ML20070M720

Forwards Responses to Open Items of SRP Discussed at 830110-12 Meetings

ML20070M720
Person / Time
Site:	Seabrook
Issue date:	01/14/1983
From:	Wingate H PUBLIC SERVICE CO. OF NEW HAMPSHIRE
To:	Knighton G Office of Nuclear Reactor Regulation
References
SBN-421, NUDOCS 8301250310
Download: ML20070M720 (25)

Text

.

SEABROOK STATION g

Engineesbg ONke:

1671 Worcemer Road M " M h

• 01701 Pubbe Service of New Hampshire (617) - s72 - 3100 January 14, 1983 SBN-421 T.F. B7.1.2 United States Nuclear Regulatory Commission Washington, D. C. 20555 Attention:

Mr. George W. Knighton, Chief Licensing Branch No. 3 Division of Licensing

References:

(a)

Construction Permits CPPR-135 and CPPR-136, Docket Nos. 50-443 and 50-444

Subject:

Open Item Responses

Dear Sir:

We have enclosed responses to the following open items which were discussed with representatives of the NRC Staff in meetings conducted on January 10-12, 1983:

NRC BRANCH SRP SECTION TMMENTS CSB 6.2.5 Hydrogen Recombiner Actuation MTEB 5.3.1 Compliance with Appendix H.

Paragraph II.B MTEB 5.3.1 Heatup/Cooldown Curves Beyond 3.9 EFPY RSB 5.2.2 Pressurizer Relief Valve Capacity (440.5, 440.6, 440.7)

ASB 9.1.1 New Fuel Criticality Analysis The enclosed responses will be included in a future Amendment to the OL Application.

Very truly yours, Y TEE ATOMIC ELECTRIC COMPANY

[ 8301250310 830114 Henry E. Wingat PDR ADOCK 05000443 Acting Project Manager A

PDR ALL/fsf cc: Atomic Safety and Licensing Board Service List 1000ElmSt.,P.O. Box 330. Manchester NH03105

• Telephone (603)669-4000. TWX7102207595

ASLB SERVICE LIST

~ ~~

O Philip Ahrens, Esquire Assistant Attorney General Department of the Attorney General Augusta, ME 04333 Representative Beverly Hollingworth Coastal Chamber of Commerce 209 Winnacunnet Road Hampton, NH 03842 William S. Jordan, III, Esquire Harmon & Weiss 1725 I Street, N.W.

Suite 506 Washington, DC 20006 E. Tupper Kinder, Esquire Assistant Attorney General Office of the Attorney General 208 State House Annex Concord, NH 03301 Robert A. Backus, Esquire 116 Lowell Street P.O. Box 516 Manchester, NH 03105 Edward J. McDermott, Esquire Sanders and McDermott Professional Association 408 Lafayette Road Hampton, NH 03842 Jo Ann Shotwell, Esquire Assistant Attorney General Environmental Protection $ureau Department of the Attornty General One Ashburton Place, 19th Floor Boston, MA 02108 G

eew-a-,m% e e,q. = 4W 9 e-

=- w we e-w wy-m,**,m.mse

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l l

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Although Sub=ction 15.6.5.3.c.1 indic:t23 that only 0.3% of the zirmnh

[ would react, for conservatism, 5% is assumed to participate in a reaction J with the coolan*: to produce hydrogen coincident with the accident.

Y flow below the operating floor promotes a flow from those regions upward to the suction of the fans at the apex of the done, with r

the containment atmosphere passing through the sprayed region.

Thus, no stratification of hydrogen is possible. The system meets the requirements of Regulatory Guide 1.7, Item C1.

The pertinent data for evaluation of the control aspect of the combustible gas control system is summarized in Table 6.2-89. yin

. order to prevent the concentration of hydrogen from reaching 4 v/o in the containment, the recombiners would be turned on at any time before, or when, the concentration reached 3.5 v/o. Utilizing the model presented in Table 1 of Regulatory Guide 1.7, it is found that, if only one recombiner is turned on fr34 hours after a LOCA

,)34 when the 3.5 v/o setpoint is reached, the hydrogen concentration begins to drop immediately. The combustible gas concentration in the containment as a function of time is shown on Figure 6.2-99, both vi h and without the operation of a single recombiner, at an efficiency of 99.9%.

The backup purge system, as described 4---

in Subsection 6.2.5.2d, could be operated at rates up to 1000 scfm.

This far exceeds that of a single recombiner which processes only 100 scfm. Although actuation of the recombiners could be delayed, the 3.5 v/o is selected to have margin for detection of the failure of both recombiners to function. In this event, the backup purge system would be started at the 3.8 v/o level. This would leave approximate 1y JgsGsfdays to start purging. Figure 6.2-100 shows h*

4 the containment hydrogen concentration with neither recombiner nor purge, and with purging at a rate of only 2 percent of the contain-ment volume per day starting at the 3.8 v/o mark,gours after_

8//

the LOCA. The actual purge rate used will be based on an analysis of the containment atmosphere following a LOCA. Projected offsite doses resulting from containment purging, if required, are described in Subsection 15.6.5.4.

6.2.5.4 Testing and Inspection a.

Hydrogen Analyzer 1

l The hydrogen analyzer is shop-tested using a gas mixture closely l

simu' acing the containment post-LOCA atmosphere expected, at the l

time the units would be placed into service, with the temperature, pressure, humidity and hydrogen concentration conditions approximated.

During pre-operational testing, the hydrogen analyzers are calibrated and checked for proper operation. System integrity will be verified during the containment leak rate tests. Periodic calibration tests will be performed in accordance with Technical Specification requirements.

6.2-79

+

SB 1 & 2 FSAR Q

TABLE 6.2-89 PARAMETERS USED TO EVALUATE CGCS PERFORMANCE Reactor Power Level, HW t 3650.6 6

Initial Volume of Containment Atmosphere, scf 2.55 x 10 Mass of Zircaloy Cladding, lb 46,000-Hydrogen in Primary Coolant, scf 1127 Hydrogen Production Rates (scfh) from Corrosion:

Time Post-IDCA Containmen t (hr)

Temperature (O )

Aluminum Zinc F

0-24 271 440.72 468.10 24-48 200 73.66 67.12 48 152 4.32 14.04 e

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Capsules U, V, W, X, Y, and Z Material Charpy Tensile 1/2T - CT l Plate R 1808-3 Longitudinal 15 3

4 Transverse 15 3

4 Weld Metal

• 15 3

4 HAZ 15

• Fabricated with same heat of weld wire and lot of flux used in vessel b31tline region circumferential weld.

l 24610:1/121382

TABLE 2 1

TYPE AND NUMBER OF SPECIMENS IN THE SEABROOK STATION UNIT NO. 2 SURVEILLANCE TEST CAPSULES Capsules U, V, W, X, Y, and Z Material Charpy Tensile 1/2T - CT Plato R 3407-2 Longitudinal 15 3

4 Transverse 15 3

4 Weld Metal

• 15 3

4 HAZ 15 0 Fabricated with same heat of weld wire and lot of flux used in vessel beltline region cin:umferential weld.

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TABLE 3 SURVEILLANCE CAPSULE REMOVAL SCHEDULE UNITS 1 and 2 Orientation Capsule of Lead Removal Expected Capsule Identification Capsules Factor (a)

Time Fluence (n/cm )

2 18 U

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4.00 1st Refueling 3.3 x 10 Y

241*

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V 61

• 3.69 9 EFPY 2.2 x 10l9(c)

X 238.5*

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1 21.5*

4.00 Stand-By Z

3 01. 5

• 4.00 S tand-By a

The factor by which the capsule fluence leads the vessels maximum inner wall fluence.

b Approximate Fluence at 1/4 wall thickness at End-of-Life.

c Approximate Fluence at vessel inner wall at End-of-Life.

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RSB SRP 52.2

NRC Question - RSB Open Item Section 5.2.2 - The applicant to demonstrate that adequate relief valve capacity is provided assuming the reactor trip is initiated by the second safety grade trip signal.

(Q.440.5,Q.440.6,Q,440.7) l e

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. Response:

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• The following provides infomation related to pressurizer safety valve sizing:

. ' Pressurizer safety valve sizing is a two-step process.

Initially, assumptions are made as to the worst anticipated transient and a valve size in excess of that required is chosen.

Secondly, all of the anticipated overpressure transients are analyzed, using the selected valve capacity.

If the results from all the transients show that ASME Code allowables are not exceeded, l

then the selected valve capacity is accepted.

This is discussed in WCAP-7769, Revision 1.

In Section 2 of this WCAP, the sizing transient is clearly stated as the loss of load transient with feed-water maintained (WCAP-7769, Rev. 1, page 2-1).

The same transient con-ducted with loss of feedwater does not result in as large a valve capacity required.

Figures 2-1 and 2-2 display the loss of load transient b th with and without feedwater being maintained.

From Figure 2-2, it any bc )bserved that the maximum safety valve capacity (0.86, normalized) is less Tor the loss of feedwater transient than that required for the transient with feed-water maintained (0.90, nomalized) and shown in Figure 2-1.

In short, more safety valve capacity is required for the loss of load transient with feedwater maintai'.ied.

i Tha additional plots shown in Figure 2-2 display the results of the transient that could be crpected if the reactor protection system is credited in tripping the reactor.

These same curves could have been plotted on Figure 2-1 but were not for the sake of clarity.

It is easily seen. that the reactor.

coolant system (RCS) is protected from overpressure -

regaNiess of reactor trip, assuming all safety valves function properly.

Should the reactor trip at the first protection grade trip (high pressurizer pressure), then only 40 percent of the total valve capacity is required.

This 40 percent capacity readily falls within that provided by two of the three safety valves.

In conclusion, the RCS is adequately protected from overpressure with only two (of the three) safety valves if the-reactor trips at the first protection grade trip setpoint.

Ono must realize that the plots shown in Section 2 of the WCAP are typical of those conducted in the early 1970's time frame.

Plant specific calculations are done and reported in the plant specific overpressure protection report required by Section III of the ASME code.

This report is certified by a registered professional engineer and details the overpressure analyses for the specific plant.

Component parameters as well as the actual reactor protection system setpoints for the specific plant are used.

As Westinghouse designed 4-loop plants are very similar in reactor coolant system components and thermodynamics, the safety valve sizing transients for virtually all 4-loop plants produce almost the same valve capacity plots.

A l

standard pressurizer safety valve size of 420,000 lbm/hr is, therefore, used by Westinghouse.

As the sizing transient is based on no reactor trip, a valve capacity plot calculated for a 4-loop plant today would almost exactly ovarlay that displayed in Figure 2-1 or 2-2 of the WCAP.

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A majcr change that shows up in plant specific analyses conducted today versu:. those in the early 1970's is rod motion delay time.

Today, 2 seconds is us !d as compared to 1 second in the early 1970's for rod motion delay time following the high pressurizer pressure trip setpoint.

This causes a notice-able :hange in the plot shown in Figure 2-2 which displays the transient with trip accurring on high pressurizer pressure.

As reactor heat exists for an additional second, the pressurizer pressure goes higher and, consequently, more valve capacity is used.

This increased capacity is displayed in the plaat specific overpressure protection report.

For example, the overpressure protection analysis for Seabrook shows that with a 2 second rod motion delay time following a loss of load transient tripping on high pressurizer pressure, a peak pressurizer pressure of approxi-mately 2550 psia is reached.

The valve capacity used is approximately 67 tercent of capacity.

(RCS pressure peaks at approximately 2640 psia.)

Rod motion delay time has less effect following the second reactor protection syst em trip setpoint.

Here, peak pressurizer pressure has already occurred and the pressure is declining.

Any additional delays in tripping the reactor at this stage have no effect on the overpressure protection afforded the RCS.

The iteam system may be overpressurized, however, should feedwater be lost.

This takes a considerable amount of time, however, and can be prevented by reactor trip from any of the trip functions (see Page 2-9 of the WCAP).

We can go back to the 2 second rod motion delay following the high pressurizer pressure trip setpoint.

Should pressurizer safety valve capacity available be less than 67 percent (2 of 3 valves), then RCS pressure will go higher.

A large margin exists between the presently calculated peak RCS pressure and the ASME code allowable of 2750 psia (110 percent).

The limiting cases of 2750 psia were calculated as part of the EPRI safety and relief valve test program and found to lie within 40-50 percent of total valve capacity.

Thus, RCS pressure should not exceed 2750 psia unless the valve capacity available fell below 40-50 percent.

Section 3 of WCAP-7769, Revision 1, is sometimes misinterpreted.

The analyses preser.ted in Section 3 are examples of the anticipated overpressure transients for the Salen Nuclear Generating Station.

They are only given here (in the WCAP) as examples of the second stage in the two-step sizing process discussed above.

Salem specific parameters are used and Salem specific plots are dis-played.

The conclusions pointed out on page 3-61 of the WCAP are applicable to Salem.

Overpressure analyses for a particular plant are detailed in the plant specific overpressure protection report.

The reference to WCAP-7769 contained in the FSAR is appropriate in that the WCAP generically addresses valve sizing.

The additional mat rial presented in Section 4 is applicable and straightforward.

Section 3, however, is instructive but should only be taken as an example of overpressure protection afforded a particular plant (Salem).

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2 RE FERENCES (1) SBU:51342, Seabrook Station - Units 1 and 2, Design and Fabrication Criteria for the New Fuel Storage Rack, UE&C, January 4,1982 (2)

" KENO-IV - An Improved Monte Carlo Program", ORNL-494, L. M. Petrie, N. F. Cross, 1975 (3) "AMPX: A Modular Code System for Generating Coupled Multi-Group Neutron Camma Libraries f rom ENDF/B", ORNL/TM-3706, N. M. Greene et. al, 1976 (4)

"123-Group Neutron Cross-Section Data Generated f rom ENDF/B-II Data for Use in the XSDRN Discrete Ordinates Spectral Averaging Code", DLC-16,

,f W. R. Cobb, 1971

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DISCUSSION' The criticality of Seabrook's new fuel storage racks has been studied as a function of moderator density. This was done in order to answer NRC l

Question 410.15 concerning the criticality of the racks in the presence of

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foam or mist.

NEW FUEL STORACE VAULT The new fuel storage vault, Figure 1, is a 17 by 5 array of storage locations with a 21" minimum center-to-center spacing between locations.

In addition, two central columns of assemblies are spaced at 33", center-to-center, to allow for f uel inspection.

The array is surrounded by 1*

thick concrete walls g

with the outer row of assemblies one foot from assembly center to the walls, (Reference (1)).

The vault is normally dry; moderator (water) is only introduced by the presence of people or by abnormal situations, like fires, which require foam or mist.

MOTHOU CF ANAI.YSIS The criticality analysis for the new fuel vault was done with the NITAWL-KENO-IV methodology and the 123-neutron energy group XSDRN library (References 2, 3, and 4).

The KENO model employed in this analysis was constructed f rom a basic unit of analysis shown in dashed outline in Figure 1 and shown in detail in Figure 2.

The model assumes that 17 x 17 Westinghouse standard fuel of 3.1 w/o U235 enrichment is loaded into all~ storage locations.

The KENO model is a 3-D model which is finite in the 5 unit direction and in the axial direction but infinite in the 17 unit direction. One foot of re flector region is included at the top and bottom of the active fuel

_leng th (143.7").

The concrete wall surrounding the array is also included on the side.

No grid structure, tie plates, or can is included in the model which is a conservative assumption, since such structures are parasitic absorbers.

RESULTS The result of the criticality analysis is,shown 'in Figure 3.

K gg.is e

plotted vs. moderator void with 0% void being the completely flooded case and 100% void the completely dry case. Zero percent (0%) void corresponds to 3

water at a density of 0.9982 gm/cm.

There are three regions to Figure 3.

The first, 0-60% void, is characterized by decreasing K gg due to decreasing moderator density and, t he re fo re,

e decreasing neutron moderation within the assemblies. The second region,60-952 void, is characterized by rapidly increasing K gg due to neutron e

moderation in the water gap between assemblics. This second region is the region of optimum moderation, and the moderator density of this region roughly corresponds to the density of fire fighting foam. The peak value of K rg in e

this region is approximately 0.91 which occurs at 95% void (.04991 gm/cm3 of H O).

In the third region, > 95% void, leakage and reduced overall moderation 2

bring K gg down again.

The final dry rack K rf is approximately 0.30 e

e which is considerably suberitical.

CONCLUSION 4

The Seabrook new fuel storage racks are below the,98 NRC criticality limit

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under conditions of optimum moderation with foam or mist.

Thus, the presence of foam or mist does not pose a hazard. Moreover, one can argue that optimum j

moderation is not achievable because:

1) optimum densities cannot be achieved with pure water at I atm 680F, and

.o 2) although fire fighting foam with an expansion factor of 8-10 can theoretically achieve such densities, it is improbable that the entire l

storage vault would be flooded with the mixture to any appreciable height.

Response to RAI 410.15 l

The Seabrook new fuel storage racks are below the 0.98 NRC criticality limit under conditions of optimum moderation with foam or mist.

This analysis was done with fuel of 3.1 w/o U235, the highest enrichment for the first cycle.

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