ML19274E489
| ML19274E489 | |
| Person / Time | |
|---|---|
| Site: | Dresden, Nine Mile Point, Oyster Creek |
| Issue date: | 03/15/1979 |
| From: | Iannni P, Marriott P GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML19274E490 | List: |
| References | |
| REF-GTECI-A-07, REF-GTECI-CO, TASK-A-07, TASK-A-7, TASK-OR NEDO-21888, NEDO-21888-01, NEDO-21888-1, NUDOCS 7903260256 | |
| Download: ML19274E489 (112) | |
Text
NEDO-21888 Class I March 1979 Revision 0 MARK I CONTAINMENT PROGRAM LOAD DEFINITION REPORT - PART B This document prepared by personnel of the Boiling Water Reactor Systems Engineering Department.
Aporoved:'
] b /> " It %
Approved:
Y P. W.
Ianni, Manager P. W. Marriott, Manager Containment Design Containment Engineering NUCLE AR ENERGY ENGINEERING DtviSION. GENE A AL E LECT 2tC COMP ANY SAN JOSE. CA LIFC ANI A 351:5 GENER AL h ELECTRIC66 rg 9 0 3 o 6 0 '
~
NEDO-21888 DISCLAIZER OF RESPOe??TBILITY This document uas prepared by General Electric pursuant to contracce with cert :in utilities owning plants utilising ?! ark I containments.
Except as otiwruise provided in such contracts, neither General Electric Company nor the individual authors:
A.
1:ake any varranty or representation, expressed or imp?ied, uith rcopect to the accuracy, completeness, or usefulness of the information contained in this document, or that the use of any information dicciosed in this document may not infringe privately owned rights; 5.
Asswne any responsibility for liability or damage of any kind uhich may result from the use of any infor~ation disclosed in thia document; or C.
Imply that a plant designed in accordance with th: recomenda-tions found in this doewnent uill be licensed by the United States Nuclear Regulatory Comission or that it vili co-mly with Federal, State or local regulations.
O ii Revision 0
NED0-21888 4.2.9 Vent Header Deflector Loads Some Mark 1 utilities will install vent header deflectors in order to mitigate pool swell loads on the vent header.
Figure 4.3.9-1 shows a typical deflector. The deflector generally consists of a pipe, sometimes with angles or tees welded on the sides, located below the header. The purpose of this section is to define the pool swell loads on the vent header deflectors.
4.3.9.1 Bases and Assumptions A combined analytical and experimental raethod was used in developing the vent header deflector loads. The experimental data base includes the plant unique QSTF test data (Reference 4.3.9-1), EPRI impact data (Refer-ence 4.3.9-2), and 1/12 Scale 3D test data (Reference 4.3.9-3).
The plant unique QSTF tests (Reference 4.3.9-1) included deflectors where they are to be utilized in the final plant configurati.n.
Figure 4.3.9-2 shows four types of deflectors tested.
Since the deflectors are located close to the initial pool surface, the load definition includes not only impact, but also acceleration and stand-ard drag loads.
In the load definition described here, these three compo-nents are defined and the$. added together to define the total applied load as a function of tine.
The development of the load definition requires two specifications:
e The transient fluid flow field past tne deflector, and The apprcpriate impact and drag equations (to predict the loads e
on the deflector from the fluid flow field information).
A detailed explanation of the procedures used to generate the transient flu id flow field (displacement, velocity, and acceleration) and the resultant de-flector loads is given in Reference 4.3.9-4 4.3.9-1 Revision 0
NED0-21888 The impact load is specified by defining the impact impulse, the peak impact load, and the shape of the impact transient. The impact impulse is defined as in Section 4.3.4:
NY I = Ag 144 c
where I = impact impulse per unit area, psi-sec K
proportionality factor = 0.3
=
M = hydrodynamic mass of the deflector, lbm h
deflector impact velocity, ft/sec V
=
A = projected area of deflector, ft Ibn - ft g = conversion constant 32.2
=
lbf - sec" O
For the deflector loads K, is based on the EPRI data (Reference 4.3.9-2),
n which is conservative because of the absence of LOCA bubbles and pool curvature in the EPRI tests.
The peak impact load is based on a correlation established in the EPRI tests (Reference 4.3.9-2);
7 P
= 7/2aV~
Max where P
= maximum load (lbf) on the structure divided by the structure Max horizontal projected area (in )
p
= water density O
4.3.9-2 Revision 0
NEDO-21888 Finally, the impact transient shape is assumed to be triangular (Figure 4.3.9-3) based on the results reported in Reference 4.3.9-2.
For those plants whose deflector is initially partially submerged there is no impact load.
Drag loads on the deflector include both standard drag and acceleration drag. Acceleration drag is calculated based on (Reference 4.3.9-5);
P.D L F
=
A 2g or FtD L + U V
F
=
A 4g c
g D
c c
where a
F = acceleration drag, lbf D = deflector diameter, ft, if it ir, a cylinder length of deflector, ft L
=
O transient fluid acceleration, ft/sec
=
D = diameter of a circle circumscribing the deflector cross section if the deflector is not a cylinder V = Displaced volume of deflector, ft D
The acceleration drag is assumed to be zero at impac t and to reach the full value specified when the undisturbed water surf ace reaches the elevation of the maximum height of the deflector. A linear variation with elevation is assumed.
4.3.9-3 Revision 0
NEDO-21888 Standard Drag is calculated using:
1
= C DL 2g pV' FD D
c where V = transient fluid surface velocity, ft/sec F
"E' D
ag e
n, c
s a function of traersion depth C
=
D (See Reference 4.3.9-4) diameter of deflector or circumscribing circle.
D
=
The load definition procedure described up to this point produces deflector loads which correspond to the average downcomer spacing.
Deflector loads at different longitudinal (z/t) stations are calculated by applytng a multi-plier to the velocity and acceleration for the average downconer spacing.
The multiplier, which is a function of z/t, is based on the 1/12 scale 3D pool velocity data (Reference 4.3.9-3).
Figure 4.3.9-4 shows the multiplier llh for velocity and acceleration vs z/t.
See Section 4.3.2 for a definition of a/i.
4.3.9.2 Load Definition The deflector loads are calculated on a plant unique basis and are presented in the Plant Unique Load Definitions.
Figure 4.3.9-5 shows a typical deflector load transient at chree locations for a typical Mark I plant in Ibf per foot of deflector. Alternately, an individual plant may select to utilize deflector load data obtained from plant unique supplementary QSTF tests, to define vent header deflector loads.
4.3.9.3 Fallback Loads on the Deflector ihe QSTS plant unique movies (Reference 4.3.9-1) indicate that there are no bulk pocl f allback loads on the deflector because the deflector and header tend to push the pool water out toward the torus walls as the pool swells upward. The movies do indicate that there would be some small froth 4.3.9-4 Revision 0
NEDO-21888 fallback load resulting from the froth located between the header and deflector. This load can be bounded by using the froth fallback load procedure explained in Section 4.3.5.
For this load definition the froth fallback dansity should be 10%.
4.3.9.4 Application of Vent Header Deflector Load The vent header deflector load should be applied in the upward direction within a range of directions including a variation of !10 from the upward vertical. The froth fallback load should be applied in the downward direction with a range of directions including a variation of 145 from vertical.
4.3.9-5/4.3.9-6 Revision 0
1 DEFLECTOR 7,
I
% X/R I
I I
I l
1 l
l X/R = 1.0 X/R = 0.0 Figure 4.3.9-1.
Typical Vent Header Deflector 4.3.9-7 Revision 0
NED0-21888 O
O j TYPE 1 TYPE 2 j
V U
PIPE PIPE WITH ANGLES TYPE 3 TYPE 4 PIPE WITH TEES WEDGE Figure 4.7.9-2.
Typical Deflector Designs O
4.3.9-8 R"*
NEDO-21888 PMAX-IMPU LSE IMPACT PRESSURE 4
DU R ATION I
TIME Figure 4.3.9-3.
Deflector Impact Loading Transient 4.3.9-9 Revision 0
1.5 1.4 1.3 1.2 us
- m E
1.1 a
C
.=
^
?.o Oa
.~_
m o.9 e
Ed 0.8 0.7 0.6 0.5 o
0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 nl:
Figure 4.3.9-4.
Lon itudinal Multiplier for Fluid Velocity s
and Acceleration O
4.3.9-10 Revision 0
SEDO-21888
=
5=
U 8
u.
s\\
\\
T)ME tmsec)
Figure 4.3.9-5.
Typical vent 11eader Deflector Load 4,3,9 11/4.3.9-l',
Revision 0
NEDO-21888 REFERENCES FOR SECTION 4.3.9 4.3.9-1 J. M. Humphrey, " Mark I Containment Program, 1/4 Scale fwa-Dimensional Plant Unique Pool Svell Test Report," General Electric Company, Report No. NEDE-21944-P, to be published.
4.3.9-2 The Electric Power Research Institute, " Water Impact Tests of Rigid and Flexible Cylinders," EPRI NP-798, May, 1978.
4.3.9-3 The Electric Power Research Institute. "Three Dimensional Pool Swell Modeling of a Mark I Suppression System," EPR1 NP-906, October, 1978.
4.3.9-4 Kennedy, W., et al., " Mark I Containment Program, '.'ent Header Deflector Load Definition," General Electric Company, Report No. NEDE-24612-P, March 1979.
4.3.9-5 Moody, F.J., " Drag Forces on Submerged Structures in Unsteady Flow," General Electric Company, Report No. NEDE-21471-P, September, 1976.
4.3.9-13/4.3.9-14 Revision 0
NEDO-21888 4.4 CONDENSATION OSCILI ATION LOAD Following the pool swell transient of a postulated LOCA, there is a period during which condensation oscillations occur at the downcomer exit.
Condensa-tion oscillation loads are caused by periodic pressure oscillations on the torus shell, submerged structures, and in the vent system. These pressure oscillations are associated with the pulsating movement of the steam-water interface caused by variations in the condensation rate.
The loads specified for condensation oscillation are based on the FSTF tests.
In these tests a prototypical segment (one bay) of a Mark I torus and vent system were subjected to ten steam and liquid blowdowns simulating a range of LOCA's.
Duration of the condensation oscillation period varies with break size and is specified in Table 4.4.1-1.
Load combinations are given in the bar charts in Section 3.0.
Condensation oscillation loads witF the largest ampli-tude occur during the DBA.
For tne IBA, bounding chugging loads are specified in lieu of condensation oscillation loads.
The purpose of this subsection is to specify the loading conditfons which may be produced during this phase of the LOCA transient.
Loads on the structures within the Mark I Con:ainment system that may be induced by condensatiot.
oscillations are defined in the following subsections:
4.4.1 Torus shell loads 4.4.2 Loads on submerged structures 4.4.3 Lateral loads on downcomers 4.4.4 Vent systems loads 4.4-1/4.4-2 Revision 0
NEDO-21888 4.4.1 Torus Shell Loads oscillating loads on the submerged portion of the torus shell during the con-densation oscillation phenomenon are caused by periodic pressure oscillations superimposed on the prevailing local static pressures.
in the following sub-sections the definition of these loads is presented along with a discussion of the bases and assumptions used and the plant unique application considerations.
4.4.1.1 Bases and Assumptions The parameters which were varied in the FSTF tests include downcomer submerg-ence, initial pocl temperature, blowdown fluid phase (liquid and vapor), and initial wetwell pressure. The break size used in the FSTF for the DBA provides steam mass fluxes in excess of those predicted for any Mark I plant. The load-ing conditions established from the FSTF data and the use of conservative application techniques result in bounding condensation oscillation loads for all Mark I plants. The FSTF, test matrix, and test results are described in Reference 4.4.1-1.
Data from the FSTF tests indicate that the amplitude of the pressure oscilla-tions induced by condensation oscillations on the torus shell is dependent on the break size and the phase of the blowdown fluid (liquid or vapor). The highest pressure amplitude was observed during test M8, the large liquid break test which simulated the DBA conditions (Reference 4.4.1-1).
Data from this particular test run of the FSTF were used as a conservative basis for the load definition for the DBA.
For the IBA, chugging loads are specified in lieu of condensation oscillation loads.
In the FSTF tests evaluated (M1, M4, and M9) with a break size equal to 25 percent of the DBA area, strong condensation oscillatians did not persist. A break size of 25 percent of the DBA area is larger than the IBA.
Some pressure oscillations did occur in these tests during air carry-over, but the pressure amplitudes were bounded by the peak values of the chugging pressures which occurred following the air carry-over period. The IBA steam condensation loads therefore are specified as being chugging loads as defined in subsections 4.5.1.2 and 4.5.4.2.
4.4.1-1 Revision 0
NEDO-21888 The onset times and durations for condensation oscillations are specified in Table 4.4.1-1.
The onset times were selected as conservatively small values based on FSTF tests.
The durations were based on an analytical study of the DBA and IBA, from which a conservatively large value was selected for the time required to depressurize the system into the chugging regime.
The analysis was conducted on a plant configuration which would produce the slowest depressurization rate.
The SBA will not result in flow rates sufficiently large to enter the condensation oscillation regime.
Plant unique loads are derived f rom FSTF data.
Flexible wall loads were measured directly in the FSTF which is prototypical of a Mark 1 plant configuration.
Pressure measurements obtained from various locations on the torus shell show that the longitudinal pressure oscillation amplitude distribution along the torus centerline is essentially uniform. These measurements also show that the varia-tion of torus wall pressure amplitude (from zero at the water surface to maximum the torus bottom dead center) can be appropriately represented as linearly at varying with elevation (see Figure 4.4.1-2).
With the spatial distributions of pressure oscillation amplitude established, the load definition can be specified in terms of pressure at the torus bottom dead center.
Pre ~ sure oscillation amplitude variation as a function of oscillation frequency is determined from integrated total pressures at the various frequencies which were derived from torus shell pressure time histories.
The basis for the pressure amplitude specitication is several segments of maxi-mum pressure amplitude data measured during FSTF tests M7 and M8.
Amplitude frequency correlations were compiled for one second intervals of data, vs and the peak average amplitudes were determined for each 1 Hz f requency band.
These amplitudes as a function of frequency are used as the load specification.
The frequency range in the pressure amplitude specification has been established to cover the range of frequencies observed i:t the FSTF tests and the range expected in all Mark I plants. Two dominant frequencies (approximately 5 and 10 Hz) were observed in the FSTF tests M7 and M8.
These dominant frequencies were also observed to vary during the tests.
The load specification frequency ranges, which were selected to conservatively bound these dominant frequency g
variances, are 4 to 8 Hz and 8 to 16 Hz respectively.
4.4.1-2 Revision 0
NEDO-21888 Alternate frequency spectra for the 4 to 8 Hz and 8 to 16 Hz frequency ranges have been specified in Figure 4.4.1-1 and Table 4.4.1-2.
For a plant unique structural evaluation, the structural responses from each 1 Hz band between 0 and 50 Hz would be summed.
The O to 50 Hz total range analyzed would include only one of the alternate frequency spectra specified for the 4 to 16 Hz range in Table 4.4.1-2.
The alternate frequency spectrum selected to ba inc.luded in the analysis would be that which produces the maximum total response.
In order to apply the FSTF measured loads to plant unique geometries, two adjustments are made.
The first adjustment is to account for FSI effects in the FSTF data and the second is to account for differences in the ratio of the pool surface area to vent cross sectional area among the Mark I plants.
Structural response effects unique to the FSTF are accounted for by the specification of a baseline rigid wall load which is given as pressure oscilla-tion amplitude as a function of frequency. This load has been derived from the measured FSTF flexible wall load by analysis with a coupled fluid-structural dynamic model of the FSTF torus.
The derivation of the baseline rigid wall load is described below:
a.
A finite element coupled fluid-structural model of the FSTF torus was excited at varying frequencies with a unit amplitude pressure source at the vent exits. The torus shell pressure amplitudes relative to the source pressure (amplification factors) were determined as a function of frequency.
b.
Using these relative amplitudes (amplification factors), the FSTF vent exit source pressures were derived from the measured torus shell pressures at the various frequencies.
The baseline rigid wall load was derived from the computed FSTF vent c.
exit source pressures by applying incompressible potential flow re la tions (spatial attenuation from the vent exits to the torus vall which is assumed to be rigid).
4.4.1-3 Revision 0
NEDO-21888 The baseline rigid wall load is then adjusted for atteni < tion due to a plant 's particular ratio of pool surface area to vent cross sectional area.
A multiplication factor to provide this adjustment has been computed from incompressible flow theory for various area ratios (see Figure 4.4.1-3). Values for the pressure attenuation between sources at the vent exits and the bottom center of the torus were determined from the product of separate two dimen-sional potential flow solutions obtained in orthogonal planes. This method of determining pressure attenuation was proven to be conservative by comparison with a solution from a three dimensional finite element model.
Pressure attenuation factors were computed for various area ratios and normalized to the FSTF value to provide the multiplication factors shown in Figure 4.4.1-3.
Since the FSTF area ratio is smaller than that of most Mark I plants, the plant unique loads in most cases will be less than or equal to the baseline rigid wall load. The plant unique load is intended to be used in conjunction with a flex-ible wall coupled fluid-structural model in the plant unique structural evaluation.
The following major assumptions are made in the definition of the condensation h
oscillation torus load:
a.
The frequency range observed in the FSTF data is assumed to be applicable to all Mark I plants. This is justifiable since the con-densation oscillation f requency is primarily controlled by thermo-dynamic and geometric conditions at the vent exit.
These conditions are very similar for all Mark I plants and were simulated in the FSTF tests.
Minor adjustments to account for vent system and down-comer acoustic frequencies were made to conservatively represent all Mark I plants.
b.
Baseline rigid wall pressure oscillation amplitude values can be derived from the flexible wall measurements in the FSTF by the use of fluid coupled-structural dynamic models and incompressible flow theory.
O 4.4.1-4 Revision 0
NEDO-21888 c.
The effect of the ratio of the pool surface area to vent cross sectional area can be derarmined since it is assumed that the pressure oscillations will attenuate from the source at the down-comer exit to the pool boundary according to incompressible potential flow theory as described in Subsection 4.4.1.1.
4.4.1.2 Load Definition The condensation oscillation load on the torus shell is a plant unique rigid wall load.
It is intended to be used in conjunction with a flexible wall coupled fluid-structural model in the plant unique st*ictural evaluation. The plant unique loads are derived from the baseline rigid wall loads as outlined in the following subsections.
4.4.1.2.1 Baseline Rigid Wall Loads Based on the considerations discussed in the previous subsection, the following baseline rigid wall condensation oscillation load definition was derived for the DBA:
Amplitude versus frequency Valuer given in Table 4.4.1-2 (Also shown in Figure 4.4.1-1).
The total spectrum to be analyzed is from 0 to 50 Hz, and is to include the one spec-trum from 4 to 16 Hz which produces the maximum total response.
Total Response Resulting responses, from applying the amplitude at each frequency given in the total spectrum to be analyzed, are to be summed.
Spatial Distribution Uniform axially along the torus center-line.
Linear attenuation with sub-mergence along the wetted perimeter of the torus cross section as shown in Figure 4.4.1-2.
4.4.1-5 Revision 0
NEDO-21888 For plant unique applications where the ratio of peal area to total downcomer area is greater than that of the FSTF, the magnitude of the condensation oscillation bascline rigid wall load is reduced. A discussion of this adjust-ment is presented in the following subsection.
4.4.1.2.2 Plant Unique Loads Since the dimensions of the torus and the number of downcomers vary from plant to plant, the magnitude of condensation oscillation loads is expected to be different for each plant. To facilitate calculation of condensation oscilla-tion loads for plant unique applications, a multiplication factor has been developed to account for the effect of the pool-to-ve-+ area ratio.
In Figure 4.4.1-3 this factor is plotted versus the pool ;o-vent area ratio.
To determine the plant unique load, the pool-to-vent area ratio is computed.
This value may be based on the average value for the entire plant for evalua-tions involving the total net vertical pressure force (e.g.,
in torus support system load evaluations).
For all other evaluations, the smallest value of pool-to-vent area ratio for any bay in the torus shall be used.
Then, from Figure 4.4.1-3, the multiplication factor for the appropriate calculated pool-to-vent area ratio is obtained. The plant unique oscillating load on the torus is determined by multiplying the amplitude of the baseline rigid wall load (Section 4.4.1.2.1) by this factor. The resulting plant unique oscillating load is to be applied to the prevailing local static pressures at the various locations on the torus shell at the appropriate times (see Table 4.4.1-1).
O 4.4.1-6 Revision 0
NEDO-21888 Table 4.4.1-1 CONDENSATION OSCILLATION ONSET AND DURATION Onset Time Du ra t ion Break Size After Break After Onset DBA 5 seconds 30 seconds IBA 5 seconds
- 900 seconds
- SBA Not Not Applicable Applicable
- For the IBA, chugging loads as defined in subsections 4.5.1.2 and 4.5.4.2 are applied to the torus shell and vent system, respectively.
4.4.1-7 Revision 0
NEDO-21888 Table 4.4.1-2 CONDENSATION OSCILLATION BASELINE RIGID WALL PRESSURE AMPLITUDES ON TORUS SHELL BOTTOM DEAD CENTER Alternate Amplitudes t be Analyzed
- Frequency Amplitudes
( PSI)
Range to be Analyzed *
(Hz)
( PSI) 1 2
3 0-1 0.29 A
1-2 0.25
~
2-3 0.32 l
3-4 0.48
?
4-5 A
1.86 1.20 0.24 5-6 1.05 2.73 0 '. 8 6-7 0.49 0.42 0.99 w
=w 2$
0.59 0.38 0.30 7-8 x
x5 52 8-9 ygg g
0.59 0.38 0.30 9-10
"'5dM$"
0.59 0.38 0.30 Jh 10-11 0.34 0.79 0.18 11-12 5555N5 0.15 0.45 0.12 h,$
hh 12-13 0.17 0.12 0.11 hhhghh 11-14 0.12 0.08 0.08 14 -15 0.06 0.07 0.03 1
15-16 0.10 0.10 0.02 16-17 0.04 A
17-18 0.04 18-19 0.04 19-20 0.27 20-21 0.20 NONE 21-22 0.30 22-23 0.34 23-24 0.33 24-25 0.26 7
- Half range (= 1/2 of peak to peak amplitude) 9 4.4.1-8 Revision 0
NEDO-21888 Table 4.4.1-2 (Continueo)
CONDENSATION OSCILLATION 3ASELINE RIGID 'n'ALL PRESSURE AFIPLITUDES ON TORUS SilELL BOTT031 DEAD CENTER Alternate Amplitudes t be,\\ulyzed*
Frequency Amplitudes Range to be Analyzed *
(ilz)
(PSI) 1 2
3 25-26 0.25 A
26-27 0.58 27-28 0.13 28-29 0.19 29-30 0.14 30-31 0.08 31-32 0.03 32-33 0.03 2 -34 0.03 34-35 0.05 35-36 0.08 l
36-37 0.10 37-38 0.07 NONE 38-39 0.06 39-40 0.09 40-41 0.33 41-42 0.33 42-43 0.33 43-44 0.33 44-45 0.33 45-46 0.33 46-47 0.33 47-48 0.33 48-49 0.33 49-50 0.33
?
- Half range (= 1/2 of peak to peak amplitude) 4.4.1-9 Revision 0
O uoTsl^oE OI-I'7'7 PRESSURE AMPLITUDE (psi) o o
uT<o>
- o ;;; z m
- 2e nm o ro a.
m I
c
- w o
E b, o m
m zH u 0:
3
<$>G
- - g I
>n r o q
r C
m am Cz mm CHf9 p
HH2m or>m oop" d.
Hmwp mmgn De m P Hm>
n
~
\\ F$5H 0:
I o
/ Ixmo Emwdo 7
EE5z 2
g gc:m 50>
I 2
2[
ya m
a o m
mm m3 mx
- O.
_ /
PRESSURE AMPLITUDE losa)
-c w
7U o
M u'A ccn w
V *
=
m.
o5 s
m C
O r
o H
Hz C
m
,o o z o 3
m m
n
=~
r-b 4
z-o m
y o
C
=.
cf. n m
~
r-z r -
n u 45
<w Woz
- RESSURE AMPLITUDE (psil n c "o *-
o d
M U
a=
w i
e a
w
=cx m
c-m C. D m
4 a
m o C-C 2
e m o z
3 C U
2
- =
w n
H
,c -
m o
u
=
E 2
z
~
=
PRESSURE AMPLITUDE (psi) e g
o H
m
'2 dom >
.p gmmr m
ir m
R>t e
C:
H m
m
-m a z o 3
C'm' omzC n
m>Hy 4
?oC m
=
Hzo 8
w omm st*
-w my
>r 7m 888IZ-OG3N
rHC$
a D
. r@e ir E
D U
T IL
)
P E
e M
D A
U T
N t
\\
O L
I P
A 1
T M
A A
A
=
L X
N )R
\\
L A
I OE
^
M C
I T S
T O
A N A
l E
C
[
L E
=
m
- a L
n R
iD o
U C
S S A i
S OE t
E D
c E
e R
RM P
S UO L
ST A
ST s
C EO s
O R B o
L P
r S
MU C
UR
=
l MO I T ae X
cd A T i u A
M (A t t ri el
=
V p m
sA X
u A
r n M
o o A
Ti t
a l
nl oi i c t s aO l
l e i r c u s s os e N
nr LE O
LC I
WP SO t r E
S O
n u
ao T
R R
P E
I P
x$
sf
. I*
n P
WA U
e n S
d o e
ni ot C u b
I i r
k t r s laiD F
E C
A 2
F R
1 U _
S 4
E E
I 4
F F
e r
Y ug i
F
- .d HIC
[< a ao
^ ^ ^
POOL TO-VENT ARE A R ATIO =
VENT EXIT CROSS SECTION AL ARE A MULTIPLICATION FACTOR
=
PROTOTYPICAL RlGID WALL LOAD AMPLITUDE 1.1 FSTF 1.0 (21.2,1.0) 09 2o oa Q
n3 0.7 4s
[
@ 06 z
z~
E m
f O
cp C'
y 0.5 eb m
2 m
ra m
9r o.4 cr.
9 d
p O.3 aas 0.2 01 i
l I
l l
I I
g 15 20 25 30 35 40 45 50 55 x
POOL TO-VENT AREA RATIO ft<
r m
o Figure 4.4.1-3.
Mark 1 Condensation Oscillation - Multiplication Factor Versus l'ool-to-Vent Area a
Ratio for Plant Unique I.oad Determination v
O O
O
NEDO-21888 REFERENCES FOR SECTION 4.4.1 4.4.1-1
" Mark I Containment Program - Full Scale Test Program, Final Report," General Electric Company, Report No. NEDE-24539-P, March 1979.
4.4.1-13/4.4.1-14 Revision 0
NEDO-21888 4.4.2 Loads on Submerged Structures Due to Main Vent Steam Condensation Oscillations Steam condensation begins after the vent is cleared of water and the drywell air has been carried over into the suppression chamber. The condensation oscillation phase induces bulk water motion and theref ore creates drag loads on structures submerged in the pool.
4.4.2.1 Bases and Assumptions The basis of the flow model for condensation oscillation load definition is derived from the work in References 4.4.2-1 and 4.4.2-2.
The major assump-tions are summarized as follows:
The total drag is the sum of standard and acceleration drags a.
b.
The submerged structures are assumed to be rigid c.
Condensation oscillation is modeled as a fluid source located at the vent exit, which induces a bulk pool motion that results in drag forces on submerged structures d.
The actual flow field can be represented by a locally uniform flow field The presence of boundaries is accounted for by using the method of e.
images f.
The strength of the fluid source for condensation oscillation is determined from wall load measurements (Reference 4.4.2-3) given in Section 4.4.1 by using potential flow theory in an incompressible fluid.
4.4.2-1 Revision 0
NEDO-21388 4.4.2.2 Load Definitien Detailed procedures to calculate drag loads on submerged structures are described in Reference 4.4.2-1.
This reference provides the methodology to utilize the unsteady velocity and acceleration flow fields from condensation oscillation to obtain drag loads on submerged structures, such as pipes, beams, etc. In the suppression pool.
The load definition first requires the establishment of the velocity and accel-eration flow fields within the torus by simulating condensation oscillations at the vent exits.
Then, drag loads on the submerged structure or sections of the submerged structure are calculated based on the previously established transient flow fields.
a.
Flow Field Establishment Condensation oscillations are described as fluid sources located at downcomer vent exits.
The source strengths are determined from wall load measurements.
By using potential flow theory and the method of images to account for the effects of solid walls and the free surface, the velocity and acceleration flow fields within the torus are catablished.
b.
Drag Loads Evaluation The drag force on a submerged structure consists of two components:
the standard drag and the acceleration drag.
The standard drag is computed on a locally uniform flow field evaluated at the local velocity. The acceleration drag load is evaluated based on an inviscid, uniform but unsteady (accelerating) flow field. The sum of these two drag forces gives the total drag load on a submerged structure.
O 4.4.2-2 Revision 0
NEDO-21888 4.4.2.3 Selection of Key Parameters for Load Evaluation Procedures This section outlines basic criteria for selecting key parameters to be used as input data to the model (Reference 4.4.2-1) so that condensation oscilla-tion loads on submerged structures can be properly obtained.
The plant specific suppression pool geometry should be identified first as follows:
a.
Torus shell dimensions b.
Torus water depth Location of submerged structure considered (Figure 4.3.8-1) c.
d.
Locations of downcomers (Figure 4.3.8-1).
Then the condensation oscillation source information calculated from wall load measurements (Reference 4.4.2-3) is assigned to the downccmer exits.
The submerged structure is divided into the appropriate number of sections for more precise calculation. The coordinates of each section center, orientation of the structure axes and the structure projected cross-sectional area should be identified. The acceleration drag volume and the standard drag coefficient for each section of the submerged structure should be obtained from Tables given in Reference 4.4.2-2.
With the input parameters identified and calculated, the submerged struc-ture drag load model is used to evaluate the resultant transient drag forces on each structure element considered.
Typical results are shown in Figures 4.4.2-1 and 4.4.2-2.
These figures show the condensation oscilla-tion drag forces on an S/RV line (with T-Quencher arms) in the radial (X) and tangential (Z) directions as a function of time.
4.4.2-3/4.4.2-4 Revision 0
NEDO-21888 CONDENSATION OSCILLATION FORCE ON S/RV LINE AT 5.5 Hz 45 i
30 -
+
15 2
0/h 5
0 8W X
Y O -30
\\
-45 0
0.1 0.2 0.3 0.4 0.5 TIME (sec)
Figure 4.4.2-1.
Sample Predicted Time History of Total X-Force on S/RV Line 4.4.2-5 Revision 0
NED0-215888 O
CONDENSATION OSCILLATION FORCE ON S/RV LINE AT 5.5 Hz 180 120 60 9
2 5
0 h
0 Z
2 a
O e
-e0 t
-120 I
I l
-180 0
0.1 0.2 0.3 0.4 0.5 TIME (sec)
Figure 4.4.2-2.
Sample Predicted Time 111 story of Total Z-Forces on S/RV Line 4.4.2-6 Revision 0
NEDO-21888 REFERENCES FOR SECTION 4.4.2 4.4.2-1 Lasher, L.
E.,
" Analytical Model for Estimating Drag Forces on Rigid Submerged Structures Caused by Condensation Oscillations and Chugging - Mark I Containments," G( ?ral Electric Company, Report No. NEDE-25070-P, March, 1979.
4.4.2-2 Moody, F.
J.,
Chow, L.
C.,
and *.:Te r, L.
E.,
" Analytical Model for Estimating Drag Forces on Rigid Submerged Structures caused by LOCA and Safety Relief Valve Air Discharges, Ramshead", General Electric Company, Report No. NEDE-21471-P, September, 1977.
4.4.2-3 J. E. Torbeck, et al., " Mark I Containment Program, Full Scale Test Program Final Report," General Eletric Company, Report No.
NEDE-24539-P, March, 1979.
4.4.2-7/4.4.2-8 Revision 0
NED0-21888 4.4.3 Lateral Loads on Downcomers The downcomers experience lateral loading during the condensation oscillation phase. The procedure for defining the dynamic portion of this loading for both a Design Basis Accident (DBA) and an Intermediate Break Accident (IBA) is presented in this section. Condensation oscillation loads are not applicable for the Small Break Accident (SBA) case.
The vent system thrust loads due to the quasi-static pressurization of the containment and the fluid flow in the vent system during the condensation oscillation regime are reported separately in Section 4.2.
4.4.3.1 Bases and Assumptions The basis for the DBA and IBA condensation oscillation lateral load definition is the data obtained from the instrumented downcomers of the Mark I Full Scale Test Facility (FSTF) during testing as described in Reference 4.4.3-1.
The load definition is developed for, and is directly applicable to, downcomer pairs which are untied, i.e.,
not connected by lateral bracing.
Considerations for braced downcomers are also provided.
The FSTF downcomer lateral loads are det.ned as Resultant Static Equivalent Loads (RSEL) which, when applied statica.ly to the end of the downcomer, will reproduce at any given time the measured bending response near the downcomer/
vent header junction. The RSEL were obtained from the measured downcomer bending moments using conversion factors determined from static calibration tests performed on the downcomers, Reference 4.4.3-1.
These RSEL are, there-fore, an actual representation of the structural-hydrodynamic interaction which occurs during condensation oscillation.
The maximum design loads for individual plants are obtained by scaling the maximum RSEL from the FSTF. The scaling factccs are derived on the basis of a cor parison of the dynamic characteristics of the downcomers of the individual plants and the FSTF.
4.4.3-1 Revision 0
NEDO-21888 Additionally, the number of RSEL reversals during the FSTF DBA and IBA condensation oscillation phases were counted and presented in the f o rm o f RSEL reversal histograms.
Since the condensation oscillation loads may be oriented in any lateral direction, these histograms were obtained for each of eight angular sectors around the downcomer end as shown in Figure 4.4.3-1.
A typical RSEL reversal histogram is shown in Figure 4.4.3-2.
The FSTF RSEL reversal histograms are scaled into a plant-unique set of histograms by first scaling the maximum RSEL reversal on the same basis as was dor.e for the maximum design loads.
Then the number of RSEL reversals are scaled by the ratio of the condensation oscillation duration specified for the plants to that of the FSTF.
For fatigue evaluation of the downcomer/ vent header junction, the plant-unique set of RSEL reversals must be converted to a set of stress reversals at the downcomer/ vent header junction which result from application of the RSEL reversals.
Net loa + can be produced on the vent header due to combination of condensation oscill mon loads on individual downcomers.
Because the FSTF downcomer pairs e
were observed to respond essentially out-of-phase in their plane, the lateral vent loads are negligible.
It can be assumed that the net loada along the axis of the vent header are bounded by the chugging loads defined in Section 4.5.3 with respect to grass multi-vent chugging considerations.
This assumption is based upon the conservative way in which the chugging phenomenon was defined, i.e., the randomness in time of the occurrence of chugs on the downcomers is ignored.
The procedure for determining condensation oscillation lateral loads on down-comers includes the following assumptions:
a.
The vent header and downcomers can be considered rigid when compared to the flexible downcomer/ vent header junction.
Therefore the down-comer responds dynamically in any given direction as a single degree of freedom system. This assumption has been verified by analytical studies (Reference 4.4.3).
O 4.4.3-2 Revision 0
NEDO-21888 b.
The FSTF lateral hydraulic laading mechanism occurring during con ~
densation oscillation is representative of the loading mechanism that would be experienced by other Mark I downcomers for the same LOCA conditions (Reference 4.4.3-1).
The FSTF RSEL reversal histograms obtained for the condensation c.
oscillation phase of the LOCA are representative of an actual plant behavior over the same duration (Ref erence 4.4. 3-1). Thus, for the fatigue evaluation of an individual plant's downcomer, the FSTF RSEL reversal histogram data must be scaled b the ratio of the condensation oscillation duration in the plant to the duration simulated in the FSTF.
d.
The condensation oscillation loads are approximated as sinusoidal in nature for the purpose of dynamic scaling. This is based on observa-tion of the FSTF condensation oscillation data (Reference 4.4.3-1).
e.
The load development procedure is applicable for loadings where the measured bending strains on the downcomer remain within the elastic range. These strains were observed to be well within the elastic range for all of the FSTF tests (Ref erence 4.4.3-1).
4.4.3.2 Evaluation Procedure The DBA and IBA loads associated with condensation oscillatioa obtained from the FSTF data will be scaled to determine the loads for other Mark. plant downcomers. For each plant, both the maximum design load and histograms of load reversals for fatigue analyses may be determined f rom the FSTF loads.
The maximum condensation oscillation design load P for an individual Mark I plant downcomer is obtained by scaling the maxinum downcomer lateral load P observed during the condensation oscillation phase of the FSTF tests.
7 The scaling law which incorporates the dynamic characteristics of ti.
FSTF 4.4.3-3 Revision 0
downcomers and the downcomers of an individual plant is given in Reference 4.4.3-1 as:
[DLF p
p 1(DLF max where DLF and DFL are the dynamic load fas< ors of the untied plant-unique and FSTF downco w r/ vent header structures, respective 1/.
The dynamic load factors depend upon the natural frequencies of the single-degree-of-freedom systems representing the down:omer/ vent header structures.
The natural frequencies in the North-South and East-West directions bound all possible downcomer frequencies and can be determined once the rotational stiffnesses of the downcomer/ vent header junction and the mass moments of inertia of the downcomer in these two principal directions have been ascer-tained. The mass moment of inertia must include the added mass of the water for the submerged portion of the downcomer. The dynamic load factors are dorer-mined using the downcomer natu"al frequency which most closely matches the condenation oscillation driving frequency range.
A maximum downcomer lateral load of 4,122 lbs was determined from the FSTF DBA condensation oscillation data to occur within 22-1/2 of the East-West direction, i.e., in the plane of the downcomer pairs, as identified by sectors 4 and 5 in Figure 4.4.3-1.
Similarly, a maximum IBA load of 2,589 lbs was observed to occur within 45* of the East-West direction, i.e.,
sectors 3 through 6 in Figure 4.4.3-1.
The maximum design load P derived from either the DBA or IBA loads must be applied to the end of the plant-unique downcomers within che above-defined respective directions such as to maximize the stresses at the downcomer/ vent header junction. The maximum downcomer lateral load in the North-South direction, i.e. along the axis or the vent header is bounded by the peak chugging load defined in Section 4.5.3.
For fatigue evaluation of the downcomers, the required stress reversals at the downcomer/ vent header junction can be obtained from the FSTF RSEL reversal histograms. Only RSEL reversals which are greater than a threshold of 5* of the maximum load range were counted for the histograms, since cycles of smaller 4.4.3-4 Revision 0
NED0-21888 magnitude do not contribute significnntly to fatigue usage. The junction stress reversals are obtained by first scaling the FSTF RSEL reversals into a plant-unique set of RSEL reversals. This scaling procedure is a two-fold process.
First, the maximum FSTF RSEL reversal is scaled on the same basis as was done for the maximum design loads.
Then, the total number of reversals must be scaled by the ratio of the condensation oscillation duration specified for the plants to that of the FSTF. Condensation oscillation durations for both the DBA and IBA are specified in Table 4.4.1-1 of Section 4.4.1 as applic-able to all plants.
Based upon actual FSTF condensation oscillation durations, the number of FSTF DBA reversals must be scaled by 1.25 and IBA FSTF reversals by 7.20.
The plaat unique set of RSEL reversals may now be transformed into a set of stress reversals at the downcomer/ vent header junction. One method which can be used to transform the loads to stresses at a particular point A near the downcomer/ vent header intersection (Figure 4.4.3-3) is through a load-stress transformation matrix. This matrix may be obtained from a detailed static finite element analysis of the downcomer/ vent header structure.
First, a unit load F wauld be applied at the downcomer end in the North direction shown in Figure 4.4.3-3 to obtain K,, F,,
c
=
a a a where c is a representative stress measure used in the fatigue analysis.
Similarly, a unit load F w uld be applied at the downcomer end in the East E
direction to obtain a similar stress measure b "*
~
b EE whicL upon combining stresses yields a
b
[ K,d K]
c = c
+c
=
E F
_ E_
4.4.3-5 Revision 0
NEDO-21888 For an arbitrary loading P having components P and P the resultant E,
stress is EP a =
where the transformation matrix E will, in general, be dif ferent for different locations of stress do:ermination.
The total nur.ber of stress reversals at a location selected for fatigue evaluation is obtained by summing the stress reversals produced at that loca-tion by the condensation oscillation RSEL reversals in each sector.
4.4.3.3 Considerations for Braced Downcomers The preceding material described the specification of downcomer lateral loads for unconstrained downcomer systems based upon experimental data obtained from the FSTF. The purpose of this section is to develop DBA and IBA condensation oscillation loads for tied downcomers based upon FSTF test comparisons between tied and untied downcomer pairs.
For Mark I plants containing structural ties between each downcomer pair which are of a ttiffness comparable to the 2-inch by 1/4-inch tie straps of the FSTF, the preceding methodology for defining a maximum DBA or IBA load on untied downcomers as outlined in Section 4.4.3.2 should be followed. The plant-unique loads may then be multiplied by reduction factors derived f rom the FSTF data.
For the DBA, a conservative reduction factor of 0.75 is specified and for the IBA this factor is 0.85.
The reduced loads are then applied to the untied down-comer in the directions previously specified. The underlying assumption to this method is that the same degree of reduction observed in the FSTF tests would also occur between tied and untied downcomers in an actual Mark I plant, and that reduction factors specified may be applied to both the plant-unique design load and the maximum stress range to which the histogram for fatigue evaluation is normalized.
O 4.4.3-6 Revision 0
NEDO-21888 For Mark I plants with downcomer lateral bracing systems which provide stiff-ness in excess of that provided by the FSTF ties, the above procedure for applying DBA and IBA load reduction may be conservatively followed.
4.4.3-7/4.4.3-8 Revision 0
NEDO-21888 SECTION OF DOWNCOMER/
HEADER STRUCTURE j
l
\\
N N
s a
m N
1 0
00 337.50 22.50 0
315 450 1
2 292.5 0
3 67.5 4
E 270 900 0
?
5 6
112.5 247.5o 7
1350 157.50 202.5
- 180 Figure 4.4.3-1.
Sectors Used to Define Directions of Lateral Loads on Downcomer's End 4.4.3-9 Revision ^
30 N] L s
0 0
25 0
270 -
- 90 20 180 5c Z
h L~
w O
E 15 14
,a Y
o e
K oo o
co 23 Z
10 g
9 7
5 5
5 5
4 4
4 4
3 2
1 l
1 0
0 10 20 30 40 50 60 70 80 90 100 V'*
RSE L (PEHCENT) 4 O
3 Figure 4.4.3-2.
Distribution of Condensation Oscillation RSEI, Reversals for a Typical Sector O
O O
O
V a
t N-I
't b"
F E Figure 4.4.3-3.
Notation Used for fransforming RSEL Reversals into Stress Reversals at a Fatigue Evaluation Location A 4.4.3-11/4.4.3-12 Revision 0
NEDO-21888 4.4.3-1
" Mark I Containment Program, Development of Downcomer Lateral Loads from Full Scale Test Facility Data, Task Number 7.3.2",
General Electric Company, Report No. NEDE--24537-P, March 1979.
4.4.3-13/4.4.3-14 Revision 0
NEDO-21888 4.4.4 Vent System Loads Oscillating loads on the vent system during the condensation oscillation phenomenon are caused by harmonic pressure oscillations superimposed on the prevailing local static pressures in the vent system.
The vent system includes main vents, the vent header, 'nd downcomers.
In the following subsections the definition of these loads is i esented on a generic basis along with a discus-sion of the bases and assumptions used.
4.4.4.1 Bases and Assumptions The basis for the vent system internal pressure loading due to condensation oscillation is the data from FSTF test M8, the large liquid break test which simulated the DBA conditions (Reference 4.4.4-1).
This is the same basis as that presented in subsection 4.4.1.1 for the tores shell loads.
Note that fur the IBA, chugging loads are specified in lieu of condensation oscillation loads (see subsection 4.4.1.1).
An inspection of the PSD's for pressure signals observed at various locations in the vent system indicates the existence cf a common dominant frequency.
The signals in the main vent and the vent header exhibit no frequency content higher than the common dominant frequency. This enables the forcing function for the main vents and vent header to be defined as sinusoidal.
Signals from pressure transducers in the downconers have two different charac-taristics. At times, some downconers show a common dominant frequency; at other times, other downconers contain higher frequency components similar to those observed on the torus shell. To be conservative, the pressure for the down-is derived f rom the FSTF downcomer data and specified as an amplitude comers as a function of frequency.
The downcomer pressure load specification is a single distribution of pressure as a function of frequency as opposed to several distributions which are speci-fled for the torus shell. Variations of the dominant frequencies that were specified for the torus shell are not specified for the downcomers. These variations are not specified because the structural frequencies of the 4.4.4-1 Revision 0
NEDO-21888 downcomers are greater than the maximum values of the variations, and the downcomers would not be affected by this frequency variation.
Major assumptions made in defining vent system condensation oscillation loads include:
a.
The vent system load values determined f rom the FSTF data are generic and can thus be applied to all Mark I plants. This is justifiable because the source strength for condensation oscilla-tions is mainly controlled by parameters such ae the downcomer diameter and the thermodynamic conditions at the downcomer exit which are very similar for all Mark I plants.
b.
The frequency range observed in the FSTF data is assumed to be applicable to all Mark I plants. This is justifiable since the con-densation oscillation frequency is primarily controlled by thermo-dynamic and geometric conditions at the vent exit.
These conditions are very similar for all Mark I plants and were simulated in the FSTF tests.
Minor adjustments to account for vent system acoustic frequency were made to conservatively represent all Mark I plants.
4.4.4.2 Load Definition Condensation oscillation loads are specified for all three components of the vent system: main vents, the vent header, and downcomers.
These loads, as determined f rom the FSTF data, are generic and are thus directly applicable to all Mark I plancc.
Main Vent and Vent Header implitude 2.5 psid.
Frequency Range The f requency producing the maximum response in the range of 4 to 8 Hz.
Forcing Function Sinusoidal.
Spatial Distribution Uniform O
4.4.4-2 Revision 0
NEDO-21888 Downcomers Amplitude Versus Values given in Table 4.4.4-1 requency (Also shown in Figure 4.4.4-1).
Total Response Resulting responses from applying the amplitude at each frequency given in Table 4.4.4-1 are to be summed.
Spatial Distribution Uniform.
The oscillating loads are to be applied to the prevailing local static pres-sures at the various locations in the components for which they are specified.
The condensation oscillation pressure load specified for the downcomers should be used to calculate only the circumferential structural response (e.g., hnop stress) of the downccmeri not system responses to lateral, thrust, or other loads which are transmitted through the downcomers to other components. The downcomer laterr.1 loads that result from the unbalanced pressure between the vent header and the downcomer exit are included in the condensation oscilla-tion lateral load definition (see Section 4.4.3), and vent system thrust loads are included in Section 4.2.
4.4.4-3/4.4.4-4 Revision 0
NEDO-21888 Table 4.4.'4-1 CONDENSATION OSCILLATION PRESSURE AMPLITUDES IN THE D0'a'NCOMERS Frequency Frequency Range Amplitude
- Range Amplitude *
(Hz)
(PSI)
(Hz)
(PSI) 0-1 0.24 25-26 0.13 1-2 0.25 26-27 0.14 2-3 0.38 27-28 0.11 3-4 0.56 28-29 0.13 4-5 1.16 29-30 0.10 5-6 2.56 30-31 0.10 6-7 0.62 31-32 0.11 7-8 0.46 32-33 0.10 8-9 0.46 33-34 0.10 9-10 0.46 34-35 0.09 10-11 0.62 35-35 0.10 11-12 0.51 36-37 0.10 12-13 0.39 37-38 0.08 13-14 0.40 38-39 0.10 14-15 0.34 39-40 0.08 15-16 0.34 30-41 0.10 16-17 0.36 41-42 0.09 17-18 0.24 42-43 0.08 18-19 0.26 43-44 0.07 19-20 0.19 44-45 0.08 20-21 0.21 45-46 0.07 21-22 0.15 46-47 0.07 22-23 0.15 47-48 0.07 23-24 0.13 48-49 0.07 24-25 0.16 49-50 0.06
- Half range (= 1/2 peak to peak amplitude) 4.4.4-5 Revision 0
3.0 2.5 2.0 e
E
[
NOTE : THE AMPLITUDE SHOWN HERE REPHESENTS ONE HALF OF THE PE AK TOPE AK AMPLITUDE o
ha 1.5 Q
p 4
a D
O w
c w
3 E
A 1.0 CC 05 II l-I o u i s i r 0
5 10 15 20 25 30 35 40 45 50 FREQUENCY (lis) x C<
-a Fir,ure 4.4.4-1.
Condensation Oscillation Pressure Amplitudes in the o
Downcomers O
O O
NEDO-21833 REFERENCES FOR SECTION 4.4.4 4.4.4-1 bbrk I Containment Program - Full Scale Test Program Final Report, General Electric Company, Report No. NEDE-24539-P,'tarch 1979.
4.4.4-7/4.4.4-8 Revision 0
NEDO-21888 4.5 CHUGGING LOADS Chugging occurs during a postulated loss-of-coolant accident (LOCA) when the steam flow through the containment vent system falls below the rate necessary to maintain steady condensation at the downcomer exits.
The corresponding flow rates for chugging are less than those of the condensation oscillation phenomenon defined in Section 4.4.
During chugging steam bubbles form at the downconer exits, oscillate as they grow to a critical size (% downcomer diameter), and begin to collapse somewhat independently in time.
The result-ing load on the torus shell consists of a low frequency component which cor-responds to the oscillating bubbles at the downcomer exits as they grow, and higher frequency components corresponding to the collapsing bubble.s.
The loads specified for chugging are based on the FSTF tests.
In these tests a prototypical segment (one bay) of a Mark I torus and vent system were sub-jected to ten steam and liquid blowdowns simulating a range of LOCA's.
Duration of the chugging period varies with break size and is specified in Table 4.5.1-1.
Load combinations are given in the bar charts in Section 3.0.
Chugging is a low flow rate phenomenon which is not break size dependent.
Therefore, the loadings presented in the following subsection apply to all the break categories specified.
Loads on the structures within tne Mark I Containment system that may be induced by chugging are defined in the following subsections:
4.5.1 Torus shell loads 4.5.2 Loads on submerged structures 4.5.3 Lateral loads on downcomers 4.5.4 Vent system loads.
4.5-1/4.5-2 Revision 0
NEDO-21888 4.5.1 Torus Shell Loads During the chugging regime of a postulated LOCA, the chugging loads on the torus shell occur as a series of chug cycles, each of which can be organized into two elements:
the pre-chug and the post-chug portions. The pre-chug portion is described as follows.
Initially, as the steam-water interface enters the pool, a relatively low frequency (%7 Hz) pressure loading is observed on the torus shell. This low frequency corresponds to the frequency of oscillation of the interface and may persist for several of these low frequency cycles. The interface eventually becomes unstable and breaks up producing a rapid underpressure as the chug occurs.
In a chug cycic, the interface break-up occurs in several, but not all of the downcomers and takes place within a time interval usually not exceeding 0.1 seconds.
Data from the FSTF tests show that for a given chug cycle chuzs will seldom occur in all of the downcomers, and the chugs in the individual downcomers are never exactly synchronized.
The post chug portion of the chug cycle is a system response ("rfng out") to the rapid under pressure caused by the break up of the steam-water interface.
The post chug occupies approximately the same time span as the pre-chug in a given chug cycle.
The resulting response of the torus shell to a chug cycle is a low frequency pre-chug oscillation prior to the post-chug, followed by a higher frequency
" ring out" of the torus shell-pool water system in response to the impulsive underpressure after the post-chug begins.
In the process of the interface collapse, steam E.bbles may also be pinched off and collapse producing a very localized short duration pressure load.
Since these.short duration loads are not transmitted to the torus support system and remain very localized, they are not considered structurally significant and have not been included in these specifications.
4.5.1-3 Revision 0
NEDO-21888 4.5.1.1 Bases and Assumptions The basis for the torus shell chugging load definition is the chugging data obtained from the FSTF. The parameters which were varied in the FSTF tests include break sizes, downcomer submergence, initial pool temperature, blow-down fluid phase (liquid and vapor), and initial wetwell pressure.
The load-ing conditions established from the FSTF data and the use of conservative application techniques result in bounding chugging loads for all Mark I plants. The FSTF test matrix and test results are described in Reference 4.5.1-1.
The FSTF blowdowns which simulated the small steam break accidents produced the most severe chugging loads, and the data from these tests were used as a basis for the chugging load specifications. The FSTF blowdowns which simulated the large steam and liquid break accidents did not exhibit large chugging like behavior: the condensation process observed during these blowdowns appeared to be more continuous.
The onset times and durations specified in Table 4.5.1-1 are based on FSTF test results and an analytical study discussed in subsection 4.4.1.1.
The onset time of chugging for the DBA is based on the time at which condensation oscillations were determined to end (see Table 4.4.1-1). The onset time for the IBA is consistent with the onset time specified in the condensation oscillation subsection, since chugging loads are specified in lieu of conden-sation oscillations. The onset time for the SBA is consistent with the time required to purge the air from the drywell. The durations specified for all break sizes were determined by the analytical modeling methods discussed in subsection 4.4.1.1.
As noted earlier, the chugging load cycles observed in the FSTF tests are divided into a pre-chug and a post-chug portion.
Initially, a pre-chug low frequency sinusoidal oscillation appears on the torus wall.
This is related to oscillations of the steam-water interface. The frequency of these oscil-lations is relatively constant in all of the FSTF chugging data but the fre-quency is related to the vent acoustic frequency, which varies from O
4. 5.1 --
Revision 0
NEDO-21888 plant to plant.
This load specification contains a range of frequencies for the application of this initial load which is consistent with the range of vent system acoustic frequencies calculated for Mark I Containments (6.9 to 9.5 Hz).
The single frequency which produces the highest system response in this range is specified to be used for evaluation purposes.
The low frequency pre-chug oscillation is followed by an impulsive load as the rapid condensation event occurs.
In the FSTF data, the torus shell loads in this post-chug portion of the chug cycle appear as a combination of pressure oscillations at the natural frequencies of the torus shell combined with oscillations at the acoustic frequencies of the downcomers (%45 Hz).
This load specification contains a range of frequencies for the application of the post-chug load which is consistent with the range of downcomer acoustic fre-quencies for Mark I Containments (40 - 50 Hz).
Figure 4.5.1-1 shows a t;pical chug cycle with the pre-chug and post-chug portions identified.
To characterize the pre-chug and post-chug torus shell behavior, Power Spectral Density (PSD) analyses have been performed on the torus wall pressure traces.
These data have been used to develop the load specification for the torus shell.
The basis of the pressure amplitude specification is several segments of maximum pressure amplitude data from FSTF tests M1, M4, and M9.
Portions of individual tests are used for the two parts of the torus load specification as follows:
Pre-chug:
Test M9 Entire Test o
Post-chug:
Tests M1, 18 Maximum Pressure Amplitude Events e
M4 and M9 In order to apply the FSTF measured loads to plant unique geometries, an adjustment must be made to account for FS1 effects in the FSTF data.
Struc-tural response ef fects unique to the FSTF are accounted for by the specifica-tion of a rigid wall load (pressure oscillation amplitude as a function of frequency).
4. 5.1-5 Refision 0
NEDO-21888 The rigid wall chugging load for Mark I Containments is derived from the FSTF pre-chug and post-chug pressure response data using the following procedure:
a.
A finite element analytical model has been developed for the FSTF test facility. This model includes the torus shell support struc~
tures and the pool water.
Torus shell pressure load amplitudes relative to chug source pressures are determined as a function of frequency using this model of the FSTF.
This defines a relationship between pressure loads on the torus wall and a pressure source at the vent exits.
b.
The measured pre-chug and post-chug frequency response from the FSTF data is then used with this relationship to develop chugging source definition. The use of the coupled fluid-structure model of the FSTF allows the removal from the chug source those elements of the chug frequency spectrum that are contributed by the structural response of the FSTF.
O c.
Using the vent source chug definition, ' rigid wall' chug pressures are defined. The rigid wall loads are obtained from the pre-chug and post-chug vent source pressures by applying a spatial attenuation factor to transfer the pressures from the vents to the torus wall.
Ine resulting distribution around the wetted perimeter of the torus cross section is shown in Figure 4.5.1-2.
This pressure distribution is based on an incompressible flow analysis of the torus geometry.
For the pre-chug portion of the chug cycle, both symmetric and asymmetric loading specifications have been developed to conservatively account for any randomness in the chugging phenomena. The symmetric loading is based on the average of the FSTF data.
The asymmetric loading is based on both low and high amplitude chugging data conservatively distributed around the torus in order to maximize the asymmetric loading.
O 4.5.1-6 Revision 0
In order to bound the post-chug portion of the chug cycle, symmetric loads based on the peak average values from the FSTF data have been specified.
Asymmetric loads have not been specified since any azimuthal response would be governed by the asymmetric pre-chug low frequency load specification.
The frequency range in the pressure amplitude specification has been estab-lished to cover the range of frequencies observed in the FSTF tests and the range expected in all >brk I plants. The vent system acoustic natural fre-quency range for all Mark I plants is 6.9 to 9.5 Hz, and is dominant in the pre-chug portion of the chug cycle. The downcomer acoustic natural frequency range for all Mark I plants is 40 to 50 Hz, and is dominant in the post-chug portion of the chug cycle.
It is intended that the pre-chug rigid wall loading be applied to a flexible wall coupled fluid structural model for symmetric and non-symmetric loading cases. The pressure distribution for the latter is illustrated on Figure 4.5.1-3.
The post-chug rigid wall loading is to be applied to flexible wall coupled fluid-structural model for the symmetric case only.
For a comparison of the relative strengths of the two portions of a chug cycle, a PSD analysis of the specified loads has been perfermed.
Comparing the power 2
(ps1 value) from a PSD of a 1 Hz band width from the pre-chug specification to a corresponding value from the entire post-chug specification yields the relative proportions of 90% pre-chug and 10% post-chug.
The major assumptions in this definition of the torus chugging loads are:
The chugging behavior observed in the prototypical FSTF tests repre-a.
sents the chugging behavior of the Mark I Containment. This is justifiable since the chugging source strength is primarily con-trolled by thermodynamic and geometric conditions at the vent exit.
These conditions are very similar for all Mark I ple ms and were simulated in the FSTF tests.
Minor adjustments to account for vent 4.5.1-7 Revision 0
NEDO-21888 system acoustic frequency were made to conservatively represent all Mark I plants.
b.
Rigid wall chugging pressure loads can be derived from the flexible wall pressure measurements from the FSTF data by the use of the coupled fluid-structural model of the FSTF and incompressible flow theory.
c.
Incompressible flow theory can be used to determine the torus cross section wetted perimeter pressure distribution from pressure sources at the vent exits.
This is justifiable because of the good agreement between the analysis and FSTF data.
4.5.1.2 Load Definition Rigid wall torus chugging load definitions to be applied as a wall load to a structural model of the Mark 1 Containment are defined below.
Because the load definitions are derived as rigid wall forcing functions it is intended h
that the torus fluid be considered in the structural evaluations.
It also is intended that the pre-chug and post-chug analyses be steady state analyses.
Pre-Chug Load Amplitude and Circumferential Two cases shall be evaluated Distribution independently:
Svemetric Distribution t2.0 psi uniform axially along the torus centerline at bottom dead center.
Asymmetric Distribution Values shown in Figure 4.5.1-3.
Vertical Cross Section Linear Attenuation with submergence Distribution along the wetted perimeter as chown in Figure 4.5.1-2.
O 4.5 1-8 Revision 0
NEDO-21888 Frequency The frequency producing the maximum response in the range from 6.9 to 9.5 Hz.
Pre-Chug Cycle Duration 0.5 seconds every 1.4 seconds for the appropriate total duration defined in Table 4.5.1-1.
These loads are to be applied about the local static pressure at the appropri-ate times in the blowdown (see Table 4.5.1-1).
Post-Chug Load Amplitude Versus Frequency Values given in Table 4.5.1-2 (Also shown in Figure 4.5.1-4).
Total Response Resulting steady state responses from applying the amplitude at each fre-quency given in Table 4.5.1-2 are to be summed.
Spatial Distribution Uniform axially along the torus center-line.
Linear attenuation with submerg-ence along the wetted perimeter at the torus cross section as shown in Figure 4.3.1-2.
Post-Chug Cycle Duration 0.5 seconds every 1.4 seconds for the appropriate total duration defined in Table 4.5.1-1.
Pre-chug and post-chug evaluations need not be combined.
These loads are to be applied about the local static pressure at the appropriate times in the blowdown (see Table 4.5.1-1).
4.5.1-9/4.5.1-10 Revision 0
NED0-21888 Table 4.5.1-1 CHUGGING ONSET AND DURATIONS Onset Time Duration Break Size After Break After Onset DBA 35 seconds 30 seconds IBA 5 seconds 900 seconds SBA 300 seconds 900 seconds 4.5.1-11 Revision 0
NED0-21888 Table 4.5.1-2 POST-CHUG RIGID WALL PRESSURE AMPLtTUDES ON TORUS SHELL BOTTOM DEAD CENTER Frequency Amplitude
- Frequency Amplitude
- Range (Hz)
(PSI)
Range (Hz)
(PSI) 0-1 0.04 25-26 0.04 1-2 0.04 26-27 0.28 2-3 0.05 27-28 0.18 3-4 0.05 28-29 0.12 4-5 0.06 29-30 0.09 5-6 0.05 30-31 0.03 6-7 0.1 31-32 0.02 7-8 0.1 32-33 0.02 8-9 0.1 33-34 0.02 9-10 0.1 34-35 0.02 10-11 0.06 35-36 0.03 11-12 0.05 36-37 0.05 12-13 0.03 37-38 0.03 13-14 0.03 38-39 0.04 14-15 0.02 39-40 0.04 15-16 0.02 40-41 0.15 16-17 0.01 41-42 0.15 17-18 0.01 42-43 0.15 18-19 0.01 43-44 0.15 19-20 0.04 44-45 0.15 20-21 0.03 45-46 0.15 21-22 0.05 46-47 0.15 22-23 0.05 47-48 0.15 23-24 0.05 48-49 0.15 24-25 0.04 49-50 0.15
- Half range (= 1/2 peak to peak amplitude)
O 4.5.1-12 Revision 0
t onc xmoob coc S
T KEA LF CP YE CR r
A l
l e
h S
s N
u O
r i
o T
T R
O P
e h
G t
U H
n C
o T
SO e
P car T
e E
r L
u C
E s
Y M
s C
I e
T I
G r
U P
H C
e E
g N
a O
r ev A
g N
u O
i h
T C
RO l
P a
G c
U i
H p
C y
N E
T R
P A
1 1
5 4
M erug i
F wT?n$ mA
'n
=
'. o. ebu xC< pH0s O
LOCAL PRESSURE OSCILLATION AMPLITUDE A
=
WET WELL AIR SPACE A
MAX; MUM PRESSURE OSCILLATION AMPLITUDE MAX (AT TORUS BOTTOM DEAD CENTER)
A
.O
_V-F R E E SU R F AC E
[
A MAX y -
m SUPPH ESSION c
POOL O
Y b
7 5
~ \\
_ a e
<I<y x
\\
u a
^
=1 A
g MAX r
C
=
Figure 4.5.1-2.
Mark I Chugging-Torus Vertical Cross Sectional Distribution a
for Pressure Amplitude O
O O
180 270-
- 90 NOTE: THE AMPLITUDE SHOWN HERE REPRESENTS ONE-H ALF OF THE PEAK-TOPEAK AMPLITUDE
-0 0
0
.6 3
NOTE: HIGHEST VALUE IN BAY SHOULD BE y
APPLIED OVER THE ENTIRE BAY a
A j
2 p
t-E e
8 y
g aq r
w w
ca 1
w 2
C Ez o
1
{
0 8
$e l
l l
_t 0
90 180 270 3GO
& (degree)
W O<
r Te oa Figure 4.5.1-3.
Mark I Chugging - Torus Asymmetric Circumferential Distribution a
for Pressure Amplitude
NOTE: THE AMPLITUDE SHOWN HERE REPRESENTS ONE41ALF OF THE PE AK-TOPEAK AMPLITUDE 1
V
=
2 w
m m
Y B
A E'
co O
O 5
10 15 20 25 30 35 40 45 50 FREQUENCY (Hd N
E E
o Figure 4.5.1-4.
Post-Chug Rigid Wall Pressure Amplitudes ou Torus o
Shell llottom Dead Center 9
e
NEDO-21888 REFERENCES FOR SECTION 4.5.1 4.5.1-1 J.E. Torbeck, et cl., " Mark 1 Containment Program, Full Scale Test Program Final Report, " General Electric Company, Report No. NEDE-24539-P, March 1979.
4.5.1-17/4.5.1-13 Revision 0
NED0-21888 4.5.2 Loads on Submerged Structures due to Main Vent Chugging Steam chugging at the downcomers induces bulk water motion and therefore creates drag loads on structures submerged in the pool.
The submerged structure load definition method for chugging follows that used to predict drag forces caused by condensa* ion oscillations (Section i. 4.2) except that the source strength for chugging is proportional to the wall load measurement corresponding to the chugging regime.
4.5.2.1 Bases and Assumptions The basis of the flow model for chugging load definition derives from the work in References 4.5.2-1 and 4.5.2-2.
The major assumptions are summarized as follows:
a)
The total drag is the sum of standard and acceleration drags b)
The submerged structures are assumed to be rigid c)
Chugging is modeled as a fluid source located at the vent exit, which induces a bulk pool motion that results in drag forces on submerged structures d)
The actual flow field can be represented by a locally uniform flow field e)
The presence of boundaries is accounted for by using the method of images f)
The strength of the fluid source for chugging is determined from wall load measurements (Reference 4.5.2-3) given in Section 4.5.1 by using potential flow theory in incompressible fluid.
Revision 0 4.5.2-1
c NEDO-21888 s
c 4.5.2.2 Selection of Key Parameters for Load Evaluation Procedure This section outlines basic criteria for selecting key parameters to be used as input data to the model (Reference 4.5.2-1) so that chugging loads on submerged structures can be properly obtained.
The plant specific suppression pool geometry should be identified first as follows:
a.
Torus shell dimensions b.
Torus water depth c.
Location of submerged structure considered (Figure 4.3.8-1) d.
Locations of downcomers (Figures 4.3.8-1).
Then the chugging source infermation calculated from wall load measurements (Rcierence 4.5.2-3) is assigned to the downcomer exits.
The submerged structure is divided into the appropriate number of sections k
for a more precise calculation. The coordinates of each section center orientation of the structure axes and the structure projected cross-sectional area should be identified. The acceleration drag volume and the standard drag coefficient for each section of the submerged structure should be obtained from tables given in Reference 4.5.2-2.
With the input parameters identified and calculated, the submerged structure drag load model is used to evaluate the resultant transient drag forces on each structure element considered. Typical results are shown in Figures 4.5.2-1 and 4.5.2-2.
These figures show the chugging drag forces on an S/RV line (with T-Quencher arms) in the radial (X) and tangential (Z) directions as a function of time.
O 4.5.2-2 Revision 0
NEDO-21888 CHUGGING (PRE 4 HUG) FORCE ON S/RV LINE AT 8.0 Hz 270 1
180 P
90 I
/
e 5 0/Y E
?'
b C
a*
-90
-180
-270 I
I I
I O
0.1 02 0.3 0.4 0.5 TIME (set)
Figure 4.5.2-1.
Sample Predicted Time History of Total X-Forces on S/RV Line 4.5.2-3 Revision 0
CHUGGING (PRECHUG) FORCE ON S/RV LINE AT 8.0 Hz 120 80
+i
+
1 40 E
)/
E E
}f e
\\
T
-40 l
1 40 i
(&
I I
I I
-120 O
0.1 0.2 0.3 0.*
0.5 TIME (sec)
Figure 4.5.2-2.
Sample Predicted Time History of Total Z-Forces on S/RV Line 4.5.2-4 Revision 0
NEDO-21888 References for Section 4.5.2 4.5.2-1 Lasher, L.
E., " Analytical Model for Estimating Drag Forces on Rigid Submerged Structurer Caused by Condensation Oscillations and Chugging - Mark I Containments," General Electric Company, Report No. NEDE-25070-P, March, 1979.
4.5.2-2 Moody, F.
J.,
Chow, L.
C.,
and Lasher, L.
E.,
" Analytical Model for Estimating Drag Forces on Rigid Submerged Structures caused by LOCA and Safety Relief Valve Air Discharges, Ramshead," General Electric Company, Report No. NEDE-21471-P, September, 1977.
4.5.2-3 J. E. Torbeck, et al., " Mark I Containment Program, Full Scale Test Program Final Report," General Electric Company, Report No.
NEDE-24539-P, March, 1979.
4.5.2-5/4.5.2-6 Revision 0
NED0-21888 4.5.3 Lateral Loads on Downcomers During the chugging phase of a postulated Loss of Coolant Accident (LOCA),
vapor bubbles are formed at the downcomer end which collapse suddenly and inter-mittently to produce lateral loads on the downcomer. The procedure for defining the dynamic portion of this loading for a Design Basis Accident (DBA),
Intermediate Break Accident (IBA) and Small Break Accident (SBA) is presented in this section. The vent system thrust loads due to the quasi-static pres-surization of the containment and the fluid flow in the vent system during the chugging regime are reported separately in Section 4.2.
4.5.3.1 Bases and Assumptions The basis for the chugging lateral load definition is the data obtained from the instrumented downcomers of the Mark I Full Scale Test Facility (ISTF) during testing as described in Reference 4.5.3-1.
The load definition is developed for, and is directly applicable to, downconer pairs which are untied, i.e., not connected by lateral bracing.
Based on FSTF observations, this load definition is also applicable to braced downcomers.
The FSTF downcomer lateral loads are defined as Resultant Static Equivalent Loads (RSEL) which, when applied statically to the end of the downcomer, will reproduce at any given time the measured bending response near the downcomer/
vent header junction. The RSEL were obtained from the measured downcomer bending moments using conversion factors determined from static calibration tests performed on the downcomers, Reference 4.5.3-1.
These RSEL are, there-fore, an actual representation of the structural-hydrodynamic interaction which occurs during chugging.
The maximum design loads for individual plants are obtained by scaling the 95th percentile RSEL from the FSTF. The scaling factors are derived on the basis of a comparison of the dynamic characteristics of the downcomers of the individual plants and the FSTF.
4.5.3-1 Revision 0
NEDO-21888 For fatigue evaluation, the number of load reversals, i.e.,
RSEL reversals, during the chugging phase were counted and presented in the form of RSEL reversal histograms.
Since the chugging loads may be oriented in any lateral d irec t ion, these histograms were obtained for each of eight angular sectors around the downcomer end as shown in Figure 4.5.3-2.
The FSTF RSEL reversal histograms are scaled into a plant-unique set of histograms by first scaling the maximum RSEL reversal on the same basis as was done for the maximum design maximum RSEL reversal on the same basis as was done for the maximum design loads. Then the number of RSEL reversals are scaled by the ratio of the chugging duration specified for the plants to that of the FSTF.
For fatigue evaluation of the downcomer/ vent header junction, the plant-unique set of RSEL reversals must be converted to a set of stress reversals at the downcomer/ vent header junction which result from application of the RSEL reversals.
During the chugging phenomenon, lateral loads (i.e., RSEL) are imposed on the ends of the dwoncomers which are random in both magnitude and direction.
Because of the random nature of these chugging forces, there is a small but finite probability that the RSEL on two or more downcomers will align in the same direction to produce a resultant axial loading on the vent system which is larger in magnitude than the maximum RSEL observed to act on a single downcomer.
When two or more downcomers chug synchronously the event has been referred to as a pool chug. The FSTF RSEL were statistically combined to determine the probability of lateral forces on two or more downcomers exceeding specific values during pool chug synchronization.
The procedure for determining chugging lateral loads on downcomers includes the following assumptions:
a.
The vent header and downcomers can be considered rigid when compared to the flexible downcomer/ vent header junction. Therefore the down-comer responds dynamically in any given direction as a single degree of freedom system. This assumption has been verified by analytical studies (Reference 4.5.3-1).
O 4.5.3-2 Revision 0
NEDO-21888 b.
The lateral hydraulic load experienced by the FSTF downcomers during chugging is representative of the load that would be experienced by other Mark I downcomers for the same condition (Ref erence 4.5.3-1).
The FSTF RSEL teversal histograms obtained for the chugging phase of c.
the LOCA are representative of an actual plant behavior over the same duration (Ref erence 4.5.3-1).
Thus, for the fatigue evaluation of an individual plant's downcomer, the FSTF RSEL reversal histogram data must be scaled by the ratio of the chugging duration in the plant to the duration simulated in the FSTF.
d.
The loading on any given downcomer as determined f rom the measured downcomer bending response is predominantly due to a chug occurring on that downcomer. The effect that chugs on neighboring downcomers have on the measured bending response can be neglected.
This assump-tion applies to excitations transferred through the vent system and is based on test observations from the FSTF, Reference 4.5.3-1.
The structural-hydrodynamic phenomenon which takes place during multi-vent chugging is however, inherent in the FSTF data and the RSEL, thus determined, include this ef f ect.
The chugging loads are approximated as triangular pulse loads for the c.
purpose of dynamic scaling. This is based on observation of the FSTF chugging data (Reference 4.5.3-1).
f.
The load development procedure is applicable for loadings where the measured bending strains on the downcomer remain within the clastic range. These strains were observed to be well within the elastic range for all of the FSTF tests, Reference 4.5.3-1.
The procedure for determining the probability of the lateral loads on groups of downcomers exceeding specific values during pool chug synchronization requires the following assumptions in addition to those mentioned above.
4.5.3-3 Revision 0
NEDO-21888 g.
The maximum chugging load on an individual downcomer can occur in any arbitrary direction with equal probability, as observed from the FSTF test results (Ref erence 4. 5.3-1).
h.
All downcomers experience chugging in phase, i.e., the randomness of chugs in time is conservatively neglected.
1.
The magnitude and direction of the chugging load on a downcomer is statistically independent from all other downcomers, as observed from the FSTP test results (Reference 4.5.3-1).
4.5.3.2 Evaluation Procedure The loads associated with chugging obtained from the FSTF data will be scaled to determine the loads for other Mark I plant downcomers.
For each plant, the maximum downcomer design load, histograms of load reversals and the maximum vent system loading produced by synchronous chugging of the downcomers may be determined from the FSTF loads.
O The load used for determining the maximum design chugging load on an individual Mark I plant downcomer is based on the 95th percentile of peak chugging loads observed in the FSTF tests.
The maximum chugging design load P for an individual Mark I plant downcomer is obtained by scaling the 95th percentile FSTF load P The scaling law which incorporates the dynamic characteristics of the FSTF downcomers and the down-comers of an individual plant is given in Reference 4.5.3-1 as:
p[(DLF]/
DLF p
1 max where DLF and DLF are the dynamic load factors of the untied plant-unique and FSTF downcomer/ vent header structures, respectively.
O 4.5.3-4 Revision 0
NEDO-21888 The dynamic load factors depend upon the natural frequencies of the single-degree-of-freedom systems representing the downconer/ vent header structures.
Mark I plant downcomer geometries, the ratio of the dynamic load factors between the FSTF and a plant-unique downcomer is essentially independent of the direction of load application, Reference 4.5.3-1.
Therefore, in this pro-cedure the dynamic load factors used in the scaling relationship shall be deter-mined on the basis of the downcomer frequencies in the East-West direction. The natural frequency in the East-West direction can be determined once the rota-tional stiffness of the downcomer/ vent header junction and the mass moment of inertia of the downcomer in this principal direction have been ascertained.
The mass moment of inertia must include the added mass of the water for the submerged portion of the downcener.
A 95th percentile load of 1,997 lbs was determined from the FSTF chugging data.
The direction of this load was observed f rom the FSTF tes s to be random in nature. The maximum design load P therefore must be applied to the end of the plant-unique downcomers in a direction such as to maximize the stresses at the downcomer/ vent header junction.
For fatigue evaluation of the downcomers, the required stress reversals at the downcomer/ vent header junction can be obtained f rom the FSTF RSEL reversal histograms. Only RSEL reversals which are greater than a threshold of 5% of the maximum load range were counted for the histograms, since cycles of smaller magnitude do not contribute significantly to fatigue usagc.
The junction stress reversals are obtained by first scaling the FSTF RSEL reversals into a plant-unique set of RSEL reversals. This scaling procedure is a two-fold process.
First, tne maximum FSTF RSEL reversal is scaled on the same basis as was done for the maximum design loads. Then, the total number of reversals must be scaled by the ratio of the chugging duration specified for the plants to that of the FSTF. Chugging durations for the DBA, IBA, and SE\\ are specified in Table 4.5.1-1 of Section 4.5.1, and are applicable to all plants. Since iden-tical durations for the IBA and SBA are specified, only the DBA and IBA need be con-considered.
Based upon actual FSTF chugging durations, for a DBA fatigue evalua-tion of the downcomers the total number of reversals must be scaled by 0.0586.
For an IBA evaluation, this scaling factor is 1.76.
4.5.3.5 Revision 0
NEDO-21888 The plant-unique set of RSEL reversals may now be transformed into a set of stress reversals at the downcomer/ vent header junction. One method which can be used to transform the loads to stresses at a particular point A near the downcomer/ vent header intersection, Figure 4.5.3-3, is through a load-stress transformation matrix. This matrix may be obtained f rom a detailed static finite element analysis of the downcomer/ vent header structure.
First, a unit load F w uld be applied at the downcomer end in the North direction as shown.
N Figure 4.5.3-3 to obtain o
=VF3 where o is a representative stress measure tc be used in fatigue analysis.
Similarly, a unit load F would be applied at the downcomer end in the East direction to obtain a similar stress measure o b
c
=KF which upon combining stresses yields N
b " (bKE F
c=a
+
E For an arbitrary loading P_ having components P and P the resultant E,
stress is a.g where the transformation matrix K will, in general, be different for different locations of stress determination.
The total number of stress reversals at a location selected for fatigue evalu-ation is obtained by summing the stress reversals produced at that location by the chugging RSEL reversals in each sector.
O 4.5.3-6 Revision 0
SEDO-21883 For the case of pool chug synchronization, the probability of exceeding a given force magnitude at least once during multi-downcomer chugging is deter-The mined from probability of exceedance curves derived from the FSTF data.
load per FSTF downcomer can be obtained from the curves for various numbers L
of downcomers once an acceptable probability level of exceedance has been established. A probability of exceedance of 10-2 is conservative for this application. The resultant load in any direction due to pool chug synchro-nization may now be determined by multiplying the number of downcomers being considered by the load per downcomer.
It is necessary to scale this resultant FSTF load by the same scale factor used previously for scaling the 95th percentile FSTF chugging load to determine the resultant lateral load on a plant-unique Mark I vent system.
4.5.3-7/4.3.5-8 Revision 0
O O
O
NEDO-21888 SECTION OF DOWNCOMER/
HEADERSTRUCTURE i
l N
J 4
m N
0 0
0 0
337.5 22.5 3150 450 1
2 292.50 0
3 67.5 4
E 0
0 270 90 5
0 247.50 6
112.5 7
1350
- 225, 0
202.5 1800 Figure 4.5.3-1.
Sectors Used to Define Directions of Lateral Loads on Downcomer's End 4.5.3-9 Revision 0
zbOaN"gm*
W 8
3 P
9'
=
o t
M L +
8 "0
1 a
N 1
r o
t 2
h 0
8 0
l 4
7 i'
2
't i
3 i
D 0
7 8
0 1
s 3
l 0
i, 6
r e
6
)
v T
e N
t E
I C
i 3
l El E
0 P
6
(
S L
E 8
S i
t g
f s
i 2
.,l 9
g 1
lt 0
!n 4
C 1
2 1
0 i
Ol 52 l
l 0
t 3
i i
l i
9 4
r t
S I
8 D
0 1
0 2
1 2
4 2
3 B
I S
04 4
5 0
e 1
r 6
i n
4 t,g 2
1 i
0 0
0 0
0 0
0 O
0 0
0 0
0 0
0
(
7 6
5 4
3 2
1 mg 2.bEuOc E32 o
e 4
u
- Y-
. [1O3 O
.IEDO-21SS8 f
0 f\\
^
y a
N f
\\
\\-
\\
smW F E A,
-9 Figure 4.5.3-3.
- otation Used for Transforming RSEL Reversals into Stress Reversals at a Fatigue Evaluatian Location A 4.5.3-11/4.5.3-12 Revi,ian 0
NEDO-21888 REFERENCES FOR SECTIO; 4.5.3 4.5.3-1
" Mark I Containment Program, Development of Downcomer Lateral Lcads from Full Scale Test Facility Data - Task Number 7.3.2",
General Electric Company, Report No. !1EDE-24537, March 1979.
4.5.3-13/4.5.3-14 Revision 0
NEDO-21888 4.5.4 Vent System Loads Pressure loadings are experienced by the vent system as a result of chugging.
These vent system loads can b'e separated into the following three components:
a.
A gross vent system pressure oscillation consisting of pressuri-zation during the pre-chug portion and depressurization during the post-chug portion of each chug cycle, b.
An acoustic vent system pressure oscillation which is excited as a result of the pressurization and depressurization of the vent system.
c.
An acoustic downcomer pressure oscillation which is excited as a result of the rapid depressurization at the downcomer exits.
The first component of pressure loading is applied over a relatively long loading cycle (1 to 2 seconds) which corresponds to the time between chug cycles. The second and third pressure load components are related to the acoustic response frequencies in the vent system and downcomer, and are defined as a periodic load with components at the acoustic frequencies of the vent system (including the downcomers) and of the downcomers themselves.
4.5.4.1 Bases and Assumptions The Mark I vent system consists of the main vents, the vent header and the downcomers. The vent system load definition is based on data obtained from the Mark I FSTF. The test facility, test matrix and test results are described in Reference 4.5.4-1.
Two tests were selected for the bases of the vent system chugging loads. Test M1 with the small steam break is the basis for both the 0.7 liz gross vent system pressure oscillation and the 40 to 50 Hz acoustic downcomer pressure oscillation load specifications.
Test M4 with the small steam break and overpressure is the basis for the 4.5.4-1 Revision 0
NEDO-21888 6.9 to 9.5 Hz acoustic vent system pressure oscillation load specification.
These tests produced the largest pressure oscillation amplitudes for these components.
The frequency range basis for the acoustic vent system and downcomer pressure oscillations are as defined in subsection 4.5.1.1.
The frequency specified for the gross vent system pressure oscillation is based on the mean of the FSTF chugging data.
The characteristics of the total vent system pressure load and the role of each of the pressure load components exhibit some change over the period of the blowdown during which chugging occurs. Early in the chugging period, both the acoustic downcomer pressure oscillation and the gross vent system pressure oscillation contributions are relatively small compared to the contribution of the acoustic vent system pressure oscillations which are dominant. Later in the chug cycle, a higher frequency post-chug acoustic downcomer pressure oscillation begins to dominate the downcomer pressure loads. The contribution of the acoustic vent system pressure oscillation h
lessens during this portion of the blowdown. The relative contribution of the gross vent system pressure oscillation generally becomes more apparent in the vent header and main vents.
The vent system data simultaneously contained contributions of some or all of the individual load components.
In order to simplify the specification, the chugging data have been analyzed to extract the maximum of each load-ing component so each vent system pressure load component can be speci-fled as if it were acting alone. This was done by selecting from the chugging data those time segments during the blowdown where each loading component was clearly dominant. The peak load is defined for each as if it alone were contributing to the observed pressure loads. The frequency range specified for these loads is based on a linear acoustic analysis of the spectrum of Mark I vent system and downcomer geometries which included all of the varia-tions in the Mark I Containment vent system designs. The result is a conserva-tive definition of the vent pressure loads.
O 4.5.4-2 Revision 0
NEDO-21888 The major assumptions made in this definition of the vent system pressure loads are:
a.
The vent system pressure loads due to chugging observed in the prototypical FSTF tests represent the pressure loads that would be expected during chugging in a Mark I Containment. This is justi-fiable because the source strength for chugging is mainly controlled by thermodynamic and geometric conditions at the downcomer exits.
These conditions are very similar for all Mark I plants and were simulated in the FSTF tests.
b.
Vent system acoustic response frequencies for the various Mark I vent system and downconer geometries can be defined by a linear acoustic analysis. This is justified by the good agreement between FS1F analysis and data.
c.
The response of the vent system to the load definitions given here will remain linear to permit the definition of bounding single frequency components of the vent system pressure loads. This is justified by the rele.eively low magnitude of the loading components.
4.5.4.2 Load Definition The vent system chugging load definition is summarized in Table 4.5.4-1.
For the duration of application refer to Table 4.5.1-1.
Because of the manner in which they were develooed, these loads are to be applied individually, not combined with each other.
They are to be applied about the local pressures at the appropriate times in the blowdown depending on the size of the break (Table 4.5-1).
The chugging load specified for the downcomers should be used to calculate only the circumferential structural response (e.g.,
hoop stress) of the downcomer, not system responses to lateral, thrust, or other loads which are 4.5.4-3 Revision 0
NEDO-21888 transmitted through the downcomers to other components. The lateral loads that result from the unbalanced pressure between the vent header and down-comer exits are included in the chugging lateral load definition (see Section 4.5.3), and vent system thrust loads are included in Section 4.2.
For other loads which may occur in combination with the chugging loads see the bar charts in Section 3.0.
O O
4.5.4-4 Revision 0
NEDO-21888 Table 4.5.4-1 VENT SYSTEM LOAD AMPLITUDES AND FREQUENCIES FOR CHUGGING Amplitude (psi)
Frequency Main Vent Load Type (Hz)
Vents lleader Downcomers Gross Vent System Use wave form in 2.5 22.5 25.0 Pressure Oscillation Figure 4.5.4-1 (0.7 Hz)
Acapstic Vent System Sinusoidal with 2.5 23.0 t3.5 Pressure Oscillation frequency varying between 6.9 to 9.5 Hz Acoustic Downcomer Sinusoidal with N/A N/A 213.0 Pressure Oscillation frequency varying between 40 to 50 Hz 4.5.4-5 Revision 0
TH E FREOUENCY IS 0.7 Hz (r = 1.4 sec) h F ALL TIME 0 25r A
y E
3 Z
y M
AMPLITUDE d
g C
a s
u i'
?
2 4
1 PERIOD r ->
eRISE TIME 0.75r +
TIME P3 e<
r-
'fs H.o
- s C
Figur-4.5.4-1.
Chugging Wave Form for Gross Vent System l'ressure Oscillation Load 9
e
NEDO-21888 REFERENCES FOR SECTION 4.5.4 4.5.4-1 J. E. Torbeck, et al.,
">1 ark I Containment Program, Full Scale Test Program Final Report," General Electric Company, Report No. NEDE-24539-P, 51 arch 1979.
4.5.4-7/4.5.4-8 Revision 0
NEDO-21888 6.4 DOWNCOMER AIR CLEARING LATERAL LOADS During the initial phase of a postulated Loss of Coolant Accident (LOCA),
the rapid clearing of air through the downcomers causes them to be sub-jected to lateral loads. The predominant load in this phase has been reported in Section 4.2 under Vent System Thrust Loads. Additionally, a minor dynamic air clearing lateral load was observed in the Mark I Full Scale Test Facility (FSTF) data. This secondary load is bounded by the condensation oscillation lateral load defined in Section 4.4.3, and therefore should not have to be considered separately. The procedure for defining it, however, is presented in this section for completeness.
6.4.1 Bases and Assumptions The bases and assumptions used for defining air clearing lateral loads on downcomers are the same as those given in Section 4.4.3.
6.4.2 Evaluation Procedure A maximum downcomer lateral load of 2,172 lbs was determined from the dynamic portion of the FSTF Design Basis Accident (DBA) air clearing data to occur within 222-1/2* of the East-West direction, i.e.,
in the plane of the downcomer pairs as identified by sectors 4 and 5 in Figure 6.4-1.
The scaling procedure used to obtain the maximum design load for an individual Mark I plant downcomer is the same as that described in Section 4.4.3.
6.4-1/6.4-2 Revision 0
NEDO-21888 SECTION OF DOWNCOMER/
HEADER STRUCTURE l
N J
\\
4 m
N A
6 00 337.50 22.50 3150 450 1
2 292.5 3
67.50 4
2700 E
900 y
5 247.So 6
112.50 7
- 225, 135 202.5 180 Figure 6.4-1.
Sectors Used to Define Directions of Lateral Loads on Downcomer's End 6.4-3/6.4-4 E""i*i " 0