ML19210E091

Informs Commission of Recent Actions Taken on Unresolved Safety Issue A-7, Mark I Containment Long-Term Program. SER Describing Basis for Safety Criteria for Mark I BWR Containments Will Be Issued Dec 1979.Criteria Attached

ML19210E091
Person / Time
Issue date:	11/06/1979
From:	Gossick L NRC OFFICE OF THE EXECUTIVE DIRECTOR FOR OPERATIONS (EDO)
To:
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ML19210D983	List: ML19210D982 ML19210E087 ML19210E091
References
SECY-79-606, NUDOCS 7911290293
Download: ML19210E091 (43)

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{{#Wiki_filter:. .y UNITED STATES NUCLEAR REGULATORY COMMISSION November 6, 1979 WASHINGTON. D. C. 20555 SECY-79-606 INFORMATION REPORT e For: The Commissioners Lee V. Gossick, Executive Director for Operations g zc4 Thru: From: Harold R. Denton, Director Office of Nuclear Reactor Regulation

Subject:

ISSUANCE OF STAFF CRITERIA FOR THE MARK I CONTAINMENT LONG TERM PROGRAM

Purpose:

To inform the Commission of the recent actions taken on the Unresolycd Safety Issue (USI) A-7, " Mark I Containment Long Term Program." Discussion: In SECY 79-409A dated September 28, 1979, I stated my inten-tion to keep the Commission advised of progress on Unresolved ~ Safety Issues and our plans for implementing the results of the staff's evaluation. This report discusses the status and implementation plans for A-7. The staff has completed its review of the generic suppression pool hydrodynamic load definition techniques and structural acceptance criteria proposed by the Mark I Owners Group. These proposed procedures are intended to be used in a plant-unique analysis of each of the 17 BWR/ Mark I suppression chamber designs to demonstrate that the planned structural modifications and/or changes in plant operating procedures will restore the intended margins of safety for the containment. As a result of this review, the staff has determined that cer-tain changes or clarifications to the criteria proposed by the 4rk I Owners will be necessary. These requirements are set forth in Enclosure 1, "NRC Acceptance Criteria for the Mark I Containment Long Term Program," which has been transmitted to each affected licensee.

Contact:

C. Grimes x28204 7911290 1431 216

The Commissioners. As stated _in our requirements, there are two aspects of the suppression pool hydrodynamic loads for which we will require additional confirmatory tests and analyses (i.e., variance in " condensation oscillation" load magnitudes and pool swell compressibility effects). We believe that the implementation of this program, including plant modifications, should not be delayed for this additional confiimatory work. Throughout the course of the Mark I Long Term Program, the staff has urged the Mark I licensees to make any necessary plant modifications at the earliest possible time to expedite the resolution of this issue. In March 1979, the staff requested that each Mark I licensee submit its planned modification schedule, based on the extent of plant modifications which were known to them at that time. With the transmittal of the staff's acceptance criteria, each licensee of the 22 operating Mark I units has been requested to provide a schedule for submittal of the plant-unique analysis which is commensurate with the December 1980 targeted completion date for the program. Submittal dates for the 3 Mark I units under construction will be established based on the projected licensing review schedules. We are continuing to discuss these schedules with the indivi-dual licensees to ensure that every effort is being made to restore the margins of safety in the Mark I facilities as quickly as possible. The staff's Safety Evaluation Report describing the bases for our criteria is scheduled to be issued in December 1979. On September 14, 1979, we discussed the rationale for our criteria with the ACRS Fluid Dynamics Subcommittee. Copies of our acceptance criteria are being transmitted to the ACRS this week. We plan to meet with the subcommittee again on November 16, 1979, to discuss our con-firmatory review efforts and our implementation progress. Due to tiie public interest expressed on this issue, we have prepared the information necessary for a Federal Reaister Notice regarding this action and have trans.nitted this material to the Office of the Secretary. A copy of the Notice is contained in Enclosure 2. 1431 217

The Commissioners - I will continue to keep you informed of our progress on the implementation of this program. 19 g Harold R. Denton, Director Office of Nuclear Reactor Regulation

Enclosures:

1. NRC Acceptance Criteria for the Mark I Containment Long Term Program 2. Federal Register Notice 1431 218

TiRC ACCEPTAfiCE CRITERIA FOR THE MARK 1 CONTAlfiMEriT L0f4G TExit PROGRA!! 1431 219

TABLE OF CONTENTS 1. INTRODUCTIO" 2. SUPPRESSION POOL HYDRODYNAMIC LOADS 2.1 Containment Pressure and Temperature 2.2 Vent System Pressurization and Thrust Loads 2.3 Net Torus Vertical Pressure Loads 2.4 Torus Pool Swell Shell Pressures 2.5 Compressible Flow Effects in Scaled Pool Swell Tests 2.5 Vent System Impact and Drag Loads 2.7 Pool Swell Impact " Drag on Other Internal Structures 2.8 Froth Impingement and Fallback Loads 2.9 Pool Fallback Loads 2.10 Vent Header Deflector Loads 2.11 Condensation Oscillation Loads 2.11.1 Torus Shell Loads 2.11.' Downcomer Loads 2.11.3 Vent fjstem Pressures 2.12 Chugging Loads 2.12.1 Torus Shell Loads 2.12.2 Downconer Loads 2.12.3 Vent System Pressures 2.13 Safety-Relief Valve Discharge Loads 2.13.1 SRV Discharge Device 2.13.2 Discharge Line Clearing Transient 2.13.3 Air-clearing Quencher Discharge Shell Pressures 2.13.4 SRV Discharge Line Reflood Transient 2.13.5 SRV Air and Water Clearing Thrust Loads 2.13.6 SRV Discharge Line Temperature Transient 2.13.7 SRV Discharge Event Cases 2.13.8 Suppression Pool Temperature Limits 2.14 Submerged Structure Drag Loads 2.14.1 LOCA Water Jet Loads 2.14.2 LOCA Bubble Drag Loads 2.'4.3 Quencher Water Jet Loads } [4 3 } i

2.14.4 Quencher Bubble Drag Loads 2.14.5 LOCA Condensation Oscillation Drag Loads 2.14.6 LOCA Chugging Drag Loads 2.15 Secondary Loads 2.16 Differential Pressure Control Requirements 3. STRUCTURAL ANALYSES AND ACCEPTANCE CRITERIA 4. REFERENCES ~ 9 ii

. 1. INTRODUCTION The purpose of the Mark I Containment Long Tem Program is to perfom a complete reassessment of the suppression chamber (torus) design to inciade suppression pool hydrodynamic loads which were neglected in the original design, and to restore the original intended design safety margins of the structure. Thir reassessment will be accomplished by a Plant-ynique Analysis (PUA) for each BWR 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 b'.en developed from the staff's review of the Long Tem Program Load Definition Report (LDR), the Plent Unique Analysis Applications Guide.(PUAAG), and the supporting analytical and experimental programs conducted by the Mark I Owners Group. These criteria specifically address the dynamic loading 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). Similarly, references to original design criteria or original loading conditions shall be defined as those criteria or loading conditions which were found acceptable by the staff during the operating license review of the FSAR. For ease of reference, LDR refers to " Mark I Containment Program Load Definition Report," NED0-21888, PUAAG refers to " Mark I Containment Program Structural Acceptance Criteria Plant Unique Analysis Applications Guide," NED0-24583, and other supporting topical reports are referred to by their report numbers. A complete set of the references used in these criteria, listed in numerical order, is presented in Section 4. 2. SUPPRESSIONPOOLHYDRODYNAMIC[0 ADS 2.1 CONTAINMENT PRESSURE AND TEMPERATURE RESPONSE The pressure and temperature transients for the drywell 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, been found to provide conservative estimates of the containment response to a LOCA, by comparison to the staff's CONTEMPT-LT computer code. Plant-specific results for each break size shall be presented in the PUA, along with the input conditions, in sufficient detail to allow the staff to perform confimatory analyses to assure proper application of these models. )hb\\

. The timing and duration of specific loads are based primarily on the plant-specific containment response analysis for the pool swell-related loads, while the condensation periods are non-mechanistically maximized. However, the duration of the generic SBA condensation loads are assumed to be limited by manual operation of the Automatic Depressurization System (ADS) at 10 minutes into the accident. Therefore, as rart of the PUA, each licensee shall identify procedures (ircluding the primary system parameters monitored) by which the operator will identify the SBA, to assure manual operation of the ADS within the specified time period. Longer time periods may be assumed for the SBA in any specific PUA, provided(l) the chugging load duration is correspondingly increased. (2) the procedures to assure manual operation within the specified time period are identified, and (3) the potential for thermal stratification and asym.etry effects are addressed in the PUA. 2.2 VENT SYSTEM PRESSURIZATION AND THRUST LOADS The vent system pressurization and thrust loads shall be defined in accordance with the procedures set forth in Section 4.2 of the LDR, with the following exception. In order to assure the proper transition between vent clearing and bubble breakthrough for those plants that propose operation with a differential pressure control, the vent clearing time shall be derived from a containment analysis assuming no drywell/wetwell differential pressure and this time shall be applied to the vent system transients calculated from a containment response with the proposed drywell/wetwell differential pressure. In addition, for clarification, in the equation for F2V in Section 4.2.lc of the LDR, P3 shall be replaced by P2. 2.3 NET TORUS VERTICAL PRESSURE LOADS The downward and upward net vertidal pressure loads on the torus shall be derived from the series of plant-specific QSTF (Quarter Scale Test Facility) 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 mar-gins be applied to each loading phase:

UP = UP + 0.215 (UPmean} mean 2 00WN = 00WNmean + 2 x 10-5 (DOWN mean where "mean" refers to the average of the QSTF plant-specific test results (ibf). These margins shall be applied to the QSTF "mean" load function prior to scaling the load function to full-scale equivalent conditions. The margin for the downward loading function )hb\\

. shell be derived in tems of a fraction of t% lead at the time of the peak downward load, and that friction stell be applied to the entire downward loading phase. The margins specified above may be reduced or omitted where minimum conservatisms (i.e., smallest parameter deviation from the nominal plant condition over the range of tested conditions) in the QSTF tested conditions for a specific plant can be demonstrated by the application of the QSTF sensitivity test series (NEDE-23545-P). The sensitivity tests may not be used to adjust the mean torus vertical pressure loads. If the plant configuration is ct.anged to the extent that a single QSTF test series no longe. represents a conservative configuration of the plant, then a new series of QSTF tests shall be performed. Appli a tion of the sens:tivity tests and interpolation between plant-specit : test series wi.) be reviewed on a case-by-case bas k For those plants that use drywell/wetwell differential pressure control as a load mitigation feature, an add't.1onM structural analysis shall be perfomed assuming a loss of the differential pressure control to demonstrate the capability of the containment to withstand his extreme condition, as specified in Sections 5.3, 5.4, and 5.6 of the PUAAG. For this analysis, a single plant-specific QSTF test run may be used to define the loading function; however, the downward and upward loading phases shall be increased by the margins specified above for the base analysis. 2.4 TORUS POOL SWELL SHELL PRESSURES The spatial distribution of the torus shell pressures during pool swell shall be defined from the plant-specific QSTF test results and the azimuthal and longitudi,nal distribution factors defined in Section 4.3.2 of the LDR. However, the QSTF results shall be adjusted to incorporate the margins specified for the net torus vertical pressure loading function as -follows : 1. During the downward loading phase, the average pool pressure shall be increased by the equivalent differential pressure, as a function of time, corresponding to the margin for the downward load. 2. During the upward loading phase, the torus airspace pressure shall be increased by the equivalent differential pressure, as a function of time, corresponding to the margin for the upward load. 3. The pressure distributions shall be maintained such that the integral of the torus shell pressures will equal the net vertical pressure function with the margins included. \\hb\\

• 4 2.5

_ COMPRESSIBLE FLOW EFFECTS IN SCALED POOL SWELL TESTS The QSTF plant-unique test series are based primarily on a " split-orif :e" vent flow scaling relationship. Preliminary calculations perfonned by EPRI and GE indicate that compressibility effects, which could not be accurately scaled in the testing program, could result in a higher loading condition at full-scale conditions than that derived from ".icaled-up" test data. The original intent of these analyses was to provide justification for the scaled 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 upward vertical pressure loads, the torus pool swell pressure distribution, the vent header pool swell impact timing, and the vent header deflector impact timing. Based on our review of the preliminary analyses perfonned by EPRI and GE, which were presented to the staff in a meeting on July 24, 1979, we concluded that there is sufficient margin in the loading functions to justify proceeding with implementation of the Mark I LTP, while this assessment continues. We will require, however, that the Mark I Owners Group complete the assessment of compressible flow effects and justify the adequacy of these load specifications prior to the issuance of the staff's Safety Evaluation Report, which is currently scheduled for December 1979. In the event that the adequacy of the load specifications cannot be demonstrated, these loading conditions will have to be reassessed. The vent header and vent header deflector impact timing do not, however, appear to be sufficiently conservative. Based on our review of the material presented thusfar concerning the EPRI 1/12-scale three-dimensional pool swell tests and the compressible flow effects analyses, we conclude that the downcomer orificing used for the " split-orifice" tests do not provide a prototypical pool swell response. Thereford, the vent header and vent header deflector impact timing shall be derived from the " main vent orifice" tests (using the same longitudinal load distribution methodology in the LDR), until a flow distribution analysis, acceptable to the staff, can justify some less severe loading conditions.

2. 6 VENT SYSTEM IMPACT AND DRAG LOADS 2.6.1 Vent Header Impact and Drag Loads The load definition procedures set forth in Section 4.3.3 of the LDR are acceptable, subject to the following clarifications:

)hb\\

. 1. The experimental data of 'ocal vent header pressure in each of the Mark I plan

• 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 water impact on the vent header shall, be detemined from the QSTF plant-unique tests. 3. The plant-unique header impact timing (i.e., longitudinal time delay) shall be based on the EPRI " main vent orifice" tests as described in Section 2.5. 2.6.2 Downcomer Impact and Drag Loads The load definition procedures set forth in Section 4.3.3 of the LDR are acceptable, subject to the following clarifications. A pressure of 8 psid is to be applied unifomly over the bottom 50 of the angled portion of the downcomer, starting from the time at which the rising pool reaches the lower end of the angled section and ending at the time of maximuni pool swell height. The pressure is to be applied perpendicular to the local downcomer surface. The structural analysis for the downcocmer impact shall either be dynamic, accounting for the approximate virtual mass of water near the submerged parts of the downcomer, or a dynamic load factor of two shall be applied. 2.6.3 Main Vent Imoact and Drao loads The impact and drag loads on the main vent shall be evaluated in the following manner: 1. Subdividethesubmergedp$rtionofthemainventpipeinto six equally wide segments (see Figure 2.6-1). If this sub-division results in al < 0.3D fewer segments may be used such that al s 0.3D. 2. Detemine the velocity and acceleration histories at Points 1 through 7 in Figure 2.6-1 from the QSTF data and appropriate corrections for longitudinal variations along the torus (at Point 7, only the initial impact velocity is required). 3. Using the velocity components nomal to the vent pipe, calculate the impact and " steady" drag pressure using the method in Section 2.7.1 (Cylindrical Structures). At Point 7 only impact force is to be considered. 1431 226

. D l bl a A y Highest pool elevation V cos a pool elevation at initial contact with V a cos a main vent a V = pool impact velocity at station a = pool acceleration at station Figure 2.6-1 Schematic Diagram Illustrating the Methodology for Main Vent Impact and Drag 1431 227

, Using the acceleration components normal to the vent pipe, calculate the acceleration drag pressure using the equation Q+F Pa" g static buoyancy ) c / DL Where Pa is the acceleration pressure averaged over the pro-3 jected area (psi), p is the density of water (lbm/ft ), D is the diameter in feet, 9 the cross flow acceleration (ft/sec2) and g is the acceleration due to gravity (ft/sec2). c 4. Sum the pressures due to impact, viscous drag and acceleration drag and multiply by D to obtain force per unit length at Stations 1 through 7. 5. To obtain a smooth loading history for the main vent as a whole, the linear interpolation method suggested for the vent header deflector in Section 3.5 of NED0-24612 may be used. 2.7 POOL SWELL IMPACT AND DRAG ON_0THER_ INTERNAL STRUCTU E S The impact and drag loads for internal structures above the suppres-sion pool (except the vent header, downcomers, and vent header deflectors), as specified in Section 4.3.4 of the LDR, shall be modified such that the structures are classified as either cylindri-cal (e.g. pipes), exposed flat surfaces (e.g., "I" beams), or grant-ings. The following load specifications for each of the three structural classifications shall be used to replace the methodology in the LDR. Any structures that cannot reasonably be classified as one of these geometries will be r,eviewed on a case-by-case basis. The longitudinal velocity distribution shall be based on the " main vent" EPRI pool swell tests, as discussed in Section 2.5. 2.7.1 Cylindrical Structures For cylindrical structures, the pressure transient which occurs upon water impact and subsequent drag is depicted in Figure 2.7-1. The parameters in Figure 2.7-1 shall be defined as follows: 1. The maximum pressure of impact P will be detemined by max 7 [)44 1 o V P = 7.0 x max \\ 9c/ where Pcax is the maximum pressure averaged over the projected 3 area Osi), o is the density of water (lbm/ft ), V is the impact is the acceleration due to gravity velocity)(ft/sec) and gc (ft/sec 1431 228

. P max en m U d iS P g P D =a> / / I TIME 7 Figure 2.7-1 Pulse Shape for Water Impact on cylindrical Targets 5 red < 5 x 10 1.5 / {/ -l. l l. l~ l l-I l i + 'i 1 5-C f' j { -t-red > 5 x 10 - D t 1.0 I i ~l t 1 i .J..__ i ..l. ...{ i ;i 1. I .l i l 4 l j i .t-- 1 i + i 1 I I t t I I f i 'I ~ t i 5 -{.-- 1 I j l 4 I i l i t..__ I j i i -i. G .q. __ I. ~ ~1 0I l I it i ~'-+- 2. u. 6. e.10 20. uo. 60. 80. 2 FRoUDE NUMEER (N = V /Dg) p Figure 2.7-2 Drag Coefficient for Cylinders Following Impact i431 229

. 2. The hydrodynamic ' ass per unit area for impact loading shall be obtained frorr.ne correlation (cylindrical target) depicted by Figure 6-8 in NEDE-13426-P. A margin of 35". will be added to this value to account for data scatter. 3. The impulse of impact per unit area shall be detemined by: h b V Ip"T T4Tg i cj hydrodynamic mass per unit area (lbm/ftgsi-sec), 4/A is the where I is the impulse per unit area ( p ) and V is the impact velocity (ft/sec). 4. The pulse duration will be detemined from the following equation: = 2I /P, T p 5. The pressure due to drag following impact shall be determined by: C 2 ) D -y D F 124g c/ ~ where P is the average drag pressure acting on the projected D area of the target (psi), Cn is the drag coefficient ag defined by Figure 2.7-2, and o is the density of water (lbm/ft ). 2.7.2 Flat - Surface Structures For flat-surface structures, the pressure transient which occurs upon water impact and subsequbnt drag is depicted in Figure 2.7-3. The parameters in Figure 2.7-3 shall be defined as follows: 1. The pulse duration (t) is specified as a function of the impact velocity: T = 0.0016W for V < 7 ft/sec r = 0.011 W for V > 7 ft/sec V where W is the width of the flat structure (feet) and V is the impact velocity (ft/sec). 1431 230

. 2. The pressure due to drag following impact shall be Gatemined by: C 2 D pV P = D 2 144 gc/ is the average drag pressure acting on the frontal where PD area of the structure (psi), CD is the drag coefficient (C-2, flat strips nomal to flow, independent of Reynolds number. 3 and e is the density of water (lbm/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. A margin of 35% shall be added to this value to account for data scatter. 4. The impulse of impact per unit area shall be detemined by: b [144g) V I = p A g c where Ip is the impulse per unit area (9-sec), M /A is the H hydrodynamicmassperunitarea(lbm/ftI),andVistheimpact velocity (ft/sec). 5. The maximum pressure (Pmax) shall be calculated from the impulse per unit area and the drag pressure as follows: 2I P + P P,3x = D 2.7.3 Gratings The static drag load on gratings in the pool swell zone of the wetwell shall be calculated for gratings with open areas greater than or equal to 60% by forming the product of the pressure dif-ferential (figure 2.7-4) and the total grating area (not only the area of the metal bars). The pressure differential curve in Figure 2.7-4 is based on a velocity of 40 ft/sec. If the maximum pool velocity in the area where gratings are located differs from 40 ft/sec, the force on the grating will be calculated as follows: F = AP x A Y grating max 40 ) To account for the dynamic nature of the initial loading, the load shall be increased by a multiplier given by: F /D = 1 + [1+(0.0064Wf)2ykfor SE Wf < 2000 in/sec. 1431 23I

. wa =3 P m max m wan. o<n..r w c.o e. I P P f' D / / / I TIME T Figure 2.7-3 Pulse Shape for Water Impact on Flat Targets _..._i_._.. ._.__i . _. 3 _.___7_.-.

~

20 .c_ s _._..g.. _ p ___ ...___.4._..._ V = 40 ft/see. ...N.....__ . _ _. g._ _ _...) 3. ... _a _g. _ ._4____. ..__..4.. 15 .p N.._..__._ Z ___-__.,.--J- __..g.___. L_. __..-i_-._. ....L__.. a.... m .__ x. g - __ _ _ g _. o .__A.__ ._ m.,4... _ _ ____._4_._..____. t_.___ .__._j._____ _.4 10 _....._-._A_. _.y L....- _ _ 4.._ a_ y_...._.

t... _ _..

3 _....t._. m w .... _ 9 ___...._t__ .. _.. - j e ..i ,m _..._,t.--- -.. _ _ _, - _ - - -.._.._--.}__.._. _j--- ..._. q ._._a . _L. ._ _ _] _._.. _._}. .. p p. 0.5 0.6 0.7 0.6 09 1.0 OPEN AREA FRACTION Figure 2.7-4 Pressure Drop Due to Flow Across Gratings PGCRORGINAL a 232

- 12 where: F = static equivalent load W = width of grating bars, in, f = natural frequency of lowest mode, Ez D = static drag load If Wf > 2000 ft/sec (not expected for gratings) the force on the bars of the gratings will be calculated by the method outlined above for flat-surfaced structures. 2.7.4 Load Apolication These load specifications correspond 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, except for the gratings. The value of the hydrodynamic mass shall be obtained from the appropriate correlation in Figure 6-8 in NEDE-13426-P. When the impact loading is primarily impulsive and talculations have already been performed in accordance with the LDR methodology, simple adjustments may be made to the LDR analyses. Under these conditions, a parabolic pulse shape, as proposed in LDR, is acceptable provided corrections are made to account for the 35% margin in the impluse and with additional corrections for the drag force imediately following impact. For structures with a natural frequency less than 30 Hz, loading can be treated impulsively (i.e., independent of pulse shape) when the conditions fall into the region above the straight line shown in Figure 2.7-5. l The following corrections must be applied to the previously calculated stresses: 1. The calculated stresses will first be multiplied by a factor of 1.35 to account for the data scatter in the impulse data. 2. The calculated stress.s will then be multiplied by an additional factor to account for the presence of drag fol. awing the impact. This factor is determined as follows: 1431 233

, TYPICAL IMPACT STRUCTURES: a -'s -~ s s s / \\ / \\V[ \\ I l l 1 J t I l %) v '/ \\ ~< W 4-- W --> 4-- W -> l. ll !} l l' T.._ l 15 - q-q. ij. i+;j ,-a=f- ..... B : - :l. -= I k.- i lL /.._ -l l- - ij -- - - ti=3-ij ' _.

j.

l - / jj;- ... 22 . __. : 4 2 a - ; -...:l f g....,... g 10 - ~ ..j. IMPULSIVE LOAD __' } REGION ~~ ,/ - ~l. l- -l e t (g f i 1.- .i -.: 1-VI i E l -1. i i /l i i i I I

5

/.i ~ il 8 /i -1 -l- -l ~i i E I/ l .i / j. .._.)

L :. ij:_g_.

0 f --j - - +'-- p-l~ -" 4. _ _ _c

-- _1 x __...

0 1 2 W (feet) Figure 2.7-5 Impulsive Impact Loading Region for Structures with a Natural Frequency Less Than 30 Hz P00R BRIllWL um nr

14 - Calculate the drag pressure,(PimpactvelocitybythemethodsNg a. d Sections 2.7.1 or 2.7.2. b. Form the ratio: Pdrag/Pmax where Pmax is the amplitude of the parabolic pulse used in the original stress analysis multiplied by 1.35. c. Detennine the dynamic load factor (DLF) from Figure 2.7-6, corresponding to the two cases: (1)parabolicpulsewithout drag and (2) parabolic pulse fcllowed by drag. d. Multiply the calculated stress by the factor DLFwith drag /DLFw/o drag 2.8 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 specification was derived from the high-speed film records of various pool swell tests and an analysis of pool acceleration following vent header impact. The 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; 2 Y ~Of P = f 144 gc where: P = froth impingement pressure (psi) f 3 pf = froth density (lb /ft ) m = froth impingement velocity (ftf)sec) V = gravitational constant (ft/sec gc 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 is 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 P00ROR;GNAL um ao

a P ma p F hv g u! / / / s r TIME i

ijp

i

_ j..r- /j. j m + 1: / A L. 2.0 - p E! M N drag ~ ~ .i~. P d' (~~ F ifi T5d ax ~;'iii._. 2b_iE +'~. [:6[d[.((j " ~~

iF 5l~; i

Mi /!0%Ld((M x-

.x : -

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-p f y' y.9f -h,:( 7 f. j 1 = =.!!- [ E- / E [ Nf3 M / [ 1 - - S - - \\ - - ;.nz # ilD X f i f A f S f =/ Vy' My1 g -l #ipipp XMi. - J {

• f il ME ![Ni[i((H. 3i.Z I

e 3 M :i ggiMM u

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f'. f =.. r. 0 ) i '.. !.d}.. 1/: 1- 'i. P,^ S [ jji[ ;l[ [ ('~~~'-~~ q ' ET.._.J. J' M ~}:=,7bM2[!-E E... .E_!5E : :! . ii::5 ~ -'D = 2 ~ ii}/!ddi%l"E~' ~ I"Eb ~. - +.: iil! " ~~ '='~2!! y -gygsgq=. atggir a: --- g:- -. :s. m i N = / % L e tiiE4.= E! F!= Y '! eel =T

• -^ -

= Natural Period i =-v.J_ J. v. : =. . :.1 :- N 71. .{ 4 ,11C_~ . ;~ Li _ .--j..-...._- f"i-q~* lE F -:l': 'j~~~ =i 0. ~ O.1 0.2 03 0.4 05 t/T N Figure 2.7-6 Effect of Drag Following a Parabolic Impact Pulse P00RORBlK a zu

, or equal to one foot, and a p portionately lower density for structures greater than one foot; i.e., p = (0.2/x) pw, where x is the dimension in feet. The load shall be r.pplied in the direction most critical to the structure sithin the 90' sector bounded by the horizontal opposite the vent header to the vertical upward as shown in Figure 2.8.1 The load shall be assumed to be a rectangular pulse with a duration of 80 milliseconds. Region II: The froth velocity shall be based on a source velocity equal to the maximum pool surface velocity, directly beneath the structure under consideration, which is corrected for subsequent deceleration from the elevation of the maximum velocity. 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, 25% water density for structures greater than one foot, and 10% water density for structures located within the projected region direct n bove the vent header. The load shall be applied in the direction most critical to the structure within the + 45* sector ~ of the upward vertical. The load shall be assumed to be a rectangular pulse with a duration of 100 milliseconds. Fallback: The froth fallback velocity shall be based on the freefall velocity fmn the upper surface of the torus shell directly above the subject structure. The froth density shall be assumed to be 25% water density, with the exception of the projected region directly above the vent header which is 10% water density. The load shall be assumed to directly follow the froth impingement load, with a duration of one second. 2.9 POOL FALLBACK LOADS The proposed load definition procedures set forth in Section 4.3.6 of the LDR for suppression pool fallback loads on internal structures following pool swell are acceptable. 2.10 Vent Header Deflector L: % The load definition procedures set forth in Section 4.3.9 of the LDR 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: 1431 237

VENT HEADER q> Tonus 1 Figure 2.8-1 Direction of Load Application for Froth Region I \\hb\\

. 2.10.1 QSTF Deflector Load-An individual plant may choosc to use deflector load data taken directly from the 1STF plant-unigt'a tests. This technique is subject to the foll wing requirements:

1. If the QSTF deflector load measurement does not have a sufficiently fast response time to resolve the initial impact pressure spike for the deflector types 1 - 3, inclusive, the loading transient shall be adjusted to include the empirical vertical force history of the spike shown in Figure 2.10-1.

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

2. The QSTF plant-unique loads shall be adjusted to account for the effects of (a) caphet time delays and (b) pool swell velocity and acceleration differences which result from uneven spacing of downcomer pairs. The longitudinal load variation shall be ualuated at the instant when the undistrubed pool surface would have reached the local elevation of the center (half-height elevation) of the deflector. The three-dimensional load variation shall be based on the EPRI " main vent orifice" tests, as discussed in Section 2.5.

3. In applying the load to the deflector, the inertia due to the added mass of water below the deflector shall be accounted for.

The added mass per unit lengtn of deflector may be estimated by: M I9 H c VW where: M = hydrodynamic inass per unit length (lb/ft) H I = total impulse per unit length associated with the impact transient (lb-sec/ft) V = impact velocity (ft/sec) w = deflector width (ft), as shown in Figures 2.10-2 through 2.10-5 gc = gravitational constant 1431 239

19 - F = Vertical upward force on deflector per unit length 8. d = Diameter of cylinder in deflector types 1 - 3 .1 V = Impact velocity \\t l r o = Water density .k;_. \\ - j t = Time from begining of impact 6.

r :

7 2 .h i pVd 2 N .i \\ l 4' \\ I .\\' l l\\ l h, l i t 2.- s\\ 1 \\ i -il l \\ l . __. _ q.. : :- - i i O. O.1 0.2 Vt / d Figure 2.10-1 Impact Force Trarisient for Addition to the Empirical Data for Deflector Types 1 - 3. j 1431 240

, 2.10.2 Analytic Deflector Loads The deflector load definition which is based on empirical expressions for impact and drag forces together with 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 drag" contributions to the load shall be computed from the correlations shown on Figures 2.10-2 through 2.10-5, for deflector types 1 - 4, respectively. For times past the periods shown, the last value shall be extended for the duration of the transient. 2. The three-dimensional load variation and timing shall be based on the EPRI " main vent orifice" tests, as discussed in Section 2.5. 3. The gravitational component of the acceleration drag shall be included in F, as defined in NED0-24612. A 4. In computing the deflector response to the load, the added mass of the water shall be accounted for, as described in Section 2.10.1.3 above. '2.11 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 phenomena and, therefore, statistical variance or load magnitude uncertainty cannot be establishbd from one test run. Although we conclude that the M8 tested conditions are conservative 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 t~ sis to proceed with implementation of the Mark I Long Tem Program. We will require that the Mark I Owners Group confirm the condensa-tion oscillation loads (i.e., torus shell loads, downcomer lateral loads, vent system pressure, and submerged drag source) by perfoming a sufficient number of additional large break, liquid blowdown tests in FSTF to establish the uncertainty in the load magnitudes. 1431 241

~ e-d-e u V 8. i l I t _, _ ._.-_{__.____---.-- i i i \\ l I l

6. \\

5 \\ i i i i i i 7 2 yd 2 \\! I I I i \\: l i i l i i

• i (l

l l i I i } l \\ i \\; (c ) from Figure 2.10-2b i _ o I (5 i 1 2. j k I i i I 9 j jt l l i-I l 'g I i i j) { I O. ' n 0.136 05 10 Vt/d Figure 2.10-2 Impact and Steady Drag Force Correlation for Type 1 Deflector J j j 1431 242

} Z-. T }. ] ' g u _ _,] - -[. If [ _ _ :.b]- PE < 5 x 10 --- 5 -El ..n { ...{ .q. {.. j} . -..q. -- -g. {. . { - { -} (C )~ ~ ~ D 1- .. _.i.....b .i.. ZZ '. :. ~~ ~ __, ; _ _; q i-1-- ~ ~ ~ T_- t 'J' }' pVd > 5 x 105___ .i_ gi i 1i. j i l' l~ ~

j. j

~J j . j- - .i .t-I llll U i I i l ll; l -l

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-l l t I i o. l f -l l { ~ 1. 5. 10. 50. 100. 2 V /gd Figure 2.10-2b (C ) f r Type 1 Deflector D j 1431 243

w I (V' d=0.707w 1 n V .7__.._,__. l I i e '+. 95 i j a.. + i l i i l l i i i 1 l I l 1 i j i. i. l f u i l i i i i p \\ N* I 2 \\ pVW 2 i

t '

I ~ V I i .__../ i s' it i. .i k e' !L ____ i o.7 3,/ ' i, j i N i / . 'I i i } s !,/ g _ _ _ _ _. _ _. _ I 7 J' i 0. t \\ .207 0. 0.136 05 0.854 1. Vt/d Figure 2.10-3 Impact and Steady Drag Force Correlation for Type 2 Deflector O lkbl

+- w - - d=0.414w 1 JL V 1 q. t ...2. i [T I l i '1 p2.90; ,x i _2 \\ i 1 i l ^ p i pVW Sa 1 2 .__1_____.___. -~- b ' O.7 1 fg s --7 .....;._....._.'..g / \\ 3 I a \\; f r._ __ 0 - i \\ .207 0, 0.136 1.0 1.60 2.0 vt/d Figure 2.10-4 Impact and Steady Drag Force Correlation for U'e 3 Deflector I I j ,w u,n

Wh

• 90 JL V

4 i. ....-5,12. -

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r... 4-3 l .jii :.2i:d..._ I 5 :n d i. 3. . i. : 5....*.. _ _ ~- - Eii-- -_i:: - :i=

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ii

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-j=~

3 EIE\\ii~ = =i =

i =~~ ~ %h 2 pVw j .;j.y .g;._. ;.jg.._ _.gg.x .... j _ _ _ Y '\\ ~ ~~ 2-i r -1 ...g ) fl \\ ~T-i if i N 1 ! :/I l i:\\. ~ I 1' /i _.:l\\ . l6. 7 i ./ !~ 1: _ . 1.:k ' + .1- .l. . ~/ l 2 . ii.' -.'_iH : __ lU,..\\.!- [j l _..:l - ...i..: 4. . _.:_\\i.. -i-O. 1. 1.5 2. 2Vt/w Figure 2.10-5 Impact and Steady Drag Force Correlation for Type 4 Deflector POOROR8NAL 1m a6

. 2.11.1 Condensation Oscillation Torus Shell 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 crafirmation: 1. Provided that the " rigid wall" load derivation technique described in NEDE-24645-P is demonstrated to be conservative, in response to Question 7 in our request for additional information (D. Eisenhut, NRC, to L. Sobon, GE, dated July 30,1979). 2. Provided that sufficient justification can be provided to exclude a condensation oscillation asymetric loading condition, in response to Question 2 in our request for additional information. 3. Provided that the uncertainty in the load magnitude is demonstrated to be less that the demonstrated conservatisms in the load specification, by the testing program 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 loading for the IBA is a continuous siausedial function with a peak amplitude and frequency range of that specified for the " pre-chug" load. We will require that the conservatism in the IBA condensation oscillation loads be demonstrated as part of the response to Question 7 in our request for additional infonnation for the specific flow regimes of interest. 2.11.2 CONDE!!SATI0flOSCILLATIONDOWilC0bERLOADS 2.11.2.1 Untied Downcomer Loads The condensation oscillation downcomer lateral loads for " untied" downcomers shall be defined as described in Section 4.4.3 of the LDR, based on the methodology in NEDE-24537-P. However, in computing the dynamic load factors I \\ DLF = P) g max \\0\\ 20

. where: P = maximum static equivalent lateral load for max plant-unique downcomer P) = maximum static equivalent lateral load in FSTF DLF = plant unique downcomer dynamic load factor DLF) = FSTF downcomer dynamic load factor the plant-unique loading condition shall be derived as follows. The plant-unique DLF shall be calculated using a damping value consistent with the requirements of Regulatory Guide 1.61, " Damping Values for Seismic Design of Nuclear Power Plants," and a natural frequency determined from the structural analysis of the downcomer - vent header system. The plant-unique driving frequency shall be specified as that frequency in the range 4 - 8hz which produces the maximwa structural response. The natural frequency and damping values for the FSTF DLF shall be conservatively established from a " pluck" test of an untied downcomer in FSTF, with a nominal water level of 3 feet 4 inches and an amplitude in the range of the response level. The driving frequency for the FSTF DLF shall be assumed to be 5.5hz. ,2.11.2.2 Tied Downcomer Loads The condensation oscillation downcomer loads for " tied" downcomers, as described in Section 4.4.3 of the LDR, are unacceptable. We will require that a load specification be derived from the maximum dynamic load components on each downcomer in a tied pair. The Toad definition and structural analysis technique shall be confinned by comparisons of the predicted structural responses to the measured strains in the FSTF vent header and tie-bar. The FSTF natural frequency and damping values shall be conservatively established by performing a " pluck" test for a tied downcomer pair in FSTF, with a nominal water level of 3 feet 4 inches and an amplitude in the range of the response level. 2.11.3 Condensation Oscillation Vent System Pressure Load The load definition procedurc set forth in Section 4.4.4 of the LDR for the oscillatory pressures in the vent system during the condensation oscillation period, are acceptable subject to confirmation by the additional testing as described above. 1431 248

. 2.12 CHUGGING LOADS 2.12.1 Chugging Torus Shell Loads The load definition and assessment procedure set forth in Section 4.5.1 of the LDR for the chugging condensation loads on the torus shell are acceptable, 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, to L. Sobon, GE, dated July 30,1979). This load specification shall be used in conjunction with a coupled fluid-structure analytical model. 2.12.2 Chugging Downcomer Loads 2.12.2.1 Untied Downcomer Loads The chugging lateral loads on untied downcomers 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 exceptions: 1. The load specification for comparison to the ASME code primary stress limits shall be based on the maximum taeasured resultant static equivalent load in FSTF. 2. The fatigue usage analysis for each downcomer shall be based on a statistical loading with a 95% probability of non-exceedance. 3. The multiple-downcomer loading to assess statistical directional dependence shall be based on a non-exceedance probability of 10-4 per LOCA. 2.12.2.2 Tied Downcomer Loads For tied downcomers, the strains in the downcomer itself shall be evaluated exactly as in the case of the untied downcomers, using tied downconer data. The strain in the tie bar shall be evaluated by assuming that one of the two tied downcomers is subjected to a dynamic load of triangular shape, with an amplitude of: RSEL (RSEL = Resultant Static F = max x ft Equivalent Load) d where RSEL is the maximum measured RSEL for an untied downcomer during chugging, f is the lowest natural frequency of vibration of an untied downcomer for the specific plant, and the duration of the load, t, shall be assumed to the 3 milliseconds. The load d direction shall be taken as that (in the horizontal plane) which result in the worst loading condition for the tie bar and its attachments to the downcomers. 1431 249

_ T.2.3 Chugging Vent System Pressure Loads The load definition procedure set forth in Section 4.5.4 of the LDR for the oscillatory pressures on the vent 5 stem during the chugging period are acceptable. 2.13 SAFETY-RELIEF VALVE DISCHARGE LOADS 2.13.1 Safety-Relief Valve Discharge Device The acceptance criteria set forth below for the quencher discharge loads and submerged structure drag load soi rce strengths are applicable only to the "T" quencher config Jration described in Secticn 1.1 of NEDE-24542-P. For plants ising other types of quencher discharge devices, the SRV disc' arge load definition, r submerged structure drag load source strength, and pool temperature limits will be evaluated on a plant-specific basis. 2.13.2 SRV Discharge Line Clearing Transient The load definition and assessment procedure, described in Section 5.2.1 of the LDR, for the pressure and thrust loads on the SRV discharge line and quencher, which is based en the methodology presented in NEDE-21864-P and NEDE-23749-1-P, is acceptable. 2.13.3 SRV Air-Clearing Quencher Discharge Shell Pressure Loads 2.13.3.1 Methodology for Bubble Pressure Prediction The load definition procedures described in Section 5.2.2 of the LDR and the methodology in NEDE-21878-P for predicting the quencher bubble pressurs are acceptable, with the following exceptions: 1. The load definition procedures set forth in Section 5.2.2 of the LDR are acceptable for SRV discharge line water-leg lengths less than or equal to 13.5 feet. In the event that the water-leg length for a particular discharge line exceeds 13.5 feet, the load prediction for a 13.5 foot water-leg length shall be used. 2. 'ae proposed methodology for predicting bubble pressures 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 sub-quent actuations. 1431 250

. 2.13.3.2 Methodolo" for Torus Shell Pressure Prediction Based on the predicted air bubble pressure-time histories, as discussed above, the torus shell pressure.; at various locations in the suppression pool shall be calcula.ted 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 supple >nent to the LDR. 2.13.3.3 Multiple - Discharge 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 accropriate pressure attenuation model, as discussed above. All bubbles shall be assumed to oscillate in-phase with the frequency ranges specified below for both first and subseauent actuations. 2, In the event that the co,mbined 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. 2.13.3.4 Frequency of Pressure Wave Form The pressure wave form predicted by the methodology described in Section 5.2.2 of the LDR within the following uncertainty ranges (stretched or compressed time scale) that will produce the maximum structural, equipment, or piping system response shall be used for the design assessment: 1431 251

. 1. First Actuati. - the frequency range shall be 0.75 time:; the minimum predicted frequency to 1.25 times the maximum predicted frequency. 2. Subsequent Actuation - the frequency range shall be 0.60 times the mimimum predicted frequency to 1.40 times the maximum predicted frequency. 2.13.4 SRV Discharge Line Reflood Trans_fent 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. 2.13.5 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. 2.13.6 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. 2.13.7 SRV Discharge Event 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 historp. The following load cases shall be considered for the design assessment: 1. A first actuation, single valve discharge shall be considered for all event combinations involving SRV events. Single valve subsequent actuations shall be considered for the SRV, SBA, and IBA event combinations, as detemined from a plant-specific primary system analysis. 2. Asymmetric SRV discharge, both first ard subsequent actuations, shall be considered for SRV, SBA, and IBA event combinations. The degree of asymetric discharge for each event combination shall be determined from a plant-specific primary system analysis designed to maximize the asymetric condition. \\ 0\\ 252

. 3. ADS valves discharging on first acutations shall be considered for the SBA and IBA event combinations, fol' wed by subsequent acutations determined from a plant-specific primary system analysis. 4. The maximum number of valves that will actuate for the SRV event combinations shall be determined from a plant-specific primary system analysis for the design basis transients, which assumes that all valves actuate at their set-Doint pressures. All first actuations shall be assumed to occur in phase, followed by subsequent actuations determined from the primary system analysis. All of the event combinations above include the earthquake events (OBE and SSE) in combination with the SRV discharge events. 2.13.8 Suopression pool Temperature Limits As part of the PUA, each licensee is required to either demonstrate that previously submitted pool temperature analyses are sufficient or provide plant-specific pool temperature response analyses to assure that SRV discharge transients will not exceed the following pool temperature limits. 1. Local Temperature Limit The suppression pool local temperature shall not exceed 200 F throughout all plant transients involving SRV operations, for any quencher device that has (a) the hole diameter equal to and (b) greater than or equal hole spacing than that of the generic Mark I T-Quencher. 2. Local and Bulk Pool Temperature The local to bulk pool temperature difference shall consider the plant-specific quencher discharge geometry and RHR suction and discharge geometry. The analysis of the plant-specific local to bulk pool temperature difference shall be supported by test data from either the existing Monticello pool temperature data or in-plant tests. The " local" temperature is defined as the temperature in the vicinity of the quencher device during discharge. For practical purposes, the wster temperatures observed in the sector containing the discharge device at shell locations on the reactor side of the torus downstream of the quencher centerline at the same elevation as the quencher 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. 1431 253

. 3. Suppression Pool Temperature Monitor System The suppression pool temperature monitoring sys' a is required to ensure that the suppression pool is within the allowable limits set forth in the plant Technical Specification. The system shall meet the following design re';uirements: a. Each licensee shall demonstrate that there is a sufficient number and distribution of pool tempera-ture sensors to provide a reasonable measure of the bulk temperature, b. Sensors shall be installed sufficiently below the minimum water level, as specified in the plant Technical Specifications, to assure that the sensor properly monitors pool temperature. c. Pool temperature shall be recorded in the control room. A sufficient number of temperature monitors shall be provided to permit the operator to establish the bulk pool temperature. Operating procedures and alarm set points shall consider the relative accuracy of the measurement system. d. Instrument set points for alarm shall be established, such that the plant will operate within the suppression pool temperature limits discussed above. All sensors shall be designed to seismic Category I, e. Quality Group B, and energized from onsite emergency power supplies. 2.14 SUBMERGED STRUCTURE DRAG.E0 ADS 2.14.1 LOCA Water Jet Loads The load definition and assessment procedure described in Section 4.3.7 of the LDR, which is based on the " Moody Jet Model" (NEDE-21472-P), : 1 acceptable subject to the follow-ing constraints and/or modifications: 1. The plant-specific jet discharge velocity, V (t), and D acceleration, ap(t) = dV /dt (t), from the QSTF plant-D specific test series shall be used as the driving soarces for the jet model. 2. Forces due to the pool acceleration and velocity induced by the advancing jet front shall be computed for struc-tures that are within four downcomer diameters below the downcomer exit elevation, even if the structure iA31 25A

. is not intercepted by the jet. The flow field shall be computed by modelling the moving jet front as a hemispherical cap centered one downcomer diameter (D) behind the " Moody" jet front positions, containing the sar.2 amount of water as the " Moody" jet, and moving with the velocity of the " Moody" jet front. The formulas for the hemisphere radius (R ) and the trajectory 3 of the hemisphere center (x ) are. c R (t) = (f +(3x (t)/D)b) 1/3 for x (t) > D f f R (t) = 9*f(t)l/3 for x (t) 1 D f 3 ( ) 2D x (t) = x (t) -0 for x (t) > D c f f x (t) = 0 for x (t) 1 D c f where x (t) is the position of the " Moody" jet front as a f function of time, as computed in NEDE-21472-P. Using formulas 1 and 2 in NEDE-21472-P and assuming an average constant acceleration of the particles contained within one downcomer diameter behind the " Moody" jet front, the cross-sectional area in this region can be approximated by: 2 (1 + 1/(1-x/x )b) A (x,t) = nD f 8 where xf(t) is the " Moody" jet front position as computed in NEDE-21472-P. The' volume contained in this portion of the jet can be obtained by integrating A(x,t) from (xf -D) to x for xf greater than D, and from x=0 to x = xf for xf f less than D. When the jet is modelled by a more realistic hemispherical cap, while conserving the total volume of the fluid, the cap radius and position is given by the equations above. The equivalent uniform velocity and acceleration at the location of the structure (x,y) shall be obtained from the time dependent potential 4)(x,y,t) induced by the jet front: 1431 255

35 - si-1 j[R)! l (x - x,) dx fR S f dR \\ _g 8 I rj 4j (x,y,t)= (_r / (dt / 2 (r j dt where' r = {(x-x )2 + y [' and y is the transverse distance 2 ofthestructurkfromthejetaxis,and(x-x)isthedistance fromthestructuretotheeffectivejetfronfcenteralongthe jet axis. The potential is the superposition of the expansion and motion of the sphere as given in any standard hydrodynamics text (e.g. Milne Thomposon, Theoretical Hydrodynamics, Fourth Edition, pp. 455-556). The local uniform flow velocity is U,(x,y,t)=vej as in NEDO-21471,while the acceleration is a(x,y,t) = BU. at This calculation need only be performed for r> R, and x > x. g If either of these conditions are not satisfied, the methodology in the LDR will bound the load and is, therefore, acceptable. 2.14.2 LOCA Bubble Drag Loads The load definition and assessment procedures described in Section 4.3.8 of the LOR, which are based on the methodology in NEDO-21471 and experimental confirmation in NEDE-23817-P, are acceptable subject to the following constraints and/or modifications: 1. Flow Field a. 05TF plant-specific test results (NEDE-21944-P) will be used. b. Model E in NEDE-21983-P will be used for the method of images simulation of the torus cross-section, c. After contact between bubbles of adjacent downcomers, the pool swell flow field above the downcomer exit elevation will be derived from the QSTF plant-specific tests. 1431 256

. 2. Drag Load Assessment a. Drag forces can be computed for circular cylinders as given in NEDO 21471, but a conservative drag coefficient of CD = 1.2 must be assumed, independent of the Reynolds number, b. Drag forces on structures with sharp corners (e.g., rectangles and "I" beams) must be computed by cons,idering forces on an 4 Lmt.x, where Lmax is equivalent cylinder of diameter Deq = 2 the maximum transverse dimension. Lmax is defined as the dia-meter of a circumscribed cylinder about the cross-pection of 2 the structure. For example, Lmax equals (a2 + b f2 for a rec-tangular cross-section of sides a and b. c. Long slender structures must be considered in segments of length (L), which do not exceed the diameter (D or Deg). Alternatively longer segments may be used as long as the equivalent uniform flow velocity and acceleration are evaluated conservatively for every point on any such segment. d. Interference effects due to the proximity of walls shall be considered for 23ch structural segment that has its center less than 1.5 diameters from a boundary. Interference effects between neighboring structures shall be considered whenever the centers of the segments are less than 3D, where D = 1/2 (D) + D ), the average diameter of the two structures. 2 For structures near walls, the multiplier (1 + Aw) shall be used to increase the acceleration drag and the multiplier (1 + Dw) shall be used to increase the standard drag. Aw and Dw that bound theory and experiments are given below as functions xw (xw = r/D - 1/2, where r is the distance from the segment center to the boundary).

0. 05 < x, < l. 0

= 0.05/x, Ag D,= 0.12/x, x < 0.05 A = 1. 0 g g D,= 2.4 For structures with neighbors that are less than 35 away and within 30' of being parallel, the multipliers (1 + Ag)dard and (1 + D ) shall be applied to the acceleration and stan I drag components respectively. Bounding expressions for AI and Di are given below as functions of x1 (x1 = rl2 D - 1 / where rl2 is the distance between segment centers). \\ n\\ 251

. 0.2 D2 \\ 0.05 < XI<2 AI=Xi' o, + o ) g DI = 0.2/XI where D' is the diameter of the structure under consideration and D2 's the diameter of the neighbor. If more than one neighbor must be considered, the AI and DI values may be sumed over the neighbor structures. For x1 less than 0.05, the two neighbor structures shall be considered as an effective single structure. The effects of wall proximity and neighbor structures may be superimposed in ordei to compute overall multipliers as follows: (1 + Aw + IAIk) k (1 + Dw + IDIk) k In situations where interference effects must be considered, but the correction techniques outlined above are not applicable, a detailed interference effects analysis shall be performed. 2.14.3 QUENCHER WAThrt JET LOADS The load definition procedure described in Section 5.2.4 of the LDR, which is based on the methodology in NEDE-25090-P is acceptable, subject to the appropriate documentation of the confimatory tests discussed in NEDE-25090-P. 2.14.4 QUENCHER BUBBLE DRAG LOADS The load definition and assessment procedures described in Section 5.2.5 of the LDR, in NEDE-21878, and irr NED0-21471-2, are acceptable subject to the following constraints and/or modifications: 1. Flow Field a. The detemination 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. b. Drag loads on the quencher arms and the SRV discharge line shall be computed on the basis of asymetric bubble dynamics. Either a full 1800 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. c. Model E in NEDE-21983-P shall be used for the method of images representation of the torus cross-section. 1A31 258

. 38. 2. Drag Load /,essment a. Drag forces for circular cylinders shall be computed on the basis of acceleration drag alone, under the condition that UmT/D < 2.74, where Um is the maximum velocity T 1s the period of bubble oscillation, and D is the cylinder diameter. For UmT/D > 2.74, the standard drag shall be included with the drag coefficient CD = 3.6 in order to bound the relevant experimental data. b. 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. 2.14.5 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: 1. Flow Field a. An average source strength shall be established by considering equal source strengths at all eight downcomers in equation B-4 in NEDO 25070. A me.ximum source strength shall be defined as twice the average source strength. For each structure,the loads shall be computed on the basis of both the average source at all downcomers and for the maximum source applied at the nearest downcomer. b. The fluid-structure inter, action 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. 2. Drag Load Assessment a. The constraints and modifications specified for the quencher bubble drag loads apply, b. These loads may be applied quasi-statically to structures, only if the highest significant Fourier comnaents occur at frequencies less than half the lowest structural frequency. 1431 259

, 2.14.6 LOCA CHUGGING DRAG LOADS The load definition and assessment procedures described in Section 4.5.2 of the LOR and the methodology in NEDO-25070 for the pre-chug drag loads are acceptable nbject to the constraints in Section 2.14.5 for the condensation oscillation drag loads. The application for the post-chug drag, loads is subject to the following constraints and/or modifications. 1. Flow Field The maximum source strength history shall be obtained by using a. the maximum measured pressure (not necessarily at the bottom center) in a Type 1 chug in equation B-4 of NEDO 25070, with For each structure, f(r) based on the single nearest downcomer. 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. The fluid-structure interaction effects shall be included for b. 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. 2. Drag Load Assessment The constraints and modifications specified for the quencher a. bubble drag loads apply., Unless the lowest structural natural frequency times the durhtion b. 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 the " spikes" or the load shall be applied in the time domain using the source time The tenn " spike" refers to the short-duration history. high overpressure peak, such as that exhibited in Figure 6.2.1-20 of NEDE-24539-P. 2.15 SECONDARY LOADS The following loading conditions may be neglected for the PUA: a. seismic slosh pressure loads b. post-swell wave loads asyninetric pool swell pressure loan c. d. sonic and compression wave loads {} 2D0 downcomer air clearing loads e.

. 40 - 2.16 DIFFERENTIAL PRESSURE CONTROL REQUIREMENTS Those licensees that use differential pressure control (AP) as a pool swell lo.d mitigation feature for the LTP, shall demonstrate conformance with th.: 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 comensurate 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 confonn 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: a. 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 b and c below. b. The differential pressure shall be established within 24 hours after placing the plant in the RUN mode, during plant startup. The differential pressure may be reduced below "X" psid 24 hours prior to a scheduled plant shutdown. c. The differential pressure may be reo.:ced to less than "X" psid 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). d. In the event that the specification in a 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 21 hours. 1431 261

. A minimum of two narrow range instrument channels shall be provided e. to monitor the differential pressure. Error in the AP measurement shall be no greater than + 0.1 psid or the allowable AP shall be increased to offset the error in the measurement. The instrument channels shall be calibrated once every six months. In the event l that the measurement is reduced to one indication, operation is pemissible 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. i 3. STRUCTURAL ANALYSIS AND ACCEPTANCE CRITERIA 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), NED0 24583, Revision 1, dated July 1979, acceptabla. The proposed criteria will provide a sufficient basis for dar.onstrating the margins of safety required for steel structures and piping in the ASME 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 fomally submitted to complete the documentation required for this program. w e

, 4. REFERENCES (Listed by Report Number) NEDE-l'426P " Mark III Confinnatory Test Program One-Third Scale Pool Swell Impact Tests Test Series 5805," General Electric Proprietary Report, August 1975. NEDO 21471 " Analytical Model for Estimating Drag Forces on Rigid Submerged Structures due to LOCA and Safety Relief Valve Ramshead Air Discharges," General Electric Topical Report, September 1977. NEDE-21472 " Analytical Model for Liquid Jet Properties for Predicting Forces on Rigid Submerged Structures," General Electric Proprietary Report, September 1977. NEDE-21864-P " Mark I Containment Program Final Report Monticello T-Quencher Test, Task Number 5.1.2," General Electric Proprietary Report, July 1978. NEDE-21878-P " Mark I Containment Program Analytical Model for Computing Air Bubble and Boundary Pressures Resulting from an SRV Discharge Through a T-Quencher Device Task Number 7.1.1.2," General Electric Proprietary Report, January 1979. NEDO 21888 " Mark I Containment Program Load Definition Report," Revision 0, General Electric Topical Renort, December 1978. NEDE-21944-P " Mark I Containment Program Quarter Scale Plant Unique Tests, Task Number 5.5.3, Series 2," Volumes 1 - 4, General Electric Proprietary Report, April 1979. NEDE-21983-P " Mark I Containmedt Program Submerged Structures Model Main Vent Air Discharges Evaluation Report, Tass Number i 5.14.1," General Electric Proprietary Report, March 1979. NEDE-23545-P " Mark I Containment Program 1/4 Scale Pressure Suppression Pool Swell Test Program: LDR Load Tests - Generic Sensitivity, Task Number 5.5.3, Series 1," General Electric Proprietary Report, December 1978. NEDE-23749-1-P " Mark I Containment Program Comparison of Analytical Model for Computing S/RVDL Transient Pressures and Forces to Monticello Data," General Electric Proprietary Report Addendum, February 1979. NEDE-23817-P " Mark I Containment Program 1/4 Scale Test Report Loads on Submerged Structures due to LOCA Air Bubbles and Water Jets, Task Number 5.14," General Electric Proprietary Report, September 1918. l 1431 263

. NEDE-24537-P " Mark I Containment Long Tenn Program - Development of Downcomer Lateral Loads from Full Scale Test Facility Data - Task Number 7.3.2," General Electric Proprietary Report, May 1979. NEDE-24539-P " Mark I Containinent Program Full Scale Test Program Final Report, Task Number 5.11," General Electric ' Proprietary Report, April 1979. NEDO 24583 " Mark I Containment Program Structural Acceptance Criteria Plant Unique Analysis Applications Guide, Task Number 3.1.3," General Electric Topical Report. December 1978. NED0 24612 " Mark I Containment Program Vent Header Deflector Load Definition, Task Number 7.3.3," General Electric Topical Report April 1979. NEDE-24645-P " Mark I Containment Program Analysis of Full Scale Test Facility for Condensation Oscillation Loading," General Electric Proprietary Report, May 1979. NEDO 25070 " Analytical Model for Estimating Drag Forces on Rigid Submerged Structures Caused by Condensation Oscillation and Chugging Mark I Containments," General Electric Topical Report, April 1979. NEDE-25090-P " Analytical Model for T-Quencher Water Jet Loads on Submerged Structures, Task Number 5.14.2 " Genercl Electric Proprietary Report, May 1979.

, ~. [ Issued for Information] 7590-01 NUCLEAR REGULA. TORY COM?C5SION [Docke'. Nos. : 50-219, 50-220, 50-237, 50-245, 50-249, 50-254, 50-259, 50-260, 50-263, 50-265, 50-271, 50-277, 50-278, 50-293, 50-296. 50-29C. 50-321, 50-324, 50-325, 50-331, 50-333, 50-341, 50-354, 50.MS, 50-366] NOTICE OF ISSUANCE OF CRITERIA FOR THE PARK I CONTAINMENT LONG TERM PROGRAM The Nuclear Regulatory r.ommission (NRC) staff has completed its review of the proposed suppression pool hydrodynamic load assessment criteria submitted by the General Electric Company on behalf of the licensees of Boiling Water Reac-tor (BWR) facilities with the Mark I presstre-suppression containment design. As a result of this review, the staff has developed requirements by which the proposed criteria are to be implemented, and has issued these requirements to the affected utilities. This action represents a key milestone in the NRC's Generic Technical Activity A-7 (Mark I Containment Long Term Program), which has been designated an " Unresolved Safety Issue" pursJant to Section 210 of the Energy Reorganization Act of 1974. This generic study originated from concerns raised during the course of the NRC staff's review of test results for an advanced pressure-suppression contain-ment design (Mark III) and operating experience related to the pressure-suppression concept. The specific concern was the capability of the containment structure (i.e., suppression chamber) to withstand the effects of suppression pool hydrodynamic loads which had not been explicitly considered in the original design of the structure. A short-term assessment was conducted, from which the NRC staff concluded that continued o)eration of the licensed BWR/ Mark I facilities would not jeopardize the public lealth and safety. However, the containment design safety margins were found to be less than that required by the Commission's regulations in General Design Criterion 50 of Appendix A to 10 CFR Part 50. Accordingly, exemp-tions from this requirement were issued to the licensed BWR/ Mark I facilities in March and April 1978, which pennitted continued plant operation until a more comprehensive long-term program could be completed. The bases for these exemp-tions are described in NUREG-0408, " Mark I Containment Short Term Program Safety Evaluation Report," dated December 1977. The purpose of the long-term program is to restore the original intended containment design safety margins. The NRC staff's requirements for the long-term program define the generic metho-dology that is to be used by the affected utilities to design any additional stractural and/or operational modifications which will restore the margins of sa nty, as determined by a plant-specific analysis. These criteria were issued to the affected utilities in letters dated October 31, 1979, to commence the }k3$ b

. 7590-01 implementation of this program. Copies of the criteria are available for public inspection at the NRC Public Document Room,1717 H Street, N.W., Washington, D.C. and the Comission's Local Public Document Room located in the vicinity of the affected nuclear power plants. In addition, these criteria will be appended to the NRC staff's generic Safety Evaluation Re-port for the Mark I Containment Long Term Program, which is currently scheduled to be issued in December 1979. For further information, contact Mr. Christopher Grimes, A-7 Task Managar, Division of Operating Reactors, U. S. Nuclear Regulatory Comission, Washington, D.C. 20555. Dated at Bethesda, Maryland, this 31st day of Cctober,1979. OR THE NUCLEAR REGULATORY COMMISSION kf W 1sen in Director Division of Operating Reactors Office of Nuclear Reactor Regulation 1431 266}}