ML20155B091
| ML20155B091 | |
| Person / Time | |
|---|---|
| Site: | Grand Gulf  |
| Issue date: | 02/28/1986 |
| From: | GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML20155B084 | List: |
| References | |
| TAC-61083, NUDOCS 8604100202 | |
| Download: ML20155B091 (50) | |
Text
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GGNS February 1986 GGNS SINGLE LOOP OPERATION ANALYSIS FEBRUARY 1986 Prepared for j
MISSISSIPPI POWER AND LIGHT COMPANY GRAND GULF 1 & 2 NUCLEAR STATIONS i
Prepared by i
GENERAL ELECTRIC COMPANY ;
NUCLEAR ENERGY BUSINESS OPERATIONS SAN JOSE, CALIFORNIA 95125 8604100202 860331 PDR ADOCK 05000416 p PDR 7 .- - , , gr vw-- - ,-- -- --
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&M- y, ,
GGNS February 1986 APPENDIX 15.C TABLE OF CONTENTS ,
Page 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 15.C.2-1 Operation 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.3.1.1.1 Core and System Performance 15.C.3-2
- 15. C . 3 '.1.1. 2 Barrier Performance 15.C.3-4 15.C.3.1.1.3 Radiological Consequences 15.C.3-4 15.C.3.1.2 Generator Load Rejection With Bypass Failure 15.C.3-4 15.C.3.1.2.1 Core and System Performance 15.C.3.4 15.C.3.1.2.2 Barrier Performance 15.C.3-6 15.C.3.1.2.3 Radiological Consequences 15.C.3-6 15 C.3.1.3 Recirculation Pump Seizure Accident 15.C.3-6 15.C.3.1.3.1 Core and System Performance 15.C.3-6 15.C.3.1.3.2 Barrier Performance 15.C.3-7 15.C.3.1.3.3 Radiological Consequences 15.C.3-7 15.C.3.1.4 Summary and Conclusions 15.C.3-7 15.C.3.2 Rod Withdrawal Error ,
15.C.3-7 15.C.3.3 Operating MCPR Limit 15.C.3-9 15.C-1 HLV:rf:rm/F07183*-2 66** 4 m-* -L.
l GGNS February 1986 TABLE OF CONTENTS (Continued)
Page 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.5.1 Break Spectrum Analysis 15.C.5-1 15.C.5.2 Single-Loop MAPLHGR Determination 15.C.5-1 15.C.S.3 Small Break Peak Cladding Temperature 15.C.5-2 15.C.6 CONTAINMENT ANALYSIS 15.C.6-1 15.C.7 MISCELLANEOUS IMPACT EVALUATION 15.C.7-1 15.C.7.1 Anticipated 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 Vibration 15.C.7-1 15.C.8 REFERENCES 15.C.8-1 1
15.C-ii HLV:rf:ra/F07183*-3
I GGNS February 1986 LIST OF TABLES NUMBER TITLE PAGE 15.C.3-1 Input Parameters and Initial Conditions 15.C.3-10, 11 for Transients and Accidents for Single-Loop Operationg 15.C.3-2 Summary of Transient Peak Value Results 15.C.3-12 Single-Loop Operation 15.C.3-3 Summary of Critical Power Ratio Results - 15.C.3-13 j . Single-Loup Operation 1
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! 15.C-iii HLV:rf:ra/F07183*-4 )
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CGNS . February 1986 LIST OF FIGURES NUMBER TITLE 15.C.2-1 Illustration of Single Recirculation Loop Operation Flows 15.C.3-1 Peak Dome Pressure vs. Initial Power Level, Turbine Trip at E0EC 15.C.3-2 Feedvater Controller Failure - Maximum Demand, Single Loop Operation 15.C.3-3 Cenerator Load Rejection with Bypass Failure, Single-Loop Operation 15.C.3-4 Seizure of One Recirculation Pump, Single-Loop Operation 15.C.5-1 Uncovered Time vs. Break Area - Suction Break, LPCS Failure i
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15.C-iv HLVarftra/F07183*-5
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GGNS 15.C RECIRCULATION SYSTEMS SINGLE-LOOP OPERATION 15.C.1 INTRODUCTION AND SUPNARY 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 single-loop operation, accidents and abnormal operational transients associated with power operations, as presented in Sections 6.2 and 6.3 and the main text of Chapter 15.0, were reviewed for the single-loop case with only one pump,in operation. This appendix presents the results of this safety evaluation for the operation of the Grand Gulf Nuclear Stations (GGNS) with single recirculation loop inoperable.
l This evaluation is performed for GE-6 fueled GGNS on an initial cycle basis and is applicable to GE-6 fueled normal annual 12 month initial cycle operation. The conditions are those of continued operation in the
]
j- operating domain currently defined in Figure 4.4.5 of Chapter 4 up a maximum power of 70.6% of r'ated.
Increased uncertainties in the core total flow and Traversing In-Core Probe (TIP) readings resulted in a 0.01 incremental increase in the Minimum Critical Power Ratio (MCPR) fuel cladding integrity safety limit during single-loop operation. No increase in rated MCPR operating limit and no change in the power dependent and flow dependent MCPR limit (MCPRf and MCPR ) are required because all abnormal operational transients p
analyzed for single-loop operation indicated there is more than enough
[ HCPR 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 12 (10CFR50, Appendix A). It is shown that SLO satisfies this stability criterion. It is further shown that the increase in neutron noise observed during SLO is independent of system stability margin.
HLV:rf:rs/F07184* 15.C.1-1
GGNS l
To prevent potential control oscillations from occurring in the recircu-lation flow control system, the flow control should be in master manual for single-loop operation.
The limiting Maximum Average Planar Linear Heat Generation Rate (MAPLHGR) reduction factor for single-loop operation is calculated to be 0.86.
The containment response for a Design Basis Accident (DBA) recirculation line break with single-loop operation is bounded by the rated power two-loop operation analysis presented in Section 6.2. This conclusion covers all single-loop operation power / flow conditions.
The impact of single loop operation on the Anticipated Transient Without l
Scram (ATWS) analysis was evaluated. It is found that all ATWS acceptance criteria are met during SLO.
The fuel thermal and mechan'ical duty for transient events occurring during SLO is found to be bounded by the fuel design bases. The Average Power Range Honitor (APRM) fluctuation should not exceed a flux amplitude of 115% of rated and the core plate differential pressure fluctuation should not exceed 3.2 psi peak to peak to be consistent with the fuel rod and assembly design bases.
A recirculation pump drive flow limit will be imposed for SLO. The highest drive flow tested during the startup test program at GGNS that meets acceptable vessel internal vibration criteria will be the drive flow limit for SLO.
HLV: rf:gc/F07184* 15.C.1-2
GGNS 15.C.2 MCPR FUEL CLADDING INTEGRITY SAFETY LIMIT j .
Except for core total flow and TIP reading, the uncertainties used in the I statistical analysis to determine the MCPR fuel cladding integrity safety limit are not dependent on whether coolant flow is provided by one or two l recirculation pumps. Uncertainties used in the two-foop operation analysis are documented in the FSAR. A 6% core flow measurement uncer-tainty has been established for single-loop operation (compared to 2.5%
l for two-loop operation). As shown below, this value conservatively reflects the one standard deviation (one sigma) accuracy of the core flow measurement system documented in Reference 15.C.8-1. The rand:.m noise component of the TIP reading uncertainty was revised for single recircu-lation loop operation to reflect the operating plant test results given in Subsection 15.C.2.2. This revision resulted in a single-loop operation l
process computer effective TIP uncertainty of 6.8% of initial cores and
'- 9.1% for reload cores. Comparable two-loop process computer uncertainty j values are 6.3% for initial cores and 8.7% for reload cores. The net l
)
effect of these two revised uncertainties is a 0.01 incremental increase in the required MCPR fuel cladding integrity safet) limit. l 15.C.2.1 Core Flow Uncertainty 15.C.2.1.1 Core Flow Measurement During Sinate-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 backflow'ng (at active pump flow above approximately 36%). 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. Inaddition,thejetpumpcoefficientis different for reverse flow than for forward flow, and the measurement of
' . reverse flow must be modified to account for this difference.
In single-loop operation, the total core flow is derived'by the following formula:
HLV: rf:ge/F07184* 15.C.2-1
GGNS ..
Total Core , f Active Loop C
/InactiveLoop k
Flow (IndicatedFlows Flow /
Where C (= 0.95) is defined as the ratio of " Inactive Loop True Flow" to
" Inactive Loop Indicated Flow". "Leop Indicated Flow" is the flow measured by the jet pump " single-tap" loop flow summers and indicators, which are set to read forward flow correctly.
- The 0.95 factor was the result of a conservative analysis to appropriately modify the single-tap flow coefficient for reverse flow." If a more exact, less conservative core flow is required, special in-reactor calibration tests would have to be made. Such calibration tests would involve: calibrating core support plate AP versus core flow during j one pump and two pump operation 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 pump operation is essentially the same as for two pump )
operation, with some exceptions. The core flow uncertainty analysis is l described in Reference 15.C.8-1. The analysis of one pump core flow uncertainty is summarized below.
For single-loop operation, the total core flow can be expressec as follows (refer to Figure 15.C.2-1):
- The analytical expected value of the "C" coefficient for GGNS is so.82.
l l
HLV: rf:ge/F07184* 15.C.2-2
_ 1 _ _ _ _ _ _ _ _ _ _ _ . _ . . _ . _ _ _ _ _ _._.._____ __.__. ..._.
6 5 -
GGNS W
- C "A ~ "I where:
W = t tal core flow, C
1 Wg = active loop flow, and W = inactive loop (true) flow.
3 By applying the " propagation of errors" method to the above equation, the variance of the total flow uncertainty can be approximated by:
2 2 ,, ,,
or a2 1 a2 a W D' C C "sys "A 1 rand rand where:
og = uncertainty of total core flow; C
og = uncertainty systematic to both loops; t
= random uncertainty of active loop only;
- c"Isrand og = random uncertainty of inactive loop only; o = uncertainty of "C" coefficient; and c
a = ratio of inactive loop' flow (W3 ) to active loop
! flow (Wg ).
HLV:rf:ge/F07184" 15.C.2-3 l
.-. , , , _ _ - - - - _ _ . _ . - - , _ _ _ _ _ _ _ _ . . . - . - - ,- -- - . - - _ . . --. __- ,. _ - t
GGNS From an uncertainty analysis, the conservative, bounding values of W sys' W A rand '
rand respectively. Based on the above uncertainties and a bounding value of 0.36* for "a", the variance of the total flow uncertainty is approximately:
2 0.36 dic = (1.6): +(1-0 36 ( .6): gl-9,3 ) ((3.5)2+(2.8)2)
= (5.0)2 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 uncertainty, the active coolant flow uncertainty ist (5.0)2 0.12 cractive = ,
- ( 1-0.12 )2 (4.1)8 =
(5.1):
coolant which is less than the 6% flow uncertainty assumed in the statistical analysis.
In summary, core flow during one-pump operation is measured in a conser-vative 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 t ecirculation pump in operation (core flow 46.3% of rated). A rotationally symmetric control rod pattern existed during the test.
l l
- 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 GGNS is ~ 0.28.
HLVirf:gc/F07184* 15.C.2-4
GGNS Five consecutive traverses were made with each of five TIP machines, i
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-signa process l
computer total effective TIP uncertainty value for single-loop operation of 6.8% for initial cores and 9.1% for reload cores. The results of the
) analysis is directly applicable to GGNS because the data collected are l typical random neutron, electronic and boiling noise during SLO for a l
l BWR.
i i
t l
l i
l l
l l
)
i
)
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l HLVirfisc/F07184* 15.C.2-5 I
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cens A J l
=8 m, "A
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! W = Total Core Flow
! Wg = Active Loop Flow !
j Wi = Inactive Loop Flow l f
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'I O M1551551PPI POWER & LIGHT ILLUSTRATION 0F $!NGLE RECIRCULATION LOOP '5.C.2-1 l
CPERATION FLOWS l
l t
GGNS 15.C.3 MCPR OPERATING LIMIT 15.C.3.1 ABNORMAL OPERATING TRANSIENTS Operating with one recirculation loop results in a maximum power output which is about 30% below that which is attainable for two-pump operation.
Therefore, the consequences of abnormal operation transients from one-loop operation will be considerably less severe than those analyzed from a 7
two-loop operational mode. For pressurization, flow increase, flow decrease, and cold water injection transients, results presented in the FSAR bound both the thermal and overpressure consequences of one-loop l operation.
i j
Figure 15.C.3-1 shows the consequences of a typical pressurization i
transient (turbine trip) as a function of power level. As can be seen, f the consequences of one-loop operation are considerably less because of the associated reduction in operating power level.
l j The consequences of flow decrease transients are also bounded by the full power aralysis. A single pump trip from one-loop operation is less severe than a two-pump trip from full power because of the reduced initial power level.
The worst flow increase transient results from recirculation flow con-troller failure, and the worst cold water injection transient results from the loss of feedwater heater. For the former, the MCPR f curve is derived from a postulated event involving runout of both recirculation loops. This condition produces the maximum pcssible power increase and hence maximum 4MCPR for transients initiated from less than rated power and flow. When operating with only one recirculation loop, the flow and power increase associated with this failure with only one loop will be less than that associated with both loops; therefore, the MCPRf curve derived with the two-pump assumption is conservative for single-loop operation. The latter event, loss of feedwater heating, is generally the most severe cold water increase event with respect to increase'in core power. This event is caused by positive reactivity insertion from core inlet subcooling HLV:rf:ge/F07184* 15.C.3-1 ;
l 1
m . -- ,- .____ _ _ . . . _ _ _ . _ _ _ . . _ - - _ . . _ . . __ _ . . _ _ . _ _ __ _ ~ . - _ , _
GGNS and it is relatively insensitive to initial power level. A generic statistical loss of feedwater heater analysis using different initial power levels and other core design parameters concluded one pump opera- l tion with lower initial power level is conservatively bounded by the full l power two pump analysis. Inadvertent restart of the idle recirculation !
pump has been analyzed in the FSAR and is still applicable for single-loop )
operation.
From the above discussions, it is concluded that the transient consequence
)
from one-loop operation is bounded by p,eviously r submitted full power analyses. The maximum power level that can be attained with one-loop operation is only restricted by the MCPR and overpressure Ifmits estab-lished from a full power analysis.
In the following sections, three of the most limiting transients of core
- flow increase, pressurizat, ion, and flow decrease events are analyzed for single-loop operation. They are, respectively:
I
- b. generator load rejection with bypass failure, (LRNBP), and i
- c. one pump seizure accident. (PS)
The plant initial conditions are given in Table 15.C.3-1.
15.C.3.1.1 Feedwater Controller Failure - Maximum Demand 15.C.3.1.1.1 Core and System Performance Mathematical Model The computer model described in Reference 15.C.8 2 was used to simulate this event. .
HLV: rf:ge/F07184* 15.C.3-2
Input Parameters and Initial Conditions The analysis has been performed with the plant conditions tabulated in Table 15.C.3-1, except the initial vessel water level is at level set-point L4 for conservatism. By lowering the initial water level, more {
I cold feedwater will be injected before Level 8 is reached result.ing in l higher heat fluxes.
l End of cycle (all rods out) scram characteristics are assumed. The safety / relief valve action is conservatively assumed to cccur with higher ,
than nominal setpoints. The transient is simulated by programming an upper limit failure in the feedwater system such that 130% of rated feedwater flow occurs at the design pressure of 1065 psig.
Results The simulated feedwater controller transient is shown in Figure 15.C.3-2 for the case of 70.6% power 54.1% core flow. The high-water level turbine trip and feedwater pump trip are initiated at soproximately 4.2 seconds. Scram occurs simultaneously from Level 8, and limits the peak neutron flux. MCPR is considerably above the safety limit so no fuel failure due to boiling transition is predicted. The turbine bypass system opens to limit peak pressure in the steamline near the safety valves to 1045 psig and the pressure at the bottom of the vessel to about <
1059 psig.
I Consideration of Uncertainties All systems used for protection in this event were assumed to have the i
poorest allowable response (e.g., relief setpoints, scram stroke time, etc.) Expected plant behavior is, therefore, expected to lead to a less i i severe transient.
l HLV: rf:gc/F07184* 15.C.3-3 i
GGN5 15.C.3.1.1.2 Barrier Performance e noted above, the consequences of this event do not result in any temperature or pressure transient in excess of the criterla for which the fuel, pressure vessel, or containment are designed; therefore, these barriers maintain
- integrity and function as designed.
15.C.3.1.1.3 Radioloaical Consequences The consequences of this event do not result in any calculated fuel failures; however, radioactive steam is' discharged to the suppression pool as a result of SRV activation.
15.C.3.1.2 Generator Load Rejection With Bypass Failure 15.C.3.1.2.1 Core and System Performance Mathematical Model The computer model described in Reference 15.C.8-2 was used to simulate this event. .
Input Parameters and Initial Conditions These analyses have been performed, unless otherwise noted, with the plant conditions tabulated in Table 15.C.3-1.
- The turbine electro hydraulic control system (EHC) power / load imbalance devicedetectsloadrejectionbeforeasessurablespeedchangetakes place.
The closure characteristics of the turbine control valves are assumed l such that the valves operate in the full arc (FA) mode and have a full stroke closure time, from fully open to fully closed, of 0.15 second.
HLV:rf:gc/F07184* 15.C.3-4
( ._
GGNS Auxiliary power is independent of any turbine generator overspeed effects and is continuously supplied at rated frequency, assuming automatic f ast transfer to auxiliary power supplies, j l
The reactor is operating in the manual flow-control mode when load rejection occurs. Results do not significantly differ if the plant had been operating in the automatic flow-control mode.
Results The simulated generator load rejection without bypass is shown in Figure 15.C.3-3.
Table 15.C.3-2 shows for the case of bypass failure, peak neutron flux reaches about 70.7% of rated and peak steamline pressure at the valves reaches 1167 psig. The peak nuclear system pressure reaches 1179 psig at the bottom of the vessel, we'll below the nuclear barrier transient pressure limit of 1375 psig. The calculated MCPR is 1.41, which is well above the safety limit.
Consideration of Uncertainties The full-stroke closure rate of the turbine control valve of 0.15 second is conservative. Typically, the actual closure rate is approximately 0.2 second. The less time it takes to close, the more severe the pressuriza-tion effect.
All systems used for protection in this event were assumed to have the poorest allowable response (e..g, relief setpoints, scram stroke time, etc.). Expected plant behavior is, therefore, expected to reduce the actual severity of the transient.
15.C.3.1.2.2 Barrier Performsncu The consequences of this event do not result in any temperature or pressure transient in excess of the criteria for which the fuel, pressure HLVarfisc/F07184* 15.C.3-5
~~
GGNS vessel, or containment are designed and, therefore, these barriers maintain their integrity as designed.
15.C.3.1.2.3 Radioloateal Consecuences The consequences of this event do not result in any calculated fuel failures; however, radioactivity is nevertheless discharged to the suppression pool as a result of SRV activation.
15.C.3.1.3 Recirculation Pump Seizure Accident 15.C.3.1.3.1 Core and System Performance Mathematical Model The computer model described in Reference 15.C.8-3 was used to simulate this event.
Input Parameters and Initial Conditions i
This analysis has been performed, unless otherwise noted, with plant conditions tabulated in Table 15.C.3-1. For the purpose of evaluating consequences to the fuel thermal limits, this transient event is assumed to occur as a consequence of an unspecified, instantaneous stoppage of the active recirculation pump shaft while the reactor is operating at -71%
M8 rated power under single-loop operation. Also, the reactor is assumed to be operating at thermally limiting conditions.
The void coefficient is adjusted to the most conservative value; that is, the least negative value in Table 15.C.3-1.
Results
- Figure 15.C.3-4 presents the results of the accident. Core coolant flow drops rapidly, reaching a minimum value of 26% rated at about 1.3 seconds.
The minimum CPR value during the transient is 1.24 and poses no threats i
HLV: rf:gc/F07184* 15.C.3 6 l
_ -. - -. - ~,- - . - . . . . , . . _ . . _ . . . , . _ . . - . _ . . _ _ . - . - .
. --~ ~-- -
GGNS l
to thermal limits.
l 15.C.3.1.3.2 Barrier Performance The consequences of this event do not result in any temperature or pressure transient in excess of the criteria for which the fuel pressure vessel or containment are designed. Therefore, these barriers maintain integrity and function as designed. l 15.C.3.1.3.3 Radiological Consecuences The ennsequences of this event do not resu'It in any calculated fuel failures.
i 15.C.3.1.4 Suneary and Conclusions The transient peak value results are summarized in Table 15.C.3-3. The Critical Power Ratio (CPR) results are summarized in Table 15.C.3-3.
This table indicates that for the transient events analyzed here, the MCPRs for all transients are above the single-loop operation safety limit value of 1.07. It is concluded the thermal margin safety limits estab-lished for two-pump operation are also applicable to single-loop operation conditions.
For pressurization, Table 15.C.3-2 indicates the peak pressures are below the ASME code value of 1375 psig. Hence, it is concluded the pressure barrier integrity is maintained under single-loop operation conditions. ;
15.C.3.2 R0D WITHDRAWAL ERROR i
The rod withdrawal error (RWE) transient for two-loop operation documented in the main text of this chapter employs a statistical evaluation of the minimum critical power ratio (MCPR) and linear heat generation rate (LHGR) response to the withdrawal of ganged control rods for both rated and of f-rated conditions. The required MC'PR limit protection for the event is provided by the rod withdrawal limits (RWL) system. ${nce this HLV:rf:gc/F07184* 15.C.3-7
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GGNS analyses covered all off-rated condition in the power / flow operating map, single-loop operation is bounded by the current technica1' specification.
The Average Power Range Monitor (APRM) rod block system provides additional alarms and rod blocks when power levels are grossly exceeded. Modification of the APRM rod block equation (below) is required to maintain the two loop rod block versus power relationship when in one loop operation.
/
One-pump operation results in backflow through 12 of the 24 jet pumps while the flow is being supplied into the lower plenum from the 12 active jet pumps. Because of the backflow through the inactive jet pumps, the present rod block equation was conservatively modified for use during one-pump operation because the direct active-loop flow measurement may not indicate actual flow above about 36% core flow without correction.
A procedure has been established for correcting the APRM rod block equation to account for the discrepancy between actual flow and indicated flow in the active loop. This preserves the original relationship between APRM rod block and actual effective drive flow when operating with a single loop.
The two-pump rod block equation is:
RB = mW 4 RB - m(100) 100 The one-pump equation becomes:
RB = mW + RB - m(100) q6y 100 where Ay - difference between two-loop and single-loop effective drive flow at the same core flow.
HLVirfsge/F07184* 15.C.3-8 l
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GGNS R8 = power at rod block in X; n= flow reference slope W= drive flow in % of rated.
top level rod block at 100% flow.
RB100 =
If the rod block setpoint (RB100) is changed, the equation must be ,
l recalculated using the new value.
l The APRM scram trip settings are flow biased in the same manner as the APRM rod block setting. Therefore, the APRM scram trip settings are subject to the same procedural changes as the rod block settings discussed above.
~ I 15.C.3.3 OPERATING MCPR LIMIT For single-loop operation, the operating MCPR limit remains unchanged from the normal two-loop operation limit. Although the increased uncer- ;
i tainties in core total 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 limiting transients have been analyzed to indicate that there is more than enough MCPR margin during single-loop operation to compensate for this increase in safety limit. For single loop operation at off-rated conditions, the steady-state operating M;PR limit is established by the MCPR,and MCPRg curves. This ensures the 99.9% statistical limit requirement is always satisfied for any postulated abnormal operational occurrence. The abnormal operating transients analyzed concluded that current power dependent MCPR, limits are bounding for single loop operation. Since the maximum core flow runout during single loop operation is only about 54% of rated, the current flow ,
limits which are generated based on the flow runout up to l dependent MCPRf rated core flow are also adequate to protect the flow runout events during single loop operation. . .
i f
HLV:rf:gc/F07184* 15.C.3-9
GGNS 3
TABLE 15.C.3-1 I INPUT PARAMETERS AND INITIAL CONDITIONS FOR l TRANSIENTS AND ACCIDENTS FOR SINGLE-LOOP OPERATION
! 1. Thermal Power Level 2708 (70.6% Rated)
Analysis Value MWt 6
- 2. Steam Flow, Ib/hr ll.06x10
- 5. Feedwater Temperature, 'F 386 3
- 6. Vessel Do:ne Pressure, psig 981
- 7. Vessel Core Pressure, psig 985
- 8. Turbine Bypass Capacity, % NBR 35 i
- 9. Core Coolant Inlet Enthalpy, Btu /lb 509.5 J
j 10. Turbine Inlet Pressure, psig 946 l 11. Fuel Lattice 8x8R 1
- 12. Core Leakage Flow, % '10.65
- 13. Required MCPR Operating Limit 1.41("'
l 14. MCPR Safe.ty Limit for incident of
! Moderate frequency l First Core 1.07 Reload Core 1.08 4
l 15. Doppler Coefficient (-)c/'F
! Analysis Data 0.132(b) l'
- 16. Void Coefficient (-)t/% Rated Voids Analysis Data for Power Decrease Events 4.0(b) l Analysis Data for Power Increase Events 14.0(b
- 17. Core Average Void Fraction, % 41.9(b)
- 18. Jet Pump Ratio. M 3.521 -
f HLV:rf:ge/F07184* 15.C.3-10
~
j .
TABLE 15.C.3-1 (Continued)
'!9 . Safety / Relief Valve Capacity, % NBR 91145 psig 102.4 Manufacturer DIKKER Quantity Installed 20
- 20. Relief Function Delay, Seconds 0.4
- 21. Relief Function Response, Seconds 0.1 '
- 22. Setpoints for Safety / Relief Valves Safety Function, psig 1175, 1185, 1195, 1205, 1215 Relief Function, psig 1145, 1155, 1165, 1175
- 23. Number of Valve Groupings Simulated Safety Function, No. 5 Relief Function, No. 4
- 24. High Flux Trip, % NBR .
Analysis Setpoint (1.22 x 1.042), % NBR 127.2
- 25. High Pressure Scram Setpoint, psig 1095
. 1
- 26. Vessel Level Trips, Feet Above Separator Skirt Bottom l Level 8 - (L8), Feet 5.88 Level 4 - (L4), Feet 4.03 Level 3 - (L3), Feet 2.16 Level 2 - (L2), Feet -2.182
- 27. APRM Thermal trip Setpoint, % NBR 9 100% Core Flow 118.8
- 28. RPT Delay, Seconds 0.19
- 29. RPT Inertia Time Constant for Analysis, '
secs. 5
- 30. Total steamline volume, ft3 4358 l
(a)0peration operating limit is given by MCPR f for a core flow of 54.1%.
(b) Parameters used in Reference 15.C.8-3 analysis only. Reference 15.C.8-2 values are calculated within the code for end of Cycle 1 condition.
These are rated condition values.
1
. l l
HLV: rf:gc/F07184* 15.C.3-11 s
~+-,-y-w.~ - . - , ,,, , , , _ _ , _ ,
TABLE 15.C.3-2 .
SUp04ARY OF TRANSIENT PEAK VALUE RESULTS SINGLE-LOOP OPERATION MAXIMUM MAXIMUM MAXIMUM MAXIMUM NEUTRON DOME VESSEL STEAMLINE FLUX PRESSURE PRESSURE PRESSURE FREQUENCY
- PARA-FIGURE DESCRIPTION (% N8R) (psig) (psig) (psig) Category GRAPH Initial Condition 70.6 981 998 974 M/A f
Feedwat.er flow 79.2 1045 1059 1045 a 15.C.3.1.1 15.C.3.2 Controller
- Failure
- (Maximum Demand) .
Generator Load 70.7 1166 1179 1167 b 15.C.3.1.2 15.C.3.3
- I Rejection With Bypass Failure 15.C.3.1.3 15.C.3.4 Seizure of Active 70.6 984 998 976 c Recirculation
- Pump
! *a = Moderate frequency incident; b = infrequent; c = limiting faults
' 15.C.3-12
- HLV: rf:gc/F07191*
l
O ye GGNS TABLE 15.C.3-3 SgdMARYOFCRITICALPOWERRATIORESULTS-SINGLE-LOOP OPERATION FWCF LRNBT PS Initial Operating Condition 70.6/54.1 70.6/54.1 70.6/54.1
(% power /% flow)
Required Two Loop Initial MCPR 1.41 1.41 1.41 Operating Limit at SLO Condition (a)
ACPR 0.07(*) 0.00 ID) 0.17 Transient MCPR at SLO 1.34 1.41 1.24 SLMCPR at SLO" 1.07 1.07 1.07 Margin Above SLMCPR** 0.27 0.34 0.17 Frequency Category Moderate Infrequent Limiting frequent incident fault incident (a)value includes option A adder ID)ACPR is less than 0.002.
" Values shown for initial cycle. Add 0.01 for reload cycles.
- Reduce margin by 0.01 for reload safety limit increase.
i HLV:rf:gc/F07184* 15.C.3-13
.-...,.~e-------- . . _ . . . . ._. ,..., .
o .. . .
1230 1220 -
g without bypass
~
E 1200 -
E N
e 2 1190 n.
1180 >
.w with bypass 3
a.
1170 -
1160 1150 70 80 90 100 110 60 Initial Power Level (1 NBR)
FIGURE MISSISSIPPI POWER & LIGHT Peak Dome Pressure versus Initial Power 15.C.3-1 Level, Turbine Trip at E0EC
- _-_ w - --w--- w-w_---___wmm ,,,_,,,,,g,, .,,.____y, , , , , , , , , _ __ _ , , , , , , , _ _ _
1 LEVEL (INC H-REF-SEP-SMIRT 2 W R SENSE D LEVEL (INCHES) 3 N R SENSE D LEVEL (INCHES) 150. Il CORE INLE I FLOH (PCT) l 1 -
l 100. ';
l ,,
I -
t i
l'
- 50. -"
A 'N N 1
- ~ 4 i ''''
0"O ~ 5. 10. IS. 20.
l TIME (SEC) -
i l
FIGURE 15.C.3-2 MISSISSIPPI FEElWATER CONTROLLER FAILURE-MAXIMUM DEMAND SINGLE LOOP OPERATION. 71%P/54%F l POWER A LIGHT i
I NEUTFION F LUX
- 2 PEAK FUEL CENTE6 TEMP 3 RVE SURFF CE HEAT FLUX 150* I
\ 11 FEEDHATEF FLOH 5 VESSEL S1 ERM FLOW
)l .
2 o 100. -
W
- u. 1 35 2 o
H
" I f
$ 50. I 3' s 3 -
M taJ
('3 1 ,
- : ti . .
1 <
.. g,....l...
- 0. 5. 10. 15. 20. .-
- TIME (SEC) .
FEEINATER CONTl!OLLER FAILURE-MAXIMUM DEPAND SINGLE LOOP OPERATION, 71%P/54%F FIGURE 15.C.3-2 MISSISSIPPI POWER A LIGHT
4 L 1 VESSEL Pd ES RISE (PSI) ll 2 STH LINE PRES RISE (PSI) l 3 TURBINE F REG RISE (PSI 1 ig core IHLtli SUB IBTU/LB1 125.
' 5 P.ELIEF Vf LVE FLOH (PCT) i G TURO STEf 11 Fl.014 (PCT) i i n. s '
i! /
!: 3 '
j -
4 r >
it
= - -
L r 2s. m l ,
~
I 5 l ,, J2h3 6 -
l
-25.
l 0* ' ' ' ' s' . 10. 15. 20.
TIME (SEC1 i
FIGURE 15.C.3-2 FEEINATER CONTROLLER FAILURE-MAXIMUM DEMAND SINGLE LOOP OPERATION, 711P/54%F
. MISSISSIPPI i~ PfMER A LIGHT i
l i . _ _ _ _ __ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _
l I VOID REE TIVITY .
- 2 DOPPLER'F ERCTIVITY i
3 SCRAM REF CTIVITY
- 1. tt 10TAL REFCTIVITY i.
(
D. NI <)
I 1
. I t _
1 s -1, m _ -
i > ..
w- _
- o _
- cr .
LLI
! a- : , ,,
l -2. ~ 2 a' ' ' ' ' ' 10. 15. 20.
O. S- -
TIME (SEC1 i
j
! FIGURE 15.C.3-2 rttownita coNinottta rattunE-MaximirM DEMAND SINGLE LOOP OPERATION, 71%P/54%F
! mssissire:
j 90WIR A LIGHT f
(
j- .
i I
1 LEVEL (INC,H-REF-SEP-SKIRT 1 2 H R SENSE D LEVEL (INCHES) l
' 3 N R SENSED LEVEL (INCHES) 200. tFCORE INLE I FLDH (PCT) i 100. .
l .
1 g .
]~
~
!, _4 -
c D. _
j
.: -e-j -100. = ' ' ' ' ' ' y. 6. 8.
- 0. 2- ,
- . TIME (SEC1 1 .
i FIGURE 15.C.3-3 GENERATOR LOAD REJECTION, BYPASS FAILURE, SINGLE LOOP OPERATION, 711P/54tr gyssg35;ppy l POWER A LIGHT i
es
i
- i j' - .
! 1 NEUTRON F LUX
!! 2 PEAK FUEL CENTER TEMP i' 3 AVE SURFF CE HEAT FLUX 150* 4 FEE 0HATEF FLOH .
' 5 VESSEL ST ERM FLOH-
^ .
8 100. 7 ll w
I W '
E 4 E [
l 5
u 50-h
- f%
3 N -a "'
- r Wp -
a 3
! 5- : g f
I b -
\
l C
. hl I' --
\
l-g Ni l D.
~
1 '- I ?J ? N j'I 1 '^L b ^
- l 4. 6. 8.
! 0. 2. .
1 TIME (SEC)
I I
i I
FIGURE 15.C.3-3 GENERATOR LOAD REJECTION, BYPASS FAILURE, SINGLE LOOP OPERATION, 71%P/54%F
[' MISSISSIPPI
/ POWER A LIQlT l l l'
i 1 1 VESSEL Pp ES RISE (PSI) 2 STM LINE PRES RISE (PSI) 3 SAFETY VF LVE FLOH (e)
)
- - . 11 RELIEF VF LVE FLOW (el.
! 5 BTPRSS VF LVE FLOH (e) 6 TUR8 STEF H FLOH (PCT)
. J .
- 200. .
I l
100. i i . -
y -
-l e s 35
" s 35 4 0* 35' ..,
l 0. 2. .
11 . 6. 8. s
! TIME (SEC) l i
l l
l l
GENERAT08t LOAD REJECTION, BYPASS FAILURE, SINGLE LOOP OPERATION, 711P/54%F FIGURE 15.C.3-3 MISSISSIPPI POWER A LIGHT
l l
i l
i m, (J
mn e=
8 C
p 1 9
- >- >- 1
>- > W >-
Hww w e -e H > >
> L3 e . e e- . E p - >-
WWO O
- g vuu u. n 2 ,
E W uJ y, w e cc m s (C LL.!
Jr J .
h O CL G -G N
- n_ c:- .
CU g )
9 o v> ~C -
. = N ff) ::2' *
- I O w
_ i -
- n .
g
/ 5 g i m
.y . -
a F E d
=
f E
s .
E
. y _
d i C
g w
= E
O f
h K
a 1
e a e a , ae a n .
U
(*) 11IA113038 5
its 3*
35 c5 zu
! I LEVEL (INC H-REF-SEP-SKIRT 2 W R SENSE D LEVEL (INCHES)
- 3 N R SENSE D LEVEL (INCHES) 150. 4 CORE INLBT FLOH (PCT) ,
l 100.
i '
t l
! 50.
m L3 ~
'. q , 3 3
- Y-
=7 _
1 i
4 -_4 4 4 1
1 0*~iiili'
! O* 10. 20. 30. 40.
!f TIME (SEC1 l
)
i i
i FIGURE 15.C.3-4 l MISSISSIPPI PUMP SEl7URE SINGLE LOOP OPERATION, 71%P/54%F
< prMER A LIGHT 1 -
I NEUTRON F LUX 2 PEAK FUEL CENTER TEMP
! 3 RVE SURFF CE HERT FLUX 150. q FEE 0HRTEF FLOW I . 5 VESSEL S1 ERM FLON 1
8 100.
. w .
- E 1 m .
ge A e --
14 O-f 1" w 50. ww r- x -
-y_
, g [
, w _
f l S. : y d
- 0. 30. 40.
0.~ ' ' ' ' ' ' ' ' '10 . 20.
i
! TIME (SEC) l FIGURE 15.C.3 4 PUMP SE170RE SINGLE LOOP OPERATION, 711P/54%F
- POWER A LIGHT
l ji
~
. I VESSEL Pd ES RISE (PSil 2 STM LINE PRES RISE (PSI) 3 TURBINE F RES RISE (PSI) 125. 11 OlFFUSER FLOH I (f)
- 5 DIFFUSER FLOH 2 (%)
6 TURB STEf H FLOH (PCT) i i
). '
- 75. .
~
6 6 l 6
- 5 5 5 5 4 4 4 -
u i "
l 25. -
3 3 3 i
7 7, 2;
-a -s g, v . . . ....
! D. 10. 20. 30. 40. ' _~
TIME (SEC) -
l 4
i i
FIGURE 15.C.3-4 PUMP SE17URE SINGLE LOOP OPERATION, 71%P/54%F
[ MISSISSIPPI POWER A LIGHT
[
i i a.
9 C
u -
8 c .
T F-
>.>H F - w
>M-wwuu a _
0 a a Mw wer e w a mw ar a c o_ c -c
- a. x ocu o
> C Ln -
-NM T d en m
n u w
m
- .n W-
=W D or N -.
R e- .
E
. to u g 5 f 5
. =
i ) o 8
- a
. y
~
E C - E i -- -. . y en R
-+
. . T . I', , , , ' , o. E
. . . , g
" O y g G
($1 11IAI13838 ._5 13 3.e h5 CB
==
GGNS ._
15.C.4 STABILITY ANALYSIS 15.C.4.1 Phenomena
~
i The least stable power / flow condition attainable under normal operating conditions (both reactor coolant system recirculation loops in operation) occurs at minimum flow and the highest achievable power level. For all operating conditions, the least stable power / flow condition may correspond to operation with cne or both recirculation loops not in operation. The primary contributing factors to the stability performance with one or both recirculation loops not in service'are the power / flow ratio and the recirculation loop characteristics. At natural circulation flow the highest power / flow ratio is achieved. At forced circulation with one recirculation loop not in operation, the reactor core stability may be influenced by the inactive recirculation loop. As core flow increases in SLO, the inactive loop forward flow decreases because the natural circula-tion driving head 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 the same time the increased core flow results in a lower power / flow ratio which is a stabilizing effect. These two countering effects may result in 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.
As core flow is increased further during SLO and substantial reverse flow i
- 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 interacts with the jet pump suction flow of the active recirculation loop and increases the jet pump flow noise. This effect increases the total core flow noise which tends to drive the neutron flux noise. .
HLV:rf:gc/F07184* 15.C.4-1
_ _. :. - ..n._____ = . = - _ . - _ - - _ _ _ _ - _ . _ . _.__s .
GGNS To determine if the increased noise was being caused by reduced stability margin as SLO core flow was increased, an evaluation was perfomed which
- phenomenologically accounts for single loop operation effects on stability (Reference 15.C.8-4). The sedel predictions were initially compared to test data and showed very good agreement for both two loop and single i loop test conditions. An evaluation was performed to determine the effect of reverse flow on stability during SLO. With increasing reverse flow, SLO exhibited slightly lower decay ratios than two loop operation.
However, at low core flow conditions with no reverse flow, SLO was slightly less stable. This is consistent with observed behavior at stability tests at operating BWRs (Reference 15.C.8-5).
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 consistent with tests data (Reference 15.C.8-4). The results of these-l analyses and tests indicate that the stability characteristics are not '
significantly different from two loop operation. At low core flows, 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 even higher core flows with substantial reverse flow in the inactive recirculation loop, the effects of cross flow on the flow noise results in an increase in system noise (jet pump, core flow and neutron I
flux noise).
l 15.C.4.2 Compliance to Stability criteria
. 1 Consistent with the philosophy applied to two loop operation, the stability )
compliance during single loop operation is demonstrated on a generic be. 3. Stability acceptance criteria have been established to demonstrate compliance with the requirements set forth in 10CFR50, Appendix A, General Design Criterion (GDC) 12 (Reference 15.C.8-6). A generic analyses which covers those fuels contained in the General Electric Standard Application for Reactor fuel (Reference 15.C.8-7) has.been performed. The analyses demonstrates that in the ' event' lief t" cycle neutron flux oscillations occur within the bounds of safety syster l HLV:rf:gc/F07184* 15.C.4-2
- - - - - + - ,---,-m-a egw- , - - - - .,, w.--- +
wt- .=w- .er -p- - r%=p,-w ew- -
rw->==-m-s---ww---rw--*wr
- m w wP-+ --gs-we-sw-o re -mww-m -4r-*wer w-s*w---w-%w-w---=-m'tw-mmww- -
-<T m- s--m-m--
l l
1 l
GGNS intervention, specified acceptable fuel design limits are not exceeded.
Since the reactor core is assumed to be in an oscillatory -mode, the question of stability margin during SLO is not relevant from a safety standpoint (i.e. the analysis already assumes no stability margin).
The fuel performance during limit cycle oscillations is characteristically dependent on fuel design and certain fixed system features (high neutron flux scram setpoint, channel inlet orifice diameter, etc.). Therefore the acceptability of GE fuel designs independent of plant and cycle parameters has been established. Only those 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 adalyses (Reference 15.C.8-6) and therefore it is demonstrated that stability compliance criteria are satisfied during single loop operation. Operationally, the effects of higher flow noise and neutron flux noise observed at high SLO core flows are evaluated to determine if acceptable vessel internal vibration levels are met and to determine the effects on fuel and channel fatigue. However, these are not considered in the compliance to stability criteria but are instead addressed on a plant specific basis. These evaluationn are addressed in j Section 15.C.7.
A Service Information Letter-380, Revision 1 (Reference 15.C.8-8) has been developed to inform plant operators how to recognize and suppresa unanticipated oscillations when encountered during plant operation.
Evaluation of additional SLO test data taken from an operating BWR in late 1983 has been completed. Results of which have been documented in revision 1 of the reference 15.C.8-6 report (NEDE-22277-P-1). These efforts combined with the analyses previously documented in References 15.C.8-4 and 15.C.8-6 provide justification that GGNS can operate at the highest achievable power with a single recirculation loop 'in o'peration.
HLV:rf:ge/F07184* 15.C.4-3
. - - - u -- -- - - - - - - -
s GGNS l
15.C.5 LOSS-OF-COOLANT ACCIDENT ANALYSIS An analysis of single recirculation loop operation using the models and assumptions documented in Reference 15.C.8-9 was performed for GGNS.
Using this method, SAFE /REFLOOD computer code runs were made for a full spectrum of large break sizes for only the recirculation suction size breaks (most limiting for GGNS). Because the reflood minus uncovery time for the single-loop analysis is similar to the two-loop analysis, the maximum planar linear heat generation rate (MAPLHGR) curves were modified by derived reduction factors for use during one recirculation pump operation.
l 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. Hot node uncovered time (time between uncovery and reflood) for single-loop operation is compared to that for two-loop operation in Figure 15.C.5-1. l The total uncovered time for two-loop operation is 174 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 177 seconds and for the 100% DBA suction break. This is the most limiting l 2 '
break for single-loop operation. In both cases, the 1.0 ft suction break has a longer total uncovered time but results in a less severe PCT response due to a later uncovery time.
15.C.S.2 SINGLE-LOOP MAPLHGR DETERMINATION 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 temperature. Therefore, as noted as Reference 15.C.8-9, the one 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 factors for single-loop" operation. .
The most limiting single-loop operation MAPLHGR reduction factor (i.e.,
l I
HLV: rf: gc/F07184* 15.C.5-1 l
I yielding the lowest MAPLHGR) for GE6 8x8 retrofit-fuel is 0.86. One-loop operation MAPLHGR values are derived by multiplying the current two-loop MAPLHGR values by the reduction factor (0.86). As discussed in Reference l' 15.C.8-9, single recirculation loop MAPLHGR values are conservative when '
calculated in this manner. ,
15.C.5.3 SMALL BREAK PEAK CLADDING TEMPERATURE Section II.A.7.4.4.2 of Reference 15.C.8-9 discusses the low sensitivity of the calculated peak cladding temperature (PCT) to the assumptions used I
in the one pump operation analysis and the duration of nucleate boiling.
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, l the calculated PCT values for small breaks will be well below the 1404*F small break PCT value previously reported for GGNS, and significantly 1 below the 2200*F 10CFRSO.46 cladding temperature limit.
l l
8 I
i HLV:rf:gc/F07184* 15.C.5-2
- - - -, - - - - - , - - . , . . , ....,,---,,..,,,,.-.-_,,-,-,,.,,,,,,e ,,,-~.,,....-.,,,--.a,, , .,. -.,-, - , ./.,4~ . -
l l
l i - 300 s
! im -
~s s ,
i N #
Eo o
'N - x
/
Y *
% l'
/
N ~%,*4 i tae -
'. E g isFT2 sntAK y no -
Ie ,. _
s 1
" ~
Til0 LOOP ANALYSIS l 6,
- -- $1NGLE LOOP ANALYSIS
[
E.
I- an -
N o
! so -
e I I I I I I ' ' '
o to ao 30 ao ao 80 70 SD so ico l SRE AC AREA (% DSAl i
l Uncovered Time vs Break Area FIGURE MIS 5ISSIPPI POWER & LIGHT Suction Break. LPCS Failure 15.C.5-1 l
,.....-.----,--..,-,-.-...,_,-.-.,-..----.-,,..,---..n.-~-,-n,n.,,..----n...-.-,-v
---,re --~--v - - - <w-~--~~~-~~~*-%~ = ~ ~ ~ ~ ~ ~ '
~~
- - ~
GGNS 15.C.6 CONTAINMENT ANALYSIS A single-loop operation containment analysis was performed for GGNS based on a tcynding analysis performed for a standard BWR6 plant.' The peak wetvell pressure, peak drywell pressure, chugging loads, condensation f
oscillation and pool and swell containment load responses were estimated ovar the entire single-loop operation power / flow region.
f The analysis shows peak drywell and wetwell pressures for the worst single loop operation condition of 34.5 psia and 21 psia, respectively.
The corresponding differential peak drywell and wetwell pressures are
' 19.8 psig and 6.3 psig which is less than the 22 psig and 9.9 psig reported in Chapter 6. The chugging loads, condensation oscillation download and pool swell velccity evaluated at the worst power /ficw l condition during single-loop operation were also found to be bounded by the rated power analysis.
l l
l l
HLV: rf:gc/F07184* 15.C.6-1
,. ^ ^ ^ :^ -
GGNS 15.C.7 MISCELLANE0US IMPACT EVALUATION
Impact Evaluation The principal difference between single loop operation (SLO) and normal two loop operation (TLD) affecting Anticipated Transient Without Scram (ATWS) performance is that of initial reactor conditions. Since the SLO initial power flow condition .is less than the rated condition used for TLD ATVS analysis, the transient response is less severe and therefore beunded by the TLO analyses. All ATWS acceptance criteria are met during SLO. Therefore, SLO is an acceptable mode of operation for ATWS considera-tions.
15.C.7.2 Fuel Mechanical Performance I
The thermal and mechanical duty for the transients analyzed have been f~ evaluated and found to be bounded by the fuel design bases.
It is observed that due to the substantial reverse flow estabitshed during SLO both the 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 *15% of rated and the core plate differential pressure fluctuation should not exceed 3.2 psi peak to peak to be consistent with the fuel rod and assembly design bases.
j
. 15.C.7.3 Vessel Internal Vibration l
A recirculation pump drive flow limit will be imposed for SLO. The highest drive flow tested during the startup test program at GGNS that show acceptable vessel internal vibration criteria will be the drive flow limit for SLO.
A preliminary assessment has been made for the expected reactor vibration level durir.g SLO for GGNS.
HLV: rf:gc/F07184* 15.C.7-1 l
~ _
GGNS I
Sefore providing the results of the assessment, it is prudent to define l
the term " maximum flow" during balanced 2-loop operation and single loop operation. Maximum flow for two-pump balanced operation is equal to rated volumetric core flow at normal reactor operating con'itions. d Maxirum flow for single pump operation is that flow obtained with the recirculation pump drive flow equal to that required for maximum flow during two pump balanced operation. For rated reactor water temperature and pressure, this maximum flow for GGN5 is about 44,600 gpm.
2 During the GE BWR-6 jet pump development tests at GE test facility HF ,
the reactor internal components were subjected to the maximum flows, as defined above, for both two-pump balanced and single-loop operating conditions. All components were found to be within acceptance limits with the exception of in-core guide tube during single-loop operation.
Due to the non-prototypical configuration of the in-core guide tube supports at HF2, it was decided that no design changes need to be made.
~
Instead, the in-core guide tube was to be monitored for vibration response at the Kuo Sheng 1 plant. Startup tests at the Kuo Sheng I plant showed all components, including the in-core guide tube during single-loop operation, to have vibration levels within acceptance limits.
I I
From the above, it can be inferred that the vibration levels of the reactor internal components for GGNS would be expected to be within acceptance limits during single-loop operation with maximum flow as I
defined above. However, since GGNS reactor internals have extensive
- instrumentation, final and definitive conclusions can be arrived after vibration data acquisition and data reduction are completed.
HLV: rf:ge/F07184* 15.C.7-2
.-._._-..-e_.,. - _ _,. _ - - , , .--n,, -- --- , _ _ , , . - . . , - + - - - - , , . ,n- , , , , _ . , , . . , _ , - - ____,,__,..,_n.m,nm _ , . , , , , , , , _ - - - - ,
GGNS 15.C.8 REFERENCES
! 15.C.8-1 " General Electric BWR Thermal Analysis Basis (GETA8); Data, Correlation, and Design Application", NEDD-10958-A, January 1977.
15.C.8-2 " Qualification of the One-Dimensional Core Transient Model for Boiling Water Reactors", NED0-24154, October 1978.
i 15.C.8-3 R. 8. Linford, " Analytical Methods of Plant Transients Evaluation for the General Electric Boiling Water Reactor", MED0-10802, April 1973.
15.C.8-4 Letter, H. C. Pfefferten (GE) to C. O. Thomas (NRC), " Submittal of Response to Stability Action Item from NRC Concerning
- Single-Loop Operation," September 1983.
15.C.8-5 S. F. Chen and R. O. Niemi, " Vermont Yankee Cycle 8 Stability j and Recirculation Pump Trip Test Report", General Electric f Company, August 1982 (NEDE-25445, Proprietary Information).
15.C.8-6 G. A. Watford, " Compliance of the General Electric Boiling Water, Reactor Fuel Designs to Stability Licensing Criteria",
General Electric Company, December 1982 (NEDE-22277-P, Proprietary Information).
i 15.C.8-7 " General Electric Standard Application for Reload Fuel",
General Electric Company, January 1982 (NEDE-24011-P-A-4).
I 15.C.8-8 "8WR Core Thermal Hydraulic Stability" General Electric J Company, February 10, 1984 (Service Information Letter-380, Revision 1).
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- . 15.C.8 REFERENCES (Cont'd) 1
" General Electric Company Analytical Model for less-of-Coolant 15.C.8-9 Analysis in Accordance with 20CFR50 Appendix K Amenchent No. 2
- One Recirculation Loop out-of-Service". MED0-20566-2 Revision 1, July 1978.
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