ML20207S507
| ML20207S507 | |
| Person / Time | |
|---|---|
| Site: | Grand Gulf, 05000000 |
| Issue date: | 09/30/1986 |
| From: | Economos C BROOKHAVEN NATIONAL LABORATORY |
| To: | NRC |
| Shared Package | |
| ML20207S501 | List: |
| References | |
| CON-FIN-A-3346, CON-FIN-A-3396 NUDOCS 8703190571 | |
| Download: ML20207S507 (28) | |
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Technical Evaluation Report (TER) for the Mississippi Power and Light Company's (MP&L)
Grand Gulf Nuclear Station, Docket Nos. 50-416 and 50-417 on SER Outstanding Issue (8) -
Mark III Containment System Issues -
The Humphrey Safety Concerns (HSC)
I.
by C. Economos September 1986 Department of Nuclear Energy Brookhaven National Laboratory Upton, New York 11973 Introduction With three exceptions, this TER addresses only those HSC's that, by previous agreement with the NRC staff, are considered BNL's responsibility. The three i exceptions are the addition of the issues related to HSC 4.4, 4.5 and 7.1. In earlier TER's (River Bend Station and Perry Nuclear Power Plant), these had been considered the responsibility of the respective Technical Monitors. Not- .
withstanding these earlier decisions, we feel that BNL input in these areas is appropriate and have therefore included our evaluation here. Note that our
- conclusion in each case is that the issues raised have been satisfactorily re-solved.
J 8703190571 870306 PDR ADOCK 05000416 E PDR
Humphrey Safety Concerns 1.1, 1.2, 1.4 and 1.5 1.1 Presence of local encroachments, such as the TIP platform, the drywell personnel airlock and the equipment and floor drain sumps may increase the pool swell velocity by as much as 20 percent.
.1.2 Local encroachments in the pool may cause the bubble breakthrough helght d
to be higher than expected.
1.4 Piping impact loads may be revised as a result of the higher pool swell velocity.
1.5 Impact loads on HCU floor may be imparted and the HCU modules may fail, which could prevent successful scram if the bubble breakthrough height is raised appreciably by local encroachments.
Evaluation The concern is that the presence of encroachments will tend to increase pool velocities and breakthrough height relative to the unencroached pool thereby causing increases in a variety of LOCA loads including bulk and froth impact and drag loads in regions above the initial pool surface.
As a result of the 1/10-scale simulations of pool swell conducted by the Con-tainment Owners Group (CIOG) as reported in References 1.1.1 and 1.1.2, it has now been established that, for the most part, the effect of encroachments is to decrease pool swell velocities relative to an unencroached or clean sup-pression pool. BNL has reviewed all of this information and has prepared a separate Technical Evaluation Report on its findings (Reference 1.1.3).
Conclusion BNL considers these issues closed for the GGNS because the 1/10-scale tests have confirmed that encroachment effects on pool swell behaviour will not give rise to loading conditions that cannot be accommodated by the existing design basis.
References
- 1.1.1 Sets of slides provided by the CIOG to the NRC from the following 1/10-l scale tests
- (a) Clinton Tests F2R, R05, and F5; (b) Grand Gulf Tests E3; (c) Perry Test D2; (d) an unidentified clean pool test (see letter of 19 March 1985 from J.E. Torbeck to R. Pender, cc: J. Kudrick of NRC).
1.1.2 " Comparison of Velocities and Thicknesses of Water Column at Contain-ment Wall for 1/10-Scale Pool Swell Tests," submitted by G.W. Smith of CIOG to H.R. Denton of NRC via letter of 15 May 1985.
1.1.3 Sonin, A. and Economos, C., " Resolution of the Humphrey Issues Relating to Pool Swell in Mark III Plants," BNL Technical Evaluation Report, February 1985.
Humphrey Safety Concern 1.3 Additional submerged structure loads may be applied to submerged structures near local encroachments.
Evaluation The loads addressed under this item fall into two categories: (1) loads on submerged boundaries (drywell wall, basemat, containment wall) and (2) loads on submerged structures proper, such as pipes and beams.
For loads on submerged boundaries the applicant's response is provided in Ref-erence 1.3.1. A two dimensional SOLAV01 simulation is used to determine the effect of encroachment. The simulation is for the GGNS geometry but with a pool width of 19 feet (corresponding to PSTF geometry) rather than the actual plant width of more than 20 feet. Drywell wall pressure is unaffected by this simulation since this is taken equal to peak drywell pressure. The calcula-tions exhibit an increase of about 15% for containment wall and basemat pres-sures. It is stated that this increase is " easily bounded" by the existing design capability.
The applicant's submittal for submerged structure loads is given in Reference 1.3.2. The new loads are developed using the velocity field deriving from the SOLAV01 simulation. The increased loads produced by the encroachments are stated to be within the code allowable limits including functional capability criteria (Reference 1.3.3) .
Conclusion BNL considers this concern to be resolved based on the stated margin between design and the loads estimated with a conservative (two dimensional) simula-tion of the encroachment effect.
References ,
1.3.1 Attachment 1 to MP&L Letter No. AECM-82/497 dated October 22, 1982 from L.F. Dale (MP&L) to H.R. Denton (NRC).
1.3.2 Attachment 1 to MP&L Letter No. AECM-82/574 dated December 3,1982 from L.F. Dale (MP&L) to H.R. Denton (NRC).
l 1.3.3 Attachment 1 to MP&L Letter No. AECM-86/0175 dated August 14, 1986 from 0.D. Kingsley, Jr. (MP&L) to H.R. Denton (NRC) .
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Humphrey Safety Concern 1.6 Local encroachments on the steam tunnel may cause the pool swell froth to move horizontally and apply lateral loads to the gratings around the HCU floor.
Evaluation The utility has performed a potential flow analysis of the flow field through the HCU floor (Reference 1.6.1). This analysis assumed steady flow, i.e., the liquid droplets had velocities equal to air and all of the froth was allowed to pass through the openings in the HCU floor. The resulting lateral pres-sures were found to be 0.45 psid on the beams and 0.30 psid on the gratings with beams located at the edge of solid floor areas experiencing a maximum horizontal force of 1.45 psid. All affected structures are stated "to be capable of withstanding the identified loads without deformation."
BNL concurs with the utility claim that the potential flow analysis is conser-vative. In fact, independent calculations conducted by the NRC indicate that over 90% of the froth will continue in the vertical direction, impact on the HCU floor and lose all its velocity. As the froth begins to fall back toward the pool, the horizontal component of the flowing air will accelerate the froth to some extent but steady-state conditions are not expected to be attained. Considering that the additional stresses are modest even with a conservative flow analysis, BNL does not feel that additional effort needs to be expended on this issue.
Conclusion BNL considers this issue to be closed.
References 1.6.1 Attachment 1 to MP&L Letter No. AECM-82/497 dated October 22, 1982 from L.F. Dale (MPAL) to H.R. Denton (NRC).
Humphrey Safety Concern 1.7 GE suggests that at least 1500 square feet of open area should be maintained in the HCU floor. In order to avoid excessive pressure differentials, at least 1500 square feet of open area should be maintained at each containment elevation.
Evaluation The applicant indicates in Reference 1.7.1 that the MPAL design provides open areas on all floors above the HCU floor greater than the open area at the HCU floor. The latter, at 1538 square feet, is greater than the required minimum.
Conclusion Based on the above response, BNL considers this issue to be closed.
References 1.7.1 Attachment 2 to MP&L Letter No. AECM-86/0175 dated August 14, 1986 from 0.0. Kingsley, Jr. (MP&L) to H.R. Denton (NRC) .
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Hunphrey Safety Concern 2.1 The annular regions between the safety relief valve lines and the drywell wall penetration sleeves may produce condensation oscillation (C.O.) frequencies near the drywell and containment wall structural resonance frequencies.
Evaluation As stated above, the concern is that additional and unaccounted for suppres-sion pool boundary loads may be produced by steam condensation at the exit of the sleeve annulus. However, the scope of this concern is expanded considera-bly by Humphrey in Reference 2.1.1. There it is speculated that, due to reso-nant coupling between the sleeve annulus C0 and the sleeve annulus acoustics, the pressure loads at higher frequencies could be amplified.
The applicant has addressed this concern using the generic approach described in Reference 2.1.2. This methodology derives sleeve annulus C0 loads by con-servatively scaling down the main vent C0 load data base. The potential for resonant amplification of these loads was not addressed.
The use of the main vent C0 data base for development of sleeve annulus loads cannot be rigorously defended because of substantial differences in geometry.
However, based on the relative size of the steam / water interface that would exist at the sleeve annulus, BNL would judge that any additional non-resonant loads that_may occur will be second order relative to main vent loads. This judgement applies to both C0 and chugging loads.
The added sleeve C0 loads that were first proposed are substantial. For exam-ple, the peak-to-peak pressure amplitude (PPA) was about 20% of that used in the main vent load definition. They are also applied uniformly in the circum-ferential direction which represents a sizable conservatism. This is because there are roughly twice as many main vents as drywell penetration sleeves.
These modifications are clearly more than second order. Thus, provided it could have been demonstrated that resonant amplification does not occur, the C0 loads which were specified would have been considered adequate.
4 In an attempt to demonstrate the absence of such a coupling, results from Gen-eral Electric's 4TC0 tests (Reference 2.1.3) were cited under the Clinton Power Station docket (Reference 2.1.4). a review of this material indicated that the contrary was the case; that is, the data implied that the type of resonant coupling suggested by Mr. Humphrey was not only possible but appar-ently had actually occurred. In fact, it can be inferred from this data that j
resonance causes about a two-fold increase in the basic C0 loads, As a result of this finding, a completely new methodology for the C0 boundary loads was proposed, again under the CPS docket (Reference 2.1.4). This meth-odology has been developed by General Electric utilizing the fiark I FSTF data base (2.1.5) . This new design loading results in a substantial increase in the pressure loading at the higher end of the frequency spectrum (20-50 Hz).
For example, in terms of an amplified response spectrum, the load intensity is about double the one first proposed. These new loads are shown to be bounded by other design basis loads. For example, on the drywell wall the sleeve C0 load when added to the main vent C0 is bounded by the chugging load l
d specification. On the containment wall, the bound is provided, with considera-t ble margin, by the pool swell boundary load.
The applicability of the. FSTF data to the sleeve annulus loads involves consid-erable uncertainty because of the great disparity in geometries between the two ,
situations. This applies not only for the steam-water interface at the respec-a tive pipe exits but, more importantly, for the acoustic path through which the mechanism that- drives the C0 phenomenon is transmitted. In fact, it is not com-pletely apparent that for the FSTF case the system has actually achieved a con-
- dition of resonant coupling. This is because of the complexity of the FSTF vent system; 1.e. the eight downcomers are connected to a vent header, which in turn i
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is connected to a main vent which then connects to a simulated drywell. Because of this complexity, it is difficult to ascertain the effective relevant vent system natural _ frequency with sufficient precision.
j Despite the uncertainties cited above, there are several factors that may be cited that compensate for any possible inadequacy. First, there is the qualita-f tive evidence from the 4TC0 tests (Reference 2.1.3) that, even when resonant coupling clearly occurs, the load amplification is limited to less than a two-fold factor. Also, the generally conservative application of the available re-
, sults provides increased confidence in the adequacy of the load method. For ex-
! ample, the pressure results observed on 24" diameter downcomers are taken over i directly and applied to the much smaller,14" diameter, SRVDL sleeves. Also, in j developing the loads on the suppression pool boundaries, the effective source (steam bubble) radius was taken equal to that of the sleeve without taking into account the actual presence of the SRVDL itself. As shown in a recent submittal '
under the CPS docket (Reference 2.1.6), this results in a margin of over 30% in the loads that were developed. BNL believes the margin would be even greater if
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. the steam bubble were modeled more ' realistically (Reference 2.1.2). Additional conservatism stems from the use of conventional acoustics for determination of i pressure attenuation from the source to the pool boundaries. Dissipative mech-anisms that are present in the suppression pool and neglected in the analysis, would further reduce the loads.
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Insofar as the chugging loads are concerned, BNL has not received a description of these, even though this information exists (Reference 2.1.7) and has been as-sessed by the Mark III Containment Issues Review Panel (Reference 2.1.8). The findings of this panel were that the proposed loads were only about 6% of main vent chugging and are easily bounded by design. BNL is satisfied that this is the case. Note that resonance effects are not expected to play any role in or influence the chugging phenomenon associated with the sleeve annulus.
Conclusions BNL considers this safety concern to have been satisfactorily resolved because
! the conservatively estimated new loads have been demonstrated to be bounded by
) other design loads, a
! References 2.1.1 Humphrey Engineering, Inc., Letter dated June 17, 1982 from J. M. Humphrey (HEI) to A. Schwencer (NRC).
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2.1.2 IPC Letter No. U-0714 dated May 25, 1984 from D. I. Herborn (IPC) to
. A. Schwencer (NRC) .
2.1.3 Bird, P. F., et al ., "4T Condensation Oscillation Test Program Final Test Report", General Electric Report NEDE-24811-P, May 1980.
2.1.4 IPC Letter No. U-600319 dated from F. A. Spangenberg
. (IPC) to W. R. Butler (NRC) .
2.1.5 Fitzsimmons, G. W. et al ., " Mark I Containment Program - Full-Scale Test Program Final Report", General Electric Report, NEDE-24539-P, April 1979.
2.1.6 IPC Letter No. U- dated from F. A. Spangenburg (IPC) to W. R. Butler (NRC) .
2.1.7 Enercon Letter No. RWE-0G-060 dated Pby 25, 1983 from R. W. Evans (Enercon) to B. R. Patel (Creare R&D).
2.1.8 Mark III Containment Issues Review Panel, " Assessment of Humphrey Concerns", CREARE R&D, Inc., Technical Memorandum TM-928, July 1984.
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Humphrey Safety Concerns 2.2 and 2.3 2.2 The potential condensation oscillation and chugging loads produced through the annular area between the SRVDL and sleeve may apply unaccounted for loads to the SRVDL. Since the SRVDL is unsupported from the quencher to the inside of the drywell wall, this may result in failure of the line.
2.3 The potential condensation oscillation and chugging loads produced through the annular area between the SRVDL and sleeve may apply unaccounted for loads to the penetration sleeve. The loads may also be at or near the nat-ural frequency of the sleeve.
Evaluation The concern bere is that the steam condensation process (C0 + chugging) at the sleeve annulus exit will give rise to loads on the SRVDL and SRVDL sleeve analo-gous to the lateral loads experienced by Mark I and Mark II downcomers during postulated LOCA blowdowns and that these structures have not been designed to accommodate them.
The applicant's specification for chugging loads is given in Reference 2.2.1.
The load has a hal f-sinusoidal time dependence with a duration of 3 m seconds and a peak amplitude of 22.4 kips. This load is derived from the Mark 11 load methodology (Reference 2.2.2) as modified by the NRC staf f's Acceptance Criteria (Reference 2.2.3). The load is developed by scaling down the peak amplitude to the outside diameter of the SRVDL and accounting for the fact that there are fewer chugs created by flow through the SRVDL sleeve annulus than exist during DBA blowdowns through the Mark 11 pressure suppression system (i .e., 20 SRVs for the GGNS vs. about 100 downcomers in a typical Mark 11 plant). Scaling down for pipe diameter is accomplished by assuming a 1.7 power dependence of the peak am-plitude on diameter. Load reduction for fewer chug sources utilizes the staff approved statistical representation for these loads (Reference 2.2.3). The re-gion of application of the load is also scaled down using a first power depen-dence on diameter.
The applicability of the Mark 11 results for the present application is somewhat uncertain due to the substantial geometric differences (straight down vs. in-clined pipe and annnular vs. circular cross section). Nevertheless, we find the approach reasonable and, in general, conservative. The use of a 1.7 power de-pendence of peak amplitude on pipe diameter is somewhat less conservative than we would have preferred since the available data (Reference 2.2.5) exhibit expo-nent values that range from 0.7 to 1.7. On the other hand, no credit is taken for the presence of the SRVDL in the steam bubble. This provides a substantial conservatism that BNL judges more than compensates for any possible non-conservatism in selecting this exponent.
The C0 lateral loads are also presented in Reference 2.2.1. In this case the load is a hannonic with amplitude equal to 630 lb and frequency ranging from 28 to 48 hg. The load is applied as a point load to both the SRVDL and SRVDL sleeve at the end of the latter and perpendicular to the pipe center line. It is stated that application of this load leads to stresses that are within code allowables and that functional capability criteria is met for all affected structures.
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These loads are developed from the same FSTF results used for C0 boundary loads (see discussion under HSC 2.1). In this case the pressure fluctuations observed in the downcomers are applied in an even more conservative manner than that used for the pool boundary loads. The additional conservatism stems from the devel-opsent of a single harmonic amplitude to include the energy content from all frequencies associated with the selected worst case pressure time history. Thi s results in an applied pressure amplitude more than 40% higher than the peak ob-served value (5 vs. 3.5 psi). Application of the specified load as a point load also represents a substantial conservatism since the unbalanced pressures on the structures are actually spread over an area roughly equal to that of the sleeve internal cross section.
Conclusion BNL considers these concerns to be resolved because of the specification of cen-servative C0 and chugging loads on the affected structures.
References 2.2.1 Attachment 3 to MP&L Letter No. AECM-86/0175 dated August 14, 1986 from 0.D. Kingsley, Jr. (MP&L) to H.R. Denton (NRC) .
2.2.2 Davis, W.M., " Mark II Main Vent lateral Loads," GE Report NEDE-23806-P, October 1978.
2.2.3 Anderson, C., " Mark II Containment Program Load Evaluation and Acceptance Criteria," NRC NUREG-0808.
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Htnphrey Safety Concerns 3.1, 3.3, 3.7 3.1 The design of the STRIDE did not consider vent clearing, condensation os-cillation, and chugging loads which might be produced by the actuation of the RHR heat exchanger relief valves.
3.3 Discharge from the RHR relief valves may produce bubble discharge or other submerged structure loads on equipment in the suppression pool.
3.7 The concerns related to the RHR heat exchanger relief valve discharge lines should also be addressed for all other relief lines that exhaust into the pool .
Evaluation The concern is that, besides the main safety / relief valves (MSRV's), there are a number of other valves that discharge fluids into the suppression pool . As a result they could produce loads analogous to those associated with MSRV dis-charges and/or 1.0CA blowdowns through downcomers. These loads have not been accounted for in plant design.
For the RHR system, flow through the heat exchanger relief valves can occur when it is operating in the steam condensing mode (SCM) . During such opera-tion, the heat exchanger is pressurized to about 200 psi by a pressure control valve (PCV). Should the PCV fail, resulting in elevated pressures, the heat ex-changer relief valves would actuate at their setpoint (about 500 psi) and vent this steam to the suppression pool via the relief valve discharge lines. Steam discharges would also be possible in the event that the relief valve itself was to fail open, although in this case, the steam flow rates would be much less.
The initial response to this concern takes credit for the line pressurizing ef-fect of the non-condensible vent flow that is bled from the heat exchanger to the discharge line during SCM operation (Reference 3.1.1). The position taken was that this bleed flow clears the line of water so that water jet and air clearing loads are negligible (Reference 3.1.2). Since then, routing of this flow to the discharge line has been eliminated (Reference 3.1.3). In any case, I
the loads were recalculated without this potential load mitigating effect in Reference 3.1.4. It is stated that the methods employed to evaluate the loads are those developed for rams head type application (Ref. 3.1.5) suitably modi-fled to account for an open ended pipe. All submerged structures were evaluated for these loads and found to be adequate.
For C0 and chugging loads, the applicant utilized the gen?ric methods employed by all the Mark III utilities. A detailed description of these methods is given in References 3.1.1 and 3.1.6. Generally speaking, the method derives from conservative application of the Mark II CO and chugging load methodolo-gies. Source tenns are developed from chugging and C0 pressure signatures selected from the Mark II data base for their conservatism. These source terms are applied to the GGNS plant without any modification to account for the dif-ference in pipe diameter between the GGNS relief valve discharge line (10 inches) and that from which the data base derives (24 inches in the test facili-ty). This is a significant conservatism since it is well established that the source strengths scale with pipe area. The pressure loads generated using these sources are shown to be bounded by MSRV or other design loads. The margins for t
E sthmerged structures, in particular, are characterized as being " considerable"
~ (Reference 3.1.6) .
The adequacy of the C0 load also needs to be judged in the context of the po-tential for unstable steam condensation; i .e.: elevated pool temperatures. -
T BNL's evaluation of this aspect is presented later under Hephrey Safety Con-cern 3.6.
A detailed description of lateral loads on the RHR discharge line due to chug-ging is given in Reference 3.1.1. The load is time dependent (half-sinusoid) with two values of peak amplitude (6.5 kips and 19.5 kips) corresponding to two different pulse durations -(6 ms and 3 ms, respectively). The load is uni-formly distributed over the lowest 4 feet of the discharge pipe. It is stated that this load specification derives from the Mark II lateral load definition 91ven in NUREG-0808 (Reference 3.1.7) and that it differs only in that the peak amplitudes have been scaled down to account for the difference in diame-ter between the RHR discharge line (10 inches) and the standard Mark II down-comer (24 inches). This is accomplished by assuming peak amplitude to scale.
with the 0.7 power of pipe diameter. Application of this load gives resultant stresses that are within code upset allowable stresses (Reference 3.1.1).
The peak amplitudes selected by MP&L are not totally consistent with the staff approved load definition given in Reference 3.1.7. This is because the appli-cant does not recognize the modification that was applied to the Mark II load method of Reference 3.1.8 to account for the stochastic nature of the chugging phenomenon. Such modification results in peak load amplitudes that vary according to the neber of chugs that would be expected during a particular accident scenario and the desired non-exceedance probability. For example, for a DBA LOCA in a Mark 11 plant with a population of 100 downcomers, a peak design load of 65 kips was employed to insure that, statistically, no member of that population is likely to experience an exceedance of this load. This value is substantially higher than the maximum tip load observed experimental-ly (35.9 kips) . It was this latter value that was scaled down by MP&L to ob-tain the design specification of 19.5 kips. .
To evaluate the adequacy of the proposed load amplitude in this context, it is necessary to estimate the neber of chugs the RHR discharge line can experi-ence without an exceedance of the stated value of 19.5 kips. Thi s i s accom-plished by using the nethod described in Reference 3.1.7. The result is a total of almost 5,000 chugs (4675) where we have utilized a more reasonable, but still conservative, linear dependence of load amplitude on pipe diameter.
Even with a conservative estimate of one second between chugs (a more realis-tic neber would be closer to two seconds), chugging could proceed for well over an hour before a single exceedance of structure capability would be ex-pected. Thus, we would not expect serious exceedance of the available load capability to occur during the relatively small number of chugs that these structures will experience during the life of the GGNS.
One other misinterpretation of the approved Mark 11 methodology used by the applicant relates to the region over which the load was applied. The correct method would require the load to be uniformly distributed over a region 1 to 4 feet from the downcomer end. Furthermore, this region needs to be scaled down to account for the difference in pipe diameter. Both of these factors would tend to increase the applied bending moment somewhat but BNL judges that
sufficient margin exists in the load specification to accarunodate this rela-tively minor deficiency.
Loads due to other steam discharges are considered to be bounded by the RHR heat exchanger relief valve loads. In general, this is a reasonable position given the relatively low flow rates and/or smaller discharge line diameters.
A lis*. of all such relief lines was provided in Reference 3.1.8 together with data enaracterizing their size and flow rates. One discharge line which has some potential for creating significant loading is the RCIC turbine exhaust line. In this case the flow rate /didneter combination results in steam flux rates that are in the chugging regime. Significant loads due to the reflood and air / water clearing loads could alt occur since this pipa has a relatively large diameter (20 inches).
The concerns we have enumerated for the RCIC exhaust line have not been ad-dressed directly by the applicant. A " white paper" has been published, how-ever, by the Containment Issues Owners Group (Ref. 3.1.9). This is intended to support the position that operating experience has shown that there are no dynamic load problems associated with operation of this system. The applicant endorses this " white paper" but with " clarifications" (Reference 3.1.10).
These clarifications highlight the fact that during low power testing " chug-ging" was observed at the RCIC exhaust sparger vents which necessitated a mod-ification of the design (the number of sparger holes was decreased). It is stated that this modification eliminated the observed chugging.
Conclusion These issues are considered to be resolved for all dynamic loads associated with the RHR heat exchanger relief valve discharge line. This judgement is based on the conservatisms used for development of design values for these loads and the margins exhibited by other design loads. This judgement is also contingent on the commitment, by the applicant, to eliminate routing of the bleed flow into this disharge line.
As for the loads associated with RCIC operation, BNL defers judgement to the NRC staff since they are in a much better position to evaluate the signifi-cance of whatever operating experience exists.
References .-
3.1.1 Attachment 1 to MP&L Letter No. AECM-83/0146 dated March 23, 1983 from L.F. Dale (MP&L) to H.R. Denton (NRC).
3.1.2 Attachment 1 to MP&L Letter No. AECM-82/574 dated December 3,1982 from L.F. Dale (MP&L) to H.R. Denton (NRC) .
3.1.3 MP&L Letter No. AECM-86/10012 dated January 28, 1986 from 0.0.
Kingsley, Jr. (MP&L) to H.R. Denton (NRC).
3.1.4 Meeting Handouts from NRC/ March I!I/GI Meeting of May 19 and 20,1983 3.1.5 " Mark 11 Dynamic Forcing Function Information Report," NE00-21061, Rev. 3, June 1978.
3.1.6 Ashley, G.K. and Leong , T.S., " An Approach to Chugging. Assessment of RHR Steam Discharge C0 in March III Containments," Bechtel Report, March 1984.
3.1.7 Anderson, C., " Mark II Containment Program Load Evaluation and Accept-ance Criteria," NRC NUREG-0808.
3.1.8 Attachment 1 to MP&L Letter No. 82/353 dated August 19, 1982 from L.F.
Dale (MP&L) to H.R. Denton (NRC) .
3.1.9 Attachment 1 to MP&L Letter No. GWS-06-143 dated &;-11 23,1985 from G.W. Smith (MP&L) to H.R. Denton (NRC).
-3.1.10 Attachment 6 to MP&L Letter No. AECM-86/0175 date August 14, 1986 from 0.D. Kingsley, Jr. (MP&L) to H.R. Denton (NRC) .
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Humphrey Safety Concern 3.2 The STRIDE design provided only 9 inches of submergence above the RHR heat eA-changer relief valve discharge lines at low suppression pool levels.
Evaluation The concern is that.because of the relatively small submergence involved, steam condensation may not be complete leading to steam bypass and failure of the pressure suppression system.
The applicant has addressed this concern using the generic approach common the all plants (Reference 3.2.1). The approach cites the full-scale data from the Humboldt Bay tests where it is shown that, over a wide range of steam flux rate, condensation was complete (i.e., no steam bypass and containment pres-surization), even with a clearance of 2 feet between the vertical vent pipe exit and the pool surface.
In the same submittal, MP&L states that the minimum submergence is 9 inches, increasing to about 4-1/2 feet at normal operating conditions. In BNL's judgement, this is sufficient to insure complete condensation of any steam discharges that may occur via the RHR heat exchanger relief valve discharge lines.
Conclusion BNL considers this issue satisfactorily resolved for the GGNS based on the full scale experimental results that have been cited.
References 3.2.1 MP&L Letter No. AECM-82/353 dated August 19, 1983 from L.F. Dale (MP&L) to H.R. Denton (NRC).
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l Humphrey Safety Concern 3.4 The RHR heat exchanger relief-~ valve discharge lines are provided with vacuum breakers to prevent negative pressure in the lines when discharging steam is condensed in the pool. If the valves experience repeated actuation, the vacuum breaker sizing may not be adequate to prevent drawing slugs of water back through the discharge piping. These slugs of water may apply impact loads to the relief valve or be discharged back into the pool at the next re-lief valve actuation and apply impact loads to submerged structures.
Evaluation The real issue here is that the various steam discharge lines may not have been equipped with properly sized vacuum breakers. This is a-credible concern in view of the historical development of the same issue for the MSRV's.
Because the potential for subsequent actuations was not fully appreciated in the early stages, the MSRV discharge lines were originally equipped with un-dersized vacuum breakers. When very high reflood elevations were encountered during tests with subsequent actuation, it became evident that this was so and much larger vacuum breakers were installed (from 1 inch to as much as 10 inch diameter or two 6 inch diameter).
In Reference 3.4.1 the applicant states that the GGNS RHR heat exchanger dis-charge lines are equipped with what appears to be 1-1/2 inch vacuum breakers.
This is contradicted by information supplied in Reference 3.4.2, where a 3/4 inch vacuum breaker is indicated. We consider even the larger of these barely adequate for the 10 inch diameter line involved. MP&L's reflood analysis re-flects this. In Reference 3.4.2 the results obtained using a conventional analytic method (Reference 3.4.3) show reflood lengths exceeding 30 feet.
This is a surprisingly high value, but according to the applicant, well below any crucial element of the piping system (the vacuum breakers are located about 5 feet above maximum reflood). Piping and piping support dynamic loads were evaluated for a second pop at this peak reflood using, once again, stan-dard methods that are acceptable to the NRC (Reference 3.4.4). It is stated that all structures were found to be adequate; all stresses were within upset allowables.
i Conclusion BMt. considers this issue adequately addressed by the applicant and therefore closed.
References 3.4.1 Attachment 1 to MP&L Letter No. AECM-82/353 dated August 19, 1982 from L.F. Dale (MP&L) to H.R. Denton (NRC).
! 3.4.2 Safwat, H.H., "GGNS RHR Relief Line Hydrodynamic Loads." Meeting Hand-l outs from NRC/ March Ill/GE meeting of May 19 and 20,1983.
l 3.4.3 Ashley, G.K. and Howard, N.M., " Vacuum Relief Valve Sizing in Condens-j ing Steam Situations," Presented at ANS 24th Annual Meeting, June 18-22, 1978.
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3.4.4 Ransom, V.H., et al.. " RECAPS /M001 Code Manual" NUREG/CR-l'826. EG&G Idaho, 1981. "Repipe Application Reference Manual ," CDC, 1980.
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Humphrey Safety Concern 3.6 If the RHR heat exchanger relief valves discharge steam to the upper levels of the suppression pool following a design basis accident, they will significant-ly aggravate suppression pool temperature stratification.
Evaluation Although the concern suggests that these discharges will occur following a DBA, the applicant has indicated (Reference 3.6.1) that the RHR will not be used in steam condensing mode during post-LOCA conditions. On the other hand, continuous steaming for an extended period under normal conditions is possible and could not only result in excessive containment pressurization via vertical thermal stratification but introduces the potential for unstable steam conden-sation leading to excessive dynamic loading on the pool boundaries.
The applicant's response to this issue was given in Reference 3.6.2. A de-monstrably conservative model of thermal deposition, stratification and pool mixing was developed and applied using MP8L plant parameters. Based on this model, it was shown in Reference 3.6.3 that even after steaming at the very high flow rate assumed in the analysis for twenty minutes, the difference be-tween the average pool surface temperature (131*F) and bulk temperature (107'F) was only about 24'F. We note also that the peak temperatures reported are just barely approaching levels that might imply unstable steam condensa-tion loads; e.g., about 130*F for a straight down pipe. Accordingly, BNL con-cludes that this scenario could safely proceed for as much as twenty minutes without the need for any mitigating action. The GGNS applicant takes the position that because operation in the SCM is so operator intensive, detection of a failure and termination of the steam flow could be accomplished within two minutes. This appears to us somewhat optimistic, but BNL is satisfied that sufficient time is available to institute a number of actions which would effectively mitigate any adverse effects of this postulated failure.
Conclusion The issue raised by the concern is considered to be satisfactorily resolved for the GGNS based on the assumption that post-LOCA operation of the RHR in a steam condensing mode will not occur.
References -
3.6.1 Yang, C.T. Informal Response to Question during NRC/ March III/GE Meet-ing, May 19-20, 1983.
3.6.2 MP&L Letter No. AECM-82/574 dated December 3,1983 from L.F. Dale (MP&L) to H.R. Denton (NRC).
3.6.3 Meeting Handout " Response to Question 9.2," NRC/ Mark III/GE Meeting, May 19-20, 1983, t
I Humphrey Safety Concern 4.3 All Mark III analyses presently assume a perfectly mixed uniform suppression pool. These analyses assume that the temperature of the suction to the RHR heat exhangers is the same as the bulk pool te perature. In actuality, the temperature in the lower part of the pool where the suction is located will be as much as 7-1/2' cooler than the bulk pool temperature. Thus, the heat transfer through the RHR heat exchanger will be less than expected.
Evaluation To complete the statement of this concern, the following should be added; ...
"and containment pressure and temperature greater than expected."
Humphrey's basis for expecting a temperature difference of up to 7-1/2*F is unclear (we assume here that Mr. Humphrey does intend Fahrenheit degrees).
BNL agrees that in the event of a postulated LOCA, the reality will be a ther-mally stratified pool. However, to decide what the difference between bulk and RHR suction temperature is requires an estimate of the degree of vertical stratification that will occur, together with knowledge of RHR suction eleva-tion.
The first of these requirements was establihed to the NRC staff's satisfaction during its evaluation of the GESSAR 11 containment loads (Reference 4.3.1).
After a lengthy, detailed, and sometimes heated review process by the various interested parties (Reference 4.3.2), the worst case _ vertical temperature pro-file proposed by the General Electric Company for design (Fig. 381-3 of Refer-ence 4.3.3) was judged acceptable. The basis for this judgement is given in Reference 4.3.1. It implies that the profile is applicable only for a stan-dard top vent submergence (-7.5 feet).
In responding to Humphrey concerns 4.7 and 4.10 (see later), the applicant in-dicates that the GGNS RHR suction is located at an elevation 10'-6" above the basemat. Comparison with the temperature profile referred to above implies .
that the RHR suction temperature is greater than bulk temperature. According-ly, this Humphrey concern is not relevant to the GGNS.
Conclusion BNL considers this issue satisfactorily resolved for the CPS.
References 4.3.1 " Mark III LOCA-Related Hydrodynamic Load Definition," NUREG-0978, August 1984 4.3.2 Transcript of the ACRS Subcommittee on Fluid Hydraulic Dynamic Effects Meeting of September 24, 25, 1981.
4.3.3 General Electric Co., 22A707, " General Electric Standard Safety Analy-sis Report," (GESSAR-II), Appendix 3B through Amendment 1, February 25, 1982.
Humphrey Safety Concerns 4.4 and 7.1 4.4 The long term analysis of containment pressure / temperature response assumes that the wetwell airspace is in thermal equilibrium with the sup-pression pool water at all times. The calculated bulk pool temperature is used to determine the airspace temperature. If pool thermal stratifi-cation were considered, the surface temperature, which is in direct con-tact with the airsjace, would be higher. Therefore, the airspace temper-ature (and pressure) would be higher.
7.1 The containment is assumed to be in thermal equilibrium with a perfectly mixed, uniform temperature suppression pool. As noted under Topic 4, the surface temperature of the pool will be higher than the bulk pool temper-ature. This may produce higher than expected containment temperatures and pressures.
Evaluation The concern is similar to that associated with HSC 4.3 above except that here the issue is the difference between pool surface temperature and pool bulk temperature. Based on the GESSAR-II temperature profile referred to previous-ly, this difference is 8'F, in rough agreement with the 7-1/2*F difference cited by Mr. Humphrey in HSC 4.2. Apparently, this was the AT he was refer-ring to and it was mistakenly cited in connection with the Bulk-to-RHR suction temperature difference.
The applicant's response to this concern is given in Reference 4.4.1. In this submittal, the issue is quantified by means of existing information and analy-ses. The results show that the effects of a 7 to 8'F difference between pool surface and bulk temperature would imply an increase in peak containment pres-sure and temperature of only 0.1 psi and 3*F, respectively. These modest dif-ferences are overwhelmed by the existing margins of 5 psi and 19'F that can be demonstrated tainment to exist response due to various conservatisms used in conventional con-analysis.
Conclusion Essentially, what this entire exercise has demonstrated once more is that the use of a mean or bulk pool temperature is an acceptable simplification which facilitates calculation of containment response. The result is not surpris-ing. BNL considers this issue to be closed.
References 4.4.1 Attachment to MP&L Letter No. AECM-82/353 dated August 19, 1982, from L.F. Dale (MpAL) to H.R. Denton (NRC).
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Humphrey Safety Concern 4.5 A number of factors may aggravate suppression pool thermal stratification.
The chugging produced through the first row of horizontal vents will not produce any mixing from the suppression pool layers below the vent row. An upper pool dump may contribute to additional suppression pool temperature stratification. The large volume of water from the upper pool further sub-merges RHR heat exchanger effluent discharge which will decrease mixing of the hotter, upper regions of the pool. Finally, operation of the containment spray eliminates the heat exchanger effluent discharge jet which contributes to mixing.
Evaluation The applicant's response to this concern was provided in Reference 4.5.1. It is referred to as Action Plant 14 This Action Plan was to utilize the fol-lowing " Program for Resolution."
- 1. Testing information will be submitted to demonstrate the effectiveness of chugging as a mixing mechanism in the suppression pool.
- 2. Analyses will be submitted to demonstrate that the suppression pool does not experience significant stratification during containment spray or fol-lowing upper pool dump.
The first of these items addresses the concern that chugging does not provide any mixing below the top vent. The test information that was supplied indi-cates that this is not correct. Measurements obtained from so-called drag disks that had been installed in the PSTF facility indicated that flow rever-sals occur periodically in both the middle and bottom vents during the chug-ging phase of the steam blowdown. Although this is a qualitatively useful finding, the attempt to quantify pool turnover time from this information can-not be taken seriously. This is because the drag disk device requires careful calibration under even the best of circumstances (steady, uniform flow).
Under the unsteady, highly non-uniform flow conditions that prevail within the vents during chugging, the notion that quantitatively correct values of flow velocities can be obtained using this procedure is not credible. Furthermore, even if one were to accept these quantitative estimates, their applicability for the case involving upper pool dump, which can increase top vent submer-gence from 7-1/2 to as much as 12 feet, would be highly suspect.
As for the second item, no analytical information has been supplied by MP&L.
Accordingly, we have developed a bounding scenario based on information devel-oped from the applicants FSAR and Reference 4.5.2. From these documents and drawings supplied by MP&L (Reference 4.5.3). BNL has determined, among other things, that the GGNS upper pool volume increases the pool depth by about four feet, that the upper pool water at the time of dump is 125'F and that the sup-pression pool bulk temperature at the time of pool dump is in the range 123*F to 127 F. We have also been able to establish that the dump is accomplished by two 30 inch diameter, gravity fed drain lines that terminate about 20 feet above the suppression pool surface. These drain lines are located at azimuth 35' and 313 which places them in close proximity to the RHR suction lines (Azimuth 32 and 328*). Also, the dump is accomplished over a period of about 5 minutes.
If during this time interval no other mass or energy addition to the suppres-sion pool occurs, the combined suppression bulk pool temperature ~would remain approximately the same. Accordingly, there would be no increask in thermal stratification relative to the pool surface-to bulk-temperature difference.
The effect of pool dump on thermal stratification as it relates to RHR suction is more difficult to estimate but as a limiting case we can spaculate that the upper pool water, because of its downward velocity'and greater density, pre-ferentially sinks to the bottom and displaces the hotter stratified layers of pool water. Except for the possibility of shert circuiting which we judge to be unlikely because of the high elevation of the GGNS RHR suction, this "ccid-water-sinks-to-bottom" scenario represents a worst case in terms of the RiiR suction temperature differential from bulk pool temperature. That is, by dis-placing the hotter layers of water upward, the RHR suction temperature is re-duced to the maximum degree possible. Our attached Figure 4.5.1 demonstrates this effect. Note especially that although the local temperature at the RHR ,
suction elevation decreases, it'still remat-rs higher than bulk pool tempera-ture. This represents an important conservatism as was pointed out during evaluation of HSC 4.3 above. Thus, upper pool dump also does nct impact nega-tively on the RHR suction to bulk temperature difference in the GGNS.
The remaining concerns implied in the statement of this Humphrey concern are also taken account of in our evaluation here. This is because our arguments have used the worst case thermal stratification; i.e.: the temperature pro-file was developed without assuming RHR operation. Thus, any RHR operation, however inefficient, will further improve the situation as we have character-ized it here.
Concl usion BNL judges that upper pool dump in the GGNS will not seriously increase pool thermal stratification. This is because the difference between upper pool and suppression pool bulk temperature at time of dump is slight and because the RHR suction in the GGNS is located at a high elevation (10.5 feet above the basemat) relative to the worst case temperature profile. ,
References 4.5.1 MPAL Letter No. AECM-82/353 dated August 19, 1982, from L.F. Dale (MP&L) to H.R. Denton (NRC).
4.5.2 Gunter, A.D. and Fuls, G.M., "Clasix-3 Containment Response Sensitivity Analysis for the Micsi sippi Power & Light Grand Gulf Nuclear Station,"
Offshore Power Systems Report No. OPS-37A15, Dec. ember 1981.
4.5.3 Attachment 4 to MPAL Letter xNo. AECM-82/497 dated October 22, 1982 from L.F. Dale (MPAL) to H.R. Denton (NRC).
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Humphrey Safety Concerns 4.1 and 4.10 4.7 All analyses completes for the Mark III are generic in nature and do not consider plant specific interactions of the RHR suppression pool suction and discharge.
4.10 Justify that the current arrangement of the discharge and suction points of the pool cooling system maximizes pool mixing.
Evaluation The concern here is that if the RHR system's geometric arrangement for the suction and return lines is not properly designed, the capability of the sys-tem to induce bulk mixing and remove thermal energy will be degraded.
The applicant has addressed these concerns via the generic approach developed for the Mark III Containment Issues Group (Reference 4.7.1). The key element of this study was the Perry one-tenth-scale tests (Reference 4.7.2). In these tests, a r.0mber of related concerns were addressed systematically. These in-cluded short circuiting, development of bulk pool motion, ability to eliminate thermal stratification and the presence of isolated recirculation zones.
BNL has reviewed this information in detail and has concluded that the Perry one-tenth-scale tests correctly simulate design conditions in Mark III plants, and, along with the other test information provided, include a sufficient l range of parameters to encompass the GGNS plant unique features, and are therefore applicable. Since the findings from these tests show that good mix-ing can be achieved, as well as the absence of short circuiting, BNL concludes that the GGNS RHR system can be expected to perform in a manner consistent with design assumptions. ,
Conclusion BNL considers the issues raised by these concerns to be satisfactorily re' .
solved for the GGNS.
References 4.7.1 Quadrex Corp., "A Survey of Tests and Analyses on the Effectiveness of the RHR System in the Pool Cooling Mode," Report No. QUAD-1-82-245.
4.7.2 Gilbert Associates, Inc., "Model Study of Perry Nuclear Power Plant Suppression Pool - Final Report," November 1977.
Humphrey Safety Concern 19.1 The chugging loads were originally defined on the basis of 7.5 feet of submer-gence over the drywell to suppression pool vents. Following an upper pool dump, the submergence will actually be 12 feet which may affect chugging loads.
Evaluation The applicant's response to this concern is provided in Reference 19.1.1. In this submittal, physical arguments and analytical procedures are used to esti-mate the pressure field that would be generated on the suppression pool bound-artes if the worst case chug from the Mark III data base were to occur with the top vent at a 12-foot submergence. The results are compared with design on an ARS basis and shown to be bounded except for local loads in the frequen-cy range 15 to 32 Hz. For these conditions an exceedance of design amounting to 35% occurs on the basemat and about 15% on the containment wall.
The applicant argues that this exceedance is not important because this is a local load affecting only the basemat and containment wall liners and that, because of the hydrostatic head to which the liner is subjected, "a negative pressure will never be imposed on the liner." Also, "since the liner is backed by concrete, no natural modes of vibration are excitable."
Without passing judgement on the merits of these arguments, BNL notes the fol-lowing: the use of an acoustic model in the analysis represents a significant conservatism; dissipative mechanisms not accounted for in such an analysis re-sult in pressure attenuation which is much greater than predicted; this has been borne out convincingly by experimental results: application of the worst case chug to all vents, which is done for local loads, also represents a very significant conservatism; in a recent submittal by the Cleveland Electric Illuminating Co. to address the concern relating to the combined effect of upper pool dump and encroachment on local chugging loads (Reference 19.1.2),
it was shown that by postulating a maximum strength chug at the central vent .
and average strength chug at adjacent vents, the design loads were capable of bounding the combined effect.
In summary, the margins inherent in the design load for chugging are very large. They can more than accommodate any increment in loading caused by off-design effects such as increased submergence due to upper pool dump.
Conclusion
! BNL considers this issue satisfactorily resolved for the CPS.
References 19.1.1 MPAL Letter No. AECM-82/353 dated August 19, 1982 from L.F. Dale, MP&L, to H.R. Denton, NRC.
19.1.2 CEI Letter dated July 11, 1984 from M.R. Edelman, CEI, to B.J.
Youngblood, NRC.
Ramphrey Safety Concern 19.2 The effect of local encroachments on chugging loads needs to be addressed.
Evaluation The applicant's response to this concern is provided in Reference 19.2.1. In this submittal, physical arguments and analytical procedures are used to esti-mate the pressure field that would be generated on the suppression pool bound-aries if the worst case chug from the Mark III data base were to occur at vents located below the GGNS TIP platform. The results pre compared with de-sign on an ARS basis and shown to be bounded except for local loads in the frequency range 12 to 30 HZ. For these conditions an ex:eedance of design amounting to 60% occurs on the basemat and about 15% on the containment wall.
The applicant argues that this exceedance is not important because this is a local load affecting only the basemat and containment wall liners and that, because of the hydrostatic head to which the liner is subjected, it will not experience a " negative pressure in the frequency of exceedance." Also, "since the liner is backed by concrete everywhere, no natural modes in this range are excitable."
Without passing judgement on the merits of these arguments, BNL notes the fal-lowing: the use of an acoustic model in the analysis represents a significant conservatism; dissipative mechanisms not accounted for in such an analysis re-suit in pressure attenuation which is much greater than predicted; this has been borne out convincingly by experimental results: application of the worst case chug to all vents below the encroachment also represents a very signifi-cant conservatism; in a recent submittal by CEI to address the staff concern relating to the combined effect of upper pool dump and encroachment on local chugging loads (Reference 19.2.2), it was shown that by postulating a maximum strength chug at the central vent and average strength chug at adjacent vents, the design loads were capable of bounding the combined effect.
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Conclusion BNL is satisfied that the issues related to this concern have been satisfacto-rily addressed by the applicant and are therefore considered closed.
References 19.2.1 MP&L Letter No. AECM-82/574 dated December 3,1983, from L.F. Dale (MP&L) to H.R. Denton (NRC).
19.2.2 CEI Letter dated July 11, 1984 from M.R. Edelman (CEI) to B.J.
Youngblood (NRC).
Additional Safety Concern BNLl*
The effects of increased submergence and encroachment on local loads are addi-tive. ARS comparisons showing the combined effect should be provided.
Evaluation A generic response was developed for this concern and included in a Cleveland Electric Illuminating Co. (CEI) submittal (Reference B.1.1). As indicated in the evaluation provided for HSC-19.1 and 19.2 above, by postulating the very reasonable assumption that a maximum strength chug occurred only at a central vent with average strength chugs at surrounding vents, local loads were re-duced to levels which were adequately bounded by design.
Conclusion The utilities have provided an adequate demonstration that the design load for chugging has sufficient margin to accommodate the increment in loading caused by the combined effects of encroachment and increased submergence. BNL con-siders this issue to be closed.
References B.1.1 CEI Letter No. PY-CEI/NRR-0123L dated July 11, 1984 from M.R. Edelman (CEI) to B.J. Youngblood (NRC).
- This additional concern, which evolved from HSC19 and was developed by BNL, is not formally recognized as a Humphrey concern. It is therefore denoted by this special notation.
Additional Safety Concern BNL3*
Upper pool dump increases the length of the water column within the main steam SRVDL. This will tend to increase pipe thrust loads during SRV actuation.
Provide an analysis of this effect.
Evaluation The applicant states in Reference B.3.1 that these thrust loads have been re-calculated. These revised loads are stated to be "within the upset allowable stresses."
Conclusion Based on the statement made by the applicant, BNL considers this concern to be resolved.
References B.3.1 Attachment 1 to MP&L Letter AECM-83/0146 dated March 23, 1983 from L.F. Dale (MP&L) to H.R. Denton (NRC).
- This additional concern, which evolved from HSC19 and was developed by BNL, is not formally recognized as a Humphrey concern. It is therefore denoted by this special notation.
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