ML19281C930

NRC Acceptance Criteria for Mark I Containment Long-Term Program

ML19281C930
Person / Time
Issue date:	02/29/1980
From:	Office of Nuclear Reactor Regulation
To:
Shared Package
ML19281C929	List: ML19281C928 ML19281C930
References
FOIA-81-82, REF-GTECI-A-07, REF-GTECI-CO, TASK-A-07, TASK-A-7, TASK-OR SECY-80-359, NUDOCS 8009090097
Download: ML19281C930 (57)

Text

1 NRC ACCEPTANCE CRITERIA FOR TiiE MARK I CONTAINMENT LONG TERM PROGRAM Revision 1 February 1980 80 000 900q 7

TABLE OF CONTENTS PAGE 1.

-INTRODUCTION.....................................................

1 2.

SUPPRESSION POOL HYDRODYNAMIC LOA 0S..............................

2 2.1 Containment Pressure and Temperature..

2 2.2 Vent System Pressurization and Thrust Loads........

2 2.3 Net Torus Vertical Pressure Loads...........................

3 2.4 Torus Pool Swell Shell Pressures............................

4 2.5 Compressible Flow Effects in Scaled Pool Swel1 Tests........

5

2. 6 Vent System Impact and Drag Loads......

6 2.7 Pool Swell Impact and Drag on Other Internal Structures.....

9 2.8 Froth Impingement and Fallback Loads........

19 2.9 Pool Fallback Loads..........................

22 2.10 Vent Header Deflector Loads...........................

23 2.11 Condensation Oscillation Loads..............................

30 2.11.1 Torus Shell Loads...................................

30 2.11.2 Downcomer Loads.....................................

31 2.11.3 Vent System Pressures...............................

32 2.12 Chugging Loads..........................................

32 2.12.1 Torus Shell Loads...................................

32 2.12.2 Downcomer Loads.......

33 2.12.3 Vent System Pressures.............................

34 2.13 Safety-Relief Valve Discharge Loads.........................

34 2.13.1 SRV Ofscharge Device..............................

34 2.13.2 Discharge Line Clearing Transient..............

34 2.13.3 Air-Clearing Quencher Discharge Shell Pressures....

34 2.13.4 SRV Discharge Line Reflood Transient.....

37 2.13.5 SRV Air and '.later Clearing Thrust Loads.........

37 2.13.6 SRV Discharge Line Temperature Transient...

37 2.13.7 SRV Discharge Event Cases........................

37 2.13.8 Suppression Pool Temper **

'e Limits.....

38 2.13.9 SRV Load Assessment by In, tant Tests......

41 2.14 Submerged Structure Drag Loads.

43 2.14.1 LOCA Water Jet Loads............

43 2.14.2 LOCA Bubble Drag Loads...

45 2.14.3 Quencher Water Jet Loads.

2.14.4 Quencher Bubble Drag Loads.......

48 48 2.14.5 LOCA Condensation Oscillation Drag Loads...

49 2.14.6 LOCA Chugging Drag Loads...

50 i

TABLE OF CONTENT 5 (Continued)

PAGE 2.'15 Secondary Loads.............................................

51

2. l's Di f ferential Pres sure Control Requi rements.................

51 3.

STRUCTURAL ANALYSES AND ACCEPTANCE CRITERIA.....................

53 4.

REFERENCES.................................................

53 t

ii

1 1.

INTRODUCTION The purpose of the Mark I Containment Long Term Program is to perform a complete reassessment of the suppression chamber (torus) 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 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 been developed from the staff's review of the Long Term Program Load Definition Report (LDR), the Plant Unique Analysis Applications Guide (PUAAG), and the supporting analytical and experi-mental programs conducted by the Mark I Lwners 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 accordar.ce with the plant's approved Final Safety Analysis Report (FSAR).

Similarly, references to original design criteria or original loading cond'tions 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," NEDO-21888, PUAAG refers to " Mark I Containment Program Structural Acceptance Criteria Plant Unique Analysis Applications Guide,"

NE00-24583, 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 2.

SUPPRESSION POOL HYORODYNAMIC LOADS 2.1 CONTAINMENT PRESSURE AND TEMPERATURE RESPONSE The pressure and temperature transients for the drywell and wetwell shall be determined 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 1

comparison to the staff's CONTEMPT-LT computer code.

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 part of the PUA, each licensee shall 1

specify procedures (including the primary system parameters monitored) by which the operator will identify the SBA, to assure manual operation of the AOS within the specified time period.

Longer time periods may be assumed for the S8A in any specific PUA, provided (1) the chugging load duration is corre-spondingly increased, (2) the procedures to assure manual operation within the assumed time period are specified, and (3) the potential for thermal stratifi-1 cation and asymmetry 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 bubole breakthrough for those plants tnat propose operation with a differential pressure control, the vent clearing time shall be derived from a containment analysis assuming no drywell/wetwell differential pressur.e and this time shall be applied to the vent system transients calculated from a

3 containment response witn the proposed drywell/wetwell differential pressure.

In addition, for clarfiication, 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 vertical pressure loads on the torus shall be derived from the series of plant-specific QSTF (Quarter Scale Test Facility) tests, in accordance with Section 4.31 of the LOR.

However, based on our review of the pool swell tests conducted by ths Mark I Owners Group and con-firmatory tests performed for the NRC by the Lawrence Livermore Laboratory, we will require that the following margins be applied to each loading phase:

= UP

+ 0.215 (UPmean) ean DOWN = DCWN

+ 2 x 10-5 (DOWNmean) mean where "mean" refers to the average of the QSTF plant-specific test results (lb ).

These margins shall be applied to the QSTF "mean" load function prior f

to scaling the load functica to full-scale equivalent conditions.

The margin for the downward loading function shall be derived in terms of a fraction of the load at the time of the peak downward load, and th fraction shall be applied to the entire downward loading phase.

The margins specified above may be reduced or omitted wnere minimum conservatisms (i.e., smallest parameter deviation from the nominal plaat 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 changed to the extent that the QSTF test series no longer represents a conservative configuration of the plant, then a new series of QSTF tests shall be performed.

4 For those plants that use drywell/wetwell differential pressure control as a load mitigation feature, an additional structural 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 single plant-specific QSTF test run may be tsed 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 longitudinal 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.

6 2.6 VENT SYSTEM IMPACT ANb ORAG LOADS 2 6.1 Vent Header Imoact and Orac Loads The load definition procedures set forth in Section 4.3.3 of the LDR are acceptable, subject to the following clarifications:

1.

The experimental data of local vent header p essure 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 water impact on the vent header shall be determined 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 Imoact and Oraq Loads The load definition procedures set forth in Section 4.3.3 of the LOR are acceptable, subject to the following clarifications.

A pressure of 8 psid is to be applied uniformly over the bottom 50* of the angled portion of +he downcomer, starting from the time at which the rising pool reaches the lower end of the angled section and ending at the time of maximum pool swell height.

The pressure is to be applied perpendicular to the local downcomer surface.

The structural analysis for the downcomer 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.

7 2 6.3 Main Vent Imoact and Drag Loads The impact and drag loads on the.1ain vent shall be evaluated in the following manner:

l.

Subdivide the submerged portion of the main vent pipe into six equally wide segments (see Figure 2.6-1).

If this subdivision resul h in al <

0.30 fewer segments may be used such that al ~ 0.30.

2.

Determine the velocity and acceleration histories at Points 1 through 7 in Figure 2.6-1 from the QSTF data and appropriate corrections for longitu-dinal variations along the torus (at Point 7, only the initial impact velocity is required).

3.

Using the velocity components normal to the vent pipe, calculate the impact and " steady" drag pressure using the method in Sect!on 2.7.1 (Cylindrical Structures).

At Point 7, only impact force is to be cons"dered.

Using the acceleration components normal to the vent pipe, calculate the accelerat:an drag pressure using the equation F

o n0 static oucyancv a

2g 144 V*

144 DL 1

Where P, = the acceleration pressure averaged ove-the projected area (ps1),

3 p = the density of water (lbm/ft ),

0 = the diameter in feet, 2

V = the cross flow acceleration (ft/sec ),

g = gradational constant (R - kmN - sec ), and 2

c L = submerged length of main vent

8 O

\\

SL I

2 e

4 y Highest pool elevation 7

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 Illustracing the Methodology for Main Vent Impact and Drag

9 4.

Sum the pressures due to impact, viscous drag and acceleration drag and multiply by 0 to obtain force per unit length at Stations 1 through 7.

- 5.

To obtain a -nooth loading history for the main vent as a whole, the linear interpolath

'-d suggested for the vent header deflector in Section 3.5 of NEDO-24e1% may be used.

2. 7 POOL SWELL IMPACT AND ORAG ON OTHER INTERNAL STRUCTURES The impact and drag loads for internal structures Pove the suppression 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 struc-tures are classified as either cylindrical (e.g., pipes), exposed flat surfaces (e.g.,

"I" beams), or gratings.

The following load specifications for each of the three structural classifications shall be used to replace the method-ology in the LDR.

Non-cylindrical structures can be conser/atively defined as equivalent flat-surfaced structures.

However, if such an approach is too conservative, a similar technique may be used with an impact data base which is appropriate for the structural configuration of interest.

The icngitudinal velocity distribution shall be based on the " main vent" EPRI pool swell tests, as discussed in Section 2.5.

2.7.1 Cvlindrical 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 determined by gx 2

P ** = 7. 0 x 1 ( la 1 o V 2\\

gc a

10 P

max e

L 2m WQ L

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r Time Figure 2.7-1 Pulse Shape for Water Impact on Cylindrical Targets 2 *.

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11 where P

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max 3

p = the density of water (lba/ft ),

V = the impact velocity (ft/sec), and 2

g = gravitational constant (ft - lbm/lbf - sec ).

c 2.

The hydrodynamic mass per unit area for impact loading shall be obtained from the 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 determined by:

I

• p 1

gc where I = the impulse per unit area (psi-sec),

p 2

M /A = the hydrodynamic mass per unit area (lbm/ft ), and H

V = the impact velocity (ft/sec).

4.

The pulse duration will be determined from the following equation:

I = 2I /P, p

5.

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

2 C

PY D

max 0"2-144 gC-where P = the average drag pressure acting on the projected area d O

the target (psi),

CD = the drag coefficient as defined by Figure 2.7-2, 3

p = the density of water (Ibm /ft ), and V

= the maximum vertical velocity attained by the pool (ft/sec).

12 (Note that different velocities are used for the determination of impact and drag loads.)

2.7.2 - Flat-Surface Structures For flat-surface structures, the pressure transient which occurs upon water impact and subsequent drag is depicted in Figure 2.7-3.

The carameters in Figure 2.7-3 shall be defined as follows:

1.

The pulse duration (t) is specified as a function of the impact velocity:

I = 0.0016W for V < 7 ft/sec I = 0.011 W for V > 7 ft/sec V

where W = the width of the flat surface (feet) and V = the impact velocity (ft/sec).

2.

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

2 PD=C9 pV,,,

T 144 gc-where PD = t e average drag pressure acting on the frontal area of the structure (psi),

CD = the drag coefficient (CD = 2, flat strips ncrmal to flow, independent of Reynolds number),

3 p = the density of water (lbm/ft ), and V

= the maximum vertical velocity attained by the pool (ft/sec).

max 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.

13 4

The impulse of impact per unit area shall be determined by:

I = "H V

l 144 g where Ip = the impulse per unit area (psi-sec),

2 M /A = the hydrodynamic mass per unit area (lbm/ft ), and H

V = the impact velocity (ft/sec).

5.

The maximum pressure (P

) shall be calculated from the impulse per unit area and the drag pressure as follows:

P

=2 p

.p 0

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 differential (cigure 2.7-4) and the total grating area (not only the area of the metal bars).

The pressure differ-ential 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:

0=@xA grating Omax

\\ 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 + (0.0064 Wf)2] ; for Wf < 2000 in/sec.

SE

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0.5 0.6 0.7 0.8 0.9 1.0 OPEN ARE A FRACTION Figure 2.7-4 Pressure Drop Due to Flcw Across Genings

15 where FSE = static equivalent load W = width of grating bars, in.

f a natural frequency of lowest mode, Hz and 0 = static drag load If Wf > 2000 in/sec (not expected for gretings) the force nn the bars of the gratings will be calculated by the method outlined above for flat-surfaced structures.

2.7.4 Load Acolication These load specifications correspond to impact on " rigid" structure:.

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.

In performing the sti ctural dyramic analysis, the drag following impact 1

(as shown in Figures 2.7-1 and 2.7-3) shall be included in the forcing function.

The transient calculation shall be continued until the maximum stress in the structure has been identified.

When the impact loading is primarily impulsive and calculations have already been performed in accordance with tne LOR methodology, (using the 1

impulse equation on page 4.3.4-5 of the LDR with K = 0.2 for cylinders and K h

H

= 0.62 for flat structures) simple adjustments may be made to the LDR analyses.

Under these conditions, a parabolic pulse shape, as proposed in LDR, is accept-able provided corrections are made to account for the 35% margin in the impulse and with additional corrections for the drag force immediately following impact.

h

16 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.

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 stresses will then be multiplied by an additional factor to account for the presence of drag following the impact.

This factor is determined as follows:

a.

Calculate the drag pressure, (Pdrag) as described in Sections 2.7.1 or 2.7.2.

b.

Form the ratic:

drag! max where P,y is the amplitude of the parabolic pulse used in the original stress analysis multiplied by 1.35.

c.

Determine the dynamic load factor (OLF) from Figure 2.7-6, corresponding to the two cases:

(1) parabolic pulse without drag and (2) parabolic pulse followed by drag.

d.

Multiply the calculated stress by the factor OLFwith drag /D Uw/o drag

17 TYFICAL IMPACT STRUCTI.7.ES:

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--r..

__. : n_ r.==== = -

- - - - -- :_=._. _.. _ =.

O.

0.1 0.2 03 0.4 03 x / T,d Figure 2.7-6 Effect of Drag Following a Parabolic !moact Pulse

19 2.8 F_ROTH IMPINGEMENT AND FALLBACK LOA 05 Froth is generated by (1) impact of the rising pool surface on the vent header"and (E) 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

Pf Pf= 144 9c where Pf = froth impingement pressure (psi) pf = froth density (lb,/fta)

V = froth impingement velocity (ft/sec) g = gravitational constant (ft - lbm/lbf - sec )

2 c

Region I:

The froth velocity shall be based on a source velocity equal to 2,5 titres 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 source vector shall be assumed to be on a line between the 45 tangent on the header and the target structure.

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) pw, where x is the dimension in feet.

The load shall be applied in the direction most critical to the structure within the 90 secte: bounded by the hori: ental 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.

I y

20 1

FROTH AEGION I VENT 4

MEACER TYPfCAL STRUCTURE

• d i

TCRUS Figure 2.8-1 Direction of Load Application for Froth Region I

22 lead to a non-conservative estimate of the froth density.

Consequently, a 1

separate, lower-bound froth source velocity must be used to determine the froth density.

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 eleva-tion of the maximum velocity.

The i.oth density shall be assumed to be 100%

water density for structures or secticns 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 directly above the vent header.

The load shall be apolied in the direction most critical to the structure within the 45 sectar of t;1e 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 from 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 1

swell are acceptable.

The drag load for pool fallback shall be assessed in a manner consistent with the LOCA bubble submerged structure drag loads (Cri-terion 2.14.2).

In additier., structures which may be enveloped by the LOCA pool swell bubbles shall be investigated to determine if bubble collapse causes a higher stress than the submerged structure drag loads.

23 2.10 VENT HEADER DEFLECTOR LOADS The load definition procedures set forth in Section 4.3.9 of the LOR are applicable only to the four defTector types shown in Figure 4.3.9-2 of the LDR, and are generally acceptable, subject to the following constraints and/or modifications:

2.10.1 QSTF Deflector Load An individual plant may choose to use deflector load data taken directly frcm the QSTF prhnt-unique tests.

This technique is subject to the following 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) impact time delays and (b) pool swell velocity and acceleration differences which result from uneven spacing of downcomer pairs.

The longitu-dinal load variation shall be evaluated at the instant when the undisturbed 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 EPP.I " 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 ceflector shall be accounted for.

The added mass per unit length of deflector may be estimated by:

M =Ig 1

g c

V

24

.~

J F = Vertical upward force on deflector per unit length 8 ' l d = Diameter of cylinder in deflector types 1 - 3

~

V = Impact velocity

'g o = Water density 6.

'y t = Time from begining of impact F

2

\\

pVd

+

2 N~;r.

r

= q = __ _3 ; - _== p =.__ --- - - --. -

w n

V 4*. 2: =_ ;-- - - - -\\q g t.----

t=--

-I v.

4

.g...=~..-

.r__.

N x __

x LN 2.

x Nx,

-- = _ _ _ _.. _ _. _ -

x I

m O.

\\

0.

0.1 0.2 Vt / d Figure 2.10-1 Impact Force Transient for Addition to the Empirical Data for Deflector Types 1 - 3.

25 where M = hydrodynamic mass per unit length (lbm/ft),

H I = total impulse per unit length associated with the impact transient (lbf-sec/ft),

~

V = impact velocity (ft/sec),

2 g a gravitational constant (ft - lbm/lbf - sec )

c 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 LOR, is acceptable, with the following modifications:

1.

The impact transient and " steady drag" contributions to the load shall be computea from the correlations shown on Figures 2.10-2 through 2.10-5, for the deflector types 1-4, respectively.

For times past the periods shown, the last value shall be extended for the duration of the transient.

Slight variations in the wedge angle or ratio w/d should be corrected using Wagner's correlation (National Advisory Committee for Aeronautics Report #1366) for wedges and the final steady drag value.

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 crag shall be included in F, as defined in NEDO-24612.

4 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.

26

<d>

F A

2 Vd l

o z.

V (C )=

rom gure 2.10-2b D

V

\\

\\

\\

0.

0.'136 A

1.0 l

Vt/d Figure 2.10-2 Impact and Steady Drag Force Correlation for Type 1 Deflector 2.

.i

,ii,,

i,,,

.i.

i i

, i,

,6 6.ie6.

,,i

,,.1 i,

i e:

.i, e

i,

,ii i,,,,,,,

,,,,n, i.,

,,i n.

n g

i

, i

,1 I

I ili.

,, e l..

,,i

,t 6

i i

i1.

.6.

, i,.,,

in o,i,,,,

m i,

,,,i,,,

o

...a.

w,

5 1.5 Re < 5 x 10 2

,.'.' x -

> i D

. x..

~

.. N,,.

, )

.,i, i

.f,

.N,,

2.,,

i i,,N..

,,o

.,i C

,e g

i

..i..

,,.,....x,

,..i.

1 x.

y.

o,

i,

q.

x :,,,

u s,i.

.i,

,x e.

L,,

, i i,

,, i i

i.

.i,,

,i

. i

,, li

],

. o so -

,6.

..i..

..a

.,x i

,.n.,

,.i,

a, x,=, x.,.

n

. e..

,i,

....x.,

f

.f

. x

, i,

i

/,

i..

N 1,,

i,,

i i.j

.,,6 x

,i, i

.,i.

/

,e,

',N,

o*S Re 3 5 x 10 A

5 i

T' -

D

,i...

, i

,,,,n, i,c m,

r r i

.i.,,

i.

.i..,

6-

.l.

e i

1, I.

e e

l.

I.

i 1

i i

i.

.i.

O.

10.

I n.n.

2 Froude Number [V /Dg]

igure 2.10-2b (C )= f r Type 1 Deflector D

27 4 w

~

A d

Y A

l V

h-1 F

Vw i

0 2

i\\

/

5

/

\\

/

\\

/

\\

\\

/

/

\\

\\

/

\\

\\

/

\\.

g

.207 0.

O.k36 05 0.854 1.0 Vt/d Figure 2.10-3 Impact and Steady Drag Force Correlation for Type 2 Deflector

28 4

W

'E d

V 1

V w

1 W

F VW 2

0s7

/

'si

\\

\\

/

g

\\

/

g

,/

\\

1 g

\\.

.'207 0,

0,136 Os83 1(33 2[0 Vt/d Figure 2.10 4 Impact and Steady Drag Force Correlation for Type 3 Deflector

29 4

w >

A V

7 F

VW o 3-

\\

\\

\\

\\

i 9

j.0 l.5 d.0 O.

2Vt/w Figure 2.10-5 Impact and Steady Drag Force Correlation for Type 2 Deflector

30 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 conden-sation oscillation regime is a harmonic phenomena and, therefore, statistical variance or load magnitude uncertainty cannot be established from one test run.

Althougti we conclude that the M8 tested conditions are conservative and prototypical for the Mark I design, a reasonable measure of the uncertainty in the loadirp function is necessary to assure the margins of safety in the containment structure.

However, based on o r assessment of the phenomenological studies conducted by the industry and the NR? Office of Nuclear Regulatory Research, we believe that the following load specifications are probably conservative and form a sufficient basis to proceed with implementation of the Mark I Long Term Program. We will require that the Mark I Owners Group confirm the condensation oscillation loads (i.e., torus shell loads, downcomer lateral loads, vent system pressure, and submerged drag source) by performing a sufficient number of additional large break, liquid bicadown tests in FSTF to establish the uncertainty in the load magnitudes.

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, provided the " rigid wall" load derivation described in NEDE-24645-P and the condensation oscillation coherence (basis for excluding asymmetric loading condition) are confirmed by the additional FSTF tests.

1 For clarification, the 1 cad 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 sinusodial function with a peak amplitude and frecuency range of that speci-fied for the " pre-chug" load.

31 2.11.2 Condensation Oscillation Downcomer Loads 2.11.2.1. Untied Downcomer Loads The condensation oscillation downcomer lateral loads for " untied" down-comers 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 P

=P y where P,y = maximum static equivalent lateral load for plant-unique downcomer, Pt = maximum static equivalent lateral load in FSTF, OLF = plant unique downcomer dynamic load factor, and OLF1 = FSTF downcomer dynamic load factor.

the plant-unique loading condition shall be derived as follows.

The plant-unique OLF 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 4-8 Hz which produces the maximum structural response.

The natural frequency and damping values for the FSTF OLF shall be conservatively established from a " pluck" test of an untied downcemer in FSTF, with a nominal water level of 3 feet 4 inches and an ampli-tude in the range of the response level.

The driving frequency for the FSTF DLF shall be assumed to be 5.5 Hz.

32 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 an each downcomer in a tied pair.

The load definition and structural analysis technique shall be confirmed 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 downcemer 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 procedures set forth in Section 4.4.4 of the LOR for the oscillatory pressures in the vent system during the condensation oscilla-tion period, are acceptable subject to confirmation by the additional testing as described above.

2.12 CHUGGING LOADS 2.12.1 Chuacino Torus Shell Loads The load definition and assessment procedure set forth in Section 4.5.1 of the LOR for the chugging condensation loads on the torus shell are acceptable.

1 For clarification, the load specification for " post-chug" loads set forth in Section 4.5.1 of the LDR shall be used in conjunction with a coupled fluid-structure analytical model.

33 2.12.2 Chuccing Downcomer Loads 2.12.2.1. Untied Downcomer Loads The chugging laterai 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 measured 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 an exceedance probability of 10 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 downcomer data.

The strain in the tie bar shall be evaluated by assuming that ene of the two tied downcomers is subjected to a dynamic load of triangular shape, with an amplitude of:

F

= RSEL (RSEL = Resultant Static Equivalent Load) max n ftd 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 d

milliseconds.

The load direction shall be taken as that (in the horizontal

34 plane) which result in the worst loading condition for the tie bar and its attachments to the downcomers.

2.12.3 Chugging Vent System Pressure Loads The load definition procedure set forth in Section 4.5.4 of the LOR for the oscillatory pressures on the vent system during the chugging period are acceptable.

2.13 SAFETY-RELIEF VALVE DISCHARGE LOADS 2.13.1 Safety-Relief Valve Discharce Device The acceptance criteria set forth below for analytically derived quencher discharge loads are applicable only to the "T" quer.cher configuration described 1

in Section 1.1 of NEDE-21864-P.

The SRV discharge load assessment procedure for otner quencher configurations (e.g., "Y" quencher), or as an alternative method for the T quencher, using in plant test data is described in Section 2.13.9.

2.13.2 SRV Discharge Line Clearing Transient The load definition and assessment procedure, described in Section 5.2.1 of the LOR, for the pressure and thrust loads on the SRV discharge line and quencher, which is based on the methodology presented in NEDE-21864-P and NEDE-23749-1-P, is acceptable.

2.13.3 SRV Air-Clearing Quencher Discharce Shell Pressure I. cads 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 pressure are accectable, with the following exceptions:

35 1.

The load definition procedures set forth in Section 5.2.2 of the LOR 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 line exceeds 13.5 feet, the load prediction for a 13.5 foot water-leg shall be used.

For discharge line volumes greater than 65 cubic feet, no additional 1

pressure amplitude increase due to the line volume trend is necessary.

2.

The 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 subsequent actuations.

2.13.3.2 Methodology for Torus Shell 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 calcula.ed by the load definition procedures described in Section 5.2.2.3 of the LOR in conjunction with the appropriate pressure atten-uation model.

For off-center T quenchers, the pressure attenuation model described in the letter to D. G. Eisenhut, NRC, from L. J. Soban, GE,

Subject:

Mark I Containment Program, Additional Information on NEDO-21888, dated September 7,1979, is acceptable for the quencher location defined therein.

For cuenchers 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.

This factor is 1.65 for local shell pressures.

However, for the determination of glebal pressure loads en the torus (for the torus supports evaluation), this multiplier may be reduced to the value required to bound the g1cbal pressure loads on the torus from the Monticello in plant tests (NEDE-21864-P).

36 2.13.3.3 Multiple Valve 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 with 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 be th first and subsequent actuations.

2.

In the event that the combined peak torus shell pressure exceeds the local predicted peak bubble pressure at the bottom center of the torus by a f' actor greater than that defined in Section 2.13.3.2, 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 LOR, 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:

1.

First Actuation - the frequency range shall be 0.75 times the minimum predicted frequency to 1.25 times the maximum predicted frequency.

2.

56 sequent Actuation - the frequency range shall be 0.60 times tne minimum predicted frequency to 1.40 times the maximum predicted frequency.

37 2.13.4 SRV Discharge Line Reflood Transient The transient analysis technique to compute the plant-specific reflood heights in the SRV discharge line following va!ve 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 Discharoe Line Temoerature 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

'he kind and number of SRV discharge events shall be based on the plant-specific system configuration and a conservative assessment of plant operational hi story.

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 determined from a plant-specific primary system analysis.

2.

Asymmetric SRV discharge, both first and subsequent actuations, shall be considered for SRV, SBA, and IBA event comoinations.

The degree of

38 asymmetric discharge for each event combination shall be determined fron; a plant-specific primary system analysis designed to maximize the asymmetric condition.

3.

ADS valves discharging on first actuations shall be considered for the SBA and IBA event combinations, followed by subsequent actuations deter-mined 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 point pressures.

All first and subsequent actuations shall be assumed to occur in phase.

The number of subsequent actuations shall be determined from the primary system analysis.

The SRV discharge event cases described abcve shall be used in the event 1

combinations described in the PUAAG.

The plant unique primary sy: tem analysis shall be performed so as to provide a conservat've estimate of the number of SRV actuations for both first and subsequent actuation events (e.g., ODYN code modified to account for sensible heat and pressure uncertainties).

2.13.8 Sucoression Pool Temoerature 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 (1) the hole diameter equal to, and (2) greater than or equal hole spacing than :nat of the generic Mark I T-Quencher.

39 2.

Local and Bulk Pool Temperature The lccal 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 poci temperature difference shall be supported by test data from either the existing Monticello pool temperature data or in plant tests.

Where in plant tests are used to establish the bulk to local pool temperature difference, the pool shall be at ambient (i.e., still) conditions prior to opening the SRV and the SRV discharge line selected for testing shall be located away from the RHR discharge nozzle.

The duration of the SRV dis-charge shall be at least ten minutes.

RHR flow shall not be initiated sooner than five minutes after the SRV is opened.

Temperature monitors shall be located on the reactor side of the torus, downstream with respect to RHR flow, at the same elevation as the quencher, and on the quencher support.

The bulk to local pool tamperature transient derived from this' test may be used directly to determine the local pool temperature transient.

In order to take maximum credit for the effectiveness of the RHR system to mix the pool, an additional in plant test may be performed where the RHR system is started at the same time the SRV is actuated.

The bulk to local pool temperature difference from the previous test shall be assumed up to the time the RHR system started and then linearly decrease in time to the minimum temperature difference from the second test.

The " local" temperature is defined as the temperature in the vicinity of the quencher device during discharge.

For practical purposes, the average water temperature 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 and at the quencher suoport may be considered as the " local" temperature.

The " bulk" temperature, on the other hand, is the temperature calculated assuming a uniform distribu-tion of the mass and energy discharged from tne SRV.

40 3.

Suppression Pool 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 Tecnnical Specifications.

The system shall meet the following design requirements:

a. ' Each licensee shall demonstrate that there is a sufficient number snd distribution of pool temperature sensors to provide a reasonable measure of the bulk tamperature.

Alternatively, redundant pool temperature 1

monitors may be located at each quencher, either on the quencher support or on the torus shell, tu provice a measure of local pool temperature for each quencher device.

In such cases, the Technical Speu..".c.Wn limits for local pool temperature shall be derived from the calculated bulk pool temperature and the bulk to local pool temperature difference transient.

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 indicated and recorded in the control 1

Where the suppression pool temperature limits are based on bulk pool room.

temperature, operating procedures or analyzing equipment shall be used to minimize the actions required by the operator to determine 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., Quality e.

Group B, and energized from onsite emergency power supolies.

41 1

2.13. 9 SRV Load Assessment by In-Plant Tests

'T'he SRV load assessment procedure for plants equipped with quencher devices other than the Monticello T quencher shall be based on a series of at least four single-valve, " cold" discharge in plant tests, performed on the SRV discharge line expected to produce the highest loads.

This is also an accept-able alternative method for the Menticello-type T quencher.

The pressure wave form and structural response measuiaments of the in plant tests shall be used to calibrate a coupled load-structure analytical model.

The load prediction portion of the model shall be that described in NEDE-21878-P as modified by the criteria.in Sections 2.13.3 and 2.13.4.

The model shall be calibrated at test conditions and then applied at design basis conditions.

2.13.9.1 Instrumentation The following instrumentation shall be used for the SRV in plan't tests, as a minimum.

1.

Columns - Unfaxial gages shall be installed on both inner and outer support columns.

2.

Shell - Strain gages, acceleremeters, and pressure transducers shall be installed on the torus shell.

There shall be a sufficient number of gages to characterize both the significant shell vibration modes (6-12) Hz and the pressure waveform.

The significant shell vibration modes can be determined from the mode shapes calculated by the analytical model.

If necessary, the shell natural frequencies and damping may be deriw a from the test data.

3.

Attached Piping - Where preliminary structural analyses indicate that certain piping systems are expected to have a significant response, instrumenting these lines at the shell attachment with acceleremeters and strain gages should be considered, in arcer to calibrate the analytical

42 modelling of these structural elements.

In such cases, accelerometers or 1

displacement transducers should also be located at points where the maximum piping response is expected.

This approach need only be applied where a more realis' tic piping response analysis is necessary; otherwise. the accepted conservative analysis techniques shall be used.

2.13.9.2 Model Calibration The analytical modelling used to predict the plant - unique structural response shall be derived from tha in plant test cata with the folicwing basic considerations:

1.

A general modelling technique should initially be developed using measured pressure waveforms (i.e., actual pressure transients) from the Monticello tests.

Adjusting the model to match the measured response from the Monticello tests will identify the significant modelling parameters for the plan anique analyses.

2.

The frequency content of all of the measured pressure waveforms from the Monticello and in plant tests may ue used to determine the maximum amplifica-tion of the structural response.

Amplifcation for first actuation and subsequent actuation events should be considered separately.

The maximum amplification shall be applied to the structural natural frequencies which occur within the range of predicted frequencies, as defined by the criteria in Section 2.13.3.4.

3.

The analysis technique used for the plant-unique analyses shall be ver.ified by comparison to (1) the peak pressure, (2) the longitudinal and circumferential attenuation, and (3) the structural response of the in plant tests, using tested conditions as inputs.

Adjustments to the analytical modelling may be made for first actuation peak pressure and structural response, based on a conservative interpretation of the in plant test data.

Adjustments may not be made for subsequent actuati0n peak pressure predictions.

Attenuation characteristics for quencher devices other than the generic T quencher

43 shall be derived from tha in plant test data (longitudinal attenuation may i

either be defined by using pressure transducers distributed in a manner similar to that used in the Monticello tests or by installing uniaxial gages en the

~

three sets of columns adjacent to the test bay).

The response of each struc-tural element shall be determined from design-basis predictions of the corrected analysis technique and the 1:aximum structural amplification.

2.14 SUBMERGED STRUCTURE ORAG LOADS 2.14.1 LOCA Water det 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), is accept-able subject to the following constraints and/or modifications:

1.

The plant-specific jet discharge velocity, V ( ), and acceleradon, D

a()*d /dt (t), from the QSTF plant-specific tests series shall be used as D

D the driving sources for the jet model.

2.

Forces due to the pool acceleration and velecity induced by the advancing jet front shall be computed for structures that are within four downccmer diameters below the downcomer 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 hemispherical cap centered one d3wncomer diameter (D) behind the " Moody" jet front positions, containing the same amount of water as the " Moody" jet front.

The formu.as for the hemisphere radius (R )

g and the trajectory of the hemisphere centar (x ) are:

R (t) =

f + 3(x (t)/0) 1/3 for x (t) > 0 s

7 f

R (t) = 0

• f( ) I for x (t) 1 0 f

s 2

2D

44 x (t) = x (t) -0 for x (t) > 0 c

7 f

x (t) = 0 for x (t) < 0 c

7

.~

where x (t) is the position of the " Moody" jet front as a function of 7

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 t,his region can be approxi-mated by:

A (x,t) = n02 (1 + 1/(1-x/x )b) f 8

where x (t) is the " Moody" jet front position as computed in NEDE-21472-P.

7 The volume contained in this portion of the jet can be obtained by integrating A(x,t) from (x -0) to x for x7 greater than 0, and from x-0 to x = x7 7

for x 7

7 less than 3.

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 $ (x,y,t) 3 induced by the jet front:

dR, R

Q3 (x,y,t) = -

y s

(x - x )

C r

c dt 2

r at where r = {(x-x ) +y}

and y is the transverse distance of the structure c

from the jet axis, and (x-x ) is the distance from the structure to the effective jet front center along the jet axis.

The potential is the superposition of the expansion and motion of the sphere as given in any standard hydrooynamics

45 text (e.g., Milne iompson, Theoretical Hvdrodynamics, Fourth Edition, pp.

455-566).

.The local uniform flow velocity is V,(x,y,t) = V$)

as in NED0-21471, while the acceleration is a(x,y,t) = bum at This calculation need only be performed for r > R and x > x.

If either s

c of these conditions are not satisfied, the methodology in the LDR will bound the load and is, therefore, acceptable.

2.14.2 LDCA Bubble Drac Loads The load definition and assessment procedures described in Section 4.3.8 of the LDR, which are based on the methodology in NED0-21471 and experimental confirmation in NEDE-23817-P, are acceptable subject to the following con-straints and/or modifications:

1.

Flow Field a.

QSTF plant-specific test results (NEDE-21994-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 ficw field above the downcemer exit elevation will be derived from the QSTF plant-specific tests.

46 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 considering forces on an. equivalent cylinder b

of diameter D

=2 Lg, where L is the maximum transverse dimension.

eq ma L

is defined as the diameter of a circumscribed cylinder about the cross-max 2

section of the structure.

For example, L,y equals (a2+b) for a rectangular 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 Deq).

Alternatively lenger segments may be used as long as the equivalent uniform flow velocity and acceleration are evaluated conservatively for every point on any such seg'r:nt.

p.

Interference effects f.e to the proximity of walls shall be considered for each structural segment that has its center less than 1.5 diameters from a boundary.

Interference effects between neighboring struc-tures shall be considered whenever the centers of the segments are less than 30, where D = 1/2 (01 + D ), the average diameter of the two structures.

2 For structures near walls, the multiplier (1 + A ) shall be used to increase the acceleration drag and the multiplier (1 + 0,) shall be used t' increase the standard drag.

A and D, that bound theory and experiments are given below as functions x (x = r/D - 1/2, where r is the distance from the segment center to the boundary).

0.05 < x,< 1.0 A,= 0.05/x, 0,= 0.12/x,

47 x < 0.05 A = 1.0 0, = 2.4 For structures with neighbors that are less than 30 away and within 30 of being parallel, the multipliers (1 + A ) and (1 + 0 ) shall be applied 7

7 to the acceleration and standard drag components, respectively.

Bounding expressions for A and 0 are given below as functions of x7 (x7=r12/0 - 1 7

7 where c is the distance between segment centers).

l2 O

0.05 < x7<2 A7 = 0.2 2

07+O2 I

D7 = 0.2/xy where D is the diameter of the structure under consideration and 0 is the 2

diameter of the neighbor.

If more than one neighbor must be considered, the A and 0 values may be summed over the neighbor structures.

For x less than 7

7 7

0.05, the two neighbor structures shall be considered as an effective single structure.

The effects of wall proximity and neighbor structures may be super-imposed in order to compute overall multipliers as follows:

(1 + A, + I AIk) k (1 + 0, + I DIk) k In situations where interference effects must be considered, but the e

correction techniques outlined above are not applicable, a detailed interference effects analysis shall be performed.

48 2.14.3 Ouencher Water 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 confirmatory tests discussed in NEDE-25090-P.

2.14.4 Ouencher Bubble Orac Loads The load definition and assessment procedures described in Section 5.2.5 of the LDR, in NEDE-21878-P, and in NED0-21471-2, are acceptable subject to the following constraints and/or modifications:

1.

Flow Field a.

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.

b.

Drag loads on the quencher arms and the SRV discharge line shall be ccmputed on the basis of asymmetric bubble dynamics.

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-21864-P.

c.

Model E in NEDE-21983-P shall be used for the method of images representation of the torus cross-section.

~

49 2.

Drag Load Assessment a.

Drag forces for circular cylinders shall be computed on the basis Df acceleration drag alone, under the condition that U T/D 1 2.74, where m

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 CD = 3.6 in order to bound the relevant experimental data.

b.

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

2.14.5 LOCA Condensation Oscillation Oraa Loads The load definition and assessment procedures described in Section 4.4.2 of the LDR and the methodology described in NED0-25070 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 maximum source strength shall be de ined 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 interaction effects shall be iacluded 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 bourdary acceleration to the local fluid acceleration.

50 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 components occur at frequencies less than half the lowest structural frequency.

2.14.6 LOCA Chucgina Orac 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 subject to the constraints in Section 2.14.5 for the conde.1sation oscillation drag loads.

The application for the post-chug drag loads is subject to the following constraints and/or modifications.

1.

Flow Field a.

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 NED0-25070, 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 then be computed on the basis of the two nearest downcomers chugging at maximum source strengths at the above established phase relation.

b.

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.

4

51 2.

Drag Load Assessment a.

The constraints and modifications specified for the quencher

~

bubble drag loads apply.

b.

Unless the lowest structural natural frequency times the dura-tion 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 history.

The term " spike" refers to the short-duration 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 slash pressure loads b.

post-swell wave loads c.

asymmetric pool swell pressure loads d.

sonic and compression wave loads e.

downcomer air clearing loads 2.16 DIFFERENTIAL PRESSURE CONTROL REOUIREMENTS Those licensees that use differential pressure control (AP) as a pool swell loaa mitigation feature for the LTP, shall demonstrate conformance with the following design criteria as part of the PUA:

52 1.

There shall be no unacceptable change in the radiological conse-quences 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 determined to be acceptable by calculation.

3.

Design and installation of the AP system shall be commensurate 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 incor-parate 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 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> after placing the plant in the RUN mode during plant startup.

The differential pressure may be reduced below "X" psid 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> prior to a scheduled plant shutdown.

c.

The differential pressure may be reduced to less than "X" psid for a maximum of four hours during required operability testing of (scecify here those safety-related systems for whicn operability tests either release.

significant amounts of energy to the suppression pool or cannot be performed with the AP established).

53 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 tn. reactor shall be in a cold shutdown condit' ion within the subsequent 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />.

e.

A minimum of two narrow range instrument channels shall be provided to monitor the differential pressure.

Error in the AP measurement shall be no greater than 2 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 that the measurement is reduced to one indi-cation, operation is permissible for the following seven days.

If all indi-cation of the differential pressure is lost, and cannot be restored in six hours, an orderly shutdown shall be initiated and the reactor sh'all be in a cold shutdown condition within the subsequent 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />.

3.

STRUCTURAL ANALYSIS AND ACCEPTANCE CRITERIA The staff finds the general analysis techniques and proposed structural acceptance criteria set forth in the " Mark I Contaic. ment Program Structural Acceptance Criteria Plant Unique Analysis Applications Guide," (PUAAG),

NEDO-24583-1, dated October 1979, acceptable.

The proposed criteria will provide a sufficient basis for demonstrating 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.

4.

REFERENCES (Listed by Report Number)

NEDE-13426-P

" Mark III Confirmatory Test Program One-Third Scale Pool Swell Impact Tests, Test Series 5805," General Electric Proprietary Report, August 1975.

54 NED0-21471

" 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-Quancher 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 Report, 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 Containment Program Submerged Structures Model Main Vent Air Discharges Evaluation Report, Task Number 5.14.1," General Elactric Proprietary Report, March 1979.

... ~

55 NEDE-23545-P

" Mark I Containment Program 1/4 Scale Pressure Suppression Pool Swell Test Program:

LDR Load Tests - Generic Sensi-tivity, Task Number 5.5.3, Series 1," General Electric Proprietary Report, December 1978.

NEDE-23749-1-9 " 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 1978.

NEDE-24537-P

" Mark I Containment Long Term 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 Containment Program Full Scale Test Program Final Report, Task Number 5.11," General Electric Proprietary Report, April 1979.

NED0-24583

" 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.

56 NEDE-24645-P

" Mark I Containmeat 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 Sub-merged Structures, Task Number 5.14.2," General Electric Proprietary Report, May 1979.

.