ML20134G309

Rev 0 to Mark I Wetwell-to-Drywell Differential Pressure Load & Vacuum Breaker Response for James a Fitzpatrick Nuclear Power Plant

ML20134G309
Person / Time
Site:	FitzPatrick
Issue date:	01/31/1985
From:	Bilanin A CONTINUUM DYNAMICS, INC.
To:
Shared Package
ML20134G312	List: ML20134G309
References
84-25, 84-25-R, 84-25-R00, NUDOCS 8508230217
Download: ML20134G309 (25)

Text

_ _ _ _

1 C.D.I. TECH NOTE NO. 84-25 e

I I

MARK I WETWELL TO DRYWELL DIFFERENTIAL t

h PRESSURE LOAD AND VACUUM BREAKEP. RESPONSE FOR THE JAMES A. FIT 2 PATRICK NUCLEAR POWER PLANT l

f Revision 0 Prepared by CONTINUUM DYNAMICS, INC.

P.O. BOX 3073 PRINCETON, NEW JERSEY 08540 Prepared under Purchase Order No. 205-YF495 f'or CENERAL ELECTRIC COMPANY

~

175 CURINER AVENUE SAN JOSE, CALIFORNIA 95125 Approved by M

(

Alan Bilanin January 1985 0508230217 050820 PDN ADOCK 05000333 P

PDR

i^'

e li lY i.;

lI lt

SUMMARY

li le

}h Hark I wetwell to drywell vacuum breaker (VB) actuation velociMes during the chugging phase of a postulated loss of coolant accident (LOCA) are ll predicted.

Data collected during the full scale test facilf ty (FSTF) test II series is used to conservatively predict the dffferential pressure load across j

the VB.

Adjustment is made for plant-unique drywell volumes with a vent 3

l dynamic model validated against FSTF test data.

The predicted differential I

pressure load is used to drive a valve dynamic model with the plant-specific l

l VB valve characteristics.

The valve dynamic model, validated against full, scale test

data, conservatively predicts actuation velocities.

These j

velocities are predicted on a plant-unique basis, and presented in this

'ij report.

i e

t 1

4 Application of the above methodology to the James A. Fitzpatrick Nuclear j

Power Plant results in a negative dif ferential pressure peak of 0.28 psid, j

applied across installed 30-inch A&M external vacuum 1,reake rs, with no j

predicted valve actuation.

I t

1 1

4 e

i

j I

i TABLE OF CONTD4TS (i

r I'I Section Page 1

i Sue:ma ry 1

i 4

1 Introduction 1-1 2

. Forcing Function Hethodology 2-1 A

3 Vacuum Breaker Methodology 3-1 l

l 1

4 References 4-1 5

l 2

1 j

A 1

4

.I 1

7 4

I l'

ii Ii

),

I t

E f

{

11 E

i

+

8

.A.

j 1.

INTRODUCTION f4 d

i l

The Hark 1 long term containment program included the construction of a a

id full scale test facility (FSTF) modeling a 1/16th sector of a Mark I torus and f

ring header, with eight downcomers.

A series of tests simulating a loss of j'

coolant accident (1.0CA) demonstrated a chugging phenomenon occurring at the j

ends of the downcomers.

Continuum Dynamics, Inc. (C.D.I.) was requesteo to d

examine the FSTF geometry and develop c vent acoustic model for predicting the dif fe rentir.1 pressure across wetwell to drywell vacuum breakers during the

..'l chugging phenomenon.

Concurrently, C.D.I.

developed a valve dynamic model that includes the ' hydrodynamic effects of pressure alleviation across the j

valve disc when the valve is partially open. These two efforts are summarized in Sections 2 and 3, respectively, of this report.

)

3 These methodologies have recently been reviewed and accepted by the Nuclear Regulatory Commission (Ref. 1).

This report documents the application of these methodologies to the James A.

Fitzpatrick Nuclear Power Plant

+

i (hereafter referred to as Fitzpatrick).

\\

f b

.t 1-1 4

l 2.

FORCING FUNCTION METHODCI.0GY i

' i This section of the report suecarizes the methodology used to define s

j plant-unique wetwell to drywell Mark 1 vacuum breaker differential pressure

j forcing functions f rom FSTF data.

Additional details of the analysis may be found in Refs. 2 and 3.

?

During the Mark 1 FSTF test series, wetwell to drywell vacuum breaker ectuation was observed during the chugging phase of a postulated LOCA.

This observation lead to.the development of a methodology defining the plant-unique l

pressure loading function acting across a vacuum breaker during the chugging l

phenomenon.

The methodology idealized the FSTF as an interconnection of simple acoustic elements and modeled the chugging phenomenon as a condensation J

1

. process occurring at the exit of each downcomer across the steam water j

interface. The FSTF drywell airspace pressure time history data was used with i

}

a vent dynamic model to coepute the consistent condensation source velocity time history during chugging. The FSTF ring header pressure time history data was then used to validate the methodology.

j For plant-unique applications the most important parameter controlling the j

magnitude of the vent pressure oscillations (and hence the VB forcing ri

(:

function) was determined to be the ratio of the dryvell volume to main vent area.

These forcing functions are specified as time histories of the differential pressure across the valve disc, using the time segment of actual

[l FSTF data that generated the most conservative condensation source strength, i

The steps taken in the development of the plant-unique forcing function model are shown in Figure 2-1.

Step 1 involves the development of analytical rij models for: the unsteady motion in the steam vent system (characterized as t

y shown in Figure 2-2); the dynamics of condensation across the steam water oj interface (schematically shown in Figure 2-3); and the dynamics of the f

suppression pool and the wetwell' airspace (idealized as shown in Figure 2-

{

4).

In the analysis the condensation source is a velocity time history representing the transport of steam into water at the steam water interf ace.

[

2-1 I

~

i.i k

Lg STEP

9

3.

1 UEVELOP A

DYNAMIC MODEL OF THE J

lf VENT

SYSTEM, STEAM WATER INTER-FACE AND POOL SLOSH WITH THE CONDENSATION RATE AT THE INTER-

.l FACE UNKNOWN.

.2.

3.j 1

j 2

USE MEASURED DRYWELL PRESSURE TO jj DETERMINE THE CONDENSATION RATE.

)

1 1j 3

WITH THE CONDENSATION RATE DETER-j MINED, PREDICT UNSTEADY PRESSURES

.t AT OTHER VENT LOCATIONS TO VAll-t DATE THE MODEL.

4 a

Il

'j I4 USE THE CONDENSATION SOURCE AT THE 3

VENT EXIT TO DRIVE DYNAMIC MODELS It OF HARK l

PLANTS TO DETERMINE ii PLANT UNIQUE VACUUM BREAKER fi FORCING FUNCTIONS.

q

.3 Figure 2-1.

Steps in determining plant-unique vacuum breaker forcing functions.

,i 2-2

b t

.i k

4

?

to

- Drywell i

.]

.j External Vocuum 1

Breaker Piping 9 % Jet Deflector Plate

,3

..i I

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8 l

I 11 8--

Main Vent

t

=

i 7

1 J.

Wetwell I

6 l

5 l

4 4

12 Airspace

i c'

b 3

2 I

d A-

~

~

$v f

a I

Downcomers a)j

)

in k

.i 1

9 1

i Figure 2-2.

Sche =atic model of the vent system depicted by 12 dynamic components.

3 A

i 1

4 1

4 2-3 t

3 i

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l i

i

{

Steam Side i

m

. r, s-

)

_ _ _ _ _ S_t e a m_ _Wa t e r_. _

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-u v w

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• dh Water Side w

dt b

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Q hll Figure 2-3.

Details of the steam water interface.

u.

t:

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2-4

,s.

g r

')

4 I

-1:4 jWetwell Airspace q) -

.a

/

+-A e

P 4

)

Pool 4 6

'7 1

')-

y t

i H

h 1

o n

ll y

y W

h w

j o

lf i

f i f my 1q u=0 f

k e

1.

A Figure 2-4.

Details of the pool dynamic model around each downcomer.

t c

2-5

1 1

r,*

For the purposes of step 1, this velocity time history is assumed to be unknown.

The steam dynamics in the vent system a re governed by one-

)

dimensional acoustic theory (in the configuration used here, element 3 in 1

Figure 2-2 is nulled).

Jump conditions across the steam water interface are the Rankine-Hugoniot relationships.

A one-dimensional model of the suppression pool (assigning an equal share of the wetwell airspace volume and fl pool area to each downcomer) was developed to account for the compression of the airspace with the lowering of the steam water interface in the downco=ers.

I

~

j For plants with external lines connecting the vacuum breakers to the main

.j vent and the wetwell airspace (elements 11.and 12 in Figure 2-2), additional l

analysis and bounding linea rized loss coefficients obtained from subscale acoustic tests (Ref. 4) are included in the vent model to conservatively j

predict the differential pressure across the VB disc.

Internal vacuum j

breakers are attached at the main vent intersection with the ring header,

{

element 7 of Figure 2-2.

The same condensation source velocity time history is assumed to act at the end of each downcomer.

.]

ai Step 2 involves determining the condensation source velocity time history C

, by using the FSTF measured drywell pressure time history data during the j'

period of most severe chugging..

o l:

}

Step 3 involves validation of the model in the FSTF by using the condensation source velocity time history determined in step 2 to predict the

/

pressure elsewhere in the FSTF. A prediction of the ring header pressure time I

J history was made and compared with experimental data.

To bound the negative f

?

pressure peaks, a load factor of 1.06 was used to multiply the predicted results to match the largest pressure data spike.

To identify the origin of h.

the nonconservatism in the vent dynamic model, the input parameters to the model were varied by wide margins without altering the results (Ref. 5).

The origin of the nonconservatism appears to result from the assumption of g

applying an averaged condensation source of each downcomer exit.

This I

4 assumption was required 'because sufficient independent data sets do not exist to determine the condensation source at the exit of each downcocer independently.

r

{

2-6 i

9

(/

v, Step 4 applies the modified condensation source velocity time history to the plant-unique vent dynamic model.

The key assumption is made that the condensation source at the end of a downcome r is plant independent.

The f

amount of steam condensed per chug per downcomer is assumed to be the same

_j between the ' FSTF and Mark ' l plants.

This ' assumption is supported by the observation that the condensation rate is fixed by local conditions at the y

vent exit, such as steam mass flow' rate, noncondensibles, and thermodynamic

7 conditions. These local conditions will vary only slightly between plants.

The only plant characteristics which are changed in a plant-unique l

calculation are the ratio of drywell volume to main vent area and the pool

/

submergence.

All lengths, areas and system flow and pool parameters a re q'

retained at their FSTF values in a plant-unique calculation.

Thus, gross depressurization, controlled by drywell volume, is corrected on a plant-unique 3

basis, while high f requency ring out at the vent natural frequency is not plant-unique and is essentially taken to be that of the FSTF.

The plant a.,

drywell may be treated as a capacitance or as an acoustic volume composed of two right ~ circular cylinders standing end to end.

The capacitance model 9:.

a results in a more conservative forcing function for Fitzpatrick.

• f For plants with external lines connecting the vacuum breakers to the main vent and the wetwell airspace, additional inputs are needed for the plant-q

' unique calculation.

These inputs include the length along the main vent from

.J the drywell to the external line (the length of element 8 in Figure 2-2), the length of the external line on the drywell side.of the vacuum breaker (the length of element 11) and the diameter of the external line.

Element 12 in Figure 2-2 is taken to be zero length bec'ause of the low signal content in the wetwell airspace.

i The plant characteristic parameters given in Table 2-1 were used to compute the differential pressure time history across the vacuum breakers in Fitzpatrick.

Figure 2-5 shows the resulting differential pressure time i

history, without s'ddition of the pressure resulting from the submergence head.

6 1

2-7 f

f

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..--_-.___-_-,_,___m

t

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j TABLE 2-1

)

.t 3

Plant Characteristic Parameters j

for James A. Fitzpatrick Nuclear Power Plant b

FSTF Parameter

i

-}

l Main vent area /downcomer area ratio 0.99 3

Main vent length 37.32 ft d

Header area /downcomer area ratio 1.47 a

5

};

Header length 15.0 ft 2

Downcomer area 3.01 ft f

Downcomer length 10.E ft 2jf Vent / pool area ratio 0.045 Plant-Specific Paraceter s-j Drywell volume / main vent area ratio 474.35 ft

)j' Submergence head 4.29 ft water

?!

I g

Lower drywell volume length 50.21 ft u

I 2

Lower drywell volume area 2199.0 ft

~

i Upper drywell volume length 51.90 ft k

2 j

Upper drywell volume area 489.4 ft l

~ Distance along main vent from drywell to external line 14.87 ft External line length 11.06 ft External line diameter 2.5 ft 2-8

3 2.

8 1.

m 1

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5 TIME CSEC)

Figure 2-5.

Differential pressure time history predicted adross a vacuum breaker located i

on the external line in Fitzpatrick.

Submergence head has not been added, a) 0 - 5 seconds.

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3.

VACUUM BREAKER METHODOLOGY This section of the report summarizes the methodology used to construct the Mark I vacuum breaker valve dynamic model including hydrodynamic effects.- Additional details of the analysis may be found in Ref. 6.

During the Mark 1 shakedown tests, the vacuum breaker displacement time history was recorded.- A methodology was developed that uses the differential pressure forcing function across the VB (computed by the. vent dynamic model) and includes the effect of torque alleviation as a consequence of fluid flow through the opened valve.

W1th the valve in an open position, the differential pressure acting across the valve disc is less than the applied pressure, because of flow across the face and around the edges of the open disc.

The purpose of the analysis is to take credit for the reduction of static pressure across the valve disc as a consequence of flow.

Hydrodynamic torque reduction is estimated using the following procedure:

1)

A linear analysis for the flow field on either side of an arbitrarily I

moving disc permits the solution for the local pressure and velocity in the vicinity of the valve disc.

l' 2)

The flow is modeled as a mathematical combination of. sources and sinks around the circumference of the open disc, with the local pressure obtained in step 1 used to evaluate the strength of the sources and sinks.

l l

3)

The complete response of the ~ valve to this resulting flow and to the applied differential pressure obtained from the vent dynamic model is I

then calculated.

In all cases, the inclusion of the hydrodynamic l

torque tends to reduce the actual differential pressure and hence load acting on the valve disc.

3-1

o t

I Comparison of the valve dynamic model with Mark 1 FSTF test data from

.?

?

blowdown SDA allows validation of the valve dynamic model (Ref. 6) since both

<j valve disc displacement and differential pressure across the valve disc were 3

measured.

l 3

1 The characteristics of the VB valve in Fitzpatrick are given in Table 7

. -l 3-1.

An application of the valve dynamic model with these characteristics and j

the differential pressure forcing function determined in Section 2 results in

'l

j no valve actuation. A summary of results appears in Table 3-2.

.i d

1

.t 6

3

. l d

y i

.y l}

y 6

i t

4 j

3-2 5

i s

a g'

s t..

o O

s j

TABLE 3-1 j

t 4

~i 1

Vacuum Breaker Valve Characteristics

.l-

'j.

for James A. Fitzpatrick Nuclear Power Plant i

11 k

A 1

Vacuum breaker type 30" A&M external

.)

2 System moment of inertia 233.2 in-lb-sec System weight 312.9 lb l

System moment arm 10.25 in j

Dise moment arm 18.0 in

),

Disc area 791.7 in2 System rest angle 0.89 rad j.

Seat angle 0.52 rad

j Body angle 1.31 rad Seat coefficient of restitution 0.8 Body coefficient of restitution 0.6 1

4j i

i I'

3 0, -

t h.-

L l

3-3

TABLE 3-2 I.

Vacuum Breaker Valve Response for James A.' Fitzpatrick Nuclear Power Plant 4

{

Haximum closing impact velocity 0.0 rad /sec 2

Maximum opening angle 0.0 rad

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3-4