ML20207S520
| ML20207S520 | |
| Person / Time | |
|---|---|
| Site: | Grand Gulf, River Bend, 05000000 |
| Issue date: | 11/30/1986 |
| From: | Economos C BROOKHAVEN NATIONAL LABORATORY |
| To: | NRC |
| Shared Package | |
| ML20207S501 | List: |
| References | |
| CON-FIN-A-3346, CON-FIN-A-3396 NUDOCS 8703190599 | |
| Download: ML20207S520 (27) | |
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- 1 Technical Evaiuation Report (TER) for the Gulf States Utilities Company's (GSU)
River Bend Station (RBS), Docket No. 50-458.
on SER Outstanding Issue (8) -
Mark III Containment System Issues -
The Humphrey Safety Concerns (HSC) by C. Economos November 1986 Department of Nuclear Energy Brookhaven National Laboratory Upton, New York 11973 Introduction With one exception, this TER addresses only those HSCs that, by previous agreement with the Technical Monitor, are considered BNL's responsibility.
The exception involves HSCs 4.4 and 7.1. In an earlier version of .this TER this issue was considered the responsibility of the NRC Technical Monitor.
Notwithstanding this earlier decision, we feel that BNL input in this area is appropriate and have included our evaluation here. Our finding is that the concern has been satisfactorily resolved.
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- . 2 Humphrey Safety Concerns 1.1, 1.2, 1.4~and 1.5 1.1 Preser,:e of local encroachments, such as' TIP platform, the drywell per-sonnel 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 height 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 (C10G), it has now been established that, for the most part, the effect of encroachments is to decrease pool swell velocities rela-tive to an unencroached or clean suppression pool. BNL has reviewed all of this infomation and has prepared a separate Technical Evaluation Report on its 'indings (Reference 1.1.1).
Conclusion BNL considers these issues closed for the RBS because the 1/10 scale tests have confimed that encroachment effects on pool swell behavior will not give rise to loading conditions that cannot be accommodated by'the existing design basis.
References 1.1.1 Sonin, A. and Economos, C., " Resolution of the Humphrey Issues Relating to Pool Swell in Mark III Plants," BNL Technical Evaluation Report, February 1985.
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c 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.
A two-dimensional SOLAV01 simulation was employed to detennine the effect of RBS enchroachments on these loads (Reference 1.3.1). An increase in the con-tainment wall load was predicted which is within the design load specifica-tion. The submerged structure loads are developed using the velocity field deriving from the SOLAV01 simulation. The increased loads produced by the en-croachments are stated to be bounded by the design basis LOCA bubble loads.
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 Enclosure to GSU Letter No. RBG-24,131, Dated July 31, 1986 from J.E.
Booker (GSU) to H.R. Denton (NRC).
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 0.85 psid on the beams and 0.24 psid on the gratings.
The additional stresses due to these lateral forces were found to be small fractions of the total stresses.
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 at-tained. Considering that the additional stresses are modest even with a con-servative flow analysis, BNL does not feel that additional effort needs to be expended on this issue.
Conclusion BNL considers this issue closed.
References 1.6.1 Enclosure to GSU Letter No. RBG-19,972 dated January 23,1985 from J.E. Booker (GSU) to H.R. Denton (NRC) .
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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 In the RBS the 1500 square feet open area requirement is satisfied everywhere except at the refueling floor which is at the 186 foot elevation (Reference 1.7.1). Here the opening is 689 square feet. The utility has conducted an analysis (using three volume nodes) which indicates that the resulting maximum pressure differential on this floor is 0.478 psid and the increase in the dry-well pressure is 1.49%.
BNL agrees with GSU that the effect of the smaller than 1500 sq. ft opening at the 186 foot level is very small. Since this elevation is 72 feet above the HCU floor and no froth will reach this level, the flow resistance provided by the 689 square foot opening to air is very much smaller than the resistance provided by the larger HCU floor opening to the two-phase froth.
Conclusion, BNL considers this issue closed, based on the above response.
References 1.7.1 Enclosure to GSU Letter No. RBG-19,972 dated January 23, 1985 from J.E. Booker (GSU) to H.R. Denton (NRC).
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Huriphrey Safety Concern 2.1 I
l The annular regions between the safety relief valve lines and the drywell wall
[ penetration sleeves may produce condensation oscillation (CO) 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 first addressed this concern using the generic approach de-scribed in Reference 2.1.2. This methodology derives sleeve annulus C0 loads by 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 sl eeves .
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.
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 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 Mark I FSTF data base (2.1.5). This new design loading results i . 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
specification. Da the containment wall, the bound is provided, with consider-able margin, by the pool swell boundary load.
The applicability of the FSTF data to the sleeve annulus loads involves con-siderable uncertainty because of the great disparity in geometries between the two situations. This applies not only for the steam-water interface at the respective 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 completely apparent that for the FSTF case the system has actually achieved a condition of resonant coupling. This is because of the complexity of the FSTF vent system; i.e. the eight downcomers are connected to a vent' header,- which in turn 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 preci-sion.
Despite these uncertainties, there are several factors that may be cited that compensate for any possible inadequacy. First, there is the qualitative evi-dence from the 4TC0 tests (Reference 2.1.3) that, even when resonant coupling clearly occurs, the load amplification is limited to less than a twofold fac-tor. Also, the generally conservative application of the available results provides increased confidence in the adequacy of the load method. For exam-ple, the pressure results observed on 24" diameter downcomers are taken over directly and applied to the much smaller,14" diameter, SRVDL sleeves. Al so ,
in 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 the steam bubble were modeled more realistically (Reference 2.1.2) . Additional conservatism stems from the use of conventional acoustics for determination of pressure attenuation from the source to the pool boundaries. Dissipative mechanisms that are present in the suppression pool and neglected in the analysis, would further reduce the loads.
Insofar as the chugging loads are concerned, BNL has not received a descrip-tion of these, even though this information exists (Reference 2.1.7) and has been assessed 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 sat-isfied 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.
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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._ i 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. Spangenber{}
(IPC) to W.R. Butler (NRC). ,
2.1.5 Fitzsimmons, G.W. et al., " Mark I Containment Program - Full-Scale Tes?.]
Program Final Report", General Electric 1 Report, NEDE-24539-P, April' 1979. \
2.1.6 IPC Letter No. U- dated from F.A. Spangenburg (IPC) to W. R. Butler (NRC). J' 2.1.7 Enercon Letter No . RWE-0G-060 dated May d5, 1983 from R.W. Evans
'C (Enercon) to B. R. Patel (Creare R&D).
t 2.1.8 Mark III Containment Issues Review Panel, " As cerns", CREARE R&D, Inc., Technical Memoranduy.ussment TM-928, July 1934. of H:,mphray : Con-c 1.. 'g.
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Humphrey Safety Concerns 2.2 and 7.3
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2.2 The potential condensatiop oscillation and chugging loads produced
, through the annular area (etween the SRVOL and sleeve may apply un-I accounted for loads 9 - the; MVDL. Since the SEVDL is unsupported from the quencher to the ins. de,0f the drywel'l wall, this may result in fail-ure of the line. 'C -
2.3 The potential condensatic.n oscillation and chugging loads produced through the annular area between the SRVDL and sleeve may apply unac-counted for loads t9 th penetration sleeve. The loads may also be at or s near the naturaK frequency o the sleeve.
Evaluation -
t s The concern here is that tie steam condensation process (C0 + chugging) at the sleeve annulus exit will give rise to loads on the SRVDL and SRVDL sleeve i analogous to the lateral loads experienced by Mark I and Mark 11 downcomers during' postulated LOCA blowdowns and that these structures have not been de-signed tC accommodate them.
(' s The appl'1 cant's specification for chugging loads i's given in Reference 2.2.1.
The,' load has a half-sinusoidal i time dependence with a duration of 3 mseconds And a peak amplitude of 22 Kips. It is stated that this load derives from the 3 ark 11 load mythodalogy of Reference 2.2.2 as modified by the NRC Staff's Acceptance Criteria (Reference 2.2.3). The load is developed by scaling down the peak amplitude to the outside diameter of the SRVDL sleeve and accounting i for the fact that there are fewer chug sources created by flow through the SRVDL sleeve annulus than exist during DBA blowdowns through the Mark 11 pres-sure suppression system (i.e.,16 SRVs for the CPS vs. about 100 downcomers in a typical Mark 11 plant). Scaling down 'for pipe diameter is accomplished by ass 3ning a 1.7 power dependence of the peak load amplitude on diameter. Load reduction for fewer chug sources utilizes the staff approved statistical rep-rescatation for these loads (Reference 2.2.3). The region of application of thi load is alsc scaled down using a first power dependence on diameter. It
\in stated thtc the stresses / loads resulting from application of these loads dre either bounded by other design loads or less than the pertinent allowa-t,l es .
The applicability of the Mark 11 results for the present application is some-what uncertain due to the substantial geometric differences (straight down g vs. inclined pipe and annular vs. circular cross section). Nevertheless, we find the approach reasonable and, in general, conservative. The use of a 1.7 power dependence of peak amplitude on pipe diameter is somewhat less conserva-tive than we would have preferred since the available data (Reference 2.2.4) exhibit exponent values that range from 0.7 to 1.7 On the other hand, no credit is taken for the presence of the SRVOL in the steam bubble. This pro-vides a substantial conservatism that BNL judges more than compensates for any possible non-conservatism in selecting this exponent.
No C0 lateral loads are specified by the applicant. In Reference 2.2.1 it is argued that since the suppression pool end of the SRVDL sleeve is truncated perpendicular to the discharge line axis, the potential for asymmetric dynamic pressure loads arising at that end is precluded. BNL believes that this
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al load specification on the SRVDL and SRVDL sleeve during postulated LOCA conditions.
Conclusion ,'
BNL considers this concern to be' resolved because of the specification of a conservative chugging load for the SRVOL and SRVOL sleeve and because the RBS configuration for these structures does not permit development of any signifi-cant unbalanced pressure loading during the C0 phase of the LOCA blowdown.
References 2.2.1 Enclosure to GSU Letter No. RBG-24,131 dated July 31, 1986 from J.E.
Booker (GSU) 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 Prograin Load Evaluation and Accep-tance Criteria," NRC NUREG-0808, 2.2.4 General Electric Letter MFN-050-80 dated February 29, 1980 from R.H.
Buchholz (GE) to C.J. Anderson (NRC).
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Humphrey 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.
3.2 Discharge from the RHR relief valves may produce bubble discharge or other submerged structure loads on equipment in the suppre,ssion 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 (MSRVs), 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 LOCA 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 pre surized to about 200 psi by a pressure control valve (PCV). Should the PCV fail, resulting in elevated pressures, the heat exchanger 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. Also, under normal operating conditions, small amounts of steam and noncondensibles are continously bled from the heat exchanger into the re-lief valve discharge line to maintain heat exchanger efficiency. As stated in Reference 3.1.1, "The flow rate ... is such that chugging will occur in the pool as long as the RHR system operates in this mode.
The applicant has supplied a plant unique response for RHR discharge pipe water clearing jet loads and for air bubble loads, including submerged struc-ture loads. The methods used are conventional and correspond, for the most part, to GESSAR and/or FSAR methods. Details of these analyses and procedures are given in Reference 3.1.2. The applicant states that, for the most part, the loads are negligible or bounded by MSRV and/or LOCA loads. Where loads were not bounded, the affected structures were reanalyzed to confirm their capability.
For C0 and chugging loads on pool boundaries and for submerged structures, the applicant utilized the generic methods employed by all the Mark III utili-ties. A detailed description of these methods is given in References 3.1.1 and 3.1.3. Generally speaking, the method derives from conservative applica-tion of the Mark II C0 and chugging load methodologies. Source terms are de-veloped from chugging and C0 pressure signatures selected from the Mark 11 data base for their conservatism. These source terms are applied without any modification to account for the difference in pipe diameter between the RBS relief valve discharge line (12 inches) and that from which the data base de-rives (24 inches in the test facility). 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 adequacy of the C0 load also needs to be judged in the context of the po-tential for unstable steam condensation; i .e. , el evated pool temperatures.
BNL's evaluation of this aspect is presented later under Humphrey Concern 3.6.
A detailed description of lateral loads on the RHR discharge line due to c:ug-ging is given in Reference 3.1.2. The load is time dependent (triangular) with a peak amplitude of 22.2 Kips and a duration of 3 ms. It is uniformly distributed over an application region extending 0.53 to 2.13 ft from the dis-charge pipe exit. It is stated that this load specification derives from the Mark 11 chugging lateral load definition on the containment downcomers. It differs in that it utilizes a triangular impulse rather than the half-sinusoid employed by the original methodology (Reference 3.1.4). Also, it differs in that the peak amplitude and region of application have been scaled down to account for the difference in pipe diameter between the RHR discharge line (12 inches) and the standard Mark II downcomer (24 inches). This is accomplished by assuming peak amplitude to scale with the 1.7 power of pipe diameter and the application region with the first power. No rationale for changing the impulse shape is provided despite the fact that the total impulse is reduced about 25%.
In Reference 3.1.2 it is stated that application of these loads lead to stresses which are acceptable relative to allowables. Also, pipe supports and penetrations were found to be qualified.
In BNL's judgement the lateral load specification proposed by GSU has several important deficiencies. We have already noted one of them above, namely, the use of a triangular impulse function without any justification for this choice. In addition, the peak amplitude used by the applicant, is not totally consistent with the stochastic nature of the chugging phenomenon as exempli-fied in the NRC approved load method. Here we refer to the fact that the peak load amplitude depends on the total number of chugs that would be expected during a particular accident scenario and the desired non-exceedance probabil-ity. As an example, for the DBA LOCA in a March II plant with a population of 100 downcomers, a peak design load of 65 Kips was employed to insure that sta-tistically, no member of that population is likely to experience an exceedance of this load during the 100 or so chugs expected during this LOCA. Since it was this value that was scaled down to derive the 22.2 Kip design amplitude, it follows that this choice is acceptable only for a structure that will not experience more than about 100,000 chugs. The applicant has not given us any l way of estimating whether this is adequate or not. A final deficiency of the l
load specification, as we see it, is the failure to recognize the randomness of chugging lateral loads in terms of their direction. For a straight down pipe this could be accounted for simply by applying the load normal to the
! pipe axis in the direction which would maximize stresses in any bracing or support structures as may exist. For the RBS discharge pipe however, since it l is equipped with a tee at the end, the proper way to account for directional
! randomness is to apply the load in such a way that both axial and torsional l stresses arise in addition to those usually associated with lateral (w.r.t.
the pipe axis) application of the load. According to Reference 3.1.5 such an
13 application may have been done but, to date, this has not been confirmed in any way.
In order to resolve this issue, the applicant needs to provide some rea-sonable upper bound estimate of the number of chugs, N, that the RHR discharge pipe is likely to experience during the life of the plant. For this purpose a chugging period of one second would be acceptable. Note that this is double the one used by the Containment Issues Owners Group in Reference 3.1.6. With N established, a suitable value of the peak load amplitude for application to the RHR discharge pipe, F, would follow from the relation F = 1.88 in N (a)
This relation derives from Equation (6) of Reference 3.1.7 by scaling down for pipe diameter and assuming that the desired non-exceedance probability, P, is equal to the reciprocal of N. To provide some indication of the order of mag-nitude of these parameters, let us evaluate the value of F needed to insure non-exceedance for the case where chugging occurs continuously throughout a 40 year plant life!! This corresponds to slightly more than one billion chugs (N 2 1.3x109) . The corresponding value for F derived from relation (a) is then about 40 Kips. Thus, with less than a two-fold increase in the peak load am-plitude, the number of chugs the discharge pipe could experience without ex-ceedance is increased by four orders of magnitude. From this perspective it would not appear that the applicant should have any difficulty resolving this concern in terms of peak amplitude.
Loads due to other fluid 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 and fluid state. A list of all such relief lines was provided in Reference 3.1.2 together with data characterizing their size and flow rates. One dis-charge line which has some potential for creating significant loading is the RCIC turbine exhaust line. In this case the flow rate / diameter combination results in steam flux rates that are in the chugging regime. Significant loads due to the reflood and air / water clearing loads could also occur since this pipe has a relatively large diameter.
The concern relating to the RCIC discharge is addressed by the applicant by citing Reference 3.1.8. Presumably, the material contained therein is the same as the " white paper" we have referred to in our earlier TERs (Reference 3.1.9). This material is intended to support the position that operating ex-perience has shown that there are no dynamic load problems associated with operation of the RCIC system.
Conclusion This issue is considered closed for water clearing, air bubble and C0 loads based on the conservatisms used to develop them and/or the wide margins exhib-ited relative to other design loads.
For chugging type lateral loads on the RHR discharge line, the applicant needs to clarify how the selected design loads have been applied and provide a clearer picture of the loading that will be experienced by this structure
. -14 during the life of the plant. Otherwise, specification of a 40 Kip load ap-plied laterally and axially and as a torque would suffice to close this issue.
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. l 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) . l 1
3.1.2 Enclosure to GSU Letter No1 RBG-24,131 Dated July 31, 1986 from J.E.
Booker (GSU) to H.R. Denton.
3.1.3 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.
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3.1.4 Anderson, C., " Mark 11 Containment Program Load Evaluation and Accep-tance Criteria," NRC NUREG-0808.
3.1.5 Communication Provided During NRC/BNL/GSU Telephone Conference, Septem-ber 9, 1986.
3.1.6 tiark III Containment Issues Review Panel, " Assessment of Humphrey Con-cerns," CREARE R&D, Inc., Technical Memorandum TM-928, July 1984 3.1.7 Lehner, J.R. and Sonin, A. A., " Determining a Lateral Load Specification During Chugging in a Mark II Containment," Structural Mechanics in Reactor Technology, Vol. J, August 22-26, 1983, pp. 75-80.
3.1.8 Enercon Services, Inc. , Letter No. JRC-0G-142, from J.R. Corn to L.A.
England, History of Development of the RCIC Exhaust Sparger, Dated April 10, 1985.
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Humphrey Safety Concern 3.2 The STRIDE design provided only 9 inches of submergence above the RHR heat ex-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 to 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.
The applicability of this data base to the RBS is somewhat questionable since, by virtue of the tees that have been installed, the steam discharge will be horizontal rather than vertical. It is stated, however, that the minimum sub-mergence of the discharge lines will be about 4 feet. With this kind of sub-mergence and the level of subcooling that would be expected during steam dis-charges through these lines (see discussion of HSC 3.6), the potential for steam bypass is, in BNL's opinion, non-existent.
Conclusion BNL considers this issue satisfactorily resolved for the RBS based on the rel-atively large minimum submergence that exists in this plant.
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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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 vac-uum 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 relief 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 MSRVs. Be-cause the potential for subsequent actuations was not fully appreciated in the early stages, the MSRV discharge lines were originally equipped with under-sized 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 2-6 inch diameter).
In Reference 3.4.1, the applicant indicates that the reflood analysis was car-ried out using an existing GE model (Reference 3.4.2). Additional details used in the analysis are presented in Reference 3.4.3. No indication is given there whether credit is taken for the effect of the bleed flow that keeps the pipe pressurized. In Reference 3.4.4 it is suggested that this is the case.
However, the applicant has provided informal assurance that this is not (Ref-erence 3.4.5).
In References 3.4.1 and 3.4.3, the applicant indicates that the RBS RHR heat exchanger discharge lines are equipped with 3/4 inch vacuum breakers. We con-sider this somewhat inadequate for the 12 inch diameter line involved. Thi s is reflected by the results reported by the applicant which indicate that max-imum reflood lengths exceeding 50 feet. This is a surprisingly high value but, according to the applicant, well below any crucial element of the piping system. Specifically, the SRVs are about 12 feet above peak reflood elevation
( while the vacuum breakers themselves, although at about the same elevation, are only partially flooded by the rising water leg (Reference 3.4.5). In any
, case, the piping and piping support dynamic loads were evaluated for a second l or subsequent actuation at this peak reflood using, once again, standard meth-ods that are acceptable to the NRC (Reference 3.4.6). It is stated that the structures were found to have sufficient margin to accommodate these loads.
Conclusion l
l BNL considers this issue adequately addressed by the applicant and therefore closed.
References I
l 3.4.1 Enclosure to GSU Letter No . RBG-19,972 Dated January 23, 1985 from l J.E. Booker (GSU) to H.R. Denton (NRC).
I l
17 3.4.2 Wheeler, A.J., Dougherty, D.A., " Analytical Model for Computing Water Rise in Safety Relief Valve Discharge Line Following Valve Closure," GE Document No. NEDE-23898-P, Dctober 1978.
3.4.3 Enclosure to GSU Letter No. RBG-24,131 Dated July 31, 1986 from J.E.
Booker (GSU) to H.R. Denton (NRC).
3.4.4 Mark III Containment Issues Review Panel, " Assessment of Humphrey Con-cerns," CREARE R&D, Inc., Technical Memorandum TM-928, July 1984.
3.4.5' Communication Provided During NRC/BNL/GSU Telephone Conference, Septem-ber 9,1986.
3.4.6 SWEC Computer Code STEHAM, FSAR Appendix 3A.
I L-
Humphrey Safety Concern 3.6 If the RHR heat exchanger relief valves discharge steam to the upper levels of the suppression pool fcilowing a design basis accident, they will significant-ly aggravate suppression pool temperature stratification.
Evaluation As stated above, the concern is that because of the presence of a layer of heated water near the pool surface, the pressure and temperature response in the containment will be underestimated by any analysis which assumes equality between the wetwell air space and the suppresion pool bulk temperatures. In this sense, the concern is similar to that expressed via HSC 4.4.
This aspect of the concern has been addressed by the applicant via the generic approach first provided under the MP&L docket (Reference 3.6.1). A demonstra-bly conservative model of thermal deposition, pool mixing and thermal strati-fication was developed and applied using GGNS plant parameters. Based on this model, it was indicated in Reference 3.6.2 that during the first 15 minutes or so of this steam discharge, the difference between the average pool surface temperature and pool bulk temperature increases at less than 1*F per minute.
In fact, at 10 minutes into the blowdown this difference is only 7-1/2*F.
Since the RBS analysis assumes that the surface temperature of the pool is always 5*F greater than the bulk temperature (Reference 3.6.3) the resulting non-conservatism will be minimal and easily bounded by the margins that have been demonstrated to exist in heat exchanger performance (Reference 3.6.4).
Based on our review of the available information we conclude that the results described above are equally applicable to the RBS. In fact, they are probably conservative since the RBS discharge line is equipped with a tee. This pro-vides improved mixing compared to steam discharge through a straight down pipe as assumed in the analysis. Thus, so long as the blowdown does not proceed for much more than ten minutes, we would not expect any severe adverse effects to result from this postulated accident scenario. The applicant takes the position (Reference 3.6.5) that operation in the SCM is operator intensive, so that detection of a failure and termination of the steam flow could be accom-plished within two minutes. This appears to us somewhat optimistic, but shut-down within ten minutes is, in our judgment, a conservative and reasonable estimate. Accordingly, we conclude that this issue is resolved for the RBS.
Despite the above conclusion, this concern is not entirely resolved. As noted under HSC 3.1, this relates to the potential for unstable steam condensation due to elevated pool temperature. For this case, the behavior of local pool temperature rather than average pool surface temperature is of interest.
According to the analytical model, the trend of this parameter depends on the total pool volume that participates in direct mixing with the steam jet. For the " worst case" estimate (mixing volume equal to 10% of the total), a local temperature rise of about 6*F per minute is predicted. This trend drops off linearly with assumed mixing volume size (i.e., 3 F per minute for 20% mixing voluma) so that a certain amount of uncertainty exists in evaluating this trend.
Based on the information we have available to us, we believe the worst case trend of 6 F/ min. to be overly conservative. For example, the response l
l L
observed in the Quad City plant during a SRV blowdown through a ramshead dis-charge device (Reference 3.6.6) averaged only 1 F/ min over the first ten minutes of the steam discharge. This was with a steam flow rate double that used in the current analysis.* In part, the great disparity between this re-sult and the analycis can be attributed to the difference in mixing geometry represented by tne ramshead in the first c;>e and the straight down discharge pipe as assumed in the analysis. The stiong influence of steam jet orienta-tion on temperature response is demonstrated by the results observed with SRV discharges through a quencher device (Reference 3.6.7). In this case, with radially directed steam flow, the temperature increased at- an average rate of 8'F/ min. with only a modestly higher flow rate (about 50%) than in the Quad City tests.
The Quad City results should be reasonably applicable to the RBS despite the fact that the RBS discharge orientation differs somewhat (by 25') from the ci rcumferential discharge used in the Quad City plant (Reference 3.6.8).
Note that this difference causes only a 9% reduction in circumferential momen-tun flux. In any case, qualitatively, we would expect an even slower temper-ature rise rate in the RBS because of the lower steam discharge rate. How-ever, in the absence of any quantitative information connecting temperature response to steam flow rate, we take the position that this represents a mar-gin of safety for the present application. This ' margin is needed to cover other uncertainties that exist such as differences in pool geometry such as that cited above.
In a cover letter to the applicant's most recent submittal (Reference 3.6.9) it is indicated that SCM operation is contemplated with suppression pool local and bulk temperatures "below 130*F." This represents a unique proce-dural change relative to all other Mark -II Utilities. These have committed to non-use of the SCM under post-LOCA conditions. The TERs we prepared for the others (Grand Gulf, Clinton and Perry) have based resolution of this issue on such a commitment.
To address the issue for the RBS, we have had to reexamine the NRC staff post-tion relative to pool temperature limit for a rams-head type device. To quote from Reference 3.6.10, this position is "... a suppression pool temperature limit has not been adequately established for the ramshead device." However, in the same reference it is further stated that "the limited amount of plant operational data may be considered as supporting data for some specific zones of mass flow and pool temperature."
The -limited plant data referred to is summarized in Figure 3.6.1 which has been excerpted from Reference 3.6.10. The notation " demonstrated" signifies that stable steam condensation was attained everywhere up to the indicated line.
2 In particular, at the steam flux rate of interest here (about 50 lbs/
ft /sec) stable steam condensation would be expected up to a temperature of 150*F.
For the reasons we have already enumerated, we expect that any steam discharge through the RBS RHR heat exchanger SRV will result in no more than a 1 F/ min
- The parameter that controls mixing is the steam momentum. However, since the exit flow areas and reservoir conditions are all roughly comparable the mass flow rate is equivalent.
. local temperature rise. Also, we do not expect this discharge to proceed for longer than ten minutes. Accordingly, if failure of the PCV valve occurs dur-ing SCM operation at pool temperatures below 130'F, we would not anticipate the occurrence of containment loads associated with unstable steam condensa-tion due to elevated pool temperature.
Conclusion The issue raised by this concern is considered to be satisfactorily resolved for the RBS. In particular, operation in the SCM would be acceptable at local suppression pool temperatures up to but not exceeding 130*F.
References 3.6.1 MP&L Letter No. AECM-82/574 Dated December 3, 1983, from L.F. Dale (MP&L) to H.R. Denton (NRC).
3.6.2 Meeting Handout " Response to Question 9.2," NRC/ Mark II/GE Meeting, May 19 and 20, 1983.
3.6.4 Attachment to MP8L Letter No. AECM-82/353 Dated August 19, 1982, from L.F. Dale (MP&L) to H.R. Denton (NRC).
3.6.5 Enclosure to GSU Letter No. RBG-19,972 Dated Januray 23, 1985 from J.E. Booker (GSU) to H.R. Denton (NRC).
3.6.6 Dougherty, D.A., " Suppression Pool Temperature Response to a Safety /
Relief Valve Discharge Through a Ramshead in Mark I and II Contain-ments," General Electric Co., Report NEDC-23689-P, Class III, March 1978.
3.6.7 .Patterson, B.J., "Monticello T-Quencher Thermal Mixing Test Final Report," General Electric Co., Report NEDC-24542-P, Class III, April 1979.
3.6.8 Sketches dated November 15, 1983 provided by the NRC to BNL via Fac-simile Phone Transmittal dated November 24, 1986.
3.6.9 Enclosure to GSU Letter No. RBG-24.131, Dated July 31, 1986 from J.E.
Booker (GSU) to H.R. Denton (NRC).
3.6.10 " Mark II Containment Lead Plant Program Load Evaluation and Acceptance Criteria," NUREG-0487, October 1978.
40 tsM M GE's RECU MENDATION l FT2-SEC
~
- \
1 Jg 160*F 4
1 -
150 - 6
<~ ',
I C s Q i-N
- END(C) *\
{ \s s EMONSTRATED s
a
\. f y 100 , .
START (C)
C START (A)
~
END (A) PLANT C PLANT A
= -
- r
- 2 -
h
~
g , , .a 50 . . . . i , , , , l , , , , , , , ,
0 50 100 150 200 STEAM lhSS Flux, (lBi/FT2-SED j Figure 3.6.1 Plant SRV Operational Data.
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 exchangers is the same as the bulk pool temperature. 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 established to the NRC staff's satisfac-tion during its evaluation of the GESSAR II containment loads (Ref. 4.3.1).
After a lengthy, detailed, and sometimes heated review process by the various 2
interested parties (Ref. 4.3.2), the worst case vertical temperature profile proposed by the General Electric Company for design (Fig. 381-3 of Ref. 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 standard top vent submergence (-7.5 feet) .
In responding to HSCs 4.7 and 4.10 in Reference 4.3.4, the applicant indicates that the RBS RHR suction is located at an elevation 3'-4-3/4" above the base-mat. Comparison with the temperature profile referred to above implies a tem.-
perature difference (Bulk-to-RHR suction) of about 10*F. This exceeds the value cited by Mr. Humphrey. On the other hand, for the GGNS with RHR suction
, elevation of 10'-6", a conservative temperature difference of about 6*F pre-vails; i.e., RHR suction temperature is greater than bulk temperature.
The applicant has indicated that RBS containment response analysis assumes RHR suction temperature to be 5'F less than the bulk pool temperature. This re-duces the nonconservatism to only 5'F. BNL also notes (Reference 4.3.5) that the RBS main vent vertical spacing is greater than standard, bringing the bottom vent in closer proximity to the basemat (2-1/4 ft vs. 4 ft for the standard plant) . We would expect this to ler.d to a flatter vertical tempera-ture profile; i.e., higher temperatures at lower elevations. Al so , we are satisfied that RHR operation will be very effective in reducing vertical stratification. This judgement is based on the results observed during in-plant SRV tests in which RHR operation wa', involved (References 4.3.6 and 4.3.7 ) . In fact, during the Kuosheng tests (Reference 4.3.7), it was found
- that RHR operation induces a favorable radial temperature stratification.
That is, higher temperature fluid is directed by the swirling motion toward the containment walls where the RHR suction strainers are typically located.
For these reasons, as well as the margins that have been demonstrated to exist
23-in heat exchanger performance (Reference 4.3.8), BNL concludes that the RBS analysis will provide conservative estimates of containment response despite the existence of vertical temperature stratification induced by LOCA blow-downs.
Conclusion BNL considers this issue satisfactorily resolved for the RBS.
References 4.3.1 " Mark III LOCA-Rel ated 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 Anendment 1, February 25, 1982.
4.3.4 Enclosure to GSU Letter No. RBG-19,972 Dated January 23, 1985 from J.E.
Booker (GSU) to H.R. Denton (NRC).
4.3.5 Gulf States Utilities Co., " River Bend Station Final Safety Assessment Report - Appendix 6A." November 1984.
4.3.6 Patterson, B. J., " Mark I Containment Program", Monticello T-Quenchers Thermal Mixing Test - Final Report," General Electric Co. Report NEDE-24542-P, April 1979.
4.3.7 NOTECH International " Final Test Report - Safety Relief Valve Discharge Test - Kuosheng Nuclear Power Station," Report ZTP-06-310, Rev. O, August 1982.
4.3.8 MP&L Letter No. AECM-82/353 dated August 19, 1982, from L. F. Dale (MP&L) to H. R. Denton (NRC).
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 pro-duce any mixing from the suppression pool layers below the vent row. An upper pool dump may contribute to additional suppression pool temperature stratifi-cation. The large volume of water from the upper pool further submerges RHR heat exchanger effluent discharge which will decrease mixing of the hotter, upper regions of the pool. Finally, operation of the containment spray elimi-nates the heat exchanger effluent discharge jet which contributes to mixing.
Evaluation The applicant has indicated that the RBS design does not incorporate an upper pool dump or containment sprays (Reference 4.5.1). In the absence of these design features, this concern reduces to that expressed via HSCs 4.3 and 4.4.
Conclusion BNL considers this issue satisfactorily resolved for the RBS based on the dis-cussion provided for HSC 4.3 (by BNL) and HSC 4.4 (by NRC).
References 4.5.1 Enclosure to GSU Letter No. RBG-19,972 Dated January 23, 1985 fran J.E. Booker (GSU) to H.R. Denton (NRC).
Humphrey Safety Concerns 4.7 and 4.10 4.7 All analyses completed 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 arrangenent for the suction and return lines is not properly designed, tl. 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 number of concerns were addressed systenatically. These included short circuiting, development of bulk pool motion, ability to eliminate ther-mal 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, include a sufficient range of parameters to encompass the RBS plant unique features, and are therefore applicable. Since the findings from these tests show that good mixing can be achieved, as well as the absence of short cir-cuiting, we conclude that the RBS 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 RBS.
References l
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, Rev. A, November 1982.
, 4.7.2 Gilbert Associates, Inc. , "Model Study of Perry Nuclear Power Plant Suppression Pool - Final Report," November 1977.
l l
I
Humphrey Safety Concern 19.1 The chugging loads were originally defined on the basis of 7.5 feet of submer-9ence over the drywell to suppression pool vents. Following an upper pool dump, the submergence will actually be 12 feet which may effect chugging loads.
Evaluation Since the applicant has indicated that the RBS design does not employ an upper pool dump (Reference 19.1.1), this concern does not apply to the RBS.
Conclusion BNL considers this issue satisfactorily resolved for the RBS.
References 19.1.1 Enclosure to GSU Letter No. RBG-19,972 Dated January 23, 1985 from J.E. Booker (GSU) to H.R. Denton (NRC).
Humphrey 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 the generic one which was origi-nally provided by MP&L for the GGNS via Reference 19.2.1. In this submittal, physical arguments and analytical procedures are used to estimate the pressure field that would be generated on the suppression pool boundaries if the worst case chug from the Mark III data base were to occur at vents located below the GGNS TIP platfarm. The results are compared with design 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 exceedance of design amounting to 60% occurs on the basemat.
The applicant argues that this exceedance is not important because this is a local load affecting only the basemat liner and that because of the hydrostat-ic head to which the liner is subjected, it will not experience a " negative pressure in the frequency range 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 fol-lowing: the loads developed for'GGNS are conservative for the RBS; as indi-cated in their respective responses to Humphrey Safety Concern 1.0, the GGNS has the largest encroachment and the RBS the smallest: the use of an acoustic model in the analysis represents a significant conservatism; dissipative mech-anisms not accounted for in such an analysis result 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 significant conservatism; in a more recent submittal by CEI to address the staff concern relative 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 cen-tral vent and average strength chugs 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 encroachment.
Conclusion BNL is satisfied that the issues related to this concern have been satisfac-torily 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).
____ -______ ______________ _ _____ ______ ________ __ _ _ _ _ _ _ _ _ _ _ _ _ .