ML20217G240
| ML20217G240 | |
| Person / Time | |
|---|---|
| Site: | Sequoyah  |
| Issue date: | 03/03/1997 |
| From: | Mark Miller, Parece M TENNESSEE VALLEY AUTHORITY |
| To: | |
| Shared Package | |
| ML19317C563 | List: |
| References | |
| 51-1266121NP, 51-1266121NP-R, 51-1266121NP-R00, NUDOCS 9708070132 | |
| Download: ML20217G240 (140) | |
Text
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51-1266121-00 Safety Evaluation of Reduced Ther=.al_ Design Flow Prephred for Sequoye.h Nuclear _ Plant Units 1 & 2 Tennes.Jee valley Authority 4
78b7 Prepared by:
Date:
Mark L. Miller, Principal Engineer 3 // ['J 7 Reviewed by:
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up; Date:
' Mark V, Parece.3 Senior Gpr Enginur Approved by:
_. /dj, [f Date: 8M/B7 V
/
t Framatome Technologies;Inc.
Lynchburg, VA 9700070132 970406 DR ADOCK 05000327 PDR
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Record of Revisions EE-Descriotion 00 Ori.ginal issue O
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. Purpose The~ Tennessee Valley Authority (TVA) plans to reduce the thermal design flow of the Sequoyah Nuclear-Plant, Units 1 &
2, to 348,000 gpm.
The purpose of this-document is to support this change in flow.
Each event in Chapter 15 of i'
the UFSAR, containment analyses, and equipment qualification aspects of the plant, are evaluated to assess the effect of-l the proposed change.
The objective is to show that all of the acceptance criteria will be met with the proposed l
change.
h For those cases-that are affected by the change in the thermal design flow, analysis is performed to quantify the effects.
These supporting analyses are presented'in the Mark-BW Fuel Assembly Topical Report for the Sequoyah Nuclear Plant -(Reference 1) where the reduced flow was.taken
- into. account.
For the events which are either bounded by these events or which are not adversely affected by the proposed changes, an evaluation is-perforded to ensure that the events will meet the relevant acceptance criteria, 4
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Background
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Framatome Cogema Fuels /Framatome Techriologies Inc. (FCF/FTI)
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performed a series of' safety analyses to support the fuel reload at'Sequoyah_(Reference 1).
As part of the reload effort, the limiting transients were analyzed assuming a 2-reduced thermal design flow of 348,000 gpm.
TVA desires to use these analyses as the basis for a change to the plant
. Technical. Specifications.
This document provides the basis for changing the thermal e
design flow in the plant Technical Specifications to 348,000 gpm.
Specifically, each accident analysis in the plant FSAR (Reference-2) is evaluated with respect to the-reduced flow.
In.each case it is determined that all applicable acceptance criteria are met.
This effort augments these studies via a l _
full examination-of' events analyzed for Chapters 6 and 15 of l
the FSAR (Reference 2).
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- 3.~ 0 Scope ~of Evaluation All of:the Sequoyah'FSAR Chapter 15 edents, containment and equipment qualification studies are examined-in this document relative to the proposed change in thermal design flow.-
Events that were analyzed in support of the sequoyah reload (Reference 1),'already consider the reduced flow and
+
l provide.a. basis for discussing the effects of the change.
The reduction in thermal design flow is validated by j
c)mparison of each FSAR event with the limiting analyces performed in the Reference 1 reload report.
In general, the i -
reduced reactor-coolant flow translates to-a reduction in the calculated margin to DNB.
The evaluations are arranged
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in the order of the FSAR Chapter 15 discussions.
The 4
Condition I, II, III and IV events are grouped together in the evaluation under similar headings.
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.4.0 Safety Evaluation 4-4;0.1 --ConditioniI Events-l
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- Condition'I events are bounded by events that are classified as Condition II, III & IV since the severity o5 Condition I
- events.is minor in comparison.
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Condition II Events Condition II events are those events shat could occur with
- mocerate. frequency.
These events will not. result in fua1
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failures and an adequate margin =to system safety limits will be maintained.
In addition, these events may not cause
. failures that could lead to Condition III or Condition IV events.
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I 4.0.2,1 Uncontrolled Rod Cluster Control Assembly l
Kithdrawal from a Subcritical Condition I
A rod cluster control assembly (RCCA) withdrawal is defined i
i as an uncontrolled addition of reactivity by a withdrawal of RCCAs that results in a power excursion.
A continuous withdrawal of an RCCA is considered unlikely.
However, the reactor protection system (RPS) is designed to terminate any such transient before fuel thermal limits are reached.
- he power excursion resulting from an RCCA withdrawal causes a primary-to-secondary heat mismatch.
As the primary system i
heats up, a positive moderator temperature coefficient would result in additional reactivity insertion.
Taken alone, the pcwt, increase, coupled with the coolant temperature increase, reduces the margin to departure from nucleate boiling (DNB).
The reactor coolant system (RCS) pressurizes until reactor trip on power range flux level (low setting) occurs.
Reactor trip reduces core power and reduces the it 5
mismatch.
RCS pressure peaks and then declint
.ae to steam relief through the pressurizer safety valves.
Long-term heat removal is established via auxiliary feedwater.
An uncontrolled RCCA withdrawal from subcritical is a l
Condition II event.
The acceptance criteria for this event are:
l 1.
Peak primary and secondary system prsasu;es shall not exceed 110% of the design pressures.
2.
Fuel clad integrity shall be maintained by ensuring that the minimum departure from nucleate boiling racio (DNBR) remains above the 95/95 DN3R limit for the correlation used.
I This event is analyzed in the UFSAR, Reference 2, assuming the operation of only two reactor coolant pumps - the minimum DNB margin is, therefore, cons 3rvatively determined because Sequoyah technical specifica?. ion. require four reactor coolant pumps to be operating in operational modes 1 and 2.
Because of this assumption the reduction of RCS thermal design flow rate, associated with tha operation of all four reactor coolant pumps, will not affect this analysis.
The licensing basis analysis for rod withdrawal from suberitical, as presented in Section 15.2.1 of Reference 2, is therefore equally appl'icable for the cperation of Sequoyah with a thermal design flow of 348,000 l
gpm.
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conclusion 1
e The margin of safety. associated with a'n uncontrolled RCCA t
withdrawal from suberitical is not affected by a reduction i
in the thermal design flow.
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4.0.2.2 Uncontrolled' Rod Cluster Assembly Withdrawal at F
Power j
An uncontrolled RCCA withdrawal at power results in an increase in core power.
Because the heat removal via the steam generators remains relat!vely constant, there is a net j
increase in reactor coolant temperature.- Unless terminated by-automatic action, this power mismatch and resultant
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coolant temperature rise would eventually result in DNB.
The RPS is designed to terminate any such transient with an
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j The uncontrolled withdrawal of an RCCA at power.cvent is a l
Condition II event with the following acceptance criteria:
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. 1.
Peak primary and secondary system pressure shall no i
exceed 110% of design value, y
l 2.
Fuel cladding integrity shall be maintained by ensuring l
that-the minimum DNBR remains above the 95/95 DNBR limit for the correlation used.
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The parameters affecting the plant response to an RCCA withdrawal from power include the reactivity insertion due J
l to rod motion, axial power distribution, moderator reactivity feedback and Doppler reactivity feedback.
These i
i parameters are unaffected by'RCS flow.
A reduction in core i
i flow may, however, reduce the predicted margin to DNB.
The uncontrolled rod withdrawal.at power event is analyzed in Reference 1 in support of Framatome Cogema Fuels tuel i
loading at Sequoyah (Section 6.2.2). -The analysisLutilized a reduced RCS thermal design flow.
The results demonstrate that the primary pressure does not exceed the design limits, j-and shows that an adequate DNB margin is retained.
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l Conclusion The uncontrolled rod withdrawal at power event was analyzed in Reference 1.
The results of the analysis demonstrate that all of the acceptance criteria are successfully met for operation of S2quoyah with a thermal design flow of 349,000
-gpm.
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l 4.0.2.3 Rod Cluster Control Assembly Misalignment RCCA misalignment accidents include:
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- 1. A dropped control rod assembly
- 2. A dropped control rod assembly bank
- 3. Statically misaligned control zod assembly l
The safety parameter of importance in this event is the DNB margin in the het channel associated with potentially increatied peaking factors. A reduction in core flow can reduce the calculated margin to DNB.
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The RCCA misalignment event is analyzed each fuel cycle as a l
part.of the core design process.
The current RCS thermal design flow is an input to the process.
Constraints in-peaking ascociated with the reduced thermal design flow are included-in the core design associated with the first loading of FCF fuel loading at Sequoyah. 'In this manner an adequate DNB margin'is assured.
i conclusion A reductioniin thermal design flow affects the calculated l
minimum DNBR associated with an RCCA misalignment event.
L The event is considered in the development of core peaking limits as part of the core design.
It is, therefore,_
assured that the acceptance criteria for this event are-successfully met for operation of.Sequoyah with a thermal design flow of 348,000 gpm, l
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4.0.2.4 Uncontrolled Boron Dilution The chemical and volume control system (CVCS) regulates both the chemistry and the quantity of coolant in che RCS.
Inadvertent operation of the CVcs resulting in injection of deborated water into the RCS would cause a reduction in RCS boron concentration.
Boron dilution in any operational mode adds positive reactivity to the core.
Depending on the time given to terminate the dilution, the resulting positive reactivity insertion could result in a power excursion and challenge core thermal margins.
The inadvertent boron dilution of the RCS via the CVCS is a condition II event and has the following acceptance criteria 1.
Peak primary and secondary system pressure shall not exceed 110% of design value.
2.
Fuel cladding integrity shall be maintained by ensuring that the minimum DNBR remains above the 95/95 DNBR limit for the correlation used.
Because the boron dilttion event is characterized by a relatively slow rise in core power, any pressure response associated with the baron dilution event is bounded by the loss of electric load (LOEL) event.
The LOEL event is initiated at power by the occurrence of an abrupt loss of load and delayed reactor trip resulting in an interruption of RCS heat removal. LOEL is analy:ed in a manner that would maximize system pressurization.
The LOEL event was analyzed in support of FCF fuel reload at Sequoyah (Section 6.2.7 of Reference 1).
The analysis assumed a reduction in thermal design flow to 348,000 gpm.
Acceptable margin to system pressure limits are demonstrated for the LOEL event.
The parameters that dominate in the plant response to this event are the minimum shutdown margin, dilution flow rate, critical boron concentration, and RCS volume.
None of these parameters are affected by a reduction in the thermal design flow. A reduced flow may, however, reduce the predicted margin to DNB for the boron dilution event.
Boron dilution is prevented during the r'efueling mode of plant operation by administrative contrdis.
Dilution may occur in the remaining plant modes and are analyzed for startup, shutdown, and power operation in Reference 2 (Section 15.4.2).
For startup and shutdown conditions the margin to DNB is not explicitly calculated; an adequatu margin to DNB is assured by determining that the operator has time to mitigate boron
dilution prior a loss of shutdown margin.
A reduction in thermal design flow has no effect on the event sequence for a baron dilution from either startup or shutdown.
The analyses presented in Reference 2 for this event is, therefore, equally applicable to plant operation with a reduced thermal design flow.
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The boron dilution at power is bounded by the RCCA l
withdrawal at power spectrum analyses.
The reactivity addition for boron dilution (2.5 pcm/s - page 15.2-15 of Reference 2) is within the range of. reactivity addition I
rates (1 - 75 pcm/s) analyzed for the ROCA withdrawal at 1
power.-
The RCCA withdrawal at power was reanalyzed in support of 1
FCF fuel reload at Sequoyah (Section 6,2.2 of Reference 1).
The analysis assumed-the reduction in thermal design flow.
l Acceptable margin to DNB was demonstrated for the RCCA withdrawal at power event over the entire range of reactivity insertion-rates.
Conclusion l
1 system pressure responses to boron dilution are bounded by
- LOEL.
Reactivity addition rates and, therefore, DNB margins for boron dilution are within the range of rates analyzed for RCCA withdrawal at power.
Both the LOEL and RCCA withdrawal from power events were analyzed in Reference 1.
-The analyses of Reference 1 indicate acceptable system pressure and-DNB margins.
It is, therefore, assured that the-acceptance criteria for the boron dilution event are i
successfully met for operation of Sequoyah with a thermal design flow of 348,000 gpm.
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i 4.0.2.5 Partial Loss of Forced Reactor Coolant Flow A partial loss of forced reactor coolant flow accident can be caused by a mechanical or electrical failure in a reactor coolant pump, or by a fault in the power supply to the pump or pumps supolied by a reactor coolant pump bus.
If the reactor is a power at the time of the accident, the l
immediate effect of the partial loss of coolant flow is a rapid increase in the coolant-temperature.
This increase could result in DNB with subsequent fuel damage if the reactor is not promptly tripped.
The necessary protection against a partial loss of coolant flow accident is provided i
by the low prjnary coolant flow reactor trip signal which is actuated in ar.y reactor coolant loop by redundant low flow signals.
The partial loss of forced reactor coolant flow is a condition II event and has the following accectance criteria:
1.
Peak primary and secondary system pressure shall not exceed 110% of design value.
2.
Fuel cladding integrity shall be maintained by ensuring that the minimum DNBR remains above the 95/95 DNBR limit for the correlation used.
A similar event to the partial loss of forced reactor coolant flow is the complete loss of forced reactor coolant flow ever.t.
A complete loss flow is more severe than a
. partial.oss of flow event in both system pressure and DNB t
response.
This is because a complete loss of flow results in both a maximized reduction in RCS heat sink and a maximized reduction in core flow prior to reactor trip.
This relationship between partial and-complete loss of flow is not affected by the assumed thermal design flow.
The complete loss of forced reacter coolant' flow event was analyzed in support of FCF fuel reload at Sequoyah (Section 6.3.2 of Reference 1).
The analysis assumed a reduction in thermal design flow.
Acceptable margin to system pressure limits and DNB were demonstrated for the complete loss of forced reactor coolant flow event.
I Conclusion This event is bounded by the complete ;1o'ss of forced reactor coolant flow event.
Although the complete loss of flow is a Condition III event, it has been successfully analyzed to meet condition II acceptance criteria. It is, therefore, assured that the acceptance criteria for the partial loss of forced reactor coolant flow event are successfully met for operation of Sequoyah with a thermal design flow of 348,000 gpm.
4.0.2.6 Startup of An Inactive Reactor Coolant Loop sequoyah Technical Specifications reqtitre that all four reactor pumps be operating while the reactor is critical.
Consequently, this is not a credible event, and it is not evaluated.
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4.0.2.7 Loss of External Electric Load and/or Turbine Trip The loss of electric load event is initiated by a loss of l
external load or a turbine trip.
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is available and reactor coolant pumps continue to operate.
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The loss of load results in an abrupt reduction in RCS heat, 1
sink and subsequent RCS heatup.
RCS heat removal is established when the secondary safety valves open.
On the primary side, the pressure will increase and the pressurizer safety valves will open to maintain the pressure within the design limits.
The RPS will act to trip the reactor prior to a loss of DNB margin.
The LOEL event is the most severe overheating Condition II event.
The acceptance criteria for this event are 1.
Peak primary and secondary syrtem pressures shall not exceed 110% of the design pressures.
l 2.
Fuel clad integrity shall be maintained by ensuring
-that the minimum DNBR remains above che 95/95 DNBR limit for the correlation used.
The LOEL event was analyzed in Reference 1, Section 6.2.7, in support of FCF fuel loading at Sequoyah.
The analysis utilized a reduced RCS thermal design flow.
System pressures do not erceed the design limits subsequent to an l
LOEL.
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Because the primary pressure increases during this event and
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.the reactor coolant pumps remain in operacion, minimum DNBR for the LOEL event is bounded by those predicted for a complete loss of coolant flow event.
The complete loss of forced reactor coolant flow event was analyzed in support of FCF fuel reload at Sequoyah (Section 6.3.2 of Reference 1).
The analysis assumed a reduction in thermal design flow.
l Acceptable margin to DNB was demonstrated for the complete loss of forced reactor coolant flow event.
Therefore, no fuel pins will experience DNB following a LOEL with reduced thermal design flow, conclusion-The minimum DNBR for the LOEL event is bounded by the DNBR for the complete loss of forced reactor. coolant-flow event.
Although the complete loss of flow is a condition III event, it was successfully analyzed to meet the.' Condition II i
acceptance criteria with a-reduced thermal design flow.
The LOEL event, itself, was reanalyzed in Reference 1.
The reanalysis demonstrates the system pressure acceptance
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criterion is met for an LOEL.
It is, therefore, assured that all.of the acceptance criteria for this event are successfully met for operation of'Sequoyah with a thermal design flow of 348,000 gpm.
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l 4.0.2.8 Loss of Normal Feedwater A loss of normal feedwater results-in a reduction in secondary liquid-inventory.
As inventory is depleted, the RCS heats up and the primary system pressurizes.
If the reactor is tot promptly tripped during this accident, l
primary plant damage could occur due to the loss of heat i
sink.
I The loss of normal feedwater flow is a condition II event and has the following acceptance criteria:
1.
peak primary and secondary system prensure shall not exceed 110% of design value.
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-2.
Tuel cladding integrity shall be maintained by ensuring that the minimum DNBR remains above the 95/95 DNBR limit for the correlation used.
j In addition, the following criteria must be met by the loss of feedwater eventi-3.
Assurance of an ultimate heat sink for decay heat removal t
4.
Assurance that the pressurir.er does not fill liquid-solid.
Reactor and turbine trip occur coincidentally in the loss of feedwater event.
In comparison, the LOEL event is initiated by an abrupt cessation of steam flow and subsequent reduction in primary heat sink capacity at full power (i.e.,
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the reactor trip is delayed).
Therefore, the peak RCS and secondary system pressure for the loss of feedwater event-
.are bounded by system pressures predicted for the LOEL event.
This relationship is unaffected by a reduction in thermal design flow.
The LOEL event was analyzed in Reference 1, Section 6.2.7, in support of FCF fuel loading at Sequoyah.
The analysis utilized a reduced RCS thermal design flow.
System pressures do not exceed the design limits subsequent to an LOEL.
Because the-primary pressure increases-during this event and the reactor coolant pumps remain in oper,ation, DNB margins associated with a loss of feedwater ar,e. bounded by those predicted for a complete loss of coolant' flow event.
The complete loss of forced reactor coolant flow event was analyzed in support of FCF fuel reload at Sequoyah (Section 6.3. 2 of. Reference 1).
The analysis assumed a reduction in
' thermal design flow.
Acceptable sargin to DNB was
demonstrated for the complete loss of forced reactor coolant flow event.
The loss of feedwater analysis in tha sequoyah UFSAR indicates that the pressurizer liquid volume peaks near the time of reactor trip on low-low st.eam generator level (Figure 15.2.8-2 of Reference 2).
This peak represents the smallest margin to pressurizer fill.
The parameters affecting this peak are heat load, average RCS temperature, secondary inventory, low steam generator level setpoint, and valve capacities.
These parameters are not affected by RCS flow.
A second, lesser, peak in pressurizer liquid volume is reported in the UFSAR late in the loss of feedwater transient.
Pressurizer liquid volume increases because of a mismatch in core decay heat production and secondary heat
- removal after two feedwater-starved steam generators boil dry.
Pressurizer level increases until the heat addition from reactor coolant pumps and core decay heat is matched by the heat removal capability of auxiliary feedwater addition to the two remaining steam generators.
After the heat production is balanced by auxiliary feedwater, the RCS begins to cool and the pressurizer level decreases.
Conclusion The loss of feedwater system pressure response is bounded by.
the LOEL event response.
The analysis in Reference 1 demonstrates the system pressure acceptance criterion is met for an LOEL.
The loss of feedwater event is bounded, in DNB, by the complete loss of forced reactor coolant flow event.
Although the complete loss of flow is a condition III event, it has been successfully analyzed in Reference 1 to meet Condition II acceptance criteria with redQced thermal design flow.
Long term heat removal and pressurizer filling is dominated by the auxiliary feedwater capacity which is not affected by RCS flow.
Therefore, it is assured that all of the acceptance criteria for this event are successfully met for operation of sequoyah with a thermal design flow of 348,000 gpm.
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4.0.2.9 Loss of Offsite Power to Statica Auxiliaries i
A loss of offsite ' power to the station auxiliaries results 4
in a reduction in secondary liquid inventory.
As inventory i
is depleted, the RCS heats up and the primary system j
pressurizes.
If the reactor is not promptly tripped during
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this accident, primary plant damage could occur due to the loss of heat sink.
g Prior to reactor trip, the loss of offsite power to the j
station auxiliaries event is identical to the loss of feedwater. event.
At reactor trip, the reactor coolant pumps trip.
The remainder of the event tests the capability of 1
i auxiliary feedwater to remove decay heat via natural circulation RCS flow.
The loss of offsite power to the station auxiliaries event 1
is a Condition II event and has the following acceptance
' criteria:
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1.
Peak primary and secondary system presture shall not
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exceed 110% of design *talue.
2.
Fuel cladding integrity shall be maintained by ensuring that the minimum DNBR remains above the 95/95 DNBR limit for the correlation used.
In addition, the following criteria must be met by the loss of offsite power event:
3.
Assurance of an ultimate heat sink for decay heat removal.
4.
Assurance that the pressurizer does not fill liquid-solid.
Reactor and turbine trip occur coincidentally in the loss of feedwater event.
In comparison, the LOEL event is initiated by an abrupt cessation of _ steam flow and subsequent reduction in primary heat sink capacity at full power (i.e.,
the reactor trip is delayed). -Therefore, the peak RCS and secondary system pressures for the loss of offsite power event is bounded-by system' pressures predicted for-LOEL.
This relationship is unaffected by a reduccion in thermal design flow.
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The LOEL-event was analyzed in Reference'1, Section 6.2.7, in support of FCF fuel-loading at Sequoyah.
The analysis utilized'a reduced RCS thermal-design.- System pressures do not exceed the design limits subsequent to an LOEL.
Both the reactor and reactor.coola.nc pumps trip coincidentally in the loss of of f s'ite power event.
In comparison, the pumps trip prior to reactor trip in the
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complete loss of coolant flow event.
Since the power to flow ratio is greatsr for the latter event, DNB margins associated with an loss of offsite power are bounded by j
those predicted for n complete loss of coolant flow event.
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i The complete loss of forced reactor coolant flow event was analyzed in support of FCF fuel reload at Sequoyah (Section 6.3.2 of Reference 1).
The analysis assumed a reduction in thermal design flow.
Acceptable margin to 7NB was demonstrated for the complete loss of forced reactor coolant i
flow event.
l The loss of offsite power analysis in che Sequoyah UFSAR indicates that the pressuri:er liquid volume peaks near the time of reactor trip on low-low steam generator level (Figure 15.2.8-6 of Reference 2).
The parameters affecting the peak are heat load, average RCS temperature, secondary inventory, low steam generator level setpoint, and valve capacities.
These parameters are not affected by RCS flow.
The second peak, late in the ' loss of of fstte power event, represents the minimum margin to pressurizer fill.
. pressurizer liquid volume increases relative a mismatch in core decay heat production and secondary heat removal af ter two feedwater-starved steam get*erators boil dry.
Pressurizer level increases until the heat addition from core decay heat is matched by the heat removal capability of i
a single motor driven auxiliary feedwater pump to the two remaining steam generators.
After.the heat production is balanced by auxiliary feedwater, the RCS begins to cool and the pressurizer level decreases.
1 The reactor coolant pumps coast down within approximately l
one minute af ter reactor and pump trip.
Following coast l
down, the RCS flow is controlled by the localized cooling mechanics of the steam geterator tubes.
A reduction in thermal design flow, therefore, has an insignificant effect on the long term removal of heat from the RCS via natural circulation cooling with auxiliary f eedwater.
In addition, the minimum margin to pressnizer fill is unaf fected by the reduced flow, l
conclusion The system pressure response is bounded,tur the LOEL event response. The reanalysis of Reference 1 demonstrates the system pressure-acceptance criterion is met for an LOEL.
The loss of feedwater event is boundsd',in DNB, by the complete loss of forced reactor coolant flow event.
Although the complete loss of flow is a Condition III event, it-has been successfully reanalyzed in Reference 1 to meet the acceptance Condition II acceptance criteria.
Following reactor coolant pump trip and coas-cdown, the reduced thermal 7
design flow has an insignificant effect on RCS flow.
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It is assured-that all of the acceptance criteria for this i
event are successfully met for operation of Sequoyah with a i'
thermal design flow of 348,000 gpm.
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4.0.2.10 Excessive Heat acmoval Due to feedwater System l
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Excessive feedwater flow could be caused by a full opening
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j of one or more feedwater regulator valves due to a feedwater control system malfunction or an operator error.
At power i
2 this excess flow causes a greater load demand on the RCS due i
i to increased subcooling in the steam generators.
At no-load l
conditions, che addition of cold feedwater may cause a l
decrease in RCS temperature and thus a reactivity insertion i
i due to the ef fects of the negative moderator coef ficient of I
l reactivity.
Overcooling for the event is mitigated by the -
j steam generator high high level trip, which closes all main I
i feedwater isolation valves, trips the main feedwater pumps, and trips the turbine.
I The excessive heat removal due to feedwater system malfunction event is a Condition II event and has the following acceptance criteria:
l
- 1.. Peak primary and seconda:y system pres %ure shall not
. exceed 100% of d9aign value.
l 4
i 4
2.
Fuel cladding integrity shall be maintained by ensuring i
that the minimum DNaa remains above the 95/95 DNBR limit l
for the correlation used.
l l
Prior to the high-high secam generator water level signal,
{
both the RCS and steam generator secondary are cooled by the 1
l-excessive feedwater and the respective system design pressures are not challenged.
Turbine trip occurs first and l
}
an imbalance in-heat production / removal is indicated in the i
i UFSAR analysis for this event that is. ass:ciated with a 2.5 i
i.
second delay _between the turbine trip on level and the 1'
reactor trip (Table 15. 2-1).
The delay between turbine trip and reactor trip for the LOEL transient is on the order of 6 to 8 seconds.
The LOEL is, therefore, bounding in system pressure response.
The LOEL event was analyted in Reference 1, Section 6.2.7, l
l in support of FCF fuel leading at Sequoyah.
The analysis _
j utilized a reduced RCS-thermal design.
System pressures do not, exceed the design limits subsequent to an LOEL.
i l _
The analysis of feedwater malfunction event is presented in Section 15.2.10 of the Segynyah UFSAR.
Rain feedwater i
_ addition to the steam generators is increased by a l
]
conservatively prescribed amount _ to characterize feedwate. -
regulator valve failure.
A reduction in RCS thermal design flow'does not affect the-initial-RCS average temperature or the reactivity feedback parameters.
The RCS cooling, and j
the increase power that results, is proportional to the j-increase -in feedwater flow and if independent of RCS flow.
Because the thermal design flow is being reduced, however, j;
t k
l_______~m.--__A.-._,.,._.._.__,__-,...m..m__
-.__......_._._._._._,.--a
the calculated margin to DNB for this event could be reduced.
The excessive feedwater event results in an increase in core power and the establishment of a new stable power level prior to reactor trip.
The event is mitigated via reactor trip and feedwater isolation on high - high steam generator level, independent of thermal design flow.
MAP limits have been developed at operational statepoints associated with overtemperature and overpower AT trip boundaries and verify l
that adequate DNB margins exist at these statepoints.
Since, for this transient, the resulting increase in power is not high enough to cause an overtemperature or overpower AT trip, there is no challenge to the fuel thermal limits.
Consequently, with respect to minimum DNBR, the excessive feedwater event is bounded by the rod withdrawal at power event.
The RCCA withdrawal at power event was analyzed in Reference 1,
Section 6.2.2, in support of FCF fuel loading at Sequoyah.
The analysis utilized a reduced RCS thermal design flow.
An adequate DNB margin was demonstrated for the RCCA withdrawal at power event.
- Conclusion The system pressure response of the excessive feedwater event is bounded by LOEL.
The reanalysis of Reference 1 demonstrated the system pressure acceptance criterion is met for an LOEL.
With respect to minimum DNBR, the excessive feedwater event is bounded by an RCCA withdrawal at power.
The reanalysis in reference 1 indicates a sufficient margin to DNB for an RCCA withdrawal at pcwer.
It-is assured that all of the acceptance criteria for the excessive feedwater event are successfully met for operation of sequoyah with a-
- thermal design flow of 348,000 gpm.
t t
k t
t J
f
-a-,e.e-----_s-
4.0.2.11 Excessive Load. Increase Accident An excessive load increase incident is defined as a rapid increase in steam generator steam flow that causes a mismatch between the reactor power and the steam generator heat removal.
The resultant cooling of the reactor primary system fluid causes an increase in reactor power due to a negative end-of-cycle moderator temperature coefficient.
It also causes a reduction in primary system pressure due to j
the contraction of the reactor coolant.
The increase in i
power and. decrease in primary system pressure produce a I
reduction in DNBR.
l The excessive load increase event is a Condition II event and has the following acceptance criteria:
1.
Peak primary;and secondary system pressure shall not exceed 110% of design value.
1 2.
Fuel cladding integrity shall be maintained by ensuring
{
that the minimum'DNBR remains above the 95/95 DNBR limit i
for the correlation used.
r An-excessive load increase-causes an overcooling and depressurization of both the RCS and steam generator secondary systems.
The respective system design pressures are not challenged.
The LoEL, therefore, bounds this event in system pressure response.
The LOEL event was reanalyzed in Reference 1, Section 6.2.7,
[
in support of FCF fuel loading at Sequoyah.
The reanalycis
~
' utilized a reduced RCS thermal design, System pressures do not excaed the design limits subsequent to an LoEL.
UFSAR Section 15.2.11 (Reference 2) reports the analysis of the excessive-load increase. The excessive load increase event results in an increase in core power and the establishment of a new-stable. power level.
The analysis results demonstrate that a reactor trip on overtemperature aT, overpower AT, or high nuclear flux does not occur for 3
, this transient. MAP limits have been developed at i
operational statepoints associated with overtemperature and
}
overpower AT trip boundaries and verify that adequate DNB j
margins exist at these statepoints.
S ine'e, for this i
transient, the resulting increase.in power is not high enough to cause an overtemperature or; overpower AT trip, there is no challenge to'the fuel' therma'l limits, t
Correspondingly,,an adequate margin-to DNB is assured.
i A reduction in thermal design flow can only affect the rate at which the primary system responds to an increase load.
The' initial core power, RCS average temperature, and
- ?
j reactivity feedback paraneters are unchanged.
The increase i
e --,---,,
,,,,-w---,nc,n,w
~n n
.-,,,--~--.-m,
,m.,
n-,,,,ww
-, +
-n,-
..--. - ~
m,-,rr-.
i t. steam tiow (to 110 percent) for this transient is a fixed, prescribed, input.
core power leve.ls would be identical,The ultimate primary cooldown and therefore, at the reduced RCS ficw.
As a result, the reactor would not and there would be no challenge to the fuel thermal limitstrip for the excessive load increase avent.
C2nabl1Lqu The system pressure response of the excessive load increase event is bounded by an LOEL. The reanalysis of Reference 1 demonstrates that the system pressure acceptance criterion is met for an LOEL.
The r.ference analysis demonstrates that a reactor trip does not occur for a 10 percent load increase.
A reduction in RCS f}ow would not affect this result.
A margin to both overtemperature oT and DNE is, therefore assured, The acceptance criteria for the excessive load increase event are successfully met for operation of Sequoyah with a thermal design flow of gpm.
348,000 4
4
=-
i 4.0.2.12 Accidental Depressucization of the Reactor Coolant i
System j
The accidental depressurization,f the RCS is initiated by
{
the inadvertent opening of a pressurizer safety valve or by
)
the failure of a vsive to close following an j
overpressurizatien transient.
The event can cause a i
reduction in reactor coolant inventory and subsequent i
reduction in RCS pressure.
If the valve is not closed, the i
continuing depressurization leads to a reactor trip on low
]
RCS pressure or overtemperature 67.
The accidental RCS depressurization event is a condition II i
event-and has the following acceptance criteria:
1.
Peak primary and secondary system pressure shall not
'7 exceed 100% of design value.
- 2.. Fuel cladding integrity shall be maintained by ensuring that the minimum DNBR remains above the 95/95 DNBR limit
_,for the correlation used.
The: plant-response tN an accidenta1 RCS deprensurization event is dictated by the RCS pressure, RCS temperature, e
moderator density and Doppler reactivity feedback, end
_ pressurizer safety valve capacity.
These parameters are not affected by thermal design flow.
For a reduced flow,,
however, a reduction in the calculated margin to DJB may result.
An accidental RCS depressurization, by definition, is characterized by a depressurization of the RCS.
The RCS design pressure is not challenged.
The steam generator secondary syrten is unaffected prior to retetor trip.
After a reactor trip signal is generated, the turbine and reactor are tripped nearly simultaneously.
In comparison, the LOEL is initiated by an abrupt cessation of steam flow prior to reactor trip, maximizing the heat load for' steaming and the-secondary pressure.
.The LOEL, therefore, bounds this event in system pressure response.
The LOEL event was analyzed in Reference 1, Section 6,2-7, in support of FCF fuel loading at Sequoyah.
The analysis' utilized a reduced RCS thermal. design.
System pressures do not exceed the design limits subsequent to an LOEL.
The accidental-RCS_depressurization event is evaluated for DNB by a statepoint analysis in Reference 3.
Selection of the bounding statepoint is straight-forward.
The DNBR will decrease throughout the period preceding reactor trip.
It will continue to decrease for a brief period following reactor trip as a result of core thermal lag.
The
-a
statepoint for DNBR is chosen by assuming that the OTDT trip
.is bypassed and the reactor trips on low RCs pressure.
DNBR is conservatively calculated for a core exit pressure of 1760 psig.
Reactor power and coolant temperature were
-ansumed to be at nominal conditions, and the reduced thermal design flow was used in tne DNB calculation.
The evaluation in Reference 3 indicates an adequate margin to DNB for the accidental depressurization transient.
Conclusion The system pressure response of the accidental RCS depressurization event is bounded by an LOEL.
A DNBR evaluation of the accidental depressurization event was conducted in Reference 3 and demonstrated acceptable margin to DNB.
It is, therefore, assured that all of the acceptance criteria for the excessive load increase event are successfully met for operation of Sequoyah with a thermal design flow of 340,000 gpm.
.e 4
.e
4.0.2.13 Accidental Depressurization of the Main Steam i
System The spurious opening or failure of a stesm ganerator relief, safety, or steam dump valve represents the most severe overcooling moderate frequency event.
Increacad steam flow resulting from the opened valve results in a depresFurization of the secondary system with an attendant reduction in RCS temperature and pressure.
FCS cooling could produce a positive reactivity insertion vit, moderator feedback and a power increase that could challenge the feel thermal limits.
The accidental main steam system depressurization ovent is a Condition II event and has the following acceptance criteriai 1.
peak primary and secondary system pressure shall not exceed 110% of design value, j
2.
Fuel cladding integrity shall be mainttined by ensuring J
that the minimum DNBR remains above the 95/95 DNBR limit for the correlation used.
I An accidental depressurization of the main steam syztem causes an overcooling and depressurization of both the RCS and steam generator secondary systems.
The respective a
system design pressures are not challenged.
The LOEL, therefore, bounds this event in system pressure response.
The LOEL event wad sne.lyzed in Reference 1, Section 6.2.7, in support of FCF fuel loading at Sequoyah.
The analysis utilized a reduced RCS thermal design.
System pressures do not exceed the design limits subsequent to an LOEL.
1 The accidental depressurization of the main steam system j
event is analyzed in Section 15.2.13 of the.Sequoyah UFSAR, Reference 2.
The results of the analysis shows that the i
relevant acceptance criteria are met for this event.
Under identical initial conditions, however, the return to power and minimum DNBR are bounded by the main steam line rupture.
The main steam line break has been reanalyzed 1.1 support of l
FCF fuel load at Sequoyah (Section 6.4.1 o.f Reference 1).
The reanalysis considers a reduction in' thermal design flow.
The reanalysis demonstrates an acceptable margin to DNB.
Conclusions The system pressure response cf the accidental depressurization of the main steam system event is bounded by an LOEL.
The reanalysis of Reference 1 demonstrates that the system pressure acceptance criterion is met for an LOEL.
The main steam line rupture bounds this event, in DNB.
t
- - -, ~
,,,y,----
-y--.-
-r,
..y,..
- - - -. - _ - _ - -. -. _.. ~.. ~. -.. - - -
i l
s r
1 i
Although the main steam line rupture is a condition IV event, it has been successfully analyzed to meet the Condition II acceptance criteria. It is, therefore, assured i
that all of the acceptance criteria for the accidental j
depressurization of the main steam system event are j
successfully met for operation of Sequoyah with a thermal
{
design flow of 348,000 gpm.
I i
I
(
I
.I 1
t i
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20 h
4 i
e 9
.w s
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n,- e r, w.
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..w.,:,.,,n,s-,env~,.~,---,--
,v-,-~,-,an.
mm,~w-a.~,
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.--va.,,,
4.0.2.14 Spurious Operation of the Safety Injection System
_at Power An error by the operator or a false actuation signal could produce spurious operation of the emergency core cooling system during full power operation. The actuation of the safety injection system will result in delivery of highly borated water to the reactor coolant system.
A reactor trip l
on the actuation of a spurious safety injection (SI) actuation can not be guaranteed.
A spurious SI, therefore, is capable of causing a negative reactivity excursion and power reduction.
The spurious operation of the SI system at power event is a moderate frequency event and has the following acceptance criteria:
1.
Peak primary and secondary system pressure shall not exceed 110% of design value.
- 2., Fuel cladding integrity shall be maintsined by ensuring that the minimum DNBR remains above'the 95/95 Dh5A limit for the correlation used.
'The' analysis of record for this event is contained in Section 15.2.14 of the UFSAR, Reference 2.
Boration of the RCS, resulting from a spurious SI, cause's a reduction in core power.
A reduction in RCS temperature and pressure occurs as a result of the core heat production and secondary heat removal mismatch.
Secondary steam flow is reduced and the secondary pressure drops relative to a reduction in steaming rate.
The pressure limits associate with both the primary and secondary systems are not, thereforo, challenged by this event.
This characteristic remains unchanged by a reduction.in_ thermal design flow.
The UFSAR demonstrates that the margin to DNB is never reduced below the initial DNB margin for the duration of the SI event.
The DNB acceptance criterion is, by definition, met for this etant.
This characteristic remains unchanged by a reduction in thermal design flow, conclusion The-spurious SI event presents neither a challenge to the system pressure limits or the DNB limit.'
This conclusion is independent of. thermal design flow. I t. i's, therefore, assured that all of the -acceptance cri'teria for a spurious are successfully met for operation of Sequoyah with a therm ^al design flow of 349,000 gpm.
-_....__m._.____m_-_____
4
- 1 4.0.3 Condition III Events i
Condition III events are generally more limiting than j
-Condition II events.
These events are expected to occur i
infrequently.
Condition III events are allowed to have some fuel failures so long as the site dose releases are within the 10CFR100 limits.
In addition These events may not cause j
failures that could lead to worse, Condition IV, event.
k i
i J
i
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1
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i i
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-.2.-.--
4 i
4.043.1 Loss of Reactor Coolant From Small Ruptured Pipes or From Cracks in Large Pipes Which Actuates Emergency Core Cooling System-A small break LOCA, by definition, is a break of a small j
pipe or a crack in a larger pipe that can trigger the emergency safety feature actuation system.
The break flow for this event cannot be accommodated by the normal makeup system.- A small break causes a reduction in liquid inventory.and a depressurization of the RCS.
Pumped-ECCS injection, and the possible injection.of the passive cold leg-accumulators, are required to mitigate the event and prevent prolonged core uncovery.
Small break LOCA is a Condition III event and the acceptance criteria are dictated by General Design criterion 35:
1.
Fuel and clad damage that could interfere with continued ef fective core cooling is prevented.
- 2..- Clad _ metal-water reaction is limitedeto negligible-amounts.
A spectrum of small breaks were analyzed in support of FCF fuel reload at Sequoyah (Section 5.9 of Reference 1).
The
_ analysis included-the effects of the reduced thermal design-flow,-and demonstrated that all of the acceptance criteria
- for the small-break transient are met.
cenclusion The small break LOCA event was reanalyzed in Reference 1.
-The reanalysis _ demonstrates that_all of the relevant acceptance criteria are met for this event with a thermal design _ flow of 348,000 gpm.
9
- * =*
u.
ta >
4,0.3.2 Minor Secondary System Pipe Breaks Minor secondary system pipe breaks are of size less than 6 inch diameter pipe breaks.
These breaks must be accommodated with a limited failure of fuel elements.
The effects of " major" secondary system pipe breaks are bounding relative to all of the relevant safety margins.
Conclusion 1
Minor secondary system pipe = breaks are not analyzed for Sequoyah since the response of the plant to those events is j
bounded by the analysis of major secondary system pipe j
breaks.
Although the major secondary breaks are Condition IV events, they are analyzed to Condition II acceptance criteria.
Assurance that all.of the acceptance criteria for 4
-this event are successfully met for operation of Sequoyah with a thermal design flow of 348,000 gpm is based on acceptable results of the major secondary break events, r.
4 i
f
?
s t
'Sl
4.0.3.3 Inadvertent Loading of a Fuel Assembly into an Improper Position The arrangement of assemblies with different fuel enrichments in the core will determine the power distribution of the core during normal operation.
The loading of fuel assemblies into improper core positions or the incorrect preparation of the fue.l. assembly enrichment could alter the power distribution of the core, leading to potentially increased power peaking and possible violation of fuel thermal limits.
The following fuel misloadings have been considered in the UFSAR:
Misloading a fuel pellet or pellets with an incorrect enrichment in a fuel rod.
Misloading a fuel rod with an incorrect enrichment in a fuel assembly.
MLaloading a fuel assembly with an incorrect enrichment or burnable poison _ rods into the~ core.
The UFSAR contains an evaluation of the inadvertent loading of a fuel assembly into an improper position event in Section 15.3.3 of Reference 2.
The evaluation concludes that:
Fuel assembly enrichment errors would be prevented by administrative procedures implemented in fabrication.
In the event that a single pin or pellet has a higher enrichment than the nominal value, the ccnsequences in terms of reduced DNBR and increased fuel and cladding temperatures will be limited to the incorrectly. loaded pin or pins.
Fuel' assembly loading errors are prevented by administrative procedures implemented during core loading.
In the unlikely event that a loading error occurs, resulting power distribution effects will either be readily detected by incore moveable detector system or will cause a sufficiently small perturbation as to be acceptable within the uncertainties allowed between nominal and design power shapes.
These conclusions are unaf fected by a ;re' duction in thermal design flow.
Conclusion Conclusions regarding the acceptable outcome of a fuel assembly misloading event are unaffected by a reduction in
4 t
i i
I thermal design flow.
The results of th. UFSAR analyses are applit.>. le irrespective of RCS flow.
It is, therefore, assured that the acceptance criteria for this event are i
3 successfully met for operation of Sequoyah with a thermal design flow of 348,000 gpm.
t t
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. 0.3. 4
. Complete Loss of Forced Reactor Coolant-Flow A complete loss of forced reactor coofant flow accident can-be caused by a coincident loss of electrical power to a?
of
-the reactor coolant pumps._
If the reactor is at power a.
the time of the acpident, the-immediate effect of the loss of coolant flow is a rapid increase in the coolant temperature.
This increase could result in-DNB with.
subsequent fuel damage if the reactor is not promptly l
tripped.
The.necessary protection against a partial loss of I
coolant flow accident is provided by the~ low primary coolant flow reactor trip signal which is actuated by redundant low flow signals.
d The complete loss of forced reactor coolant flow is a Condition III event.
The event is, however,. analyzed to Condition II acceptance criteria:
1.
Peak primary and secondary system pressure shall not exceed 110% of design value.
I 2.
Fuel cladding. integrity shall be maintained by ensuring l
that the minimum DNBR remains above the 95/95 DNBR limit for the correlation used.
l The complete loss of coolant flow event was analyzed in-Reference 1,-Section 6.3.',
in support of FCF fuel loading l
at Sequoyah.
The analysis utilized a reduced RCS thermal design flow and demonstrated that all Condition II acceptance critaria are met for this event.
Conclusion r
i.
The. total loss of "ed
..t flow was analyzed in Reference 1 with reduced thermal eesign flow.
The analysis demonstrated chat all of the acceptance criteria for this event are successfullf met for operation af Sequoyah.with a thermal design flow of 348,000 gpm.
d i
4 4
t -
t i
j' a:
1 F
6
~
~
- -. - -. ~.-
Q 4,0.3.5 Waste Gas Decay Tank Rupture The waste gas decay tanks contain the' gases ven cd troa the RCS, the volume control tank, and the liquid holdup tanes.
Sufficient volume.is provided in each of four tanks to score i
the gases evolved during a reactor shutdown.
The waste gas accident is defined as an unexpected and_ uncontrolled release to the atmosphere of the radioactive xenon and 4
krypton gases that are stored in the waste gas storage system.
The UFSAR analysis assumes the rupture of a single
. waste gas decay tank.
i The-waste gas-decay tank rupture accident is a Condition IV event.
The acceptance criteria are:
1,-The offsite dose shall not exceed 10CFR100 limits, i -
- 2. The dose to control room personnel shall not exceed 5 rem.
~
The existing analysis for this event are contained in the j
UFSAR, Section 15,5,2'of-Reference 2.
The analysis concludes that all doses resulting from a waste gas decay tank rupture are well within the limits and that the acceptance criteria are met, The parameters important to
-t; dr ie calculations for a waste gas decay tank rupture are
- he,;1 activity concentration and site-specific dispersion i ac :>r s.
Neither of these parameters are affected by a ch.:u a in thermal' design flow, The Reference 2 analysis results, therefore, remain applicable.
C2nclusion' The results of the-UFSAR analysis for the waste gas decay tank rupture-are unaffected by RCS flow, It is, therefore, assured that the acceptance criteria for this event are successfully met for operation of Sequoyah with a thermal design flow of 348,000 gpm.
i 0
a w
c==
c
4.0.3.6 Single-Rod Cluster Assembly Withdrawal at Full Power p
The althdrawal of a single RCCA from its inserted bank results in both a reactivity increase and increased power peaking in the region of the core surrounding the withdrawn
~
i RCCA.
The reactivity increase causes the neutron flux to i
increase and produces a localized increase in peaking.
Subsequently, thermal power, coolant and fuel temperature, and system pressure increase.
Reactor trip on overtemperature AT provides protection for this event.
The peaking asymmetry associated with the withdrawn RCCA can,
[
however, can lead to localized fuel failures.
An analysis of the single RCCA withdrawal event is reported in Section 15.3.6 of the Sequoyah UFSAR (Reference 2).
l There are two parts of the analysis, the system analysis and the peaking analysis.
The system analysis for the bank withdrawal accident evaluates very low to very high l
reactivity insertion rates.
-The reactivity insertion rate of'a' single rod witudrawal accident is wifhin the range of i
reactivity rates analyzed for the bank withdrawal analysis.
Therefore, the core response from a single rod withdrawal
_ accident is already analyzed by the bank withdrawal analysis.
The results for the bank withdrawal accident demonstrate that no DNB occurs.
l Since the peaking for the single rod withdrawal accident can be higher than the peaking for the bank withdrawal accident, the second part of the evaluation is performed.
It is
[
conservatively assumed that the minimum DNBR of the bank withdrawal accident is at the limit even though the analysis demonstrates that it only approaches the limit for certain i
4 reactivity insertion rates.
Thus, any peaking increase caused by_-the single rod withdrawal event above that limit established by a-bank withdrawal event is assumed to fail fuel.
The number of pins from a single rod wichdrawal that exceeds the. limit must be less than 5 percent of the total fuel rods in the core to assure that the results of the CFSAR analysis j
are bounding and remain applicable.
The estimated pin failures for a single rod withdrawal event are evaluated on a cycle,by cycle basis.
Conclusion A reduction in thermal design flow coul'd' affect the calculated minimum DNBR associated with a single RCCA withdrawal event.
The reduction in thermal design flow has been evaluated for the rod bank withdrawal event and found to be acceptable.-The first part gf the single rod withdrawal analysis is the same as the bank rod withdrawal
event and-is acceptable.
The second part of the single rod withdrawal analysis is a peaking analy. sis and is not
- dependent upon the design flow.
The conclusions associated with the UFSAR analysis of this event are confirmed.
It is,
- therefore, assured that the acceptance--criteria for this event-are successfully met for operation of Sequoyah with a thermal design flow of 348,000 gpm.
e 1
9 f-4 s
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4 s&*-
t 4
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4.0.3.6 Steam Line Break Coincident with Rnd Withdrawal at-Power The steam line break coincident with rod withdrawal at power-is an event that is classified as condition III event for Sequoyah.
The event occurs at_ full power conditions, initiated by a steam line break which causes rod withdrawal to occur.
The high neutron flux and the overtemperature AT reactor trips are disabled in this event.
The overpower AT and low steam line pressure trips provide the necessary protection.of the reactor.
.The acceptance criteria for this event are 1.
The thermal margin limits -(DNBR) as specified in the Standard Review Plan shall be met.
2.
Fuel centerline temperatures as specified in the Standard Review Plan shall not be exceeded.
3.
" Reactor pressure boundaries shall not exceed the ASME Code design limi~ts on pressures.
The steam line break coincident with rod withdrawal at power event was analyzed in support of the FCF fuel loading (Section 6.4.6 of Reference 1).
The analysis took into consideration reduced thermal design flow.
It was demonstrated that all acceptance criteria were met for this event.
Conclusion The steam line break coincident with red withdrawal at power i
eveut-was reanalyzed in Reference 1.
The reanalysis demonstrated that all of the relevant acceptance criteria for this event are met for this event with a thermal design flow of 348,000 gpm e
a e
- 1
)
- 4.0.4
' Condition IV Events.
l i
Condition'IV events are more limiting than Condition III events.
Condition IV events are_not expected to take place but are postulated because their consequences would include-the potential for the release of significant amounts of radioactive material.
These represent the most limiting design cases.
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F 4.0.4.1.
Major Reactor Coolant System Pipe Ruptures (Loss of Coolant Accident)
The Sequoyah nuclear power plant is designed to withstand thermal effects caused by a loss of coolant accidenL of size up to and' including the double-ended guillotine break of the largest RCS pipe.
The reactor core and internals,- together with the emergency core cooling sy, tem (ECCS) are designed so that the reactor can be shutdow. safely and-the core geometry retained to permit long term cooling.
The ECCS is designed so that, even with the assumption of the most severe single active failure, the acceptance criteria for this event are met.
The large break LOCA event is a Condition IV event, subject to the following 10CFR50.46 acceptance criteria:
- 1. The calculated' peak cladding temperature must be less than 2200 degrees F.
2..The maximum local cladding oxidation must be less than 17
- percent,
- 3. The maximum amount of core-wide oxidation does not exceed 1 percent of the fuel cladding.
- 4. The cladding remains amenable to cooling.
5.
Long-term cooling is established after LOCA recovery.
Of the many parameters affecting the plant response to a large break, the initial core fluid and fuel temperature distribution are most affected-by a reduction-in thermal design flow.
The large break LOCA was analyzed in support of FCF fuel loading in sections 5.1 through 5.8 of Reference 1.
The analysis accounted for the reduced thermal design flow and demonstrated that all of th'e acceptance criteria are met.
Conclusion The large break LOCA event was analyzed in Reference 1.
The analysis demonstrated that all of the relevant acceptance criteria are met.
It is therefore assured that all of the acceptance criteria for this event are successfully met for operation of Sequoyah with a thermal design flow of 348,000 gpm.
4.0.4.2 Major Secondary System Pipe-Rupture A major secondary system pipe rupture is generally defined i
as a guillotine break of the main steam line.
A steam line break results in the blowdown of the affected steam generator and severe ove.rcooling of the primary system.
The event is initiated from hot zero power, the worst operational mode for overcooling.
With a negative moderator temperature reactivity coefficient, the primary system cooldown results in a reduction in core shutdown margin sad possible return to power.
If the most reactive control rod is assumed to be stuck in its fully withdrawn position, the core can become critical and return to power.
The return to power, with the large local flux peak in the region of the stuck control rod, could result in fuel pins experiencing DNB.
Major secondary system piping failures are classified as design basis, Condition IV events.
This means that some i
fuel pins can experience DNB as long as the offsite doses remain within 10CFR100 limits.
However, the major secondary system rupture event is analyzed to' Condition II acceptance criteria:
1.
Peak primary and secondary system pressures shall not exceed 110% of the design pressures.
2.
Fuel clad integrity shall be maintained by ensuring that the minimum departure from nucleate boiling ratio (DNBR) remains above the 95/95 DNBR limit for the correlation used, i
The main steam line break event results in a cooling and depressurization of both the secondary system and the primary RCS.
The event does not, therefore, pose a threat to system p'ressure limits.
The LOEL event remains bounding in system pressure.
Of the parameters that most affect the plant response to main steam line break, the primary to secondary heat balance is most affected by a reduction in thermal design flow.
Since the heat transfer coefficient is inversely related to thermal design flow, overcooling would effectively be reduced, as would the return to power.
On the other hand, the calculated margin to.DNB could be adversely affected by the reduction in flow.
The main steam line break was analyzed'in support of FCF fuel loading at Sequoyah (Section 6.4.1 of Reference 1)
The system analysis was conducted with an RCS flow that was higher than the thermal design flow - the flow associated with a clean, unplugged, steam generator.
In this way the heat transfer rate, overcooling, ind return to power were maximized.
The hot channel thermal-hydraulic analysis was
then conducted utilizing the reduced thermal design flow to conservatively calculate margin to DNB, The analysis demonstrated that all of the acceptance criteria =for this event are met.
Conclusion The major' secondary system pipe break event was reanalyzed in Reference 1.
The reanalysis demonstrated that all of the
-relevant acceptance criteria are met.
.It is, therefore, assured that all of the acceptance criteria for-this event are successfully met for operation of sequoyah with a thermal design flow of 348,000 gpm, e
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4.0.4.3 Major Rupture of a-Main Feedwater pipe A major rupture of a main feedwater line represents.a raoid
' decrease in heat removal capability of the secondary sys' ten because it reduces the supply of feedwater to the steam generators.
Main feedwater is, in fact, assumed to be lost to-all of the steam generators at the time of rupture.
' Reverse blowdown of the affected steam generator results in a relatively rapid reactor trip signal on low-low level in
_that generator.
Following reactor trip, the main feedwater line break is characterized by excess heat removal from and cooldown system as the affected steam generator continues to blow down through the break-When the low steam line pressure setpoint-is reached the main steam isolation valves are closed and safety injection to the RCS is initiated.
The loss of steam generator inventory and rising pressure steam pressure cause primary temperatures to rise.
Successful termination of the transient ~is achieved when the auxiliary feedwater supplied to.the steam generators is sufficient to remove core decay heat,
~
The main feedwater line break is a Condition IV event,
-subject to the following acceptance criteria:
1.
Peak primary and secondary system pressures shall not exceed 110% of the design pressures.
2.
Fuel clad integrity shall be maintained by ensuring that the minimum DN3R remains above the 95/95 DNBR limi: for the correlation used.
- cover the 3,
Liquid in the RCS shall be sufficient reactor core at all times.
The main feedwater line break event resulcs in a depressurization of the steam generator. secondary.
Isolation of the steam-lines occurs on low steam pressure and is delayed following reactor trip.
In comparison, the LoEL is caused by an. abrupt loss in steam flow before the i
reactor trip-The secondary heat load is, therefore,
. comparatively higher for the.LoEL event.
The secondary pressurization is, therefore, bounded by che LoEL transient This relationshio is unaffectWd by a reduction in response.
thermal _ design flow.
The crimary pressure response-during a LoEL bounds that for p
the main feedwater line break becauce the reactor trip is delayed with respect to the LoEL.
In contrast, reactor trip and turbine trip occur at the same time on low-law steam generator level in the main feedwater line break, resulting i
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in a lower _ peak pressure as compared with the LOEL event.
This_ relationship is unaffected by RCS flow rate.
4 LOEL was analyzed in support of FCF fuel reload at Sequoyah (Section 6.2,7 of Reference 1).
The analysis assumed a reduced thermal design flow, and acceptable margin to the j
primary pressure acceptance criteria was demonstrated.
~
Ultimately, the feedwater line break event is mitigated via reactor trip and feedwater isolation on a low-low steam 1
generator level.
This result is independent of thermal design flow.
MAP limits were developed at operational i
statepoints associated with overtemperature and overpower AT trip boundaries and verify that adequate DNB margins exist at these statepoints.
Since the transient does not l
trip on overtemperature or overpower AT, there is no challenge to the fuel thermal limits.
No DNB occurs and the feedwater line break is, therefore, bounded by the rod withdrawal at power event which relies on the overtemperature AT for protection.
The uncontrolled rod withdrawal at power event was analyzed i
in Reference 1 in support of Framatome Cogema Fuels fuel loading at Sequoyah (Section 6,2.2).
The analysis utilized a reduced RCS thermal design flow and demonstrated that adequate DNB margin is retained.
The feedwater line break is also analyzed in the UFSAR to demonstrate that long-term cooling is maintained.
Long-term decay heat removal capability is related to secondary side liquid inventories and auxiliary feedwater flow.
These parameters are unaffected by RCS flow rate.
Conclusion The feedwater line break event is bounded, in pressure, by the LOEL event.
LOEL was analyzed with reduced thermal design flow and system pressures were demonstrated to remain within 110% of design values.
There is no challenge to DNB for this event as the reactor trips prior to the approach to the safety limits.
Furthermore, with respect to DNS, the feedwater line break is bounded.by the rod withdrawal at power event.
The red withdrawal at power event was analyzed with a reduced thermal design flow and margin to DNB was demonstrated.
Long-term decay heat removal is related to steam generator secondary liquid inventory and auxiliary feedwater flow capacity.
These are n',t.affected by primary o
system flow, and the existing UFSAR analysis remains bounding.
Therefore, it is assured that all of the acceptance criteria for this event are successfully met for operation of Sequoyah with a thermal design flow of 348,000 gpm.
4.0.4.4 Steam Generator Tube Rupture The steam generator tube rupture is a' design basis accident that considers postulated failure of a single steam
-nerator tube.
After the rupture, the RCS depressurizes s a mass transfer from the primary system to the steam generator secondary.
The reactor is tripped, main feedwater flow is isolated, and the SI system is actuated on the low pressurizer pressure reactor protection signal.
The primary system event is effectively terminated when makeup flow vi the ECCS matches the rate of coolant loss through the failed steam' generator tube.
Tube leakage is terminated when the operator depressurizes the prima'y system below the steam r
pressure of the affected steam generator.
The acceptance criteria for the steam generator tube rupture (SGTR) event are:
1.
The calculated doses do not exceed 10CFR100 guidelines.
The-RCS activity at the initiation of this_ event is a prescribed value, una'ffected by a reduction in RCS flow.
The offsite doses for a steam generator tube rupture event are dependent on the primary to secondary break flow.
Break flow is dependent on RCS pressure and temperature and the tube cross-sectional area.
It is also dependant on the actions of the reactor operator in'the mitigation of the event.
Since none of these parameters are affected by a reduction in thermal design flow, the UFSAR analysis of this event, sections 15.4.3 and 15.5.5, are unaffected by the reduction ~.
Conclusion 1
The parameters governing the plant response and offsite dose b
. release associated with a steam generator tu e rupture ara independent of RCS flo.w.
The existing UFSAR analyses for this event demonstrat'e acceptable results for this transient and are equally applicable for operation at a reduced RCS
^
flow.
It is, therefore,-assured that all of the acceptance criteria for this event are successfully met for operation of Sequoyah with a thermal design flow of 348,000 gpm.
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4.0,4.5 Single Reactor Coolant Pump Locked Rotor The locked rotor accident is analyzed considering a
- postulated seizure of an RCP rotor.
The locked rotor is more limiting than a reactor coolant pump shaft break as it presents a greater resistance to RCS fle'...
RCS Flow is rapidly reduced as a result of a locked rotor.and the reactor trips on low flow.
The rapid flow reduction results in a decrease in the DNBR.
Following reactor trip and subsequent rod insertion, the DNBR increases.
This event is classified as a condition IV event with the following acceptance criteria:
1.
Pcak RCS pressure shall not exceed 110% of the design value.
2.
Peak clad temperature shall not exceed 1800 F.
4.
Maximu.n calculated dose. is less than 10CFR100
~ guidelines.
The locked rotor event was analyzed in Reference 1, Section C 4.4, in support of FCF fuel loading at Sequoyah.
The reanalysis utilized a reduced RCS thermal design flow and
-concluded that all of the relevant acceptance criteria are met.for this event.
Conclusion The complete loss of coolant ficw event was reanalyzed in Reference 1 with reduced thermal design flow.
The reanalysis demonstrated that all of the relevant acceptance criteria are met for this. event with a thermal design flow of 348,000 gpm.
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4.0.4.6 Fuel Handling Accident This_ event takes place in the spent fu'el pit _ floor.
A spent fuel assembly is dropped on the pit floor and results in the rupture of the cladding of all fuelsrods.
This event uis not af fected by a change in thermal design flow since it does not take place in the nuclear steam supply system, The fuel handling analysis in section 15.4.5 of the UFSAR, Reference 2, therefore, remains applicable for Sequoyah.
gpnclusion Thermal design-flow changes have no effect on the consequences of this event.
The UFSAR analysis for this event demonstrates acceptable results and is equally
-applicable at reduced thermal design flows.
It is, therefore, assured that all of the acceptance criteria for-this event are successfully met _for operation of Sequoyah with a thermal design. flow of 348,000 gpmr e
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4.0.4.7 Rupture of a Control Rod Drive Mechanism Housing (Rod Cluster Control Assembly Ejection)
This' accident is defined as the mechanical failure of a i
control rod mechanism pressure housing resulting in the i
ejection of a-RCOA and drive shaft.
The consequence of this mechanical failure is a rapid positive reactivity insertion together with an adverse core power distribution, possibly leading to localized fuel rod damage.
4 The ejection of an RCCA is classified as a limiting fault, Condition IV and is subject to the following acceptance criteria:
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Average fuel pellet enthalpy at the hot spot below 225 cal /gm for unirradiated fuel and 200 cal /gm for irradiated fuel.
2.
Fuel melting will be limited to-less than 10% of the i-fuel volume at the hot spot even if the average fuel
. -pellet enthalpy is below the-limits-of criterion 1 above.
3.
Peak reactor coolant pressure less thaa that which would cause stresses to exceed the faulted condition
,[
-stress limits, The Sequoyah UFSAR analysi,s of ejected rod, section 15.4.6 of Reference 2, contains sufficient information to verify conformance of a given fuel cycle is within the applicable j
- bounds, Section 6.4.5 of Reference 1 describes how a j
. comparison is performed ea h cycle to ensure the applicability of the UFSAR analysis, utill:ing codes and L
methods developed at FCF.
This means that the ejected rod event is evaluated as a part i.
of the fuel design process.
This evaluation takes into account the reduced thermal design flow.
Ehsuring that parameters such as ejected rod worths, power peaking, delayed neutron fraction and fuel: pin failure census for each cycle are bounded by those in the UFSAR analysis ensures that the-UFSAR ejected rod analysic, which
' demonstrates that the relevant acceptance criteria are met, is applicable to that cycle.
Conclusions t
'd'-for in the Reduced thermal design flow is accounte evaluation of ejected rod that is performed for each cycle design.
This evaluation is performed v.ia a comparison of plant parameters to those used in the UFSAR analysis and an j
independent pin failure census.
Conservative assurance is
_provided that the UFSAR' transient'" remains bounding and that-the acceptance criteria are met.
Therefore, it is assured r
r
/
- t. hat all of the acceptance criteria for this event are met for operation of Sequoyah with a thermal design flow of 348,000 gpm ;
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4 4.1 Environmental-Consequences 4.1.1 Environmenta2 Consequences of a Postulated Loss of AC Power to Plant Auxiliaries The analysis of the environmental consequences of a postulated loss of AC power to the plant auxiliaries is presented in the UFSAR,'section 15.5.1.
The analysis assumes a prescribed primary-to-secondary leakage, a
prescribed primary _ coolant activity, a prescribed iodine partition. factor, and a prescribed steam generator blowdown rate.
In that the analysis does not allow variability in the plant response to the transient, but chooses conservative values to bound them, there is no effect of the thermal design flow on the UFSAR consequence evaluation for this event.
Conclusion Reducing the thermal design flow *o 348,000 gpm at Sequoyah does not adversely affect the consequences of this event.
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~4.1 2 Environmental-Consequences of a Postulated Waste Gas Decay Tank Rupture An evaluation of this event is contained in section 4.0.3.5 of this event.
~ Conclusion.-
The.results of the UFSAR analysis demonstrating an adequate margin to safety for the waste gas decay tank rupture are unaffected by RCS flow. Reducing the thermal design ficw to 348,000 gpm at Sequoyah-does not adversely affect the consequences of this event.
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- Coolant Accident-The analysis of the environmental consequences of a postulated. loss of coolant accident is presented in the UFSAR, section 15.5.3.
The analysis-assumes a prescribed i -
mass release.-
In that the analysis does not allow
-variability in the plant response to the transient, but-
. chooses conservative values to bound them, there is no effect of the thermal design flow on the UFSAR consequence 4
i evaluation for this avent.
Conclusion Reducing the thermal design flow to 348,000 gpm at Sequoyah p
goes not adversely affect the consequences of this event.
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.Environmentall Consequences of a Postulated Steam Line Break The analysis of the environmental consequences of a postulated steam line break is presented in the UFSAR, section is.5.4.-The analysis assumes a prescribed. primary-to-secondary-leakage, a prescribed primary coolant activity, a; prescribed iodine partition factor, and a. prescribed steam-generator blowdown rate.
In that the a'nalysis does not allow variability in the plant response to the transient,.
but chooses conservative values to bound them, there is no effect of the thermal design flow on-the UFSAR consequence evaluation for this event, conclusion Reducing the thermal design flow to 348,000 gpm at Sequoyah does not adversely affect the consequences of this event, i
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4.1.5 Environmental Consequences of a Steam Generator Tube Rupture An-e"al'uation of this event is contained in section 4.0.4.4 of this document.
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-Conc us on The results of the _UFSAR analysis demonstrating an adequate margin to-safety for_the cteam generator tube rupture are unaffected by RCS flow. Reducing the thermal design flow to 348,000 gpm at Sequoyah does not adversely affect-the consequences of this event.
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. 4~.1.6 Environmental Consequences of a Postulated Fuel
- Handling Accident An evaluation of'this event is contained in section 4.0.4.6 of this_ document.
Conclusion The results of the UFSAR analysis demonstrating an adequate margin.to safety for the. fuel handling event are unaffected by RCS flow. Reducing.the thermal design flow to 348,000'gpm at Sequoyah does not adversely affect the consequences of this event.
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. Ejection Accidet.t--
- !An evaluation of this event is-contained in section 4.0.4,7 of this event.
].
Conclusion The results of the UFSAR analysis demonstri. ting an adequate margin to safety.fr-'he rod. ejection event are unaffected
-by RCS flow. Reduc, "g the thermal design fl w to 348,000 gpm-at Sequoyah does not adversely a'f fect the ccnsequences of this event.
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4.2 Containment Analysis and Equipment Qualification The UFSAR identifies two events that are analyzed to
-determine peak containment building pressure: the large break loss of coolant accident and the main steam line
- break, Critical parametcrs for the LOCA event are the initial primary mass and-internal energy in the primary system.
The important parameters for the steam line breax are secondary mass, secondary internal energy and primary-to-secondary heat transfer.
Mass and energy release associated with LOCA rely on the break size, the initial RCS pressure and energy content of j
the fuel, the coolant and the internal structures.
With a reduction in thermal design flow, the' primary pressure, coolant and fuel average temperature, and structural s
i temperatures are unchanged.
The containment pressures and temperatures subsequent to a large break LOCA'will therefore be equivalent with a reduced thermal design flow.
l The.' mass and energy release associated with a steam line break are a function of the internal energy and primary-to-secondary heat transfer.
With a reduction in thermal design j
flow in conjunction with an equivalent RCS average coolant e
temperature, the secondary side would of necessity initially be at a slightly reduced pressure and temperature to i
maintain a heat balance.
This is advantageous in the=early i
phases of a steam generator blowdown.
In the long term, the reduced thermal design flow effectively retards the primary-i to-seconcary heat transfer rate via a reduction in heat transfer coefficient.
The energy release would, therefore, be lower for the steam line break, although this beneficial effect would be extremely small.
The containment pressures and temperatures subsequent to a main stem line break will, therefore, be less with a reduced thermal design flow.
i-For equipment qualification outside of containment, a steam line break is postulated.
The same line oftreasoning
-applies here.
A reduced thermal design flow would result in an initially lower energy release and, in the long term, would retard the primary-to-secondary heat transfer rate.
The vapor. temperatures in compartments cutside of containment following a steam line break would be less with j '
a reduced thermal design flow.
-Conclusion f
Parameters affecting the large break L'oCA mass and energy release-rate considered in the UFSAR chapter 6 containment
. analyses are unchange'd by a reduction in thermal design flow.
Lower compartment pressures and temperatures would be predicted both inside and outside of containment as a result of steam'line break, considering'a reduction in thermal 3
design flow. Reducing the thermal design flow to 348,000 gpm i
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E at;sequoyah does not adversely affect the consequences of' this event.
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5.0 General Conclusions
- The accident.s -in the Sequoyah -FSAR were evaluated to determine the effects of a reduction in RCS thermal design flow to 348,000 gpm. - The evaluation determined that all NRC-approved acceptance criteria continue to be met with the proposed changes in RCS-flow.
Certain limiting accidents were analyzed for FCF reload cuel (Reference 1).
Those analyses incorporated the revised RCS flow and showed that all acceptance criteria were met.
Those events that were not reanalyzed for the FCF reload fuel were systematically examined to determine how the proposed change affects the transient response and the acceptable result of the licensing basis analysis.
The probability of occurrence of any of che Chapter 15 transients will not be increased due to the changes proposed.
No new transients'will be created as a consequence of the changes proposed.
The probability of l
malfunction of equipment important to safety will not be l-increased in-any way due to the proposed dhanges.
Because all NRC-approved acceptance criteria continue to be met with the proposed changes in thermal derign flow, the margin of safety is not reduced.
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6.0 References
- 1. FTI-Topical Report BAW-10220P, " Mark-BW Fuel ASJembly Application for Sequoyah Nuclear Units 1 and 2."
- 2. Sequoyah Final Safety.Malysis~ Report, Updated through Amendment 12.
3.
FTI Calculation 32-1240666-00, 'Sequoyah Maximum
- Allowable Peaking Limits," JL Griffith i
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l 16. - The're is no justification that 'a complete-loss of flow is ~
more limiting than a partial loss of flow as stated in the submittal.
Please provide the justification.
Response
Both partial and complete loss of flow transients were analyzed in the McGuire/ Catawba application report for Mark-BW reload, sections 4. 3.1 and 4. 3.2 of BAW-10173, respectively.
The results of these analyses clearly demonstrate that because the complete loss of flow exhibits a more severe reduction in core flow as compared with the partial loss of flow event, the DNB response for the complete loss of flow is bounding (see Figure 16-1).
McGuire/ Catawba and SQN are similar in design (all are 4-loop Westinghouse plants) with identical power densities (Mark-BW cores rated at 3411 MWth), similar flows, etc.
It is, therefore, concluded that the comparative results apply equally to SQN - the complete loss of coolant flow event bounds the partial loss of coolant flow event with respect to DNBR.
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17.
For a number of transients (loss of flow), a delayed neutron fraction corresponding to end of life (ECL) is chosen.
Describe why this is conservative.
Response
For events, such as the loss of flow event, that ure posed to conservatively predict a minimum margin to DNB, the power I
response prior to reactor trip is maximized.
One of the meaures taken to maximize power response is to utilire the most limiting reactivity feedback parameters, e.g. moderator and Doppler feedback, to suit the transient even if this requires mixing beginning-of-life and end-of-life coefficients.
Power response, however, is related to both reactivity and delayed neutron fraction.
A min! mum, EOL, delayed neutron fraction maximizes the power response to a given reactivity addition.
This method of modeling reactivity feddback is used frequently in FCF safety analysis and is consistent with the approved safety analysis methodology of BAW-10169.
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18.
Describe how the stuck rod is modeled for_each transient and identify which analyses assume a stuck rod and which ones do not.
Response
All of the events analyzed in the reload applications topical l
assume a stuck rod.
The most reactive rod is assumed to be stuck out of the core and is reflected in both the development of the rod worth for reactor trip or in the characterization of pla.,e shutdown margin.
In addition, the steam line break analysis explicitly models localized peaking and reactivity feedback near the stuck rod in the process of predicting DtTB margin.
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19.
The transient analysis presented does not discuss single failure assumptions in any detail.
For each transient, state the limiting single failure chosen and why it is limiting.
Response
The following six events are were reanalyzed for use of FCF fuel at Sequoyah.
1.
RCCA withdrawal at full power 2.
Loss of electric load 3.
Loss of forced flow 4.
Locked rotor 5.
Main SLB at subcritical condition 6.
Steam line break coincident with rod withdrawal at power Each one of these events is discussed here with respect to the single active failure assumption used.
RCCA Withdrawal at Full Power The only active safety-grade system that mitigates this event is the reactor protection system.
A single failure of the RPS will not stop it from performing its design function ( i.e. insert control rods).
Loss of Electric Load The only active safety-grade system that mitigates this event is the reactor protection system.
A single failure of the RPS will not stop it from performing its design function (i.e. insert control rods).
Although pressurizer safety valves are passive devices and are exempt from the single failure requirement, a single pressurizer safety valve was failed to maximize the primary system pressure for this event.
.It should be noted that the auxiliary feedwat'er system, which is an active safety system, is required for long-term core heat removal. This system is also designed with the defense-in-depth concept, allowing for single failures of various components within the system.
This system is not modeled in our analyses because'it does not mitigate the system pressure or core DNB responses and because FCF fuel does not' affect the ability of the system to meet its design function.
Loss of Worced Flow and Locked Reactor Coolant Pume Rotor The only active safety-grade system that mitigates these events is the reactor protection system.
A single failure of the RPS will not stop it from performing its design function (i.e. insert FCF Non-Proprietary
Although pressurizer safety valves are passive devices and are exempt from the single failure requirement, a single pressurizer safety valve was failed to maximize the primary system pressures for these events.
Main Steam Line Break The active safety-grade systems that mitigate this event are the reactor protection system, the steam and feedwater isolation systems, and the safety injection system.
A single failure of the RPS will not stop it from performing its design function (i.e. insert control rods).
Failure of a single main feedwater isolation valve to close cannot prevent isolation of feedwater to the affected steam generator because an isolation signal closes the feedwater control valves and trips the feedwater pumps.
Failure of the feedwater pumps to trip cannot prevent feedwater isolation because the feedwater isolation valves will close.
Consequently, the only single active failure that can affect the core response to main steam line break is the failure of a single train of safety injection.
This failure minimizes the injection of boron into the reactor coolant system.
This is the single active failure assumed in the FCF analysis of this event.
Steam Line Break Coincident with Rod Withdrawal at PoweI This eve'nt results in an increase in power and a coincident decrease in core DNBR until the reactor trips and control rods are inserted.
The only active safety-related component that mitigates this event is the reactor protection system.
A single failure of the RPS will not stop it from performing its design function (i.e. insert control rods).
FCF Non-Proprietary
20.
For the main steam line break analysis (p. 6-54), the volume of water between the cold leg piping and the first check valve is considered to be at 0
ppm boron l
concentration.
The current Final Safety Analysis Report (FSAR) analysis assumes the volume of water in the piping from the RWST to the cold leg piping (a much bigger volume)
-is all at 0 ppm.
Please describe how you can assure that the water in.the piping between the RWST and the first check valve is at least 1950-ppm.
Response
For the FCF main steam line break analysis it is assumed that the coolant in the four afety injection lines downstream of the second emex
%1"e removed from the cold leg, has a boron concentratar o' ) ppm, not the first set which is incorrectly alluded to on page 6-54.
Modeling an unborated purge volume accounts for possible dilution via the diffusion of RCS coolant back into the lines plus potential leakage of the first set of check valves.
Once the unborated water is purged, boron is assumed to be injected into the RCS at a concentration of 1950 ppm via a single intermediate head pump.
This is a reasonable method of modeling post-accident boren injection in that it both is conservative and reflective of the actual plant configuration.
Even though boron injection is modeled for the analysis of steam line break, the return to power associated with the limiting break is not mitigated by boron.
The limiting break, a complete severance of the steam line upstream of the pipe flow reducer with offsite power available, is reported in Section 6.4.1 of the reload topical report.
The results clearly show that the core power turns over prior to the injection of boron.
Core power is controlled instead by the dry-out of the broken steam generator and the subsequent cessation of both RCS cooling and positive reactivity feedback.
Subsequent to the original FCF analysis, sensitivity studies have been performed related to the assumption of boron delivery to the RCS for the main steam line break event <
The sensitivity studies examine the extent that leakage / diffusion of RCS coolant can occur in the safety injection lines before the results of the limiting case are adversely'affected.
The studies show that coolant with a
' boron concentration equivalent to that'of the End of Life RCS (0 ppm) can exist in the safety injection system piping back as far as three-check valves removed from the cold leg up to the discharge of the safety injection pump (see Figure 20-1).
A boron concentration of 1950 ppm is assumed from the-RWST to the discharge of the SI pumps with 0 ppm from the pump discharge to the RCS.
Based on these conditions,
the limiting break (a complete severance of the steam line upstream of the pipe flow reducer with offsite power available), as reparted in Section 6.4.1 of the reload topical remains valid and limiting.
The sensitivity studies performed prove that the original FCF analysis remains bounding and the return to power associated with the other non-limiting steam line break cases will also remain less than the return to power associated with the limiting case assuming Oppm boron back to the discharge of the SI pump.
Boron injection is conservatively modeled in the FCF analysis of the main steam line break for Sequoyah, even though the assumption does not match the assumption of the original FSAR analysis in that:
a conservative volume of unborated liquid is assumed to be purged prior to the injection of borated water into the RCS I
The boron injection assumption also reflects the physical arrangement of the plant in that:
the SI system is isolated from the unborated charging lines when the charging system is aligned for normal makeup the SI system always draws suction from the RWST during normal operation, which is borated to a minimum of 2500 ppm check valve leakage, which could potentially dilute boron in the SI lines, is detectable via pressurization as a result of leakage back into liquid-solid lines Again, the limiting steam line break presented in the Sequoyah reload topical report does not require the addition of boron for the mitigation of the core power transient for this event.
The SI pump casing is vented every 31 days per Technical Specification 4.5.2.b.1.
This in turn causes the upstream piping back to the RWST to be vented.
Venting occurs until a steady stream of water is noted at the pump casing vent, j
The pressure source is the head difference between the RWST and the SI pump.
The RWST is verified to have at least 2500 I
ppm boron every 7 days per Technical Specification 3.5.5.
The piping downstream of the SI pump is,also vented for the hot and cold leg injection paths.
The procedure requires that the flowpath through the SI pump, the pump discharge check valve and each test header isolation valve (63-21, 23, 121 and 167) be open for ten minutes, for a total of forty minutes of venting through the piping header.
Again, the pressure source is the head difference between the RWST, the SI pump and the discharge piping (see Figure 20-1).
l 1
l l
The Technical Specification minimum RWST volume is 375,000 l
gallons.
The volume of water between the SI pumps supply I
check valve and the RWST tank entrance is approximately 4250 l
gallons.
The approximate volume of water between the SI pump supply check valve and the discharge of one SI pump is 65 gallons.
These are the volumes assumed to be at 19s0 ppm I
of boron.
The approximate volume of water between one SI pump discharge and a typical secondary check valve inside containment is 160 gallons.
The approximate volume of water between a typical secondary and primary check valve is 60 gallons.
The approximate volume between a primary check valve and the hot / cold leg piping is 20 gallons.
These are the volumes assumed to be at 0 ppm.
The venting procedures previously described for the downstream piping of the SI pump require venting each of the four paths for 10 minutes each.
Based on an average flow rate of 10 gallons per minute for each vent path, a minimum of 400 gallons-is displaced in the section of piping-between the RWST and the secondary check valves inside containment, An additional volume is also displaced in this suction r
{
piping due to the venting of each SI pump casing.
This
-displaced water is replaced with RWST inventory with a minimum of 2500 ppm boron every 31 days.
The system is a water solid closed system during operation.
Any leakage past any of the check valves will result in the rapid pressurization of the piping volume upstream of the leaking check valve.
This pressurization will only require a relatively small amount of liquid to equalize the pressure, thereby limiting the potential for the dilution of the contained volume.
This condition is therefore self-
- limiting, Upstream of the SI pump suction but downstream of the common SI pump suction check valve, a code relief valve is installed (rel.ief setpoint of 220 psi),
Downstream of the SI pump discharge, three code relief valves (relief setpoint of 1750 psi) are installed, one on each injection flowpath.
Continued leakage through multiple check valves would be required to create a leak path through any of the above relief valves. The following is an excerpt from the S64uoyah FSAR section 5.2.7.8, ECCS Intersystem leakage:
Leakage from the RC system into lok pressure porcions of several ECCS lines is prevented ky2che use of two check valves in series.
The check valves are tesced for leakage in accordance with che applicable surveillance inscructions.
The probabilicy of a major leak through any-pair of check valves will cherefore be limited to approximacely 5.5 x so.9 per reactor year.
This probabilicy-is low enough co eliminace any concern for a
l--
major intersystem leak into low pressure ECCS Systems.
However, means are available to concinuously monitor and l
alarm intersystem leakage across the incerfaces between the RC System and the following:
Cold Leg Accumulacors (CLA), Chemical and Volume control Syscom (CVCS), Safecy Injection System (SIS), and RHR Syscem.
Leakage into those syscems can be detected both by monitoring for l
signs of incoming leakage and by monitoring che RC System for signs of outgoing leakage.
Intersystem leakage across the two check valves in each of che four SIS cold leg injection lines or across the two check valves and one normally closed gate valve in
.each of the four SIS hot leg injection paths would increase the pressure in those segments of the lines.
A separate pressure sensor is provided in each of the two SIS pump discharge lines with indication continuously available in the McR.
The two pump discharge lines are connected with a ' normally open crossover line so a pressure increase in this segment would be detectable by either sensor..Three pressure relief valves are also provided for these SIS lines.
When the pressure in the line reached 1750 lb/in2g, the relief valves would discharge a total of 60 gal / min to che pressurizer relief tank.
Discharge into this tank would increase the tank level, pressure, and temperature.
A level sensor is provided on the tank having both continuous indication and alarm available in the NCR.
In 3ddition, verification by sample that a sufficient boron concentration exists in the SI pump suction piping will be perfo med if required by Abnormal Operating Procedure AOP-R 05, "RCS Leak and Leak Source Identification".
This will ensure that a sufficient boron concentration exists-in the suction piping if leakage past three check valves occu.s.
Based on the system design and the Technical Specification-re, quired venting surveillance (with a 31 day frequency) of the SI pump and piping,-it is technically acceptable to allow the assumption of a 1950 ppm borbh concentration in the SI pump and suction piping for the : main steam line break
. event. -Assumptions regarding the modeling of boron
' injection for the main steam line break event (i.e.,
assumption of Oppm from the reactor vessel to the discharge of the SI pump) will be clearly stated in revisions to-the-FSAR related to the Mark-3W fuel load at Sequoyah,
EL 740-RwSt EL 706*
63-536 g
1750 PSI 63-582 PI 63-558 63-547 A
63-559 63-549 g
-Q j Et 669' v:
vi o
v:
],.
p{A g.,.63-857 63-526 pj 63-535 63-560 63-553 INIIIAL COLD LEG s'50 PSI 8
63-48 63-550 63-5 63-561 63-553 INJ[CilON PATH 63-562 63-555 SIP 63-563 63-557
'[
/)
(
()
,b 63-558 Ra PI
- 220 PSI 63-22 b 58 a
63-537 PI 63-641 63-545 A
63-9 I
j 63-644 63-543 W
Y
(
( ()
b 6*
6 3-t S6 g3 57, ma PI A
63-47 63-12 63-2 IP 63-167 (
J
}
10 NOLD UP TANK 63-Fi 63-8.
.3-2, Q SIMPLIFIED FLOW PATH RWST/ SIP / VESSEL Figure 20-1
21.
Explain why there is no flow to the intact steam generators after 20 seconds into the main steam line break transient analysis (Figure 6.4-13).
Shouldn't en'ergency feedwater start injecting when main 2eedwater is isolated?
Response
The Sequoyah Nuclear Plant design criteria for auxiliary feedwater (SQN-DC-V-13.9.9 Revision R6, revised through 3/7/96) require that the maximum flow contribution from the AFW system to containment during a main steam line rupture must not exceed 2250 GPM.
A conservatively high auxiliary feedwater flow rate (2350 GPM) is assumed to be delivered only to the f aulted steam generator to increase the severity of the core cooldown during the main steam line break transient.
The assumption is consistent with the analysis of record (see FSAR Section 15.4.2.1.2).
This~ is why there is no flow to the intact steam generators (SGs) af ter main feedwater flow is isolated af ter 22 seconds as per Table 6.4-1.
In fact, the intact SGs become a heat source shortly after closure of the MSIVs (at 10 seconds as per the Table 6.4-1).
Thus, distributing a part of the emergency flow to the intact SGs would have resulted in less severe overcooling.
t s-to i
4 h
FCF Non-Proprietary l
.---m.._,.
~.
. -. ~...
22 o A acatistical core design (SCD) methodology is'used to analyze some of the transients and used to derive some safety limits, peaking limits, and departure from nucleate boiling ratio (DNBR) limits.
Describe how the non-SCD transients are used to provide input to the same limits.
Also, identi'y the conditions for their application to this reload.
Response
The DFB-based safety limits and DNB-based peaking limits were developed using the SCD methodology as defined in BAW-10170P-A.
The SCD methodology is applicable because the system and core conditions for these calculations remain within the allowable parameter ranges defined in Table 2.2 of BAW-10170P-A, as shown below.
Parameter Units Allowable Parameter Range Q
Reactor _ Thermal (v of 3411. MWt)
(
)
Power W
RCS Flow
(% of Nominal
(
)
Flow)
P System Pressure psig
(
)
T Inlet Subcooling
- F
(
)
Similarly, those transients that have predicted system and core conditions within the range of the parameters shown above were analyzed with the SCD methodology.
However, those transients that have predicted system and core conditions falling outside the range of the parameters shown above are analyzed using non-SCD methods (where the parameter uncertainties are treated in a deterministic manner).
The non-SCD transients include the steam line break event and the suberitical rod withdrawal event.
The results of these transients are not used in establishing DNB-based safety and/or peaking limits.
However, the non-SCD analyses are used to evaluate the relevant core parameters for-the Sequoyah core and establish that the appropriate acceptance criteria are met.
For the steam line break, it was shown that DNB does not occur and that the fuel.and cladding temperatures are acceptable (Section 6.1.4 of BAW-10220)
For the subcritical rod withdrawal, the non-SCD evaluation established that the reference analysis remained bounding and applicable (Section 6.2.1 of 9AW410220).
- FCF Non Proprietary
23:
Is rod bow explicitly accounted for in the DNBR methodology?
If so, where is it accounted for (retained margin, DNBR penalty, peaking factor adjustments) ?
l Responce Yes, the impact of potential fuel rod bow is explicitly accounted for in FCF's DNBR methodology, As discussed in Section 4.1.1,7 of BAW-10172P " Mark-BW Mechanical Design Report", the Mark-BW fuel design has several features (non-rigid grid attachment and keyable grids) that make its fuel rod bow performance similar to that of other FCF fuel designs.
In BAW-10186P " Extended Burnup Evaluation", FCF presented new data that extended the rod bow data base for FCF fuel to ( ).
The topical report concluded that the rod bow correlations from BAW-10147PA-R1 " Fuel Rod Bowing in Babcock and Wilcox Fuel Designs" are applicable at extended burnups and apply to the Mark-BW.
BAW-10147PA-R1 has shown (
). The average power hot channel
-factor, Fj, has a value of 1.03 and is combined statistically with other uncertainties to establish the statistical unsign limit (SDL) DNBR used with the statistical core design method (discussed in Section 7.2 of BAW-10220 and Section 3.2.2 of BAW-10170P-A).
For non-SCD analyses, the average power hot channel factor, which includes the rod bow effect, is incorporated into the LYNXT model as a direct multiplier on the hot pin average power.
Copies of pertinent pages from BAW-10220P and BAW-10170P-A have l
been attached.
A total of six (6) attachment pages follow.
4 4
- M TCt Non-Proprietary 1
etrn u nm u T v L>ponse e=
Wue s, u sn 4$
PcedMed b^}o b
~ IO Radial peak (Fj) 1.64
=
Axial Peak (F,)
1.35
=
Axial Peak Location = 0.5 x/L (steady-state) 0.7 x/L (transient)
=
Dundle Average Peak = 1.557 where x/L = normalized axial location along the heated length It should be noted that the radial peak of 1.64 corresponds to a maximum allowable radial peak of 1.70 when a 4% total rod power uncertainty factor is included.
_7.1.4 Core Conditions A summary of general core conditions used in the SON thermal-hydraulic analyses is provided on Tables 7.1-1 and 7.1-2.
7.1.5 Engineering Hot channel Factors Engineering hot channel factors (HCF's) are penalty factors that are used to account for the effects of manufacturing variations on the maximom linear heat generation rate and enthalpy rise.
2.1.5.1 Local Heat Flux Engineerina Hot Channel Factor The local heat flux engineering hot channel factor, F is used in the evaluation of the maximum linear heat generatiol,n rate.
This factor is determined by statistically combining manuf acturing variances for pellet enrichment and weight and has a value of 1.03 at the 95% probability level with 951 confidence.
As discussed in References 7-2 and 7-3, relatively small heat flux epikes auch as those represented by Fj have no effect on DNB, therefore this factor is not used in DNBR calculations.
7.1.5.2 Average pin Power Engineering Hot channel Factor The average pin power f actor, Fj, accounts for the effects of variations in fuel stack weight, enrichment, fuel rod diameter, and (fin __ pitch on hot pin average power)
This factor, which has a value of 1.01, is combined statistically with other uncertainties to establish the statistical design limit (SDL) DNBR used with the statistical core design method (discussed in Section 7.2).
Since Fj is incorporated into the statistical design limit (SDL), this factor is not included in the LYNXT model used for SCD analyses.
For non-SCD analyses, Fj is incorporated into the LYNXT model as a multiplier en the hot pin average pcwer.
7-3
AHa w.d v. tasP.ssa Y t%2sf!K 23 D
EdMcot by k WIO220NP Non Propriotary Table 7.2-1
. Statistical Core Design Application Summary i
i Sequoyah Plant Specific fincertainties Plant Uncertainties Variable Name Uncertainty D,tstributlen Q
Core Power
(
)
i W
Core Flow
(
)
P Core Pressure
[
]
T Core Inlet Tettperature
(
)
R Measured F5
(-
1 FCF Analysis Uncertainties Variable Name Uncertaintv Distrlbution W
Core Bypass Flow
(
l
-)t-Hot Channel Factor
(
))
R Bundle Spacing
(
)
t A
Axial Peaking Factor
(-
3 Z
Axial Peak Location
(
l D
BWCMV Uncertainty
(
)
D LYNXT Code Uncertainty
(
)
D RSM to LYNXT Fit
(
3 I
4
'Also applies to the Westinghouse VyNAGE SH 9b 7-12
/414uIwu,d to F-osponsa 19 %e4f.w 33 Pe, knc,d Pa3e. bm 6 A ld - 10 M O - A Non Propriotary Coding follows finding of the Xi (and any transfornations).
- Thus, 1,XbR Xg (coded) = 2A 3 R t
where XbR is the RSM mean, and DXg is the full RSM range for the variable X.
3.2.2.
Uncertainties for Procacation The uncertainties propagated through the RSM to arrive at a final statistical Design Limit (SDL) will define the new LYNXT design analysis.
That is, if an uncertainty (such as a 4 degree temperatore error) has peen propagated in determining the final SDL, it need no longer be considered in the LYNXT design case, and the nominal value can be used.
The RSM, through which the uncertainties are propagated, results in a minimum core DNB ratio, D, for each core state input vectort Q, W, P,T,R,A, and 2.
All applicable uncertainties, including those on D, will be treated.
If an uncertainty is not treated, such as the 0.95 VMFT flow factor, it nust be considered directly in the core thermal-hydraulic analysis.
Further, if an uncertainty is treated at an inferior level, the remaining portion of that uncertainty will be compounded in the core analysis.
Thus, for instance, if a 30 psia pressure error has been propagated, but it is later found that the actual error is 45 psia, the remaining 15 psia must be compounded in the core analysis.
For this
- reason, conservative values of each applicable uncertainty are treated.
s 3-6 i
i
y-Pu he.d Pap 6 BM -l0 N -A Non Propriotary Finally, for a slight added conservatism, appropriate uncertain-l ties are propagated with a uniform distribution when it is questionable that a normal distribution can be justified.
A brief discussion of each uncertainty.to be propagated follows.
i 1.
Q (CORE POWER) - 1 Uncertainty - A two percent normal random heat balance error is treated.
2.
W (CORE FLOW) 2 Uncertainties A 2.2 percent RCS flow uncertainty and a 1.5 percent core bypas's uncertainty are treated.
Each uncertainty is separately propagated as random unifora errors.
3.
P (CORE PRESSURE) - 1 Uncertainty - The pressurizar pressure uncertainty of 30 psia is treated as random uniform.
4.
T (INLET SUBCOOLING) - 1 Uncertainty - This error is derived from the uncertainties in inlet temperature from both the rea,ctor (turbine) and rod control systems.
The total error of 4 degrees F is treated singly as a random uniform error.
5.
R (RADIAL PEAKING FACTOR) 3 Uncertainties Three uncertainties on the radial, all random and normal, are treatedt an upper-- limit of 5
percent (from either calculational or neutron detector errors) percenthot) 3,.
[ channel factor on rod-power,jana a 1.s percent increase in hot rod power due to local and bundle spacing.
6.
A (AXIAL PEAKING FACIOR) 1 Uncertainty
-A 2.5 percent random normal uncertainty from calculational sources is treated.
t 7.
2 (AXIAL PEAK LOCATION) - 1 Uncer'tainty - An upper limit
\\
8 inch nodal uncertainty on axial peak location is treated as random normal.
3-7
+
4 4
1 1
a
+
g p p;g; ggp;;igT; ale 3KCz3 Pufroed P)
Fro 7 B4v-to l70 -6 Non Proprietary 8.
D (Uncertainties
'n DNBR) 3 Uncertainties The random normal uncertainties on DNDR to be treated are the BWCMV correlational uncertainty of 0.1667 (which is K95 times the 0.1020 measured to predicted standard deviation),
a 5 percent code (LYllXT) uncertainty, and a 2.5 percent uncerJainty on the RSM to LYNXT fit.
The RsM uncertainty is the RsM standard error of (
) *' normalized to the midpoint DNBR of 2.533 and multiplied by K95.
This value [
Jwas arbitrarily increased to 2.5 percent.
These uncertainties and their distributions must be verified for each separate application (core).
All uncertainties are at the 95 percent, one sided level.
A summary of the above is shown in Table 3-4.
f 4
0 e
3-8
JugKi+~+;~Fi((ca up Ye [sb'. d,as % 25 PerWCef P p. F - S A W-lo 70-A Non Propriotary TABLE 3-4 UNCERTAINTIES TOR PROPAGATION. SUlciARY VARIABLE UNCERTAINTY LEVEL DISTRIBUTION Q
!!aat balance 2.0 %
Normal W
Reactor Coolant Sfstem 2.2 %
Uniform W
Core Bypass 1.5 %
Uniform P
Pressurizer 30 psia Uniform T
Control System 4
F Uniform
-R Calcul'ational/Datector 5.0 %
Normal y
Hot Channel Factor 3.0 %
Norma R
Spacing 1.5 %
Normal A
Calculational 2.5 %
Normal Z
Calculational 8 inches Normal D
Correlation (BWCMV) 0.1678 Normal D
T.H Code (LYNXT) 0.05 Normal D
RSM to LYNXT Fit 0.025 Normal a
t s
3-11
24:
Why is the RCS pressure used for DN3R purposes (Table 7 1-2) 4 chosen to be 2280 psia when the nominal RCS pressure is 2235 ps.ig?
Response
The RCS pressure quoted for DNB analyses is the pressure that sets the boundary condition for the analyses, that being the core exit pressure.
A 2280 psia core exit pressure used for core DNBR predictions (Table 7.1-2) is based on a nominal RCS pressure of 2235 psig (at the pressurizer) with adjustments for gage pressure and pressure drop from the pressurizer to the core exit.
I t
Nominal
+ 15 psi adjustment + 36'psiv.sssure drop from the Cue Exit
=
Pressure RCS Pressure to gage pressurizer to the core exit 30 psi 2235 psig
+ 15 psi 2280 psia
+
=
i i
p t
t 4
4 4
9 FCF Non-Proprietary.
4 25:
Please provide a b 6.ification for the Sequoyah-specific uncertainties chose. to calculate the statistical design limit.
Response
The Sequoyah-specific uncertainty allowances incorporated into the determination of the statistical design limit (SDL) for core power, RCS flow rate, RCS pressure, inlet temperature, and I
measured FJ are justified as follows:
Parameter Uncertainty Batts Core Power
()
This value represents the maximum allowance for calorimetric measurement uncertainty and is consistent with the maximum steady state uncertainties outlined in Sequoyah FSAR Section 15.1.2.2 Core Flow
()
This value represents a bounding
- generic uncertainty" for RCS flow measurement which is consistent with Technical Specification
(
3/4.2.5 (Table 3.2-1).
The value is conservative with respect to actual flow measurement techniques used to verify minimum RCS flow values which have 2.4% uncertainty (i.e. elbow-tap data normalization to baseline calorimetric flow measurements).
Core Pressure
()
This value represents the maximum-allowance for steady' state system pressure fluctuations and instrument measurement errors.
It is consistent with the maximum steady state uncertainties outlined in Sequoyah'FSAR Section 15.1.2.2.
Core Inlet
()
This value represents the maximum Temperature allowance for deadband and instrument measurement errors.
It is consistent:with the maximum steady state uncertainties outlined in-Sequoyah'FSAR Section 15.1.2.2.
.a FCFNon Proprietary
Measured Fa".
()
This value represents a bounding uncertainty for measurement of the enthalpy hot channel factor using the incore neutron detectors and is consistent with the uncertainty required by Technical specification 4.2.3.2.c.
Table Q25-1 provides-a listing of the generic uncertainty values extracted direccly from Table 3-4 of BAW-10170P-A.
Table Q25-1 also provides a listing of the Sequoyah-specific uncertainties.
As the table shows, for the sequoyah plant-specific application of the statistical core design methodology, the uncertainty on core flow is [] and the uncertainty on the measured Fa", is ().
All other uncertainty values are the same as those used in BAW-10170P-A, Since two of the uncertainty values were different
~
than those used for the generic SDL,. a new SDL calculation was required. Results--from.this SDL calculation produced a maximum hot pin SDL which is bounded by the generic value [] documented
~
in BAW-10170P-A.
To maintain consistency with BAW-10170P-A, the generic value was adopted for use in the Sequoyah analyses.
I i
?
6
?
l FCF Non-Proprietary
?..--d I
4.1i--r-r-1--e
- ev+-=een-ow-+-&
n-w rm,we
--$mrm---.<'.--.ne
---s-e=-s-e
- ww ww. r,
-.-,---o
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e,-wsu W -eg www
-rw-,--++,=v-w.y
_.._ _._.. _ _ __._._____.-._.. _._ _ _._ _.m._
f t
Table Q25-1
[
Statistical Core Design Application Summary Sequoyah Plant Specific Uncertainties Plant Uncertainties l
l Generic Uncertainties Sequoyah From Specific Variable Name BAW-10170P-A Uncertainties Q
Core Power 2%
()
W Core Flow 2.2%
()
e P
-Core Pressure 30 psi
()
T Core Inlet Temperature 4*F
()
R Measured FA",
5%
()
FCF Analysis Uncertaincies l
Analysis Generic Uncertainties Uncertain,Les Used For From Sequoyah Variable Name BAW-10170P-A Verification W
Core Pypass Flow 1.5%
()
R Hot Channel Factor 3%
()
l.
R Bundle Spacing 1.5%
()
l j
A Axial Peaking Factor 2.5%
()
I Z
Axial Peak Location 8 inches
().
D BWCMV Uncertainty 16.78%'
()
()
D LYNXT Code Uncertainty 5%
D RSM to LYNXT Fit 2.5%
()
e
(
FCF Non Proprietary l
t
26:
The critical heat flux correlation used for non'-LOCA applications is approved for both the Framatome and Westinghouse fuel; however, it has not been approved for mixed core applications.
Provide an appropriate penalty to the limiting rod that will bound the misapplication of the critical heat flux correlation.
Additionally, the analysis provided does not include the Westinghouse standard fuel in the mixed core penalty.
Provide a justification why this is not accounted for.
Response
The BWCMV CHF correlation is approved for application to homogenous cores of Mark-BW or Westinghouse VANTAGE SH fuel, In addition, FCF has demonstrated that !
) This conclusion is based on the fact that the BWCMV CHF performance level is applicable to the VANTAGE 5H fuel design and that the hydraulic ef fects of transition cores are conservatively evaluated in the method used to predict steady-state and transient DRBRs.
Specific mixed core analysis results that support this conclusion were submitted by letter to the NRC on 1/10/97 (1) and are discussed in greater detail in the response to Question 28.
The presence of Westinghouse standard fuel assembly reinserts in Sequoyah transition cores is not significant because: 1) the number of standard fuel assemblies planned for use is relatively low, and 2) the Westinghouse standard fuel assemblies will have relatively low-powered operation as reinserts and will therefore be non-limiting.
Also, the mixed core penalty analysis doe's not include the Westinghouse standard fuel because the effect of these assemblies is bounded by the effect calculated for the transition from VANTAGE SH to Mark-BW fuel.
The transition penalty applicable to the transition from Westinghouse standard fuel to the Mark-BW was shown in BAW-10178 to be (
)
This penalty was calculated with the same mixed core method and modeling technique as used to calculate the corresponding penalty for transition from VANTAGE SH to Mark-BW, In both cases the penalty was determined for the bounding case, (
)
Thermal-Hydraulic Methods for the Transition from Vantage 5H to Mark-BW Fuel at TVA's Sequoyah Plant, JHT/97-2, January 10, 1997 FCF Non-Proprietary
l 27:
Describe background information and the bases for those studies related to the Trojan Plant which concluded that a
() transition core DtGR penalty should be applied to the Mark-BW when it is being inserted into a Westinghouse i
standard core with respect to the hydraulic compatibility of j
the Mark-BW fuel design with the Westinghouse standard i
design.
Identify the similarity or difference in relation to the transition core DNB penalty between the sequoyah and Trojan reloads.
j
Response
i Troian Transition Penaltv The analysis methods used to determine and apply the transition core penalty for the introduction of the Mark-BW at Trojan are the same methods used in the analyses supporting the transition to Mark-BW fuel at Sequoyah.
The principle followed in the mixed core analyses was to determine the minimum DNBR performance for the core by examining all bounding mixed core combinations and using the most limiting result to establish the transition core penalty.
The bounding combinations included:
Configuration Configuration Number Description 1
Full Core Resident Fuel Design 2
[]
-3
()
4 Full Core of Mark-BW Fuel A [] transition core penalty was applied to the Mark-BW at Trojan.
The penalty was determined by examining the DNBR differences between Configurations #2 and #4 for statepoints that set the DNB-based safety limits and operating limits as well as the limiting DNB transient.
More specifically the transition penalty analysis for the Westinghouse standard to Mark-BW transition at Trojan was performed at the following conditions:
core safety Limits 6tttepoints:
()
- eE FCF Non Proprietary
l Limiting Transient l.
Conditions:
()
i The () penalty was based on the 6DNBR for ( ).
The ADNBR was calculated as follows:
Full Core Mark-BW DNBR
()
Single Mark-BW in STD Core DNBR 11 Difference
()
Penalty - ()
Although the transition core penalty would have been smaller than
() if the DNB reload analyses had considered the actual cycle-specific core distribution of fuel designs With a Configuration somewhere between Configurations #2 and #3, the conservative application of the () transition core penalty assured DNB protection.
It should be noted that the Trojan transition core analysis was based on a total pressure drop difference of () between the Westinghouse standard and the Mark-BW.
(Pertinent pages from BAP-10178p, which provide the details from that analysis, have been attached.) ()
This change relative to the Trojan application is discussed further in the response to Question 31.
Shquovah Transition Penalty f
The [] transition core penalty (applied to the Mark-BW results using Configuration #4) was determined for the Mark-BW introduction at Sequoyah, in a similar manner.--Specific mixed core analysis results-that support the () transition core penalty
-were submitted by letter to the NRC on 1/10/97 (1) and are discussed in greater detail here.
For each fuel type, the following discussions begin with a summary of how peaking limits were developed and then proceed to a description of how the transition penalty was developed.
Mark-BW Peakina Limi ts
- Peaking limits for the Mark-BW fuel-were established using the
().
The following conditions were evaluated:
1 i
4 k
4
. FCF Non Proprietary 9
w g-g gqir yv a g 9 T**--*
e v
w w-r-g-r'tr'a-*-'
g-rst wr 9-w
+-5 y
we rw'-wov-=
-r r
w-%
--sv-1--
Vv--'-*eir&P---
-vmp-y--w
j
\\
Core Safety Limita Conditions:
()
Peaking:
[]
Limiting Transient I
Condit ions :
()
Peaking:
[}
Mark-BW Tranel tion Penal ty The transition penalty determination for the Mark-BW involved i
evaluating the ADNBR when the peaking limits determined with ( )
(configuration 2).
The following conditions were evaluated:
Core Safety Limits conditions:
()
Peaking:
()
Limiting Tranalent conditions:
()
Peaking:
()
The () penalty was the maximum calculated difference and was based on eDNBR for ().
The ADNBR was calculated as follows:-
e FCF Non Proprietary
~
Full Core Mark-BW DNBR
()
Single Mark-BW in V5H Core DNBR 11 Difference
()
Penalty = []
Vantaae SH Peaking Limite vantage SH peaking limits were determined using the 12-channel LYNXT model:in both a full core vantage SH configuration and in a single Vantage 5H in a Mark-BW core configuration.
Final peaking limits were established by overlaying the two sets, taking the most limiting combination, and_ setting that as the limit.
The following conditions were analyzed:
Core Safety Limits Conditions
()-
Peaking:
()
. Limiting Transient conditions:
()
Peaking:
[]
Vantaae SH Transition Penaltv Since the Vantage 5H penalty limits were established-using (),
the hydraulic effects of any potential fuel mix are accounted for in the final limits.
Therefore, ().
1.-
LettE*.
J. H. Taylor, Framatome Technologies te USNRC, Thermat Hydraulic Methods for the Transition from Vantage SH to Mark.?W Fuel at TVA's Sequoyah Plant,
- JHT/97-2, January 10, 1997
- FCF Non Proprietary
-,c Attachment to Response to Question 27 ECE MON-PROPRIETARY Pertinent Page From BAW-10178P a
4
- 7. MIXED CORE ANALYSIS i
For any new fuel design.that is being introduced on a reload basis,
-compatibility must be establinhed with the' existing, or resident, fuel in the core.
Therefor'e, when a..ew or modified fuel design, i
having J.ifferent hydraulic characteristics from the resident fuel is' introduced, a transition core analysis is performed.
For each mixed core configuration during the transition, performance of each fuel type is eveluated relative to a reference analysis.
This 3
-refertnce analysis is typically based on a full core of the new i
fuel.'
To determine the performance of each fuel type relative to the reference
- analysis, each fuel type le-modeled in a
conservatively bounding mixed core configuration. Theso mixed core analyses are. used to determine the transition penalty that is applicable to each fuel type.
A typical transition core analysis determines the effect of the mixed core on parameters such as minimum DNBR, core pressure drop, fuel assembly lift, and lateral flow velocities, i
The Mark-BW fuel ass'embly, described in BAW-10172P (Reference 6),
was designed to be hydrauli v.11y compatible with the Westinghouse Standard (STD) 17x17 fuel design.
During the transition to Mark-BW fuel at the Trojan Nuclear Plant, the resident fuel (i.e. the fuel being displaced by Mark-BW fuel assemblies) will be the STD design.
Therefore, the following discussion relates to the
-transition from the STO fuel design to the Mark-BW.
-7,r.
Flow Testina To' verify the compatibility between the Mark-BW reload fuel and the Westinghouse STD fuel, a series of fIcw tests was performed in 7-1
Attachment to Response to Question 2 Pertinent Page From BAW-10178P ECT NON-PROPRIBThRY 1988, using BWFC's transportable flow test rig (TFTR).
The TFTR is a " cold-flow" loop, with typical test conditions encompassing water flow rates of 500 to 1500 gpm at pressures from i")
to 160 psig and temperatures from 130'F-to 140*F.
Pressure crep testing was performed on individual fuel assemblies, providing pressure drop characteristics for the. STD fuel design, as well as for the Mark BW both with and without a debris-filtering bottom nozzle design.
4 Prior to testing the Westinghouse STD assemblies, a preliminary sat of flow tests was performed to qualify the TFTR, establish a benchmark for comparison to later
- tests, and provide the incremental pressure drop for the Mark-BW debris-filtering bottom nozzle.- These tests were performed with the same Mark-BW prototype assembly that had been previously tested at the B&W Alliance Research Center, Alliance, Ohio, in the Control Rod Drive Line (CRDL) flow loop.
The CRDL test spanned a wide range of pressure, temperature, and flow conditions encompassing those that occur during_ reactor operation.
Core thermal-hydraulic analysis models L
for the Mark-BW fuel assembly
- design, including
ra loss coefficients' for individual components, were oevel.oped from the CRDL test results.
The measurements obtained with the TFTR demonstrated that the TFTR reproduces the CRDL results at comparable Reynolds numbers.
Once the TFTR qualification runs were completed, an additional test program was carried out to test the Westinghouse STD design.
For that test program, indiv_idual flow tests were performed on four STD assemblies.
The following paragraph describes the test sequence I
used in the STD flow test.
mi i
4 m
7-2 1
j-J
Attachment to Response to Question 27 FCT MON-PROPR1ETARY Pertinent Page From BAW 10178P
\\
9 The results of the TFTR test progratas demonstrated that the total pressure drop of the Westinghouse STO design is approximately{
]than that of the Mark-BW with a debris-fiJtering bottom nozzle.
A plot of unrecoverable pressure drop versus axial elevation for the two designs is provided on Figure 7-1.
7.2. Mixed Core DNBR Analysis For Westinghouse STD to Mark-BW transition cycles, core DNBR safety and operating limits and DNBR margin during transients are based on analysis of the final full-core Mark-BW configuration.
Similarly, the reference reload safety analysis as documented'in BAW-10176p (Reference 11) is performed for the Mark-BW fuel design.
To ensure that these full-core analyses remain applicable to the Westinghouse STD/ Mark-BW transition cycles a mixed core analysis is pe Hormed.
This mixed core analysis quantifies the transition.
cycle penalty that must be applied to either the resident or the new fuel designs.
To maintain the applicability of the full-core
- analyses, the transition core penalty is assessed against the retained thermal margin that is incerporated in the DNB analysis through the use of the thermal design limit (TDL, per BAW-10170P-A, Reference 8).
The transition core penalty is a DNBR penalty that is determined initially for a conservatively bounding mixed core configuration.
7-3 i
l
Attachment to Response to Quest on 27 Pertinent Page From BAW10178P.
ECE NON-PROPR1EThaY' For applications where.c is' desirable to reduce the penalty to a value less--than this-bounding value, a core-speciffe calculation is performed,- with a model that represents the actual transition cycle core geometry.
The calculation process is'as follows:
l l
l l
I l
?
L.
For reload licensing safety evaluations',
the transition core penalty is assessed against-the retained.thernal margin that is incorporated into the safety analyses throbgh the use of a therr.a1 design limit that'is greater than the statistical design limit.
Wh
_ _ - - - - - - - _ - - - - - =
Attachment to Response to Question 27 FCi NON-PROPR1ETARY Peninent Page From BA W-10178P Iri cp ;cific applic2tions the penalty applied is designated as either the generically bounding or the cycle-specific transition core penalty. The generic transition core penalty for introduction of' Mark-BW fuel into a core of Westinghouse STD assemblies is provided on Table B-3 in Appendix B.
- 1. 3. Mixed core Pressure D'ron. Lift and crossflov For the transition,from one fuel design to another, the mixed core pressure drop is bracketed by the homogeneous core pressure drop values for each of the fuel designs.
Therefore, mixed-core pressure-drop calculations are not required.
For fuel assembly hydraulic lif t force determination, the limiting core configuration is the *ull core configuration with the highest core pressure drop.
For the Westinghouse STD to Mark-BW transition,{
]
As part of the mixed core methodology, an evaluation is per formed to determine the limiting configuration for lif t forces on each assembly type.
This evaluation is based on the relative pressure drops of the two fuel designs.
Knowing the relative pressure drop of the two assembly types, the flow diversion characteristics of the core can be determined, and in turn, the core configurations which produce the highest lif t forces in 'the two fuel typ.es can be determined. Mixed core pressure drop and hydraulic lift force calculations are determined by modeling the mixed core with LYNX 1, LYNX 1 is also used to calculate inter-assembly crossflow velocities, which a.re used to evaluate the potential for flow-induced vibration.
Since the total pressure drop values of the Mark-BW and the Westinghouse STD design [
] Therefore, no mixed core lif t calculations are necessary.
Inaddition,[
]and fuel rod diameters are identical, lateral crossflow is very small.
Calculations for the Mark-BW LTA's, 7-5
Attachment to Nsponse to Question 27 ECW 808-PROPRIEThay Pentnent Page From BAW10178P designed for use in Trojan Cycle 13, showed lateral flow velocities
{
].whichiswellwithinthe[
]limitthat' is used:as an evaluation criterion.
7.4.
Mixed core Analysis Summary The - Mark-BW fuel _ assembly has been determined to have a total i
pressuredropthatis}
] than the Westinghouse STD fuel design.
Based-on bounding analyses that address the small differences in pressure distribu' tion that exist, a Mark-BW fuel assembly in a mixed Mark-BW/STD core will be subject to a.small I
transition penalty.. The STD assemblies, having a slightly lower pressure drop and the same fuel rod diameter as the Mark-BW, vill have no transition penalty.- The transition-penalty for the Marka BW 4 provided on Table B-3 in Appendix B.
7.5. Methods and Models for Resident Fuel The Mark-BW is similar n design.to the Westinghouse STD 17x17 fuel L
' design, with the primary difference being that the Mark-BW has zircaloy intermediate-spacer grids while the STD design has inconel
. grids.
Table 7-1 provides a comparison of important parameters for the two fuel designs.
From the thermal-hydrau)ics perspective the two designs are identical with-the exception of the hydraulic form loss coefficients-for the-spacer grids and nozzles.
Therefore the
_ thermal-hydraulic core models described in section 2. apply also to the STD fuel, when the appropriate form loss coefficients-are used.
The BWCMV CHF correlation (BAW-10159P, Reference 9) includ' es in its data base 17x17 CHF data representative.of the STD fuel design and BWCMV has been approved for application :: this fuel (Reference
- 12). The statistical core design thermal'-hydraulic analysis method documented in BAW-10170P-A was developed independently of fuel type,- therefore the scD method applies;;: the STD fuel when the 9
appropriate'_ uncertainties are i n c o r p o r'a.e d.
The treatment of
-uncertainties in the development of.a statistical design limit is 7 - 6,,,
Attachment to Response to Quest on 27 FCE NON-PROPRZEThRY Peninent Page From BAW10178P-discussed in section 3.1 of this report.
From that discussion it can be seen that the only uncertainty parameter dependent on the fuel design is the engineering hot. channel factor.
A suitably conservative value]
]has-beenselected for this parameter, such-that the SDL applies to the STD fuel.
7.6. DNBR Penalties Aeolicable to the Resident Fuel For Westinghouse STL to Mark-BW transition cycles, core DNBR safety
'and operating limits and DNBR marg}h during transients are based on analysis of the final full-core Mark-BW configuration.
This I-full-core analysis is based on the statistical core design thermal design limit that includes 10 percent thermal margin.
As stated in Section 7.2, there is no transition core DNBR penalty on the l
Westinghouse Standard.
There. is, however, a 1.5 percent DNBR penalty to account for fuel rod bow in the Standard fuel.
The 10 percent margin t *7t is available in the thermal design limit more than offsets the rod bow penalty.
9 4
7-7
Attachment to Response to Quest $n 37 ECE NOM-PROPR1EThRY Pertinent Page From BAW-10178P Table 7-1 Comearison of Mark-BW and W STD Desie_n__
17x17 B&W=
17x17 g Mark-BW Standard Fuel - Assembly Fuel Assembly Parameter De'sion Desien Fuel Assembly Length, in.
159.8 159.8
. Fuel Rod: Length, in.
151.500 151.630 Assembly Envelope, in.
8.425 S.426 Fuel Rod Pitch, in.
.496
.496 Humber of Fuel Rods / Ass'y_
264 264 Number of Guide Thimblas/ Ass'y 24 24 l.
Number of Instrumentation Tube / Ass'y 1
1 Fuel. Tube Material Zircaloy-4 Zircaloy-4 Fuel Rod Clad O.D.,
in.
.374
.374 Fuel Rod Clad Thickness, in.
0.024 0.0225 Fuel / Clad Cap, mil 6.5 6.5 Fuel Pellet Diameter, in.
0.3195 0.3225 Fuel Pellet Density, % TD 95 95 Guide Thimble Material Zircaloy-4 Zir aloy-4 4
Inner Diameter of Guide Thimbles (upper part),.in.
O.450
-0.450 Outer Diameter of Guide Thimbles (upper part), in.
0.482 0.482 Inner Diameter of Guide Thimbles (lower part), in.
0.397-0.397 Outer Diameter of Guide Thimbles (lower part), in.
0.4,29 0.429 Inner Diameter of Instrument Guide Thimbles, in.
0.450 0.450 Cuter Diameter of Instrument Guide Thimbles, in.
0.482 0.482 7-8.
Anachment to Response to Question 37-ECE NON-9RO981EThRY Pertinent Page From BAW-10178P Table 7 Connarison of Mark-BW and W STD Desian (Cont'dl 17x17 B&W 17x17 g Mark-BW Standard Fuel Assembly Puel-Assembly Parameter Desian Desian composition of Grids 2 Inc.-718 2 Inc.-713 End Grids End Grids l
1 Zircaloy-4 non-mixing Grid 5 Zircaloy-4 6 Inc.-718 Mixing Grids Mixing Grids Top Nozzle Holddown Springs 3-Leaf 3-Leaf r
a G
t
-6 7-9
Attachment to Response to Question 27 Pertinent Page From BAW10178P ECE NON-PROPRIETAnt A
. Figure 7-1 Unrecoverable Pressure Drop for Full Cores of
- Mark-BWs and PGE Standard Lattice Assemblies i-euuimumme e N
e 4
I N
4 4
=_ M i
- u' 7-10
--- ~5ME8nt to Response to Quenibn 27-__----- - -
' Pertinent Page From BA W10178P FCT NoN-PROPR1ETARY Tcblo B-3 cont.)
Statistical Core Design (Application Summary Retained Therwal Marain The statistical design = limit (SDL) for the Trojan core based on the uncertainties listed on the previous page has.been calculated as 1.345 BWCMV.
Also, for the Trojan analyses the TOL is 1.50 BWCMV.
The retained thermal margin-based on these values is as follows:
Retained Thermal Margin (%) = 1. 5 0 -
- 1. 3 4 5_ X 10 0 = 10 %
1.50 Penalties & Offsets to be Assessed Against The Retained Thermal Margin l
Penaltv/ offset Value Mark-BW W STD i
Transition Core Penalty m
Rod Bow 0%
1.5%
o -
I O
4 B-9 1
28:
Provide the final conservatively bounding mixed core configuration for SQN mixed core DNBR analysis and the transition penalty based on assuming that the center hot assembly is either a single Mark-BW fuel assembly in a core of the VANTAGE SH or a single VANTAGE 5H in a Mark-BW core.
Also, provide.the result of the DNBR analysis using plant-and cycle-specific core loading configuration and the same limiting power distribution input in the above analyses.
Show that the VANTAGE SH to Mark-BW design peak difference will offset any transition core effects on the VANTAGE SH and provide the description of the retained thermal margin in relation to the transition core penalty.
Response
The method of determining the DNB-based safety limits and peaking limits for.the mixed core situation at Sequoyah is described in detail in the discussion of Sequoyah transition core thermal-hydraulic analysis methods submitted by letter to the NRC on 1/10/97 (1].
FCF determined the DNB-based safety limits and peaking limits by examining the performance of the resident VANTAGE SH and the Mark-BW in the four configurations defined in the response to Question 27:
Configuration Configuration Nurnber Description 1
Full Cor: Resident Fuel Design 2
[]
3
[]
4 Full Core of Mark-BW Fuel For determining DNB-based safety limits and peaking limits, FCF used thC 12-channel LYNXT model depicted in Figu.es Q28-1 with the pin-by-pin power distributions listed on Tcble Q28-1.
For configuration 2,
[].
For configuration 3,[].
One ground rule established for the thermal-hydraulic DNB analyses was that the existing reactor core safety limits would
- be maintained for Sequoyah.
Therefore, the DNB assessment was composed of statepoint evaluations which, demonstrated that the minimum DNBR remains above the thermal design limit (TDL) of 1.50 using SCD methodology.
The DNB-based rea'ctor core safety limits were evaluated using each of the above ' configurations.
Results showed () thermal margin to the TDL for the existing DNB-based safety limits.
4 FCF Non Proprietary
The VANTAGE SH DNB-based peaking limits were developed using Configurations #1 (full VANTAGE SH core) and #3 ().
Final peaking limits were established by calculating allowable peaks with both of these configurations and then selecting the most limiting values.
The Mark-BW DNB-based peaking limits were developed using Configuration #4 (full Mark-BW core).
The transition core penalty for the Mark-BW was developed by evaluating the DNBR difference between this model and () core model (Configuration #2).
The () transition core penalty for the Mark-BW bounds the most limiting of the DNBR differences determined by that comparison.
One of tne requests in Question 28 is to " provide the result of the DNBR analysis using plant-and cycle-specific core loading configuration and the same limiting power distribution input in the above analyses."
Since the 12-channel LYNXT model (depicted in Figure Q28-1) combines all the fuel bundles outside of the hot bundle into a single LYNXT channel, that model'is not conducive to modeling the cycle-specific core loading.
Therefore, the DNB option in the 31-channel LYNXT model was used.
This confliJuration models the cycle-specific core loading pattern exactly.
The 31-channel model, depicted in Figure Q28-2, is a one-eighth core model in which each channel represents one fuel assembly.
The use of a one-eight core model is reasonable because the reload fuel cycle design is one-eighth symmetric.
Since the smallest channel is one-einhth of a fuel assembly no individual subchannels are modeled.
This approach is acceptable because the results are being used only to demonstrate the relative effects of different core models and power distributions.
The first set of cases run with the 31-channel model was set up to approximate the bounding mixed core model that had been used l
for the 12-channel DNB analysis.
That is, comparing the full core models with the hot bundle in the center core-location, to
[].
The peaking distribution used for this part of the analysis is shown on Figure Q28-2.
That distribution has a 1.600-hot bundle peak which is slightly higher than the hot bundle average power defined for the 12-channel DNB model.
Core conditions for these runs were ().
It should be noted for these calculations that the Mark-BW DNBR values were determined using BWCMV [] while the VANTAGE SH DNBR values were determined using BWCMV ().
The maximum bundle peak for all of the 31-channel calculations was set at 1.600 in order to permit the direct comparison of results from different-configurations.
These cases.were analyzed to provide a basis for comparison to those following, which used the cycle-specific core configuration.
Resul'ts are discussed below.
- =.'
FCF Non-Proprietary
The second set of cases is based on modeling the cycle-specific core loading configuration (depicted on Figure Q28-3) with modified cycle-specific power distributions depicted on Figures l
Q28-4 and 5.
For the cycle-specific core loading models, all three assembly types that~will be in core (MK-BW, V5H and W STD) are modeled in their actual locations.
The cycle-specific power distributions represent the most limiting time in life for the two respective fuel-types:
0 EFPD for the Vantage 5H fuel and 182 EFPD for the Mark-BW fuel.
As shown on Figures Q28-3b and Q28-4b, the cycle-specific power distributions (l_ The purpose of this-set of cases was to evaluate the appropriate scaled power distribution on a full core of each fuel type and on a model with the cycle specific core configuration and to isolate the DNBR difference for the two.
Again, core conditions were based on the nominal operating point.
DNBR results for the Mark-BW are provided in the following table.
Mark-BW Transition Core DNBR Analysis Results Power Distribution Type Design Modified
~
Cycle-Specific Power Distribution Figure Figure Q28-2 Figure Q28-4b LYNXT Model Figure Figure Q28-2 Figure Q28-3 l
Hot Bundle Location in LYNXT Model Channel 1 Channel 11 l
MDNBR for Full Mark-BW Core
[]
{}
MDNBR for Hot MK-BW in S1C9 Core
[]
[]
l Configuration l
MDNBR for []
()
[]
l These results show the following:
[]
These results indicate that the method FCF has used to develop core peaking limits and the transition penalty for the Mark-BW fuel for the initial Sequoyah transition core is bounding and conservative.
For future reload cores, ()
.c FCF Non-Proprietary
~
DNBR results for the VANTAGE SH fuel are provided in the l
following table.
Vantage SH Transition Core DNBR Analysis Results Power Distribution Type Design Modified Cycle-Specific I
Power Distribution Figure Figure Q28-2 Figure Q28-
' 5b LYNXT Model Figure Figure Q28-2 Figure Q28-3 l
Hot Bundle Location in LYNXT Model Channel 1 Channel 17 l
MDNBR for Full V5H Core
[]
()
MDNBR for Hot V5H in S1C9 Core
()
[]
Configuration lyDNBR for ()
()
[]
l These results also show that the DNBR difference is greater when
[] is compared to a full core model, using-a design peaking distribution, than when the cycle-specific model is compared to a full core model, using an adjusted cycle-specific power distribution.
More importantly, for the Vantage SH, these results show that the DNBR response for the cycle specific loading pattern is very similar to the DNBR response for a full core of Vantage SH fuel.
Therefore, the method FCF has used to develop core peaking limits for the Vantage 5H fuel [] is conservative and applicable.
It should be noted that these. tables were included in the response to this question to specifically address the request to
" provide the result of the DNBR analysis using plant-and cycle-specific core loading configuration and the same limiting power distribution input in the above analyses.*
The tables were provided to emphasize the relative DNBR differences between the full core models, the limiting design models () and the cycle specific core loading models.
The DNBR's listed are based on core conditions at the nominal operating' point.
These tables are not intended to represent DNBR differences over the full range of
~
operating space allowed by the core des'ign.
Those DNBR differences, []
Details of those calculations are discussed in the response to Question 27.
The design radial peaks for the VANTAGE SH and Mark-BW are 1.62 and 1.70, respectively.
By maintaining the VANTAGE SH fuel at FCF Non-Proprietary
- _ _ -. - ~.
its current design peak, it is ensured that the resident fuel will retain its current safety margins and that the safety margins for the Mark-BW and VANTAGE SH are similar.
This difference i. design radial peaks is not required to support the absence of a VANTAGE SH fuel design transition core _ penalty, since the method of evaluating the VANTAGE SH performance in bounding core configurations yielded DNBR results that require'()
transition core penalty on the resident fuel.
However, the difference in design peaks adds additional assurance that the VANTAGE SH will operate at lower peak thermal power levels than 4
the Mark-BW fuel.
e The retained thermal margin for the Sequoyah DNB analysis resides j
in-the thermal design limit (TDL).
The TDL (1.50) is 10% higher than the 1.345 statistical design limit (SDL) using the BWCMV CHF correlation, Retained Margin in the Thermal Design Limit
()
Transition Core Penalty for the Mark-BW
[]
Bypass Correction (7.0% to 7.5%)
()
l
-Balance of Retained Thermal Margin for the Mark-BW
()
l t
Similarly, the VANTAGE SH fuel design has an equivalent []
retained thermal margin:
Retained Margin in the Thermal Design Limit
[]
Transition Core Penalty for the Vantage SH
[]
Rod Bow Penalty for the Vantage 5H
[]
Bypass Correction (7.0% to 7.5%)
()
Balance of Retained Thermal Margin for the Vantage 5H
[]
l 1.
- Letter, J. H. Taylor, Framatome Technologies to USNRC, Thermal-Hydraulic Methods for the Transition from Vantage SH to-Mark-BW Fuel at TVA's Sequoyah Plant, JHT/97-2, January 10, 1997 A
- e.
FCF Non-Proprietary
d Figure Q28-1
=
- LYNXT 12-Channel Model Hot Bundle 1/8th Core m meen
.,.,.g.q<)..o.oo.o 1
3
- --- -e =
- ***'OO O
.s OOOO
~~
' ~
',, oeoo T
-1
~
~
OOOO e-~
O.O O O o-
- O.,0 O Pm 13 e Remamder of Hot Bundle O* O i
- ** '
- a * "* ' a' 'a ca'-
i Charmel 11 e Remainder of Mad Buruse
,Oj c"**""'*****'""'c
menanau ?
l Table Q281 Pin By Pin Power Distribution For
- 12. Channel LYNXT DNB Model -
Re!ativo An Paske 5 -
Adark.RW hatinnbmiam Fumi Ner unal umrimum Nc...-i. ;
umvimum MCD NonMn Af'n hon Mf'n h1 1 Maa 1 700 1 man 1 R?O h2 1 AAA
-1M7 1MM
'MM Pin 't i Mad 1 Ra?
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Figure Q28-2 Design Power Distribution For Fuel Mix Scoping Study With Umiting Bundle at the Center Core Location (H-08)
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Figure Q28-4a Sequoyah 1 Cycle 9 Power Distribution Umiting Mark BW For Any Time in Cycle H
o V
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12 Lates Merk-GW s Locaton E00 992
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Figure Q28-4b Sequoyah 1 Cycle 9 Power Distribution Showing
- Limiting Mark BW Scaled to Design Bundle Peak H
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FCF Non-Proprietary.
i
.. Figure Q28 Sa.
Sequoyah 1 Cycle 9 Power Distribution Um! ting Wntage SH For Any Time in Cycle.
H o
F E
D C
8 A
i.
l 1 042
- ties t 2s3 1 257
- 1 246 1 108
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1266 1205 t142 1 026 205 Pomer Deentmhan Estracted Frorn to S1CS Feel Fuel Cyde Dougn at 0 F7PQ 1 265 L107 478 Landeg Wntage 5H at Locaton E10 1047 364 8
v 13 Figure Q28 5b Sequoyah 1 Cycle 9 Power Distribution Showing Umiting Vantage SH Scaled to Design Bundle Peak w
G F
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.797-056 8
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~13
..a FCF Non-Proprietary
29:
Provide-clarification of the limited use of Westinghouse' standard reinserts in a SON transition core application and provide justification that the transition penalty will bound the SQN application.
Response
- The last full batch of Westinghouse standard fuel assemblies was discharged Crom-the_Sequoyah cores in 1994.
Since that time, the Westinghouse standard fuel assemblies have been reinserted in limited numbers as an economic measure and as a means to reduce peaking next to adjacent fresh fuel.
Since the standard fuel assemblies available to TVA have undergone earlier irradiation, they can only be used in limited numbers because of their reduced energy potential.
For the first Mark-BW transition cycle at Sequoyah 1 there will be ten Westinghouse standards assemblies in core.
TVA expects to. continue using the Westinghouse standard fuel assemblies in this limited fashion for future cycles.
FCF has showr, based on hydrau',ic testing that the standard fuel assembly has a pressure drop (
) than aither the resident VANTAGE SM or Mark-BW design.
Therefore, the standard fuel
[
. assemblies will receive, on the average, slightly more flow than L
the other two designs in the mixed core.
Combining this effect
}-
with the fact that the standard fuel assemblies will be used in f-low power locations in the core clearly demonstrates that the standard fuel assemblies will have appreciably more thermal margin than either the VANTAGE SH or the Mark-BW.
As discussed in BAW-10178 (the transition core report for the Westinghouse standard to Mark-BW transition at the Trojan plant),
because the Westinghouse standard has (
) than the Mark-BW there is no need for a transition core penalty on the Westinghouse standard.
However, that report established a (1 penalty for the Mark-BW during that transition.
As outlined in the response to Question 26, that (1 penalty is bounded by the (1 penalty being imposed on the Mark-BW in the Sequoyah transition core.
9
~
1 i
FCF Non-Proprietary
--.+ _
4 3 0 :-
Provide the basesifor obtaining 2% of an increase in lift force for the limiting transition core configuration (one VANTAGE 5H in a Mark-BW core).
Also, provide the data for lateral crossflow velocities for the mixed core configuration and an acceptable criterion for lateral crossflow.
Response
The basis for the 2% increase in lift force is the' increase in-pressure drop seen when the total unrecoverable pressure drop for the single VANTAGE 5H assembly in the Mark-BW core is compared to the total unrecoverable pressure drop for the VANTAGE SH assembly in the same core location in the full-VANTAGE SH core, ()
As FCF demonstrated with detailed mixed core hydraulic evaluation results submitted to the NRC by letter dated 1/10/97 (1), the i
predominate hydraulic difference between the two assembly types is the pressure drop across the grids. ()
This flow diversion results in the small increase seen in the VANTAGE SH lift force 3-in the mixed core, Plots of crossflow velocities for the limiting mixed core configurations were submitted to the NRC in the 1/10/97 letter.
i Those plots showed that the maximum crossflow velocity in the mixed core is less than () and the corresponding span average crossflow velocity is less than (). The design. acceptance criterion for crossflow, to preclude any adverse flow induced vibration effects, is that the average crossflow velocity across an axial span between two spacer grids must be less than (),
Therefore, it can be seen that the lateral crossflow velocities for the mixed core configuration are well within the acceptance criterion,
[]
1.
- Letter, J. H. Taylor, Framatome Technologies to USNRC, Thermal-Hydraulic Methods for the Transition from Vantage.5H 1
to Mark-BW Fuel at TVA's Sequoyah Plant, JHT/97-2, January 10,-1997 i
J e
FCF Non-Proprietary 1
1
31:
Form loss coefficients for the fuel subcomponents were
-determined using measured pressure drops.
A LYNXT hydraulic model using those form loss coefficients showed that the total pressure drop of the Westinghouse VANTAGE SH design is approximately []- than that of the Mark-BW and the Westinghouse standard fuel assembly is approximately () in the pressure drop than the current Mark-BW.
Provide the detailed-analysis with respect to the overall impact on the mixed core DNBR analysis based on these () pressure drops.
Also, describe how the Figure 3.2 is generated and its application to the mixed core DNBR calculation if flow is much greater than 383,000 gpm.
s
Response
In BAW-10220P and the follow-up submittal of 1/10/97 (1), FCF reported that the Westinghouse standard fuel assembly is approximately [] in total pressure drop than the current Mark-BW.
While developing responses to these questions, it was determined
.that this () difference was in error.
Review of FCF's Westinghouse standard hydraulic analyses showed that the Westinghouse standard 'is only () in total pressure drop than the current Mark-BW.
However, discovery of this difference does not invalidate any of the transition analyses presented to date.
The following discussion summarizes the effects of the pressure drop differences that are present in the current transition core and also addresses similarities and differences between this transition core and the transition core at-the-Trojan plant documented in BAW-10178P.
T l
The VANTAGE SH fuel design exhibits about a () total fuel assembly pressure drop than the Mark-BW.
However, as was demonstrated through the detailed mixed core calculations j
submitted to the NRC by letter dated 1/10/97 (1), the mixed core DNBR prediction is influenced more by the axial distribution of the hydraulic differences than absolute difference in total assembly pressure drop.
Results from the hydraulic testing of the VANTAGE SH fuel design showed that [].
The pressure drop difference in the lower part of the assembly is further exaggerated by the first intermediate gridfon the Mark-BW, a non-mixing grid, being adjacent to a mixing grid at that location on the VANTAGE 5H.
However, as Figure 3-2 of BAW-10220 shows, as elevation increases, 1
L FCF Non-Proprietary
-()
The impact of these component hydraulic differences is illustrated by the hot channel mass velocity plots included with the 1/10/97 letter.
Those plots show that in the mixed core [].
(As-discussed in the response to questions 26 through 28, a
bounding () penalty has been applied to the first transition core).
In the response to question 27, the () transition penalty applied to the Mark-BW during the Trojan transition is shown to have been based on a () total pressure drop difference between the Westinghouse standard and the Mark-BW.
(Copies of pertinent pages from BAW-10178P are attached to the response to Question 27.)
However, as has been demonstrated for the Vantage SH assembly, the driving force behind DNB effects in the mixed core is not the total p essure drop difference but the grid pressure drop difference.
In the mixed core data package provided to the NRC by letter dated 1/10/97, it was shown that the pressure drop across the Westinghouse vantage 5H mixing grid is () than the pressure drop across the Mark-BW mixing grid and that the pressure drop across i
the Westinghouse standard mixing grid was [] than the pressure drop across the Mark-BW mixing grid at the time of the Trojan transition.
The mixed core studies that were performed for 1
Trojan showed that the [] Mark-BW to Westinghouse standard difference resulted in a [] DNB penalty.
For the Sequoyah mixed core studies, it has been shown that a () Vantage SH to Mark-Eii mixing grid pressure drop difference results in a [] DNB penalty.
Therefore, there is good agreement between the level of grid pressure drop difference, [], and the resulting DNB_ penalty, [].
As discussed above, the Westinghouse standard is.now shown to be
[] in total pressure drop than the current production Mark-BW assembly. () This [] -relative to the Trojan case is not significant since as stated in BAW-10220 and in the responses to Questions 26 and 29, in the Sequoyah core there will be only a.
limited number of Westinghouse standard fue} assemblies, and the hydraulic 4
FCF Non-Proprietary
_.~ _._.
characteristics of the mixed core are dominated by the Vantage SH.
Furthermore, FCF has shown that the model that was used to develop the () transition core penalty [] is conservative and that a penalty based on the actual core configuration would be much lower.
Therefore, the [] DNB penalty is adequate for the SQN 1 Cycle 9 transition.
Figure 3-2 was developed to demonstrate the pressure drop difference between the resident VANTAGE SH fuel design and the
+
Mark-BW design at nominal plant operating conditions (i.e. the best estimate flow rate).
The figure is generated by analyzing a homogenous core of each fuel type and plotting the corresponding total core pressure drop versus core height-for the two cases on l
a single figure.
Pressure drop conditions have been calculated for higher flow rates (greater than 383,000 gpm) for hydraulic lift analyses.
For those cases, the trend between the two fuel designs' axial pressure drop profiles remains the same but at elevated pressure drop magnitudes.
As presented, Figure 3-2 is not a direct reflection of conditions evaluated in.the transition DNB analyses, since those analyses are performed at the thermal design flow of 360,100,gpm.
However, the flow diversion trends discussed earlier in this response are valid regardless of flow 4
- rate, i
1.
Thermal-Hydraulic Methods for the Transition from Vantage SH to Mark-BW Fuel at TVA's Sequoyah Plant, JHT/97-2, January 10, 1997 i
4 FCF Non-Proprietary
32.
The horizontal seismic and LOCA loads were calculated for the mixed core for Mark-BW fuel and Westinghouse Standard fuel. Why was Westinghouse V 5H not used?
Additionally, explain how the mixed core calculation was performed to assure conservative results.
Resoonse Framatome Cogema Fuels (FCF) has calculated the horizontal seismic and LOCA loads for the mixed core for Mark-BW and Westinghouse Vantage 5H fuel assemblies as described on page 8-6 of BAW-1v220P, Rev. O. The results of the calculations were compared with the seismic and LOCA loads results of the full Mark BW core configuration. The resulting changes in spacer grid impact loads are minor [ ] as stated on page 8 6 of BAW-10220P, Rev 0. The spacer grid impact loads for all the faulted conditions are within the elastic load limit. Therefore, the requirement of a core coolable geometry is met.
The core configurations as indicated below and shown in Figure 32-1 were analyzed for a 5- -
assembly model. These conSgurations are based on providing combinations of E Vantage 5H and FCF Mark-BW fuel assemblies in mixed cores. The configurations are identical to those analyzed for the E Standard and FCF Mark-BW mixed core analysis (Figure 32-2).
V5H V5H V5H V5H V5H Base Case V5H V5H BW V5H V5H -
Core ConfI BW V5H V5H V5H BW Core Conf 1I 1
V5H BW V5H BW V5H Core ConfIII The all-V5H case is given because that establishes a base line loading for V5H assemblies. A goal of the mixed core analysis is to demonstrate the adequacy of the FCF fuel assemblies when mixed whh E resident fuel assemblies. It is also shown that the loads on E V5H and E Standard in a mixed core are equal or less than the all V5H or Standard baseline.
Based on previous calculations as discussed in Section 4.2 of B AW-10133P, Rev.1 (NRC approved topical), the five fuel assembly model gave the highest impact load, and as the number of FAs increased beyond five, the load decreased. For the Sequoyah plant, five FAs do not exist in any row of the core. The minimum number of fuel assemblies that exist in one row is seven.
However, for additional conservatism, the Sve-FA row was se,tected for the Sequoyah plant. As reported in Figure 5 of Enclosure 2 of B AW-10133P, Rev. I by using the five-fuel assembly core model, instead of the seven-fuel assembly model, resulted in the peak impact load higher by approximately ( )%.
As additional conservatism to achieve a maximum load on Mark-BW fuel assembly, the core 1
7 FCF Non-Proprietarf 4
1
configuration consisting ofE Vantage 5H, E standard, and Mark BW fuel assemblies as shown in Figure 32-3 was analyzed. This configuration is conservative because the Mark BW fuel assemblies are placed on the periphery as this location which has been shown in the previous cases to experience the maximum impact loads. Also, E Standard assemblies, which have Inconel intermediate spacer grids with a higher grid stiffness than the Zircaloy grids of the Mark-BW, are placed next to Mark-BW fuel assemblies, which increases the impact loading on the Mark-BW assemblies. The results of this analysis are discussed in the last paragraph of the response to Question 33.
4 i
l i-i i,.
i 4 -
)
i a
FCF Non-Proprietary i
?
-7
. ~..
Figure 32-1 Vantage SIUMark-BW Mixed Core Configuration FCF Non-Proprietary i
l 9
Figure 32-2 l
E Standard / Mark IlW Mixed Core ConDguration I
l 1-e I
l I
FCF Non Proprietary l
Figure 32-3 lY Standard / Vantage Sil/Slark llW 511xed Core Configuration
~
G I
FCF Non Proprietary 1
.\\
33, is a full core analysis performed for the combined LOCA/ Safe Shutdown Earthquake (SSE) loads for mixed core applications? Compare the critical loads (crushing Ivads) for the three types of fuel that will be in the reactor and describe how the results are effected by the differences.
Response
An analysis described in the response to above Question 32 demonstrates the adequacy of the FCF hiark BW fbel assemblies when mixed with.E Vantage SH fuel assemblies. For E Standard fuel assemblies, a mixed core bounding analysis as discussed on page 8 6 of B AW 10220P, Rev' O was performed for a mixed E and FCF fueled core. The resulting changes in spacer grid impact loads as discussed on page 8 7 of BAW 10220P, Rev. O are minor [ ] and well within the spacer grid elastic load limit (cmshing load). Hence, the requirement of a core coolable geometry is met for all combinations of Westinghouse (Vantage SH and Standard) and FCF hfark BW fuel assemblies The clastic load limit for the hfark BW zircaloy intermediate spacer grid at 600'F is ( ) Ibs. This valae was determined through testing. The buckling strength (cmshing strength) of the Westinghouse Vantage SH and Standard spacer grids were determined through a comparison of basic grid geometry and the use of test data. The calculated crusning strength of the Westinghouse Vantage SH and Standard spacer grids at 600'F are:
Pvm = ( ) lbs.
Psw = ( ) lbs.
An analysis described in Section 4.2.1.3.2 of Sequoyah FS AR demonstrates the adequacy of the E Vantage SH and E Standard fuel assemblies under seismic and LOCA loadings. For hiark-BW fuel assemblies, a seismic and LOCA analysis of a mixed core bounding configuration (hfark BW, W Standard, and E Vantage 5H) was performed, which showed that the hiark BW fuel assembly loads are within the clastic _ load limit. The grid maximum impact force for the seismic plus LOCA for this worst case configuration is [ ] lbs. Hence, the requirement of a core coolable geometry is met.
A base core configuration of all h! ark BW fuel assemblies was analyzed. The maximum impact force experienced by this full Afark-BW core was [ ] lbs. In comparison, the maximum grid impact force in the three assembly core configuration, shown in Figure 32 3 and described in the last paragraph of the response to question 32, is [ ] lbs. The presence of the Westinghouse fuel assemblies increased the maximum grid impact forces on the Afark-BW by [ ]%.
4 FCF Non Pruprietary
i i
1 34.
The methodology approved for trie mixed core structural analysis is contained in BAW-10133.
There is no reference to this mi:.hodology in the submittal.
Please verify that l
this methe4> logy was used.
Response
-A reference is made to the NRC approved FCF Mark-BW 17X17 fuel assembly topical report BAW-10172P. This topical report provides the structural evaluations specific to the Mark-BW fuel assembly.
lioweve r, this topical report refers to BAW-10133P, Rev.1. (tac approved) for the methodology used for the fuel assembly i
seismic and LOCA mechanical response analysis l
4 4
4 4
i FCF Non-Proprietary
m-35.
On p. 8-7 the stated design criteria (with a reference to the Standard Review Pian (SRP]) for the LOCA combined with the SSE does not include control rod insertabilityt however the SRP does require control rod insertability for this event.
Correct the criteria and verify that control rod insertability is maintained for the combined loads.
Response
The acceptance criteria of the Standard Review Plan (NUREG 0800),
Section 4.2, states that loads from the worst-case LOCA that requires control rod insertion must be combined with the SSE loads, and control rod insertabiljty must be demonstrated for I
that combined load.
The FCF design basis LOCA does not require control rod insertion.
Control rod insertion ~is only required for small break LOCA.
The FCF-supplied fuel assemblies will not inhibit control rod insertability for the design basis combined seismic and LOCA loads, as the spacer grid impact loads _for all the faulted conditions and core configurations are within the spacer grid crush load limit.
This assures that the FCF-supplied assemblies will not inhibit control cod insertability for small break LOCA, because the design basis combined seismic and LOCA loads'are substantially higher than those that would result from a small break LOCA.
FCF has conservatively complied with the SRP acceptance criteria.
e FCF Non Proprietay i
i l
36:
The submittal states that the target burnups for SON are 62,000 MWD /mtU for the peak rodi however, the safety evaluation for the Mark-BW fuel only approves the fuel up to burnups of 60,000 MWD /mtU for the peak rod.
Verify that the peak burnups will not exceed approved values and describe how each of the other limitations contained in the Safety Evaluation for SAW-10172 (Section 6.0 Conclusione) are met.
Response
Peak fuel assembly and fuel rod burnups will not exceed the approved values, as defined in BAW-10172 or in later documents (such as SAW-10186, currently under review).
Other limitations in the SER for BAW-10172 (Section 6.0 Conclusions) are as follows:
"Those licensees.that use the Mark-BW fuel design for relcad applications are required to submit the following plant-specific analyses: rod pressure, cladding collapse, DNB analysis, and fuel melting._
In addition, DNB analyses of mixed cores containing Mark-BW and Westinghouse fuel designs must be performed using an approved mixed core methodology."
The regtested plant-specific analyses have been performed with approved methods, as discussed in Sections 7.6 (Mixed Core Analysis) and 7.7 (Fuel Thermal Performance Analysis).
Detailed results of the plant-specific mixed-core DNB analyses have been provided by letter submitted to the NRC dated 1/10/97.
Fuel thermal performance analyses that predict fuel rod temperature and internal pressure conditions during core operation have been performed for the UO: rods with TACO 3 (BAW-10162 P - A) and the Gadolinia fuel rods with GDTACO (B AW-1018 4 P-A).
These analyses are used to determine the centerline melt limit and maximum fuel rod burnup limit based on the fuel rod internal pressure.
The TACO 3 code and internal gas pressure analysis methodology have been applied to fuel rod operation with pressure greater than reactor coolant system pressure as described in the NRC-approved topical report BAW-10183P-A.
'The results of the plant specific analyses are presented below for both the uranium dioxide and urania-gadolinia fuel rods.
The cladding collapse analysis that predicts fuel cladding creep collapse has been performed for the UO and gadolinia rods with CROV (BAW-10084P, rev. 3).
This analysis'is used to show that cladding creep collapse will not occur during the life of the fuel.
Fuel Rod Internal Pin Pressure The criteria for fuel rod internal pressure, stated in the safety evaluation of BAW-10183P-A, requires "the rod internal pressure for a limited number of rods will be permitted to exceed the RCS pressure, but will not exceed the smaller of the following the FCF Non-Proprietary
proprietary limit above RCS pressure (), or that pressure which would cause cladding liftoff to occur at s'.gnificant LHGRs.
The number of rods which will exceed the RCS pressure is limited by a core protection criterion which declares tbat the number of such rods, which also experience DNB in overpower transients, shall not exceed 0.01% of the rods in the core."
The predicted fuel rod internal pin pressure determined with the approved TACO 3 methods shows that both the UO and gadolinia fuel rod designs are acceptable to a burnup of () mwd /mtU.
l Maximum Mark-BW Predicted P
Fuel Rod Internal d/ t Design Pressure (psia)
U0
()
( l 3
Urania-( l'
()
Gadolinia Fuel Melt Limit The beginning-of-life (BOL) linear heat rate to melt limit bounds the time-in. life predictions.
The Melt Limit is calculated at various burnaps using the approved TACO 3 methodology.
The BOL limit, which bounds all of the calculated values, has been applied over the entire burnup range.
1 P
Fuel Rod Timit d/ tU)
(kW/ft)
UO:
()
BOL Urania-()
BOL Gadolinia Cladding Creep Collapse The cladding creep collapse griterion in.BAW-10084P-A, rev. 3 requires that " creep collapse will not occur during the life of the fuel."
Creep collapse of the UO and gadolinia fuel rods, calculnted with the approved CROV methods, will'not occur within s
a burnup of () mwd /mtU which bounds the design fuel red burnup 1
() in the Sequoyah core.
FCF Non-Proprietary
+
l t
37.
The latest TS change submittal indicates that TS X.2.3.d has a 48 hour5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br /> requirement to lower the OTAT limit.
Why is this a 48 hour5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br /> requirement and not a 12 hour1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> requirement as stated in BAW-10163.
Response
The TS required reduction of the OTAT limit requires an adjustment to the K term in the OTAT setpoint equation.
3 This adjustment involves an update to the SON Eagle 21 protection system parameters.
The_ update of the K i parameter requires the same actions as necessary_for the update of-the K._ and f (AI) terms in the OPAT equation.
The activities required to complete the update of these parameters in the four protection channels takes approximately 16 hours1.851852e-4 days <br />0.00444 hours <br />2.645503e-5 weeks <br />6.088e-6 months <br /> considering proper planning of the update and coordination with affected plant organizations.
An additional period of approximately 16 hours1.851852e-4 days <br />0.00444 hours <br />2.645503e-5 weeks <br />6.088e-6 months <br /> is allotted for testing to ensure that the update _has not adversely affected the accuracy of the oTAT or OPAT functions.
These activities. combined require approximately 32 hours3.703704e-4 days <br />0.00889 hours <br />5.291005e-5 weeks <br />1.2176e-5 months <br /> of the proposed 48 hours5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br />.
The remaining 16 hours1.851852e-4 days <br />0.00444 hours <br />2.645503e-5 weeks <br />6.088e-6 months <br /> will provide for contingencies that can not be foreseen.
30f the three terms proposed in the TS change, only the K.
term is-c:urrently required to be adjusted in a TS action requirement.
The time currently allowed for the cdjustment of this term in TS 3.2.1 Action a is 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />.. This action allows subsequent power operation past the 72-hour limit provided the K. term is properly adjusted to reduce the trip setpoint., The proposed TS change has reduced the allowed adjustment time for this term to 48 hours5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br /> consistent with the proposed times for the other two-terms.
This is based on 48_ hours being sufficient time to adjust these terms because the same activities apply.
Adjustment of the OPAT function in the= latest revision of the standard TS for Westinghouse Electric Corporation (NUREG 1431) continues to utilize a 72-hour action time.
The basis for this time is that it is sufficient considering the small likelihood'of a severe transient in this tin.e period and the associated actio:t to promptly reduce thermal power.
This basis is applicable to the OTAT function as well as the OPAT function.
~ The sample TS included in thd FCF topical report BAW 10163P-A was not intended to provide specific requirements for acceptable TS action times.
The 12 hour1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> provision for the adjustment of the above described OTAT and'OPAT terms in the sample TS were provided as an example of how the' action could-be applied. -It did not consider i
37 s
i i
plant activities:that are necessary to-implement such adjustments.
In addition, these times were not used to prescribe maximum times that are acceptable from a. safety
(
standpoint.
The change of these action times to 48 hours5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br /> does not invalidate the FCF analysis described in BAW 10163P-A because this analysis did not consider this time as part of the analysis and did not consider these-times to be a value determined in this analysis.
In conclusion,-the proposed TS change to use a 48-hour action requirement to adjust OTAT and OPAT functions is actually a reduction in this time when compared to similar current TS requirements.
This time duration is reasonable based on the time required to properly implement such-adjustments, a reasonable _ action time considering the potential for needing these functions for accident mitigation, and a similar standard TS requirements--that allow longer times to implement similar actions.
Therefore, the_ proposed action time of 48 hours5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br /> to implement
-adjustments to_the__oTAT and'OPAT_: functions istacceptable and does not impact nuclear safety.
e t
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k 38:
In reference to Section 8.2.2, provide allowable vertical loads for guide. thimbles, fuel rods, lower nozzle, and holddown spring.
It appears that the holddown spring load of () lbs exceea the design limit of () lba shown in Table 8-1.
Why the sesimic loads not included in the calculation of vertical loads?
Response
Guide Thimbles Two forms of necant formulas, one in terms of compressive stress
'j and one in terms of the lateral displacement are used to determine the allowable buckling load for the guide thimble.
Analyzed in this manner, guide thimble buckling is not a question of determining how long the column can remain straight and stable under an increasing lead, bitt rather how long the column can be permitted to bend under the increased load, if the allowable deflection or stress is not exceeded. The guide thimble buckling loads determined in this manner are-considered very conservative.
The allowable buckling ~ loads are span dependent.
The allowable loads for guxde thimble span buckling are as follows:
Load to produce maximum allowable compressive yield stress for
() lbs ZR-INC span Pit.31,
=
I
() lbs ZR-ZR span Pit....t.
=
Load to produce a midspan lateral displacement of () in, for
~
ZR-INC span Pit.31. = () lbs
[] lbs ZR-ZR span Pit..31.
=
The midspan lateral displacement of () in. is conservatively based on not affecting control rod insertion or trip performance.
The criteria for the guide thimble buckling allowables are discussed in Section 8.1.2.2 of BAW-10220P.
t Fuel Rods The. allowable vertical fuel rods load is, based on using Euler's formula. The calculated critical load i,s.[] lbs.
Lower Nozzle The maximum vertical impact force for the LOCA condition is ()
lbs.
The maximum impact force value on the lower nozzle for the LOCA condition reported in Table 8*4 of BAW-102220P FCF Non Propnemy
5 should have been () lbs.
Nevertheless, this force of () lbs is i
well below the impact forces observed in the fuel assembly drop i
tests, maximum of () lbs, in which no damage to the fuel assembly was observed.
l The lower core plato has been previously analyzed by Westinghouse l
up to a vertical load of () lbs without exceeding the allowable i
stresses for the material _as determined by the ASME Code.
As the vertical LOCA impact force of () lbs is under the load of () lbs, it is concluded that the lower core plate can withstand the vertical LOCA loads imposed on it.
Holddown Springs Three types of holddown margins are considered.
- % margin to FA lif tof f - (normal operation)
% margin to set (pump overspeed)
% margin to spring solid deflection (pump overspeed)
The. minimum holddown spring load at BOL, 100'F is () lbs. -End-of-life (EOL) holddown loads are different from-BOL due to increased spring-load from fuel assembly irradiation growth.
Holddown margins for beginning-of-life (BOL) conditions are provided since these provide the design limiting conditions for %
margin to fuel asssembly lif tof f (normal operation) and % margin to set (pump overspeed).
At 120%
pump overspeed condition, the fuel assembly will experience some liftoff.
The liftoff will be minimal and the holddown spring deflection will be less than the worst-case normal operating cold shutdown condition.
The holddown spring ultimate stress margin is not calculated, since the holddown spring reaches the solid deflection, before reaching the ultimate stress limit.
The total compression load required to bottom out the holddown springs against the upper nozzle at 650'F is () lbs.
This load is higher than the hol'ddown springs worst case LOCA load of () lbs.
The holddown springs do not fail due to reaching the solid-contact with the upper nozzle.
- Hence, during.a LOCA the fuel assembly does not contact the upper core plate, and a concern of the springs failure does not arise.
The margins listed below are for the best estimate flow rate and the mechanical design flow rate.
Evaluation Condition Calculated Margins Calculated Margins Best Estimate Flow Mechanical Design Flow Rate Rate 1.
Normal Operation
.BOL, 100'F
()
()
BOL, 14 0'F
()
[}
~
FCF Non Proprietary
l l.
Bob,500'F
()
()
BOL,100% P
[]
()
BOL, 118% P
()
()
1 2.
120% Pump overspeed (Margin to set, Margin to solid Deflection) t BOL,100% P
{}
()
[
BOL, 118% P
()
()
j vertical Seismic only the LOCA effect was anayzed in the vertical direction as the seismic excitation in this direction is not expected to cause fuel r
assembly liftoff, thus the vertical seismic loads will be the same as normal operating loada, i
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' FCF Non Proprietary -
~
. _. ~.. _ _ -
39:
A list of transients and the number of their loading cycles is provided in Table 8-2.
In reference to Section 8.1.2.1, provide maximum cumulative fatigue usage factors (CUFs) for the holddown spring and the clamp screw.
j
Response
The holddown spring and the clamp screw were analyzed for the total fatigue usage factor using the procedure outlined in the ASME Boiler and Pressure Vessel Code, section III.
The cumulative usage factor for the design loading cycles provided in Table 8-3 of BAW 10220P is determined to be () for the holddown rpring and () for the clamp screw.
Based on the code, the summation of the usage factors for all should not be greater than 1 for all system design transients.
Since this was shown to be the case, the holddown springs and the clamp screw will retain their structural integrity throughout the fuel assembly design life cycle.
4 i
a FCF Non Proprietary
40:
Discuss the effects of flow-induced vibration on the instrument guide thimble.
Response
Vibration tests were onducted in France in the HERMES T loop on two Mark-BW fuel assembly prototypes in a double assembly loop configuration.
The tests were performed at flow rates representative of reactor startup and normal operation conditions.
The vibration measurements focused on the assembly grids which provide the vibrational characteristics of the fuel assembly structure.
The level of vibrations measured in bending remain below () pm ([]
i in) regardless of the flow.
The integrated displacement is between () pm for the nominal flow.
The measured amplitudes were
- low-and no significant effect of the vibration on the fuel assembly was observed.
Since the instrument sheath and guide thim' 'es are part of the fuel assembly skeleton, the flow induced vac*ation behavior of the instrument sheath and guide thimbles can be a. ictly correlated to the measurement on the' fuel assembly structure.
The instrument sheath has the same lateral fit at each grid as the guide thimbles.
The outside diameter and the inside diameter of the 3
instrument sheath are identical to the guide thimbles except for a reduction in diameter below the dashpot-for the guide thimbles.
So, the instrument sheath and guide thimbles will experience the same lateral motion as the grids.-
For both the guide thimbles and the instrument sheath designs, each grid provides sufficient tube support to limit rod vibration and maintain cladding wear to within acceptable limits in the same manner.
This has been proven by () hour life & wear testing, in addition to proven in reactor experience.
Both testing and operating experience have shown fretting and rod vibration not to be a concern with the Mark-BW design.
V 4
FCF Non Proprietary
41:
The 1992 edition of ASME Code,Section III, Nuclear Power Plant Components, was used for the evaluation of fuel assembly.
Specify which subsection was used.
Is the Code used consistent with the code of record at Sequoyah?
If not, reconsile the differences and provide justification for use of the Code in accordance with 10CFR 50.55a.
Response
j The stresa intensity limits for most of the fuel assembly components for Condition I and II events (normal and upset conditions) are taken from the ASME Code Section III, Subsection NG.
The stress intensity limits for most of the fuel assembly L
components for Condition IV events (faulted conditions) are taken from the ASME Code Section III, Appendix F.
All the stress intensity limits used in the fuel assembly stress analysis are primarily obtained from the 1992 edition of the ASME Code.
These stress intensity limits remain the same as discussed in Section 4.2 (Mechanical Design) of the Sequoyah FSAR.
t i
4 FCF Non-Proprietary
r-42:
Provide methodology, damping values, input motions used in the sesimic analysis of the fuel assembly.
Response
The fuel assembly responses reaulting from seismic excitation and LOCA were analyzed using the general procedure outlined in FCF topical report BAW-10133P, Rev.
1.
The specific models used for the Mark-BW fuel assembly are provided in 3AW-10172P.
Both reports have received the NRC approval for referencing in licensing applications.
A reference is made to BAW-10172P in the BAW-10220P submittal.
Topical report BAW-10172P refers to BAW-10133P, Rev.1 for the methodology used for the fuel assembly seismic and LOCA mechanical response analysis.
A brief description of the horizontal model and methodology could be found
.in Section 4 1.2.1 of BAW-10172P.
~~
-The fuel assembly structural damping values determined experimentally for various-loading conditiens used are presented below.
The fuel assembly damping values were derived from fuel assembly tests in water which are described in Appendix A of BAW-10133P. Rev. 1.
These tests were conducted in still water and in flowing water for severhl initial displacement amplitudes, in which damping was determined from the free oscillation measurements. The fuel assembly structural damping values reported I
in Table 3-3 of BAW-10220P were obtained from tests performed in Y
air.
Damping values used in the SSE and LOCA analyses are conservative values for actual flow conditions.
These damping values for fuel assembly f aulted conditions are Faulted Condition Damping values
(% critical)
()
Values are from NRC-approved Topical BAW-10133P Rev 1 LOCA
(}
Table C-3 A fuel assembly structural damping value corresponds to a minimum
. operational rate expected dur.ing SSE and LOCA conditions.
The Safe Shutdown Earthquake (SSE) and Loss-of-Coolant Accident (LOCA) time history motions of the upper and lower core plates and the upper elevation of the upper core baffle plate were provided by Tennessee Valley Authority (TVA) for the.Sequoyah Nuclear Units 1 and 2.
These core plate motions were provided on floppy disks j
and the seismic input motions for two orthogonal directions, namely for the X and Z directions are illustrated in Figures 1 through 6 (TVA Proprietary).
Reference 1:
Letter - M. J, Lorek (TVA) to L. W. Newman (FCF).
S-013, dated October 19, 1994.
y FCF Non-Proprietary
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