ML19274G168

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Draft NRC Acceptance Criteria for Implementation of Mark I Containment long-term Program
ML19274G168
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Issue date: 08/02/1979
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Office of Nuclear Reactor Regulation
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NUDOCS 7908300316
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{{#Wiki_filter:, 01 INTRODUCTION The purpose of the Mark I Containment Long Tem Program is to perfom a complete reassessment of the suppression chamber (to:us) design to include suppression pool hydrodynamic loads which were neglected in the original design, and to restore the original intended design safety margins of the structure. This reassessment will be accomplished by a Plant-Unique Analysis (PUA) for each Eifa plant with a Mark I containment, using load specifications and structural acceptance criteria that are appropriate for the life of the plant. The following acceptance criteria have been developed from the staff's review of the Long Tem Progrsa Load Definition Report (LDR), the Plant Unique Analysis Applications Guide (PUAAG), and the supporting analytical and experimental programs conducted by the Mark I Owners Group. The:'s criteria specifically address the dynamic lo n g conditions. Unless otherwise specified, all other loading conditions and structural analysis techniques (e.g., dead loads and seismic loads) will be in accordance with. the plant's approved Final Safety Analysis Report (FSAR). CONTAINMENT FRESSURE AND TD4?ERATUEE RESPCNSE The pressure and teaperature transients for the d:ywell and wetwell shall be detemined by the use of the analytical models and assumptions set forth in Section 4.1 of the LDR. These techniques have, in the past, bean found to provide conservative estimates of the containment response to a LOCA, by ccmparison to the staff's CONTEMPI-lT ccmputer code. Plant-specific results, for each break size, shall be presented in the PUA, along with the input conditions, in sufficient detail t allow the staff to perfom confirmatory analyses to assu:e proper application of these models. 2032 97' 7008300 g / g ~

) 02 For the Design Basis Accident (DBA), the mass and energy release rates from the primary system are to be calculated with the Homogeneous Equilibrium Model (HEM), applied in a non-mechanistic reactor system which does not take credit for pressure reduction in the piping during the early portion of blowdown and conserva.tively assumes all liquid flow during most of the remainder of the blowdown. By comparison of the mass and energy release rate predictions of the GE m: del to those of a conservative RElAP-4 analysis, we cen:1ude that the GE model will provide a conservative prediction of the critical flow rates for a postulated double-ended recirculation line break for BWRs with a Mark I containment system. The staff has determined that the application of Hm to calculate the mass and enerEy release rates from the primary system will not necessarily provide conservatively hi h release rates for the Intermediate Break Accident (BA) 6 and the Small Break Accident (SBA). However, the purpose of the IBA and SEA is to provide a spectrum of event combinations, where the primarf loading conditions are steam condensation and SRV discharge loads. The primary loading ccndition affected by the HEM model, and the assumed break sizes for the I3A and SBA, is the containment pressure and temperature For the IBA and SEA, the containment response is of secondary response. importance to the lomHeg condition and the primary leading conditions are calculated independent of the containment respense. On this basis, we conclude that application of the Hm model for the IBA and SBA event combinations is acceptable. The timing and duration of specific loads are based primarily on th ,1;.at-specific centainment response analysis for the pool swell-related loads, while the condensation periods are ncn-mechanistically maximized. However, 2032 2 " a

the duration of SBA condensation loads are assumed to be limited by manual operation of the Autematic Depressurization System (AIS) at 10 minutes into the accident. Therefore, as part of the PUA, each licensee sF M identify the procedures (' including the primary system parameters monitorad) by which the operator will identify the SBA, to assure manual operstien of the AIS within the specified time period. VENT SYSTEM PRESSURIZATION AND THRUST LOADS The vent system pressurization and thrust loads during pool swell shall be defined in accordance with the procedures set forth in Section 4.2.1 of the LDR. These loads have been derived by assuming a distribution of flow losses which will maximize the vent system pressures, and an examination of the QSTF test data to maximize mass flow effects. The vent clearing time is significant, especially for those plants that propose operation with differential pressure control, because it establishes a transition in the vent load calculations. Therefore, the vent system pressures and the vent clearing time shall be detennined by the GE containment response model, which incorporates a virtual mass (equivalent to an extended downcomer length) in the vent clearing model. These procedures will provide conservative estimates of both the initial vent system pressure transient and the subsequent vent system flew effects and are, therefore, acceptable. 2032 m

04 NET TORUS VERTICAL PEESSURE IDADS The downward and upward net vertical pressure loads on the torus shall be derived from the series of plant-specific @TF tests, in accordance with Section 4.3 1 of the LDR. However, based on our review of the pool swell tests conducted by the Mark I Owners Group and confirmatory tests performed for the NRC by the Lawrence Livermore Laboratory, we will require that the following margins be applied to each lonMng phase: UP +0.215(UPmean) UP = mean DOWN,,,+ 2 x 10-5 (regg )2 DOWN = where "mean" refers to the statistical average of the QSTF plant-specific test runs. These margins shall be applied to the UTF "mean" load function prior to scaling the load function up to full-scale equivalent conditions. The margin on the upward ~tessure load includes 15% to bound the uncertainties arising from comparisons of all the two-dimensional and three-dimensional upward load test data. The remainder of the upward load maqin and the margin for the downward load reflect the randemness observed in the UTF test results. The margins specified above may te reduced or emitted where minimum censervatisss in the @ TF tested conditions for a specific plant can be demonstrated by the application of the @TF sensitivity test series (NZIE-23545-P). The sensitivity tests may not be used to adjust the mean torus vertical pressure loads. If the plant configuration is changed to the extent that the QSTF test series no len6er represents a conservative configuration of the plant, then a new series of QSTF tests sF41 be performed. 2032 F"

O5 For those plants that use drywell/wetwell differential pressure control as a load mitigation feature, an additional stmetural analysis shall be performed assuming a loss of the differential pressure control to demonstrate the capability of the containment to withstand'this extreme condition, as specified in Sections 5 3, 5.4, and 5.6 of the PUAAG. For this analysis, a singin plant-specific @TF test run may be used to define the loM*ng function however, the downward and upwari loading phases shall be increased by the margins specified above for the base analysis. TORUS SHELL PRESSURE - POOL SWEL_L The spatial distribution of the torus shell pressures during pool swell shall be defined from the plant-specific QTF test results and the azimuthal and longitudinal distribution factors defined in Section 4 3 2 of tne LDR. However, the @TF results shall be adjusted to incorporate the mard.ns specified for the net torus vertical pressure loading function; i.e., the avera6e pool prassure shall be increased by the magin specified for the downward load during the downward loading phase and the airspace pressure shall be increased by the margin specified for the upward load during the upward loading phase. Although the distribution factors have been based on averaged tes". results (44 tests for the azimuthal and 24 for the longitudinal), he conclude that the margins in the base load function adequately covar the uncertainty in the local pressure definition. DRAT am y

.s 06 CCMPRESSIBLE FLCW EFFECTS IN SCALED 100L SWLL TESTS The QSTF plant-unique and sensitivity test series are based on a " split-orifice" vent flow scaling relationship. Preliminary calculations performed by EPRI and GE indicate that compressibility effects, which could not be accurately scaled in the testing program, could result in a hi6 er loading h condition at full-scale conditions than that derived from " scaled-ur" test data. The original intent of these analyses was to provida L,ustification for the sea.ed flow distribution in the EPRI 1/12-scale, three-dimensional pool swell test program. The loading functions predominantly affected by this finding are the torus downward and upwari vertical pressure loads and the vent header pool swell impact timing. Based on our review of the preliminary analyses performed by EPRI and CE, which were presented in a meeting with the staff on July 2f+, 1979, we conclude that there is sufficient margin in the torus vertical pressure loads (previously specified) and in the header impact timing techniques to accomodate this uncertainty. Further, we do not believe that the Mark I implementation should be delayed during the time that it will take to resolve this concern. We will require, however, that the Mark I Owners Group complete the assessment of compressible flow effects and justify the conservatism in these load specifications prior to the issuance of our Safety Evaluation Report, which is currently scheduled for December 1979 In the event that this conservatism cannot be demonstrated, these two loading conditions will have to be reassessed. DRFI mm

07 VENT SYSTD4 IMPACT AND DRAG IDADS A. Vent Header Impact and Drag Loads The load definition procedures $let forth in Section 4 3 3 of the LDR are acceptable, subject to the fonowing clarifications:

1. The expcrimental data of local vent header pressure in each of the Mark I plants shall be obtained from the QSTF plant-unique tests.
2. The specification, for each Mark I plant, of the pressure inside the vent header relative to that in the torus airspace at the time of rater impr. t en the vent header shan be determined from the QSTF plant-unique tests.
3. The plant-unique header impact timing (i.e., longitudinal and circumferential time delays) shall be documented in each pUA, and are acceptable subject to confirmation from the assessment of compressible flow effects in scaled pool swen tests as previously discussed.

s 1,. Downcomer Impact and Drag Ioads The load definition procedures set forth in Section 4.3 3 of the LDR are acceptable, subject to the fonowing clarifications. A pressure of 8 psid is to be applied uniformly over the botton 50 of the angled portion of the downcemer, starting frca the tine at which the rising pool reaches the lower end of the angled secticn and ending at the time of maximum pool swen height. The pressure is to be applied perpendicular to the local downecmer surface. The structtral analysis for the downcener impact shall either be dynamic, accounting for the approximate virtual mass of water near the sutz:erged parts of the downecmer, or a dynamic load facter of two shall be applied. 2032 '77

A8 C. Main Vent Impact and Drag Loads For the main vent, the acceptance criteria specified for impact and d.m leads on other structures above the pool apply, except that the structure shall be subdivided into smaller sections and the impact and drag loads calculated separately on each subdivision. 2()3 2 ? ? ')

09 ""dfI. E P001 SVELL IMPACI' AND LRAG ON OTHER INTETAL STRUCTURES The impact and d.ra6 loads for intemal structuns, other than the vent header, downcomers, and vent header deflectors', as specified in Section 4 3 4 of the LER shall be modifie'd on the# asis of a cylindrical (e.g., b pipes) or an exposed flat surface (e.g., "I oesss). The load specification for these two 6eometries is as follows: A. Cylindrical Structures For cylindrical structures, the pressure transient which occurs upon water impact and subsequent drag is depicted in Figure 1. The parameters in Figure i shall be defined as follows:

1. The pulse duration ( T ) i specified as a function of the impact velocity:

T = 0.013 D for V $ 16 ft/sec t = 0.21 (D/V) for V> 16 ft/sec where D is the cross-sectional diameter of the teget in fee't.

2. The hydrodynamic mass per unit area for impact loading shall be obtained from the correletaan (cylindrical target) depicted by Figure 6-8 in NEDE-13426-P.
3. The impulse of impact per unit area shall be determined by:

bI } V I

  • p T

144 g ) where I is the impulse per unit area (psi-sec), g/A is the 2 hydrodynamic mass per unit area (1bm/ft ) and V is the impact velocity (ft/sec). 0b A

10 4, The pressure due to drag following impact shall be determined by: D / P9 i PD= I 144 Sef where P is the average drag pressure acting on the projected area D of the target (psi), C is the d ag coefficient as defined by Figure 2, D 3 and [ is the density of water (lbm/ft ).

5. The maximum pressure (P

) shall be calculated from the impulse per unit area, the pulse duration, and the drag pressure as follows: (I - 0.5 T P P D P = ,p t I, where P is the maximum pressure averaged over the projected area, &]d I,is a dimensionless empirical parameter defined by Figure 3.

6. The line connecting P to point A in Figure 1 is defined by:

1 - exp[a(1-t/t )] line A = (P -P + D D where a is a peakedness parameter defined by Figure 4, t is time (sec), P is the maximum pressure averaged over the projected area, and P is the average drag pressure acting on the projected D area. O]

R G. ~ B. Flat - Surfaced Structures For flat-surfaced structures, the pressure transient which occurs upon water impact and subsequent drag is depicted in Figure 5. The parameters in Figure 5 shall be defined as follows:

1. The pulse a-ation ( t ) is specified as a function of the impact velocity:

T = 0.0016 W forV{ 7 ft/sec t = 0. 011 W for V > 7 ft/sec y where W is the width of the flat structure (feet) and V is the impact velocity (ft/sec).

2. The pressure due to drag following impact shall be determined by:

IO D / V D" I 144 g where P is the average drag pressure acting on the frontal area D of the structure (psi), the drag coefficient C = 2 (flat strips D normal to flow, independent of Renolds number), and [ is the 3 density of water (1bm/ft ).

3. The hydrodynamic mass per unit area for impact loading shall be obtained from the correlation (flat targets) in Figure 6-8 in NEDE-13426-P.
4. The impulse of impact per unit area shall be determined by:

y T((144g \\ p" where I istheimpulseperunitarea(psi-sec),.g/Aisthe hydrodynamic mass per unit area (lbm/ft 1, and V is the impact velocity (ft/sec). 2032 m;

5. The maximum pressure (P

) shall be calculated from the impulse per unit area and the drag pressure as follows: 2I P ,p P = max t D C. Load Application For both cylindrical and flat structures, a margin of 35% shall be added to the loads derived above to obtain a conservative design assessment load specification. This load specification corresponds to impact on " rigid" structures. When performing the structural dynamic analysis, the " rigid body" impact loads shall be applied; however, the mass of the impacted structure shall be adjusted by adding the hydrodynamic mass of impact. The value of the hydrodynamic mass shall be obtained from the appropriate correlation in Figure 6-8 in NEDE-13426-P. 2032 '99

13 P a max m o 1.c. aoe C. E N o 00 m E. point A o> P D I p P s / 9 4 tira 7 Figure 1. Pulse Shape for Water Impact on cylindrical Targets . i + e i ,t .Ii ,l .l. l l l. i i, i. ^ e i i 1 .i i i o i i, ,o, 'l ,l[! N ',,e 2. .,vi e i ,,e , i i s ,,, i. i, i i.. X i,i i. i. , i i .. i., ,3 x ,i s . i. C ,x. D . i. ,, i, e 1 i .x, . 4. i .l .i e 1 i . i x i e i i I, + i ,6 .l ., i i i i 8 i*' i i I t I T i i. ,4 i ,i g. 9 e 1 .i. 1* I .i i i i. . i. .i 4 t . i i i i. .i. i i i 4 i i .i. i . I I a, e 6. .i i.. , i. t.i e i i i i .a i - e. i i i i o t I I 4 l l 4 2. 4. 6. 8, 5 2. 4. 10 d"G70 i n/ UJf Re(D) Figure 2. Drag Coefficient for Cylinders Following Impact

14

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-= = fur:=t_.4._t t-- = r .r 4 n.= i._._ = i -t______1__.__ _. _ _1 - = E 5-n - ~ rn= :2gg-jE! ._ rg_71,_..;;=r- ;run- _-t 7~ = .= 1 n;.n = n=:=n==; u=rn==r = -- ~ __ = - -- t := =.__. _ t .___.t. i.__.__.t.__. _ _.._. +. _.__.i __ = t-- --- r =t -- ---4 ---+ ---t 3*=__ : __i _. _ g { + _.. _ _ - +.

x-t = 7 -

i:" b = ---l- -* d! EI U-Ex=--! 5 -- x - _Agx=_.n=t-

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na4 -- pr:rg:__==- t + + 1 7 t intttlaeci - + :==t : _ _ __.

=_ __;==: r --

-t - :2 ----- __{. --- -n 6 + +1. gut.er_C.,=__D_ _i.mo,n_s_ i_ c. _ni e._ _._ _........ _... s_ s___P.u l_se_ _"Sh_, apo _rPa. rma. c.t_e r_. _._. _._ 4 =- _t e 2032 2M

15 EP ~~ ?. *** e a E E-Ee E I yPD P / 4 / / / I Ig time Figure 5. Pulse Shape for Water Impact on Flat Targets i 2032 - b

~ FROTH IMPINGEMENT AND FALLBACK LOADS Froth is generated by (1) impact of the rising pool surface on the vent header and (2) bubble breakthrough, as described in Section 4.3.5 of the LDR. The following load specfication was derived from the high-speed film records of various pool swell tests a" .. analysis of pool acceleration following vent header impact. Th.: impingement loads for Region I and Region II and the froth fallback loads, as described in Section 4.3.5, shall be defined as follows: p g P f 144 9c where: P = froth impingement pressure (psi) 7 pf = froth density (lb,/ft ) V = froth impingement velocity (ft/sec) g = gravitational constant c Region I: The froth velocity shall be based on a source velocity equal to 2.5 times the maximum pool surface velocity prior to vent header impact, which i; corrected for subsequent deceleration due to gravity starting at the 45 tangent on the bottom of the vent header, as shown in Figure 4.3.5-1 of the LDR. The froth density shall be assumed to be 20% water density for structures or sections of structures with a maximum cross-sectional dimension of less than or equal to one foot, and a proportionately lower density for structures greater than one foot; i.e., p = (0.2/x) p,, where x is the dimension in feet. The load shall be applied in the direction most critical to the structure within the 90 sector bounded by the horizontal opposite the vent header to the vertical upward. The load.shall be assumed to be a rectangular pulse with a duration of 80 milliseconds. 2032 2M

l L Region II: The firth velocity shall be based on a source velocity equal to the maximum pool surface velocity prior to vent header impact, which is corrected for subsequent deceleration hre the elevation of the bottom of the vent header. The froth density shall be assumed to be 100% water density for structures or sections of structures with a maximum cross-sectional dimension less than or equal to one foot, 2 5 water density for structures greater than one foot, and 10% water density for structures located within the projected region directly above the vent header. The load shall be applied in the direction most critical to the structure within the t5 sector of the upward vertical. 4 The load shall be assumed to be a rectan6ular pulse with a duration of 100 milliseconds. Fallback: The froth fallback velocity shall be based on the freefall velocity from the upper surface of the torus shell directly above the subject structure. The froth density shall be assumed to be 2 % water density, with the exception of the projected regien directly above the vent header which is 10% water density. The lead shall be applied in the direction most critical to the structure within the t 5 sector of the 4 . vertical downward. The load shall be assu=ed to directly follow the froth impingement load, with a duration of one second. 2032 2 "

VENT 1EAIER IEFIECTOP LOAIE The load definition procedures set forth in Section 4.3 9 of the I M are applicable only to the four deflector types'shown in Figure 4 3 9-2 of the LDR, and are generally acceptable, subject to the following constraints and/or modifications: A. h1 individual plant may choose to use deflector load data taken directly from the UTF plant-unique tests. This technique is subject to the followin6 requirements:

1. If the @TF deflector load measurement does not have a sufficiently fast response time to resolve the initial inpact pressure spike for the deflector types 1 - 3, inclusive, the lomiing transient shall be adjusted to include the empirical vertical force history of the spike shown in Figure 1.

This impulse need not be applied for the type 4 deflector.

2. The @TF plant-unique loads shall be adjusted to account for the effects of (a) impact time delays and (b) pool swell velocity and and acceleration differences which result from uneven spacing of the downcomer pairs. The correction technique sr4 1 be evaluated for the instant when the undisturbed pool surface would have reached the local elevation of the center (half-height elevation) of the deflector. The proposed three-dimensional load variation and timing is acceptable, subject to confirmation frem the assessment of compressible flow effects in scaled pool swell tests, as previously discussed.
3. In applying the load to the deflector, the inertia due to the added mass of water below the deflector shall be acccunted fer. The added mass per unit length of deflector may be estima+ d by:

2032 M"

19 I MaVw where: I = total impulse per unit length associated with the impact transient, V = impact velocity w = deflector width (as 'shown in Figures 2 - 5)

3. The deflector load definition which is based on empirical expressions for impact and drag forces together with semi-empirical, plant-specific definition of the pool swell velocity and acceleration transients, as described in Sections 4.3.9.1, 4 3 9 3, and 4.3 9.4 of the LDR, is acceptable, with the following modifications:
1. The impact transient and " steady d:-ag" contributiens to the load shall be ecmputed from the correlations shown on Figures 2 - 5.

for deflector types 1 - 4, respectively. For times past +Jw periods shown, the last value shall be extended for the duration of the transient.

2. The proposed three-dimensional load variation and timing is acceptable, subject to confirmation from the assessment of compressible flow effects in scaled pool swell tests, as previously discussed.
3. The gravitational ecmponent of the acceleration drag shall be included in F, as defined in NEDO 24612, A
4. In computing the deflector response to the lead, the added mass of the water shall be accounted for, as described in tre previous lead definition technique.

0 3 2 z'.0 9

20 F = vertical upward force on deflector per unit length d = diameter of cylinder in deflector types 1-3 V = impact velocity P = water density 'l t = time from beginning of impact s s I 7.o b y f$ 4 A D 6 6 i 6.'f g d'E g Yd* s y Figre 1. Impact Force Transient for Addition to the Empirical Data for Deflector Types 1 - 3 2032 .'o1

21 DRAFT r g m erd 3 l ~1*O ff f f ( r a e h t ] V i .4 ,g .l. .A

  • 1 M t
i..

= k .n c v. O es s 't Q Eb ~ ~ ~ ~ - ~ .... : g y i I D 0,5~ / o. y Vf t 4 7 l Firre 2. hpact & Steady Drag Force Correlation for Type i Deflector 2032 20!

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23 .g. V l b ?' ' g N' O'IW 5~- 4.95 F 2 $$uf f.I,/ ff f 7 ( / i T ' .1 (. /

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~ Vb _k ' O 6 o.s n t.o d 6 r. Q t t . h. g .f I D** '9' i d OQ d 2 Figre 3. Impact & Steady Dra6 Force Correlation for Type 2 Deflector 2032 2*5

u drift I r Q' w-w- j L i y 'W' 9F f' d,0.4H W" i r _ g pq N'g /,,,, " O 90 y l I i t. l-si I s-I i f i-I. t l / l i I -07 . / [, .,i ye , l, l a g v,... ,'1 l .t. Q g s. ~ d 'y ' ' :..$.. -l,..,\\. A -, - 7 ~ ' ~ t, " II 3 } D

  • D

~ w [u S f o o Ju _a Fig m 4 Impact & Steady Idag Force Ccrrelation for Type 3 Leflector 203? .?o: +e--

25 RWI W i 6 e 6 8 . ;'9 f. 4 ~ W.' fr/ irlW - $,y \\r S Y ov' 3.w-.g i ) a. c.7 t 6 ,t l g t.__ l.' fyf J Ii 't- \\ t 2 w-uL. e + : i f. . h. 4 e- .s f s. i I Fign=e 5 Impact & Steady Drag Force Correlatien for Type 4 Deflector 203? m w-

gg CONDENSATION OSCILLATION LOADS The following criteria have been developed in consideration of the fact that ~ the " condensation oscillation" loads (i.e., high vent flow rate with low air content) have been derived from a single FSTF test run (M8). The condensation oscillation regime is a hamonic fonction and, therefore, stastical variance or load magnitude uncertainty cannot be established from one test run. Although we conclude that the M8 tested conditions are coaservative and prototypical for the Mark I design, a reasonable measure of the uncertainty in the loading function is necessary to assure the margins of safety in the containment structure. However, based on our assessment of the phenomenological studies conducted by the industry and the NRC Office of Nuclear Regulatory Research, we believe, that the following load specifications are probably conservative and fom a sufficient basis to proceed with implementation of the Mark I Long Tem Program. We will require that the Mark I Owners Group confirm this assessment by performing at least two additional large break, liquid blowdown tests in FSTF. A commitment to perfom these tests and the associated schedule will be required before the issuance of the staff's generic Safety Evaluation Report, which is currently scheduled for December 1979. 2032 N

CONIENSATION OSCILLATION TORUS SIELL LOADS The load definition and assessment procedures set forth in Section 4.4.1 of the LDR for the condensation oscillation loads on the torus shell are acceptable, subject to the following confirmatio'n:

1. Provided that the " rigid wall" load derivation technique described in NEDE-24645-P is demonstrated to be conservative, in respense to Question 7 in our request for additional infom ation (D. Eisenhut, NRC, to L. Soben, GE, dated July 30,1979).
2. Provided that sufficient justification can be provided to exclude a condensation oscillation asymmetric leading condition, in response to Questien 2 in our request for additional informatien.
3. Provided that the uncertainty in the load magnitude is demonstrated to be less than the demonstrated conservatisms in the load specification, by the testing pro 6 ram described above.

For clarification, the load specification set forth in Section 4.4.1 of the LDR shall be used in conjunction with a coupled fluid-structure analytical model. The condensation oscillation londi,5 for the I3A is a continuous sinusoidal function with a peak amplitude and frequency range of that specified for the " pre-chug" load. We will require that the conservatism in the I3A condensation oscillation loads be demonstrated as part of the ..spense to Questien 7 in our request for additional infomation for the , specific flow regimes of interest. CCNIENSATION CSCTTLATION DCVNCCMER LATERAL LOADS A. Untied Downcener Lateral Leads The condensation oscillation downcomer lateral loads for untied downceners shall be defined as described in Section 4.4.3 of the LLR, based en the methodology in NE3-24537-P. However, in computing the dynamic load factors 7 1 1 'l ) Q s, ') u e

28 max " i i P,y = maximum static equivalent lateral load for plant-where unique downcomer, P = maximum static equivalent lateral load in FSTF y DLF = plant-unique downcomer dynamic load facter DLF = FSTF downcomer dynamic load factor 1 the plant-unique loading ccndition shall be derived as follows.

1. The damping factor assumed in the DLF shall be 2% (for both the plant-unique and FSTF) if the value of % for the plant is larger than that in F3TF, and 5% (for both plant-unique and FSTF) if the S-{for the plant is less than that in FSTF.
2. Assume that A is in the range between 4 and 8 Hz, and specify the lonMng condition as that which produces the maximum response in this range.

As used here, A is the driving frequency associated with the condensation oscillation lateral loads and to is the natural frequency of free lateral vibration of the downcomer (the ene closest to resonance with A, if the two major frequencies differ). E. Tied Downecmer Lateral Loads For tied downcomers, the structural response shall be assessed by a dynamic analysis of the tied downcomer pair and its tie bar. The applied load on each downecmer pair will be sinusoidal, with a frequency in the range 4 - 8 Hz (the worst loadin6 case shall be taken as the design condition). For the DBA event ecmbination, the amplitude of the load on each downcomer shall te 1360 lb, to occur synchreneusly on each of the two downcemers. f The direction of the load en each downcemer is to be within + 22 5 of the plane of the downecmer pair, as in the case of the untied downceners 2032 'oo

29 and the two downcomers shall be assumed to move away from and toward the torus axis synchronously, althou6h their angular direction within the i 22 5 sector is random. The design strains en the downcomer and the tie bar are to be taken as the worst load ccmbination, given the 4 - 8 Hz frequency range and the i 22 5 angle of application. For the IBA case, the specification is the same as that for the DBA case, except the amplitude of the load on each downecmer will be 854 lb and the direction of application shall be random within i 45 f of the plane of the downcomer pair. CONIENSATION OSCILLATION VENT SYSTEM FRESSURE LOADS The load definition procedures set forth in Section 4.4.4 of the LDR for the oscillatorf pressures in the vent system during the condensation oscillation period, are acceptable subject to confi=tation by the additional testing as described above. DR\\fi 2032 299

30 CHUGGING TORUS SHELL LOAE I The load definition and aes3ssment procedure set forth in Section 4.5.1 of the LDR for the chu66 ng condensation loads on the torus shell are acceptable, 1 provided the " rigid wall" load derivation technique is demonstrated to be conservative in response to Question 7 in our request for additional information (D.Eisenhut,NRC,toL.Soben,GE,datedJuly 30,1979). This load specification shall be used in conjunction with a coupled fluid-structure analytical model. CHUGGING DOWNCmER LATERAL IDAM A. Untied Downcomer Lateral Loads The chugging lateral loads en untied downcemers shall be defined as described in Section 4.5 3 of the LDR, which is based on the methodology in NEDE-24537-P, with the following exceptiens:

1. The load specification shall be based on the appropriately scaled maximum measured resultant static equivalent load (RSEL) in FSTF, rather than on the 95th percentile RSEL.
2. The multiple downcomer loading shall be based on a nonexceedance

-2 probability of 10, rather than 10 B. Tied Downecmer Lateral Loads For tied downcemers, the strains in the downcomer itself shall be evaluated exactly as in the case of the untied downcomers. The strain in the tie bar shall be evaluated by assuming that one of the two tied downcomers is subjected to a dynanic lead of triangular shape, with an amplitude of: 2 0 3 ^? sn" RSEL ymax.T t '~ ~ fd

31 where BSEL is the maximum measured RSEL for an untied downcomer during shuggin6, T is the lowest natural frequency of vibration of an untied g downcener for the specific plant, and the duration of the load, t '

  • d be assumed to be 3 milliseconds. The load direction shall le taken as that (in the horizontal plane) which results in the worst loading condition for the tie bar.

CHUGGING VENT SYSTEM PRESSURE LOADS The load definition procedure set forth in Section 4 5 4 of the LDR for the oscillatory pressures en the vent system during the chugging period are acceptable. DRAFI 2032 W'

32 A 3AFETY-RELIEF VALVE DISCHARGE DEVICE The acceptance criteria set forth' below for the quencher discharge loads and submerged structure drag load source strengths are applicable only to the "T" quencher configuration described in Section 1.1 of NEDE-24542-P. For plants using other types of quencher discharge devices, the SRV discharge load definition, submerged structure drag load source strength, and pool temperature limits will be evaluated on a plant-srecific basis. SRV AIR-CLEARING QUENCHER DISCHARGE SHELL PRESSURE '.0 ADS A. Methodology for Bubble Pressure Prediction The load definition procedures described in Section 5.2.1.3 of the LDR and the methodology in NEDE-21878-P for predicting the quencher bubble pressure are acceptable, with the following exceptions:

1. The load definition procedures described in Section 5.2 of the LDR are applicable only to SRV line submergences less than 13.5 feet.

In the event that the submergence exceeds 13.5 feet, additional justification will be required for the proposed load definition procedures.

2. The proposed methodology for predicting bubble p:mssures due to SRV subsequent actuations is not acceptable. The pressure amplitude predicted for the SRV first actuation shall be used in conjunction with the bubble frequency range for subsequent actuation, as specified below, for structure, equipment, and piping assessment in response to events containing SRV subsequent actuations.

2032 3na

D AN 33 Imf 2 B. Methodology for Torus she 1 Pressure Prediction Based on the predicted air bubble pressure-time histories, as discussed above, the torus shell pressures at various locations in the suppression pool shall be calculated by the load definition procedures described in Section 5.2.2.3 of the LDR in conjunction with the appropriate pressure attenuation model. For quenchers located on the torus center-line, the pressure attenuation model described in Section 2.4 of NEDE-21878-P in conjunction with the bounding factor presented in Section 3.2 of NEDE-21878-P shall be used. The load adjustment and attenuation factors proposed for the "off-center" T-quencher configuration presented in a meeting with the staff on May 30, 1979, are acceptable. We will require, however, that this load specification and its bases be documented in a supplement to the L2. C. Multiple - Discharg+ Loads The torus shell loads due to multiple SRV actuations shall be calculated as follows:

1. The peak values of bubble pressure due to a single valve actuation shall be combined by linear superposition (ABSS method) with the appropriate pressure attenuation model, as discussed above.

All bubbles shall be assumed to oscillate in-phase with the frequency ranges specified below for both first aid subsequent actuations.

2. In the event that the combined peak torus shell pressure exceeds 1.65 times the local predicted peak bubble pressure due to a single valve actuation, the resultant torus shell peak pressure for the design assessment may be taken at the lower value.

}}}} 4 fl 3

D. Frequency of Pressure Wave Form The pressure wave fonn predict'ed by the methodology described in Section 5.2 of the LDR within the following uncertainty ranges (streched or compressed time scale) that will produce the maximum structural, equipment, or piping system response shall be used for the design assessment:

1. First Actuation - the frequency range shall be 0.75 times the minimum predicted frequency to 1.25 times the maximum predicted frequency.
2. Subsequent Actuation - the frequency range shall be 0.60 times the minimum predicted frequency to 1.40 times the maximum predicted frequency.

2032.';0A

35 SRV DISCHARGE LINE CLEARING TRANSIENT The load definition and assessment procedure, described in Section 5.2.1 of the LDR, for the pressure and thrast loads on the SRV discharge line and quencher, which is based on the methodology presented in NEDE-21864-P and NEDE-23749-9, Addendum 1, is acceptable. SRV DISCHARGE LINE REFLOOD TRANSIENT The transient analysis technique to compute the plant-specific reflood heights in the SRV discharge line following valve' closure, as described in Section 5.2.3 of the LDR. and based on the methodology in NEDE-23898-P and NEDE-21864-P, is acceptable. SRV AIR AND WATER CLEARING THRUST LOADS The load definition and assessment procedure for the quencher and quencher support thrust loads, described in Section 5.2.6 of the LDR, is acceptable. SRV DISCHARGE LINE TEMPERATURE TRANSIENT The transient analysis technique to compute the maximum temperature loads on the discharge line and quencher device, as described in Section 5.2.7 of the LDR, is acceptable. DRlFI 2032 w

36 SRV DISCHAPSE EVETP CASES The kind and number of SRV discharge events shall be based on the plant-specific system configuration and a conservative. assessment of plant operational history. The following load cases shall be considered for the design assessment:

1. A first actuation, single valve discharge shall be censidered for all event ecmbinations involving SRV events. Sin 61e valve subsequent actuations shall be ccasidered for the SRV, SEA, and DA event combinations.
2. Asymmetric SRV discharge', both first and subsequent actuations, shall be considered for SRV, SEA, and EA event ccmbinations.

'Ihe de6:ee of asymmetric discharSe for each event combination shall be detemined from a plant-specific primary system analysis designed to maximise the asymmetric condition.

3. AIS valves discharging en first acutation shad be considered for the SEA and EA event ccchinations, followed by subsequent actuatiens determined from a plant-specific primary system analysis.
4. All valves discharging shall be considered for the SRV event combinations, followed by subsequent actuations determined from a plant-specific primary system analysis.

All of the event combinations above include the earthquake events (OBE and SSE) in combination with the SRV discharge events. 2032 jne

V SUFrxssSION POOL TDIPERATUFS LIMITS As part of the PUA, each licensee is requ' red to either demonstrate that previously sulmitted pool temperature analyses oi provide plant-specific pool temperature response analyses to assure that SRV d.ischage transients will not exceed the following pool temperature limits. A. Local Pool Temperature Linit The suppression pool local temperatt.re shall not exceed 200 F throughout all plant transients involving SRV operatier.s. B. Local and Bulk Pool Temperature The local to bulk pool temperature difference sb.11 be based on either the existing Menticello pool temperature data or plant-specific in-plant tests,andshallconsider(1)thequencherconfiguration,(2)theSRV discharge locations in the pool, and (3) the Rh3 suction and dischar6e geometry. The " local" temperature is defined as the temperature in the vicinity of the quencher device during discharge. For practical purposes, temperature measurements from locations on the reactor side of the torus, downstream of the quencher end-cap holes, and at the same elevation as the discharge device, may be considered local temperatures. The " bulk" temperature, on the other hand, is the temperature calculated assuming a uniform distribution of the mass and energy discharged from the SRV. 2032 307 1 w~ e

33 C. Suppressior. Fool Temperature Monitor System The suppression pool temperature monitoring system is required to ensure that the suppression pool is within'the allowable limits set forth in the plant Technica) Specification. The system shall meet the following general design requirements:

1. Each licensee shall demonstrate that there is a sufficient number and distribution of pool temperature sensors to provide a reasonable measure of the bulk pool temperature.
2. Sensors shall be installed sufficiently below the minimum water level, as spe".ified in the plant Technical Specifications, to assure that the sensor properly monitors pool temperature.

l

3. Pool temperature shall be monitored on recorders in the control room. Two sensors from each sensor group shall be recorded.

The difference between the measurement reading and the actual local pool temperature shall be within 1 2 F.

4. Instrument set points for alam shall be established, such that the plant will operate within the suppression pool temperature limits discussed above.

5 All sensors shall be designed to seismic Category I, Quality Group B, and energized from onsite emergency power supplies. 2032 (n'> s e f

39 LOCA WATER J1" JAM The load definition and assessment.procedu:e descirbed in Section 4.3 7 of the LDR, which is based on the 'Hoody Jet Model" (NEE-21472-P), is acceptablesubjecttothefollowingconstraintsand/ormodifications: A. The plant-specific jet di charge velocity, V (t), and acceleration, g a (t) = dVgdt (t), from the QSTF plant-specific teet series shall D be used as the driving sources for the jet model. B. Forces due to the pool acceleration and velocity induced by the advancing jet front shall be computed for structures that are within four downcomer diameters below the downcener exit elevation, even if the structure is not intercepted by the jet. The flow field shall be computed by modelling the moving jet front as a hettispher' cal cap centered one downcomer diameter (D) behind the " Moody" jet front positions, containirig the same amount of water as the " Moody" jet, and moving with the velocity of the " Moody" jet front. The fo2xulas for the hemisphere radius (R,) and the trajectory of the hemisphere center (x)a:e: R(t)=f(f+3(x(t)/D)Y)1/3 fe QD D 3 f U (t) )

  1. # *f(t) { D D

f R (t) = 7 ( s 2D x (t) = x (t) -D for x (t) > D c f f x (t) = 0 for x (t) $ D f c where x (t) is the position of the " Moody" jet front as a function f of time, as computed in NE2-21472-P. The equivalent unifom velocity and acceleration at the location of the st:ucture (x,y) shall be obtained from the timt dependent potential M (x,y,t) induced by the jet front: 2032 w)

t R dR h (x,y,t) = r

  • -g

(* ~ *c) dt 5 4r/ dt (r / fx-x) +y and y is the transverse distance of the where r = structure from the jet axis, and (x-x ) is the distance frem the c structure to the effective jet front center along the jet axis. The local uniform flow velocity is: U(x,y,t)=V@3 as in NEDO 21471, while the acceleration is a(x,y,t) = 3% This calculation need only be performed for r> R, and x) x. If either of these conditions are not satisfied, the methodology in the LDR will bound the load and is, therefore, acceptable. LOCA EUBBIE DRAG 10AIS The load definition and assessment procedures descirbed in Section 4.3.8 of the LDR, which are based on the methodology in NEDO 21471 and experimental confirmation in NEIE-23817-P, are acceptable subject to thefollowingconstraintsand/ormodifications: A. Flow Field

1. QSIT plant-specific test results (NEIE-21944-?) will be used.
2. Model E in NEIE-21983-? will be used for the method of ima6es simule. tion of the torus cross-section.
3. After contact between bubbles of adjacent downcomers, the pool swell flow field above the downconer exit elevation will be derived from the QSTF plant-specific tests.

20?>2 371

41 B. Drag Icad Assessment

1. Drag forces can be computed for circular cylinders as given 5

= 1.2 must ELO 21471, but a conservative drag coefficient of CD be assumed, independent of the Reynolds number.

2. Drag forces on structures with sharp corners (e.g., rectangles and "I" beams) must be ccuputed by considerir4 forces on an equivalent cylinder of diameter D

= (2 L ), where L is the maximum transverse dimension.

3. Long slender structures must be considered in segments of lergth (L), which do not exceed the diameter (D or D

).

4. For structural segments the centers of which are separated from each other by less than Oese diameters of the larger structure, interference effects shall either be considered in detail or a bounding load shall be established by multiplying both the acceleration and standard drag loads by a factor of four.

QUEICER WATER JET IAAIS The load definition procedure described in Section 5 2.4 of the LDR, which is based on the methodology in NEDC-25090-P, is acceptable, subject to the appropriate documentatien of the confirmatory tests discussed in NEIE-25090-P. Q1EiCER EUB3LE DRAC LOAIS The load definition and assessment procedures described in Section 5 2 5 of the LOR, in NECE-21878, and in NED0-21471-2, are acceptable subject to the following constraints and/or modifications: 2032 51

DRE A. Flow Field

1. The determination of the charging, formation, and rise of the oscillating bubbles is subject to the same conservative factors that are used for the quencher torus shell pressure loads, as described in NEDE-21878-P.
2. Drag loa /.s on the quencher arms and the SRV discharge line shall be ccmputed on the basis of asymmetric bubble dynadics.. Either a full 180 phase shift shall be considered for full strength bubbles on opposite sides of these structures, or a more detailed assessment of the asymmetry of the bubble source strengths and phasing must be obtained from the experimental information in NEDE-21878-P.
3. Model E in NEDE-21878-P shall be used for the method of images representation of the torus cross-section.

B. Drag Load Assessment

1. Drag forces for circular cylinders will be computed on the basis accelerationdragaloneundertheconditionthatU,T/Di2.74, where U,is the maximum velocity, T is the period of bubble oscillation, and D is the cylinder diameter.

For U,T/D > 2. 74, the standard drag shall be included with the drag coefficient D = 3.6 in order to bcund the relevant experimental data. C

2. The constraints specified for the LOCA bubble drag load assessment also apply to the quencher bubble drag loads, with the exception of the drag coefficient.

) t].7 ) ?1) I >. e

43 LOCA CONDENSATION OSCILLATION DRAG LOADS The load definition and assessment procedures described in Section 4.4.2 of the LDR and the methodology described in NEDO 25070 are acceptable subject to the following constraints and/or modifications: A. Flow Field

1. A maximum source strength shall be established by using the deduced " rigid wall" pressure (p,) at the bottom center of the torus (equation B-4 in NEDO 25070, with f(r) evaluated by equation B-7).

An average source strength s?lall be established by considering equal source strengths at all eight downcomers in equation B-4 in NEDO 25070. For each structure, the loads shall be computed for both the flow field produced by the average source applied at all downcomers and the flow field resulting from the maximum source applied to the two nearest downcomers on one side of the structure alone.

2. The fluid-structure interaction effects shall be included for any structural segment for which the local fluid acceleration is less than twice the torus boundary acceleration.

This may be accomplished by adding the boundary acceleration to the local fluid acceleration. B. Drag Load Assessment

1. Tia constraints and modifications specified for the quencher bubble drag loads apply.
2. These loads may be applied quasi-statically to structures, only if the highest significant Fourier components occur at frequencies less than half the lowest structural frequency.

2032 313

DRAFT LOCA CHUGGING DRAG LOADS The load definition and assessment procedures described in Section 4.5.2 of the LDR and the methodology in NEDO 25070 are acceptable subject to the following constraints and/or modifications: A. Flow Field

1. The maximum source strength history shall be obtained by using the maximum measured pressure (not necessarily at the bottom center) in a Type 1 chug in equation B-4 of NEDO 25070, with f(r) based on the single nearest downcomer. For each structure, the phasing between the two nearest downcomers that maximizes the local acceleration shall be established. The local acceleration shall then be computed on the basis of the two nearest downcomers chugging at maximum source strengths at the above established phase relation.
2. The fluid-structure interaction effects shall be included for any structural segment for which the local fluid acceleration is less than twice the torus boundary acceleration.

This may be accomplished by adding the boundary acceleration to the local fluid acceleration. B. Drag Load Assessment

1. The constraints and modifications specified for the quencher bubble drag loads apply.
2. Unless the lowest structural natural frequency times the duration of the " spike" in the source strength is greater than 3, the loads shall be applied dynamically. Either sufficient Fourier components will be included to bound tne " spikes" or the load shall be applied in the time domain using the sucree time history.

2032 3'4

SECCNDARY IDADS The following loading conditions may be neglected for the PUA:

1. seismic slosh pressure loads
2. post-swell wave loads
3. asymmetric pool swell pressure loads
4. sonic and compression wave loads
5. downcomer air clearing loads SCNIC AND COMPRESSION WAVES Immediately following the postulated instantaneous rupture of a large primary system pipe, a sonic wave front is created at the break location and will propagate throu6h the drywell and into the vent system. The subsequent compression of air in the drywell and vent system will cause a compression wave to be generated in the water leg inside the downcomer.

This compression wave propagates throu6h the poo. and results in a differential pressure loading on sutnerged structures and the toras wall. These loading conditions were observed in the FSTF data. The desi6n of FSTF was such that a conservative estimate of the leading condition could be established, because the simulated drywell volume did not allow s1 nificant attenuation of the wave front. The maximum observed loads 6 were approximately 20 psid in the drywell and 10 psid in the torus, with a maximum duration of less than 5 milliseconds. This loaMng condition preceeds all other dynamic loads, and is insignificant in comparison to the other dynamic loads. In addition, a more realistic attonuation of the wave front, based on the actual configuration of the drfwell would result in even lower loads than those obserted in FSTF. 2032 5!c

46 On this basis, we conclude that neglecting the sonic and compression wave loads is acceptable. SEISMIC SLOSH Seismic motion induces suppression pool waves which can (1) impart an oscillatory pressure loading on the torus shell, and (2) potentially lead to uncovtry of the downcemers, which would result in steam bypass of the suppression pool and potential overpressurization of the torus, should the seismic event occur in conjunction with a LOCA. To assess these effects, the Mark I Owners Group undertook the development of an analytical model which would provide plant-specific seismic wave amplitudes and torus srall pressures. Thismodelwasbasedon1/30-scale"shaketest" data for a Mark I torus geometry. Based on the results of the plant-specific analyses, the Mark I Owners Group concluded that (1) the seismic wave pressure loads on any Mavk I torus are insi nificant in comparison to the other suppression pool 6 dynamic loads, and (2) the seismic wave amplitudes will not lead to downcocer uncovery for any Mark I plant. This conclusion was based on the maximum calculated pressure loads and the minimum wave trough depth relative to the downconer exit. We have reviewed comparisons of the analytical predictions with scaled-up test data, the small-scale test program, and the seismic spectrum envelope used in the' plant-specific analyses. Based on this review, we cenclude that the seismic slosh analytical predictions will provide reasonably conservative estimates of both the wall pressure loading and the wave amplitude, for the ran6e of Mark I plant conditions. 2032

47 Since the maximum local wall pressures were found to be less than 0.8 psi at a 95 upper confidence limit, the Mark I Owners Group has proposed that the seismic slosh loads may be neglected in the structural analysis. We a6:ee that the seismic sloch loads are insi6nificant by comparison to the other suppression pool dynamic loads. On this basis, we conclude that ne6 ecting seismic slosh loads in the FUA is acceptable. 1 The results of the slosh wave amplitude predictions indicate that, within the local area of maximun amplitude and with maximum suppression pool drawdown (resulting from ECCS system flows), the slosh waves will not cause downcemer uncovery. We have reviewed the assumptions used in these analyses and conclude that they are sufficiently conservatis t-provide assurance that seismic slosh will not result in downcener uncovery. p0ST-p00L SWELL WAVES Following the initial pool swell transient, pool wave action will result from centinued flow through the vent system. This wave action will result in pressure loads on the torus walls. For the period immediately following the downward and upward vertical pressure transient, the Mark I Owners Group has concluded that this wave action is inherently included in the @TF pressures and are negligible. Although the scaling relationships by which the @TF was designed are not applicable following bubble breakthrough, we ag ee that the @TF results provide a reasonable estimate of the wave loads during this period. We, therefore, conclude that the post-pool swell wave loads may be neglected in the pUA. 2032 3i7

48 Durin6 the subsequent condensation period, the pool wave action is inherently included in the condensation oscillation and chu661ng load specifications. These loads were derived from wall pressure measurements from full-scale steam condensation tests, and, therefore, a separate load specification for condensation wave loads is unnecessary. ASLvMETRIC VENT SYSTEM FLOW The effects of asymmetric flow rates in the vent system have been considered with respect to unequal vent flows (e.g., vent blockage) and unequal vent flow composition, to evaluate the potential for asymmetric pool swell loading conditions. The three-dimensional pool swell tests conducted by EPRI for the Mark I Owners Group, and the confinatory three-dimensional pool swell tests conducted for the NRC by the Lawrence Livermore Laboratory, included specific tests to assess the effects of partial and full blockage of one main vent. The results of these tests indicate that the distribution of the pool swell pressure loads are relatively insensitive to the main vent block 36e, because the vent header tends to equalize in pressure and, therefore, equalize flow throu6h the downcomers. Due to the configuration of the Mark I vent system, the main vent entrance is the principal location where flow blockage could occur, if at all, and, therefore, flow blocka6e assumptions have not been considered for other locations in the vent system. To assess the effects of potential aspmetric vent flow composition, we have considered the extreme case of localized steam flow throu6h the vent system, without benefit of any steam-air mixing. For the DBA, the earliest 2032 319

d) time th. a steam front could reach the towncomer exit is shortly before the peak upward load. However, an additic.ual time delay would occur before the steam could reach the bubble pool interface and affect the local p sure loads, due to the circulation of t'he steam in the existing air bubble. In addition, as the steam ccndenses reducing the bubble growth, the air compression, which is the major contributor to the upwari lonMng phase, would tend to equalize. This would result in a reduced potential for an any= metric 7.oading condition and lessen the severity of the pool swell loacs, We have also considered this extreme case for smaller breaks to assess the potential for localized pool heatin6, leading to overpressurization of the wetwell. Based on the resultL of FSTF, we conclude that the increased vent flow rates accompanying higher energy deposition in the pool will provide sufficient mixing to prevent overpressurization of the wetwell. Based on the results of this assessment, and the extreme nature of the assessment, we conclude that neglecting asymmetric pool swell conditions for the PUA load definition is acceptable. DCW'iCCfER SUEMERGENCE AND TEEEMAL TPATIFICATION One methed of suppression pool hydrodynamic load mitigation thec the Mark I Owners Group has adopted for the LTP is reducing the initial sulnergence of the downecmer in the suppression pool to a minimum of three feet. The pool volume (i.e., themal capacity) of the original design would be maintained. This approach, however, raises conce=s rega::-A' the increased 7 potential for downconer uncovery and steam condensation capability, both of which could lead to wetwell overpressuri::ation. 2032 <!o

The potential for downcener uncovery was addressed by the preceeding seismic slosh assessr ent. This assessment was perfomed at the most extreme conditions that could potentially lead to downcomer uncovery and was predicated on a minimum three foot downcomer subergence. Condensation capability of the suppression pool, for the spect um of IDCAs, is a function of the local pool temperature in the vicinity of the downcomer exit. FSTF test results and foreign test data have shown that themal strAification occurs, and becomes more severe as the downcomer suburgence is reduced. The most severe themal stratification has been observed in low flow tests with a quiescent pool. In actual plant conditions, the Reisidual Heat Removal (RHR) system and SRV discharge will provide sufficient long-tem pool mixing to minimize themal stratification. As previously discussed, for asymmetric vent system flows, we have detemined that, for the short-term, the increased vent system flow rates with hi6 er energy deposition will prevent overpressurizatien. h This assessment included cont,ideration for vertical themal stratification as well. In addition, the analytical predictions of the wetwell pressure and bulk temperature response have been found to be conservative by comparisen to FSTF test data for plant-simulated initial ccnditiens. The local temperature variation in the pool which has been observed in the test data are not significant tc the structure, and, therefore, need not. be considered in the st=uctural analysis. Based on this assessment, we conclude that a minimum initial downcomer subergence of three feet is acceptable, and there is sufficient conservatism in the centainment respense analysis techniques to accomodate the effects of themal stratification. 03cs

51 DOWNCOMER AIR CIEARING LATERAL LOAIS During the initial pahse of a LOCA, the. rapid clearing of air from the vent system causes the downcomer to be subjected'to a lateral load as bubbles are being formed in the pool, in addition to the thrust loads on the vent system previously discussed. Conservative estimates ef the air clearing lateral loads were obtained from the FSTF data. The Mark I Owners Group has proposed to neglect the & clearing lateral load because it is bounded by the repetitive steam cond.usatien lateral loads on the downcomer. We concur with this assessment ad, therefore, conclude that neglecting the air clearing lateral load on it s downcemers is a.coptable. 2032 Di 8

I DIFFERENTIAL PRESSURE CONTROL REQUIREMENTS Those licensees that use differential pressure control (AP) as a pool swell load mitigation feature for the LTP, shall demonstrate conformance with the following design criteria as part of the PUA:

1. There shall be no unacceptable change in the radiological consequences of an accident as a result of the inclusion of the AP system.
2. Steam bypass of the suppression pool via the AP system shall be eliminated by appropriate system design, or such bypass shall be demonstrated to be acceptable by calculation.
3. Design and installation of the AP system shall be coc:mensurate with other operational systems in the plant,
4. When the AP system involves the addition of containment isolation valves, the additional valves shall be included in the plant's Technical Specifications and the valve design and arrangement shall conform to the requirements of General Design Criterion 56 in Appendix A to 10 CFR 50 and the regulatory positions in Standard Review Plan Section 6.2.4.

Subsequent to the PUA, a license amendment shall be submitted to incorporate the following Technical Specification requirements for the AP system:

1. Differential pressure between the drywell and suppression chamber shall be maintained equal to or greater than "X" (where X is the plant-specific differential pressure and values less than one psid will not be credited for load mitigation), except as specified in 2 and 3 belev.
2. The differential pressure shall be established within 24 hours after placing the plant in the RUN mode, during plant startup.

The differential 20.32 m

DRJIT pressure may be reduced below "X" psid 24 hours prior to a scheduled plant shutdown.

3. The differential pressure may be reduced to less than "X" paid for a maximum of four hours during required operability testing of (specify here those safety-related systems for which operability tests either release significant amounts of energy to the suppression pool or cannot be performed with the AP established).
4. In the event that the specification in 1 above cannot be met, and the differential pressure cannot be restored within six hours, an orderly shutdown shall be initiated and the reactor shall be in a cold shutdown condition within the subsequent 24 hours.
5. A minimum of two narrow range instrument channels shall be provided to monitor the differential pressure. Error in the A P measure-ment shall be no greater than 1 0.1 psid. The instrument channels shall be calibrated once every six months.

In the event that the measurement is reduced to one indication, operation is permissible for the following seven days. If all indication of the differential pressure is lost, and cannot be restored in six hours, an orderly shutdown shall be initiated and the reactor shall be in a cold shutdown condition within the subsequent 24 hours. 2032 m

5$ STRUCTURAL ANALYSES AND ACCEPTANCE C E.:n The staff finds the general analysis techniques and proposed structural acceptance criteria set forth in the " Mark I Containment Program Structural Acceptance Criteria Plant Unique Analysis Applications Guide," (PUAAG), NEDO 24583, Revision 1, dated July 1979, acceptable. The proposed criteria vill provide a sufficient basis for demonstrating the margins of safety required for steel structures and piping in the AGME Boiler and Pressure Vessel Code and for concrete structures in the American Concrete Institute Code. Revision 1 to the PUAAG was presented to the staff in a meeting on June 29, 1979. We will require that this revision be formally submitted to complete the documentation required for this program. ALLOWABLE STRESS LIMITS The structural acceptance criteria set forth in the PUAAG which will be used to evaluate the acceptability of existing Mark I containment systems or to provide the basis for any plant modifications to withstand suppression pool hydrodynamic loading conditions are generally contained in Section III of the ASME Boiler and Pressure Vessel Code through the Su=mer 1977 Addenda. The application of these stress limits to the Mark I design will provide adequate margins of safety to insure the containment structural integrity for all anticipated loading combinations and to insure that the containment and attached piping systems will perform their intended functions during those loading conditions expected to occur as a result of a LOCA or SRV discharge and are, therefore, acceptable. 2032 Pd

DRWT Additionally, the ratio of the dynamic collapse load to the static collapse load was established for the torus shell during LOCA pool swell pressure loads and for the vent header during pool swell. impact loads. These values, in conjunction with Code Case N-197, were used to establish the allowable stress values for the torus shell and the vent header local stresses. The staff and their consultants have reviewed the analyses used to establish these factors and find them acceptable. PIPING FUNCTIONALITY Recent studies by Battelle have demonstrated that theoretical collapse and piping wall buckling limits can potentially be exceeded for certain components in ASME Class 2 and 3 piping systems if Level C and 0 stress limits are used. Since the piping of concern in the Mark I systems is carbon steel and does not fall in the classification of diameter to thickness ratios that are greater than 50, the area of concern with respect to exceeding the theoretical collapse moments is piping elbows. For A-106 Grade B piping the Level 0 stress limit, 2.4 S, is 36 ksi or approximately yield for the h material. Since the study has shown that the theoretical collapse load could be exceeded in elbows when the stress calculated by Code Class 2 and 3 rules reaches approximately 3/4 of the material yield stress, the theoretical collapse load could be exceeded by approximately 1/3.

However, this theoretical limit and the supporting test data is based on a deflection limit of a few times the yield point, which would not significantly alter the piping flow area.

In addition, when a segment of a continuous piping system begins to yield a load redistribution will occur to other areas that would limit the deflection. 2032

Therefore, the staff concludes t !, '< r he dynamic loadings under consideration in the Mark I reassessment, the linear elastic analysis of piping systems using ASME Code Class 2 and 3 Limits provides reasonable assurance that sufficient margins exist to prec'lude excessive piping deflections which could impair flow in safety related piping systems and, therefore, consider these limits accpetable. DAMPING The damping values used in the analysis of dynamic loading events will be those specifisd in Regulatory Guide 1.61. Since these values are specified for seismic analysis of structures and components for OBE and SSE conditions, the values used will be consistent with the stresses expected under similar loading conditions. The staff considers the use of the Regulatory Guide 1.61 damping values acceptable for the Mark I dynamic analyses. OPERABILITY OF ACTIVE COMPONENTS Active components, as defined in Section 2.2.9 of the PUAAG, shall be considered to be operable if Service Limits A or B are met, unless the original design criteria establishes more conservative limits. If the original component design criteria establish more conservative limits, confomance with these more conservative limits shall be demonstrated even if Service Limits A e B are met. If the original component design criteria are silent with respect to operability limits, satisfaction of Level A or B Service Limits shall be considered as sufficient to demonstrate operability. 032 <9<

s1 Active components which do not satis y ervice Limits A or B, and therefore either Service Limits C or D are satisfied, require demonstration of cperability. If the original component design criteria for operability exist, confomance with those criteria shall be' demonstrated. If the f orignial component design criteria are silent with respect to operability limits, operability limits shall be established and confomance with those i criteria shall be demonstrated. The operability requirements are necessary to assure the active safety-related components will be able to perform their intended functions. It is the staff's position that loads calculated by elastic analysis which produce stresses in excess df the material yield stress can produce excessive deformation in a component w5ich can cause interference 6f mechanical motion. We recognize that the designation of Service Limits A and B do not, by themselves, guarantee the operability of active components. However, the scope of the Mark I Conatinment Long Term Program is directed toward the effects of the incremental load increase due to the definition of suppression pool hydrodynamic loads, and the restoration of the original intended design safety margins. The criteria for operability specify that the original component design criteria must be met where they are more conservative than the Service Limits A and B. We believe that these criteria are reasonable and practical and are sufficient to accomplish the objectives of the program. 2032 <r'

COMBINATION OF STRUCTURAL RESPONSES l' The structural responses resulting from two dynamic phenomena will be combined by the absolute sum method. Time phas.ing of the two responses will be such that the combined state of the stress results in the maximum stress intensity. However, as an alternate, the Cumulative Distribution Function (CDF) method may be used if the absolute sum does not satisfy the structural acceptance criteria. The CDF abcissa value corresponding to an ordinate value of 84% (i.e., the combined stress intensity value corresponding to an 84% probability of nonexceedance) will be used to compute a. eduction factor which will be applied to the stress intensity computed by the absolute sum method. An 84% probability of nonexceedance corresponds to a mean plus one standard deviation for two dynamic responses. The CDF method is more conservative than Criterion 2 of the Newmark-Kennedy Criteria proposed for use in the Mark II Containment Program. The rationale for the use of this methodology is similar to that contained in NUREG-0484 and is, therefore, acceptable. 2032 soo e}}

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