ML20054L149

Forwards Response to NRC 820430 Request for Addl Info Re Specific Questions Which Will Be Incorporated Into PSAR Amend 69,scheduled for Submittal in Jul

ML20054L149
Person / Time
Site:	Clinch River
Issue date:	06/30/1982
From:	Longenecker J ENERGY, DEPT. OF, CLINCH RIVER BREEDER REACTOR PLANT
To:	Check P Office of Nuclear Reactor Regulation
References
NUDOCS 8207070193
Download: ML20054L149 (59)

Text

_.

Department of Energy Washington, D.C. 20545 Docket No. 50-537 HQ:S:82:061 JUN 3 01982 Mr. Paul S. Check, Director CRBR Program Office Office of Nuclear Reactor Regulation U.S. Nuclear Regulatory Commission Washington, D.C. 20555

Dear Mr. Check:

RESPONSES TO REQUEST FOR ADDITIONAL INFORMATION l

Reference:

Letter, P. S. Check to J. R. Longenecker, "CRBRP Request for

! Additional Information," dated April 30, 1982 This letter formally responds to your request for additional information contained in the reference letter.

Enclosed are responses to Questions CS760.ll, 21, 27, 32, 44, 46, 48, 51, 52, 55, 82, 90, 93, 94, 95, 97, 102, 106, 107, 108, 109, 114, 115, 117, 118, 120, 121, 122, 127, 129, 130, 132, 135, 139, and 140; which will also be incorporated into the PSAR Amendment 69 scheduled for submittal in July.

Sincerely, J n R. Longene er Acting Director, Office of the Clinch River Breeder Reactor Plant Project Office of Nuclear Energy Enclosures cc: Service List Standard Distribution Licensing Distribution

{0 8207070193 820630 PDR ADOCK 05000537 A PDR L -- _ _ _ _____________________ _ ----------__ - --.---

PIga 8 (82-0374) [8,22] #99 Ouestion CS760.11 Discuss how the Doppler coefficient is affected by changes made in going from the homogeneous to the heterogeneous core arrangement. Discuss the uncertain-ties of the Doppler coef ficient and how it impacts under cooling and reactivity insertion events.

Resoonse The following table summarizes typical CRBRP regionwise flooded Doppler constants in the heterogeneous and homogeneous core designs for clean-core conditions at the beginning of cycle one.

Doppler Constant, -T dk/dT x 10 4 Heteroceneous Core Homcaeneous Core fuel 25.8 43.0 (Inner core) 12.9 (outer core) inner blanket 44.0 ----

radial blanket 11.8 7.0 ax1al bIankets _L6 _ad TOTAL (isothermal) 84.2 67.3 The lower fuel Doppler in the heterogeneous core is a result of the smaller number of fuel assemblies (156 vs 198 in the honogeneous core) and their placement in the heterogeneous core, the higher fuel enrichment and lower fertile-to-fissile ratio in the fuel, and the harder average neutron energy spectrum. The higher total core (Isothermal) Doppler constant in the heterogeneous core is a result of the higher total (fuel plus blankets) heavy metal loading. The Doppler uncertainty, in the temperature range around hot-full-power, is 110% (Id) as discussed in Section 4.3.2.3.1 of the CRBRP PSAR.

PSAR Sections 15.1.4.1 and 15.1.4.2 provide analyses for the worst case overpower (Section 15.2) and undercooling (Section 15.3) events, respectively.

These include the results of a sensitivity study for both types of transients considering uncertainties in Doppler feedback (see Tables 15.1.4-1 and 15.1.4-7). The maximum and minimum values of Doppler coefficient used in these sensitivity studies are summarized below for the models described in Section 15.1.4 and Doppler information in Table 4.3-16:

EVENT MAXIMUM DOPPLER MINIMUM DOPPLER Undercooling -0.0118 -0.0019 Overpower -0.0035 -0.0019 As Indicated in Section 15.1.4 minimum Doppler is the worst case for the overpower event and maximum Doppler is the worst case for the undercooling event. The Doppler values used in the analyses appropriately reflect 1) a 30%

uncertainty (3a) in its calculated value and 2) time in core life effects.

This establishes the worst case conditions to be used in the nuclear kinetic /

1hermal-hydraulic evaluations for the fuel and blanket hot rod calculations.

QCS760.11-1 Amend. 69 July 1982

kge - 4 GGs-@330 G,F2] O@h Ouestion CS760 21 In Section 15.2.2.2.1 It is stated that gradual core radial motion in response to normal temperature changes is discussed in Section 4.2.2.4.1.8. However, thi s sed ion i s not i n the PS AR. LikewIse, the Iast three paragraphs on page 15.2-2a appear to be mispi aced.

Resnonse in PSAR Section 15.2.2.2 all ref erences to Section 4.2.2.4.1.8 have been changed to Section 4.2.2.4.3.

The latter part of this question, deal ing with page 15.2-2a, is addressed in the response to Question CS760.16.

QCS760.21 Amend. 69 July 1982 fM-@29$f4 __ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

Prga - 5 [82-0321] 8,22 #90 15.2.2.2 Sudden Core Radial Movement 15.2.2.2.1 Identification of Causes and Accident Descriotion The event to be considered here involves core radial motion which occurs rapidly and is difficult to accurately predict. This is in contrast to normal core radial motion which occurs gradually and predictably in response to normal temperature changes and Irradiation Induced material swelling and creep. The latter type event is discussed in Secton 4.2.2.4.3.

The type of sudden core radial motion to be evaluated has been termed " stick-slip" motion. Stick-slip motion ref ers to a situation in which the reactor assemblies are restrained from moving radially by interassembly frictional forces at the assembly load planes (stick) and then suddenly move to a new position dictated by current temperature and Irradiation environment as the interassembly frictional forces are suddenly removed or reduced (slip). If it is postulated that sticking occurs while the reactor assemblies are bowed away from the core centerline, a sudden positive reactivity insertion can take place as the assemblies slip to an inwardly bowed shape (towards the oore centerline). Such an event is unlikely since the bulldup of interassembly f ri ctional forces which would be roquired to cause sticking would occur only when the assemblies are in a compact inwardly bowed state. If the assemblies are bowed outward away from the core centerline, the interassembly gaps would be larger and then the probability of sticking would be minimal. On the other hand, if because of thermal and irradiation ef fects the assemblies due to manufacturing tolerances and frictional forces.

If the assemblies are prevented from achieving a compact state due to interassembly frictional ef f ects, it is possible that a seismic event could overcome the frictional ef fects and allow the reactor assemblies to take on a more compact state. This is considered to be the only realistic initiating mechanism for a stick-slip type event.

if the stick-slip event occurred, the reactivity insertion would cause temperature rises of the f uel, cladding, and coolant. The power rise would trigger a primary control system scram if the limits of Section 15.1.3 were exceeded.

15.2.2.2.2 Event Evaluation Model. Assumotions. and Conservatisms To determine the maximum possible reactivity insertion, the following analysis steps were followed:

1. Predict the dif ference in core assembly positions and bowing between l refueling and full power.

l

2. Determine the reactivity worth f actors associated with radial motion of each core assembly.

15.2-43 Amend. 69 July 1982 ,

1 -

7 Pege - 6 [82-0321] 8,22 #90

3. Frcm the predictions of maximum possible radial motion and worth f actors, determine an upper limit for possible reactivity insertion f rom stick-siip.

To predict the core assembly positions and bowing at ref ueling and f ull power conditions, a finite element model was constructed of a radial row of core assembl i es. The reactor environmental conditions were then applied along with material characteristics to give bowing and position curves like those of Fi gures 4.?-88 through 4.2-92. Ref er to Section 4.2.2.4.3 for f urther detail s l of the cors assembly bowing analysis. Comparison of the bowing shapes for Figures 4.2-88 through 4.2-92 shows an inward bowing at f ull power (100% power to flow ratio). The reactor assemblies were assumed to stick in the ref ueling posttion (at 0% power to fIow ratio) and to then siip suddeniy to the f ulI power position.

Conservative nominal compaction reactivity worth coef fIclents were determined by using the assumptions that alI control rods would be parked above the core at the beginning of an equil lbirum cycle. The worth coef f icients are shown in Tabl e 4.3-14.

The above procedure results in a prediction of approximately 60 for the maximum val ue of step reactivity insertion (see Section 4.2.2.4.3). l The above upper limit is considered to be conservative for the following reasons:

1. In the analysis, all the gaps in the core were compressed compleely out whereas core compaction tests (1) Indicate that not all gaps will be compressed out in a real core. This is due to manuf acturing tolerances as well as f rictional ef f ects in the core.

2. The analysis assumptions were that sticking of the core assembiles would occur where the assemblies are in their maximum outwardly bowed configuration. More realistically the sticking would not occur until

substanti al inwardly directed thermal bowing had already occurred and f orces had begun to buil d-up between assembl ies. Thus, part of the bowing reactivity change can be expected to occur gradually which will be compensated f or by Doppl er and thermal expansion ef f ects. This would reduce the maximum possible step reactivity change.

l 3. The inherent vibrational motion of the core assemblies when flow is

( passing through would tend to prevent sticking. This would aid in allowing smooth translation of the core assemblies in response to thermal bowing.

l l 1. W. C. Kinsel, "FTR Core Ccmpaction and W Ithdrawal Tests," May 1973, HEDL-TE-73-58, UC-7 9 e, g, h.

l 15.2-44 Amend. 69 July 1982

P gs 1 (W82-0416) #111 Question CS16Hi21 In CRBRP-ARD-0308, Feb.1982, a new model of the upper internals structure (UIS) is described verbally but no details are given. Please provide the applicable equations and estimate the effect of the new model on temperatures.

Resoonse The DEMO upper Internals structure (UIS) model is based upon a lumped nodalization scheme which represents the thermal-hydraulic characteristics of the collector and 29 chimneys comprising the UlS, as well as the collector radial gap. The nodalization scheme consists of 10 axial nodes and 3 radial nodes (2 metal, I sodium) for the chimneys and a single node for the collector. The UlS thermal-hydraulic nodeling is briefly described below o Thermal Nodel

1) Hest transfer through the chimney wall is assumed to occur b',

simple conduction.

2) Convective heat transfer between the sodium in the chimney and the chimney wall is calculated using the Lyon correlation for fully turbulent flow and heat transfer in pipes:

Nu = 7.0 + 0.025(Pe)0.8  ;

where Pe = Peclet number for 3% of rated chimney mass flow.

3) Convective heat transfer between the sodlum outside the chimney (within the shear web) and the chimney wall is calculated using the Maresca and Dwyer correlation for turbulent flow through unbaffled rod bundles:

Nu = 6.66 + 3.126(P/0) + 1.184(P/D)2 + 0.0155(TPe)0.86 where P = pitch D = chimney inside diameter Y=E/2 9 9 r

QCS760.27-1 Amend. 69 July 1982 o

Pag 2 2 (W82-0416) #111 o Hydraulics Model The pressure drop from the collector node to the reactor outlet nozzle through the upper path (chimney) and the lower path (radial gap) must be equal. Thus, ICHWCH WCH!+NCH + NPU " IGAPW GAPWGAP!*NPL where WCH = chimney flow WGAP = radial gap flow H = chimney natural head CH H = natural head for upper path from PU chimney to outlet nozzle Hp = natural head for lower path from collector through radial gap to outlet nozzle And GAP

• WCORE - NCH Solution of the above two equations provides values for chimney and radial gap flows. The U!S natural heads are calculated using upper plenum and chimney everage sodium temperatures.

Assumptions used in the development of the Ul5 model and their corresponding justifications are given below:

1) Sodium in the collector mixing chamber is assumed to be completely mixed and at a uniform temperature. Spatial deviations in collector temperature distribution existing at steady state conditions should be suf ficiently mitigated during the initial portion of the natural circulation transient to justify a one-node collector representation.

2) The convection heat transfer coef ficient between the sodium in the chimney and the chimney wall was assumed to remain constant with changes in sodium flow rate. The addition of a flow-dependent convection coefficient was found to have minimal effects on core flows and temperatures.

QCS760.27-2 Amend. 69 July 1982

@cg2 9 M @-@496D #199

3) The twenty-nine chimneys in the UlS are modeled as one chimney. Each chimney can be approximated as.two concentric cylinders which divide the flow path into Inner and outer regions. Chimney structure and configuration have been accounted for in both flow area and heat transfer area calculations.

4) Sodium on the inside and on the outside of the shear web region of the UlS is assumed to be well mixed and at a uniform temp'erature.

5) The loss coeffIclents for the UlS chimney and radial gap pressure drop correlations used in the analysis reported in CRBRP-ARD-0308 were determined from experimental results to be:

psid f

CM

= 2.047 x 10 -8 (Ibm /sec)2

-7 PsM f (f r a 1 inch gap)

GAP = 6.567 x 10 (Ibm /sec)2 Additional analyses using other gap heights revealed that the total reactor flow was not sensitive to the gap height.

Previous DEMO (Rev. 4) modelIng of the upper Internals structure lumped the above core structures (upper fuel assemblies and UlS) into one metal-sodium pair of nodes. The revised model described above provides a more detailed representation of the thermal-hydraulic characteristics of the UlS.

Comparison of natural circulation transient results using the two models shows that a lower minimum reactor cass flow occurs with the revised UlS model. A proliminary evaluation of the effect of the UlS model was made by adding The model to the simulation used in the analysis reported in CRBRP-ARD-0132 ("A Preliminary Evaluation of the CRBRP Natural Circulation Capability", November, 1977). The peak temperatures were seen to increase by approximately 70 F.

QCS760.27-3 Amend. 69 July 1982

Pogo I (82-0396) [8,22] #105 Question CS760.32 Light-water reactor experience Indicates that the reliability of auxiliary feed systems which are normally throttled is worse than auxiliary feed systems which initiate at full flow rates. Are there any safety-related reasons which preclude using full auxiliary feed flow at initiation? -

Resoonse The auxillary feedwater control valves are normally open. With SGAHRS initiation, the AFW isolation valves are opened and the AFW control valves are modulated in position to control the steam drum water level. The AFW flow to each steam drum depends on the water level of that drum.

There Is no apparent safety related reason to preciude fulI fIow at initiation. The diversity and redundancy in the CRBRP design results in a highly reliable design and as such the need for full flow auxiliary feed flow at initiation is considered unnecessary.

l l

l l

QCS760.32-1 July 1982

m

Pzg2 - 4 [82,0357] 8,22 #92 Ouestion CS760.44 in Section 15.5.2.1, it is noted that if an assembly does not freely drop into an open lattice position and the triple rotating plugs are subsequently operated, ".... additional severe damage can be inflicted to the assembly, to the adjacent assemblies, to the IVTM, and to the reactor upper internals." If such an incident does occur, what are the plans to assure that reactor operation does not commence with damages core assemblies or upper Internals structures in the vessel?

Resoonse The accident hypothesized in PSAR Section 15.5.2.1, is prevented from occurring by design. Interlocks are provided to prevent this event. IVTM grapple actuation to release a core assembly is prevented unless the assembly '

has been lowered to its setdown elevation (reference PSAR Section 9.1.4.4.2).

Plug rotation is prevented unless the IVTM grapple is fully withdrawn, and lowering the grapple is prevented if the plugs are rotating (reference PSAR Section 9.1.4.4.1 ). Furthermore, Control system logic operates without dependency upon these interlocks.

A specific procedure has not been outlined for recovery from this event. The likely approach to responding to this event is: refueling operations would be stopped immediately and the situation reported to plant management, who would '

assign responsibility for directing recovery operations. Also, the reactor vessel cover gas would be monitored immediately for detection of fission gases to determine !f cladding failure of any core assemblies had occurred. (Ref.

PSAR Sec. 9.8 for cover gas analysis and Sec. 7.5.4 for failed fuel monitoring). The refueling operations leading up to the event would be reviewed to determine the location and extent of possible contact inside the reactor vessel. Removeable core components might be inspected in the FHC If determined necessary. If Inspection of reactor Internals were necessary,

special inspection equipment would be obtained. The capability is provided for complete unloading of the reactor core and draining of reactor vessel sodium if necessary to facilitate inspection and repairs.

l QCS760.44-1 Amend. 69 July 1982 M ._.- . , - , _ ~ - - . - - - - __-

Pcg3 - 6 [82,0357] 8,22 #92 i

Ouestion CS760.46 in the evaluation of cover gas release during ref ueling, one cause of this event was identified as separation of the AHM from an open floor valve during a seismic event. During such a seismic event, additional fue.1 rod failures t

.may occur above the 1% level. Furthermore, the fission gas released by these  :

f ailed rods would not be processed by the RAPS bef~o re release to containment.

Given this sequence of events, what would be the ef fect on th'e of fsite doses as a function of the number of seismically falled fuel rods? Alternatively, demonstrate that the event sequence de. scribed above is so improbable as to beyond design basis. ,

Ecsponse The failure of fuel rods as the result of a seismic event during refueling is beyond the design basis. Refueling preparation and termination operations involving mating the AHM at the IVTM port (when the subject cover gas release event could occur) will be performed with the upper internals structure in its lowered position providing mechanical holddown of the core assemblies.

During operation and the portion of refueling with the AHM mated to the port, the reactor core is a compact unit and there is no mechanical damage to the fuel assemblies due to a seismic event.

Fuel pin failures from a seismic event during operation will be the result of a power transient initiated by a reactivity insartion caused by seismically induced control assembly movement. During ref ueling, the reactor core will be suf ficiently subcritical to prevent a power transient induced f uel pin failure.

I i

I QCS760.46-1 * '

Amend. 69 July 1982

Pags 1 (82-0397) [8,22] #106 i

Ouestion CS760.48 Section 15.6.1.5 contains analysis of p,stulated intermediate HTS pipe breaks and resulting sodium fires in the steam generator building. Please provide the detailed design Information regarding the high capacity veriting which is required to prevent overpressurization.

EDSDKG1D The design Information requested is provided in PSAR Section 6.2.7.

4 l

QCS760.48-1 Amend., 69 July 1982

Ouestion CS760.51 Assess the impacts of leak rates beyond the EBL in the Intermediate loop.

Additionally, here, the consequences of the sodium spray fire will be greater since the atmosphere is deinerted (i.e., 20% 0 2Instead of 2-3%). The time

. dependence of the leak rate itself is an imporTant factor i.i determining the course and effects of the spray fire. The analyses must substantiate that the consequences of the leak rate and spray fire are conservatively included.

Resoonse The Project has defined the IHTS Design Basis Leak (DBL) as that equivalent to the flow from a sharp-edged circular orifice whose area is equal to one-half the pipe diameter times one half the pipe wall thickness. This DBL is based on the results of Inserv!ce Inspection, pipe fabrication and installation quality assurance measures, fracture mechanics analyses and tests and leak detection provisions. These conditions show that a sudden large failure approaching the complete severence of an IHTS pipe is not credible.

As discussed in Section 15.6.1.5 of the PSAR, conservative assumptions have been used to maximize the ef fects from a IHTS Design Basis Leak spray fire. A leak was postulated to occur in the IHTS hot leg with the IHTS system operating at maximum normal operating temperature and pressure. Sodium discharge from the postulated leak continues at maximum flowrate, (m 1000 GPM) until the IHTS loop and pump tank has been drained (n/8.5 minutes) with subsequent plant shutdown on a plant protection primary-secondary flow mismatch signal. Subsequent to lHTS pump shutdown, sodium discharge continues as a result of static head driving pressure. This scenario results in the maximum challenge to the SGB Integrity from a postulated IHTS design basis leak. No action is assumed to be taken by the operator to mitigate the IHTS sodium leakage even though extensive leak detection Information would be available in the main control room to confirm the occurrence of significant sodium leakage in the Steam Generator Building (SGB).

In summary, the selected IHTS DBL is conservative and appropriate for assessing the capability of the SGB. No larger leak should be considered in the design and evaluation of the plant.

QCS760.51-1 Amend. 69 July 1982

4 ,

-- w-,,

Pcge 2 (82-0397) [8,22] #106 Ouestion CS760.5?

Section 15.7.1.2.1 states that the description of f ailure consequences of saf ety-related air supplies will be described only in the FSAR. Is this acoeptable? There is essentially no technical information provided in the present report. These are anticipated events. Please provide technical Information related to this section.

Resoonse The air supply system for the CRBRP is not a saf ety-related system although snple system redundancy and capacity is provided. This air supply system f urnishes compressed air to the plant systems. The system is described in Section 9.10. Systems requiring an air supply for their saf ety-related operations are provided with saf ety-related accumulators such that the f ailure of the compressed air system will not result in the loss of any safety f unction f or the duration required. Other saf ety-related val ves are designed to move in a preferred direction with the loss of air supply.

Section 15.7.1.2 has been revised to clarify the f ailure ef fects of the compressed ai r system.

4 l

QCS760.52-1 Amend. 69 July 1982 4e

P go 1 (82-0402) [8,15] #54 15.7.1.2 Loss of instrumentation or Valve Air 15.7.1.2.1 Identification of Causes and Accident Descriotion The system design precludes the loss of air supply to safety-related valves or instruments due to a single credible event. However, multiple f ailures, or a single f ailure occurring at the time of a design basis event, could cause loss of instrumentation or val ve ai r. Among such single f ailures are check valve mal function caused by val ve seal f ailure. Table 15.7.1 provides a listing of safety-related valves which requires a compressed air supply and their pref erred operating directions.

15.7.1.2.2 Analvsis of Effects and Consecuences The systems supplying compressed air to safety-related valves or Instruments will be designed such that a single credible f ailure will not cause Interrup-tion of the air supply. The Instrument air system is designed to supply clean, dry, and oil-free air for plant Instrumentation and control. The air receiver tanks are designed to the ASME Boller and Pressure Yessel Code, Section Vill, Division 1. Piping is designed to ANSI B31.1.0. Piping which penetrates the reactor contai nment wal ls, and the contai nment Isolation valves are ASME Section III, (Sections 3.9.2 and 6.2.4). Intercooler and after-coolers are designed to TEMA Class R.

All active saf ety-related, air operated valves will be designed to move in a pref erred direction with the loss of air supply. Tabl e 15.7.1.2-1 Identifies the saf ety-related valves requiring compressed air and the normal and f ailed positions and f unction perf ormed. Valves required to be operable for a saf e shutdown are equipped with saf ety-related accumulators. Each saf ety-related system is redundant.

There is no compressed air supplied to saf ety-related instrumentation such that the loss of compressed air would result in a loss of the instrumentation saf ety-related f unction.

15.7.1.2.3 Conclusions Based on the preceeding discussion, the compressed air system will be designed to prevent any adverse effects on the safe operation of the plant due to loss  ;

of instrument or valve air. '

15.7-4 Amend. 69 July 1982

Page 1 (82-0406) [8,19] #55 TABLE 15.7.1.2-1 ACTIVE SAFETY-RELATED VALVES OPERATED BY COMPRESSED AIR Normal Failed Position Val ve Operating After Loss of System Number Position Compressed air Function Primary Sodium Removal and HV001A Opened Closed Containment Isolation Decontamination System HV044A Opened Closed Contalrment Isolation (Nuclear Island General HV004B Opened Closed Contairment Isolation Purpose Maintenance System) HV085A Opened Closed Containment Isolation LN085B Opened Closed Containment isolation HV086B Opened Closed Contairment isolation i

Emergency Chilled Water NV353 Opened Failed Open System isolation NV354 Opened Failed Open System isolation NV400 Opened Falled Open System isolation NV401 Opened failed Open System Isolation NV403 Opened Failed Open System isolation NV404 Opened Failed Open System isolation NV409 Opened Failed Open System isolation NV410 Opened Falled Open System isolation NV141AC Opened Falled Open System isolation NV141AD Opened Failed Open System Isolation NV141BC Opened Failed Open System Isolation NV141BD Opened Failed Open System isolation 15.7-4a Amend. 69 July 1982

Page 2 (82-0406) L80 19] 855 Normal Falled Position Val ve Operating After Loss of Sy stm Number Position Compressed Air Function Emergency Chiiled Water (cont'd.)

A0V165 Opened Closed Containment isolation ADV166 Opened Closed Containment isolation ADV167 Opened Closed Containment Isolation A0V168 Opened Closed Contalrunent Isolation A0V211 Opened Closed Containment isolation 1

A0V212 Opened Closed Containment isolation A0V79 Opened Closed Contalrunent Isolatton A0V80 Opened Closed Contalrunent isolation ADV415 Opened Closed Contalnment IsoIation A0V418 Opened Closed Contalrunent Isolation Auxillary Liquid Metal System EVST Na Cooler Outlet NaK Loop 1 HV359' Open FaiI as Is Systom isoIatIon Loop 2 HV420' Open Fal1 as Is System Isolation EVST NaK Loop 1 Isolation HV3578 Closed FalI as Is Contalnment IsoIation EVST NaK Loop 1 Isolation HV358' Closed Fall as Is Contalrunent Isolation I EVST NaK Loop 2 Isolation HV415* Closed Fall as is Containment isolation EVST NaK Loop 2 Isolatf or. HV416* Cl osed Fall as is Contalrunent isolation

' Air stored In an accumulator for emergency operation of the valve.

15.7-4b Amend. 69 July 1982

rage 3 (ut-040t0 L#el>J 83)

Normal .Falled Position Val ve Operati ng After Loss of System Number Position Compressed Air Function inert Gas Receiving and RPHV001(1) Opened To RAPS Process Effluent Processing System RPHVOO2(1) Opened Closed Containment isolation RPUV015A(1) Opened Closed System Isolation RPUV015B(1) Opened Closed System Isolation RPUV018(1) Opened Closed System isolation RPUV019(1) Opened Closed System isolation APHV001(2) Opened Closed Containment Isolation APHV002(2) Opened Closed Containment isolation NGHV351 A(3) Opened Closed Contaltssent Isolation NGHV351 B(3) Opened Closed Contalrnent Isolation CGHV501(4) Opened Closed Contalrunent isolation GHV301(4) Openad Cl osed Containment isolation (1) See Figure 11.3-4; (2) See Figure 11.3-63 (3) See Figure 9.5-8; (4) See Figure 9.5-2 Evaporator Water Dump 53WDV001-004 Closed Closed System isolation Superheater Outlet 53SGV106-108 Closed Closed Relief (Power Operation)

Evaporator Outlet 53SGV100-103 Closed closed Rollef (Power Operation)

Steam Drum Outlet 535GV104-105 Closed Closed Ret lof (Power Operation) 15.7-4c Amend. 69 MlrLI G2

Page 4 (63-0406) Lo,15J 855 Normal Falled Position Val ve Operating After Loss of System Number Position Compressed Air Function Heating Ventilation and ARA 0V046A Opened Closed Containment isolation Air Conditioning System ARA 0V0468 Opened Closed Contalrment isolation ARA 0V046C Opened Closed Contalrment isolation ARA 0V047A Opened Closed Containment isolation ARA 0V047B Opened Closed Contaltment Isolation ARA 0V047C Opened Closed Containment Isolation ACA0V064A Opened Closed System isolation ACACOV064B Opened Closed System isolatica ACA0V122A Opened Closed System isolation ACA0V122B Opened Closed System Isolation ACA0V123 A Closed Open System isolation ACA0V1238 Closed Open System Isolation Floor Drain System ADV34 Opened Closed Contalrment Isolation A0V67 Opened Closed Contairment Isolation 15.7-4d Amend. 69 July 1982

P$ge3(82-0397)[8,22]#106 Ouestion CS760.55 in Section 15.7.1.4, Off-Normal Cover Gas Pressure, the relief valve setpoint pressure is 15 psig, while the elastomer seal system is also designed for 15 psig and it is stated that the dip seals of the reactor vessel ' closure would be " upset" at this pressure.

a. Why not design the seals for a higher pressure or change the relief valve setpoint to a lower value?

b. What type of valves are being used? What is the f ailure f requency (failure to close following a discharge)? Are they subject to common cause/ mode f ailures?

c. What are the consequences should the seals f all without any increase in the buffer gas flow rate?

Resoonse

a. PS AR Section 15.7.1.4.2 has been amended to clearly demonstrate that the ,

design pressure of the elastomer seals is 300 psid well above the relief val ve setpoint.

b. Two cover gas pressure relief valves are provided f or overpressure protection of the reactor and overflow vessels' gas spaces. These valves are located in Cell 107B on the equalization piping connecting the two vessels. The pressure relief valves are spring loaded saf ety type valves designed to ASME Section lil Class 2 requirements, and each valve is protected from the sodium vapor environment by a rupture disk located upstream of the val ve. Each pressure relief valve / rupture disk assembly is equipped with a normally open blocking valve located upstream of the assembly. The blocking valve is used to isolate each assembly from the reactor cover gas boundary during testing of the relief valve. It can be used to isolate the line in the event the pressure relief valve does not reset following a pressure relief operation. The back pressure on both pressure relief valves is ref erenced to the Cell 107B atmospheric pressure.

c. A bounding analysis of the consequences associated with seal fa!Iure has been provided in the response to Question 001.81 Incorporated into the PS AR in Amendment 2.

QCS760.55-1 Amend. 69 July 1982

P ge 2 (82-0402) [8,15] #54 15.7.1.4 Off-Normal Cover Gas Pressure In the Reactor Coolant Boundarv 15.7.1.4.1 Identification of Causes and Accident Descriotion As described in Section 9.5.1, the cover gas system serving the Reactor and Primary Heat Transport System maintains a pressure in the gas'spacefof the Reactor Coolant Boundary of 6" 2" of H 20. There is a constant sweep flow into the cover gas spaces and through the shaft seals of the primary pumps.

This in-ieakage is accommodated by two parallel pressure regulators in the Iine between the RAPS and the primary system overflow tank, which is maintained at the same pressure by a gas pressure equalization line connecting the pumps, reactor vessel, and overflow tank. The makeup regulators and the regulators controlling the bleed f rom the overflow tank to RAPS are both controlled from the same pressure signal. Failures of the pressure regulators (primary and redundant) or operator error, could cause deviation from the normal operating pressure of 6" i 2" W.G. l 15.7.1.4.2 Analvsls of Effects and Conseauences

a. Under pressure: If the pressure regulators (including redundant regulators) between the overflow tank and the RAPS system f all open, the pressure in the cover gas spaces within the Reactor Coolant Boundary will go sub-atmospheric since gas from the overflow tank will flow intotheRAPSvgcuumvessel. Since the vacuum vessel volume is approximately 300 ft (at 8 psia minimum) and the combined gas volume of the react 5 vessel, three primary pumps and the overflow tank is about 4500 ft , the reduction in pressure is modest: about 1 psi.

Such a reduction in pressure would have no adverse affect on the primary system. The change in NPSH available to the pump would not be signifIcant, and the seals in the reactor and pump closures would not be materially affected.

b. Over pressure: If the regulators between the overflow tank and the RAPS should close and the regulators control lIng fIow to the reactor vessel should, at the same time, f all open, the cover gas pressure in the reactcr coolant boundary would increase. Any potential problem is mitigated however by: 1) the time required to establish any significant overpressure, and 2) pressure relief dev ces on the overflow tank. Asmentionedabove,thevolumeofthgcovergasspace within the reactor coolant boundary is about 4500 ft at normal operating conditions. Since the gas makeup system wilI be designed to limit the makeup rate to about 50 SCFM, it would take at least an hour to double the cover gas pressure in the reactor coolant boundary. An annunciator in the control room will alert operators to take appropriate action (such as isolation of the makeup gas regulators) long before any appreciable overpressure will be realized. In addition, relief valves set to relief at 15 psig, will limit the pressure even if no operator action is taken prior to reaching this pressure. The discharge of cover gas from the rellef device wilI be modest and CAPS action will preclude any hazard to the public.. Even if the pressure does increase to 15 psig, there will be no af fect on reactor vessel level or pump tank level performance.

15.7-6 Amend. 69 July 1982 as-awe

P;ge 3 (82-0402) [8,153 #54 The pressure boundary margin seals will resist pressure in excess of 300 psid without failure, if the pressure in the reactor vessel should increase to 15 psig, these seals would remain intact. Cover gas would bubble through the dip seal and be trapped in the riser annulus between the dip seal and inflatable elastomer seals.

The primary system (and reactor vessel) design pressures have been established on the basis of a 15 psig cover gas pressure, and therefore the system, from a structural standpoint, is unaf fected by any overpressure which could occur.

If the primary system gas pressure should drift up due to one of the postulated f ailures, the 10 ps! minimum P between the Intermediate and primary sodium in the IHX would decrease; however, this loss of 6P would be monitored and annunciated and appropriate action would be taken.

15.7.1.4.3 Concigion Off-normal cover gas pressures in the Reactor Coolant Boundary will not cause a safety problem. Underpressure would be limited to approximately 1 psi below the normal operating pressure of 6 inches W.G. Overpressure conditions would l be limited to 15 psig by relief actions and would take about an hour to achieve. Even if such an overpressure condition were to exist, there will be no deleterious effect on the integrity of the Reactor Coolant Boundary. Since radiation dose rate builds up slowly and adequate radiation monitoring is provided, the radiation consequences would be small to the operating staff and are trivial to the public.

I 15.7-7 Amend. 69 July 1982 Mdbwi69)

Prge 25 W82-0320 [8,22] 59

~

Ouestion CSZfdLE2 One page 4.4-57, it is explained that the THDV conditions are more conserva-tive than the PEOC and therefore, represent the " worst bound" of plant conditions. On page 4.4-13, it is stated that a maximum sodium temperature of 1550 ~ 'er transient conditions provides an adequate margin to bolling. The Iimit assumes THDV conditions and a 750 F inlet temperature. For these conditions, Table 4.4-3 shows f uel and radial blanket temperatures of 1571 F*

and 1580 F, respectively, which are well in excess of the 1550 F limit.

Explain why those temperatures can exceed the prescribed Iimit?

• Note: Table 4.4-3 shows the maximum transient temperature in a f uel assembly as 1571 F f or assembly number 46. According to Figure 4.4-9, assembiy number 46 is an Internal blanket assembiy.

ReSMDSD We anphasize again that the value of 1550 F is not a limit. It is a conserva-tive guidel ine used to guide the core orif icing. The answer to this question is simil ar to the answer to the preceding Question CS760.81. Maximum temperatures were calcul ated f or three representative and l imiting assembl ies having i nitially assumed f l ows. A maximum steady-state temperature which corresponds to a 1550 F transient temperature was then cal cul ated. In fact, Tabl e 4.4-3 shows that f or the f uel and radial bl anket assembly, the steady-state temperature cat cui ated w Ith the assumed fI ow, while for the inner bl anket where the maximum transient calculated temperature was 1498 F, the steady-state temperature corresponding to 1550 F is higher. Thus, to satisfy the 1550 F guideline, the assumed assembly flow must be increased f or the f uel and radial blanket assembly but decreased in the inner blanket case.

Actually, the three assembiles considered were only representative worst cases and the intent was to establish criteria for the steady-state temperature so as not to exceed a transient temperature of 1550 F. These val ues were used in determining the TELTs for each assembly.

Regarding the note, Table 4.4-3 has been amended to remove this inconsistency.

QCS760.82-1 knend. 69 July 1982

m .

TABLE 4.4-3 .

i COOLANT LIMITING TEMPERATURES FOR TELT CALCULATIONS (TEPPERATURES IN *F)

STEADY STATE TEMP.

HETEROGENEOUS CORRESPONDING J0 STEADY STATE TEMP.

CORE MAXIMUM llLTER0GENLOUS CORE CORRESPONDING TO Typical Worst Case TRANSIENT TEMP. MAXIFUM TRANSIENT 1550*F MAXIMUM for Assembly Type (rpRE-2M CALCULATED) TEMP. (r0RE-2M) TRANSIENT TEMP. M __

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1331 1316 1252 First Core h Fuel Assembly 1571 1261 Second Core Inner Blanket Assembly 1498 1247 1282 1193 First Core 1207 Second Core 1331 1310 1232 Radial Blanket Assembly 1580

> t 51 t c-y Temperatures at TICV. 3a. 750*F Inlet Temperatures for Em

'< m PE0V. Zo i ."

P ga 1 (82-0399) [8,22] #108 Ouestion CS760.90 The major concern for the PHTS and the IHTS is in maintaining and assuring the integrity of the sodium piping. The main focus is on the definition of the size of the pipe break which must be considered as part of the design basis for the primary loop hot leg and Intermediate loop pipes. The applicant has an on-going offort in this area to substantiate the " leak before break" hypothests.

a. Sodium-to gas leak detection systems exhibiting adequate sensitivity for operating in this type of environment have been designed. However, please  ;

provide us with evidence that their long-term performance has been adequately demonstrated. Provide evidence ihat they have been tested on large systems such as will be used for CRBR. What data is being used to verify their rellabilIty?

b. Please provide us with the details regarding the pre- and in-service inspection program.

c. Please provide us with the details regarding the material surveillance program.

d. Since considerable effort is being made on Improved weld materials and welding techniques, please provide us with all long-term data on the CRBR weld compositions which demonstrate that adequate residual ductility is available during the lifetime of the plant.

Resoonse

a. Long-term performance testing of prototypical CRBRP Aerosol-Type Leak Detectors (sodium lonization & plugging filter aerosol) was initiated in April, 1977 at EBR-Il and is currently in progress. These detectors, on tests for over 5 years, have demonstrated the capability of those devices to meet CRBRP performance objectives.

The FFTF has a permanently installed sodium-to-gas leak detection system.

The performance of that system demonstrate reliable functioning under actual plant operating conditions.

Verification testing of the aerosol gas sampling leak detection system was performed with a prototypical section of CRBRP large diameter piping in both an air and an inerted environment. This was an insulated pipe 24 inches in diameter with a 8 foot long annulus and a 1 Inch gap between the pipe and the Insulation. Aerosol sniffer tubes and a collection manifold representative of the CRBRP design were used to transport the aerosols to the detectors during weeping leaks (100g/hr) at typical 01BRP operating temperatures. These leaks were detected welI within the time period which could cause significant corrosion damage to the piping.

Data obtained from the development tests (screening, characterization and optimization, mockup, natural circulation, and verification tests conducted at Al-ESG, the long-term performance tests conducted at EBR-Il and performance of the FFTF have sodium-to-gas leak detection system have been used in verifying the reliability of this equipment. An analysis of QCS760.90-1 Amend. 69 July 1982 m_arv.w

Page 2 (82-0399) [8,22] #108 the sodium-to-gas leak detection system to reliably detect small leaks in the PHTS has been performed. The assessment results are that the design provided a highly reliable system function.

b. The details of the pre- and in-service Inspection program are provided in PSAR, Appendix G.

c. The details of the material surveillance program are provided in W ARD-D-0185. (Reference 2 of PSAR Section 1.6).

d. Data on the effect of long-term exposure on weld compositions were evaluated and included in the CRBRP " Integrity of Primary and Intermediate Heat Transport System Piping and Containment" report, WARD-D-0185.

(Reference 2 of PSAR Section 1.6).

In this report, it was recognized that additional information would be required, and, consequently, a long-term thermal aging ef fects program on CRBRP prototypic welds was identified. This program is contained in Appendix C, Volume 2, of the referenced report.

QCS760.90-2 Amend. 69 July 1982

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Pags 3 (82-0399) [8,22] #108 Ouestion CS760.93 Given the presence of a small, undetected leak, the erosion of external surf aces of the piping by leaking sodium and its reaction products may be minimized, but not eliminated by the nitrogan/ oxygen inerting atmosphere.

However, while such an environment surround; the primary piping under normal operation, most of the Intermediate pipes are in a normal air atmosphere.*

How has this been accounted f or in detail for specification of the pipe break /Isak sizes?

Resnonse The Project is currently compiling additional Inf ormation regarding piping Integrity in the Intermediate heat transport piping system. The effects of ,

operation in a normal air environment are being addressed and have been 1 considered in the specification of the design basis leak for the intermediate Heat Transport piping. The Project will submit an amended response to this ,

question by October,1982.

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QCS760.93-1 Amend. 69 July 1982

Pcge 4 (82-0399) [8,22] #108 Ouestion CS760.94 Provide analyses which consider a spectrum of postulated pipe breaks of ,

dif ferent sizes (up to double-ended); a) consider critical locations in the primary hot leg (e.g., at the pump discharge, and in the intendediate loop piping and b) are run under varying initial and transient conditions to provide suf ficient assurance that the entire range of potential thermal /

hydreulic consequences to the system have been assessed, c) are based on '

analytical techniques and computer codes which are verified to conservatively bound the ef fects of such pipe breaks.

Resoonse The response to this question is contained in the response to Questions CS760.37 and 760.51.

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1 QCS760.94-1 Amend. 69 July 1982 l

Pzgm 1 (W82-0419) #114 Question CS760.95 The main steam isolation valves (superheater outlet isolation valves) play an important role in many potential events including station blackout and main steam lino break. Under what conditions are these valves closed? Which

. systems close them? What is the OSIS system referenced in Table 5.5-5 and under what circumstar.ces will it close the valve?

Resoonse There are two conditions which result in automatic closing of the superheater outlet isolation valve and superheater bypass valve for any single loop. The first is a large or intermediate-sized sodium water reaction pressure rellef system (SWRPRS) trip signal. The conditions that result in a SWRPRS trip are given Ir, PSAR Section 7.5.6. The second condition which results in automatic closure Of a superheater outlet isolation valve and superheater bypass valve is a low superheater outlet pressure (< 1100 psig). This function is identified in PSAR Table 5.5-5.

The outlet steam isolation subsystem (OSIS) closes the superheater outlet isolation valves and superheater bypass valves in all three loops as a result of either a high steam-to-feedwater flow ratio signal or a low steam drum level signal. The OSIS is discussed in PSAR Section 7.4.2.

PSAR Table 5.5-5, Section 7.4.2.1.3, and Figure 7.5-6 (Sheets 3, 4, 5, and 6) change pages are provided to clarify the OSIS function.

l l QCS760.95-1 l Amend. 69 July 1982 l

Paga - 1 (8,05) #49 TABLE 5.5-5 SGS PUMP AND VALVE DESCRIPTION ACTUATING PUMPS ACTIVE INACTIVE SIGNAL Recirculation Pump X N/A VALVES Pump Suction isolation X Manual (Remote)

Evaporator inlet isolation X SWRPRS Evaporator Inlet Water Dump X SWRPRS Evaporator Outlet Relief X SWRPRS**, High Pressure Evaporator Steam Drum Rellef X High Pressure - Steam Drum Superheater Inlet isolation X SWRPRS Superheater Relief X SWRPRS**,High Pressure Superheater Superheater Outlet Isolation X SWRPRS**, OSIS/SGAHRS or Low Super-heater Outlet Pressure Superheater Bypass Valve X SWRPRS**, OSIS/5GAHRS, or l

Low Super-heater Outlet Pressure Steam to SGAHRS HX X Manual (L.O.)*

Water from SGAHRS HX X Manual (L.O.)*

Steam to SGAHRS Auxiliary FW Pump X Manual Feedwater from SGAHRS X Manual (L.O.)*

Main Feedwater SGB lsolation X SWRPRS**, High Steam Drum Level, Low Steam Drum P~

r essure, Cell Temp and Humidity Main Feedwater Drum isolation X High Steam Drum Level Main Feedwater Check Valve X Simple Check Main Feedwater Control X High Steam Drum Level, Cell Temp and Humidity l Startup Feedwater Control X High Steam Drum Level, Cell

! Temp and Humidity l Evaporator Outlet Check Valve X Check Valve l Superheater Outlet Check Valve X Check Valve ,

SWRPRS**, SGAHRS Initi ation, Steam Drum Drain isolation X Low Steam Drum Pressure 1

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• • This function is not safety active l

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Amend. 69 July 1982

p;g3 1 W82-0425 (8,7) 27 7.4.2.1.2 Eauloment Design A high steam flow-to-feedwater flow ratio is indicative of a main steam supply leak down stream from the flow meter or insuf ficient feedwater flow. The superheater steam outlet valves and superheater bypass valves shall be closed with the appropriate signal supplied by the heat transport in'strumentation system (Section 7.5). This action will assure the isolation of any steam system leak common to all three loops and also provide protection against a major steam condenser leak during a steam bypass heat removal operation.

7.4.2.1.3 Initiating Circuits The OSIS is initiated by the SGAHRS initiation signal. The SGAHRS Initiation signal is described in 7.4.1.1.3. This initiation signal closes the superheater outlet isolation valves in all 3 loops when a high steam-to-feedwater flow ratio or a low steam drum level occurs in any loop. In each Steam Generator System loop, the three trip signals for high steam-to-feedwater flow ratio and the low steam drum level are input to a two of three logic network. If two of three trip signals occur in any of the 3 loops, the OSIS is initiated, and all 3 loops are isolated from the main superheated steam system by closure of the superheater outlet isolation valves and superheater bypass valves.

7.4.2.1.4 Bvoasses and Interlocks Control interlocks and operator overrides associated with the operation of the superheater outlet isolation valves have not been completely defined.

Bypass of OSIS may be required to allow use of the main steam bypass and condenser for reactor heat removal. In case the OSIS is initiated by a leak in the feedwater supply system, the operator may decide to override the closure of certain superheater outlet isolation valves.

7.4.2.1.5 Redundancy and Diversity j Redundancy is provided within the initiating circuits of OSIS. The primary l

trip f unction takes place when a high steam-to-feedwater flow ratio is sensed by two of three redundant subsystems on any one SGS loop. The low steam drum i

7.4-7 Amend. 69 July 1982

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Paga 5 (82-0399) [8,22] #108 Ouestfon CS760.97 What, if any, protection or control systems are capable of detecting a recirculation pump trip directly?

Resnonse The Plant Protection System does not detect a recirculation pump trip directly. However, a reactor trip is initiated via PPS when a recirculation pump trips f rom High Evaporator Outlet Temperature.

The recirculation pump trip can be detected and verified at the following locations:

Discharge pressure indicator located on local SGS panel. <

- Suction pressure Indicator located on local SGS and main control panels.

- Dif ferential pressure Indicator located on local SGS and main control panels.

Grodp alarm for low dif ferential pressure located in the main control room.

QCS760.97-1 Amend. 69 July 1982

Pag) 1 (82-0401) [8,22] #110 Ouestion CS760.102 in Section 5.5.2.3.4 (Steam Generator Module), the presentation on accident analysis takes credit f or improved methods of welding the tube to the tube-sh eet. However, the PSAR Indicates that this weld is in a developmental stage.

If the weld method is important to saf ety (e.g. failure f requency), please provide details of the method and any supporting evidence that Indicates its superiority over previous methods.

Resoonse The weld method employed in the tube-to-tubesheet welds of the CRBRP Steam Generators is an in-bore butt weld. It was selected to avoid the crevices which exist if a f ront f ace fillet weld would be used. The weld method as well as the welding equipment has been utilized bef ore and as such were not the subject of the development program. The development program was aimed at the Improvement of the weld quality and dependable repeatability of the process. The measures taken to this end are described in Section 5.5.2.3.4 of the PS AR as fol lows:

For the steam generator tube-to-tubesheet welds, the ASME Code requirements (NB-4000 and NB-5000) were supplemented by requirements of RDT E15-2 and additional requi rements. Requirements imposed on the tube-to-tubesheet welds above those of the Code include:

o Vacuum-Art Remelt or Electroslag Remelt - material is specified to

, reduce impurities and improve properites f or tubesheet forgings and tubes.

o Post weld heat treatment range defined to optimize resistance to caustic stress corrosion cracking.

o Hel lwn leak test.

o Penetrant test requirement limiting def ect size to much less than that of the Code.

o Weld geometry requirement limiting concavity, convexity and wall th i nni ng, o Micro-focus radiographic examination - developed to radiograph tube-to-tubesheet welds with improved resolution.

All of the above measure were taken to assure high quality welds. The actual

" weld development" is the weld procedure development required to quality the procedure, equipment and personnel as required by the Code.

l QCS760.102-1 Amend. 69 July 1982

Pags 2 (82-0401) [8,22] #110 in addition, the following ef forts were undertaken to improve upon available commerical quality standards to achieve the highest quality, dependable welds obtainable:

4

1. The tube-to-tubesheet preliminary weld development ef forts ' covered work on

. CRBRP steam generator tube-to-tubesheet welding up to the beginning of wel d qualification. This included laboratory weld development, the check-out and verification of the process under manuf acturing conditions, a statistical evaluation of the process to establish acceptance criteria, the associated quality assurance procedures and the development and procurement of appropriate welding power supplies.

2. Definition of tight weld geometry acceptance criteria.

3. Post-wel d heat treatment thermal stress eval uation.

4. Investigation to determine the likelihood of cracking of the tube-to-tubesheet welds during PWHT.

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QCS760.102-2 Amend. 69 l - - - - - - _ _ _ _ . __ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _

Pcgs 1 (82-0405) [8,5] #46 autogeneous buttwelded, tube-to-tubesheet joints. The shell and tube material is 2-1/4 Cr-1 Mo steel. There are two evaporator modules and one superheater module per loop. The evaporator modules operate with a recirculating ratio of 2:1.

.The steam generator design requires that each loop (two evaporators and one superheater) develop 325 MWT at rated full load. The design life of the steam generator module Is 30 years with items that cannot be reason ~ ably expected to last 30 years being replaceable during in-service inspection periods. The steam generator is designed to withstand the normal, upset, emergency and f aulted operating conditions in accordance with CRBRP Criterion 26. The steam generator module is also designed to withstand the loading combinations indicated in Section 3.9.2.2.

The materials used in design of the steam generator module major components are as follow:

Pressure Boundarv (2-1/4 CR-1 Mo-Ref. 3. Vol . 1. Section 2.2):

Shell Forgings - SA 336, Class F22A Shell Plate -

SA 387, Grade 22, Class 1 Tubesheet Forgings -

RDT M2-19 with optional provision 2 and Code Case 1557-2 Tubing - RDT M3-33 as modified to limit silicon and carbon content for weldability and carburization considerations (Reference 5).

I The steam generator module supplier will provide procedures for welding and heat treating in accordance with the requirements specified in the Code as modified by RDT E15-2NB (see Section 5.5.1.2). Welding qualification is l controlled by the Code as modified by RDT F6-5 and RDT E15-2NB (see Section 5.5.1.2).

The weld method employed in the tube-to-tubesheet welds of the CRBRP Steam Generators is an in-bore butt weld. It was selected to avoid the crevices which exist if a front face fillet weld would be used. The weld method as well as the welding equipment has been utilized before. A development program was aimed at the Improvement of the weld quality and dependable repeatability of the process. The measures taken to this end are described as follows:

For the steam generator tube-to-tubesheet welds, the ASME Code requirements (NB-4000 and NB-5000) were supplemented by requirements of-RDT E15-2 and additional requirements. Requirements imposed on the tube-to-tubesheet welds above those of the Code include:

o Vacuum-Arc Remelt or Electroslag Remelt - material is specified to reduce impurities and improve properties for tubesheet forgings and tubes, o Post wold heat treatment range defined to optimize resistance to caustic stress corrosion cracking.

o Helium leak test.

5.5-9a Amend. 69 July 1982

Paga 2 (82-0405) [8,5] #46 o Penetrant test requirement limiting defect size to much less than that of the Code.

o Weld geometry requirement limiting concavity, convexity and wall thinning.

o Micro-focus radiographic examination - developed to radiograph tube-to-tubesheet welds with improved resolution.

Material Integrity prior to placing the steam generators in service will be assured by complying with the ASME Code Section lli which requires weld radiography, tubing ultrasonic testing, plate ultrasonic testing, tubing hydraulic testing, component pressure testing and helium leak testing.

Material considerations are Indicated in Sections 5.5.1.4 and 5.5.3.11.

Section 5.5.3.1.5 Indicates the tests being conducted to support the steam generator design. It is not anticipated that back-up materials will be required.

5.5-9b Amend. 69 July 1982

Pag) 2 (82-0358) [8,22] #93 Ouestion CS760.106 in the event of pipe breaks, what would cause the various isolation valves to close? It is stated in Section 5.6.1.2.1 that in at least one such break it is necessary for the operator to close an Isolation valve to save the Inven-tory of the PWST. How many such postulated breaks required operator interven-tion? How does the operator determine the break location?

Resoonse in the event of a large pipe break, in a steam generator loop or in an AFW loop downstream of the AFW check valves, the AFW isolation valves to the affected SGS loop will automatically close following the steam drum depressurization to <200 psig. All postulated breaks that do not allow steam drum depressurization will require operator action to isolate the AFWS If the breaks are large enough to initiate SGAHRS.

The Information available to the operator to isolate these pipe breaks is described in PSAR Section 5.6.1.2.1.1 " identification of Active and Passive Components which inhibit Leaks".

l QCS760.106-1 Amend. 69 July 1982

P gs 1 (82-0414) [8,05] #47 5.6.1.2 Design Descriotion 5.6.1.2.1 Design Methods and Procedures 5.6.1.2.1.1 Identification of Active and Passive Comoonents'which Inhibit Leaks The equipment of the SGAHRS is shown schematically in Figure 5.1-5. Valves and pumps with*n the SGAHRS are classified as active or inactive, and the'r operating mode is given in Tables 5.6-5 and 6.

In the event of a pipe break in the auxiliary feedwater portion of SGAHRS, continued heat removal capability will be assured by the multiple loop feature of the SGAHRS and heat transport system.

If a large pipe break occurs in any portion of a steam generator loop, this will result in a reactor shutdown and an AFW Initiation. Automatic isolation of the AFW supply to the ef fected loop will occur within approximately 2 minutes when the steam drum pressure f alls below 200 psig. Operator action as a backup is available.

If af ter an AFW initiating event, a pipe break were to occur in the auxillary feedwater piping between the steam drum and the isolation valves immediately downstream of the control valves, the flow in the effective loop will increase until limited by the control valve (at approximately 110% of rated flow). A flow limit alarm in the control room will alert the operator to the f act that corrective action is necessary. Following the control valve flow limit al arm, the operator verifles a leak f rom Information provided by the following Instrumentation:

a) Saf ety-related steam drum level and pressure Indication are provided on each loop to assist in making a break determination. An inability to recover level or maintain pressure on any steam drum with a corresponding flow limiting alarm on AFW provides a break Indication.

l b) The Steam Generator Building (SGB) Flooding Protection Subsystem l annunciates abnormal SGB temperature, humidity, and sump level in the control room to alert the operator to pipe breaks that could compromise SGAHRS operation (see Section 7.6.5).

c) The plant trip signal: high or low steam to main feedwater flow ratio or low steam drum level. A trip of this type will direct the operator's l

attention to the steam / water-side of the plant.

Operator action in the control room will close two AFW supply Isolation valves to isolate the defective loop. In addition to the above, automatic isolation will occur when the AFW flow remains above 150% for 5 sec. (Indicating a flow limiter failure). Due to the flow limiting capalblity of the control valves, the leakage flow will be minimized and proper flow to the two remaining steam drums will continue even though one loop has suffered a pipe break.

5.6-4 Amend. 69 Ju!y 1982

Pcge 3 (82-0358) [8,22] #93 Ouestion CS760.107 How soon must the operator intervene and under what conditions to avoid necessary depletion of the PWST inventory?

Resoonse The conditions under which the operator must intervene in the event of pipe breaks are described in the response to Question CS760.106. The time available for operator Intervention to isolate a pipe break is discussed in PSAR Section 5.6.1.2.1.1.

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i QCS760,107-1 i Amend. 69 July 1982 -- _

.. o,o, Prg) 3 (82-0401) [8,22] #110 Ouestion CS760.108 Does the operator have any way of knowing the water levels in steam drums, in the PACC loops?

Resoonse The main control room operator can continuously monitor the steam drum water level in each loop en main control panel Indicators. In addition, steam drum level Indication is available in the main control room on the PDH&DS. Steam drum level Indication also is provided on the three SGAHRS panels at the remote shutdown station in the SGB Intermediate bay. The signals to the Indicators mentioned above are generated by three, independent, saf ety-related dif ferential pressure detectors connected to each steam drum. In addition, each steam drum is equipped with a multiport sight glass which can be observed locally.

There are no level detectors on the PACC loops. However, the PACC common condensate return line contains a venturl flow meter the output of which is indicated on the main control panel, PDH&DS and the SGAHRS panels. During normal plant operation, the water level in the PACC condensate return piping essentially :s the same as the water level in the steam drum. During PACC operation, water level in the PACC condensate return lines will be higher due to the flow induced pressure drop in the lines. The height of the water level will depend on the heat being removed by the PACC and, therefore, the condensate flowrate back to the recirculation pump suction as described in PS AR Secti on 5.6.1.3.2. The PACC is being designed to ensure that the condensate flow from the tube bundles will be stable. Theref ore, condensate flow measurement is adequate indication of proper PACC operation and no PACC loop level measurement is required.

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QCS760.108-1 Amend. 69 July 1982

Pagt 4 (82-0358) [8,22] #93

. Question Cs760.109 More information is required concerning the turbine drive for the turbine driven auxillary feedwater pump. How will pump performance depend on the pressure and temperature of the steam driving the turbine?

'Resoonse The SGAHRS turbine drive is designed to produce a rated 2000 hp at 4000 rpm when supplied with a steam at 1000 psig, 546 F and 96% quality at turbine inlet. The required turbine inlet conditions are regulated and maintained by a pressure control valve which reduces the steam pressure from the steam drum operating pressure (1475 psig during superheater venting,1550 psig during steam drum venting) to 1000 psig at turbine Inlet.

Normally, following SGAHRS venting, the SGS pressure will be 1400 psig as controlled by the PACC heat rejection rate. When the PACC assumes the total heat load (~1 hour following SGAHRS Initiation) the need for feedwater is limited to makeup leakages. Under the ref ueling conditions the pressure will drop as the SGS temperature is reduced to 400 F. At refueling conditions the makeup of water leakage is provided by the main feedwater system, in the event the main feedwater system becomes unavailable, the motor driven auxiliary feedwater (AFW) pumps provide a redundant alternate source of makeup water. Each of the two motor driven AFW pumps has f ulI capability of providing makeup water for these conditions. There is no need to operate the turbine driven auxiliary feedwater pump at steam drum presures below 1000 psig.

QCS760.109-1 Amend. 69 July 1982

Page - 2 (8,22) #104 Ouestf on CS760.114 Please provide direct heat removal service design details including:

a. Air blast heat exchanger and overflow heat exchanger design .

elevations.

b. Length of the piping

c. Design temperatures at Inlet and outlet locations of the heat exchangers.

Resoonse a) Design elevations:

1) Air Blast Heat Exchanger (ABHX) centerline of the Nak outlet line is elevation 777'-0 5/8".

2) Overflow Heat Exchanger (OHX) centerline of the sodium outlet nozzle is elevation 757' - 10 1/4".

b) The length of the Direct Heat Removal System piping can be determined

per Ref. CS760.114-1.

c) Design temperatures:

1) Structural design temperature for the ABHX is 650 0F at 100 psig internal pressure.

2) Structural design temperature for the OHX is 650 F at 100 psig Internal pressure.

Ref. 760.114-1 :

Letter HQ:S:82:29, J. R. Longenecker to P. S. Check, dated June 25, 1982.

QCS760.114-1 Amend. 69 July 1982

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p0ge 2 W82-0298 (8,22) 43 Ouestion CS760.115 The auxiliary feedwater system is an Integral part of the decay heat removal system, yet for some transients (e.g., steam generator tube leaks) It is necessary to shut down the auxillary feedwater supply.

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a Does the requirement to be able to isolate the auxiliary feedwater system Impair the availability when needed?

b. Can a single failure or spurious isolation signal lead to a simultaneous isolation of all auxillary feedwater trains?

Rosconse

a. The design of the AFW subsystem of the SGAHRS is such that isolation of AFW to one loop of the Steam Generator System will not impair the availability of the AFW to the other two loops. The AFW subsystem is capable of removing plant decay and sensible heat under all conditions in which at least one heat transfer loop remains intact. Isolation of one AFW loop will, theref ore, not impair safe shutdown of the plant.

b. The AFW subsystem design provides separate piping, valving,and controls for each steam generator loop. This arrangement ensures separation of each loop in perf orming its f unction. As such a single failure or spurious isolation signal will not lead to simultaneous isolation of all AFW trains.

QCS760.115-1 Amend. 69 July 1982

Page - 4 (8,22) #104 I

Ouestion CS760.117 in the PSAR the presentation of the Direct Heat Removal Service (DHRS) does not contain enough detailed inf ormation f or analysis at this time. The DHRS decay heat removal capabilities cannot be adequately assessed but on the basis of .Ilmited inf ormation available some very preliminary simplified analysis has been conducted. The DHRS design is such that the sodium Intake nozzle only penetrates the reactor vessel wall (not the thermal liner). The sodlum flows into the intake nozzle partly from tha hot upper plenum through the thermal i j liner ports (2.625 inches below normal operating sodium level) and partly from the bypass flow from the cold lower plenum. The DHRS return nozzle penetrates  ;

Into the hot upper plenum where the returning cold sodium mixes with the hot sodium and depends on the primary pony motors f or force the sodium flow l through the primary loop. -

Our concerns regarding DHRS to loss of heat sink (LOHS) events includes

a. The DHRS Intake nozzle doesn't penetrate into the upper plenum. The sodium level f al ls below the thermal liner ports f or an extended period of time, how would level be recovered and if it cannot be recovered what are the consequences?

b. From our DHRS preliminary studies, the tanperature of the sodium returning to the upper plenum from the DHRS is about 6000K, and during l loss of normal heat sink events, the sodium temperature exiting f rom the core is about 900 0K. What is the thermal mixing in the upper plenum? Is there any thermal stress concentration? Is there any flow reversal ?

c. Please provide the details of the 1/21 scale plenum tests which are mentioned i n Q001.580.

f Resnonse l

a. There is no credible combination of events which would result in ,

sodium level f alling below the thermal liner ports f or an extended period while DHRS is operating to remove decay heat; however, DHRS is  !

designed with the following makeup capability in the event that overflow has been Interrupted f or any reason: Prior to reactor scram, the overflow vessel is one-hal f ful l (approximately 17,000 gal lons).

The makeup pumps continue to transfer sodium back to the reactor.

Pump flow is increased to 560 gpm total (280 gpm per pump) upon initiation of DHRS. The 17,000 gallons of sodlum in the overflow tank Is equivalent to about 89 inches of available makeup level in the reactor.

b. The fluid mixing in the upper plenum during DHRS operation has been determined experimentally in the 0.248 scale Integral Reactor Flow Model (IRFM). Richardson number modeling was used to simulate CRBRP shut-down conditions for both three-loop and one-loop flow. The test results Indicate that the plenum is well mixed with only 6% of the DHRS flow short-circuiting the primary loops during three-loop QCS860.117-1 Amend. 69 July 1982

us uso, Pcge - 5 (8,22) #104 operation and 10-11% short-circuiting during one-loop operation. The DHRS design allowable is 20%. A second Indication of the fluid mixing is f rom the smalI magnitude of the thermal striping measurements.

The highest peak-to-peak f luctuation, measured during three-loop testing, was 18% of the dif ference between the core exit temperature and the DHRS return nozzle (sodium makeup nozzle) temperature.

Because of the fluid mixing in the plenum the only area with a potential thermal stress concentration is at the sodium makeup nozzle.

There will not be any flow reversal through the core with any primary loop pony moters operating. If one or two of the loops are Inoperative there could be reverse flow, without safety consequence, through those loops.

c. In the response to Q001.580 it was stated that the adequacy of the geometric locations of the DHRS makeup and overflow nozzles was demonstrated in a 1/21 scale Outlet Plenum Feature Model Test, and that this adequacy would be confirmed in f uture testing in the 0.248 scale IRFM. The IRFM tests have been canpleted. Documentation of test results will be available in August 1982.

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QCS760.117-2 Amend. 69 July 1982

Page - 6 (8,22) #104 Ouestion CS760.118 Provide a Failure Modes and Ef fects Analysis (FMEA) for the PHTS, lHTS, the DHRS, the PACC's, or other systems involving shutdown heat removal and natural circulation?

Resoonse Failure Modes and Ef fects Analysis (FMEA's) f or the PHTS, lHTS, DHRS, PACCs, and other systems involving shutdown heat removal and natural circulation are currently being conducted. Current Project FMEAs are not sufficiently finalized to provide adequate or appropriate information at this time. This information will be available at the OL state.

QCS760.118-1 Amend. 69 July 1982 n a a tv w

Page-1(6,22)#1h7 Ouestien CS760.120 Given a complete loss of the PHTS and trip, what time is available to either recover the system or to put the DHRS Into operation before (A) unacceptable loss of level in the vessel, (B) dry out of the Steam Generator's or (C) boiling in the hot channel or the core?

Response

The Project interprets the NRC question to be one which questions the operator response time to initiate the DHRS assuming all heat transfer is lost at the IHX at the time of scram.

This information can be found in the response to NRC Question 760.38.

i QCS760.120-1 Amend. 69-

p:ge 1 W82-0358 (8,22) 94 Ouestion CS760.121 Has a reliabill*y . lysis on the PACC system been initiated? What methods and models are presently used to determine:

o corrosion impact o heat transfer deterioration o monitoring

- parameters during transients

- frequency

- testing

- maintenance o frequency of demand o analysis of operation

- nominal

- off-normal possibilities of overcooling varying steam supply loss of power to fan changing steam conditions Resoonse Yes; both quantitative and qualitative reliability analyses on the PACC system have been initiated. The analyses utilizes failure state modeling and Failure Mode and Effects Analysis (FMEA).

o Corrosion Imoact The impact of corrosion on the reliability of the PACC system is being considered. In the quantitative reliability evaluation of the PACC, corrosion is considered as one of several causative f actors for PACC leakage and is included in the PACC system leakage failure rate estimate. Corrosion is considered as a possible failure cause. The impact of corrosion would be minimized by inservice inspection and by stringent water chemistry requirments.

o Heat Transfer Deterioration The impact of heat transfer deterioration on the reliability I of the PACC is being considered. Degraded heat transfer l capability has been considered as a PACC f ailure mode. The results how that the fouling f actor is not significant.

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o Monitoring The PACC reliability analyses have not taken credit for the benefit of monitoring specific PACC parameters during transients.

l QCS760.121-1 Amend. 69

p ga 2 W82-0338 (8,22) 94 In the quantitative reliability evalution of the PACC, consideration of monitoring f requency, testing, and maintenance is being given. Monitoring frequency and testing impact on the PACC reliability is considered by evaluating the PACC model using a one-week inspection and testing Interval. The Impact of maintenance on the PACC is -

considered in terms of repair times used to estimate PACC unavailability, in the FMEA, monitoring f requency, testing and maintenance

• are being considered (and in most cases eliminated) as possible failure causes.

o Frecuenev of Demand The Impact of frequency of demand on PACC reliability is being considered through tie use of Shutdown Heat Removal System (SHRS) shutdown initt aiors. This is, the PACC f ailure state model is evaluated using shutdown initiators which place a demand on the PACC.

o Analysis of Ooeration in quantitative reliability analysis, nominal operating conditions are being considered. The loss of power to the fans is included in the PACC evaluations. Items: (a)

Possibilities of overcooling, (b) varying steam supply, and (c) changing steam conditions, are not appropriate for the failure state model evaluation.

QCS760.121-2 Amend. 69 July 1982 e5a-awn

p;ge : LAK-WJMWW Ouestion CS760.122 What data base was used to determine the design characteristics of the PACC and its operational abilities?

Response

The data base being used for the design of the PACC is described in the response to NRC Question 760.035.

QCS760.122-1 Amend. 69 July 1982 82-0420

P:ga - 4 (8,22) #107 Ouestion CS760.127 The cooldown limit of 150 F/hr is quite high compared to LWR limits. What is the limiting component at this cooldown rate? Are the associated thermal stresses based on perfect mixing in the upper plenum? ,

Resoonse t

No component has been identified as being limiting for the normal cooldown from 40% power. In this regime near perfect mixing in the outlet plenum exists and the thermal stresses are calculated on this basis.

QCS760.127-1 Amend. 69 July 1982

v. us,v Pega - 5 (8,22) #107 Ouestion CS760.129 Does the estimated frequency of reactor scram (around 10 per year) take into account spurious trips introduced by the severely reduced secon.dary trip setti ngs?

Resoonse The f ollowing definition of a reactor trip f rom the PSAR* Is provided to define the items included for determining the f requency of reactor scrams:

U-1 Reactor Trlo This transient includes real scrams due to malfunctions (including rapid reactivity transients) which cause a PPS trip level to be exceeded, and spurious scrams covering those situations in which a PPS trip level is not actually exceeded but a scram occurs due to a f ault in the PPS, control system, or plant instrumentation.

The ef fect on plant availability is considered when determining the setpoints for the Primary and Secondary RSS trips. The settings f or the secondary trip 4 set points will not lead to excessive spurious trips.

• PS AR Appendix B, Section B.I .2.1 QCS760.129-1 Amend. 69 July 1982

P:gs 1 (h82-0426) #117 Ouestion CS760.130 What is the basis for the 3% per minute rate of change of power limit during load follow operations? Can you provide data that demonstrate that fuel cladding mechanical Interaction will not lead to excessive cladding f ailure during load follow? Have control system Interactions been considered in determining time delays to scram? Please provide a discussion with the detalIs of the analysis.

Resoonse The CRBRP plant is designed to provide the capability to load follow. The reactor vessel, piping and other systems are being designed to accommodate thermal stresses and heat up rates consistent with a maximum 3% per minute power rate change. However, the applicant has no plans to operate CRBRP in the load follow mode during the first 328 full power days (first core f uel load) because the capability to sustain cyclic rapid power increases in the fuel has not been demonstrated. The 3% per minute is also an upper limit and slower ramp rates (slower heat up rates) can be accommodated by the system.

The automatic control system precludes reactivity ramps exceeding 13% of fulI power during steady state operation.

Later core loads of CRBRP may be cycled in a load follow type of operational mode. Initial tests to demonstrate this capability have been perf ormed in EBR-i l (run 112 with 1%/sec power change). Additional power cycle tests are planned as part of the operational reliability transient testing progrm in EBR- l l . At this time, only calculations with the LIFE computer code can be perf ormed to indicate the adequacy of the f uel to sustain load follow operati on. However, load cycle tests in EBR-ll and load follow tests in f oreign reactors (see Ref erences, QCS760.130-1 and 2) Indicate no obvious rod damage due to reactor load fol Iow operation.

Load follow power rate changes can be firmly addressed af ter the irradiation testi ng has been perf ormed. The rate of power change increases may depend on the power swing and the duration of operation at the lower power.

Power increase rates af ter extended low power operations will be determined i

based on the planned Operational Reliability Testing (ORT) progre transient tests in EBR-ll and FFTF experience. The steady state and transient progran l

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l QCS760,130-1 Amend. 69

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P ge 2 (WB2-0426) #117 plans will be provided to NRC via a summary description document bef ore the end of FY82 which will include the ORT program transient tests.

Control System Interactions have been considered in the Chapter 15 Safety Analysi s. For f ast transients, PPS action occurs prior to Control System respon se. For slow transients, Control System response tends to mitigate the event consequences. In general, a conservative accident scenario is postulated in PSAR Chapter 15 analyses by assuming no Control System action.

During power operation, the time response of the Reactor Control System is approximately 20 seconds; the time responses of the Sodium Flow Control Systems are also approximately 20 seconds. For most transients, the time it takes the PPS to recognize a scram condition and initiate a scram is on the order of 0-5 seconds; thus, the control system interactions do not cause appreciable delays in the time to scram. Additionally, control system Interactions generally act to decrease reactor power or increase sodium flow during transients.

References QCS760.130-1 R. Lal lement, " French Experience Concerning the Reactor Behavior of Breeder Fuel Elements",1981 ANS/ ENS Meeting on Reactor Saf ety Aspects of Fuel Behavior.

QCS760.130-2 T. Rousseau, et. al ., " Fast Neutron Reactor Fuel El ements and Power Grid Duty Cycling", Internationel Conf erence on Fast Breeder Reactor Fuel Perf ormance, Monterey, Cal ifornia,1979.

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QCS760.130-2 Amend. 69 July 1982

Pcg2 - 6 (8,22) #107 Ouestion CS760.132 No discussion of pressurization transients is given in 5.7.3.

Please Indicate the margin in the design for overpressurization resulting f rom elther postuiated accidents in Section 5 or . steam generator tube f alIures, whichever is applicable.

Resoonse The Sodium-Water Reaction (SWR) Design Basis Leak (DBL) as discussed in Section 5.5.3.6 is the transient which imposes highest pressures in the Intermediate Heat Transport System (IHTS). The Sodium-Water Reaction Pressure Rellef Subsystem (SWRPRS) as described in Sections 5.5.2.4 and 5.5.2.6 is an overpressure protection system designed to ASME Code Section lil, Division I, Class 3. In the event of a large SWR, the SWRPRS functions autanatically to limit lHTS pressures to below emergency (Level C) limits.* The SWRPRS rupture disks are nominally rated at 325 PSID so that sustained overpressure is limited to about 325 PSIG. At major components, transient pressures as high as 395 PSI A (Table 5.5-11) are associated with the SWR DDL. The portion of the IHTS that will be exposed to the above described event is designed to ASME Code, Section Ill, Division I, Class 1. The margine relative to design limits wilI be determined upon completion of component Final Stress Reports.

• The sodium / water reaction event is classified as f aulted for the af fected stean generator unit, for the Interconnecting piping to the other steam generator units in the loop, for the injected reaction products separator tank (s), af fected rupture discs and Interconnecting pressure relief piping.

For the rest of the plant, the event is classified as an Emergency Event.

QCS760.132-1 Amend. 69 l

July 1982

Pag 3 - 3 (8,22) #95 Ouestion CS760.135 The saturated steam line rupture is identified as the most severe event in terms of the temperature transient seen by the evaporator and IHX but there does not appear to be any evidence that the designs can accommodate such a severe thermal shock.

Pleasa provide evidence that these components can accept a 5000F change over a 10 minute time span. Please provids a list of critical components and the acceptable thermal transients associated with each of them.

Resoonse The IHX design specification identifies temperature and flow transients that the IHX must withstand. The saturated steam line rupture event is identified as one of the events that the IHX shall be designed f or. The thermal and stress analysis of the IHX has been completed and the vendor has certified that the IHX can accommodate all specified thermal transients including the saturated steam line rupture event.

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l QCS760.135-1 Amend. 69 bk .. .... _ _ - - - _ _

Page - 10 (8,22) #95 l

l Ouestion CS760.139 Please Indicate the f unctional requirements for the Battery Backup System for the following areas:

o Total power requirements valves turbine control other o design life  !

o maximum length of time under battery operation Has any reliability study been done for this system?

Resoonse The CRBRP DC power system, which consists of Class IE and non-Class 1E power supplies, is described in the PSAR Section 8.3.2. The batteries are sized in accordance with IEEE Standard 485-1978 and include temperature correcilon and aging factors. Each battery is sized for its maximum expected duty cycle including a design margin for load growth. The battery chargers are sized to provide DC power to all continuous loads and also to charge the batteries from a totally discharged state to full charge within twelve (12) hours.

A list of Class IE (Division 1, 2 and 3) DC loads is provided in Tables 8.3-2A, 8.3-2B and 8.3-20 of the PSAR. These tables include the power requirement and duration of operation for each load. The Class IE loads which require uninterruptible AC power supply are powered by the corresponding DC  !

battery through an inverter (DC to AC).

The DC system design for CRBRP is presently under development. The complete list of Class IE and non-Class IE loads (DC and AC), which are supported by the DC batteries upon loss of the Plant AC power, will be included in the FSAR Chaptor 8. As a minimum this list will provide load description, power requirement and maximum duration of operation for each load, as well as total power requirements for each DC bus.

Ratings of the Class IE and non-Class 1E DC batteries and DC to AC Inverters are shown in the PSAR Chapter 8, Figure 8.3-2.

The qualification life of the station batteries will be in excess of 15 years.

It is expected that each battery will be replaced once during the 30 year life of the plant.

Normally, the batteries will be on float with the continuous DC load suppiled by the battery chargers. On loss of AC power, the batteries will supply power to the connected DC loads for a period of at least two (2) hours without recharging, except for the security battery which will require recharging af ter 15 minutes.

A reliability study for the DC system will be performed and the results will be included in the FSAR.

QCS760.139-1 Amend. 69 July 1982

F Paga 1 (t:2-0427) #116 Ouestion CS760.140 Some portions of the hot leg piping require further inelastic calculations (WARD-D-0185, Pg. 4.1-18 ) . Have these calculations been perf ormed? Please provide the results of these calculations.

Response ,

Inelastic analyses of the CRBRP primary heat transport system hot leg between the pump and Intermediate heat exchanger have been perf ormed. These results confirm the Integrity of the piping that was f orecast based upon elastic analysis results. The results of this analysis is given in Reference QCS760.140-1.

References QCS760.140-1 A.K. Dhalla, "A Procedure to Evaluate Structural Adequacy of A Piping System in Creep Range", ASME Publication PVP-63, Anerican Society of Mechanical Engineers, New York,1982.

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