ML20024D916
| ML20024D916 | |
| Person / Time | |
|---|---|
| Site: | Oyster Creek |
| Issue date: | 08/31/1979 |
| From: | GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20024D913 | List: |
| References | |
| NEDO-24195-A-04, NEDO-24195-A-04-ERR, NEDO-24195-A-4, NEDO-24195-A-4-ERR, NUDOCS 8308080422 | |
| Download: ML20024D916 (60) | |
Text
[ .3 NUCLEAR ENERGY BUSINESS OPER ATIONS o GENERAL ELECTRIC COMPANY SAN JOSE, CALIFORNI A 95125
'C , GENER AL $ ELECTRIC APPLICABLE TO:
NEDO-24195 PUBLICATION NO.
79NED288 ER N M And ADDMDA T' I' E' NO' SHEET TITLE GENERAL ELECTRIC RELOAD 4A no, FUEL APPLICATION FOR February 1983 DATE OYSTER CREEK NOTE: Correct allcopies of the applicable ISSUE DATE August 1979 publication as specified below.
REFERENCES (SECTION, PAGE IfJSTRUCTIONS ITEM PA R AG R APH. LIN E) (COR RECTIONS AND ADDITIONS) 01 Pages viii, viii-a Replace with new pages viii, viii-a and x.
and x 02 Page 4-1 Replace with new page 4-1.
03 Pages 5-5, 5-6, Replace with new pages 5-5, 5-Sa, 5-6, 5-7, 5-8 5-7 an.4 5-8 and 5-8a.
04 Page 5-10 Replace with new page 5-10. 05 Page 5-32 Replace with new page 5-32. 06 Page 5-39 Replace with new pages 5-39 and 5-39a. 07 Pages 5-40 and Replace with new pages 5-40, 5-41, 5-41a and 5-41b. 5-41 08 Page 5-42 Replace with new page 5-42. 09 Page 5-51 Replace with new pages 5-51 and 5-51a. 10 Page 5-58 Replace with new pages 5-58 and 5-58a. 11 Page 5-59 Replace with new page 5-59. 12 Pages 5-59a and Insert new pages 5-59a and 5-59b. 5-59b 13 Pages 5-75a, 5-75b Replace with new pages 5-75a, 5-75b and 5-75c. and 5-75c l 14 Page 5-75d Insert new page 5-75d. 15 Pages 5-76a, 5-76b Replace with new pages 5-76a.1, 5-76a.2, 5-76b and 5-76c and 5-76c. I 16 Page 5-76d Insert new page 5-76d. 17 Page 5-76e Insert new page 5-76e. fy 18 Pages 5-77a, 5-77b Replace with new pages 5-77a, 5-77b and 5-77c. )
~(/ and 5-77c 8308080422 830722 PAGE Of 2 PDR ADOCK 05000219 P PDR
e . . l NUCLEAR ENERGY BUSINESS OPER ATIONS o GENERAL ELECTRIC COMPANY SAN JOSE, CALIFORNIA 95125 C) GENER AL $ ELECTRIC APPLICABLE TO: l'UBLICATI NNO. NEDO-24195 79NED288 ERRATA And ADDENDA T. l. E. NO. TITLE 4A NO. FUEL APPLICATION FOR OATE February 1983 OYSTER CREEK NOTE: Correct allcopies of the applicable ISSUE OATE AUE" t 19I9 publication as specified below. REFERENCES INSTRUCTIONS ITEM (SECTION PAGE PAR AGR APH, LINE) (CORRECTIONS AND ADDITIONS) 19 Page 5-78 Replace with new page 5-78. 20 Pages 5-78a and Insert new pages 5-78a and 5-78b. 5-78b 21 Page 5-79 Replace with new page 5-79. 22 Page 5-79a Insert new page 5-79a.
) 23 Page 5-80 Replace with new pages 5-80, 5-81 and 5-82.
24 Pages A-1/A-il, Replace with new pages A-1/A-il, A-2, A-3, A-3a, A-2, A-3, A-7, A-7, A-9 and A-ll. A-9, and A-11 25 Appendix B Insert new Appendix B. w) PAGE 2 Of 2
NED0-24195 LIST OF ILLUSTRATIONS (Continued) O Figure Title Page 5-6 Increase in Heat Generated as a Function of Distance 5-65 from the Gap 5-7 Idealization of Flux Spike 5-66 5-8 if' as a Function of a Gencral Profile 5-67 5-9 Central Peak Heat-Flux Distribution (TS No. 65) 5-67 5-10 Central Peak Heat-Flux Distribution (NS No. 76) 5-68 5-11 Critical Quality Versus Boiling Length for Tests 65 and 76 5-68 5-12 Damping Coefficient Versus Decay Ratio (Second Order Systems) 5-69 5-13 Accident Reactivity Shape Functions at 20*C 5-70 5-14 Accident Reactivity Shape Functions at 286*C 5-71 5-15 Doppler Reactivity Coefficient vs Average Fuel Temperature as a Function of Exposure and Moderator Condition 5-72 5-16 Scram Reactivity Function for Cold Startup 5-73 5-17 Scram Reactivity Function for Hot Startup 5-74 [ 5-18a Water Level Inside the Shroud and Reactor Vessel - Pressure Following a 4.66 ft2 Recirculation Discharge . Line Break, Emergency Condenser Failure 5-75a q 5-18b Water Level Inside the Shroud and Reactor Vessel Pressure Following a 1.0 ft2 Recirculation Discharge - es Line Break, Emergency Condenser Failure 5-75b _ E es 5-18c Water Level Inside the Shroud and Reactor Vessel j Pressure Following a 0.3 ft2 Recirculation Discharge ., i Line Break, Emergency Condenser Failure 5-75c q
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l 5-18d Water Level Inside the Shroud and Reactor Vessel Pressure Following a 0.10 ft2 Recirculation Discharge l , Line Break, Emergency Condenser Failure 5-75d q m 5-19a.1 Peak Cladding Temperature Following a 4.66 ft Recirculation Line Discharge Break, Emergency Condenser ., Failure (LBM) (E >20,000 mwd /ST) 5-76a.1 q 5-19a.2 Peak Cladding Temperature Following a 4.66 ft l Recirculation Line Discharge Break, Emergency Condenser g Failure (LBM) (E <1000 mwd /ST) 5-76a.2 _g l _ 5-19b Peak Cladding Temperature Following a 1.0 f t 2 _
. Recirculation Line Discharge Break, Emergency Condenser N t
Failure (LBM) 5-76b _2 k i 5-19c Peak Cladding Temperature Following a 0.3 f t 2 l Recirculation Line Discharge Break, Emergency Condenser _R i Failure (LBM) 5-76c _S vili - L j
NEDO-24195 LIST OF ILLUSTRATIONS (Continued) O Page () Figure Title
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2 5-19d Peak Cladding Temperature Following a 0.3 ft Recircu-lation Line Discharge Break, Emergency Condenser Failure (SBM) 5-76d 5-19e Peak Cladding Temperature Following a 0.10 ft 2 Recirculation Line Discharge Break, Emergency Condenser Failure (SBM) 5-76e 5-20a Fuel Rod Convective Heat Transfer Coefficient at the - Highest Power Axial Node for a 4.66 ft 2 Recirculation - Line Discharge Break (LBM) 5-77a _ 9/79 5-20b Fuel Rod Convective Heat Transfer Coefficient at the 1/80 Highest Power Axial Node for a 1.0 ft2 Recirculation 9779 Line Discharge Break (LBM) 5-77b _ 5-20c Fuel Rod Convective Heat Transfer Coefficient at the 1/80 Highest Power Axial Node for a 0.13 ft2 Recirculation - Line Discharge Break (LBM) 5-77c _ 9/79 g s 5-21a Normalized Core Average Inlet Flow Following a Maximum Recirculation Line Discharge Break (4.66 f t Z) 5-78 5-21b Normalized Core Average Inlet Flow Following a Maximum Recirculation Line Discharge Break (1.0 f t2) 5-78a 5-21c Normalized Core Average Inlet Flow Following a Maximum (~] V Recirculation Line Discharge Break (0.3 f t2) 5-78b 5-22a Minimum Critical Power Ration Following a 1.0 ft 2 Recirculation Line Discharge Break 5-79 5-22b Minimum Critical Power Ratio Following a 0.3 ft 2 Recirculation Line Discharge Break 5-79a 5-23 Normalized Power Versus Time 5-80 5-24 Peak Cladding Temperature Versus Break Area 5-81 9/79 l 5-25 Oyster Creek Control Rod Withdrawal Sequence 5-82 i l l l '% viii-a
NED0-24195 LIST OF TABLES (Continued) Table Title Pm 5-12 LOCA Analysis Figure Summary 5-59 5-13 Single Failures Considered in the Oyster Creek LOCA Analysis 5-59 _ 5-14a MAPLHGR Versus Average Planar Exposure k* (Fuel Type: P8DRB239) 5-59a - 5-14b MAPLHGR Versus Average Planar Exposure (Fuel Type: P8DRB265L) 5-59a 5-14c MAPLHGR Versus Average Planar Exposure k" (Fuel Type: P8DRB265) 5-59b 5-15 Single Failures Considered in the Oyster Creek LOCA Analysis 5-59b _ O O X
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- 4. STEADY-STATE HYDRAULIC MODELS V
Core steady-state thermal-hydraulic analyses are performed using a model of the reactor core. This model includes hydraulic descriptions of orifices, lower tieplates, fuel rods, fuel rod spacers, upper tieplates, the fuel channel and core bypass flow paths. The orifice, lower tieplate, fuel rod spacers, upper tieplate and, where applicable, holes in the lower tieplate are hydraulically represented as being separate, distinct local losses of zero thickness. The fuel channel cross section is represented by a square section with enclosed area equal to the unrodded cross sectional area of the actual fuel char.nel. For thermal-hydraulic analysis, General Electric uses one of the fuel assembly designs documented in Reference 4-9 in place of any non-GE assembly present in the core. The General Electric-supplied bundle which has a fuel rod geometry and enrichment most like that for a non-General Electric-supplied bundle is used in place of that non-General Electric-supplied bundle. Chosen replacement bundle designations for the reference cycle are given in the reference cycle supple-ment. Additional information is presented in Appendix B. . /, . () The flow distribution to the fuel assemblies and bypass flow paths is calculated on the assumption that the pressure drop across all fuel assemblies and bypass flow paths is the same. This assumption has been confirmed by measuring the flow distribution in boiling water reactors (References 4-1, 4-2, 4-3). The components of bundle pressure drop considered are friction, local, elevation and acceleration (see Subsections 4.1 through 4.4). Pressure drop measurements made in operating reactors confirm that the total measured core pressure drop and calculated core pressure drop are in good agreement. There is reasonable assurance, therefore, that the calculated flow distribution throughout the core is in close agreement with the actual flow distribution of an operating reactor. An iteration is performed on flow through each flow path (fuel assemblies and bypass paths), which equates the total differential pressure (plenum to plenum) across each path and matches the sum of the flows through each path to the total core flow. The total core flow less the control rod cooling flow enters j the lower plenum. A fraction of this passes through various bypass paths documented in Subsection 4.5. The remainder passes through the orifice in the fuel support (experiencing a pressure loss), where more flow is lost through 4-1
NED0-24195 The rod-by-rod R-factor distributions used for the bounding statistical
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analysis are summarized in Reference 5-25. The basis for the additive con-stants used to determine the P8x8R R-f actor is documented in Reference 5-2. Results of the analyses show that at least 99.9% of the fuel rods in the core are expected to avoid boiling transition if the MCPR for P8x8R reloads is 1.07 or greater. 5.2 MCPR OPERATING LIMIT CALCULATIONAL PROCEDURE A plant-unique MCPR operating limit is established for General Electric-supplied
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fuel to ensure that the fuel cladding integrity safety limit for that fuel is not violated for any moderate frequency transient. This operating requirement is obtained by addition of the absolute, maximum CPR value for the most limit-ing transient from rated conditions postulated to occur at the plant to the fuel cladding integrity safety limit. Core-wide rapid pressurization events (Turbine Trip w/o Bypass, Load Rejection (vO w/o Bypass, Feedwater Controller Failure) are analyzed using the system model documented in References 5-27 and 5-28. The ODYN code contains a one-dimensional representation of the reactor core which is coupled to the recir-culation and control system model. The integrated model is based on one-dimensional reactor kinetics, multi-noded thermal-hydraulic and heat transfer relationships, and mechanical kinetic equations of the equipment. ODYN con-tains a refined reactor core description and a detailed steamline model to simulate pressure dynamics during a transient. Improvements made in ODYN are { documented in Reference 5-29. The improvements consist of a new model used in the calculation of mass flux through the safety / relief valves and the replacement of the coupled five-equation thermal-hydraulics model with a similar model in which the momentum equation is decoupled. The improved ver-sion yields the same results as the previous version, within expected numeri-cal uncertainty (Reference 5-30). Further updating of ODYN was performed to correct a minor coding error (Reference 5-35). For the slower core-wide tran-sients, Loss of Feedwater Heating is analyzed using either the steady-state / three-dimensional BWR Simulator Code (Reference 5-7) or the REDY Transient model (v) 5-5
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f-wg (Ref erences 5-3, '5-4 and 5-5), as described in Reference 5-31. A worse, usually
\-'# maximum, power condition is assumed with thermally limited fuel conditions. The i philosophy with respect to using the equipment performance components of the 4
transient models for design and safety evaluations is to consider the performance 1 of key components at their adverse tolerances. Circuitry delays in the reactor protection system as well as other key equipment circuit delays are assumed at the maximum specified values. The speed of all the control rod drives follow-ing a scram is assumed to be at the plant technical specification value. Field data have shown considerable. conservatism in this key component performance. The setpoints for the safety / relief valves both in the safety and relief func-tion and for pressure scram are assumed at their specified limits. The assumed ' setpoints used in the analysis are given in Table 5-3. Other equipment per-formance such as relief and safety valve opening characteristics, recirculation hk pump drive train inertia, and main steam line isolation valve closing times 1 are all assumed to be at adverse tolerances. End-of-cycle (EOC) conditions for nuclear data are used (except where specific
- exposure dependent evaluations are performed, Subsection 5.2.2) to provide a
[) varying level of conservatism associated with core exposure aspects. The nuclear data which are re-evaluated for each reload analysis are the scram reactivity function, void reactivity function and Doppler reactivity function. These parameters are calculated within the ODYN code for each cycle, as described in References 5-27 and 5-28. The basis for using EOC Doppler and ' void reactivity functions in the analyses is outlined in Reference 5-6. A discussion of the above significant parameters follows. 1 i O 5-Sa
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NEDO-24195
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Scram reactivity is the worth of control rods as a function of time or posi-V tion following the scram signal. The scram reactivity insertion is normally lowest at the EOC (all rods out condition) because there are no stubbed rods to insert negative reactivity more quickly than the remaining blades of the control rod bank. The scram reactivity function is related to dynamic per-formance when expressed (as plotted) in the form of Ak/kS. While the scram reactivity characteristic is uniquely determined during evaluations of rapid pressurization events, the scram reactivity used for slow transients (Loss of Feedwater Heating) is a typical reload curve (see Reference 5-25). A factor of 0.8 is applied for use in the analyses for added conservatism. This curve is applied generically for slow transients because these transients are insensitive to the rate of negative reactivity insertion during a scram. This insensitivity occurs because the rate of positive reactivity insertion from a reduction in core void fraction is very slow compared to the time required for complete control rod insertion. _ The void reactivity coefficient is an important parameter, not only in trans-q ient analysis, but also in core stability. The core average void coefficient \ must be negative; however, it must not be so negative as to yield such a strong positive reactivity feedback during void collapse events that core and vessel limits are threatened. Conversely, events with void increase must produce suf-ficient negative feedback to maintain operation within safety limits. A trans-ient index which is used to assess the void reactivity characteristics is the dynamic void coef ficient. This parameter is defined as the core physics void coef ficient multiplied by the average full power void fraction and divided by the delayed neutron fraction. The presence of U-238 and, ultimately, Pu-240 contributes to yield a strong negative Doppler coefficient. This coefficient provides instantaneous nega-tive reactivity f eedback to any fuel temperature rise, either gross or local. The magnitude of the Doppler coefficient is not dependent on gadolinium posi-tion or concentration in any bundle because gadolinium has very little effect on the resonance group flux or on U-238 content of the core. The core physics p U 5-6
NEDO-24195 ( Doppler coefficient is divided by the delayed neutron fraction to define a dynamic Doppler coefficient (K D
), which most effectively correlates the dynamic response of the plant to the Doppler reactivity feedback.
Additional conservatism is introduced into the analysis to account for biases and uncertainties in the derivation of the nuclear data and its application to m The conservatism factors used to account for these m the REDY transient model. biases and uncertainties are summarized in Table 5-4. These values, in effect, represent a maximum level of conservatism in the transient analysis upon which compliance to existing safety limits is generally based. The reactor core behavior for General Electric-supplied fuel during the rod withdrawal error transient is calculated by doing a series of steady-state three-dimensional coupled nuclear thermal-hydraulic calculations using the three-dimensional BWR simulator (Reference 5-7). This approach assumes that the transient is very slow such that there is sufficient time for heat transfer and void redistribution to equilibrate and also that the neutron n (.) flux and heat flux are in equilibrium with each other. This calculation is achieved by maintaining a constant eigenvalue via increasing core average power as a control rod is withdrawn incrementally. Descriptions of the transient events are given in Subsection 5.2.1. Inputs to l these transients which are specific to the Oyster Creek plant are given in Tables 5-3 and 5-5. Transient inputs which are specific to the reference cycle are~ o e given in the reference cycle supplement. Because the transient model establishes 3 operating conditions, only licensing basis values are given in Table 5-5. l Actual values used in the analyses will be within the tolerances shown in the table. The transient descriptions given in Subsection 5.2.1 are used as a basis for the typical analyses performed for plant reloads. Oyster Creek l analyses will dif fer in certain aspects f rom the typical calculational pro-cedure due to the utility selected margin improvement option of exposure-l dependent limits. A description of this option and its ef fect upon the transient analysis is given in Subsection 5.2.2. ATWS pump trip is also p assumed in the analysis of Oyster Creek. L] i 1 5-7
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. The operating limit MCPR for General Electric fuel for rapid transients is calcu-lated by using the SCAT computer program (Reference 5-8). Inputs to this pro-gram consist of the transient analysis results, steady-state flow distribution (from Section 4), bundle power, axial distribution, and gap conductance. The axial power distribution used in the analyses is given in Table 5-6. Nonvarying plant initial conditions for the GETAB analysis are given in Table 5-7. Reload-dependent plant initial conditions for the GETAB analysis and the resulting reload " g MCPR operating limit for General Electric fuel in the reference cycle are given in , N the reference cycle supplement. The initial MCPR
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assumed for transient analyses is usually greater than or equal to the GETAB operating limit. Figure 5-3 illustrates the effect of the initial MCPR on transient ACPR for a typical BWR core. This figure indicates that the change in ACPR is approximately 0.01 for a 0.05 change in initial MCPR. Therefore, nonlimiting GETAB transient analyses may be initiated from a MCPR below the operating limit because the higher operating limit MCPR more than offsets the increase in ACPR for the event. This may also be applied to limiting transients if the dif ference between the operating limit and the initial MCPR is small (0.01 or 0.02). Densification power spiking is not considered in establishing l the MCPR operating limit. Justification for this is presented in Sub-section 5.2.3.
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The deterministic ACPR value which results from ODYN/ Improved SCAT evaluations (for all rapid pressurization transients) must be adjusted such that a 95/95 ACPR/ICPR licensing basis is calculated (i.e., 95% probability with 95% con-fidence that the safety limit will not be violated). The SER which describes these requirements and procedures is given in Reference 5-32. The method of applying these adjustment factors is given in Reference 5-33. Each utility has the choice of operating under one of the following basic options: k e Option A - Under Option A, an NRC-imposed factor of 1.044 is applied to the MCPR for each event to account for code uncertainties. o Option B - Under Option B, General Electric - developed adjustment factors (Reference 5-32) are applied to the ACPR/ICPR ratio. These factors are a result of a statistical analysis of transient response based on an improved CRD scram insertion time distribution. The 5-8
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factors are a function of-plant type and event. Since these adjust- .m ment factors take credit for conservatism in the scram speed assumed for the transient analysis, each plant operating under Option B must demonstrate that its scram speed is within the dis-N tribution used-in the statistical analysis. This conformance pro-cedure is described in Reference 5-32 and will be incorporated in the technical specifications of each plant operating under Option B. The adjusted MCPR values are given in the Supplemental Reload Submittal. The operational limit MCPR must be increased for low flow conditions. This is because, in the BWR, power increases as core flow increases, which results in a corresponding lower MCPR. If the MCPR at a reduced flow condition were at the 100% power and flow MCPR operating limit, a sufficiently large inadvertent flow increase could cause the MCPR to decrease below the Safety Limit MCPR given in Subsection 5.2.1. Therefore, the required operating limit MCPR is increased at reduced core flow rates by a -flow f actor, Kg , such that: O Required MCPR =K
- MCPR Operating Limit Operating Limit - f @ 100% core flow The flow factor (K ), as a function of the core flow rate, the flow control f
mode, and (in the manual flow control mode) the maximum possible core flow n 5-8a
NEDO-24195 will be a slow, exponential' curve. An analysis.to the linear approximation, however, will be conservative, since it overpredicts the power level for any given exposure. In Ref erence 5-9, evaluations were made to .40% power level points on the linear curve. The results show that the pressure margins from the limiting pressurization transient and the MCPR operating limits exhibit a larger margin for each of these points than the EOC full power, full flow case. MAPLHGR limits for the full power, rated flow case is conservative for the coastdown period, since the power will be decreasing and rated core flow will be main-tained - (Subsection 5.5) . Therefore, it can be concluded that the coastdown operation beyond full power operation is conservatively bounded by the , analysis at the EOC conditions. 'In Reference 5-34 this conclusion is confirmed 2 s for analyses done with ODYN. - 5.2.1 Transient Descriptions Eight nuclear system parameter variations can pose potential deleterious effects to the Nuclear Steam Supply System. The parameter variations are as follows: , (1) Nuclear system pressure increase - threatens to rupture the reactor coolant pressure boundary from internal pressure. Also, a pressure increase collapses the voids in the moderator. This causes an inser-tion of positive reactivity which may result in exceeding the fuel cladding safety limits. (2) Reactor vessel water (moderator) temperature decreases - results in an insertion of positive reactivity as density increases. Positive reactivity insertions threaten the fuel cladding safety limits because of higher power. (3) Positive reactivity insertion - is possible from causes other than nuclear system pressure or moderator temperature changes. Such reactivity insertions threaten the fuel cladding safety limits because . of higher power. 5-10
-NEDO-24195 Approximate Event Elapsed Time (g) APRM 115.7% power signal scrams reactor (conser-
. vative; in startup mode, APRM scram would be operative + IRM). r (h) Scram. terminates accident. < 5 see To limit the worth ofLthe rod which could be dropped, the rod worth minimizer system (RWM) is used below 10% power to enforce the rod withdrawal sequence. - n The RWM is programmed to-follow the control rod sequences shown in Figure 5-25. .m So The rod drop accider.t design limit restricts peak enthalpies in excess of 280 cal /gm for any possible plant operation or core exposure. 5.5.1.2 Model Parameters Sensitivities s v Although there are many input parameters to the RDA analysis, the resultant peak fuel enthalpy is most sensitive to the following input parameters: (1) cteady-state accident reactivity shape function; i (2) total control rod reactivity worth; i (3) maximum interassembly local power peaking factor.p P . P p repre-sents the maximum local peaking factor normalized over the four bundles surrounding the dropped control rod. Mathematically, i i 4 BP 1 P = Max P i i = 1,2,3,4 (5-14) F L , 4 , I , }[ BP m j=1 s
- O 5-32
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R NEDO-24195 _2R 5.5.2 Loss-of-Coolant Accident (3 b This analysis of the Oyster Creek Nuclear Generating Station loss-of-coolant accident (LOCA) is provided to demonstrate conformance with the ECCS accept-ance criteria of 10CFR50.46. The objective of the LOCA analysis contained herein is to provide assurance that the most limiting break size, break loca-tion and single failure combination has been considered for the plant. The documentation contained in this section is intended to satisfy these require-ments. i The general description of the General Electric (GE) LOCA evaluation models is contained in Reference 5-18. Applicability and approval for pre-pressurized reload fuel are given in Reference 5-25. Model changes are described in References 5-20 and 5-21, which were approved by the USNRC (Reference 5-19). The analysis utilizes the short-term thermal-hydraulic model (LAMB) and the transient critical power model (SCAT) in addition to the long-term thermal-hydraulic model (SAFE) and the core heat-up model (CilASTE). (~\ This LOCA analysis differs from previous BWR/2 LOCA analyses in the following (_/ ways:
- 1. Core flow coastdown and core depressurization are now calculated with the LAMB and SCAT codes, as opposed to no modelling of coastdown.
In order to use the LAMB computer code for a non-jet pump plant, LAMB n co inputs were developed to execute the code for a recirculation line ;7 break. These inputs reflect the geometry of the non-Jet pump plant. The LAMB code only allows for two recirculation pump loops; therefore, the five loops were modeled such that the intact loops are combined into one loop, and the broken loop is modeled as the second loop.
- 2. For the large break region, dryout time is now calculated using the SCAT code, as opposed to assuming a set dryout time.
The Oyster Creek LOCA analysis reported here was performed as an independent,
,s - self-contained analysis similar to that performed as a lead plant analysis.
b - 5-39
NED0-24195 7-s, . 5.5.2.1 Input to Analysis V l A list of the significant input parameters to the LOCA analysis is presented in Table 5-10. 5.5.2.2 LAMB ANALYSIS i This code is used to analyze the short-term blowdown phenomena for large postulated pipe breaks (breaks in which nucleate boiling is lost before the water level drops and uncovers the active fuel). The LAMB output (core flow as a function of time) is input to the SCAT code for calculation of blowdown heat transfer. The LAMB results presented are: e Core Average Inlet Flow Rate (normalized to unity at the beginning of the accident) following a Large Break. 4 5.5.2.3 SCAT ANALYSIS Q N This code completes the transient short-term thermal-hydraulic calculation for large breaks. The GEXL correlation is used to track the boiling transition in 4 time and location. The post-critical heat flux heat transfer correlations are , built into S Jr, which calculates heat transfer coefficients for input to the core heatup code (CHASTE). The SCAT results presented are: e Minimum Critical Power Ratio following a Large Break, e Convective Heat Transfer Coefficient following a Large Break. 5.5.2.4 SAFE ANALYSIS This code is used primarily to track the vessel inventory and to model ECCS performance during the LOCA. The application of SAFE is identical for all 5-39a J
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NEDO-24195 1R J;;
,, break sizes. This code is used during the entire course of the postulated (j accident. SAFE calculates reactor system pressure, ECCS flows (which are pressure dependent), and hot fuel node uncovery time. Reflooding times are not calculated in non-jet pump analysis. All peak cladding temperatures are turned over by action of core spray heat transfer.
The SAFE results presented are: e Water Level inside the Shroud. e Reactor Vessel Pressure. 5.5.2.5 CHASTE ANALYSIS This code is used, with suitable inputs from the other codes, to calculate the fuel cladding heatup rate, peak cladding temperature, peak local cladding oxidation, and core-wide metal-water reaction for large breaks. The fuel 73 model in CHASTE considers transient gap conductance, clad swelling and rupture, -' and metal-water reaction. The empirical core spray heat transfer and channel wetting correlations are contained in CHASTE, which solves the transient heat transfer equations for the entire LOCA at a single axial plane in a single fuel assembly. Iterative applications of CHASTE determine the maximum permissible planar power required 2 to satisfy the 10CFR50.46 acceptance criteria. 7 The CHASTE results presented are: e Peak Cladding Temperature and Peak Local Oxidation versus Time. e Peak Cladding Temperature and Peak Local Oxidation versur Break Area. _ e Peak Cladding Temperature and Peak Local Oxidation versus Planar $ s Average Exposure for the most limiting break size. - e Maximum Average Planar Heat Generation Rate (MAPLHGR) versus Planar r~, Average Exposure for the most limiting break size.
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5-40
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NED0-24195 9/79 _ 1/80
,5.5.2.6' Methods In the following sections, it will be useful to refer to the methods used to analyze large and small breaks. These methods are described below.
i 5.5.2.6.1. Large Break Methods (LBM) LAMB / SCAT / SAFE / CHASTE Break sizes: 0.3 ft2 < A < DBA. Heat transfer coefficients: Figures 5-20a, 5-20b, and 5-20c. SCAT is omitted for those break sizes which do not result in significant positive coastdown flow (40% DBA to DBA). Under these circumstances, the duration of nucleate boiling following the recirculation line break is calculated by CHASTE using the GE dryout correlation, which is based on no-flow conditions. This correlation relies on experimental data from many different test-assemblies with various axial power shapes and different geometries whose purpose was to investigate the no-flow dryout phenomenon. The Ellion pool boiling heat transfer corre- Q l lation is used from the time that the decay heat transfer from the dryout E I correlation ceases to apply until the time that the hot fuel core node is g-g (-} uncovered.- 4 5.5.2.6.2 Small Break Methods (SBM) SAFE / CHASTE Break sizes: A < 0.3 ft . Heat transfer coefficients: nucleate 2 ? boiling prior to core uncovery,' core spray, and 1000 Btu /hr-ft *F after rods wet. Peak cladding temperature and peak local oxidation are calculated in CRASTE. 5.5.2.7 Break Spectrum Calculations i For convenience in describing the LOCA phenomena, the break spectrum has been separated into two regions: (1) small breaks, and (2) large breaks. The potentially limiting single active failures considered are given in Table 5-13. The selection of the transition break between Large Break Methods (LBM) and . . Small Break Methods (SBM) (i.e., 0.3 ft2) is made in order to achieve similar 3 - '~ coastdown flow, core uncovery, and rated spray time characteristics to those 5-41 F
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NED0-24195
~
exhibited at the transition break for jet-pump plants. The transition break has been analyzed with both the large and small break methods with the same single failure to allow comparison between the methods. The analysis of the transition break is shown in Figures 5-19c, 5-19d, and 5-22b. The large break region is defined as that portion of the break spectrum between
.the transition-break and the DBA. The DBA is defined as the complete severence of the largest pipe in that portion of the system which yields the highest peak
-cladding temperature when the most limiting single failure is~ assumed. The most limiting single failure in this region is the Emergency Condensor failure.
5.5.2.8 Large Break Analysis In the large break region between the 60% and DBA break sizes, there exist
-two single failures (ADS and EC) which produce break spectrums which are nearly equal. This is due to the fact that in this break region both failures have no impact on the hot node uncovery time and the time of rated core spray. The DBA break is the limiting break in the large break region (0.3 < A < DBA). This O is due to the earliest core uncovery and no coastdown flor characteristics for the DBA. The DBA break with the Emergency Condenser single failure produces a.
PCT of'2*F higher than the DBA with the ADS single failure. Therefore, the EC single failure is considered the limiting failure. In the region between 60% DBA and 0.3 ft2 the EC failure is more limiting than the ADS failure by an even larger margin. I In the range of 40% DBA to DBA, credit is not taken for coastdown flow due to the rapid decrease in core inlet flow calculated by the IMB code for this ) region. In the region between 40% and 0.3 ft2 break, coastdown flow has a ! significant impact in delaying boiling transition, and, therefore, the results f of IMB and SCAT are used to calculate PCT. I 5.5.2.9 Small Break Analysis I
- In this region, the vessel depressurizes relatively slow. The EC is the most severe single failure in this region compared to the ADS failure because the l
- % 1 ADS single failure results in a later hot node uncovery time. The break i
5-41a l l
NEDO-24195 spectrum in this region shows that the 0.1 ft break is the most limiting break O, size. The limiting break in this region shifts from the previous 0.13 ft to 2 2
-0.10 ft because of the use of the SBM. The previous worst break of 0.13 ft was' determined by the LBM without credit for coastdown core flow. .The dif-ferences in these models cause the time between uncovery and rated core spray to have a greater effect than the higher decay heat at time of uncovery for 0.13 ft2 The time between uncovery and rated spray is approximately 15 sec-
.onds greater for the 0.'1 ft2 break-compared to the 0.13 ft2 break. This larger time between these two events is responsible for the 0.10 ft 2break being more limiting than all other break sizes in this region.
5.5.2.10- Break Spectrum Conclusions A summary of the analytical.results is given in Table 5-11. Table 5-12 lists the figures provided for this analysis. The MAPLHGR values for each General Electric fuel type available to Oyster Creek are given in Table 5-14. Peak' cladding temperatures and oxidation fractions are also given in this table. The results of the complete break spectrum show that two break sizes may be lm limiting depending on fuel exposure. The exposure-dependent limiting break is due 'o the combined effects of gap conductance and rod internal pressure. For fuel in the low exposure region ($1000 mwd /ST), the DBA break size is limiting because of the PCT limit of 2200 F. Because of a gradual improvement in gap conductance after 1000 mwd /ST, the small break (0.1 f t ) becomes limiting due to PCT at the mid-exposur'e range (5000 mwd /ST to 15,000 mwd /ST). At higher exposure (120,000 mwd /ST), as fission gas release increases, the DBA break l becomes limiting due to peak local oxidation fraction. The. break spectrum summary curve is shown in Figure 5-24. This figure gives the maximum PCT and the maximum local oxidation, over all exposures, as a function of break size. O 5-41b l
NEDO-24195
- 9/79 1/80 .- 5.5.3 thin Steamline Break Accident Analysis 2/83
%) The analysis of the main steamline break accident depends on the operating thermal-hydraulic parameters of the overall reactor (such as pressure) and over-all factors affecting the radiological consequences (such as primary coolant activity). Insertion of reload fuel will not change the radiological conse-quences of this event. Analytical results supporting this are presented in c) co Reference 5-26. _ 2T 5.5.4 Loading Error Accident Calculational Methods One of the events which has been evaluated in BWR safety analysis reports is fuel bundle loading error. A loading error in the core configuration is defined as: (1) a General Electric fuel bundle in an improper location (mislocation), or (2) a General Electric fuel bundle in an improper orientation (i.e., mis-oriented - rotated 90* or 180*). -~ Proper orientation of fuel assemblies in the reactor core is readily verified by visual observation and assured by verification procedures during core loading. Five separate visual indications of proper fuel assembly orientation exist: (1) The channel fastener assemblies, including the spring and guard used to maintain clearances between channels, are located at one corner of each fuel assembly adjacent to the center of the control rod. (2) The identification boss on the fuel assembly handle points toward the adjacent control rod. (3) The channel spacing buttons are adjacent to the control rod passage area. 0\ V 5-42
NED0-24195 (] '5-15 R. C. Stirn, C. J. Paone, and R. M. Young, Rod Drop Accident Analysis for V Large Boiling Water Reactors, Licensing Topical Report, July 1972 (NEDO-10527, Supplement 1). 5-16 R. C. Stirn, C. .J. Paone, and J. M. Haun, Rod Drop Accident Analysis for Large Boiling Water Reactors, Addendwn No.' 2, Exposure Cores, Ltcensing Topical Report, January 1973 (NEDO-10527, Supplement 2). 5-17 Fuel Densification Effects on General Electric Boiling Water Reactor Fuel, August 1973, Supplement 6 to NEDM-10735. 5-18 Analytical Model for Loss-of-Coolant Analysis in Accordance uith 10CFRSO Appendix K, January 1976 (NEDE-20566-P and NED0-20566) . 5-19 Letter, K. R. Coller (NRC) to G. G. Sherwood (GE), " Safety Evaluation for General Electric ECCS Evaluation Model Modifications", dated April 12, 1977. 5-20 Letter, A. J. Levine (GE) to D. F. Ross (NRC) dated January 27, 1977,
" General Electric (GE) Loss-of-Coolant Accident (LOCA) Analysis Model Revisions - Core Heat Code CHASTE 05".
5-21 Letter, A. J. Levine (GE) to D. B. Vassallo (NRC), dated March 14, 1977,
" Request for Approval for Use of Loss-of-Coolant Accident (LOCA) Evalu-ation Model Code REFLOOD05".
5-22 Letter, N. G. Trikouros (GPU) to J. F. Kilty (GE), dated November 13, 1975, R d No. S&L-3305, "0yster Creek Nuclear Generating Station Revised Single g Failure LOCA Analysis". - 5-23 Letter, R. E. Engel to D.'G. Eisenhut (NRC), " Fuel Assembly Loading Error", November 30, 1977. 5-24 Letter, D. G. Eisenhut (NRC) to R. E. Engel (GE), MFN-200-78, May 8,1978. 5-25 " General Electric Standard Application for Reactor Fuel", NEDE-240ll-P*. 5-26 Attachment to Letter, S. Bartnoff to A. Giambusso, "0yster Creek Nuclear o Generating Station Docket Number 50-219 Loss-of-Coolant Accident R Analysis Re-evaluation Additional Information," April 28, 1975. 5-27' " Qualification of the One-Dimensional Core Transient Model for BWR's", "" October 1978 (NEDO-24154, Vol. 1 and 2). 5-28 " Qualification of the One-Dimensional Core Transient Model for BWR's", $ October 1978 (NEDE-24154-P, Vol. 3). N 5-29 Letter, J. F. Quirk (CE) to P. S. Check (NRC), "0DYN Improvements," September 25, 1981. _
- Reference refers to the revision of NEDE-24011-P-A which is approved by the NRC as of the date the specific analysis is initiated.
5-51
. . = . . _ _ -_ - - _
'NEDO-24195
~
5-30 Letter, J. F. Quirk (GE) to T. P. Speis (NRC), "0DYN Improvements", () October 13, 1981. 5-31 Letter, R. E. Engel (GE) to T. A. Ippolito (NRC), " Change in GE Methods for Analysis of Cold Water Injection Transients", September 30, 1980. 1 5-32 Letter, R. C. -Tedesco (NRC) to G. G. Sherwood (GE), " Acceptance for m Referencing General. Electric Licensing Topical Report NED0-24154/ -f; NEDE-24154P", February 4, 1981. 5-33 Letter, R. H. Buchholz (GE) to P. S. Check (NRC), "0DYN Adjustment Methods for Determination of Operating Limits", January 19, 1981. s 5-34 Letter, R. E. Engel (GE) to T. A. Ippolito (NRC), "End of Cycle Coastdown Analyzed With ODYN/TASC", September 1, 1981. 5-35 Letter, H. C. Pfefferlen (GE) to D. G. Eisenhut (NRC), " Correction of ODYN Errors", June 8, 1982. .
.O i
l i f l l 5-51a
i NEDO-24195 Table 5-10 ~S s SIGNIFICANT INPUT PARAMETERS TO THE LOSS-OF-COOLANT ACCIDENT ANALYSIS Plant Parameters:
' Core Thermal Power 1969 Mut, which corresponds to 102% of rated power ,_ n m
Vessel Steam Flow 7.4 x 106 lbm/h, which corresponds , to 102% of rated power Vessel Steam Dome Pressure 1020 psig
- Recirculation Line Break Area 4.66 ft2 (DBA), 1.0 ft2, 0.3 ft for Large Break -mN
((hk-Recirculation Line Break Area 0.3 ft , 0.10 ft - m for Small Break m- I
' Number of Drilled Bundles None SE Fuel Parameters:
Peak Technical Specification - Linear Heat Design Axial Initial Fuel Bundle -Generation Rate Peaking Minimum Fuel Type Geometry (kW/ft) Factor CPR* P8DRB239 P8x8R 13.4 1.57_ 1.30 ', g
- P8DRB265L P8x8R 13.4 1.57 1.30- {; EI P8DRB265H P8x8R- 13.4 1.57 1.30
*To account for the 2% uncertainty in bundle power required by Appendix K, the ,_
SCAT calculation is performed with an MCPR of 1.2745 (i.e., 1.30 divided by 1.02) for a bundle with an initial MCPR of 1.30. 4 I e w A 1 1
^
4 i 5-58
HED0-24195 Table 5-11 -' Sl} DIARY OF BREAK SPECTRUM RESULTS
.)
e Break Size Core-Ulde e Location Peak Local Metal-Water e Single Failure PCT (*F) Oxidation (%) Reaction (%) e 4.66 ft2 2200 17.0
- e Recirculation Discharge o Emergency Condenser (LBM) e 1.0 ft2 2064 *
- e Recirculation Discharge e Emergency Condenser (LBM) e 0.3 ft 2123 *
- es n
e Recirculation Discharge c? g s e Emergency Condenser (LBM) N -, e 0.3 ft 2042 *
- I 1
\_ ' e Recirculation Discharge o Emergency Condenser (EBM) e 0.10 ft 2200
- 0.40 e Recirculation Discharge ,_
e Emergency Condenser (SBM) - R
*Less than limiting case %2 5-58a
NED0-24193~
/
Table 5-12 LOCA ANALYSIS FIGURE
SUMMARY
Large Break Model Small Break Model DBA 1.0 ft2 0.3 ftZ 0.3 ft2 0.1-ft2 Water Level Inside Shroud and 5-18a 5-18b 5-18c 5-18c 5-18d Reactor Vessel Pressure vs. Time
-Peak Cladding Temperature and 5-19a.1/ 5-19b 5-19c 5-19d 5-19e h m
Peak Local Oxidation vs. Time 5-19a.2 Heat Transfer Coefficient vs. 5-20a 5-20b 5-20c -- -- n Time - m, N Normalized Core Average Inlet 5-21a 5-21b 5-21c -- --- Flow vs. Time Minimum Critical Power Ratio -- 5-22a 5-22b -- -- vs. Time Normalized Power vs. Time 5-23 -- -- -- -- O Peak Cladding Temperature and 5 '
;7 Peak Local Oxidation vs. Break ,_
Om Area - Table 5-13 SINGLE FAILURES CONSIDERED IN THE OYSTER CREEK LOCA ANALYSIS
'(Reference 5-22)
Single Break Location Failure Available Systems i Recirculation Line 1 EC 2 CS + 0 EC + 5 ADS 1 ADS 2 CS + 1 EC + 4 ADS i Feedwater and 1 EC 2 CS + 1 EC + 5 ADS Steam Lines 1 ADS 2 CS + 2 EC + 4 ADS i Core Spray Line 1 EC 1 CS + 1 EC + 5 ADS 1 ADS 1 CS + 2 EC + 4 ADS EC = Emergency Condenser CS = Core Spray () ADS = Automatic Depressurization System > 5-59
NEDO-24195 i Table 5-14a MAPLHGR VERSUS AVERAGE PLANAR EXPOSURE Plant: Oyster Creek Fuel Type: P8DRB239 Average Planar s Exposure MAPLHGR PCT 0xidation (mwd /t) (kW/ft)_ (*F) Fraction 200 9.5 2198 0.095 4 I 1000 9.5 2198 0.095 5000 9.5 2194 0.085 10,000 9.5 2193 0.085 15,000 9.5 2194 0.085 20,000 9.0 2092 0.169 + 25,000 8.9 2048 0.169 30,000 8.9 2049 0.170 35,000 8.9 2050 0.170 40,000 8.8 2049 0.170 j t m
- Table 5-14b 2 s
MAPLUGR VERSUS AVERAGE PLANAR EXPOSURE 1 Plant: Oyster Creek Fuel Type: P8DRB265L l Average Planar Exposure MAPLHGR PCT 0xidation 4 (HWd/t) (kW/ft) ("F) Fraction I 200 9.5 2198 0.095 1000 9.5 2198 0.095 5000 9.5 2198 0.086 ! 10,000 9.5 2198 0.086 15,000 9.5 2198 0.086 20,000 9.0 2081 0.170 25,000 8.8 2049 0.169 30,000 8.8 2045 0.170 35,000 8.8 2049 0.170 i 40,000 8.8 2050 0.170
- (::)-
4 ^ 5-59a p.- , -, .g - , , . ~ . . , . _ . . . .,w . .,.,_,.,_.r__ - . , _ . rm#.__,m .. . . , _ , , _ . . , , , y,,_,_. ,,_,..,.c_,..,- ,_ m._ . . _ . , ,
NEDO-24195 Table 5-14c p MAPLilGR VERSUS AVERAGE PLANAR EXPOSURE ( Plant: Oyster Creek Fuel Type: P8DRB265 Average Planar Exposure MAPLilGR PCT 0xidation (mwd /t) (kW/ft) (*F) Fraction 200 9.5 2198 0.095 m 1000 9.5 2198 0.095 h 5000 9.5 2198 0.086 10,000 9.5 2198 0.086 15,000 9.5 2198 0.086 20,000 8.9 2078 0.170 25,000 8.8 2050 0.170 30,000 8.8 2044 0.166 35,000 8.8 2049 0.170 40,000 8.8 2050 0.170 , m 5-59b
lI l llIi tI t[ -i ' . .!. i ;ii6' - l ' i 9' g,N* $v.mW O 0- 3aEy< 0 6 g 0 4 0 3 0 2 0 1 00 0 _ - - - 4 2 f t 6 _ 6 4 _ a _ g n _ i w oe ll NG _ l r l u ol E N KI AD _ Fi a TO 0 eF T I O DF L F _ F a - 0 3 r ur EE A A se RR T_ B ss en CR re OO NF _ Pd n l o E _ eC s sy T O N _ ec V n I _ rg e or _ t e cm O D e 0 0
)
( E c o s aE R , e r. U O _ 2 M d a ne R I T ar H S _ d ue B E D _ on I ri S hL N I L E _ E R S eg e U h r V E L _ S S E t a h R ec _ P M d s i i E sD _ T S Y e - 0 0 n I n S 1 o _ l i et va el L u c rr _ ei t c ae _ WR _ a _ 8 1
- , 5 9 - - - - e O r 0 0 0 O 0 0 O u 0 0 0 o 0 0 g 2 0 8 c 4 2 i .
1 1 F O 1$ $- awn $ i vi ? i;
- II Il! <! 1 i1 ,1' Il ijtl ;] .i1i!] lIi!1I. lll'.
i O O O i l 1200 60 i i 1 4 1000 - 50 i rJOTE: NO CHEDIT TAKEN l FOR REFLOODING J 1 I SYSTEM PRES $t,RE i 200 - 40
.s
$ 2 z
i m e 3 w M i u h TAF a 30 3 b 1 4 u m 800 mm. - --------------------- R4 me , cr E ,_ ta ~ i a i g- LEVEL INSIDE SHROUD g j j O .I > 400 - - 20 , BAF 2@ - - to i I o l 0 ' 9 100 200 300 400 500 ' ! i TIME (sec) ( i Figure 5-18b. Water Level Trside the Shroud and Reactor Vessel Pressure following a 1.0 ft i Recirculation Discharge Line Break, Emergency Condenser Failure
1 O O O 4 I I r l i i i ] 1200 60 l i I 1000 NOTE: NO CREDIT TAKEN - 50 4 FOR REFLOODING 1 ,i i 1 800 -
- 40
$ 2 2 M
) ] SYSTEM PRESSURE t I c ] w 8I 3 I U $ 800 m TAF > N i O E
,_ -- - -- --.--.------ - au 30 $
E g N ! a. w *
- C3 1
1 w M. w r 4 3 w 3: l 400 - 20 BAF m----- - 4 t 200 - LEVEL INS!DE SHROUD - 10 t i O l l ' - O 1 0 100 200 300 ' 400 j TIME (sec) I Figure 5-18c. Water Level Inside the Shroud and React.or Vessel Pressure Following a 0.3 ft 2 . Recirculation Discharge Line Break, Emergency Condenser Failure _ l 4 i
J O o O I % 2 1200 - NOTE: NO CREDIT TAKEN FOR REF LOODING - 50 i 1000 1
~ # ,
g i
* - LEVEL INSIDE SHROUD SYSTEM PRESSURE _
m )
- O i j - o
]- w d b I ) x r_ _ _ _ _ .; _ g i i $ __ e % u l g g-a ______ p 4 3 u k ' W 3: 3w - 20 40; - BAF -
- - I i
! - 10 4 ' 200 - I I o I I 300 400 ! o 200 0 100 j TIME bc) i i ' i Figure 5-18d. Water Level Inside the Shroud and Reactor Vessel Pressure Following a 0.10 ft _ l Recirculation Discharge Line Break, Faergency Condenser Failure i
O O O 50 2500 - C gm - e_ FUEL 6 g ROD z r ROD RE-WET o 4 PERFORATES
- z 5 < m 9 C o
2 x w u, w F SPRAY mo % 8 y 1500 - 00 m COOLING O
- h o HIGH POWER AXIAL
( " 4 PLANE UNCOVERED k d ROD RE WET E 3, _
- to FUEL ROD
^
ONSET OF BOILING TR ANSITION
' ! ' 'O O 10,000 100,000 1.0 to 100 g,000 TIME (sec)
Figure 5-19a.1. Peak Cladding Temperature and Peak Local Oxidation Following a 4.66 ft Recirculation Line Discharge Break, Emergency Condenser Failure (LEM) (E > 20,000 mwd /ST) _
O O O 1 3000 : -- - 60 - 2500 - - 50 i D 2000 - ROD RE. WET - 40 SPR AY COOLING G 5 P 1 E E < 2 vi s 9 rn [ $ 1500 - - 30 5 y8 ri i o 24 M*
"5 O
=
3 2 , 5 \_ HIGH POWER AXIAL M l U PLANE UNCOVERED $
< 1000 - -
20 E ONSET OF BOILING ROD RE-WET g _ TRANSITION - 10 I I I I o o 1.0 10 100 1,000 10.000 100,000 TIME (sec) Figure 5-19a.2. Peak Cladding Temperature Following a 4.66 ft Recirculation Line Discharge Break, Emergency Condenser Failure (LBM) (E< 1000 mwd /ST) _ m
O O O 60 3000 2s00 - 50 ROD RE-WET
=
- c 20m -
2
= -
tr z I R 9 5 t
! # 1sm -
- 20 5 Y 2
$ 8 ~ is
.I a a 2 o M w
?
w 4 m, 5 SPRAY COOLING e j y y 20 5 g iOm - 1 ROD RE-WET l d HIGH POWER AXIAL PLANE UNCOVERED - 10 500 - ONSET OF BOILING i TRANSITION l 0 0 10 100 1,000 10,000 1.0 TIME (sec) 2 Figure 5-19b. Peak Cladding Temperature and Peak Local Oxidation Following a 1.0 ft Recirculation i Line Discharge Break, Emergency Condenser Failure (LO!) - i,
O O O 60 3000 1 1 4 t 2600 -
- 50 l
i I I SPRAY ROD R E-WET l E COOLING - 40 i o, 2000 - Q z 3 Q Q 5 c. o 1 5 5
- 30 0 Z I Y 1500 -
4 3 y t'1 C { u $ g %. o y i - 8 a u a i o 5 4 us 0 e I o HIGH POWER AXIAL ' La I M PLANE UNCOVERED
- 20 6
a. 1000 - 1 ROD R E-WET I 500 - 10 { ONSET OF BOILING TRANSITION i I ! ! O ! O 1.0 to 100 1,000 10,000 1 TIME (sect 9 Figure 5-19c. Peak Cladding Temperature and !'eak Local Oxidation Following a 0.3 ft' Recirculation Line Discarge Break, Emergenc Condenser Failure (LBM) - t
O O O k 9 .i 2000 60 l i 2500 - - 50 i i i l 1 ROD RE-WET 3 ( 2000 - - 40 ] E 7
~
3 SPRAY l Q COOLING g i b $ k e , g 1500 - - 30 g 2
- O J M L
E O 5 O 8I A. O .J w 5 x o e o HIGH POWER AXIAL $ a. g $ u
$ PLANE UNCOVERED y i w 1000 - AND ONSET OF -
20
! BOILING TRANSITION I
ROD R E-WET 500 - - 10 l I I O O 1.0 10 100 t,000 10,000 TIME (sect Figure 5-19d. Peak Cladding Temperature and Peak Local Oxidation Following a 0.3 ft' Recirculation Line Discharge Break, Emergency Condenser Failure (SIN) _ m
O O O
, so 2.oo -
.0 SPRAY i
COOLING 2000 - e E - o_ ae g ROD RE-WET { a H e H
< 4 E O g V
E 1500 - 30 E M 2 O g 4 s s . F
" 0 z
b k N 0 s = W 0 e ka
- u 1000 -
e j - 20 5 l ROD RE-WET i 500 - HIGH POWER AXIAL - [ - 10 PLANE UNCOVERED a AND ONSET OF '. BOILING TR ANSITION i 0- 0 l 1.0 10 100 1,000 t o,og;
- TIME (sec) l F1gure 5-19e. Peak Cladding Temperature and Peak Local Oxidation Following a 0.10 ft Recirculation Line Discharge Break, Emergency Condenser Failure (SBM) _
t l I a
NED0-24195 _ 5 10 O ONSET OF BOLLING TRANSITION
=
I 4 - 10 e 4 o d 103 - a e 2 I w , 8 - e i
% 10 4
E HIGH POWER AXtAL PLANE UNCOVERED I ~ 10 I I I I l 3o 0 0 10 20 30 40 50 60 TIME (sec) Fuel Rod Convective Heat Transfer Coefficient at the High O Figure 5-20a. Power Axial Node for a (4.66) f t 2 Recirculation Line Discharge Break (LBM) _ 5-77a
NED0-24195 l 5 10 O 4 _ ONSET OF BOILING 10
/ TRANSITION ac 4
0 L 103
~
4 b c 5 c w~ M
$ E 8
5 O n z 4 10 2 - HIGH POWER AXIAL 6 g / PLANE UNCOVERED e i 10 - i l i j I I O 10 20 30 40 50 00 TIME (sec) Figure 5-20b. Fuel Rod Convective Heat Transfer Coefficient at the Highest Power Axial Node for a 1.0 ft2 Recirculation Line Discharge O Break (LBM) _ 5-77b
NEDO-24195 O 10 10* -
# ,7 NSETO OF BOILING
{ p TRANSITION L b 5 103 _ z w 5 i w 8 3 HIGH POWER AX1AL
$ PLANE UNCOVERED m O 5 2
- e N
g10 E E I - 10 t I I I I I i f i 0 10 0 to 20 30 40 50 80 70 80 TIME (sec) Figure 5-20c. Fuel Rod Convective Heat Transfer Coefficient at the High Power Axial Node for a 0.3 ft Recirculation 2 Line Discharge Break (LBM) _ 5-77c
i O O O , 1.0 i l l 1 i i 4 t 0.6 - i E l 3 i u. l w ' E O O O w 5 2: l a m i y 1 E8 mi s E 5 % h l < c I 0.2 - g i l i i 1 i
-o.2 -
3 I I i j 0 1 2 3 4 5 l TIME AFTER BREAK (seconds) Figure 5-21a. Normalized Core Average Inlet Flow Following a Maximum Recirculation Line Discharge Break (4.66 ft2) _ i i
i O O O J .i I 1.0 1 l i i 0.6 - E i 3 l W c I O Z O m N O j' u y N O I b CD
]
4 $ b i I m 2 c-i s e c , O TWINDOW j (G = 100,000 lb/hr-ft ) IN 1 1 a 1
\
i r - l t i'
-0.2 l l 1 O 8 16 24 TIME AFTER BRE AK (secondd rigure 5-21b. Normalized Core Average Inlet Flow Following a Ifaximum Recirculation Line Discharge Break (1.0 ft2) 1 1
T< dOtN Ycw D 9~C O n o - i t a l
\
2 u 3 c Mt> i r t c p e R m u Hg M m i yG x g T t a M a
' g I n i
\ 4 w 2 o
- l l
)
k o m F o c w e s o ( l K F f A ) O . E R B R t2 et l f n A: E T F I e0 3 A g( E a 6 M rk 1 I T ea _ ve Ar . B e M re og C r a I dh ec
- zs ii lD a
me rn oi NL
\ 8 c
1 2 5 _ e _ r u g i F O 0 l \ 0 6 2 1 1 O O A s u5 0 %341 Oz wh@ s ll, ,.]li i'
O O O i l
~
2.0 4 l l , NOTE 1: CPR = 1 AT SPACER 2 1 1.8 - 9.29 kvulft AVERAGE PLANAR LINEAR HEAT GENERATION RATE (APLHGR): 82 PERCENT 1.6 - OF MAXIMUM AVERAGE PLANAR LINEAR HEAT GENERATION RATE (MAPLHGR) O P NOTE 2: CPR = 1 AT SPACER 3 ' < 1139 APLHGR ' E 1.4 ' 100 PERCENT OF MAPLHGR w Z (n M l .J U N < y O } e o 1.2 b g h d w i , x ~ u - o NOTE 2
]
E 1.0 -
- E T WINDOW G = 100,100 lb/hr-ft 23 j OH 1 - IN 0.8 - '
, 0.6 -
0.2 - e i e e i e i i s 0 1 2 3 4 5 6 7 8 9 10 TIME (sec) l I i Figure 5-22a. Minimum Critical Power Ratio Following a 1.0 ft2 Recirculation I Line Discharge Break - i
O O O I I 5.0 4.5 - window l j (GIN* I "
- O I 4.0 -
l 1 1 3.5 - l 9
, E 1 a:
-! 5 30 - t E i 2 a j w
~
E i $_ 2.5 a 4
- u t
- N co o
i e w m
. 5 j g2.0 - C e
1
? w i E a :E 1.5 -
l i L 1.0 - i 4 j 0.5 - 1
- 1 I I ! ! I j 0
- O 5 10 15 20 25 30 t
TIME (sec) Figure 5-22b. Minimum Critical Power Ratio Following a 0.3 ft Recirculation Line Discharge Break
- ~
zm ! wb *v,
~NmW
- J O
0 0 0 0 1 0 i 0 0 1 0 e I 0 m 1 i T s u
) s c r e
s e ( T V N r E D e I w C o I o C P t A O R E T F d i e z A l E a _ MI m T r o N I 0 3 1 2 5 e r u g _ i F _ 1 l 0 1
- - - 0 2 o 0 0 8 s. 4 1 O O 0 0 cw*2 $ O O
Y=O il ; i ii i ::1' J!j4) : !
J O O O h SMALL BREAK METHOD LARGE BREAK METHOD M i i l i i e I 2200 - I E ' w : E
! R I i
y 2100 - 8 w m I I
' i E
3 i i 5m - i o i x g 2
< ~
11D V i E - i
; u 8 , w 8,
,,00 _
, e 15 8 9 2 j $ i x w y
i e- e 12.s 0 i - u 9
; I O 8
10.0 j I O h
! - 7.5 5
? z lii
~
l - 5.0 . a' I e i
; I l t i 1 1 I I I I i i I , l 0.07 0.08 0.10 0.12 0.14 0.16 0.18 0.2 0.3 1.0 4.66 2 ! 8REAK AREA (ft 3 I
Figure 5-24. Peak Cladding Temperature and Local Peak Oxidation Versus Break Area, _ Recirculation Discharge Break, Emergency Condenser Failure i, l
NEDO-24195 SEQUENCE A O~< 51 2 1 47 5 4 5 4 5 43 1 2 1 2 1 2 30 3 6 3 6 3 35 2 1 2 1 2 1 31 4 5 4 5 4 5 4 27 1 2 1 2 1 2 23 6 3 6 3 6 3 6 19 2 1 2 1 2 1 15 5 4 5 4 5 11 1 2 1 2 1 2 07 3 6 3 6 3 03 2 1 02 06 10 14 18 22 26 30 34 38 42 46 50 SEQUENCE 8 51 1 2 1 47 5 4 5 4 43 1 2 1 2 1 V 39 6 3 6 3 6 3 35 1 2 1 2 1 2 1 31 4 5 4 5 4 5 27 2 1 2 1 2 1 2 23 6 3 6 3 6 3 19 1 2 1 2 1 2 1 15 4 5 4 5 4 5 11 1 2 1 2 1 l 07 3 6 3 6 03 1 2 1 02 06 10 14 18 22 26 30 34 38 42 46 50 NOTE: l The maximum rod worth is determined by the Rod Drop Analysis, and the RWM is programmed to assure the l 280 cal /gm limit is met. These groups may be sub-i divided to assure compliance with this limit. Figure 5-25. Oyster Creek Control Rod Withdrawal Sequences 5-82 D
NEDO-24195 O
, APPENDIX A OYSTER CREEK REFERENCE CYCLE SUPPLEMENT August 1980 Revised February 1983 0
) i O A-i/A-il
NEDO-24195 ,/~] 2. CALCULATED CORE EFFECTIVE MULTIPLICATION AND CONTROL SYSTEM WORTH - NO VOIDS, 20*C (3.3.2.1.1 and 3.3.2.1.2) (/ BOC k eff Uncontrolled 1.111 Fully Controlled 0.946 Strongest Control Rod Out 0.987 R, Maximum Increase in Cold Core Reactivity with Exposure Into Cycle, Ak 0.000
- 3. STANDBY LIQUID CONTROL SYSTEM SHUTDOWN CAPABILITY (3.3.2.1.3)
Shutdown Margin (Ak) ppm (20*C, Xenon Free) 600 0.043
- 4. TRANSIENT ANALYSIS INPUTS (3.3.2.1.5 and 5.2)
EOC D h Void Coefficient N/A* (c/% Rgo) -6.48 / -8.10 , Void Fraction (%) 35.92 _$ Doppler Coefficient N/A (c/*F) -0.222/ -0.211 Average Fuel Temperature (*F) 1184 Scram Worth N/A (S) -37.64 /-30.11 Scram Reactivity vs Time Figure 2
- 5. CETAB TRANSIENT ANALYSIS INITIAL CONDITION PARAMETERS (5.2)
EOC _ Fuel Design P8x8R Ex8 Peaking factors (local, radial 1.20 1.28 and axial) 1.711 1.624 1.40 1.40 m m R-Factor 1.051 1.098 s Bundle Power (MWt) 5.744 5.462 3 87.89 89.97 Sundle Flow (10 lb/hr) [} Initial MCPR 1.32 1.30 _
*N = Nuclear Input Data A _= Used in Transient Analysis A-2
NEDO-24195
- 6. CORE-WIDE TRANSIENT ANALYSIS RESULTS (5.2.1)
\ss' ACPR Exposure 4 Q/A 20 Transient (mwd /t) (% NBR) (%) P8x8R Ex8 Figure "
Turbine Trip EOC 517 119 0.25 0.23 Figure 3 j{ without Bypass N Loss of 100*F BOC to E0C 116 115 0.13 Figure 4 Feedwater Heating Feedwater Controller Failure EOC 297 116 0.18 0.16 Figure 5 'hk
-N
- 7. LOCAL ROD WITHDRAWAL ERROR (WITH LIMITING INSTRUMENT FAILURE)
TRANSIENT
SUMMARY
(5.2.1) Limiting Rod Pattern: Figure 6 Reactor Rod Position ACPR MLHCR (kW/ft) Power (%) (Feet Withdraen) P8x8R P8x8R 104 4.0 0.08 16.56 7 x, 105 5.5 0.14 16.62
\/ 106* 6.0 0.16 16.62 107 9.0 0.18 16.62 108 9.5 0.18 16.62 109 10.0 0.18 16.62
- Indicates APRM rod block setpoint selected.
l
~
- 8. CYCLE MCPR VALUES (5.2) l l
l Non-pressurization Events Exposure Range: BOC to EOC P8x8R $ l c% Loss of Feedwater Heater 1.20 l Fuel Loading Error 1.25 t i Rod Withdrawal Error 1.23 1 ! f^N ( ./ A-3
NEDO-24195 Pressurization Events Exposure Range: BOC to EOC Option A Option B g P8x8R Ex8 P8x8R Ex8 e7 Turbine Trip w/o Bypass 1.38 1.36 1.33 1.31 Feedwater Controller Failure 1.30 1.28 1.23 1.21
- 9. OVERPRESSURIZATION ANALYSIS
SUMMARY
(5.3) s1 y Plant Transient (psig) (psig) Response , MSIV Closure 1271 1320 Figure 7 21 (No Scram) ( ()~x i I t I ("N 4 )
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NEDO-24195 , t !O i t i i 4 j i l APPENDIX B , t THERMAL HYDRAULIC MODEL FOR NON-GE FUEL , l (EXXON 8x8 TYPE VB) i i
- O I
L i l l 1 6 I . l l
- 9 B-1/B-il
- ._ . . . . . . . - _ = . _- . - - . . . . .. . -
NEDO-24195 1,
-A General Electric has developed a thermal hydraulic model for non-GE fuel LJ (Exxon 6x8 Type Vil) for use in transient analyses with a mixed core of GE and non-GE fuel. The thermal hydraulic model for the non-GE fuel design is based
.; on the geometry of the non-GE fuel and pressure drop data for the non-GE fuel . Other thermal hydraulic characteristics of non-CE fuel were assumed to be identical CE fuel. The geometry of the non-GE fuel is tabulated in Table 15-1 i and the thermal hydraulic model assumptions are listed in Table 15-2. I The data and assumptions.used to develop the model were provided by GPU. Nuclear. General Electric assumes no responsibility in determining the appropriateness i of the modeling of the non-GE fuel, the applicability of GE methods to non-GE i fuel or the usefulness of these results. 2 The transient results for non-GE fuel are reported in Appendix A along with the 4 CE fuel. The calculation methods used are described in Section 5.
- :o
- 1. .
t O t i i O 1 l B-1 1
NEDO-24195
,- Table B-1
(,) EXXON 8x8 FUEL
- Geometrical Dimensions
- 1. Fuel Cladding Dimensions
- a. Outside Diameter 0.5015 in,
- b. Inside Diameter 0.4295 in.
- 2. Non-Fueled Rods
- a. Outside Diameter 0.5015 in,
- b. Inside Diameter N/A
- c. Water Flow Rate at Rated Conditions N/A
- 3. Channel Dimensions
- a. Inside Dimension 5.278 in.
- b. Wall Thickness 0.080 in.
(_)
- c. Corner Radius 0.380 in.
- 4. Fuel Pellet Diameter 0.4195 in.
- 5. fuel Column Length 144 in.
- 6. Spacer Axial Location e Total 7 spacets
*0yster Creek FDSAR Amendment 76
,e m i B-2
l NEDO-24195 (~N; Table B-2 v THERMAL-HYDRAULIC MODEL ASSUMPTIONS The non-GE bundle was assumed to be identical to a GE 8x8 bundle with a Mod 1 spacer design with respect to the following:
- 1. Non-spacer local pressure loss coefficients.
- 2. Friction pressure loss coef ficients.
- 3. Two-phase multipliers.
- 4. Exposure dependent bypass leakage fraction.
- 5. Exposure dependent surface crudding.
The non-GE bundle was assumed to differ from a GE 8x8 bundle with a Mod 1 spacer design with respect to:
- 1. Bundle geometry.
- 2. Spacer loss coefficient.
i o t The non-GE bundle was modeled with a spacer loss coefficient of 0.31, which was calculated to best fit the bundle pressure drop data supplied by GPU Nuclear. f~h L) B-3/B-4}}