 22 0.19 0 15 0 21 0 20 O!2 C 76 0 79 0 03 C 73 0 27 6 0 09 0 11 0 18 0 15 0 19 0 09 0 16 0 22 029 0 26 0 73 0 66 0 06 0 55 0 51 7 0 09 0.11 0 10 0 17 0.15 0.16 0 18 0 24 0 36 O!4 0 65 0 88 0 60 0 64 0 00 0 0 0 09 0 16 0 17 0.21 0 22 0 24 0 15 052 0 79 1 06 1.11 I ll 0 63 C e3 9 0 16 0 19 0.21 0 30 0.26 0 29 0.36 052 0 75 1.14 154 2 20 1.74 i.50 1 35 1 to 0.21 0 26 0 45 0.39 0.52 0.26 0.53 0.78 1.14 1.15 351 3 19 4 06 2.37 1.94 11 0.15 0 47 0 58 0 63 0.75 0.72 0 64 1.10 1.54 3 49 5 46 5 92 5.71 452 1 47 F 7 12 0 48 064 0 44 0.78 065 0 87 1.06 2.19 3.19 5 92 6 17 6.76 4 74 L A 13 0 24 0 63 0.75 0.82 0 05 0 68 1.10 1.73 4 04 5 70 6 75 6 15 2 34 14 0.26 0 61 0,72 0.55 0 63 0 63 1,57 2.36 451 4 73 2 34 15 0 27 0 49 0 58 0.81 134 1.93 1 47 Y M
- Peak Assem0ly Power k A 4-53
l l l Figure 4-22 I' FSAR Section 15.4.8 - Control Rod Ejection BOC liFP MARP Curves M l l l i
=
0. 7 ~ l c . 1
= '
0 i 1 l I M 4-54
. . . . _ -- . . . _ . - . - - . . _ . - . _ - - . . . - . . - ~ .
Figure 4-23 FSAR Section 15.4.8 - Control Rod Ejection BOC llZP MARP Curves Gud I I w
- c. .l 5
b '
@ i 3 l
.?<
m. f. l 4-55 i
Figure 4-24 FSAR Section 15.4.8 - Control Rod Ejection EOw HFP MARP Curves e - O.
~n a
E a 8 e W. 4-56
_1 ;;;- j;:ggg, m_ a ~ l N 4-57
j Figure 4-26 PSAR Section 15.4.8 - Control Rod Ejection BOC HFP Pins in DNB by Assembly R P N M L K J H 0 f E D C D A 1 0 0 0 0 0 0 0 2 0 0 12 0 0 .0 0 0 48 4 0 3 0 0 0 0 55 0 0 0 149 0 0 6 0 4 0 -c o 0 0 0 0 4 0 7 0 30 34 5 0 12 0 0 109 150 0 0 0 244 264 32 167 163 0 6 0 0 56 0 149 0 0 0 70 4 264 0 264 109 63 7 0 0 0 0 0 0 0 0 10 256 73 264 46 264 131 0 0 0 0 0 0 0 0 0 43 264 264 225 264 204 169 9 0 0 0 3 0 70 10 41 240 264 264 264 264 264 180 i , 10 - 0 0 148 0 244 4 254 264 264 264 264 264 264 264 165 l II O 49 I 6 264 264 65 264 264 264 264 264 264 264 0 12 4 0 0 : 31 0 264 233 264 264 264 264 264 219 13 0 6 36 163 264 52 264 264 264 264 264 256 28 14 0 34 162 106 264 264 264 '264 264 219 28 15 0 61 130 168 180 164 0 Total Pins in DNS = 18806 (36.9IE) 4-58
Eigure 4-27 PSAR Section 15.4.8 - Contr01 P.od Ejection DOC HZP Pins in DNB by Assembly R P M L J t; L H 0 f [ 0 ( g 4 I O O O O O O O 2 0 0 0 0 0 0 0 0 0 0 0 3 0 0 0 0 0 0 0 0 0 0 0 0 0 4 0 0 0 0 0 0 0 0 0 0 0 0 0 5 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 6 0 0 0 0 0 0 0 0 0 0 0 0 C 0 0 7 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 8 C 0 0 0 0 0 0 0 0 0 0 0 0 0 0 9 0 0 0 0 0 0 0 0 0 0 17 162 0 0 0 10 0 0 0 0 0 0 0 0 0 1 251 264 264 46 0 11 0 0 0 0 0 0 0 0 250 21 264 264 264 257 0 12 0 0 0 0 0 0 0 157 264 264 264 264 220 13 0 0 0 0 0 0 0 0 264 264 264 263 70 14 0 0 0 0 0 0 0 42 251 220 68 15 0 0 0 0 0 0 0 Total Pins in OND = 5464 (10 72%) 4-59
\
Figure 4-28 FSAR Section 15.4.8 - Control Rod Ejection EOC HFP Pins in DNB by Assembly A P N M L K J H 0 f E D C D A l 1 0 0 0 0 0 0 0 1 2 0 0 0 0 0 0 0 0 0 0 0 3 0 0 0 0 0 0 0 0 0 0 0 0 4 0 0 0 0 0 0 0 0 0 0 0 0 0 5 0 0 0 0 0 0 0 0 0 0 0 0 0 5 0 6 0 0 0 0 0 0 0 0 0 0 12 0 29 0 0 7 0- 0 0 0 0 0 0 0 0 0 0 0 0 0 0 8 0 0 0 0 0 0 0 0 0 0 0 0 0 0 56 9 0 0 0 0 0 0 0 0 0 i8 12 253 70 170 til 10 0 0 0 0 0 0 0 0 18 113 264 264 264 261 135 11 0 0 0 0 0 9 0 0 15 264 264 264 264 262 0 12 0 0 0 0 0 0 0 254 264 264 264 264 2i? 13 0 0 0 0 29 0 0 64 264 264 264 256 38 14 0 0 3 0 0 0 169 260 262 217 38 15 0 0 0 54 111 134 0 l j rotai pins in one - 735 i < i 4 43s> I 4-60
Figure 4-29 FSAR Section 15.4.8 - Control Rod Ejection EOC li3P Pins in Lt/B by Assembly R P N f1 L K J H G F E D C D A i 0 0 0 0 0 0 0 2 0 0 0 0 0 0 0 0 to 36 0 3 0 0 0 0 0 0 0 0 0 39 176 123 0 4 0 0 0 0 0 0 0 0 0 68 5 170 62 5 0 0 0 0 0 0 0 0 0 0 149 47 166 69 0 0 0 ') 0 0 0 0 0 0 0 0 0 0 t3 0 0 7 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 0 5 0 1 2 0 9 0 0 0 0 0 0 0 0 0 4 58 207 111 107 67 10 0 0 0 0 0 0 0 0 4 39 264 264 264 251 164 11 0 13 45 69 150 0 0 0 60 264 264 264 264 264 70 12 31 182 6 148 0 0 0 207 264 264 264 264 260 13 0 127 171 107 13 0 16 l10 264 264 264 264 145 14 0 63 69 0 0 107 3 250 264 260 145 15 0 0 0 0 67 160 69 Total Pins in DNS = 9970 (19.57S) 4-61
Figure 4-30 Core Coolant Volume Expansion Rate for IEP, BOC Case 1 e O O u;. b= E C $M C
=
0. M LLI
= .,
Time (seconds) 4-62
ROD EJECTION 2800 -- r
-~- - - - --~" - ' - - - ~ ~ ~'
2700 -
- - - - - - - - - - - - - - - - --- - - - ~ -
m m b w 5 2600 -- - - -- - - - - - - - - - - - -- - - - u M ' d 1,1 o u& h
<o N 3 h I
cn " g E'
, 2500 .
-- - - ~ -- - - - - - -
S. . 2400 - - - - - - - - - - - - - - - - - - - - - - - - ~ ~ - - 2399 .,,,7,,7.,.., ,, , , , . , - , .,,.7, ., ,_,,,r,,.,. . , . , , , , . , , , , , . , . , , , , , , , . . , . , , , , , , , , , , . 0.0 0.5 1.0 1.5 2.0 2_5 3.0 35 TIME (SECONDS)
y Sn
+<
l 5.0 STEAM LINE DREAK ANALYSIS I 5.1 Oyerview 5.1.1 Description of Steam Line Break Accident The steam line break transient is described in FSAR Section 15.1,5 (Ref erence 5-1). The steam release arising from a break in a main steam line would result in an initial increase in steam flew, with a subsequent decrease during the accident as the steam pressure falls. The energy removal from the Reactor Coolant System (RCS) causes a reduction of coolant temperature and pressure, in the presence of a negative moderator temperature coefficient, the cooldown results in an insertion of positive reactivity. If the most reactive control rod is assumed stuck in its fully withdrawn position after reactor trip, the core might become critical and return to power. A return to power following a steam line rupture is a potential problem mainly because of the high power peaking factors which exist assuming the most reactive control rod 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. 5.1,2 Acceptance Criteria A major steam line break is classified as an ANS Condition IV event, a limit- ; ing fault. Minor secondary system pipe breaks are classified n ANS Condition III or infrequent events. The analysis is performed assuming a stuck control rod, a single failure in the engineered safety features, and with considera-tion of both offsite power maintained and offsite power lost. The following two criteria must be satisfied. First, the core must remain in place and intact. The analysis submitted herein meets this criterion by showing that the 95/95 DNB limit of Reference 5-2, Section 4,4 is satisfied. Future analyses using these same methods might meet the criterion by demonstrating continued core cooling capability based on an acceptable fuel damage model and result. Second, radiation doses must not exceed the guidelines of 10CFR100. These dose limits are 25 rem whole body and 300 rem thyroid. The Condition III and IV criteria regarding overpressurization are not challenged by a steam line break transient. 5-1
5.1.3 Analytical Approach The steam line break transient requires a limiting set of physics parameters
- to be determined for use as initial and boundary conditions. These parameters are input to a McGuire/ Catawba RETRAN-02 (Reference 5-3) model for the system thermal-hydraulic analysis. The RETRAN-02 analysis generates the core state-point conditions which correspond to the transient time of minimum DNBR.
Neutronics ' codes such as EpRI-NODE-P (Reference 5-4) or SIMULATE-3P (Reference , 5-5) are.used to generate core power distributions corresponding to the statepoint conditions. The core power distribution along with the core thermal-hydraulic boundary conditions from the RETRAN-02 analysis are then input to a McGuire/ Catawba VIpRE-01 (Reference 5-6) model to calculate the minimum DNBR. If this value were below the DNBR limit, then a fuel rod census would be performed to determine the number of fuel rods in DNB and therefore the fraction of gap activity released. The dose consequences of this release would then be evaluated. 5.2 Simulation Codes and Models 5.2.1 System Thermal-Hydraulic Analysis 5.2.1.1 Selection of a Bounding Unit Differences between the McGuire and Catawba units are discussed in Section 3.1,6 of Reference 5-7 for the steam line break transient. The most_important differences with respect to steam line break are the steam generator type and the Auxiliary feedwater System runout protection. The steam generator type influences the transient via the initial liquid inventory. A higher faulted steam generator liquid inventory delays tube uncovery and emptying, thereby enhancing primary-to-secondary heat transfer and producing a more severe RCS cooldown. The Auxiliary Feedwater System runout protection influences the transient via the pump alignment and the maximum open position limits on the control valves. Other factors influencing auxiliary feedwater flow delivered to the faulted steam generator are pump discharge piping resistance and pump capacity, 6-2 w
Of McGuire and Catewba, the Catawba units currently have the higher faulted steam generator delivered auxiliary feedwater flow. Of the two Catawba units, Unit 2, because of its counterflow preheat steam generators, currently has the larger initial liquid inventory, Catawba Unit 2 is, therefore, chosen as the most conservative unit of the four for analyzing the steam line break tran-sient. 5.2.1.2 Modifications to Base Plant Model Renodalization of Reactor Vessel .
)
s g b 5-3
--_-___-_-- _ _ - _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ _ _ _ - _ _ - _ _ - - _ _ - _ _ _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _-____ a
Renodalization of Steam Generator Secondary-5.2.1.3 Break Modeling The. full cross-sectional area of .he 34" main steam line is 5.4 ft'. The area of the flow restrictor at the steam cenerator outlet is 1.4 f t*,
]ThisanalysisusestheMoodycriticalflowmodel. For the
- l. timeframe of interest, the break flow is always 1imited by critical flow.
5-4 l
1 l I 5.2.2 Nuclear Analysis l The transient system response during a steam line break accident is sensitive to core reactivity versus temperature and the Doppler Temperature Coef ficient. The core thermal-hydraulic response is sensitive to the three dimensional core power distribution. Therefore, the nuclear analysis for this event must specify pre-break core physics characteristics and post-break power distribu-tions based on the calculated system response. 5.2.2.1 Core Physics Parameters The k-effective versus temperature curve (Figure 5-2) and Doppler temperature coef ficient are selected such that a limiting return to power is expected to occur in the RETRAN analysis. This curve represents the effect or. reactivity of an isothermal cooldown f rom the technical necification shutdown margin limit. The Doppler coef ficient was chosen to be -3.5 pcm/ F for this analy-sis. The conservatism of the k effective versus temperature curve and Doppler coef ficient will be confirmed each cycle as described in Section 5.5. 5.2.2.2 Powet Di stribution s I
] SIMULATE-3PorPDQ(Reference 5-4)are used to calculate the peak pin to assembly average ratio for the hot assembly.
This pin to assembly factor is applied to the assembly average power calcu-lated for the limiting RETRAN statepoints. The three-dimensional power distribution and the system analysis results are then combined for the thermal-hydraulic evaluation. 5-5 1 ________________ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - - _ _ _ _ _ _ _ _ .~
1 I l 1 5.2.3 Core Thermal-Hydraulic Analytis 5.2.3.1 VIPRE Code Description The VIPRE-01 code (Reference 5-6) is used for the steam line break core thermal-hydraulic analyses. VIpRE-01 is a subchannel thermal-hydraulic computer code. With this subchannel analysis approach, the nuclear fuel element is divided into a number of quasi one-dimensional channels that communi: ate laterally by diversion crossflow and turbulent mixing. Given the geometry of the reactor core and coolant channels and the boundary conditions or forcing functions, VIPRE-01 calculates core flow distributions, coolant conditions, fuel rod temperatures and the minimum departure from nucleate boiling ratio (MONBR) for steady-state conditions and for transients. VIPRE-01 accepts all necessary boundary conditions that originate either from the RETRAN system transient simulation or the core neutronics simulation. Included is the capability to impose different boundary conditions or dif-ferent regions of the core model. For example, dif ferent core region inlet temperatures, flow rates, heat flux, and even different assembly and pin radial powers or axial flux shapes can be modeled in steady-state or transient modes. 5.2.3.2 Analysis Methodology model calculates the statepoint local coolant properties and the DNBR. The critical heat flux (CHF) correlation used to evaluate the DNBR is the Westinghouse W-35 correlation (Reference 5-6, Appendix D). The W-35 CHF correlation has been recently approved by the NRC for analysis with system 5-6
pressures as low as 500 psia (Reference 5-8). The N cBeth CHF correlation (Reference 5-6, Appendix D) may also be used for the analysis. Two stnady-state cases are analyzed: the first case with of f site power available, and the second case with offsite power unavailable. A statepoint DNBR calculation is performed instead of a transient DNSR calculation since the steam line break accident is a slow transient and a statepoint consisting of the limiting surface heat flux and inlet boundary conditions provides conservative DNBR results. Model Description
~
Axial Power Distributions M 5-7
, Radial Power 01stributions
~~ ., ~.
5.3 Transient Analysis 5.3.1 Initial Conditions Pressurizer Pressure Since this transient is being evaluated for minimum DNBR, a low initial pressurizer pressure is used. The low initial pressure causes an earlier safety injection actuation since the transient starts closer to the setpoint. This is compensated for in the safety injection setpoint as described below. Nominal pressurizer pressure with any control rods withdrawn is 2235 psig. The initial condition uncertainty allowance for reduced pressurizer pressure is 3D psi. The-initial condition for this transient is, therefore, 2705 psig. Pressurizer level . A low initial pressurizer level minimizes RCS inventory during the transient. This minimizes core outlet: pressure and is, therefore, conservative for evaluation of minimum DNBR. This effect more than compensates for the
.slightly quicker boration when the safety injection fluid mixes with the
' smaller RCS mass. The hot zero power programmed pressurizer level is 25%.
The initial. condition uncertainty allowance for reduced pressurizer level is 9%. The initial condition for this transient is, therefore, 16%. 5-8
.RCS Temperature Since this transient is being evaluated for minimum DNBR, a high initial RCS temperature is used. A slightly greater reactivity insertion results from starting from a high initial temperature since the slope of the k-effective j
vs. temperature curve is greater at higher temperatures. ihe hot zero power i programmed RCS temperature is 557 F. The initial condition uncertainty l allowance for increased RCS temperature is 4 F. The initial condition for this transient is, therefore, 561 F. RCS Flow Since this transient is being evaluated for minimum DNBR, a low initial RCS flow is used. The effect of lower flow on DNBR more than offsets the decrease in primary-to-secondary heat tran3fer. The Technical Specification minimum measured flow assumed for this analysis is 382,000 gpm. The Catawba flow
-measurement uncertainty is 2 d , which is larger than the corresponding McGuire value. The initial condition for this transient is, therefore, 373,596 gpm.
Steam Generator Water Inventory - Since the primary to-secondary heat transfer is the driving force behind the excessive PCS cooldown and depressurization, steam generator inventory is maximized to provide the largest cooldown capacity and to prolong the time prior to U-tube uncovery and heat transfer degradation. The normal hot zero P , powermassisapproximately; jlbmpersteamgenerator. The initial condition uncertainty allowance for increased steam generator level is 8%. In this region of the steam generator, this is equivalent to an additional Ibm. The initial condition for this-transient is, therefore, approximately
}}bm.
Core Power Initial core heat output would' result in a lower temperature decrease since this energy would have to be removed in addition to that stored in the RCS 5-9
.l i
i l fluid and metal. 4 This_would result in a milder transient and would be. nonconservative. The core is, therefore, initially at hot zcro power, here
-defined as=10-' times full powe,'.
-)
Steam Generator Tube plugginq I Assuming no steam generator tube plugging maximizes the steam generator heat l transfer area and minimizes the RCS loop flow resistance. Both of these q effects enhance primary-to-secondary heat transfer and are, therefore, conser- I vative. These effects more than offset the slight decrease in RCS inventory l which would result from plugged tubes. Therefore, no tube plugging is assumed i for this analysis. Core Bypass Flow Core bypass flow is assumed to be 9% of total core flow. 5.3,2 Boundary Conditions 5.3.2.1 Availability of Systems and Components Reactor Coolant pumps The reactor coolant pumps are assumed to trip when offsite power is lost. For portions of the analysis during which offsite power is maintained, all reactor coolant pumps are-assumed to be operating. Pressurizer pressure Control No credit is taken for pressurizer heater operation. This assumption enhances the RCS depressurization and is therefore conservative for.the evaluation of minimum DNBR. 5-10
pressurizer Level Control No credit is taken for the automatic operation of the Chemical and Volume Control System (CVCS) to attempt to increase RCS mass and thereby maintain pressurizer level and pressure. The charging and letdown flows are assumed to isolate simultaneously and to be balanced prior to isolation. Not taking credit for CVCS action to maintain pressure is conservative for the evaluation of minimum DNBR. Condenser Steam Dump The condenser steam dump valves are initially assumed to be open slightly to release the steam generated by the relatively small heat input to the RCS from the reactor coolant pumps. These valves are assumed 'o be closed after reactor trip. However, since the flow through these. valves is very small compared to break flow, the opening or closing these valves has an insignifi-cant effect on the analysis. Main Feedwater The main feedwater pumps take suction from the hotwell pumps via the conden-sate booster pumps. Both of the latter sets of pumps are run from offsite
- power.
When offsite power is lost, both of these types of pumps trip, causing the main feedwater pumps to trip on low suction pressure. It_ir assumed that this process takes no more than 5 seconds. For events in which offsite power is maintained,'no main feedwater pump trip is assumed. For all cases, no credit is taken for feedwater isolation on low-low RCS average temperature coincident with reactor trip. Auxiliary Feedwater All three auxiliary feedwater pumps are assumed to start on loss of offsite power'and deliver flow to all four steam generators. This is conservative since it maximizes the secondary heat sink. 5-11 1
, . . - - - - - ~ . - .- ~ - - . _
Offsite Power ,
?As instructed by Section-15.1.5 of Reference 5-2, the assumptions regarding the loss.of offsite power and the timing of such a loss were studied to determine their effects on the consequences of the accident. Analyses were perform ' with offsite power both maintained throughout the transient and lost during the. transient. The core is ultimately shut down by borated water from the high'and intermediate-head safety injection pumps. In the absence of offsite power, the pumps are powered from emergency buses energized by diesel generators. The diesels start on either a safety injection signal or an undervoltage condition on 'the emergency buses (indicative of the loss of f offsite power). Since delaying diesel generator start delays borated water delivery, and is therefore conservative, the loss of offsite power is timed to coincide with the safety injection actuation.
Safety injection Pumas The injection of borated water introduces negative reactivity and is therefore a benefit. The injection of cold, unborated water is a penalty, however, since it-makes the cooldown more severe. Because of tir.s, the single f ailure, the loss of one train of safety injection, is timed so coincide with the point at which the high-head safety injection piping is purged of unborated water.
-5.3.2.2- Response Times
; Pumped Safety injection Flow L
A delay is assumed from the SI setpoint being reached until the SI signal is
! generated; An additional delay is assumed from the diesel generator start signal until the first load group, which includes the high-head safety injec-tion pump discharge valves, is sequenced onto the emergency bus. A third delay is assumed from the sequencing of the first load group onto the emer-gency bus until delivery of unborated water to the RCS. The total of.these
, three delays is 33 seconds. For the case in which offsite power is maintained, the corresponding delay is 19 seconds. l L 5-12
.,, .- , . . - - - , ,,,- -c-- - - - -
Feedwater Isolation Valves l Following the receipt of a safety injection signal, an additional delay is assumed to generate a feedwater isolation signal and complete closure of the isolation valves. The total response time for the feedwater isolation function is 12 seconds. Main Steam _ Isolation Valves Following the receipt of a steam line isolation signai, an additional delay is _ _ _ assumed to close the main steam isolation and main steam isolation bypass valves. The total response time for the steam line isolation function is 10 seconds. Auxiliary Feedwater Pumps Since cold auxiliary feedwater flow into the steam generator makes the cool-down more severe, no time delay is assumed between the loss of offsite power and the delivery of flow to the steam generators. 5.3.2.3 Flow From Interfacing Systems Safety Injection Safety injection flow is varied as a function of RCS pressure. The limiting head-flow curves among the high and intermediate head pumps are adjusted to conservatively account for pump head degradation. Main Feedwater At hot zero power, the main feedwater control valve is closed, and the feed-water is delivered to the steam generator upper nozzle through the main feedwater control bypass valve. In assessing the amount of main feedwater flow during a steam line break, the following aspects must be considered: automatic control of pump speed, automatic control of bypass valve position, and line resistance of the piping to the upper nozzle. The speed controller 5-13
- - _ _ _ _ - _ _ _ - _ - - _ _ _ _ - _ - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - ^
, - ._. - - . .- . - - . ~ - - . . _- . , .
will initially attempt to-reduce pump speed. -No credit is taken in-the analysis for a flow reduction due to this effect. Rather than model the bypass valve controller in detail, the analysis conservatively assumes that the valve instantaneously travels to its full open position. A lower limit is placed on the: Upper nozzle piping resistance in this configuration. The flow boundary condition is then conservatively increased as steam generator pres-sure decreases, by assuming that feedwater pump discharge pressure remains constant at the initial.value corresponding to the low resistance limit.
'uxiliary_Feedwater Auxiliary feedwater flow is varied as a function of steam generator pressure.
The limiting head-flow curves among the motor and turbine-driven pumps are adjusted to conservatively account for installed-pump performance being better than the-curves and for pump motor speed being higher than predicted. 5.3.2.4 Engineered Safety Features Actuation Setpoints Safety Injection Safety injection is. assumed to be actuated at 1700 psig pressurizer pressure or at 700 psig steam line pressure, whichever occurs earlier. Steam Line Isolation Steam.line isolation is assumed to occur at 700 psig steam line pressure. No
- credit is taken for steam line isolation on high containment pressure for breaks-inside containment.
~l
_ Dynamic Compensation of Steam Line Pressure Sianal I No credit is taken for the lead / lag compensation on the steam line pressure signal .for' actuation of steam _line _ isolation or_ safety injection. This I results in later actuation for each function causing, respectively, prolonged blowdown of the intact steam generators and delayed boration. Both of these 1 5-14
i I effects make the transient more severe, and this modeling therefore bounds both the presence and the absence of thc lead / lag compensation. 5.3.2.5 Boron Injection Modeling Transpoy The boron transport model used is described in Section VII.2.5 of Reference 5-3. The boron is assumed to be soluble in the transport medium and to have no direct effect on the fluid equations. The basic equation computes the time rate of change of boron mass in a control volume from the net inflow from connected volumes plus the net generation within that volume. Purge Volumes Purge volumes from tne outlet of the refueling water storage tank to the inlet of the RCS are separately calculated for both the high and intermediate-head safety injection pumps. These piping volumes are assumed to be initially at a concentration of 0 ppm. Borated water is assumed to reach the RCS only after an amount of unborated water equal to the purge volume has been injected. This purging is done separately for the high and intermediate head pumps. Concentration The boron concentration in the injection water is an assumed 1900 PPM Refueling Water Storage Tank Technical Specification lower limit value minus a 1% concentration measurement error, or 1881 ppm. 5.3.2.6 Core Kinetics Modeling Poir,'. Kinetics The RE M N point kinetics model is used for the system thermal-hydraulic analysis. The particular option employed uses one prompt neutron group, six delayed neutron groups, eleven delayed gamma emitters, plus U-239 and Np-239. The point kinetics model is adequate for this application since the system 5-15 _ _ _ _ _ _ - - - ___. - - - - - - - - --------- --------- ----- ------ - - - - - ---- - - - - - - ~ - - - - ~
analysis does not require detailed modeling of power distribution effects. The power distributions used in the system analysis are determined to be
-conservative as discussed below.
-The effective delayed neutron fraction and the prompt neutron lifetime values are chosen to minimize the ratio of the former to the latter. This ratio is a RETRAN inout. Minimizing it increases the neutron power spike when prompt criticality is achieved.
T_emp_erature Feedback i The basis for the *.emperature feedback is a relationship between the reactivity vs. temperature curve and i, which is input to the point kinetics model. Axial Power Distribution The axial power distribution for the RETRAN analysis is simply the energy depositien fraction for each of the three axial core conductors. These fractions approximate the axial power distribution calculated by the three [ dimensional core model described in Section 5.2.2.2. The RETRAN axial powar distribution at the peak heat flux statepoint is more top peaked than the. distribution calculated by the three. dimensional model. This approach is conservative since it results in a more severe return to power. i l i 5-16
- i. !
l
kn Radial Power Distribution ~ 4 This-approach is conservative since it results in a more severe return to power. Control Rod Reactivity l Since the steam line break transient is a concern chiefly because of power peaking in the vicinity of a stuck rod, the control rods are assumed to begin the, transient outside of the core; i.e., the reactor is initially not tripped. Manual action by the operator is assumed to immediately trip the reactor. This assumption is conservative since any cooldown prior to rod insertion
- would introduce positive reactivity which would increase core power, This
- would increase RCS-stored energy and cause decay heat generation, both of which cause a less severe cooldown. The amount of negative reactivity intro-duced by rodL insertion is sufficient to make the core subcritical by the technical = specification shutdown. margin.
- Boron Reactivity-The' negative reactivity inserted by boration is modeled by
; core boron concentration.
This concentration is multiplied by a boron worth to give a reactivity, 5.4 Results and Conclusions
- 5.4.1 Primary and Secondary System Response Sensitivity studies were performed to demonstrate that the 1.4 ft* break size is limiting. The steam'line break transient is analyzed both with offsite power maintained and with offsite power lost coincident with safety injection 5-17
. . .- - - ~ - - . . . -.- - . - . - . _ . . . - . - - _ _ ~ . . - -
I
-1 actuation. --The event sequences for~the two cases are presented in lables 5-1 and 5-2. Figures 5-4 through 5-14 correspond to the case with offsite power . maintained and 5-15 through 5-25 to the case with offsite power lost.-
1 Offsite. power Maintained Steam line pressure in the faulted steam line (Figure 5-4) decreases af ter the break occurs. The depressurization rate initially increases after steam line isolation occurs, since beyond this point only the f aulted steam generator is i supplying steam to the break. The depressurization rate then decreases as the -l i steam line continues to blow down towards atmospheric pressure. Steam line pressure in the intact steam line (shown on the same figure) also decreases 1 until steam line-isolation occurs. Beyond this point the intact steam generators, and therefore their associated steam lines, experience a slight pressurization. l Th? cold leg temperatures (Figure 5-5) closely follow the pressures in tho ; respective steam lines. The hot leg temperatures (Figure 5-6) follow the cold i 1 leg temperatures until the return to power occurs. A larger difference ; between the hot and cold leg temperatures develops beyond this point due to
. the core heat output.
I Core baron concentration (Figure 5-7) is zero until af ter the unborated water is purged <from the safety injection piping. Thereafter, it slowly increases as the borated safety injection water mixes with the unborated RCS inventory. The temperatures drive the core reactivity transient shown in Figure 5-8. Reactivity. initially drops to the-technical specification shutdown margin on reactor trip as the rods fall into the core. The positive eactivity inserted i due to the decreasing temperatures causes total reactivity to increase until
- prompt criticality is momentarily achieved, The fuel temperature feedback caused by the' sudden powar increase causes reactivity to decrease rapidly to l
- near 2ero. Reactivity decreases slowly as power increases due to increasing fuel temperature feedback. Reactivity decreases further with the addition of borated water from the Safety Injection System.
5-18 o
l l The neutron power transient (Figure 5-9) caused by this reactivity transient, is zero until prompi criticality occurs. At this point power spikes up and then immediately decreases sharply due to the negative Doppler feedback. Power then increases in equilibrium with reactivity until just after boron reaches the core. This is follrved by a slow decrease toward shutdown. The core heat flux (Figure 5-10) is ,,milar to the core power with two exceptions. First, there is some heat flux generated prior to prompt criticality by removal of stored energy from the fuel. Second, the power spike at prompt criticality is too brief to be reflected in the heat flux. Pressurizer level (Figure 5-11) decreases rapidly until the pressr 12er empties. It stays at zero until enough water inventory is addet oy the Safety injection System to offset the contraction of the original inventory due to the cooldown. Pressurizer pressure (Figure 5-12) decreases relatively slowly until the pressurizer empties. The decrease is more rapid until the satura-tion pressure is reached in the hottest parts of the RCS. Thereafter, pres-sure increases slowly as inventory addition f rom the Safety injection System of f sets inventory contraction f rom the cooldown. Break flow (Figure 5-13) initially decreases as the steam line pressure decreases. After steam line isolation, flow from the intact loops stops. Beyond this point flow decreases with decreasing pressure. The core mass fluxes (Figure 5-14) increase with '!me since tha reactor coolant pumps provide essentially constant volumetric flow which, with the decreasing RCS temperatures, is equivalent to an increasing mass flow rate. _0ffsite Power Lost Although Figures 5-4 through 5-14 depict the case in which offsite power is maintained, the discussion is generally applicable to Figures 5-15 through 5-25, the case in which offsite power is lost at safet/ Injection. Important exceptions are noted below. Neutron power (Figure 5-22) does not begin a sustained decrease until af ter boron from both the high-head and intermediate-head safety injection pumps has reached the core. 5-19
- _ - _ _ _ _ - _ _ _ - _ - _ - - - _ - _ - - ---------A
The core mass fluxes (Figure 5-25) decrease beyond the point at which of f site power is. lost due to the coastdown of the reactor coolant pumps. The system transient response for each case is reviewed to select the statepoint(s) for the power peaking and DNBR analysis. Values provided for each statepoint include neutron power, core heat flux, core outlet pressure, 5.4.2 Core Response 5.4.2,1 Axial and Radial power Distributions-Using the limiting statepoints from the RETRAN analyses discussed in Section 5.4.1, n ial and radial power distributions are calculated as described in Section 5.2.2.2. :The axial power distribution for the offsite power main-tained case has a top peaked shape, whereas the offsite power lost case has a bottom peaked shape. This effect is caused by the difference in moderator temperature feedback resulting from the large difference in RCS flow. Typical , values of the, mar .num axial peaking factors for the peak radial location are jfor the offsite power maintained and offsite power lost cases, respectively. Figures 5-26 and 5-27 show the asymmetric core assembly radial power distributions. The cold quadrant, which contains the stuck rod, has a more highly peaked assembly radial power distribution than the rest of the core. Typicalmaximumhotassemblypinradialpowerpeakingfactorsare[ for the offsite power maintained and offsite power lost cases, respectively. 5.4.2.2 : Minimum DNBR Results Using the limiting statepoints from the RETRAN analyses discussed in Section 5.4;1, together.with the power distributions discussed in Section 5.4.2.1, the VIPRE model is used to calculate the core local fluid properties 5-20
and MDNBR. The MONBRs predicted by the W-35 CHF correlation are greater than 1.45 for both the offsite power maintained and offsite power lost cases .Therefore, the criterion that tiie core remain in placa and intact, as discussed in Section 5.1.2, is met. Because this criterion is met, the current FSAR dose analysis, which assumes no DNBR-related fuel failures, remains valid. 5,5 Cycle Specific Evaluation The cycle-specific reload evaluation for the steam l'ie break accident focuses on the conservative core physics parameters input _to the system transient modeling. Each reload cycle is evaluated to determine whether the reactor is suberitical-at the core and system conditions corresponding to the limiting peak heat flux statepoint of the system transient. There.is a high degree of confidence that each reload core will be bounded since the system model was developed with
+
The minimum shutdown margin allowed by the technical specifications
+ A conservativa reactivity versus temperature response
- A conservative Doppler coefficient.
If the_ cycle-specific reactivity check shows the reactor to be subcritical with respect to the core assumed in the existing licensing basis analysis, including a stuck rod, then the response predicted by the system analysis bounds the reload core. If the' reload core is not subcritical at these conditions, two approaches are available to obtain acceptable steam line break analysis results: redesign the reload core, or reanalyze the transient. References 5-1 Catawba Nuclear Station Final Safety Analysis Report, 1988 Update. 5-2 Standard Review Plan, Volume III, NUREG-0800, NRC, Revision 2, July 1981. 5-21
5-3 RElRAN-02: A Program for f ransient Thermal-Hydraulic Analysis of Complex Fluid flow Systems, EPRI NP-1850-CCM, Revision 4. EPRI, November 1988. 5-4 Nuclear Physics Methodology for Reload Design, DD".-NF-2010A, Duke Power Company, June 1985. 5-5 Huclear Design Methodoingy Using CASMO-3/ SIMULATE-30, DPC-NE-1004, Duke Power Company. _ 5-6 VIPRE-01: A Thermal-Hydraulic Code for Reactor Cores, EPRI NP-2511-CCM-A, Pevision 2, EPRI, July 1985. 5-7 Thermal-Hydraulic Transient Analysis Methodology DPC-NE-3000, Revision 0, Duke Power Company, July 1987. 5-8 January 31, 1989 letter from A. 5. Thadani (NRC) to W. J. Johnson (West-inghouse), " Acceptance for Referencing of licensing Topical Report, WCAP-9226-P/9227-NP, ' Reactor Core Response to Excessive Secondary Steam l Releases". 5-22
lable 5-1 Sequence of Events for 1.4 f t' Split Breal With Offsite Power Maintained [y_en t. Time _(seconds) Break occurs 0.01 Operator manually trips reactor Pressurizer level goes offscale low 22 - i 51 actuation on low pressurizer pressure 35 iteam line isolation on low steam line pressure 36 Criticality occurs 46
$1 pumps begin to deliver unborated water to RCS 52 High-head Si lines purged of unborated water 119 One train of 51 fails Peak heat flux occurs 120 Intermediate-head 51 lines purged of unborated water 191 5-23
E i I l 1 Table 5-2 Sequer.ce of Events for 1.4 f t* Split Dreak t With Offsite power Lost at $1 Actuation ' l l Event Time (seconds) Dreak occurs '0.01 i Operator annually' trips reactor pressurizer level goes offscale low 22 SI actuation on low pressurfrer pressure 35 Offsite power lost , Reactor 1,oolant pumps begin to coast down Steam-line isolation on low steam line pressure 36 i Criticality occurs 52
'S! pumps begin_to deliver unborated water to RCS 66-High-head $1_ lines purged of unborated water 134 One train of 51 fails t
Intermediate-head Si lines purged-of unborated wster 223 Peat heat flux occurs 228 d 5-24 '
s Figure 5-1 RETRAN Reactor Vessel Model he b 4
+
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Figure 5-2 K-effective versus Moderator '"emperature 1.07 1.0C *
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Figure 5-26 Typ1 Cal Radial Powe:- Distribution Offstte Power Available, Peak Heat Flux M D
- A N L K J H C- F E D C 8 A i 0 II O 11 0 17 0 20 0 20 021 0 15 2 0 11 0 07 0.16 0 10 0.19 0.12 0.22 0.13 0.24 0.11 0.20 -
3 0.11 0 26 0 18 0 14 0 31 0 15 0 19 0 19 0.46 0 23 0.31 0 48 0.21 4 0.07 0.18 0 19 0.50 0.47 054 0.26 0.73 0.72 0 84 0.34 0.34 0.14 5 0.11 0 16 0.14 0 50 0.55 0.73 0 60 0.77 0.80 1.17 1.02 1 00 C 30 0.41 0.32 6 0.17 0 10 0.31 0 47 0.73 0.37 0.64 0.41 0.91 0.68 1.58 1.07 0.74 0.28 0 57 7 0.17 0.19 0.15 0 54 0 60 0 64 0.87 1.24 1.32 1.33 157 1.44 0 43 0.68 C 69 8 0.20 0.12 0.19 0.26 0.77 0 di 1.24 0.81 2.36 l .T / - 83 1.01 0.78 0.56 1 03 1
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- 0.67 13 0.20 0.48 0.34 0.30 0.74 0.43 0.78 1.00 2.59 1.22 1.74 2.75 1.23 14 0.21 0.14 0,41 0.28 0.68 0.56 1.32 0.00 1.52 0.67 1.23 15 0.32 0.57 0.69 1.03 1.23 1.43 1.00 P 1
- Stuck Rod Locatii k J 5-50
l' l l l Figure 5-27 Typwal Raatal Power Distribut10n Of f stte Power Lost, Peak Heat F1" A D N M L K J .-i 0 F 0 C D A 1 0.26 0 43 0.41 0 40 0.42 0.45 0.27 2 0.20 0.18 0 40 0 29 0 45 0.25 0.47 0.30 0 43 0.20 0 22 3 0.20 0.47 052 0.35 0 68 0.35 0 39 0.36 0.73 0 39 0 57 0.52 0 22 4 0.18 C 52 0 39 0 87 0 93 1.13 0.50 1 20 1.01 0 97 0.44 0.59 0 21 5 0.26 0.40 0.35 0 87 1 06 1 45 1.30 1.36 1.39 1.6 I I.23 1.03 0 43 0.50 0.34 6 0.43 0.29 0.68 0.93 1,45 0.70 1.43 0.71 1.60 0.85 182 1.18 0 86 0.38 0.59 7 0.41 0.45 0.35 1.13 1.30 1.43 1.32 1.36 1.59 1.88 1.83 1.56 0.49 0.66 0.62 8 0 40 ' .25 0.39 0.50 1.36 0.7 I l.36 0.98 1.08 1.27 2.44 0.93 0,73 0 47 0.74 9 0.42 0.47 0.36 1.20 1.39 1.60 1.59 1.98 2.82 3.82 312 2.74 0.84 1.06 0.94 r 1 10 0 45 0.30 0.73 1 01 1.61 0.85 1.88 1.27 3.82 3 63 4 25 2.51 1.78 . 0.74 1.08 L A - 1I 0.27 0.43 0.39 0.97 1.23 1.82 1.83 2.44 3 22 4 25 3.01 2.36 0,96 1.07 0 69 12 0.20 0.57 0.44 1.03 1.18 1.56 0.93 2.74 2.51 2.36 1.07 1.40 0.50 , 13 0.22 0.52 0.59 0.43 0.86 0,49 0.73 0.84 1.78 0 96 1.40 1.27 0.55 14 0.22 0.21 0.50 0.38 0.66 0.47 1 06 0.74 1.07 0.50 0.55 15 0.34 0.59 0.62 0.74 0.94 1.08 0.69 f 1
- Stuck Rod Location k M 5-51
'6.0 DROPPED ROD ANALYSIS 6.1 Overview -6.1.1 Description of Dropped Rod Accident The dropped rod ace.ident is described in FSAR Section 15.4.3 (Reference 6-1). The scenarios of concern consist of all single and multiple dropped control rods for rods orioi-' ting in the same group. Beginning from a full power initial condition ( sower power levels are less limiting), one or more rods drop into the core and cause a prompt reduction in reactor powcr. The Rod Control System, in the automatic control mode, detects a mismatch between reactor and turbine power and responds by withdrawing the controlling rod group, Control-Bank D. With the Rod Control System in manual, Control Bank D does not withdraw, and the reactcr power decreases to a new equilibrium power level. lhe power mismatch also results in a reduction in the average core moderator temperature, which typically adds positive reactivity due to the presence of a negative moderator temperature coefficient (MTC). The combination of rod withdrawal and decreasing temperature can cause the reactor to return to full power and even exceed the initial power level. Since the core power peaking is increased by the dropped rod (s), the potential exists for the DNBR limit to be approached. 6.1.2 Acceptance Criteria The dropped rod accident is classified as an-ANS Condition II event., an anticipated transient, Therefore, it must be demonstrated that the DNBR limit is not exceeded. The other Condition II criteria regarding overpretsurization or propagation to a Condition III event are not challenged by a dropped rod transient. 6.1.1 Analytical Approach The dropped rod accident requires a large set of physics parameters to be determined for use as initial and boundary conditions. These parameters are input to a RETRAN-02-(Reference 6-2) McGuire/ Catawba model (Reference 6-3) for 6-1
. . - -. . . . ~ . - _ . -.
1 the system thermal-hydraulic analysis. The RETRAN analysis generates the core
- statepoint con ('tions which correspond to the transient time of minimum DNBR.
EPRI-NODE-P (Reference _6-4) or SIMULATE-3P (Reference 6-5) is used to generate power distributions corresponding to the possible dropped rod combinations. The power peaking analysis uses either the pre-drop or the post-drop thermal boundary conditions. The core power distribution along with the core thermal-hydraulic boundary conditions from the RETRAN analysis are then input to a VIPRE-01 (Reference 6-6) McGuire/ Catawba model (Reference 6-3) to calculate the minimum DNBR. 6.2 Simulation C, odes and Models 6.2.1 System Thermal-Hydraulic Analysis The McGuire/ Catawba RETRAN model described in Section 3.2 of DPC-NE-3000 (Reference 6-2) is used for the dropped rod analysis. A one-loop mo~ ' is sufficient sirce little loop asymmetry develops during this transient. A Catawba Unit I model 'is selected due to the higher primary system T-ave used in the core thermal-hydraulic analysis. There are no differences between the McGuire and Catawba units which are significant in the context of a dropped rod transient. L 6. 2 .' 2 Nuclear Analysis l
. The dropped rod transient is modeled using EPRI-HODE-P to predict three--
dimensional power distributions and_ core reactivity. The analysis is based on L several cycles at various burnups. Each core analyzed contains 193 Westing-I house optimized fuel assemblies. However, the behavior of the important physics parameters and the bounding values selected for this analysis would not_ change for cores containing Mk-BW fuel. 6-2
6.2.3 Core Thermal-Hydraulic Analysis Methodology The VIPRE-01 code is used for the dropped rod cora thermal-hydraulic analyses. VIPRE thermal-hydraulic boundary conditions are obtained from the RETRAN system transient simulation. RETRAN predicts core inlet flow, inlet tempera-ture, outlet pressure, and heat flux for a set of cases based on the dropped rod worth and limiting burnup condition. A core neutronics simulation code provides the usial shape and radial power distributions. A
)VIPRE _
model is used to calculate the limiting statepoint local coolart properties and DNBR. The critical heat flux (CHF) correlation used to evaluate DNBR is the 9&W BWCMV CHF correlation (Reference 6-7). The C "9E analysis employs the statistical core design DNBR limit o f 1. 55 ( Re f arance 6-8). Model Descriptic, 6.3 Transient Analysis 6.3.1 Initial Conditions The VIPRE evaluation of the minimum DNBR resulting from the dropped rod transient uses the statistical core design (SCD) methodology. Consequently, the following RETRAN initial conditions are specified as nominal values since the uncertainty is factored into the SCD design limit. 6-3
Power level = 100% FP
- DCS flow = 382,000 gpm
- Pressurizer pressure = 2235 psig
- RCS T-ave = 590.8 F
- Core bypass flow = 7.5%
A discussion of the non-SCD parameters and the basis for selecting their initial condition values follows. Pressurizer level A low initial pressurizer level reduces the initial core outlet pressure and minimizes the transient pressure response, which is conservative for DNBR. The full power programmed pressurizer level for the Catawba Unit 1 model is 60%. The initial condition uncertainty allowance for reduced pressurizer level is 9%. Therefore, the initial pressurizer level for this analysis is 51%. SG NR Level A low initial steam generator narrow range level minimizes the initial steam generator inventory. Catawba Unit I has model D3 steam generators that have a programmed level that varies with reactor power. A low initial level serves to maximize effects due to changes in feedwater flow. This parameter has no significant impact on the results of the dropped rod transient. The full power programmed steam generator narrow range level is 66.5% The initial condition uncertainty allowance is 8%. Therefore, the initial steam genera-tor level for this analysis is 58.5%. Averese Fuel Temperature Maximum average fuel temperatures for an equilibrium 390 EFPD fuel cycle with Mark-BW fuel at BOC, MOC and EOC conditions are used. The average fuel temperatures used are at BOC, MOC and EOC, respectively. 6-4 l l {
l SG Tube Plugging Assuming no steam generator tube plugging maximizes the initial steam line pressure and results in a more limiting transient. Therefore, no tube plugging is assumed for this analysis. This parameter has no significant impact on the results of the dropped rod transient. 6.3.2 Boundary Conditions 6.3.2.1 Physics parameters The important physics parameters required by the RETRAN and VIpRE models for the dropped rod ana_1ysis are discussed below. These parameters are evaluated for the dropped rod scenarios over a range of conditions to ensure that the selected values bound current and future reload designs. The RETRAN analysis uses values for each of these parameters that are consistent in terms of a beginning, middle, or end-of-cycle condition. RETRAN VIPRE
- Dropped rod worth X Control Bank D worth X Core tilt following rod drop X Moderator temperature coefficient X
- Doppler temperature coefficient X
- Ef fective delayed neutron fraction X
- Radial peaking factor X
- Axial peaking factor X Dropped Rod Worth The dropped rod worth ranges up to ]pcm. This worth exceeds the worth of all possible combinations of dropped rods from the same rod group.
6-5
Control Bank D Worth Control Bank D worth ranges from pcm at the rod insertion limit as
- J a function of burnup.
FAH Versus Worth Power peaking increases in those areas of the core opposite the dropped rod (s) and is generally greatest for those cases in which three rods are dropped. The effect of a dropped rod on FAH is a function of burnup, dropped rod worth, and the number of dropped rods. Enveloping FAH responses derived from assem-bly average power are presented in Figures 6-2 to 6-4. Axial Shape A bounding axial shape at each burnup condition, based on a top peaked power distribution, is chosen for the thermal-hydraulic analysis. Core Tilt Followino Rod Drop Fifty-three full-length control rods of two designs, Ag-In-Cd and boron
. carbide (B.C), are analyzed. All combinations of rods in each group of the control banks and shutdown banks are dropped into the core from the rod insertion limit (RIL) and the all-rods-out (ARO) position to determine which dropped rod cases would result in the worst excore tilts and power peaking.
Figure 6-5 illustrates the effect of increasing dropped rod worth on the induced tilt. In general, a set of three dropped rods results in the most severe combination of excore tilts and power peaking. The reactor response to a dropped rod transient depends on the core tilt-detected by the excore detectors since the excore detector signal is an input to the Rod Control System. Dropped rods cause the power level to decrease in the vicinity of the dropped rod and to increase in areas away from the dropped rod. The excore tilt for.the four quadrants is modeled by using the assemblies closest to the excore detector to generate'a det -tor response. 6-6
l l Moderat,or Temperature Coefficient I Several fuel cycles were reviewed to determine realistic but conservative moderator temperature coefficients (MTCs) to use in the dropped rod analysis. Conservative slopes were chosen for the MTCs versus moderator temperature at several burnup statepoints and boron concentrations. An MTC was conserva-tively chosen at the burnup ttatepoints analyzed for the HFP, nominal condi-tion moderator temperature. The HFP moderator temperature and the slope of the MTC versus moderator temperature curve are used to determine MTCs at various power levels occurring during the dropped rod transient. The MTC assumed is a least-negative or most positive value depending on the core burnup and moderator temperature. This assumption minimizes the negative reactivity feedback that is available when the post-drop power level increases as Control Bank 0 is withdrawn. Dggpler Temperature Coefficient The Doppler temperature coefficient is selected as a least-negative value. This assumption minimizes the negative reactivity feedback that is available when the post-drop power level increases as Control Bank 0 is withdrawn. Effective Delayed Neutron Fraction The effective delayed neutron fraction is not a very important parameter since the transient response near the DNBR statepoint is slow. A minimum value of beta-effective is used. 6.3.2.2 Reactor Protection System The RETRAN analysis takes credit for a reactor trip only on low pressurizer pressure. The trip setpoint is reached only for higher dropped rod worth cases at beginning of-cycle. The overtemperature and overpower AT trip setpoints may be reached for some dropped rod events. No credit is taken for the negative flux rate trip function, low-low steam generator level trip or main steam isolation on low steam line pressure. 6-7
6.3.2.3 Power Range Nuclear Instrumentation
~~
6'3.2.4 Rod Control Systen
'The Rod Control System is explicitly modeled in the RETRAN analysis. The controller uses a power mismatch signal and a temperature error signal to determine-Control Bank 0 Insertion or withdrawal and rod speed. The power
-mismatch signal is a difference between turbine power (impulse pressure) and the auctioneered-high NI flux indication. The temperature error signal is a difference between a reference temperature based on turbine power and the auctioneered-high primary loop T-ave indication. -Due to the importance of Control . Bank.0 withdrawal on the dropped rod analysis, the worst case single failure has been determined to result in the Nl flux indication auctioneering low, This failure causes the maximum post-drop power levels by accelerating the onset and increasing _the rate of Control Bant 0 withdrawal.
6.3.2.5-- pressurizer Pressure and Level Control Since the' dropped rod transient is a DNB transient, pressurizer sprays and the. pressurizer p0RV are assumed to function' in order to minimize primary pres-sure. Pressurizer heaters are assumed not to function. -Pressurizer level control =is assumed to be in manual, and is not important for this transient. 6-8
l 6.3.2.6 Main Feedwater and Turbine Control 6.4 Results and Conclusions 6.4.1 System Transient Results The dropped rod event is analyzed at beginning, middle, and end-of-cycle conditions. The typical transient re'sponse at baginning and en'd-of-cycle conditions are discussed in detail below, and the results of the bounding cases are presented. 6.4.1.1. Typical Beginning-of-Cycle Response 100 pcm Case Reactnr power (Figure 6-6) initially decreases rapidly in response to the dropped rod. Then power recovers as the Rod Control System withdraws Control Bank 0 and power overshoots the initial level. Raactor power reaches a maximum value of 115.3% and starts to decrease again before the Rod Control System terminates rod withdrawal. Also shown on Figure 6-6 is the NI signal that is input to the Rod Control System. While core power reaches a maximum of 115.3%, the Ni signal is only about 72%. Control Bank D position is shown in Figure 6-7. Bank D motion results from the combination of power mismatch and temperature error signals. Average loop temperature (T-ave) as shown in Figure 6-8 initially decreases about 2 F and then increases about 7 F. The T-ave response is determined by the balance between core power and the steam load. Pressu,rizer level, shown in Figure 6-9, responds to the change in T-ave. The pressurizer pressure response (Figure 6-10) is dictated by changes 6-9
~ . -in pressurizer level and by the actuation of pressure mitigation equipment. -Pressure initially decreases due to the impact of the dropped rod on reactor power, and then increases with power. Pressurizer sprays and PORVs actuate to minimize primary pressure, which is conservative with respect to ONBR. The PORVs continue to cycle until T-ave begins to decrease. The limiting statepoint occurs at the time of maximum heat flux. }JLqh_er Worth Cases As the dropped rod worth is increased above 100 pcm, the initial reduction in core power is larger, and consequently the initial reductions in T-ave and pressurizer pressure are larger. For beginning-of-cycle (B0C) cases at or above[ pcm, a reactor trip occurs on-low-l'w pressurizer pressure and DNB is not a-concern. In general, the transient responses to dropped rod worth cases at }ptmhavelowerstatepointvaluesforcorepower, pressurizer pressure, and T-ave. The dif ference ir, statepoint conditim s mainly results-from the response of the Rod Control System to the comoined power mismatch and temperature error signals.
6.4.1.2 Typical End-of-Cycle Response 400 ocrr Case Reactor power (Figure 6-11) initially decreases rapidly in response to the dropped rod, then increases due to the negative MTC and to the Rod Control System. withdrawing Control Bank D. Reactor power recovers, overshoots its initial value reaching a maximum value of 109.8%, and starts to decrease before the Rod Control System terminates rod withdrawal. Also shown on Figure 6-11 is the NI signal input to the Red Control System. Control Bank 0 ! position is shown on Figure 6-12. T-ave (Figure 6-13) initially decreases approximately 7'F, then increases about 3'F to the point of maximum heat flux. H T-ave. decreases until core power exceeds the steam load. Pressurizer level l l (Figure 6-14) also follows the trend of T-ave, reaching a minimum of about 42% at 20 seconds. Pressurizer pressure (Figure 6-15) initially decreases with the drop in T-ave and reaches a minimum of approximately 2150 psig before increasing to a maximum of about 2280 psig. Pressurizer spray actuates at 6-10 l
l 1 aoproximately 47 seconds and continues for the remainder of the simulation. The limiting statepoint occurs at the time of maximum heat flux. Lower Worth Cases For dropped rod worths lower than 400 pcm, the power overshoot is slightly higher. The initial power reduction is less, and therefore, the reactivity addition due to moderator feedback and Baitk D withdrawal has less of a deficit to offset. As in the BOC cases, the most important boundary condition is the response of the Rod Control System. I litcher Worth Cases As the Popped rod worth increases above 400 pcm, the power overshoot de-creases in magnitude. The most significant change with increasing worth is that pressurizer pressure is significantly lower at the statepoint. The substantial contraction of the primary coolant immediately following the rod drop, and the associated depressurization, have not resulted in pressurizer level and pressure recovering to the initial values at the statepoint time. 6.4.1.3 Limiting Statepoint Selection Cases Analyzed Dropped rod cases are analyzed at beginning, middle, and end-of-cycle (EOC) conditions. The dropped rod worth for cases analyzed at a given time in tycle is in increments of 100 pcm untti a reactor trip on low pressurizer pressure occurs. Reactor trips on low pressurizer pressure occur in the[ }pcm beginningof-cyclecaseand[' " pcm middle-of-cycle (MOC) case. The maximum dropped rod worth analyzed is 'pcm.
- a. .
Beginnino-of-Cycle Cases The trends of the key results for the spectrum of BOC cases are shown in Figures 6-16 through 6-18. The peak core power (Figure 6-16) decreases with increasing dropped rod worth. T-ave (Figures 6-17) decreases with increasing 6-11
rod-worth, in the )pcm case, pressurizer pressure (Figure 6-18) at the statepoint does not follow'the trend of decreasing pressure with increasing dropped rod worth. During this case, reactor power continues to increase for several minutes after Bank D is fully withdrawn due to the positive MTC-applied. During this relatively long transient pressurizer pressure recovers to above its initial value. JHdle-of-Cycle _ Cases
~
The trends of the key results for the spectrum of MOC cases are shown in Figures 6-19 through 6-21. The peak core power (Figure 6-19) decreases with increasing dropped rod worth. T-ave (Figure 6-20) also decreases with de-creasing rod worth, -Above pcm,- pressurizer pressure (Figure 6-21) at the peak power statepoint does'not recover to the initial value. End-of-Cycle Cases The trends of the key results for the spectrum of end-of-cycle cases are shown in Figures 6-22 through_6-24. The peak core power (Figure 6-22) decreases with increasing dropped rod worth. T-ave (Figure 6-23) also decreases with decreasing rod worth. Above pcm, pressurizer pressure (Figure 6-24) at the peak power statepoint does not recover to the initial value. Limitinq_ Cases For each dropped rod worth the RETRAN analysis results were compared and the limiting burnup condition was determined. The selection of the limiting case is based on the product of the peak core power and the associated radial peaking factor for that dropped rod worth and burnup. The burnup with the I largest value of this' parameter is then evaluated with respect to the other important DNB parameters (pressure, temperature, flow) to confirm the limiting burnup condition. -The limiting burnup condition for each dropped rod worth is then analyzed with VIPRE_to determine the minimum DNBR. The limiting cases were determined to be as- follows: ; l- 6-12 l { 1 1
l Dropped Rcd Burnup Worth (pcm) Condition
~
Figures 6-25 through 6-27 show the trends of neutron power, T-=ve, and pressurizer pressure for the limiting analyses, in all cases the limiting statepoint occurred at the time of peak core powor. 6.4.2 Core Response 3 6.4.2.1 Statepoint Conditions RETP.AN results yield the following statepoint conditions f or the limiting cases analyzed. Dropped Heat Core Core Outlet Core _ Rod Flux Inlet Flow Pressure Case Worth (pcm) (MBTV/hr.f 8t ) Temperature (oFj QiLBM/hr.ft2 ) (psia) C eu l 6-13
6.4.2.2 Thermal-Hydraulic Results for DNBR Tne dropped rod event is analyzed utilizing the VIPRE-01 code and a SCD DNBR limit of 1.55. The VIPRE model DNBR rasults for the limiting cases with the given axial shapes (Figure 6-28) and radial power peaking (Figures 6-2 throuah 6-4) are greater than 1.55 for each case. 6.5 Cycle Sp_ecific Evaluation The reference dropped rod analysis is verified to be bound'ng by comparison of several cycle-specific physics parameters against values assumed in the reference analysis. Physics parameters that are checked for each reload core are:
- Initial FAH
+ Axial Flux Shape Moderator and Doppler Temperature Coefficients
- Maximum Dropped Rod Worth Available Control Bank Worth for Withdrawal Several reload cycles were analyzed in order to determine bounding inputs for the reference dropped rod analysis. The results of these analyses established
' bounding curves, versus the number and worth of dropped rods, defining both the' increase in radial peaking and limiting excore detector responses. Thasa inputs are considered independent of the reload core design and will not be checked on a cycle-specific basis. While the above physics parameters are not expected to change for e reload core, they are checked to ensure that the reference dropped rod analysis remains valid. For each reload core, the maximum core FAH for the pre-dropped condition is verified to be less than L.I lfor allowed rod insertions. Moderator and doppler temperature coefficients are also verified to be I conservative by comparison against the coefficients used in the reference analysis. The axial shapes assumed in the reference analysis will be checked for_all dropped rod combinations. The maximum allowable dropped rod worth is 6-14
r verified to be less than the maximum worth analyzed ( i pcm). Also, the available control bank worth for withdrawal is verified to be less than the values assumed in the reference analysis, in conclusion, the reference dropped rod analysis is applicable to the reload core if the cycle-specific physics parameters are determined to conservatively bound the values assumed in the reference analysis. In the unlikely event that any of _the reference analysis input physics parameters do not bound a _given reload design, there are several recourses. The reload core can be redesigned, the dropped rod analysis can be reevaluated with cycle-specific inputs, or a new reference analysis can be performed with updated limiting values. References 6-1 Catawba Nuclear Station Final Safety Analysis Report, 1988 Update. 6-2_ RETRAN-02: A Program for Transient Thermal-Hydraulic Analysis of Complex Fluid Flow Systems, EPRI NP-1850-CCM, Revision 4, EPRI, November 1988. 6-3 Thermal-Hydraulic Transient Analysis Methodology, DPC-NE-3000, Revision 0, Duke Power Company, July 1987. 6-4 Nuclear Physics Methodology for Reload Design, DPC-NF-2010A, Duke Power Company, June 1985. 6-5 Nuclear Design Methodology Using CASN3-3/ SIMULATE-3P, DPC-NE-1004, Duke Power Company. 6-6 VIPRE-01: A Thermal-Hydraulic Code for Reactor Cores, EPRI NP-2511-CCM-A, Revision 2, EPRI, July 1985. 6-7 BWCMV Correlation of Critical Heat' Flux in Mixing Vane Grid Full-Assemblies, BAW-10159-P, May 1986. 6-8_ Core Thermal-Hydraulic Methodology-Using VIPRE-01, DPC-NE-2004, Revision 0, Duke Power Company, December 1988. 6-15
.F lo.ure 6- 1.,
VIPRE Model L - f
'O 6-16
Figure 6-2 ! F-Delta-H versus Dropped Rod North Beginning of Cycle i 3 d.
~
Dropped Rod Worth (PCM) 6-17
Figure 6-3 F-Delta-H versus Dropped had Worth Middle of Cycle. Io c3 a. Ml Dropped Rod Worth (PCM) I 6-18 I
1 l l Figure 6-4 F-Delta-H versus Dropped Rod Worth End of Cycle n 8 e. Dropped Rod Worth (PCM) 6-19
4 1 Figure 6-5 i t
~
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Appendix A REACTOR VESSEL THERMAL MIXING EVALVATION Part A: Forced Circulation Mixino A.1 Backcround A thermal mixing test-was performed at McGuira Nuclear Station Unit 2 to
. determine the degree of coolant mixing.in the reactor vessel during
.q A-1
k A.3 Instrumentation / Equipment C;nfiauration The instrumentation used in the therm 31 mixing test consisted of the wide range hot and cold leg resistance temperatura detectors (RTDs) and selected core exit thermocouples (CETCs). rach hot leg or cold lag is equipped with one wide range, thermowell+ mounted RTD. The loop B hot leg thermowell is located upstream of the surge line and the loop A cold leg RTD ii located upstream of the normal charging such that pressurizer outsurges and charging do not directly impinge on these RTDs and possibly adversely affect the RTD temperature response. The response time of the wide range RTDs is estimated to be approximately 20 seconds. A total of 26 CETCs were used during the teu to obtain core exit temperature measurements. The CETCs utilizad are fairly evenly distrib-utad, with approximately 7 CETCs per core quadrant. The positions of the CETCs with respect to core locations and the orientation of the loops with respect to the core are shown in Figure A-1. The response time of the CCTCs is relatively f ast; it is less than one second based on avail-able references. A.4 Data Acquisition Wide range hot leg and cold leg RTD data was recorded using the Operator Aid Computer (OAC) transient mnnitor at a one second frequency and on the OAC general program at a five second frequency. CETC data was recorded using the OAC general program at the five sacond f raquency. A.5 -Definitions 1 NI A-2 i
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Part B: Natural Circulatten Mixing A.10 Rach reund The two transients on which the natural circulation reactor vessel , thermal mixing estimate is based are the September 1, 1981, McGuire Unit I natural circulation startup test and the January 29, 1985, Catawba Unit 1 station blackout startup test. A.11 Apy n ach in order to quantify reactor vessel thermal mixing in the abse re of , forced circulation, data analysis techniques similar to those in Part A are used, but with the following exceptions: Since the natural circulation data is from plant events not designed to measure mixing, there are data limitations including the lack of frequent core exit thermocouple data.
- Since the RCS loop flow rates are relatively low, the flow measure-ment devices are not useful for determining the magnitude. There-fore, the flow rate can only be estimated. This makes transit times only approximate.
A.12 Instrumentation / Equipment Configuration The discussion in Part A is applicable for natural circulation, except as noted above concerning core exit thermocouple data. A.13 Acquisition Wide range hot and cold leg temperature data was recorded using the OAC-transient monitor at a one second frequency for the Catawba event and a one minute frequency for the McGLtre event. A-8
l A.14 Definitions i The discussions in Part A are applicable for natural circulation. wem 4 m h' h' A-9 i.
l l l
=
m W N A-10
Figure A-1 - Mixing Test Thermocouple Locations
& CWitt COMMT R P N H L K J B ,G T E D C B A
= .
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i Figure A-2 to A-7 i r i
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-i t
( i l A-12 to.A-17 p.. i s.
DPC NE 3001 Docketed Correspondence
+ January 29.1990 - Subinitial Letter e Septernber 14, 1990 - Ruponse to NRC Question
- February 13.1991 Response to NRC Questions e June 3,1991 - Response to NRC Questions
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; j JUKE POWER January 29, 1990 U. S. Nuclear Regulatory Commission Washington. D.C. 20555 Attention: Document Coatrol Desk
Subject:
McGuire Nuclear Station Docket Nos. 50-369. 50-370 Catawba Nuclear Station Docket Nos. 50-413, 50-414 Safety Analysis Physics Parameters and Multidimensional Reactor Transients Methodology, DPC NE-3001-P Gentlumen: - Please find enclosed for your review fifteen copies of the proprietary Topical Report DPC-NE-3001-P. " Safety Analysis Physics Parameters and Multidimensional Reactor Transients Methodology." This report describes the Duke Power Company methodologies fort (i) simulating the FSAR Chapter 15 events characterized by multidimensional reactor transients, and 11) systematically confirming that reload physics parameters important to Chapter 15 transients and~ accidents are bounded by values assumed in the licensing analyses. . The multidimensional reactor transients described are the rod ejection accident, the main steam line break, and the dropped rod transient. Also, included are four copies of three reports describing the EPRI computer code ARROTTA. This code is used in the analysis of the rod ejection accident. The reports are ARROTTA: Advances Rapid Reactor Operational Transient Analysis
- Computer Code. Computer Coda Documentation Package Thuory Manual.
ARROTTA Validation and Verification - Standard Benchmark Set. EPRI Research Project 1936-6, July 1989. Prepared by S. Levy, Inc. ARROTTA-HERMITE Code Comparison, EPRI NP-6614, December 1989.
~
The first report describes the theory incorporated'into the computer software. The second.provides validation and verification of the computer software to various numerical benchmarks. The third report provides comparison of a rod ejection transient analyzed using ARROTTA to one analyzed using HERMITE which has already been reviewed and approved by the NRC. .These reports are included to facilitate review of the rod ejection section.
. _ _- _ _ .- = _ _ _ _ _ _ _ _ _ _ ~ _ _ _ _ . - - . . - _ _ _ . _ _ _ _ _ _
U. 5. Nuclear Regulatory Commissten p
,lanuary 29. 1990 Page 2 The objectives of this report are to document the methods to be used to verify that-the key physics paraceters calculated for a reload core are bounded by values assumed in the Licensing Chapter 15 analyses, and to describe the methods used to analyze three complex rSAR Chapter 15 accidents. Rasults of the multidimensional reactor transient analyses are presented to demonstrate the methods and to serve as substitute FSAR analyses.
In accordance with 10CFR 2.790, Duke Fcwer Company requests that this report be considered proprietary; Information supporting this request is included in the attached affidavit. A non proprietary version will be submitted following receipt of the Safety Evaluation Report. If you have any questions, or require more information. please call Scott Gewehr at ( 704) 373-7581. t Very truly yours, c -
.?;;;= l h A f $ f
/
Hal B. Tucker SAG 206/lcs ! xc: (w/o Attachments) Mr. Darl S. Hood Project Manager Office of Nuclear Reactor Regulation U. S. Nucinar Regulatory Commission Washington D.C. 20555 Dr. Kahtan Jr '
, Project Manager Offico of Nu. teactor Regulation U. S. Nuclear alatory Commission Washington, D.C. 20555 Mr. S. D. Ebneter. Regional Administrator U. S. Nuclear Regulatory Commission 101 Marietta Street, NW. Suite 2900 j Atlanta, Georgia 30323
( l Mr. Robert C. Jones, Acting Branch Chief L Reactor Systems Branch. l Office of Nuclear Reactor Regulation ! U. S. Nuclear Regulatory Commission Washington, D.C. 20555 l l m._ ._ . _ _ _ _ _ _ _ . _ . . - - - - - -- -- - - - -
ll __
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Ouke Pourt Comoacy Hn B h w PO Bat 331H b oce Presternt Chancar, h C l8212 %citar PmJuctm 1:04133 43J1 DUKEPOWER September 14, 1990 U. S . Nt
- m i agulatory Commission ATTiis D e ..4t Control Desk Washington, D.C. 20555
Subject:
McGuire Nuclear Station Docket humbers 50-369 and -370 Catawba Nuclear Station Docket Numbers 50-413 and -414 Topical Report DPC-NE-3001 By letter dated January 29, 1990, Duke submitted the subject Topical Report y
" Safety Analysis Physics Parameters and Multidimensional Reactor Transients Methodology" for review.
During a July 23, 1990 meeting between representatives from Duke, NRC Staff, and Brookhave National Laboratory, a question was raised by the staff regarding the stope of review required. Attached is the response to that question. Please note that a portion of the response is proprietary, and should be withheld from public disclosure. Included with the January 29, 1990 letter is an affidavit which supports the proprietary designation. If there are any quer.ti'n.., please call Scott Gewehr at (704) 373-7581. Very truly yours, 1 - o h - Ib ' l k ,q Hal B. Tucker SAG /231/lcs
U. S. Nuclear Rsguletory. Commission September 14. 1990 Page 2 xc Mr. Tim' Reed, Project Manager Office of Nuclear Reactor Regulat1 n9 U. S. Nuclear Regulatory Commission Washington, D.C. 20555 Dr. Kahtan Jabbour. Project Manager Office of Nuclear Reactor Regulation U. S. Nuclear Regulatory Commission Washington, D.C. 20555 Mr. S. D. Ebneter, Regional Administrator
-U. S. Nuclear Regulatory Commission 101 Marietta Street, NW, Suite 2900 Atlanta, Georgia 30323 Mr. Robert C. Jones Reactor Systems Branch Office of. Nuclear Reactor Regulation U. S. Nuclear Regulatory Commission Washington, D.C. 20555 e
a 4 9 l l I i 1. l
Q. Are there parts of ARROTTA that do not need to be reviewed? A. Yes. Duke Power uses ARROTTA only for the rod ejection transient which is a very rapid transient. Therefore, the ability to model fission product poisoning (iodine. xenon, promethium and samarium) does not need g to be reviewed. Also due to the rapidity of the event moderator feedback effects are not u ry important. Therefore ARROTTA's ability to model moderator feedback does not need to be reviewed. 7' ARROTTA's neutronic theory is identical to QUANDRY's, however the numerical solution technique is dif farent. ARROTTA also utilizes the same thermal hydraulics routines found in BEAGL. The material $gik? properties have been modified to match those of RETRAN and the steam q.{; properties enhanced to allow supercritical properties. Since Doppler , ;i ; ' feedback terminates the power excursion, the fuel pin thermal model is g important, however, the fluid modeling has a secondary ef fect. Because the input cross sections have been adjusted in many ways to make e the licensing model liciting { the cross sections and ARROTTA's ability to match measurements is of
].thesourceof minor i=portance.
_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ -- - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - -- ~-
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,, .3 i DUKE POWER d ~ ? C' G r au m.0 :.3 February 13, 1991 U. S. Nuclear Regulatory Commission ATTN: Document control Desk Washington, D.C. 20555
Subject:
McGuire Nuclear Station Docket Numbers 50-369 and -370 Catawba Nuclear Station Docket Numbers 50-413 and -414 Responses to Questions on Topical Report DPC-NE-3001 On January 29, 1990, Duke submitted the subject Topical Report, " Safety Analysis Physics Parameters and Multidimensional Reactor Transients." By letter dated December 24, 1990, the NRC staff prcvided questions regarding the subject Topical Report. Attached are formal responses to the staff's questions. If there are any further questions, please call Scott Gewehr at (704) 373-7581. Very truly yours, e ' % M. S. Tuckman SAG /252/lcs xc: (W/ Attachments) Mr. T. A. Reed, Project Manager Office of Nuclear Reactor Regulation U. L. Nuclear Regulatory Comission Washington, D.C. 20555 Mr. R. E. liartin, Project Manager Office of Nuclear Reactor Regulation U. S. Nuclear Regulatory Comission Washington, D.C. 20555 Mr. S. D. Ebneter, Regional Administrator U. S. Nuclear Regulatory Comission 101 Marietta Street, NW, Suite 2900 Atlanta, Georgia 30323 Mr. Robert C. Jones, Acting Branch Chief Reactor Systems Branch Office of Nuclear Reactor Regulation U. S. Nuclear Regulatory Comission Washington, D.C. 20555 s
r , I (-
-l 1
- 1. k*ill tne DPC netnocs ce appliec to cores includinc fuel from cultiple fuel vencors ? ;f so, j ustify the use of '.'IPRIOl, ARROTTA and !
RETRAN-02 anc tne selectec- options / data f or ttis application? Re sponse: i The DPC netnocs cescribed in DPC-NE-3001 will be applied to reloac
. cores wnich may include fuel from dif ferent :uel vencors. Due t: tne relative si=11arity of current PWR fuel designs, it is anticipated that the DPC nethods vill re=ain valid. At present, the Westincncuse opti=1:ec fuel assecoly (OFA) design and the 35W Mark-BW fuel design have been analy:ec. .The neutronic dif ferences in these fuel desitns are acco=mocatec by determining _ values of saf ety analysis physics parameters that conservatively bound both fuel types, or by explicitly .I analvring a specific reload design. Fuel design data input to tne VIPRE-01, ARROTTA, and RETRAN-02 codes are selected to be consistent with one of these approaches. The VIPRE-01 and RETRAN-02 analyses
=ocel botn fuel types to ensure that the icoact of fuel desicn cata is i exolicativ calculatec. In cne ARROTTA rod ejection analysis. ene r e la t ively snail neutronic dif ferences between 0FA and Mark-BW fuel designs are insignificant when compared to the- conservative adjustments race to the cross sections in order to model a highly peaked ccre with bounding physics parameters. In addition, the selection of coce opcions is not af fected by a dif ferent fuel design.
The fuel assen:1y design data employed in the analyses vill be , consistent with the fuel types comprising the reload core. I i
f Discuss the conclusi:n that a VIPRE-01 unif or= pellet power profile is conse rva t ive f or D'!3R calculattens for fuel :cmoerature sat calculations anc non-conservative botn E0L ano EOL). Re sponse i guk .
=
i
Fuel Pellet Radial Ps.ser Profiles Used for Rod Ejection Analyses l l 0 C. l l L l - i l !~ , RIR o 1 4
. . ... . . . . - . - ~ ~ . . .- .~ - .. - - . - - - ..
9g 3.- Tnat -is tne ef fect of neglecting t..e time-cepencence' of the assen:1y-wise: flow distribution and assembly cross flo- in ARROTTA!
Response
The 'ef fect f neglecting the tire cependence of-.the assecoly-wise flow
~
distributiens and assembly cross flow'in ARROTTA is negligible. :n the HIP. cases. -_ the water is not significantly heated until af ter the transient poyer level peaks and starts to decrease. Bulk boiling does not occur until well af ter VIPRE ;tedicts DNB to occur. For the HTP cases,' hulk boiling never occurs in the ARROTTA model. Thus, the reactivity ef feces of the rod ef ection' transient-are relatively unaf fected 'by 'the changes to the water properties, which neans that , cross flow or changes in the flow distribution have little effect upon
- the transient results.
f k 4 kl -I
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Are all :ne ,:ey safety para =eters calculacea vich approved codes anc ' methods? Re sponse:
- Yes, Seett:n d.0 states that ~ f or generation of key safety parameters
- s. for relcad : ores approved codes and methods.will be used. The currently approveo codes and metnods are described in DPC-SE-2010A.
As the 3RC-approves other Duke Fever fopical Reports, those methods and' coces =av be substituted for the ones described in DPC-NE-2010A. t l l ; I
- t. !
l
-l l _
l l l l
I l 5. ' hat M.F.0TT! cersion and epticas are ceins; usec in :ne DPC reference analy s e s i
!! :nese are not the same verstentcoti:ns used in the coce benenzarv.tr.:
- cmear ::ns
;ustif :ne applicaollity ei tr.e cene . arnins;
.i:ensin ::iculati ns.
Re sponse : The DPC ref erence analysis used version L.32 ct idI.0TTA. The same version was usea f:r the riEPJil!E cc:.carison. A cc=carlson of the options is given :elcw. usea t . SPC and the optiens usea for the riE?. MITE cccparison The dif ferences shown are neglis;1hle. ARROTTA-HEF. MITE Parameter OP C-N E-3001 Comearison
- Coc: e n t s m
se su i ___ _ _ _ . - - - - - - - ~ - ~ - _ . _ _ _ _ . _ - . _ _ . _ . _ - - - - - - . - . - - - - - - - - - - - - - ~ - ~ ~ - - - - - ~ ~ - - ' - - '
.~ -. -.- . . - . , . , - . - . . - . . - . - - .. . . - . - . - . . . - . . - - . .
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l'" 6. ? cvide ene" appendices' t'o 'the ARROTTA coce desetiption report (Volume-1),- The code descriptien provided is' a draf t report. :-:a s this code .cescript1_on' received final approval? -Please provide t..e Iisal :oce cascription report. L
- Response:
4, ,w .J See revised response in June 3, 1991 letter s
)
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l l l i The s;x key safety paraceters of Table a-4 do not provide a c:=plete character 1:ati:n of the three-cimensional ARROTTA reference calcula tio n. For exacole. two core = having identical v.ey safety: para:eters out, due t: core loauing, f uel burnup. roc insertt:n or aenon cistribution, can nave different axial and radial power distributions wnich affect both the reactivity insertion anc the feeoback and scram reactivities. Discuss the ability of these selected key saf ety parameters to completely characterite the three-dimensional rod ejection event. 'tha t uncertainty is introduced by this specifi: selection and definition of the key safety parameters.
Response
The sLx key safety parameters of Table u-4 are very si=ilar c: the current list of parameters described in the McGuire and Catawoa FSARs for the rod ejection transient. In addition to these parameters. Techt.ical Specificatiens establish li=1ts on shutdown tartin. : ore power level. reactor system pressure, core flow, anc control ::d insertion. 'la lu e s for these parameters were set conservatively in the red ejection analyses. Other para =eters are established conservatively to assist in providing limits on initial conditions for the rod ejection transient. These include the Power-Axial Flux Difference (AFD) operating limits and F.q. The Power-AFD li=1ts restrict the initial axial power snapes and xenon distributions that must be considered. The F, limit restricts the initial pin peaking allowed in a reload core. ' 30th of these Ibnits are established through Technical Specifications and are monitored through Technical Specification surveillance requirements. Cycle-specific analyses are perfor ec to verify the bounding nature of the parameters assumen in the reference analyses. The DNB anc Fq - checks pertermec account fer cycia-specific radial and axial :ower distributions anc rod worths. The DNB and Fq checks are basec on static post-ejectec power distributions which are conservatively calculated by neglecting both coderator and Doppler feeabacks. For additional details concerning the calculation of the key safety parameters, refer to the answer of Question 13. Radial and axial power distributions also i= pact reactivity insertions and trip reactivities. Technical Specifications govern the a: cunt
- ds ::n be insert:d ca a function ci reactor power anc tncref:re limi ts the amount of reactivity that can ce inserted from a rec
' ejection event. The amount of excess shutdown reactivity available for insertion (shutdown cargin) is set by Technical Specifications at 1300 pcm. For tne rod ejection analysis, the shutdown margin was pessi=istically reduced below 1300 pcm (to 250 pcm) by assuming the simultaneous occurrence of a stuck rod coincident with an ejected rod.
The rate in which reactivity is inserted to the core is governed by the trip reactivity curve, Figure 4-7. The trip reactivity curve is
- alculated assuming a bottom peaked power distribution to delav reactivity insertion for as long as possible. This curve is verified for each reload core.
n
' LThe- cocaination c? checks on the_ six key parameters and Technical Specification lidi:s. -and the cycle- specif t: calculations serve to-
!! idefine' an acceptable envelope of reload core characteristics.- ' Changes to Tecnntcal' 3pecif t:sti:ns by requirement-are- sue::ttea to anc L reviewco'by tne ,';RC. Changes in the six key parameters outside of the b i range used. in this analysis would require an evaluation, reanalysis of the: transient. Or a redesign of the reload core.- l 1 l r I. I: i ? , l
- s. -
- i. ?rovide a descript::n of tne coce usec c precare :ne c o=c o si tio n-ce p e ncer, t cross sections fer ARROTTA. Since ene capaoility i nis coce has n:t ocen exercised in :ne :enen=arking co=carisons, ;revide ene appr:oriate c:de validati:n i:: :ne ::c ej ection application.
Response
All. cross sections were calculated by the CASMO-3 progra=. A series of auxiliarv programs transform the fuel cross sections into a datacase suita:le it: ne 5!OTRAN progra=. For eaca point in core life. SIGTRAN prepares the cross section data for ARROTTA by transfor=ing cne cross sections from the database into. tne ARROTTA co= position cepencent functions. An ARROTTA composition may be applied to dif ferent nodes in the core that have identical enrich =ent and burnable poison content Out with slightly dif ferent fuel exposure. he nodas are-ce=bined into compositions by scanning througn the core to find the noces that maten within some criteria and averstine those noces into a single co=cosica:n. Nodes are averages oniv if tne e nric hmen t anc EP content are identical. If tne nu:cer of compositions is too high. : ore sweeps are =ade with incre=entally less restrictive criteria until tne nucoer of composita:ns is acceptaole. [. correctly place ea]ch co= position into the core.SICTRAN also to creates the ARR The entire cross section process was verified at DFC by forcing SIGTRAN to create cross sections for selected CASMO-3 exposure points. ARROTTA was then forced to evaluate these cross section sets at the CASMO-3. state coints (fuel anc =ocerator temperature. :Oron concentration, e t c. ) . Finally , the cross sections procuced by ARROTTA sere' co=carea bacx to the original CASMO-3 cross sect::ns. The ability co creoare a =odel of a reactor : ore was cestec as described in Question E,
.SIGTRAN has oeen validated and certified in accordance with Duke Power Design Engineering Quality Assurance procedures.
f. low if :ne nu=cer Of ARROTTA :uel crosr secticn c::::tsitions deter:1:ec :: ce.acequate ser ::celine :ne three-ci:ensional fuel turnu: :istri:utien?
?.e spons e s
- n the tevelor:ent of the EOC :ocal for the reference analysis,
,several preli=inary ARROTTA :ocels were createc. Oc carison of ..
~
l I' I ( l l l i l 1
.' 0. The e:ettee rod location D-1 is near tne core bouncarv anc tencs to
=axt: :e leakage and to r.ini:1:e the statial reglen approacning licit 3. For tne same key saf ety parameters, vnat is the increase in cne :::er :: recs in DNB f tr :entrally 1:catec ejectec ::c?
]T} Resoonse: The !*cr.n: cal specifications allow only bank D to be inserten at full power anc allow bank D to be inserted the furthest at hot :ero power. Bant ; consicts of I rods: the center (H-08) location and tne four locations sy=me tric to D-12. Because of core loading consicerations. A the D-12 location typically has a higher vorth than the H-08 location. Thus. ;: cation D-12 vill al=ost always be tne location of the hignest wortn ej ected red,
~~ j because of these adjustments, the calculations snov DNB occurring in the two quadrants adjacent to the cuacrant with the ejected rod, !! H-08 vere chosen as the ejectec rod location, the power cistribution adjustments vould have been = ace only in' the center of the core, unich would eli=1nate this effect. Because of these power dirtribution adjustments, the DNB response f rom ej ect; g D-12 is conservative with respect to H-08.
Rean ac check cases vill be execured using approved codes and =ethods to evaluate the ej ection o f control rods f re= H-08, D-12, and f ro: the concr:1 rod banks that are not fully inserted at HIT. These cases vill evaluate ej . ttac control r:c vneths anc cost-ej ected cover distr;:utions to ceter=ine ene nu=cer of pins in DNB. These values. wel. as the otner Key safety parameters, vill be calculatec anc [ carec to tne reference analysis. If the parameters for a given
.loac cycle exceec tnose used in tne saf ety analysis, then either an evalua tio n , a new safety analysis, or a core redesign is performed.
, . . . . .- ..._. .. - - . . - - . - ... - _. ,. ..- -. .. - . . .. . . . - - ~
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lysis sensitivity s t.udies of Ref erence 4,1 -i (. I Response:-
- l. ; e peport :i:leci "E*R Rod E;eccion Accident: ARROTTA Sensit< e --
es . . nesearen ?:oject 2941-T. January 1991, is encloseci, r .. j~ \ t l I: l i. i r i 1 l-l: L F l
.2. ' Discuss ene selection of the VIPRE-01 fuel :her=al conouctivity and .
gap concuctance used t: insure conservative calculaticas of both the . f uel :e=cerature and the DNBR/ pressure in:: ease calculaticns.
Response
)
l I i l 1 1 l l 1 44 REFERENCES
- 1. C. W. Stewart, et a'l. VIPRE-01: A Ther:a1 Hydraulic Code for Reactor Cores. "olu=e 1: Mathematical Moceling, I?RI.
NP-2511-CCM-1, Rev. 2, August 1989.
.3.
How safe tyare the=cred para ejectton t ers voren, :CC, DTC, S . ?q, anc pin census Key ce te r=inec? Are these calculated for tne state with the rod' inserted or ejectec? Are tne MTC anc DTC calculated irttnertally? Is tne rod wortr. :alculated with feecoacr.? .re enese parameters been aajusted? calculatec for the state in vnich tne cross secttons nave
Response
Ejected rod worths for tne reference calculation are deter Inec using the computer code ARROTTA. Statt eigenvalue calculations are performed for the initial-conditi:n and post ejected statepoints. The initial condition for the ejectec rod worth calculation is established by positioning control banks at their rod insertion limit and perfor=1ng an eicenvalue calculation. Next, a second eigenvalue. calculation.is perfor=ed with the ejected rod fully withdrawn. For the full power cases, fuel and cocerator ' temperature feedbacks are held constant at their initial cencition values. Moderator and Doppler temperature coefficients (MICs and DTCs) are typically calculated by isotner= ally perturbing the moderator temperature caiculating or fuel temperature from an initial valuq anl then - the resultine chance in core reactivity i I 4 The ef fective delayee neutron fraction. S , is calculated in ARROTTA f or eacn composition using 6 groue delayed neutron data. E is set at its respective 6 group delayedlimit for each neutron data. red ejection transient by modifying the is shown in Table 4.4. The value of 8 used in each transient M' i j
=
l 1
A descr :::. n of the pin census perfor=ed :o ceters ne t .e nu=cer of-fuel-; :s in D: B is described en pages 4-14 and 4-15 in DPC-tie-3001-?. a
m
-14. Are tne':necxlist parameters calculatec in exactly the same canner for
.both 't..e reference analysis and :he cycle-specific analysis?
If not,
'j ustif:. :nis inconsistency.
Response; Checklis t phystes : odes paraand eters will be calculated using approved steady state methods. ' ejectic:. checklist Jor all transients otner than rod to the raierence analyses. parameters will be calculated in a similar canner The rod ejection ref erence analyses used ARROTTA to calculate the key paramete:s anc to set them ac. the appropriate bounding values. . The calculation of Doppler and moderator coef ficients using ARROTTA is perfornec slightly differently frem that using the steady state physics codes cue to the ARROTTA coding. In the steady state codes, the coef ficients are calculated by individually perturhing the fuel or moderater te=cerature and calculating the resultinc reactivity chance. In ARRCT!.i. :ne mocerator and Doopler coef ficients are calculated as describac in the resconse to Question 13. Calculations of DTCs and MTCs by. either. =e thod yield similar results. Other key paramerers are calculatac in a similar manner in both the reference analyses anc the cycle-scecific checks. t 6
- - e-w J % - p *- M
. . , . . _ - . _ _ . ,_ -.. . .m. _ ... _ .. ~ . - . . . _ . . _ . . . - . . . . . _ _ . . . - _ _ = . _ _ . . - - . . . _ ._
; i$, is the c:rcle-spec:iic pin census compar ea to a static pin census for
= the rererence- core ? _ f not , justify this inconsistency.
Response
The dose analysis was conservatively perfer=ed assuming 50% of the fuel pins fail as 'a result of D:!B. Static pin censuses for the reload cycle vill: te ' conpared against the 50% criteria. Static pin censuses
.for the IOC reference ~ analyses were perfor=ed and found to be more
- conse rva tive tnan the reference transient cases with Doppler feedback
' and are.also counceo by' the 50% failed pin value used in the dose analy sis. ,
4_ t
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l l l
- 36. !rovide tne cetails of the :se calculations of Section 4.6.
Responss: The dose the dat' analysis for the rod ejection accident vas perfornec using and assumpticas given in FSAR Chapter 13 for McCuire and Ca tawba except noted in the topical report. In this analysis the dose contr.Mitssn fro: the contain=ent release was cathematically codeled basis LOCA, assuming except the tha rod t ejection accident was similar to the cesign the source ters was limiten to the gap activity of the fraction of the assemblies experiencing DN3. This source was it.stantaneously released to containment. All fuel neole gas release was assu=ed to be released into containment and iodines were deposited in the sump water and containment atmosphere as I proportioned in the LOCA analysis. A 95; filtration ef ficiency for the Annulus leakage Ventwas fraction 11stion Syste= vas assumed. A 7% unfiltered bypass also assumed. assumed to be The contain=ent leakrate was nalf tnis a=ount f or tne re=aincer or the maximum the Technical Scecification 30 days. the first dav and Inside containment. conservative crecit was taken for todine ret .al by the ice condenser arid containment spray based on assumotJ -a given in Standard System (ECCS) Review Plan Sections 6.5.2 and 6.5.4. I=ergency Core Cooling release patns. leakage and the Hydrogen Purge Syste: vere considered as The secondary side dose contribution was calculatec by assuming maximum steam release. Technical Specification prt:ary-to-secondary leakage for the A 0.10 iodine partition f actor and a 1.0 noole gas partition factor were assu=ed in the steam generators. The entire source volu coolant released se for fthe rom the fuel side secondary was dose assumed to be clxed with the reactor contribution. and seconcary coolant were assumea to be at Technical SpecificationThe pri ary limit activity levels at the start of the accident. ' The dose calculation perfor:ec for the topical report has been revised for the Catswoa 1 Cycle 6 reload report to incorporate changes unich more accurately represent station response to the accident. The revised results continue to meet the applicable acceptance criteria. Assumptions and methodology provided by the Stancarc Review Plan a(NVREC-08001 naly s is. The and Reg. Guide 1.77 are incorporatec into the revised calculations, and source term is based on DPC-NE-3001 Chapter 4 DNB no fuel melting is assumed. Cap releases are fetermined as before. So ECCS leakage or hydrcgen : urge release containment was considered. Technical Specificati:n containment frcm leakage was assumed and the source term was released into the containment instantaneous ly . Credit spray removal of iodine was taken as before.for ice concenser and containment The secondary side steam release was extended to 8 hours with releases comparable to other secondary side release accident assumptions. partition factor was set at 0.01 tor iodira The steam generator and 1.0 for noble gas in accordance with the Stancard Review Plan. Technical Specification 11:1c concentrations beginning of the accident ofasiodine were assumed in the ::alant before. theat
1 i l i I De '.enon-incuced top _ 1 cycle. spec fic analysis s to . P wer1'distributien used 'n th
. deep 17 inserted rod), _ustify$1 su erd (e. .. for a 3 "PP1 tations.
Response
^
- -t0D-Peaked axial power distr 'ution is conservat**'8 for ene R1 im W
l l .'
l l l
- 18. The present VIPRE-01/ARROTTA nocel assunes Mark-BW desien data includist dimensions and loss coefficients.
"alidate: f or new fuel designs? Sov vill this nodal be !
Response
The roc ' ejection analysis results presented in DPC-NE-3001 are based on the c'2rk-BW fuel design for the thernal analysis, and on the OTA design for the neutronic analysis. As stated in the response to Questi::-1, the neutronic differences between OFA and Mark-BW are minor. Since the cross sections were significantly modified to ootain a highly-peaked power distribution with bounding physics paraneters, the ARROTTA analysis is conservative for either _ fuel type. The VIPRE-01 thermal-hydraulic analysis was explicitly analyzed for both fuel types. The Mark-BW results are presented in DPC-NE-3001 since this fuel type is the new one. -; The results for the OFA fuel type are slightly worse but meet all acceptance criteria. It is expected that future f.:e1 desiens will be suf ficiently similar to Mark-BW so tnet ene. nocel will remain valid. For any fuel design change the inoact en all analyses vill be evaluated and a reanalysis performed as necessarv. I. k i f h t 4 +
i t-
- 19. Provide the uncertainty estimates and basis for each key safety parameter. k*ha t c:ditional uncertainty estimates will be used te account for the uncertainties introt'uced by (1) the VIPRE-01, RETFri-02 and AFIOTTA codes, modeling and assunoriens and (2) tne selection and definition of the enecklist parame t ers ?. How will these uncertainties be incorporated in the fuel renperature DNB, and licensing predictions of the rod ejection, steam 1.ine break und dropped rod events?
Response
Se+ revised response in June 3, 1991 letter Y
- 20. _ Has the fuel temperature calculation in ARRO*!A been conservatively
, modeled to minimize the Doppler feedback?
Response
SWI M' n i-l l
- 11. Regulatory Guide l.77 recoc=enos a low-powerec calculation as well as the hot-tero-power red ejection calculations. 5.'ha t reference analysis wil1 ~ be perf ormed for the low-powered case ? How will the possibility of a positive FTC he evaluated?
Respnnse: Regulatory Guide 1.77 Section 3. recon = ends analysis of the rod ejection accident for at least the following three initial conditions: o Hot stancby o Low power o Full power The analyses documented in DPC-NE-3001 do not cover the hot standby case from the Regulatory Guide. Hot standby at McGuire and Catswba is defined as the core being sube" 'tical by at least l~ak/k. The reactivity inserted to the co in excess of 13 at hot standby conditions is considerably less than the ej ected red worth assumed in the low power case. There for e , there are no severe consequences f rom a rod ej ection transient from hot standby concitions. The DPC-NE-3001 low power case initial condition is critical at 10 ~9 times nominal power. The choices of core power initial conditions for the DPC-NE-3001 rod ejection cases are conristent with the McGuire and Catauba FSAR Chapter 15 analyses. The SERs for McGuire and Catawba, NURECs-0422 and -0954, both conclude that this approach is "in accordance with, or more conservative than, those reco= mended in Regulatory Guide . l .77". Therefore, the DPC-NE-3001 choice of initisi power levels ' 1s consistent with the approved licensing bases of the McGuire and Catauba Stations. In addition, it would be expected that a full power case would bound results from positive, but lover power levels for at least two reasons. Firs t , the fuel temperature coef ficient. becomes nore negative as core power decreases, even though no' credit was taken for this in DPC-NE-300 7, i.e. . the same coefficient was used for the full and low power cases. Seconc. four reactor coolant pumps are always required to be operable per -he McGuire and Catawba Technical Specifications while the reactor is at power. The ratio of core heat flux to required core flow is therefore highest'at full power, where a high value of this ratio minimizes the margin to DNB. -The analysis was perforced at the Technical Specification li=1ts on MTC: 0.0 pcm/F.at HFP to +7.0 pec/F_ at HZP. Thus, the ef fects of a positive MTC are already included in the
. analysis.
l l-
+
- 2. Previde :ne canis i:r the assutec ::: ejecti:n velocity ans scra delav tt:e?
Fesponse: The assuced red ejection velocity is c:nstant based en the active core lenet.. divided by a rod ejection ti:e Of 0.1 second. Die 0.1 secend value is the current TSAR assu=pticn. The scra= delay ti=e is 0.5 secones. This is also an TSAR value anc a Technical Specif1:ati:n s urveillanc e. .-
- 23. Is any :recit :sken f or tne assutea rupture of the contrcl red housing in tne rod ejection pressure calculati:n?
Response' 1
"o. :!o 1:ss of coolant and depressuri:stien are codeled due to the ejected sco.
l l l ' 24. '.tha t nocific.a t:cns will be made to the ref erence analyses in t'.c case that future reloaas include new fuel desicns involving chances to parametern not inclucec in the key caf e ty para eter :necklist unich makes the retcrence analvs2s less b oun c ir. g !
Response
Bef ore any :uel a olies with signifi: ant desien :1f ferences from the current .:csirns would be included in a reloac, a complete safety evaluation would be perf ormed. This evaluatien vould determine whether the reterence analyses remained valid or unether reanalycis would be necessarv. 3e intent of the key saf ety paracoter e.hecklist is t o ident11v vnether reanalysis is r.ecessarv. Se reference analyses include a set of conservative assumptiens that are ev'- s sd to remain bouncing f:r the foreseeabic future. Se reload sas
- evaluation t rocess described in DPC-NE-3001 will identif y any >
situations anc appropriate action vill be taken. For parameters rot included in the key safety parameter checklist. it is concluded t5at their tn11cetive tenact on the safety analv:.ts as a result of the er.pectea c2nor : f fcrences tietween reloacs will :e insignificant. T
- 25. So the initial conditions (core power, flow, pressure, te=perature, etc.) and core protection setpoir . s used in the transient analyses include an allavance f or uncertainty ? If not, how vill this
;ncertainty be a cc c::=oca t e c ?
Responsa i This question was partially answered in the response to Questien 19 ! which stated that all of the initial conditions have included an l appropriate allowance f or uncertainty. The uncertainty associated with the Reactor Protection System setpoints is explicitly required and is documented in the Catauba Technical Specificatiens. Be McGuire Technical Specifications de not include this information, however the ssme analyses are applicable to both stations. A total uncertainty allowance plus scue cargin has been factored into the analysis values. _ _ __ ___ _ _ _ _ . _ - _ . _ - - - - - - - - - - - - - - - -~ _
i i DUKE POWER June 3, 1991 U. S. !!uclear Regulatory Commission ATTil Document Control Desk Washington, D. C. 20555
Subject:
McGuire Nuclear Station Docket Humpers 50-369 and -370 , Catawba Nuclear Station Docket Numbers 50-413 and -414 Response to Request for Additional Information Relative to Topical Report DPC-NE-3001 By letter dated February 13, 1991, Duke Power Company submitted responses to NRC Staff questions relative to Duke-proprietary Topical Report DPC-NE-3001. As a result of telephone conversations among the report reviewer, NRC staf f, and Duke, revisions to two of the responses (6 and 19) have been prepared and are attached. Also attached are responses to a second set of questions which have been received. Note that the second set of 21 questions has been renumbered (26-46) consecutive to the original 25 questions to avoid confusion. Please find also (Attachment 2) marked-up revision pages to the Topical Report. The report will be reprinted in its entirety upon receipt of the Safety Evaluation Report. Ploese note that both the question responses and the revision contain proprietary information and should be withhold from public disclosure. Art affidavit supporting this designation was contained in the original submittal of the report, dated January 29, 1990. If there are any questions, please call Scott Gewehr at (704) 373-7581. Very truly yours, M. S. Tuckman resp 3001/ sag
I fluclear Regulatory Commission June 3, 1991 ' Page 2 i cc: Mr. T. A. Reed, Project Manager Office of liuclear Reactor Regulation U. S. lluclear Regulatory Commission Mail Stop 91!3, OWFli Wachington, D. C. 20555 Mr. R.E. Martin, Project Manager Office of 11uclear Reactor Regulation U. S. !!uclear Regulatory Commission Mail Stop 9113, OWFli Washington, D. C. 20655 Mr. S. D. Ebneter, Regional Administrator U.S. liuclear Regulatory Commission - Region II 101 Marietta Street, liW - Suite 2900 Atlanta, Georgia 30323 Mr. R. C. Jones Reactor Systems Branch Office of 11uclear Reactor Regulation U. S. lluclear Regulatory Commission Washington, D. C. 20555
l l (6) Provide tne acpchcices to Ine ARECTTA cc:o coscription report (Volume-A). The coce cescription provideo is a draft report. Has this ::de descriptien received final approval? Please provide the final ::ce description recort. Respense: The appenclx describing the ARROTTA cede tneory was transmitted to the NRCJ ' ocument Control LesK in a letter f rem d. E. Tucker cated Maren 20, 1990. Appencices A-1 and A-2 were omitted in that transmittal. Enclosed are appencices A-1, A-2, ano A-3 f rom Kore S. Smith's Master's Thesis at the Massacnusetts Institute of Tecnnology: "An Analytic Nodal Metnod for Solving tne No-Group, Multidimensional, Static and Transient Neutron Dif f usion Equaticns," dat ed March 1979. These three appendices will be incorporateo into Appendix A of the ARROTTA theory manual. - The final ARROTTA Volume-1 code report is currently expected to go to the printer near the end of May and be releasec in the middle of June. EPRI has indicateo nat there are no substantive changes from the draft version proviced with the Duke Power sutmittal. ( 19 ) provide the uncertainty estimates and basis for eacn key safety parameter. What additional uncertainty estimates will be used to acccunt for the uncertainties introduced by (1) the VIpp2-01, PITRAN-02 and ARROTTA codes, :nodeling and assumptions and (2) the selection and definition of the checklist parameters? How will these uncertainties be incorporated in the fuel temperature, DNB, and licensing predictions of the rod ejection, steamline break and dropped rod events?
Response
The methodology of OpC-NE-3001 addresses the issue of uncertainty in models and input by applying explicitly determined uncertainty allowances and by aading conservative margins to reload-typical values. This approacn 13 applied to both key parameters and to many other parameters, ae.d ensures suf ficiently conservative results. Uncertainty modeling and allowances for the steam line break, rod e]ection, and dropped red analyses are listed in Tables 1-3 and are based on the following approach. Initial Csnditions: The initial conditions in the analytical models include appropriate allowances for uncertainty. These are summarized below and include allowances on power level, temperature, pressure, flow, pressurizer level, etc. For analyses using the statistical core design (SCD) core ther:nal-hydraulic design approacn, the uncertainties in the DNB-related parameters are incorporated in the DNBR limit and are not repeated in the simulations. i Boundary Conditions: Boundary conditions are modeled with attention to the impact en the transient responsa and the end result. These include parameters such as control rod drop time, ECCS flowrate, Reactor protection System setpoints and delays, etc. Each parameter value is conservatively selected and many include an explicit uncertainty allowance.
l Code oct1:r.s: Iacn cf the codes in tne retnodolcgy includes different modeling cpti:ns, tne use of wnicn is de te r:1nec Dy the user. The selection process is analyisir specific. Examples are tne models f or the fuel-c2aading gap, ::re enormal f eectacx, and time step selection. These uptions are seiectea cased on sensitivity studies, ::mparisons to data or reterence analyses. Or cased on user experience. For those options wnich impact a given analysis, attention is paid in :ne selection process to ensure a : ens e rvative result. The selected rodaling options are descrinco in tne report. Core Thermai-Hycraulic Models: Uncertainty in predicting local fluid conditiens and critical heat flux is addressed by :ne statistically-based DNBR limit. Specific uncertainties in fuel assemoly design and manufacturing variances are accounted for in the SCD ONBR limit or in the hot channel facters if the SCD approach is not used. The SCD limit also includes an uncertainty allowance for the VIPRE-01 code. Core Physics Models: The core power distributions include 95/95 reliability f actors which are documented in DPC-NE-2010A and have been approved cy 'ne NRC. Axial flux difference, scram wor ;h. and rod position values all include explicit uncertainty allowances. Key safety analysis physics parameters, such as moderator and Doppler temperature coefficients, offective delayed neutren fractian, and e]ected rod worth (refer to Table 2-1 of DPC-NE-3001) are detemined using the bounding parameter approach. Based on current reload designs and expectations for , future reloads, values which should not be exceeded are eclected for key parameters. Margin exists between cycle-specific values and the values assumed in the transient analyses. Due to cycle-to-cycle variations in these parameters, the margin may vary. However, the integral effect of margins in all parameters ensures a conservative result. Additional margin exists between the analysis results and the acceptance criteria. The above apprcach for introducing sufficient conservatism into the analyses is consistent with the current licensing basis. Uncertainties are determined for some but not f or every key safety parameter. For all important parameters margin exists between the values assumed in the analysis and calculated values for current cores and those expected for future reload cores. Code uncertainties are not determined (note VIpRE-01 code and nuclear reliability f actor exceptiens) . Each code has been validated to benchmark data as part of the code development and model development processes. Based on these efforts it has been concluded that each code is suited for the applications in DpC-NE-3001. In the event that any key safety analysis physics parameter associated with a given reload exceeds the value assumed in the transient analyses, a 50.59 evaluation will be performed to detemine wnether an unreviewed safety question exists. ,If this evaluatirn has a positive finding, then a re, analysis will be performed and submitted for NRC review. Alternatively, the reload core can be redesigned. s
1 l
!able i Steam Line Breax A.*.alysic Uncertainty Allcwances
- ETTx4 Ana . n is
- n2tial Ocnditions:
o Pressuricer pressure: -30 psig e Pressurizer level: at e RCS temperature -4 T e RCS flow -2.2% e SG level: +8% e core cypass flow: +1.5% Boundary Conditions: e Maximum auxiliary feeowater flow e Maxirum credible main f eedwater flow e Maximum safety in]ection flow wnen unborated e Minit:mm safety in]ection flow wnen borated e Bounding long time for f eedwater isolation e Bounding long time for main steam isolation e Bounding large purge volume for safety injection piping
- Boron concentration in safety injection water: -1%
"a e Bounding moderator density reactivity feedback e Bounding fuci temperature reactivity feedback e Minimum Tech Spec shutdown margin "a
e Minimum boron worth -10% e Minimum credit for reactor vessel thermal mixing e Maximum delays in engineered safeguards actuation Core Power Peakino Analysis e Axial peak uncertainty:( 3 e Radial pwak uncertainty:C 3 eRadialpeakuncertaintyduetotilt:( e Shutdown margin calculation } o -10% rod worth uncertainty o worst stuck red Core Thermal-Hydraulic Analysis e Enthalpy rise hot channel f actor to account for manuf acturing tolerances: +3% e Subchannel flow area reduction: -2% e Hot assembly flow reduction: -5% NOTE: "*"" indicates key physics parameters
+
TeLble 2 Rod Ejection Analysis Uncertainty Allowances
- AR":*"'A Ana l /t i s Typical e Parameter Value Assumed Reloac Value
*" Ejected red worth: HFP/90C 200 pcm SS pcm HFP/EOC 200 pcm 79 pcm HZP/BOC 720 pem 327 pcm HZP/E00 900 pcm 411 pcm Seta-effective: HFP/BOC 0.0055 0.0061 -
HFP/EOC 0.0040 0.0052 HZP/BOC 0.00551 0.0061 HZP/EOC 0.0040 0.0052
"* _OTC: HFP/BOC -0.9 pcm/F -1.18 pcm/F HFP/EOC -1.1 pcm/F -1.45 pcm/F HZP/BOC -0.9 pcm/F -1.51 pcm/F HZP/EOC -1.1 pcm/F -1.85 pcm/F MTC: "JP/BOC 0.0 pcm/F -12.1 pcm/F
' HFP/FOC -10.0 pcm/F -32.3 pcm/F HZP/BOC +7.0 pcm/F -2.8 pem/F HZP/EOC -10.0 pcm/F -17.5 pcm/F F-Q (total pin HFP/BOC 4.12 3.00 peak) HFP/EOC 4.88 3.40 HZP/BOC 16.62 6.30 ) HZP/EOC 23.55 9.50 Pin Census .J eRadiatpeakuncertainty:("J) e Axial peak uncertainty:[ o E _ Core Tnermal-Hydraulie Analyses e Enthalpy rise hot channel factor to account for manufacturing tolerances: +3%
. Sabchannel flow area reduction: -24 e Hot assembly flow reduction: -2%
~
1
OE* FAN Araly9:s_ e Pressurl::or presnure: +60 psi
- Pressu - :r level: -95
- Fressut. ,ec safety valve modeling o +M drift in lift :stpoint a M accur::ulation ,
L , NOTE: * *" indicates key physics para.teters M _--- - - - - - , - - - - - - - _ - - - - - - - _ - - - - - - - , - - - - - - - - - - - - - - - - - - - _ . - _ _ - - - - - . - - - - - - - _ - _ . - - - - - _ _-_u__ -- -
l l Table 3 Drcrpes R0d Analycis Uncerto:nty All:wances PETPN4 /*nalynit e Uncertainties in power level, flow, rypass flow, pressure, and temperature and f act n eo into the SCD DiBR approacn e Pressuri;:er level: -9% o SG lovel: -85 Core Phycice and Power Peakino Analyses
- Baunding drcpped rod worth e Bounding available rod worth for withdrawal
- Bounding core til*.
e Bounding MTC e Eounding DTC e Bounding ceta-effective e Bounding radial peax e Bounding axial peax e Power peaxing uncertainties are f actored into the SCD DNBR approach Core Therm 1 , Hydraulic Analysis e Hot assemoly flow reducticn: -Si e other uncertainties covered by the SCD CNBR approach NOTE: indicates key physics parameters
;6) In a 2.3.22)? recriticality :f so, analysis perf ormed f or the RCP startup event (Section physics safety parabeter isn't ino f crminir.um this event? snutdown margin ccrsidered a key Rerponce:
No, tocause that McGuire and Catawba Technical Specifications 3.4.1.1 require operating.during power cperaticn and startup all f our reactor coolant pumps be W y under two conditions: Therefore, reactor coolant pump start is procedurally allowed
- a. An isothermal, i.e.,
reactor coolant loops are cero at core power, situation in which 1) all the same temperature concition as each other and as the core and 2) backwards forced circulation is provided in the inactive loops by a portion of the flow from the operating pumps. In such e situation there is essentially no temperature transient initiated difference by a puntp in therestart. RCS and there can be no reactivity
- b. ;
Because three loop cperation at McGuire and Catswba is not 'd, the " nominal N-1 ;cep operation values," refecred to in . son 15.4.4.2, Item 1 of each plant's FSAR, do not exist. The trip of any single pump above 48% power (the Technical Specication P-8 interlock) would cause a reactor trip. Three pump operation at power is therefore an unlikely phenomenon since the permissible initial condition, a single pump trip from less than 48% power so that the reacter does not trip, is very rare. Three pump operation in therefore a transient condition prior to restart of the fourth
- pump to satisfy the Technical Specification requirements.
' plant Since the analysis. begins such a transient at power, there is no recriticality (27) Why isn't the prompt neutron lifetime a key safety parameter for the uncentrolled ROCA withdrawal? [
Responser a. Uncontrolled Bank Withdrawal at power, FSAR Section 15.4.2. The core kinetics response in the Uncontrolled Bank Withdrawal at Power (UCBW) transient is dominated by the moderator temperature and , fuel temperature feeuback. For a given set of feedback parameters, a withdrawal rate sensitivity is perfomed cvor the spectrum of rates from very slow to the maximum possible. A change in the prompt neutron lifetime caur.es a small change in the kinetics response, which manifests itself as a small change in the rod withdrawal rate at which the MDNBR occurs, but the value of the MDNBR is unaffected., b. Single Uncontrolled Rod Withdrawal at power. FSAR Section 15.4.3d. Since the reactivity insertion during the transient is lianited to the worth of a'aingle rod, the limiting statepoint occurs af ter the rod has been fully withdrawn. At the limiting statepoint, the net reactivity due to red withArawal and feedback effects is nearly 4 4
l l ! cero. and tne time intervci bring wnicn t*.e prcmpt eutren ..u t re
.; s i g n i f i ca r.t has passed.
fncontrolled Bang Withdrswal at c:bcritical. FEAR Secti:n ;5.4.; A censitivity stucy has Leen cerformea to demonstratt. that the prc pt neutron .ifetime is nt a key safety pararoter for the ancontrolled bang withdrawal at suberit' al transient. The stucy
;nowed that a sensitivity f acter of( exists for in1. parareter (a[ ) change in the parreter results an a in the tranutent result). A censitivity factor of (this }cnange small magnitude contirms that the pt ept neutren is not a key safety parceter for the uncontrolled banx withdrawal at subcritical transient.
(:0) Why isn't the decay heat a key saf e'y parameter for the loss-ot f eed-water / loss et offsite power and LOCA events? Fesponse: Decay heat ;s a key safety parsmeter for these transients as well es other concern. transients where post-tr:p overnesting of the RCS or tne fuel is of In Duke Power terminology decay heat is considerec to te a boundary condition rather than a physics parameter. It is for that reason that decay heat is not included in Table 2-1 of CPC-NE-3001. A conservative decay heat boundary condition will be assumed in those analyses !cr which it is important. (29) Is the application of VIFRE-01 and FITRAN-02 in the rod ejection, steamline break and dropped red analysis censistent with the limitations of their approval (e.g., Reference 4-11 in the case of VIPRE-01)? Respor.se : The limitations, on tho' approval of RETRAN-02 are given for the MOD 002 code version in Reference 1. Although the code chor.ges resulting from the MOD 003 and MOD 004 versions of the code were recognized in Reference 2, that document left the limitations of Reference 1 essent.1 ally unchanged. The approval of PITRAN-02 M00005 has not been issued. The limitations in Reference 1 were reviewed with respect to the analyses in DPC-NE-3001. The conclusions regarding the pertinent limitations are given below. The letters correspond to those in the original list:
- a. The neutronic space /tme effects of rod ejection are not s mulated with RETRAN-02. For the steam line break and dropped rod transients the multidimensional space / time etfects are conservatively treated as explained in DPC-NE-3001, b.
The neutronics model is not started from subcritical or zero fission power. For steam line break, the initial condition is critical at 10 ' times nominal pcNer. Shutdown margin is simulated by an i:maediate reactor trip of an amount of negativo reactivity equal to that shutdown margin.
- c. The generalized transport model included in MOD 005 version has been used for boron transport moceling in the steam line creas transient
l l I anif
're conservative cpp11:sta:n of this rodel witn respect to purge vetumes and assumed toren :encantrations is discusseo ;n sect;;n 5.2.2.$ of CPC-NE-3001.
- f. The enequilibrium pressurizer mocet is used in all transients 'n OPC-NE-iOO1.
m. The core neat transter ;n the steam line creak and dr:pped red transients is restricted to situations in which single pnase er pre-CHF regimes dominate. The transient core heat transf er in the RETPN1-02 analysis of rod e]ection, as described in Section 4.5 of DPC-NE-3001, is not simulated. o. The steam line break transient was analyzed both with and without ~ wall neat conductors in the pressurl:er. The analysis without these concuctors produced the lower core exit pressure at the ONB statepoint and this analysis is theref ore presented in CPC-NE-3001. The dropped rod analyses used
} to simulate wall heat concuction the{ in the pressurizer as descr: red in Section 6.2.1 cf :PC-NE-3001. The RETPAN-02 red e;ect;on analysis used the( }modelingastne dropped rod analysis.
q. The suitaD111ty of the def ault Westinghouse single phase homologous pump curves for the McGuire and Catawba reactor coolant pumps is deconstrated in Reference 3.
- u. The dropped rod and RETRAN-02 rod ejection analyses use no .
applications justified in Reference 3. of the bubble rise model which were not already Theuseofthe( jsteam generator secondary model presented in Reference 3. The otner ~ aspects of a( }are discussed in the response to Qt.estion 32.
- v. A dominant flow direction is provided by sustained forced or natural circulation in all loops in all transients analyzed in DpC-NE-3001.
- x. The steam generator model used for the dropped red and tha RETRAN-02 rod ejection al.alyses is the[ ]modeldescribedinReference 2,
wnich has the phase separation in the[
]
- y. Application of the RETRAN local conditions heat transfer model in DpC-NE-3001 is [
}The local conditions model provides a more
reaA;;;;; 1pproach than ino stancarc nemogenous model, anc is appliec for that reason. 2. As part :t tne pre-executien review cf the . nitialisation for a part::ular transient, the analyst ensures that the feecwater flow is ad]usted so that the RETRMbO2 fill enthalpy cias correcticn results in a negligible enange to the user input enthalpy. Similarly, when there ir a enange neat transfer in the occurs, conditions e.g., due under wnlen primary-to-secondary to steam generator tube plugging or a change in power level, the(
}until the RETRAN-02 haat transfer area adjustment correction results in a negligible change to the user input area.
- 1) Validation of the secondary-side transient modeling including two-phase effects is addressed by the benchmarxing perform 3d and documented in Reference 3. (
5
} HEM modeling is adequate f er these conditions, ii) The pressuriter does not D PC-!!E- 3001.
fill for any of the transients analyzed in seconds. The rod ejection transient is st.:ulated for only a few Although the pressurizer level is still rising at the end o f t his t ime , the shatp decrease in energy deposition into the RCS which accompanies rod insertion, the increase in energy removal from the RCS wh!.ch accompanies steam line safety valve lif t, and the loss of RCS inventory out the break, will reverse the level trend prior to filling the pressurizer. SBLOCA following The rod ejection accident resembles a reactor trip. The qualification of the McGuire/ Catawba RETRAN-02 pressurizer model for pressurizer emptying was presented in deference 3. The limitations on the approval of VIPRE-01 are given in Reference 4. The limitations in Reference 4 were reviewed with respect to the VIPRE-01 analysis in DPC-NE-3001. The conclusions regarding the partinent limitations are given below. The numbers correspond to those in the original list (Page 28 of Reference 4): (1) In the rod ejection analysis, the transient VIPRE-01 fuel pin conduction model is utilized. In the fuel temperature /entnalpy calculations, the heat transfer can reach the film boiling regime. On Page 4-13 of the report, heat transfer correlationc used for the four major sogaients of the boiling curve are shown. The use of these neat transfer correlations was determined by sensitivity study to assure acceptable and conservative fuel temperature results. In Raference 4, 1* states on Page 28: "The application of VIPRE-01 is limited to PD. licensing calculations with heat transfer regime up to CHF. Any use of VIPRE-01 in BWR calculations or post CHF calculations will require prior NRC review and approval." Thus, Duka Power Company requests NRC's review and approval for use of VIPRE-01 in post CHF calculations in the rod ejection analysis. i
in the steam line creax ano dropped rod anaayses, tne FITRA!'-02 heat fit.x rsoundary conditicn is used and the 71 PRE-01 fuel conduction moael is not e~ployec. This means that heat is added directly from
- ne cladding surface to the fluid as a boundary condition in the calculation. Co r.s equ e ntly , the heat r pu re d.
transfer solution is not (2) :n the rod e]ecticn analysir., the W-35 critical heat flux (CHF) currelation is used to define the peak of the boll nq curve, and the mar.imum DNBR value f or wnich transition boilina cccurs is set to be 1.30 (Page 4-13 of the report). The historical W-3S CHF correlation limit is 1.30, which was det e rmined using closed channel core themal-hydraulic rnethods. The W-3S CHF correlation limit utilizing VIPRE-01 has not been determined. Since VIFRE-01 is an cpen channel core themal-hydraulic code, it is obvious that a lower DNBR limit value could be obtained if VI?RE-01 were to be used to determine the correlation limit. assumed in the analysis. Therefore, a limit of 1.30 is conservatively In the DNBP evaluation of the reo ejection analysis, the EMCMV CHF correlation is utilized to generate the maximum allowable radial pean ng (MARP). The correlation limit utill:1ng VIPRE-01 has been determined to be 1.21 (Reference 5). The NRC is currently reviewing Reference 5. The Dt"3R limit utilized in the rod ejection analysis is 1.331 (1.331 = 1.10 x 1.21 where the 1.10 factor adds 10% raarg in ) . In the steam line break VIPRE-01 analysis, the W-3S CHF correlation is utill:ed for the DNBR calculation. As described above, the W-35 CHF correlation limit utilising VIPRE-01 has not been determined. However, a DNBR limit of 1.45 (1.30 x 1.115 where the 1.115 f actor adds margin) is utilized in the analysis (Page 5-21 of the report). In the dropped rod VIPPI-01 analysis, the EWCMV CHF correlation is utiller,d for the CNBR calculation. The BWCMV Statistical Core Design (SCD) limit of 1.55 (Page 6-3 of the report) is employed in the analysis. This SCD limit of 1.55 has been determined utilising the VIPRE-01 code (Reference 5). (3) For the rod ejection analysis,, the specific modeling assumptions, choice of two-phase flow models and correlations, heat transfer correlations, thermal-hydraulic correlations, and conservative factors, etc. are described in Sections 4.2.2.2, 4.2.2.3, and 4.2.2.4 of the report. For the steam line break and dropped rod analyses, the core thermal-hydraulic models utilised are described in Sections 5.2.3.2 and 6.2.3 of the report. The choice of two-phase flow models and correlations, thermal-hydraulic correlations, and conservative facters, etc. ara described in Reference 3. The NRC is currently reviewing Reference 3. (4) In the rod ejection cnalysis, the VIPRE-01 iransient mode is employed in the fuel temperature, DNBR, and coolant expansion rate _ _ __ _ __ ._--___ - - - - - ~ - - - - - - - - - - - - - - - - - - ---- --
l l l
- .cuis:;:nn. *ne p r:f f.;e f;: z ul.cco m o vold oce; ,;uen is i.evy
_.a EFR: mocels) ; ret uti..:eo n tne analysis. {
- ceic are usea as escrice: en face 4-6 :: tne report. ]
.u s . tne
- ansient time crap s;:e ;s a a concern.
- octn :ne steam .ine creax and dr pped rod analyses. :no 7 :F F.E-01
- easy-state mode : empi:y in :no CNBR calculation. hus, :ne transient time step 31:0 1: again not a concern.
- 5) :.xe Power C:mpany nas acided by the quality assurance procedures descr::ca in Section 1.6 of Ref erence 4.
'30) ;;es 01tawc1-2 have a larger steam generator inventory than McGuire. !! -
not. .w in this nonconcervatism in the steamline orcax analyst.s acceuntec f r? *
'esponse:
'le s . u acre power, snien .s *no :n.t;al concition f:r :ne c: cam 1.ne creax inalysis presentec .c ;PC-NE-2001, Catawoa Unit : Oces nave a Larger steam generator liculd ass than eitner McGuire unit. Chis is cecause the programmed level at Catawoa Unit remains constant, at the full ;:wer value, as power level doccesses, while progra:mned leval at the other tnree units decreases as power level decreases.
'21) What re the modeling differences between tae DpC and *;estingnouse analyses of tne steamline creax and dropped rod events?
Response
The relevant non-preprietary Westinghouse tog ical reports, WCAP-9227 and WCAP .';95,. along with the McGuire/ Catawba FS.\R descriptiens, were ~ reviewea to determine the moceling dif ferences be'. ween tne Duke Power and Weutin;nouse analyses of tne steam 11
- breas and dropped red events.
The f:llowing list incompicte is for two main reasons: ')
. the Westin;nouse methodology is presented in insufficient detail to determine all differences and 2) some differences which are discussed by Westin;nouse are deleted from the non-proprietary versions of the reports.
As Duke understands them, the differences for steam line break are:
'. ) Westingnouse takes creo. for a reduction in :e primary-to-secondary neat transfer af ter the steam generator U-tuces encover.
- ) Westingneuse takes no penalty for cold, possibly uncorated flow from
..e intermediate head safety in3ection pump (s) disenarge ":1 ping.
- 3) Westinghouse takes a penalty for 1.unediate auxiliary feedwater flow delivery, while Duke Power assu::es i:xnediate delivery after the e ar. lies t actuation signal, in this case safety injection.
l i _ _ _ _ _ _ _ _ _ _ . _ - - - - ~
i I
- 4) .est;n;nouse a:ea : .e
-eenn;;al 5;ecificat;:n .;mit :n refuel:n:
.at er :torage , tang coren c:ncentrat;;n without taxing a penalty f:r
- ncentration measaroment e-....
I)
..3:ntained s t ; .;nouse andassu.
lost, ,es colays Of M anc 43 secencs. for offsite tower rescectivety, netween tne generation f a l
;a:ety .njection sicnai ans :ne teint at wnl:n tne ECCO valves reacn l ne;r :inal pusitiens
- ..
- :eed. The Duke and tne nign t;ead saf ety injection p=p is at
- r. iection 5.3.2,2 ot Power : rresponcing assumptions are specified
- PO-ME-3001. In Revision 1 tho folicwing sen:ence will be added to the enc cf the relevant se::::n: pocagraen of this For the case 'a wnicn offsite power is maintained, the actresponding delay is 19 secunds.
- 6) sest.nenouse as surne s that the ;;ss of offsite power occurs
- w;aneous sarety signal. with the steam line creak and the initiat10n of the Althougn :nis is an approximate statement 01nce the i:r er tvent causns the .atter, the two events are in close
- 20:ession in the Westingnouse anaivsis.
As shown in Table 5-2 of
;PC-SE-3001, there is a 35 second tefference cetween the times of ne :wo events in the Duke Power ana.ysis. Therefore, as explained
;n sect:,on 5.3.2.1 of DPC-NE-3001, Cub Power censervatively assumes
.nat tne loss of of f site power is conci.rrent with saf ety in3ection.
7) West:,ngnous e assumes e<Tullibrium xenon conditions when calculating tne avail;.ble shutdown margin. transient xenon conditiers which typically re Duke Power calculations assume in a the aveslable shutdown margin by as much as( sult ]pcm.reductionin 8) he destingnouse k-e f f ect.1v'e versus temperature curve includes the effects of pressure, temperature and a rodded core with the most reactive rod at its fully withdrawn positicn. The Duke Power
<-ef f oetive versus temperature curve was con v rvatively generated attn respect to the acove mentioneo condit.ons in addition to
;ne:. ding the reactivitypenalty(
3 As Duke understands them, the differences for dropped rod are: 1) West;ngnouse assumes a linear variation of Control Bank D worth versus carnup to cotain :ne uiddle-of-cycle value. Duke Power expi;:ltly evaluates the wortn at this condition. 2) The Westinghouse analysis a ssu:nes eyele-specific values for the centr:. temperature rod worth availanle coefficient for withdrawal and for the moderator (hTC). The Duke Power analysis assumes cour. ing values for both the control rod worth available for witnd. awal and MTC. 3) The ass =:ng:entroJ a rod worth available 7 for withdrawal is calculated
!from the HFP roc insertion limit in
_ _ ___ _ ----- ~
l tne Duke Ecwer analysis. The Westingneuse analysic calculates tne control roc sorth available for withdrawal frem the HFP rod insertien timit.
- 32) What errer is introduced by the steam generator ~.odeling . sed in the steam line creax analysis?
Response
The( )McGuire/ Catawba steam generator secondary model described in OPC-NE-3000 is generally used by Duke Power Company in system trancient analyses because it has two main advantages ovtr a simpler model:
- 1) Multiple nodes allow a more m realistic mass and energy distribution within the steam generator, making it easier to accurately predict physically significant quantities cuch as total cteam generator liquid and vapor masses, U-tube bundle region void fraction, and preneater subcooling.
11) Having separate, geometrically accurate nodes inside and outside the U-tune bundle / wrapper boundary enables a meaningful calculation of the steam generator level indications at the plants, which are based on differential pressure. These etfects are not important for a steam line break at zero power:
- 1) Because of the ery low steaming rates prior to the accident, the steam generator invento., below the .4xture level is much more nearly saturated single phase ligaid. With much more uniform mixture thermodynamic ce ditions and a negligible U-tube bundle void fraction, a{ secondary model is much more accurate at S
predicting tne above parameters. At zero power, the feedwater flow r is not introduced into the preheater region, but rather into the upper downcomer. This gives this flow a chance to mix with the fluid already in the steam generator before entering the preheater. As a result, the preheater, at zero power , is not significantly subcooled compared to the remainder of the steam generator.
- 11) The steam generator level indications are not used as inputs for either steam automatic or :ero line break at manual actions in the mitigation of the limiting power. Therefore it is not necessary that the steam generator secondary nodalization used for this analysis be capable of modeling the level indications.
(33) Discuss the cause of the observed mixing during the McGuire forced circulation tests, and the applicability of these results to steamline breaks in the other three loops as well as the the Catawba units.
Response
)
I
7 Y 9 e ammp , (34) The flow behavior observed in the McGuire tests, and used to determine the thermal mixing, involves mixing between the loops. Discuss the 2 adequacy of the RETRAN-02 model to calculate the ecserved mixing.
Response
m.
'M>
Mi (35) 21scuss :n detail the method used to convert the coserved limiting ther:tal mixing (LM) of Appendix-A into a flow representation.
- esponse:
~
;t snould be noted that this relation is valid only for a four loop clant.
(36) :n the steamline break analysis the assumed isothermal cooldown reactivity is less than the actual distributed temperature cooldown _____,_,__.__u---->-- - - - - - " - ~ '
reactivity due to spatial weighting. How is this accounted f or? noncenservatism
Response
The react:vaty insertion f rom a distributed temperature coolder is more lir.iting inan the reactivity insertion f rom an isothermal cooldown due to spacial aeaghting. However, conse rvatisms employed in the reactiviti weighting used in the RETRAN-02 temperature feedback model, and in the develcpment of the isotherv.a1 k-ef fective versus temperature curve, more than corupensate for the non-conservatisms introduced by not explicitly modeling tne cooldown asymmetrically. The reactivity versus temperature curve was generated assuming virtually [ This assumption results in an overprediction(of feedranex medel the reactivity inse}rtion.The RETRAN-02 temperature -
} The ;ombination of these conservatisms ensures a conservative reawtivity insertion during the cooldown.
(37) Locating :no stuck rod close to the faulted loop results in increased " power peaxing, but also results in a reduced inlet temperature. Since these effects have an opposite mpact on DNBR margin, what is the effect of assumino the stuck rod is located away from the faulted loop?
Response
C
}Thecombinationofincreased power peaking and low flow ensures a conservative DNBR evaluation even though the inlet temperature in the f aulted loop is lower than the inlet temperature in the intact loop. _
7 4 (30) In view of the large axial paaking that occurs in the steamline break, what error is introduced in the axial heat flux by the numoer of axial tones in RETRAN-02?
Response
rn .. . { l The errer ir.tr:cuted in the axial heat fbxfromusing( in t r.e EETPAH-02 core mocol ;s udged :: ce relatively small for the } , folicwing reasons: - e t (39) The themal mixing approach res21ts in a less bottom peaked power distribution, wnich is nonconservative for the steamline break with the loss et offsitt power. How is this nonconservatism accounted for?
Response
=
9 l9 l
1 \ (40) In the cycle-specifi: the return to criticality wnich will be performed::teamline creax analysis descrice the e not cuberatical at* ine limiting conditions. if the reload core is Responce: The cycle-specific reactivity check will reveal whetner the reload core will be suocritical with respect to the core previcusly analyzed. If the reload core is determined not the transient will be reanalyzed. to be subcritical, it will be redesigned or Therefote, in Revision 1 of CPC-NE-3001 the last paragrapn of Section 5.5 has been revised to read: If the cycle-specific reactivity check Mows the reactor to te suberitical wAth respect to the core assumed in the existing licensing basis analysis, includAng a stuck rod, then the response predicted by the system analysis bounds the reload core. If the relcad core is not suberitical at these conditions, two approaches are available to obtain acceptable steam line break analysis results: redesien the reload core or reanaly e the transient. (41) In the drcpped rod analysis, the most negative temperature coefficients provide the maximum positive reactivity insertion f rom the cooldown cut also provide the greatest negative feedback during thn power excursion. What for theare the bounding dropped-rod analysis temperature at BOC andcoefficients EOC? (most positive / negative) Responoe: As stated on p. 6-7 of OpC-NE-3001, the least negative (most positive) value of MTC is assumed for BOC, MOC, and EOC. also assumed. These values The least negative DTC is were selected following sensitivity studies which clearly showed that minimizing the negative feedback during the power excursion produced the limiting results. This effect is more significant than the positive reactivity insertion from the cooldown, - since of Bank surplus D. positive reactivity remains available from the withdrawal overshoot. It is the withdrawal of Bank D that causes the power (42) In the droppari the excore detector response been accounted for? rod analysis, has the effect of a w
Response
yes. 6-5, As stated on pp. 6-6 and 6-8 of DpC-NE-3001, and as shown in Figure the effect conservatively modeled.of power tilt on the dropped rod analysis has been Due to the importance of Bank D withdrawal on this event, the condition analyzed -wt.s System. a failure in the Rod Control This signal (lowest power quadrant) postulated failure results in the minissa excore flux the Rod Control System. Since appearing to be the core average power in the indicated power is low, Bank D withdrawal is maximized and worsens the power overshoot. Figure 6-5 shows_that the tilt factor input to the analysis bounds predicted tilts. It is noted that for large summed dropped rod worths, there is no tilt _ ~~ _ -_-__ -- --- -- -- -
l l since a sy=etric pattern af ::cc .TJst os droppec to ::tain norths Cf
- nat magnitude.
>42) The contro; system response nas a major e:fect en tne c:nsequences of a droppuo rod.
How tre : nservative control system para.eters selected anien bounc all Duke Power plants and cycles including uncertaintics? Respcnse: 1 [ There are no significant dif ferences between the control systems in the four Duke Power units in the context of dror+ed r:d transients. The control systems are moceled assuming nominal setpoints. No mayor conservat systems. we assumptions are made which dominate the etfects of control The Rod control System is assumed to fall in a manner which results in the lowest excore flux signal (rather than the nignest) being used as an indication of core power level. The limiting power tilt is assumed. This power oversnoot. assumption maximizes withdrawal of Banx D and worsens the The Rod Control System rod withdrawal stops are also assumed to ce inoperative. These rod stops terminate rod withdrawal snenever any of the following conditions exist:
- Auctions,ered high NI power greater than 103%
Bank D at'223 stepo withdrawn
- Urgnnt f ailure alarm for Bank C (prevent: Band D witndrawal if a Bank D rod drops}
This assumption allows the power overshoot to continue above IC3% power. It is noted that this failure modo can be related to the failure that inputs the lowest excore power signal to the Rod Control System. Many other process parameters are used in the control logic. However, the impact cf the two ma]cr assumptions discussed above is so dominant as
- o relegate other considerations to insignificance. -
(44) When are the pre-drop and post-drop thermal boundary conditions used in the dropped rod power peaxing analysis?
Response
=
MD
- taa l
l l \
> 5) ;; a full-core neutronics ca.cutation perferred f:r it.e :: ppeo-rod event? !! not, ciscuss the ef fects of tnis approx:: nation.
- esponse:
Yes. Tull-core three-dirrens tenal calculations were perferred to calculate dropped rod worths, racial peaxing factors, execre detector responses and axial snapes for the reference droppeo rod analysis. (46) how are the offects of crossfi-v Detween adjacent channels treated in the VIPRE-01 model usad in the droppec-rod event? channel modelDNBR The{dr]opped-red the snown in figure 6-1 of CPC-NE-3001 is utilired in analysis. tnat described in Reference 3. This( } channel model is identical to In Ref erence 3, model development, model
- ustification, code options, and input selectial havu been described in detail.
The moceling of crossflow between add;acent channels in the droppec-rod analysis is identical to that in Helo.tence 3.
REFERENCES:
1. Septemoer 2, 1984 letter form C. O. Thomas (NRC) to T. W. Schnat: (UCRA),
" Acceptance for Referencing ot EPRI ND-1850-CCM, 'RETRAN-02 A Program for Transient Thertnal-Mydraulic Analysit of Complex Fluid Flow Systems,'"
Enclosure 2 to the attached SER. 2. Enclosure to October 19, 1988 letter from A. C. Th6dani (NRC) to R. Furia (UGRA), " Acceptance for Referencing Topical Report EPRI NP-1850-CCM-A, Revisions 2 and 3 Regarding RETP.AN-02 MOD 003 and MOD 004." 3.
" Thermal Hydraulic Transient Analysis Methodology," DPC-NE-3000, Ouke Power Company, Revision 1. Febraury 20, 1990.
4. Letter from C. E. Rossi (NRC) to J. A. Blaisdell (UGRA), " Acceptance for Ref erencing of Licensing Topical Report, VIPRE-01: A Thermal-Hydraulic Analysis Code for Reactor Cores, EPRI-NP-2511-CCM, Vol.1-5, May 1986."
- 5. " Duke Power Company McGuire and Catawba Nuclear Stations Core Thermal-Hydraulic Methodology Using VIPRE-01," DPC-NE-2004, Decemoer 19B8.
__ __ ____ _ _ - - - - - - - - - - - - - - - - - - ~}}