ML20151P233
| ML20151P233 | |
| Person / Time | |
|---|---|
| Site: | Fermi  |
| Issue date: | 04/30/1987 |
| From: | GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20151P196 | List: |
| References | |
| A-2666, A00-02666, MDE-56-0386, MDE-56-0386-R01, MDE-56-386, MDE-56-386-R1, NUDOCS 8808090273 | |
| Download: ML20151P233 (42) | |
Text
1 l
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l 9CE-5610386 Revision 1 l ORF No. A00-02666 FERMI-2 SINGLE LOOP OPERATION ANALYSIS APRIL 1987
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Prepared for DETROIT EDISON COMPANY ENRICO FERMI ATOMIC POWER PLANT UNIT 2 Prepared by GENERAL ELECTRIC COMPANY NUCLEAR ENERGY BUSINESS OPERATIONS SAN JOSE, CALIFORNIA 95125 3828283888?ljgt P
, EF-2 FSAR APPENDIX 15.C TABLE OF CONTENTS -
Pace 15.C RECIRCULATION SYSTEM SINGLE-LOOP OPERATION 15.C.1-1 15.C.1 INTRODUCTION AND
SUMMARY
15.C.1-1 15.C.2 MCPR FUEL CLADDING INTEGRITY SAFETY LIMIT 15.C.2-1 15.C.2.1 Core Flow Uncertainty 15.C.2-1 15.C.2.1.1 Core Flow Measurement During Single-Loop Operation 15.C.2-1 15.C.2.1.2 Core Flow Uncertainty Analysis 15.C.2-2 15.C.2.2 TIP Reading Uncertainty 15.C.2-4 15.C.3 MCPR OPERATING LIMIT 15.C.3-1 15.C.3.1 Abnormal Operational Transients 15.C.3-1 15.C.3.1.1 Feedwater Controller Failure - Maximum Demand 15.C.3-2 15.C.3.1.2 Turbine / Generator Trip 15.C.3-4 15.C.3.1.3 Summary and Conclusions 15.C.3-6 15.C.3.2 Rod Withdrawal Error 15.C.3-6 15.C.3.3 Operating MCPR Limit 15.C.3-8 15.C.4 STABILITY ANALYSIS 15.C.4-1 15.C.4.1 Phenomena 15.C.4-1 15.C 4.2 Compliance to Stability criteria 15.C.4-2 15.C.5 LOSS-OF-COOLANT ACCIDENT ANALYSIS 15.C.5-1 15.C.S.1 Break Spectrum Analysis 15.C.5-2 15.C.S.2 Single-Loop MAPLHGR Determination 15.C.5-2 15.C.S.3 Small Break Peak Cladding Temperature 15.C 5-3 15.C-1
EF-2 FSAR TABLE OF CONTENTS (Continued)
Page :
i 15.C.6 CONTAINMENT ANALYSIS 15.C.6-1 ,
15.C.7 MISCELLANEOUS IMPACT EVALUATION 15.C.7-1 15.C.7.1 A.iticipated Transient Without Scram Impact Analysis 15.C.7-1
- 15. C,. 7. 2 Fuel Mechanical Performance 15.C.7-1 15.C.7.3 Vessel Internal Yibration 15.C.7-2 15.C.8 REFERENCES 15.C.8-1 15.C-li
EF-2 FSAR
}ISTOFTABLES MUMBER TITLE PAGC 15.C.3-1 Input Parameters and Initial Conditions 15.C.3-9,10,11 15.C.3-2 Sequence of Events for Figure 15.C.3-1, Feedwater Controller Failure, Maximum Demand With Full Bypass and Reheater Flow 15.C.3-12 15.C.3-3 Sequence of Events for Figure 15.C.3-2, Turbine / Generator Trip With One-Half Bypass and Reheater Flow 15.C.3-13 15.C.3-4 Summary of Trar.sient Peak Value and CPR Results 15.C.3-14 15.C.3-5 Summary of CPR Results 15.C.3-15 O
15.C-tii
EF-2 FSAR
~
LIST OF FIGURES NUMBER TITLE PAGE 15.C.2-1 Illustration of Single Recirculation Loop Operation Flows 15.C.2-5 15.C.3-1 Feedwater Controller Failure - Maximum Demand, Full Bypass and Reheater Flow, 75% Power /60% Core Flow 15.C.3-16,17, 18,19 15.C.3.2 Turbine / Generator Trip One-Half Bypass and Reheater Flow, 75% Power /601 Core Flow 15.C.3-20,21, 22,23 15.C.5-1 Uncovered Time vs. Break Area - LPCI IV Failure 15.C.5-4 15.C.5-2 Water Level Inside the Shroud vs Time 15.C.5-5 After Break - DBA, Recirculation Suction Break, LPCI IV Failure, Single Loop Operation 15.C-iv il
- ._. . ~
~
EF-2 FSAR l
15.C RECIRCULATION SYSTEMS SINGLE-LOOP OPERATION I
15.C.1 INTRODUCTION AND
SUMMARY
Single-loop operation (SLO) at reduced power is highly desirable in the event recirculation pump or other component maintenance renders one loop inoperative. To justify sing'ie-loop operation, accidents and abnormal oper-stional transier.ts associated with power operatioM as presented in Sec-tions 6.2 and 6.3 and the main test of Chapter 15.0, were reviewed for the single-loop case with only one pump in operation. This appendix presents the results of the safety evaluation for the operation of the Enrico Fermi Atomic Power Plant Unit 2 (Fermi-2) with single recirculation loop inoper-able. This safety evaluation was performed for P8x8R fueled core. The transient safety analysis was performed on an initial cycle basis consistent with that for the FSAR. The analysis shows that the transient consequences for SLO (ACPR) are bounded by the full power analysis results given in the l FSAR. The conclusion drawn from the transient analysis results presented in this report is applicable to the reload cycle operation as well as initial cycle operation for Fermi-2. The conditions are those of continued opera-tion in the operating domain currently defined in Figure 4.4.5 of Chaptor 4
- up to maximum power of approximately 70% of rated.
l Increased uncertainties in the core total flow and Traversing In-Core Probe (TIP) readings resulted in a 0.01 incremental increase in the Minimum l
Critical Power Ratio (MCPR) fuel cladding integrity safety limit during ,
! single-loop operation. No increase in rated MCPR operating limit and no l change in the flow dependent MCPR limit is required because all abnormal operational transients analyzed for single-loop operation indicated that there is more than enough MCPR margin to compensate for this increase.in MCPR safety limit. The recirculation flow rate dependent rod block and scram setpoint equation given in Chapter'16 (Technical Specifications) are adjusted for one-pump operation.
Thermal-hydraulic stability was evaluated for its adequacy with respect to General Design Criteria li (10CFR50, Appendix A). It is shown fAat this stability criterion is satis (led during SLO. It is further shown that the 15.C.1-1
, EF-2 FSAR increase in neutron noise observed during SLO is independent of system stability margin. '
To prever,t potential control oscillations from occurring in the recir-culation flow control system, the operation mode of the recirculation flow control system must be restricted to operate in the manual control mode for single-loop operation.
The limiting Maximum Average Planar Linear Heat Generation Rate (MAPLHGR) reduction factor for single-loop operation is calculated to be 0.90.
The impact of single-loop operation on the FSAR specifications for containment response, including the containment dynamic loads, was evalu-ated. It was r";nfirmed that the containment response under SLO is within the present design values.
The impact of single-loop operation on the Anticipated Transient Without Scram (ATWS) analysis was evaluated. It is found that all ATWS acceptance criteria are met during SLO.
The fuel thermal and mechanical duty for transient events occurring during SLO is found to be bounded by the fuel design bases. The Average Power Range Monitor (APRM) fluctuation should not exceed a flux amplitude of
- 15% of rated and tha core plate differential pressure fluctuation should l not exceed 3.2 psi peak to peak to be consistent with t'he fuel rod and assembly design bases.
A recirculation pump drive flow limit is imposed for SLO. The highest drive flow that meets acceptable vessel internal vibration criteria is the drive flow limit for SLO. The pump speed at Fermi-2 should be limited to 75% of rated under single-loop operating conditions.
l l
15.C.1-2
EF-2 FSAR 15.C.2 MCPR FUEL CLADDING INTEGRITY SAFETY LIMIT Except for core total flow and TIP reading, the uncertainties used in the statistical analysis to determine the MCPR fuel cladding integrity safety limit are not dependent on whether coolant' flow is provided by one or two recirculation pumps. Ur:ertainties used in the two-loop operation analysis are documented in the FSAR. A 65 core flow measurement uncertainty has been established for single-loop operation (compared to 2.5% for two-loop operation). As shown below, this value conservatively reflects the one standard deviation (one sigo ) accuracy of the core flow measurement system documented in Reference 15.C.8-1. The random noise component of the TIP reading uncertainty was revised for single recirculation loop oper9 tion to reflect the operating plant test results given in Subsection 15.C.2.2. This revision resulted in a single-loop operation process computer effective TIP uncertainty of 6.8% for initial cores and 9.1% for reload cores. Comparable two-loop process computer uncertainty values are 6.3% for initial cores and 8.7% for reload cores.
The net effect of these two revised uncertainties is a 0.01 incremental increase in the required MCPR fuel cladding integrity safety limit.
15.C.2.1 Core Flow Uncertainty !
l 15.C.2.1.1 Core Flow Measurement During Single-Loop Operation The jet pump core flow measurement system is calibrated to measure core flow when both sets of jet pumps are in forward flow; total core flow.is the !
sum of the indicated loop flows. For single-loop operation, however, some inactive jet pumps will be backflowing (at active pump speeds above approxi- 3 mately 40%). Therefore, the measured flow in the backflowing jet pumps must !
be subtracted from the measured flow in the active loop to obtain the total core flow. In addition, the jet pump coefficient is different for reverse flow than for forward flow, and the measurement of reverse flow must be '
modified to account for this difference.
l l
1 15.C.2-1 '
EF-2 FSAR In single-loop operation, the total core flow is derived by the follow-ing formula: '
[TotalCoreI ,[ActiveLoop\-C
( Flow j iIndicated Flow 4 (IndicatedFlowj The coefficient C (=0.95) is defined as the ratio of "Inactive Loop True Flow" to "Inactive Loop Indicated Flow". "Loop Indicated Flow" is the flow measured by the jet pump "single-tap" loop flow sumers and indicators, which are set to rsad forward flow correctly.
The 0.95 factor was the result of a conservative analysis to appropri-ately modify the single-tap flow coefficient for reverse flow.* If a more exact, less conservative core flow is required, special in-reactor calibra-tion tests can be made. Such calibration tests would involve: calibrating core support plate AP versus core flow during one-pump and two-pump opera-tion along with 100% flow control line and calculating the correct value of C based on the core support plate AP and the loop flow indicator readings.
15.C.2.1.2 Core Flow Uncertainty Analysis The uncertainty analysis procedure used to establish the core flow uncertainty for one-;; ump operation is essentially the same as for two-pump operation, with some exceptions. The core flow uncertainty analysis is described in Reference 15.C.8-1. The analysis of one-pump core flow uncer-tainty is sumarized below.
For single-loop operation, the total core flow can be expressed as
- follows (refer to Figure 15.C.2-1)
i WC*NA~ I
'The analytical expected value of the "C" coefficient for Fermi-2 it 0.84.
I 15.C.2-2
{ .
EF-2 FSAR Where:
j WC = total c re flow, ,
Wg = active loop flow, and j Wy = inactive loop (true) flow.
By applying the "propagation of errors" method to the above equation, the variance of the total flow uneartainty can be approximated by:
2 3 o* o* 1 o* t a l a* a*
( 1-a )2 W
C
- W sys
+ Wg rand
+M ( Wy rand Cj where:
og = uncertainty of total core flow; og = uncertainty systematic to both loops; og = random uncertainty of active loop only; og = random uncertainty of inactive loop only; o
c
= uncertainty of "C" coefficient; and a = ratio of inactive loop flow (Wy ) to active loop flow (W
A*
From an uncertainty analysis, the conservative, bounding values of og og og and oC are 1.6%, 2.6%, 3.5%, and 2.8%,
sys' A
- I rand rand respgetively. Based on the above uncertainties and a bounding value of 0.36 for "a", the variance of the total flow uncertainty is approximately:
- This flow split ratio varies from about 0.13 to 0.36. The 0.36 value is a conservative bounding value. The analytical expected value of the flow split ratio for Fermi-2 1s s 0.33.
15.C.2-3
EF-2 FSAR sc = (1..). .c.a.:= -(2.m. .< u u ' -
@.-.(2.-:
= (5.0%)'
When the effect of 4.1% core bypass flow split uncertainty at 12%
(bounding case) bypass flow fraction is added to the total core flow uncer-tainty, the active coolant flow uncertainty is:
c8 (5.0%): ' O.12 38 active = . (4*15)' =
(5*15)'
coolant + (1-0.12 ,
which is less than the 65 flow uncertainty assumed in the statistical analysis.
In sunenary, core flow during one-pump operation is measured in a conservative way and its uncertainty has been conservatively evaluated.
15.C.2.2 TIP Reading Uncertainty To ascertain the TIP noise uncertainty for single recirculation loop operation, a test was performed at an operating BWR. The test was performed at a power level 59.3% of rated with a single recirculation pump in opera-tion (core flow 46.3% of rated). A rotationally symetric control rod pattern existed during the test.
Five consecutive traverses were made with each of five TIP machines, giving a total of 25 traverses. Analysis of this data resulted in a nodal TIP noise of 2.85%. Use of this TIP noise value as a component of the process computer total uncertainty results in a one-sigma process computer total effect TIP uncertainty value for single-loop operation of 6.8% for initial cores and 9.1% for reload cores.
i 15.C.2-4
t.r-z P E M Q
s' CORI l- -
/
We w2 l
l N
C TOW Con nos Wg = Active Loop Flow , j h';
- Ina:tive Loop Flow l
l DETEli EDISDN 3LLUSTitAT]DN OF $ltlGLE RECIRCULATION LOOP* FIGURE OPERATION FLOWS 15.C.2 1 15.C.2-5
. , EF-2 FSAR 15.C.3 MCPR OPERATING LIMIT 15.C.3.1 Abnormal Operational Transients Operating with one recirculation loop results in a maximum power output which is about 305 below that which is attainable for two-pump operat. ion.
Therefore, the consequences of abnormal operational transients from one-loop operation will be consic4rably less severe than those analyzed for two-loop operation. For pressurization, flow increase, flow decrease, and cold water injection transients, the results presented in Chapter 15 bound both the thermal and overpressure consequences of single-loop operation.
The consequences of flow decrease transients are bounded by the full power analysis. A trip of the recirculation pump during single-loop operation is less severe than a two-pump trip from full power because of the reduced initial power level. Thus, the results presented in Chapter 15 bound the pump trip transient consequence of single-loop operation.
The worst flow increase transient results from a recirculation flow controller failure, and the worst cold water injection transient results from the loss of feedwater heating. For the former event, the K curve is f
derived from a postulated runout of both recirculation loops. This condi-tion produces the maximum possible power increase and kence maximum ACPR for transients initiated from less than rated power and flow. During operation with only one recirculation loop, the flow and power increase associr,ted with this failure with only one loop will be less than that associated with both loops; therefore, the K curve f deriv 6d with the two-pump assumption is l conservative for single-loop operation. "
The latter event, loss of feedwater heating, is generally the most severe cold water event with respect to increase in core power. This power increase is caused by positive reactivity insertion from increased core inlet subcooling and it is relat.ively insensitive to initial power level. A 15.C.3-1
EF-2 FSAR
- statistical loss of feedwater heating analysis using different initial power levels and other core design parameters concluded one-pump operation with lower initial power level is conservatively bounded by the existing full .
power two-pump analysis. Inadvertent restart of the idle recirculation pump has been analyzed in Chapter 15.4.4 and is still applicable for single-loop operation, i
In the following sections, results of the two most limiting pres-surization transients analyzed for single-loop operation are presented.
They are respectively:
- b. Turbine / Generator Trip (T/GT)
The plant initial conditions are given in Table 15.C.3-1, 15.C.3.1.1 Feedwater Controller Failure - Maximum Demand This event is postulated on the basis of a single failure of a master feedwater control device, specifically one which can directly cause an increase in coolant inventory by increasing the total feedwater flow. The most severe applicable event is a feedwater controller failure during maximum flow demand. The feedwater controller is assumed to fail to its upper limit at the beg:nning of the event.
A feedwater controller failure during maximum flow demand at 75% power and 60% flow during single recirculation loop operation produces the sequence of events listed in Table 15.C.3-2. The computer model described in Reference 15.C.8-2 was used to simulate this event.
The analysis has been performed with the plant conditions tabulated in Table 15.C.3-1, except with the initial vessel water level at Level 4 (instead of normal water level) for conservatism. By lowering the initial
{
water level, more cold feedwater will be injected before Level 8 is reached resulting in higher heat fluxes.
1 15.C.3-2
EF,-2 FSAR The safety analysis condition is at 75% rated thermal power and 60%
rated core flow, which represents single recirculation loop operation at 1005 pump speed on the 1055 rod line. End of cycle (all rod out) scram characteristics are assumed. The safety-relief valve action is co'nserva-tively assumed to occur with higher than nominal setpoints. The transient is simulated by programing an upper limit failure in the feedwater system such that 130% of rated feedwater flow occurs at the reactor dome pressure of 972 psig.
Feedwater Controller Failure - Maximum Demand with Bypass in Service
. The simulated feedwater controller failure transient is shown in Figure 15.C.3-1. The high water level turbine trip and feedwater pump trip are initiated at approximately 8.1 seconds. Scram occurs simultaneously from stop valve closure, and limits the neutron flux peak and fuel thermal transient. The turbine bypass system opens to limit peak pressures in the steam supply system. In addition to turbine bypasr Flow after the turbine trip, steam flow through the reheater lines is also available. The reheater flow is assumed to be 10% NBR steam flow for two seconds after bypass valves open, followed by ramping to zero flow for three more seconds. Events caused by low water level trips, including initiation of HPCI and RCIC core cooling system functions are not included in the simulation. Should these events occur, they will follow sometime after the primary effects have occurred, and are expected to be less severe than those already experienced by the system.
Table 15.C.3-4 gives a summary of the transient an'alysis results. The calculated MCPR is 1.16, which is well above the safety limit MCPR of 1.07 so no fuel failure due to boiling transition is predicted. The peak vessel pressure predicted is 1094 psig and is well below the ASME limit of 1375 psig.
l l 15.C.3-3
. EF-2 FSAR Feedwater Controller Failure - Maximum Demand, with Bypass Out-of-Service An evaluation has been performed with the main steam bypass function assumed to be unavailable (i.e., out-of-service). The steam flow'to the reheater line is assumed for the evaluation.
Table 15.C.3-5 sumarizes the CPR results. The calculated MCPR is 1.15 whi:h '7: considerably above the safety limit MCPR of 1.07.
15.C,J.1.2 Turbine / Generator Trip A Turbine / Generator trip, with closure of the turbine stop valves and turbine control valves, can be initiated by a variety of turbine or nuclear system malfunctions. Some exar.ples are moisture separator and heater drain tank high levels, large or excessive vibrations, operator lock out, low condenser vacuum and reactor high water level. The fast closure of the turbine stop valves and turbine control valves causes e sudden reduction of steam flow which results in a nuclear pressure increase.
The turbine stop valves and turbine control valves initiate a scram trip signal for power levels greater than 30% of rated. The closure
, characteristics of the turbine control valves are assumed such that the valves uperate in the full arc (FA) mode and have a full stroke closure time of 0.20 seconds.
All plant control systems maintain normal operation unless specifically designated to the contrary. The pressure relief system which operates the relief valves independently when system pressure exceeds relief valve instrumentation setpoints is assumed to function normally during the time period analyzed. Auxiliary power would normally be independent of any turbine-generator overspeed effects and continuously be supplied at rated frequency as automatic fast transfer to auxiliary power supplies occurs.
l c
15.C 3-4 L __ _ . _ _ _ - _ _ - - _ _ - _
EF-2 FSAR The computer model described in Reference 15.C.8-2 was used to simulate this event. The safety analysis initial condition is at 75% rated thermal !
power and 605 rated core flow.
1 Events caused by low water level trips, including init.iation of HPCI and RCIC core cooling system functions are not included in this simulation.
If these events occur, they will follow sometime after the primary concerns of fuel margin and overpressure effects have passed, and will result in effects less severe than those already experienced by the reactor system, and will provide long-term reactor inventory control.
Turbine / Generator Trip with One-Half Bypass The analysis has been performed with the plant conditions tabulated in Table 15.C.3-1, except that the turbine bypass function is assumed to be one-half capacity which is approximately 13% NBR. The reheater line flow is also available as in the case of the FWCF transient.
Table 15.C.3-4 summarizes the transient analysis results. The peak neutron flux reaches about 111% of rated and average surface heat flux peaks at about 102% of its initial value. The peak vessel pressure predicted is 1146 psig and is well below the ASME limit of 1375 psig. The calculated MCPR is 1.19 which is considerably above the safety limit MCPR of 1.07.
Turbine / Generator Trip with Full Bypass Failure An evaluation has been performed with the full turbine bypass function assumed to fail. The reheater line flow is also available as in the case of the FWCF transient.
Table 15.C.3-5 sumarizes the CPR results. The calculated MCPR is 1.18 l which is considerably above the safety limit MCPR of 1.07.
1 l
l l
l 15.C.3-5
EF-2 FSAR O
15.C.3.1.3 Sumary and conclusions The transient peak value results and the Critical Power Ratio (CPR) results are sumarized in Table 15.C.3-4 and Table 15.C.3-5. They indicate that the MCPRs for all transients are above the single-loop operation safety limit value of 1.07. It is concluded that the operating limit MCPRs estab-11shes for two-puy operation are also applicable to single-loop operation conditions.
For pressurization, the transient peak pressures are well below the ASME code value of 1375 psig. Hence, it is concluded that the pressure barrier integrity is maintained under single-loop operation.
From the above discussions, it is concluded that the transient conse-quences for one-loop operation are bounded by previously submitted full power analyses.
15.C.3.2 Rod Withdrawal Error The rod withdrawal error at rated power is given in the FSAR. .These j analyses are performed to de:nonstrate, even if the operator ignores all instrument indications and the alarm which could occur during the course of the transient, the rod block system will stop rod withdrawal at a minimum critical power ratio (MCPR) which is higher than the fuel cladding integrity safety limit. Modification of the rod block equation (below) and lower power assures the MCPR safety limit is not violated.
One-pump operation results in backflow through 10 of the 20 jet pumps while the flow is being supplied into the lower plenum from the 10 active jet pumps. Because ot' the backflow through the inactive jet pumps, the present rod block equation was conservatively modified for use during l .one-pump operation because the direct active-loop flow measurement may not l indicate actual flow above about 40% core flow without correction.
t l
15.C.3-6
EF-2 FSAR A procedure has been estabitshed for correcting tha rod block equation to account for the discrepancy between actual flow and indicated flow in the active loop. This preserves the erfginal relationship between rod block and actual effective drive flow when operating with a single-1 cop.
The two-pump rod block equation is:
RB = sk + RB100 - a(100)
The one-pump equation becomes:
AB = sW + RB100 - m(100) - W where AW =
difference between two-loop and single-loop effective drive flow at the same core flow. This value is expect-ed to be 8% of rated.
RB = power at rod block in %;
m u flow reference slope for the rod block monitor (RBM)
W = drive flow in % of rated.
RB 100
= top level rod block at 100% flow.
If the rod block setpoint (RB100) is changed, the equation mu:;t be recalcu-lated using the new value.
The APRM trip settings are flow biased in the same manner as the rod block monitor trip setting. Therefore, the APRM rod block and scram trip settings are subject to the same procedutal changes as the rod block monitor trip settings discussed above.
15.C.3-7
EF-2 FSAR
- 15.C.3.3 Doerating MCPR Limit For single-loop operation, the' operating MCPR limit remains u,nchanged from the normal two-loop operation limit. Although the increased uncertain-ties in core flow and TIP readings resulted in a 0.01 incremental increase in MCPR fuel cladding integrity safety limit during single-loop operation (Section 15.C.2), the limitir.g transients have been analyzed to indicate that there is more than enough MCPR margin during single-loop operation to i compensate for this increase in safety limit. For single-loop operation at 1 lower flows, the steady-state operating MCPR limit is established by the X l f
curve. This ensures the 99.9% statistical limit requirement is always l satisfied for any postulated abnormal operational occurrence. Since the f
mar,1 mum core flow runout during single loop operation is only about 60% of l rated, the current flow dependent K fcurve which fs generated based ca the flow runout up to rated core' flow are aise adequate to protect the flow l .
runout events during single-loop operation. l l
m mm e
h e
1 15.C.3-8
EF-2 FSAR l
TABLE 15.C.3-1 IMPl1T PARAMETERS ^AND INITIAL CONDITIONS
- 1. Thermal Power Level, fWt 2470
- 2. Steam Flow, Ibs per hr 10.17 x 10 0 l
- 3. Core Flow, 1bs per hr 60 x 10 6
- 4. Feedwater Flow Rate, Ib per s e e.. 2824
- 5. ,Feedwater Temperature. *F 390
- 6. Vessel Dome Pressure, psia 987
- 7. Vessel Core Pressure, psis 993 ;
- 8. Turbine Bypass capacity 5 NBR 26.0
- 9. Core Coolant Inlet Enthalpy, 8tu per 1b 512.1
- 10. Turbine Inlet Pressure, psia 957
- 11. Fuel Lattice C (P8x8R)
- 12. CoreAverageGapconductance, 8tu/sec-ft *F 0.1744
- 13. Core Leakage Flow, 1 11.99
- 14. Required MCPR Operating Limit Kf* rated CLMCPR
- 15. MCPR Safety Limit 1.07
- 16. Doppler Coefficient, c/'F **
- 17. Void Coefficient, c/5 Rated Voids *
- 18. Core Average Void Fraction, 5 **
- 19. Scram Reactivity, SoK e
- The values are calculated within the code (Reference 15.C.8 2) for end of Cycle 1 conditions.
15.C.3-9
. EF-2 FSAR
. TABLE 15.C.3-1 (Continued)
I
- 20. Control Rod Drive Speed, Position versus Time Figure 15.0-1 l
- 21. Jet Pump Ratio M 3.57 l
- 22. Nuclear Characteristics used in ODYN analysis E001 l
- 23. Safety / Relief Valve Capacity, 5 NBR e 1121 psig 90.85 Manufacturer Target Rock Quantity Installed 15 j
- 24. Relief Function Delay, second 0.4 l l
- 25. Relief Function Stroke Time, seconds 0.15 l
- 26. Setpoints for Safety / Relief Valves Safety / Relief Function, psig 1121, 1131, 1141 1
- 27. Number of Valve Groupings Simulated Safety / Relief Function, No. 3
- 28. High Flux Trip, 5 NBR Analysis set point (121 x 1.02), 5 NBR 123.4
- 29. High Pressure Scram Setpoint, psig 1101.0
- 30. Yessel Level Trips, Feet Above Separator Skirt Bottom Level 8 - (LB), feet 5.917 Level 3 - (L3), feet 1.75 Level 2 - (L2), feet (-)4.71
- 31. Simulated Thermal Power (STP) Scram Trip Maximum safety set point, 5 NBR (117 x 1.02) 119.3 Time constant, seconds 6.0 Flow reference set point slope Figure 15.0 5 32 High Pressure Recirculation Pump Trip Pressure set point, psig '
1135.0 Delay time, seconds 0.3 l
l 15.C.3-10
(
4 i
, EF-2 FSAR TABLE 15.C.3-1 (Continued)
- 33. Pump + Motor Inertia Time Constant Seconds *** 2.0 .
- 34. Total Steam 11ne Volume, ft* 4658.0
- 35. Reheater Bypass Flow, 5 NBR 10.0 (Figures 15.0-3 and 15.0-4) m The inertia time constant is defined by the expression:
2nJno t= gT o , where t = inertia time constant (sec)
J, e pump motor inertia (1b-ft')
n = rated pump speed (rps) g = gravitational constant (ft/sec')
T, = pump shaft torque (1b-ft) 15.C.3-11
EF-2 FSAR TABLE 15.C.3-2 SEQUENCE OF EVENTS FOR FIGURE 15.C.3-1 FT.EDWATER CONTROLLER FAILURE, RAXIMUM DEMAND WITH FULL BTPA55 AND REMEATER FLOW Time-See Event 0 Initiate simulated failure of the 130 percent upper limit on feedwater flow.
8.06 L8 vessel level setpoint trips main turbine and feedwater pumps.
8.06 Turbine trip initiates closure of turbine stop valves and fast closure of turbine control valves.
8.06 Turbine trip initiates bypass operation.
8.08 Main turbine stop valves reach 90 percent open position and initiate a reactor scram trip.
8.09 Fast closure of turbine control valves initiates a reactor scram trip.
8.16 Main turbine bypass valves opened.
8.21 Turbine control valves closed.
8.26 Turbine stop valves closed.
10.06 Isolation valves on reheater lines start to close.
13.06 (Est) Reheater flow decays to zero.
15.C.3-12
EF-2 FSAR .
TABLE 15.C.3-3
$EQUENCC 0F EVENTS FOR FIGURE 15.C.3-2 TURBINE / GENERATOR TRIP ,
WITH ONE-HALF BYPASS AND REHEATER FLOW Time-see Event 0 Turbine / Generator trip initiates closure of main turbine stop valves and control valves.
O Turbine / Generator initiates bypass operation (one bypa.ss valve only).
0.02 Main turbine stop valves reach 90 percent open position and initiate a reactor scram trip.
0.1 Turbine bypass valves start to open to regulate pressure.
0.15 Turbine control valves closed.
0.2 Turbine stop valves closed.
2.0 Isolation valves on reheater lines start to close.
4.34 Group 1 relief valves actuated.
5.0 Reheater flow decays to zero.
7.5 Group 1 relief valves closed.
8 P
s 15.C.3-13
EF-2 FSAR TABLE 15.C.3-4 SUPNARY OF TRANSIENT PEAK VALUE AND CPR RESULTS a
FWCF T/GT1/2BN Initial Power / Flow (5 Rated) 75/60 75/60 Peak Neutron Flux (5 Rated) 102.0 111.0 Peak Heat Flux (5 Initial) 108.0 102.0 Peak Dome Pressure (psig) 1075 1128 Peak Vessel Bottom Pressure (psig) 1094 1146 Initial MCPR 1.30 1.30 c
Transient MCPR 1.16 1.19 Safety Limit MCPR (for SLO) 1.07 1.07 Margin to Safety Limit 0.09 0.12 a
FWCF = Feedwater Controller Failure to Maximum Demand with Reheater Flow and Bypass in Service b
T/GT 1/2 BP = Turbine / Generator Trip with Reheater Flow and Half-Bypass
- Includes Option A adder.
15.C.3-14 l
- EF-2 FSAR TABLE 15.C 3-5
$UMMARY OF CPR RESULTS a b FWCFNBP T/GTN8P_
Initial Power / Flow (5 NBR) 75/60 75/60 C d Rated OLMCPR 1.31 1.25' I
SLO,OLMCPR 1.40 1.34 Transient MCPR9 1.15 1.18 Safety Limit SLCPR 1.07 1.07 Margin to Safe'ty Limit 0.08 0.11
- a. FWCFNBP = Feedwater controller failure to maximum demand with Reheater Flow and with bypass out-cf-service,
- b. .'/GTNBP = Turbine / Generator Trip with Reheater Flow and with full bypass failure.
- c. Rated OLMCPR = operating limit minimum critical pow'er ratio for rated power flow condition.
- d. For Bypass out-of-service,
- e. For Bypass in-service. '
l g. Includes option A adder.
15 C.3-15 1
1 NEUTRON F ' LUX 2 PERK FUEL CENTER TEMP 150* -- = 3 RVE SURFI CE HEAT FLUX 11 FLEDH01EF FLOH
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, _ , EF-2 FSAR 15.C.4 STABILITY ANALYSIS 15.C.4.1 Phenomena The primary contributing factors to the stability performance with one recirculation loop.not in service are the power / flow ratio and the recircu-lation loop characteristics. As forced circulation with only one recircu-
'lation loop in operation, the reactor core stability is influenced by the inactive recirculation loop. As core flow increases in SLO the inactive jet pump forward flow decreases because the driving head'across the inactive jet pumps decreases with increasing core flow. The reduced flow in the inactive loop reduces the resistance that the recirculation loops impose on reactor core flow perturbations thereby adding a destabilizing effect. At ine same time the incres mi core flow results in a lower power / flow ratio which is a stabilizing effect. These two countering effects result in slightly decreased stability margin (higher decay ratio) initially as core flow is increased (from minimum) in SLO and then an increase in stability margin (lower decay ratio) as core flow is increased further and reverse flow in the inactive loop is established. I As core flow is increased further during SLO and substantial reverse flow is established in the inactive loop an increase in jet pump flow, core flow and neutron noise is observed. A cross flow is established in the :
annular downcomer region near the jet pump suction entrance caused by the reverse flow of the inactive recirculation loop. This cross flow 11teracts with the jet pump suction flow of the active recirculation loop and in-creases the jet pump flow noise. This effect increases the total core flow noise which tends to drive the neutron flux noise.
To determine if the increased noise is being caused by reduced stabil-ity margin as SLO core flow was increased, an evaluation was performed which phenomenologically accounts for single-loop operation effects on stability, as sumarized in Reference 15.C.8-3. The model predictions were initially i
i 15.C.4-1
_ _ . _ _ . _ _ _ _ _ _ _ _ _ _ _ . _ - - _ - - _ _ - - - - ~ - - - - - - - - - - - ~ - - - - - - - - - - - - ' '
EF-2 FSAR compared with test data and showed very good agreement for both two-loop and '
single-loop test conditions. An evaluation was performed to determine the effect of reverse flow on stability during $LO. With increasing reverse flow, SLO exhibited slightly lower decay ratios than two-loop operation.
However, at core flow conditions with no reverse flow, SLO was slightly less stable. This is consistent with observed behavior in stability tests at operating BWRs (Reference 15.C.8-4).
In addition to the above analyses, the cross flow established during reverse flow conditions was simulated analytically and shown to cause an increase in the individual and total jet pump flow noise, which is consis-tent with test data (Reference 15.C.8-3). The results of these analyses and tests indicate that the stability characteristics are not significantly different from two-loop operation. At low core flow, SLO may be slightly less stable than two-loop operation but as core flow is increased and reverse flow is established the stability performance is similar. At higher '
core flow with substantial reverse flow in the inactive recirculation loop, the effect of cross flow on the flow noise results in an increase in system noise (jet pump, core flow and neutron flux noise).
15.C.4.2 Compliance to Stability Criteria Consistent with the philosophy applied to two-loop operation, the stability compliance during sin;1e-loop operation is demonstrated on a generic basis. Stability acceptance criteria have been established to demonstrate conpliance with the requirements set forth in 10CFR50, Appen-dix A General Design Criterion (GDC) 12 (Reference 15.C.8-5). The generic stability analysis has been performed covering all licensed GE BWR initial core and reload core fuel designs including those fuels contained in the Generel Electric Standard Applicatior, fcr Reload Fuel (GESTAR Reference 15.C.8-6 through Amendment 10).* The analysis demonstrates that in the event limit cycle neutron flux oscillations occur within the bounds of safety system intervention, specified acceptable fuel design limits are
=The reload fueT"designs contained in GESTAR include fuel designs through the GE8x8E design (including barrier fuel).
15.C.4-2
EF-2 FSAR not exceeded. Since the reactor core is assumed to be in an oscillatory mode, the question of stability urgin during SLO is not relevant frori; a safety standpoint (i.e., the analysis already assumes no stability margin).
The fuel performance during limit cycle oscillations is characteris-tically dependent on fuel design and certain fixed system features (high
- neutron flux scram setpoint, channel inlet orifice diameter, etc.). There- 1 fore the acceptability of GE fuel designs independent of plant and cycle parameters has been establishcc. W y th m parameters unique to SLO which '
affect fuel performance need to be evaluated. The major consideration of SLO is the increased Minimum Critical Power Ratio (MCPR) safety limit caused by increased uncertainties in system parameters during SLO. However, the increase in MCPR safety limit (0.01) is well within the margin of the limit cycle analyses (Reference 15.C.8-5) and therefore it is demonstrated that stability conpliance criteria are satisfied during single-loop operation.
Operationally, the effects of higher flow noise and neutron flux noise observed at high SLO core flow are evaluated to determine if acceptable vessel internal vibration levels are met and to determine the effects on fuel and channel fatigue, and are not considered in the compliance to stability criteria.
Service Information Letter-380, Revision 1 (Reference 15.C.8-7) has been developed to inform plant operators how to recognize and suppress unanticipated oscillations when encountered during plant operation.
As a result of the above analysis and operator rec'ormendations, the NRC staff has approved the generic stability analysis for application to single-loop operation (Reference 15.C.8-8) provided that the recomendations of l
$1L-380 have been incorporated into the Plant Technical Specifications.
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EF-2 FSAR 15.C.5 \
LOSS-DF-COOLANT ACCIDENT ANALYSIS
, If two recirculation loops are operating and a pipe break occurs in one of the two recirculation loops, the pump in the unbroken loop is assumed to immediately trip and begin to coast down. The decaying core flow due to the pump coastdoun results in very effective heat transfer (nucleate boiling) dur.ing the initial phase of the blowdown. Typically, nucleate boiling will be sustained during the first five to nine seconds after the accident, for the design basis accident (DBA).
If only one recirculation loop is operating, and the break occurs in the operating loop, continued c:re flow is provided only by natural circu-lation because the vessel is blowing down to the reactor containment through both sections of the broken loop. The core flow decreases more rapidly than in the two-loop operating case, and the departure from nucleate boiling for the high power node might occur one o* two seconds after the postulated accident, resulting in more severe cladding heatup for the one-1 cup operat-ing case.
In addition to changing the blowdown heat transfer characteristics, 1esing recirculation pump coastdown flow can also affect the system invento-ry and reflooding phenomena. Of particular interest are the changes in the high-power node uncovery and reflooding times, the system pressure and the time of rated core spray for different break sizes. One-loop operation results in small changes in the high-power node uncovery times and times of rated spray. The effect of the reflooding times for various break sizes is also generally semil.
An analysis of single recirculation loop operation using the models and assumptions documented in Reference 15.C.8-9 was performed for Fermi-2.
Using this method, SAFE /REFLOOD computer code runs were made for a full spectrum of large break sizes for only the recirculation suction line breaks (most limiting for Fermi-2). Because the reflood minus uncovery time for 15.C 5-1
EF-2 FSAR the single-loop analysis is similar to the two-loop analyt,is, the maximum planar linear heat generation rate (MAPLMGR) curvas were modified by derived reduction factors which account for'the degraded blowdown heat transfer during one recirculation pump operation.
i The LPC1 loop selection logic does not specifically affect the analysis. This logic trips the recirculation pumps and then waits until after pump coastdown to determine which recirculation loop to inject into. 1 The ability of the LPCI loop selection logic to adequately perferm ti:is ,
function is unaffected by operation with a single recirculation loop.
15.C.5.1 Break Spectrum Analysis SAFE /REFLOOD calculations were performed using assumptions given in Section II.A.7.3.1 of Reference 15.C.8-9. Not node uncovered time (time between uncovery and reflood) for single-loop opeastion is compared to that ,
for two-loop operation in Figure 15.C.5-1. '
The total uncovered time for two-loop operation is 148 seconds for the 100% DBA suction break. This is the most limiting break for two loop operation. For single-loop operation. the total uncovered time is 149 seconds for the 1005 DBA suction break. This is the most limiting break for single-loop operation. Figure 15.C.5-2 shows the water level response inside the shroud under single loop operation for the 100% DBA suction break.
15.C.S.2 Single-loop MAPLHGR Determination 4
The small differences in uncovered time and reflood time for the limiting break size would result in a small change in the calculated peak cladding temperaturer. Therefore, as noted in Reference 15.C.8-9, the single and two-loop SAFE /REFLOOD results can be considered similar and the generic alternate procedure described in Section II.A.7.4 of this reference was used to calculate the MAPLHGR reduction f actors for single-loop operation. The .
most limiting single-loop operation MAPLHGR reducticn factor (i.e., yielding the lowest MAPLHGR) is 0.90. Single-loop operation MAPLHGR values are 15.C.5-2
. EF-2 FSAR i
derived by multiplying the current two-loop MAPLHGR values by the reduction factor 0.90. As discussed in Referehce 15.C,d-0, single recirculation loos MAPLHGR values are conservative when calculated fri this manner. This ,
! MAPLHGR multiplier is applicable to P8x8R fuel in the initial core. For '
reload situations, the M4PLHGR must be assessed for each cycle to determine if it is still applicable because the single-loop MAPLHGR multiplier was based on the calculated peak cladding temperature from the two-loop analysis for the initial core fuel.
15.C.5.3 Small Break Peak Cladding Temperature Section II.A.7.4.4.2 of Reference 15.C.8-9 discussas the low sensi-tivity of the calculated peak cladding temperature (PCT) to the assumptions used in the one-pump operation analysis and the duration of nucleate boil-ing. As this slight increase ($ 50'F) in PCT is overwhelmingly offset by .
the decreased MAPLHGR (equivalent to 300'F to 500'F PCT) for one-pump operation, the calculated PCT values for small breaks will be well below the 1847'F small break PCT value previously reported for Fermi-2, and signifi-cantly below the 2200*F 10CFR50.46 cladding teg erature limit.
15.C.5-3
EF-2 FSAR e
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[F-2 FSAR 15.C.6 CONTAINMENT Analysis .
The range of the power / flow conditions which are included in the SLO operating domain for Fermi-2 was investigated to determine if there would be cny impact on the FSAR specifications for containment response, including the containment dynamic loads. The SLO operating conditions were confirmed to be within the range of operating conditions which have previously been considered in defining t:1e containment pressure and temperature response and containment dynamic loads for two-loop operation. Therefore, the contain-sent response for Fermi-2 with single-loop operation has been confirmed to be within the present design values.
15.C.6-1 R_m
EF-2 FSAR b
15.C.7 MISCELL ANEDUS IMPACT EVALUATION 15.C.7.1 Anticipated Transient Without Scram (ATWS) Impact Evaluation The principal difference between single-loop operation (SLO) and normal two-loop operation (TLD) affecting Anticipated Transient Without Scram (ATW5) performance is that of initial reactor conditions. Since the SLO initial power flow condition is less than the rated condition used for.TLO ATWS analysis, the transient response is less severe and therefore bounded I
by the TLD analyses.
It is concluded that if an ATWS event were initiated at Fermi-2 from the SLO conditions, the results would be less severe than if it were initi-ated from rated conditions.
15.C.7.2 Fuel Mechanical Performance Evaluations were performed to determine the acceptability of Fermi-2 single-loop operation on P8x8R fuel rod and assembly thermal / mechanical performance. Component pressure differential and fuel rod overpower values were determined for anticipated operational occurrences initiated from SLO conditions. These values were found to be bounded by those applied in the fuel rod and sciembly design bases, It is observed th:1 due to the substantial reverse flow established during SLO both th". Average Power Range Monitor (APRM) noise and core plate differential pressure noise are slightly increased. An analysis has been ,
carried out to determine that the APRM fluctuation should not exceed a flux amplitude of 215% of rated and the core plate differential pressure fluc-tuation should not exceed 3.2 psi peak to peak to be consistent with the fuel rod and assembly design bases.
15.C.7-1 L
c d-2 FSAR i O
15.C.7.3 Vessel Internal Vibration Feral-2 is a BWR/4-251" plant which is categorized as Non-Prototype Category 1 according to the NRC Reg. Guide 1.20. The prototype plant for Fermi-2 is Browns Ferry-1. The prototype plant is completely instrumented and extensively tested during preoperational and start-up testing at various power levels and recirculation flow rates, including single-loop operation.
In addition, test results from Browns Ferry-2 and 3 and Peach Bottom-2 and 3 are available which are of the same design as Browns Ferry-1. j i
In Fermi-2 jet pumps there is a diffuser adapter which is not present l st the Browns Ferry-1 jet pumps. In the La Sa11e-1 (BWR/5-251) jet pumps i there is an identical diffuser adapter. La Sa11e-1 jet pumps were extensively tested during preoperational and startup testing at various power levels and recirculation flow rates, incluaing single-loop operation.
The single-loop operation test results of La Sa11e-1, Browns Ferry-1, 2 and 3, and Peach Bottom-2 and 3 were used to conservative:1y infer the allowable pump operating speed at Fermi 2. Dynamic analyses were performed on the jet pumps of each plant to obtain their natural frequencies, mode shapes and generalized forces. Based on these dynamic characteristics the effect on jet pump vibration due to the addition of the diffuser adapter was established. These effects were applied to the Browns Ferry test results under single-loop operation and the expected level of vibration at Fermi-2 was calculated. Based on this conservative calculation, the maximum allow-able pump speed during single-loop operation is 75% of' rated (Reference 15.C.B-10).
It is recomended that Fermi-2 not perform single-loop operation with pump speed exceeding 75% of rated pump speed.
l l
15.C.7-2 i
EF-2 FSAR 15.C.8 REFERENCES 15.C.8-1 "General Electric BWR Thersal Analysis Basis (GETAB); Data, Correlation, and Design Application", NED0-10968-A, January 1977, 15.C.8-2 "Qualification of th One-Dimensional Core Transient Model for Boiling Water Reactors" NEDO-24154, October 1978.
15.C.8-3 Letter, H.C. Pfefferlen (GE) to C.0. Thomas (NRC), "Submittal of Response to Stability Action Item from NRC Concerning Single-Loop Operation," September 1983.
15.C.8-4 5.F. Chen and R.0. Niemi, "Vermont Yankee Cycle 8 Stability and Recirculation Pump Trip Test Report". General Electric Company, August 1982 (NEDE-25445, Proprietary Information).
15.C.8-5 G.A. Watford, "Compliance of the General Electric Boiling Water Reactor Fuel Designs to Stability Licensing Criteria" General Electric Company, October 1984 (NEDE-22277-P-1, Proprietary Information).
15.C.8-6 "General Electric Standard Application for Reload Fuel", General Electric Company, April 1983 (NEDE-24011-P-A-6).
15.C.8-7 *BWR Core Thermal Hydraulic Stability", General Electric Company, February 10, 1984 (Service Information Letter-380, Revision 1).
15.C.8-8 Letter, C.0. Thomas (NRC) to H.-C. Pfefferlen-(GE), "Acceptance for ,
Referencing of Licensing Topical Report NEDE-24011, Rev. 6 Amendment 8. Thermal Hydraulic Stability Amendment to GESTAR II,"
April 24, 1985.
15.C.8-1
EF-2 FSAR 1
1 .
s 15.C.8 REFERENCES (Cont'd) 15.C.8-9 "General Electric Conpany' Analytical Model for Loss-of-Coolant Analysis in Accordance with 10CFR50 Appendix K Amendment No. 2 -
One Recirculation Loop Out-of-Service" MED0-20566-2 Revision 1, July 1978.
l 15.C.8-10 "Fermi-2 Single Loop Operation Vibration Analysis". EQDE-32-0486, April 1986 DRF A12-00033.
o 6
15.C.8-2 k- ._ _ _ _ _