ML20053A505
| ML20053A505 | |
| Person / Time | |
|---|---|
| Site: | Big Rock Point File:Consumers Energy icon.png |
| Issue date: | 05/21/1982 |
| From: | Sacramo R CONSUMERS ENERGY CO. (FORMERLY CONSUMERS POWER CO.) |
| To: | |
| Shared Package | |
| ML20053A496 | List: |
| References | |
| ISSUANCES-OLA, NUDOCS 8205260160 | |
| Download: ML20053A505 (42) | |
Text
_.
's 4
-4 UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION BEFORE THE ATOMIC SAFETY AND LICENSING BOARD 1
I In the Matter of
)
)
Docket No. 50-155-OLA 4
l CONSUMERS POWER COMPANY
)
(Spent Fuel Pool
)
Modification) i (Big Rock Point Nuclear Power Plant) )
i TESTIMONY OF RAYMOND F.
SACRAMO CONCERNING l
CHRISTA-MARIA CONTENTION 8 AND O'NEILL CONTENTION IIIE-2 i
My name is Raymond F.
Sacramo.
I reside at 17204 l
Chiswell Road, Poolesville, M.tryland.
I have been employed
).
with NUS Corporation, an engineering consulting firm in Gaithersburg, Maryland, since October 1, 1977.
A resume which i
f describes my background and qualifications is attached, i
i l
INTRODUCTION i
Christa-Maria Contention 8 and O'Neill Contention IIIE-2 were 1
consolidated and rewritten by the Board to read as follows:
The occurrence of an accident similar to TMI-2 which would i
prevent ingress to the contain-t ment building for an extended period of time would render it i
impossible to maintain the i
expanded spent fuel pool in a l
safe condition and would result i
in a significantly greater risk l
to the public health and safety
)
than would be the case if the
{
increased storage were not allowed.
i i
i 8205 2s o WO
s 6
4 i
l !
1 1
I My testimony summarizes the results of an analysis, i
]
attached as Exhibit A, which shows that at the maximum boiling 1
temperature of 237' that could occur in the Big Rock spent fuel pool the structural integrity of the pool concrete, liner i
j and storage racks is adequate.
l In its Memorandum and Order of February 19, 1982, i
the Licensing Board questioned the accuracy of the 237 F l
maximum boiling temperature of the Big Rock Point spent fuel j
pool and the effects of a 247 F maximum temperature on pool structural strength.
The Board also questioned the integrity l
of the pool when subjected to point loading from the storage racks.
Specifically,-the Licensing Board adopted the argument advanced by the Intervenors concerning Christa-Maria Conten-e tion 8 and O'Neill Contention IIIE-2, which can be restated as follows:
I j
Is the concrete in the spent fuel storage i
pool strong enough to (i) withstand the j
global effects of a loss in overall l
strength due to being subjected to 247'F l
and (ii) resist a temperature of 247'F and i
point loading from the storage racks.
I i
My testimony addresses both of these issues.
In addition, I summarize the result of a study which evaluates the effects of increased concrete degradation in the Big Rock Point spent fuel pool.
s
.h i 3 i
'i MAXIMUM POOL BOILING TEMPERATURE 1
The accuracy of the 237*F maximum boiling tempera-ture which could occur in the Big Rock Point Spent Fuel Pool was verified in Dr. Prelewicz's testimony filed on May 10, i
j 1982, concerning pool thermal hydraulic conditions, i
l The results of detailed calculations on fuel pool natural circulation cooling summarized in Dr. Prelewicz's testimony showed that the water entering the bottom of the fuel racks would be approximately 212*F.
This rc3 ult verified l
l the initial assumption of the previous analysis performed by l
Dr. Prelewicz.
On the basis of this assumption, Dr. Prelewicz 4
i j
showed that a maximum boiling temperature of 237'F would occur
{
at.276 inches below the top of a fuel bundle stored in the racks.
Therefore the maximum temperature of 247* referred to q
f by the Board is not attainable and should not be used in the structural integrity analysis of the spent fuel pool concrete, liner and. storage racks.
In fact, the use of 237'F is conservative for analyzing the concrete strength and the j
thermal gradient of the pool walls and floor and the strength l
of the pool liner since, due to the natural circulation, the coolant in contact with the liner would be closer to 212*F.
Nonetheless, the value of 237'F was used as the base tempera-1 l
ture for establishing the effects of coolant boiling on pool
- strength, i
i
t 4
1
! 4 4
SPENT FUEL POOL STRUCTURAL ANALYSIS I
I The structural integrity of the spent fuel pool I
concrete, liner and storage racks were investigated for the i
loading condition which would develop during the accident scenario defined in the testimony of David P.
Blanchard concerning Christa-Maria Contention 8 and O'Neill Contention j
l IIIE-2.
This loading condition considers that the pool is filled to its proposed capacity of 441 assemblies during a boiling condition of the coolant which exists for an extended l
period.
In order to determine the effects of this accident 1
l scenario, the following areas were addressed:
3 o
Effects of the accident environment on I
degrading the pool concrete strength.
o Strength load-carrying capacity definition of the pool geometry.
o Thermal gradient, hydrostatic and deadweight loading conditions across the various walls and floor during the accident.
l' o
Accident condition bending moment and shear forces imposed on the walls and floor.
I o
Design margins under the resultant accident loading conditions.
t o
Pool liner strength and stress conditions during the accident.
o Fuel rack strength considerations during
[
coolant boiling.
l l
l I
f
.. EFFECTS OF THE ACCIDENT ENVIRONMENT ON DEGRADING POOL CONCRETE STRENGTH Special consideration was initially given to the potential concrete strength degrading which could result from the accident environment elevated temperature of 237*F.
Technical papers and test data on the topic of concrete strength degration were reviewed.
In this review, four major strength properties were investigated: compressive, bond, tension, and shear strength.
Compressive and bond strength both relate to the bending moment resistance of the pool walls and floor.
More specifically, compressive strength is a governing property of the concrete while bond strength relates to the ability of the steel reinforcement bars to stay attached to the concrete while resisting bending moment loads.
Adequate bond strength is established by assuring that a given length of reinforce-ment bar (known as the development length) is provided on either side of where the bending moment is being applied.
Shear and tension strength both relate to the 1
ability of the walls and floor to resist transverse forces.
More specifically, shear relates to a failure condition where cracks would propagate vertically through the walls or floor thickness.
This type of failure is considered to be a true shear failure.
A tension failure on the other hand would 1
. tepresent a failure condition where cracks would propagate at a 45* angle through the walls or floor thickness.
This type of failure is considered a diagonal tension mode of failure.
If shear failure were to occur in the pool walls and floor, the condition of failure would lie somewhere between true shear and diagonal tension.
American Concrete Institute standards, such as References 1 and 2, specify varying levels of shear capacity depending upon the application of the corcrete and reinforcement.
Concrete Degradation Literature and Test Data Review Several papers published on the effects of tempera-tures on concrete compressive strength were reviewed. One example cited in the Exhibit A analysis is G.
N.
Freskakis, R.
C.
- Burrow, E.
B.
Debbas, " Strength Propcrties of Concrete at Elevated Temperatures," April 2, 1979 (Reference 3).
This paper presents a summary of published test data on concrete strength properties at elevated temperatures.
The summary presented in this paper, which was based on over fourteen different aggregates and mixtures, provides both upper and lower bounds on the reduction in 28-day concrete strength at temperatures up to 1600 F.
The 28-day aging period has been established as a standard by the American Concrete Institute.
l l
l
. From this published data the relevant data for the maximum boiling temperature condition of 237*F'was selected.
For all of the aggregates tested, the data shows a range from no reduction in strength to a maximam reduction of 25% at 237 F.
Concrete compressive strength does, however, in-crease during aging beyond 28 days.
A test program carried out at Oak Ridge National Laboratory, which is published in
" Uniaxial Compressive Strengths of Concrete for Temperature Reaching 1033 K," Nuclear Engineering and Design", 45 (1978),
pp. 439-448 (Reference 4) and also summarized in the Reference 3 report, considered such aging conditions.
Eight-to nineteen-month-old concrete cylinders of a limestone aggregate mix were tested.
(As indicated below limestone aggregate was l
judged to be used in the Big Rock Point spent fuel pool.)
The l
cylinders were increased to their tested temperature at a rate I
of 30*F/hr. and maintained at the test temperature for 14 days.
The compressive strength at failure for the various test temperatures was compared to the 28-day compressive strength at normal temperature (approximately 70*F).
The results showed that the compressive strength at 237*F was approximately 120-130% of the 28-day-old strength at normal temperature.
Consequently, concrete aging compensates for the effects of temperature degradation at this temperature level.
-~g4,-m-,--%.-,
_nm.,,,,7%
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y,-m--,m
,r--,wnw.._..,,r--.,..
p,..._m-.w-,r.w_m,-.,m,_,_,
g.-,-.--_
cp-.p.
+
1 ;
l The remaining temperature consideration is the j
prolcnged effect of temperature on concrete strength.
Even though concrete strength reduces as the temperature condition prolongs, it stabilizes under steady-state conditions.
i Steady-state conditions will occur approximately eight days after the coolant boiling condition begins.
The test results j
are, therefore, reasonable for the accident environment as j
postulated by Mr. Blanchard.
I Irradiation effects tend to degrade concrete com-1 I
pressive strength.
At the Big Rock Point Plant, spent fuel is l
randomly placed in the fuel pool racks.
The following analy-i sis is based on placing new spent fuel in one location, i.e.,
I the center of the pool floor.
This assumption is conservative when compared to the actual method of spent fuel storage 1
described above for the Big Rock Point Plant.
4
{
The irradiation would be of two types:
that pro-I duced by neutrons and that produced by gamma rays.
Neutron j
fluxes emanating from spent fuel are normally low.
These j
fluxes are further reduced due to the neutron absorbing l
capability of the water.
A conservative total neutron ex-
)
,posure estimate for this center pool floor area after a 11 2
{
service life of 40 years would be on the order of 10 n/cm.
Test results contained in an Oak Ridge National Laboratory report (Reference 5), showed that the compressive strength i
I i
1,
[
's t
i properties of a reactor shield made from Barytes-Haydite concrete did not degrade at such irradiation levels.
Serpentine concrete specimen test results summarized in
" Behavior of Special Shielding Concretes and of Their Consti-tuents Under Neutron Irradiation," (Reference 13), showed i
19 2
irradiation dosage levels of 10 n/cm produced only slight changes in concrete compressive strength.
An upper-bound estimate of gamma irradiation for a 10 service life of 40 years is 10 Rads.
Test results sum-
'l marized in " Gamma Radiated Damage of Structural Concrete l
Immersed'in Water," (Reference 14), showed that specimens receiving such gamma ray irradiation levels on the average sustaine'd no degradation *.
The upper-bound loss in compressive 1
strength was only 10%.
No appreciable strength reduction in l
concrete is, therefore, expected to occur from the effects of l
^
either neutron or gamma irradiation over c 40-year service i
life of the pool.
In summary, based on the above data, the pool concrete strength at Big Rock Point during and after the accident hypothesized by Mr. Blanchard is considered to remain above the 28-day compressive strength at normal temperature (approximately 70 F).
Nevertheless, for the purpose of conservatism, a lower-bound 80%, 28-day compressive strength level at 237*F was used as the design criterion.
Independent
3
. l support for the compressive strength reduction cited above is found in an Oak Ridge National Laboratory report and an American Concrete Institute Special Publication technical
]
paper, References 5 and 6.
1 The second strength property investigated was shear strength, which, as stated previously, would include test data on both true shear and diagonal tension modes of failure.
Shear strength testing performed at the Oak Ridge National Laboratories (Reference 7) indicat5d that a reduction of 25%
l for limestone aggregate concrete could be anticipated at a temperature of 237*F.
A range of tension strength reduction is summarized from an Oak Ridge National Laboratory report (Reference 5) and an American Concrete Institute Special i
Publication (Reference 8).
In Reference 5, tension strength testing was cited which concluded that a 4-7% reduction would occur over the temperature range of 200*C (392 F) to 250 C (482 F).
Test results given in Reference 8 show tension strength reductions of 10% and 20% respectively for silica and limestone aggregate concretes at temperatures of 237*F.
It was also concluded in Reference 8 that compressive and tension strength reductions were similar for _ilica concrete while the tension strength reduction of limestone concrete was larger than the compressive 3trength reduction.
Based on the shear and tension strength reductions illustrated above the maximum
_ _ _ _ _ _ _ _ _. _ _ _ _... -. _ _ -.. _ _ _ _ _.. ~,. _
__ =-
i I !
25% reduction from the Oak Ridge National Laboratory testing was used in the Exhibit A report.
Accordingly, the minimum i
shear strength expression from the Reference 2 American Concrete Institute Standard was used to derive shear strength capacities for the walls and floor, and only 75% of these values were used as an acceptable limit.
The final strength property investigated was bond strength.
As stated previously, bonding represents the strength which assures that the reinforcement bars stay 1
attached to the concrete while the concrete and reinforcement 4
are rcsisting bending moment loads.
Bond strength testing performed by Oak Ridge National Laboratory (References 7 and
- 9) indicated a 15% reduction in strength for reinforcement bars pulli:tg away or becoming detached from limestone aggre-gate concrete specimens.
These tests were performed in accordance with the Standard test method contained in j
" Standard Task Method for Comparing Concretes on the Basis of 4
the Bond Developed with Reinforcing Steel," (Reference 10).
Results of the testing summarized in the Reference 9 Oak Ridge i
National Laboratory report led to a recommendation that rein-forcement bar development length requirements of the Reference 2 Standard should be increased by 15% for the design of
)
reinforced foundations operating under temperatures reaching 300'F.
Bond strength testing performed in Japan, Reference 8, indicated that reinforcement bars in silica aggregate concrete I
4 4
(
. P l
chowed bond strength reductions of 50 to 60% for temperatures varying from 212 F to 572*F.
l In the Exhibit A analysis, the Oak Ridge National 1
Laboratory testing was considered to be more representative of the pool conditions for two reasons.
(1) Based on my review of Walter H.
Flood and Company Test Reports for Bechtel Corporation, cited at page 4-1 of Exhibit B, it is my l
judgement that limestone aggregate was used for the construc-tion of the Big Rock pool.
(2) Because of the primary emphasis on standard testing methods used and the recommenda-tions for application of the data to American Concrete Insti-tute Standard banding strength requirements.
This recommen-l dation was applied in the Exhibit A analysis by checking critical pool locations to assure that sufficient reinforce-ment development length existed.
i STRENGTH LOAD CARRYING CAPACITY DEFINITION OF THE POOL GEOMETRY The next step in the analysis requires the discus-sion of the load-carrying capacities which will resist the resultant bending moment and shear forces.
In the case of bending moment loads, it should be noted that although the concrete compressive strength does
- _ _ _ _. ~. _ _ _ - _ _ -. _.. _ _ _ _ _. _ _ _ _ _ _ _ _ _, _ _ _ _ _ _ _ _. _ _ _. _
. - = - -
. affect the total load-carrying capacities of the walls and floor, the major factor is the yield strength of the c
reinforcement or rebar.
The spent fuel pool walls and floor reinforcement which has a yield strength of S
= 40,000 psi at y
70*F would be reduced to 37,175 psi at 237'F.
Steel yield t
etrength and other material properties at temperature were taken from the American Society of Mechanical Engineers Boiler and Pressure Vessel Code, Reference 11.
The bending moment load-carrying capacities, using the strength design method, were developed for the various pool walls and floor locations considered.
The strength design method, a method used by American Concrete Institute, allowed the stress of the concrete reinforcement, under tensile loading, to reach 90% of the reinforcement yield strength.
Under bending moment loading conditions, reinforced J
l concrete would fail when a load slightly larger than that i
which causes the tension reinforcement to exceed the yield l
l strength is applied.
A summary of the condition of failure is 1
[
as follows.
When a slight additional load is applicd beyond the load causing yield in the reinforced steel, the steel l
stretches a considerable amount.
The increased steel i
(
l deformation in turn causes the neutral axis of the concrete i
cross-section to shift toward the compressive side.
The shift in neutral axis reduces the area under compression and thereby l
l
{
. increases the unit compressive stress.
This process of addi-tional steel deformation, shift in neutral axis, and compres-sive area reduction continues until the concrete fails. When steel does not have a sharp yield point, the load required for failure will increase, making calculations based on 90% of yield strength conservative.
It should also be.noted that the ultimate strength design method assumes that the tension concrete has already failed and offers no resistance *to the bending loads.
Throughout the spent fuel pool, the concrete is double reinforced near the pool side and outside wall sur-faces.
The reinforcement or rebar is also placed vertically and horizontally through the walls and in both horizontal directions for the floor.
Consequently, as many as four bending moment strength capacities per wall and floor location were developed.
Two load-carrying capacities were developed l
for both the horizontal and vertical reinforcement.
The first I
set of two load-carrying capacities were based on the final resultant moments causing tension in the pool side reinforce-ment.
The second set considered tension in the outside wall reinforcement.
The shear strength capacities were based on the area of concrete and have one value for a given wall and floor location.
1
i LOADING CONDITIONS ACROSS THE VARIOUS WALLS AND FLOOR 9
Having established the load-carrying capacities of the Big Rock Point spent fuel pool, I determined the loading conditions imposed on the pool.
Four critical or bounding cross-sections were investigated.
The topmost portion of the i
east wall which is 2'-0"- thick represents the thinnest wall cross section of the pool.
This wall cross section is critical since it offers the least load carrying capacity.
The worst wall temperature gradient occurs at 20 hours2.314815e-4 days <br />0.00556 hours <br />3.306878e-5 weeks <br />7.61e-6 months <br /> into l
the coolant boiling transient.
After 20 hours2.314815e-4 days <br />0.00556 hours <br />3.306878e-5 weeks <br />7.61e-6 months <br />, the temperature of the outside wall increases while the inside wall surface remains constant at the boiling temperature of the coolant.
The 6'-9" thick north wall would develop the i
largest bending moments due to the thermal transient.
- Again, i
after 100 hours0.00116 days <br />0.0278 hours <br />1.653439e-4 weeks <br />3.805e-5 months <br /> the temperature of the outside wall increases while the inside wall surface remains constant at the boiling I
temperature of the coolant.
The south wall, below elevation 624'-0", varies from 5'-9" thick on the east side to 3'-6" on the west side.
The 3'-6" portion of the wall is the thinnest section at the lower elevation and maintains the same wall thickness down to the ground floor elevation.
This section of wall would provide the least load carrying capacity for i
e e-~$
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x-
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m.--
+ -
w-+-
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w'
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. hydrostatic loading conditions.
The maximum temperature gradient occurs at 50 hours5.787037e-4 days <br />0.0139 hours <br />8.267196e-5 weeks <br />1.9025e-5 months <br /> into the thermal transient.
The 6'-0" thick pool floor supports the deadweight loads of the concrete, coolant, racks and contained fuel assemblies.
The maximum temperature gradient occurs at 100 hours0.00116 days <br />0.0278 hours <br />1.653439e-4 weeks <br />3.805e-5 months <br /> into the thermal transient.
Thermal loading resulting from these temperature gradients were developed based on the input discussed in Dr. Prelewicz's affidavit of September 25, 1981 and they are defined by the following equations:
2'-0" thick east wall (z /t) 2. 7 = 128 (z /t) 2. 7 AT AT
=
g 6'-9" thick north wall (z /t) 2. 4 = 150 (z/t) 2.4 AT AT
=
g 3'-6" thick south wall AT (z /t) 2.0 = 142 (z /t) 2. 0 l
AT
=
g l
6'-0" thick floor (z /t) 2. 5 (z /t) 2. 2 = 150 AT AT
=
z where AT
= the difference in temperature between the poolside l
and outside surfaces of the wall or floor; l
l t
l
o AT
= the difference in temperature between the outside z
surface and any location z through the wall or floor; z
= distance from the outside wall to any location through the wall or floor; t
= the thickness of the wall or floor.
A concrete surface subjected to a curvilinear gradient, as those defined by the above equations, will develop a moment due to the gradient.
Also, a resisting moment caused by forces (PT) acting over a distance (L) results from the restraints at the corners.
The restraints are the other walls and floor of the pool.
A further effect is the possible cracking of the concrete in tension.
This cracking results from an increase in the stresses in the tension side reinforcement due to a shift in location (L) of the resisting forces PT*
The tension stresses in the reinforcing steel which would develop due to thermal gradients and the equivalent moments required to develop the stress were based on analyti-cal methods contained in the Reference 12 Standard.
For the mechanical loading conditions (deatseight and hydrostatic), the analysis of fuel pool walls and floor was performed using thin plate theory.
Thin plate theory is based on shear deflections being small compared to the
..exural deflections.
This condition is normally maintained by limiting the thickness of the plate to about one-quarter of the least transverse dimension.
For reinforced concrete, the flexural deflection is controlled by the placement of the rebar while the shear deflection is governed by the concrete.
In all cases, the location of the reinforcement, nearest to the outside surfaces of the wall, is within one-quarter of the least transverse dimension.
For reinforced concrete, the flexural deflection is controlled by the placement of the rebar while the shear deflection is governed by the concrete.
In all cases, the location of the reinforcement, nearest to the outside surfaces of th'e wall, is within one-quarter of the least transverse dimension.
Also, since the reinforcement is extremely ductile as compared to the concrete, one can assume that the shear deflections are small compared to the flexural deflections.
The walls and floor will also act as integral members with each other in distributing the loads.
Loads would be distributed in a proportional manner to the moments of inertia of the various walls and floor comprising the pool.
A simplifying and conservative assumption was made that all walls and floor would act independently.
This assumption was verified as being conservative by comparing resultant moments and shears obtained in this manner to those obtained when assuming integral behavior of walls and floor.
1 J
7
. The pool walls were analyzed based on three sides fixed and the top side free.
The pool floor was analyzed on the basis of simply supported as well as fixed sides to obtain a bound on the moment and shear values.
Hydrostatic and deadweight loads were based on the pool filled to the maximum elevation when boiling begins.
This is conservative since under the sc'enario conditions defined coolant is allowed to boil off to a level which will keep the fuel assemblies covered.
The applied hydrostatic loads on the walls were linear, increasing from zero at the 2
top surface to approximately 1935 lb/ft at the bottom pool elevation.
Deadweight condition on the pool floor was based-on a uniform loading of approximately 3,830 lb/ft.
The resultant moments and shear forces were developed considering the uniform load being applied to a plate having the floor dimensions and thickness.
The 3,830 lb/ft uniform load includes the weight of the coolant, fuel racks, contained fuel assemblies and concrete floor.
The weight of the coolant was 2
1,935 lb/ft It was assumed that the weight per cell for the higher density stainless steel racks, approximately 1,030 2
lb/ft would be applied over the entire pool floor surface.
This is conservative, since the weight of the aluminum rack per square foot is less, and approximately 20% of the pool floor is not covered by a rack.
A uniform load of 1
. 2 approximately 865 lb/ft represents the weight of the pool floor concrete.
WALLS AND FLOOR BENDING MOMENTS, SHEAR FORCES AND DESIGN MARGINS Resultant bending moments and shear forces due to the above loading conditions which develop during the hypothesized accident s7enario are summarized in Table 1 for locations of maximum bending and shear.
Locations defined in Table 1 can be correlated to wall and floor locations by the number index shown below:
rop of WALL n
r a+
='
(FLaca Pt.AW VIEW) tt0 l
_a 3
m~ c, ww J4.,
b/g n
b
=
=
J These resultant bending moments were algebraically i
combined to define the direction of bending.
A conservative absolute summation was then performed and compared with the loading-carrying capacity in the defined direction.
Minimum margins in bending and shear for each wall and the floor are given below:
Moment Margin Shear Margin East Wall 81.9%
85.4%
North Wall 74.9%
68.9%
South Wall 64.5%
37.8%
I Floor 37.2%
27.1%
f In summary, the spent fuel pool concrete will with-i stand the loads resulting from the accident scenario postu-lated by Mr. Blanchard.
POOL LINER STRENGTH AND STRESS CONDITIONS l
The spent fuel pool stainless steel liner acts as a secondary seal to protect the concrete from the pool water.
--q.-
e.-
,,,-a-r%, - -
,-mw,,ww, y-m-
,----*,.u-2-r,---
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ww----.
w*F
==
ar-s
. 4
'Jhe liner was analyzed for (i) thermal loads due to relative j
thermal expansion between the liner and concrete walls, (ii) hydrostatic loading of the water, and (iii) contact loads between the liner and rack legs assuming the racks were i
completely loaded with fuel assemblies.
The only liner anchor locations are at the extreme top of the spent fuel pool.
Plates 1/4" thick are welded to the liner at the anchor elevations which provide additional plate thickness to prevent the liner from tearing around the anchors.
In all cases, the anchors are located above the spent fuel pool coolant elevation.
This means that shear i
failure of the anchors would not break the water seal provided 1
by the liner.
Structural integrity of the anchors is, there-i i
fore, not required to prevent coolant from leaking to the concrete side of the liner.
Critical locations, where tearing would cause coolant leakage to the concrete side of the liner, l
l are at the wall edges and rack leg locations.
These locations j
were analyzed for an extreme loading condition, that is, not allowing the liner to bow during thermal expansion.
During the coolant heat-up period, the liner will expand approximate-l ly 0.19 inch more than the concrete at the coolant boiling 1
temperature of 237'F.
Compressive forces would then develop at the wall edges, and bowing of the liner along the length will occur.
Buckling of the plate, however, will not occur l
l
+ wr f e er ge-Tv mw
. - - > - - = >
sa-
-r
-w-
-.si..yw-e-p-,
w'wsq-im--vy-w--see-Wry,a.
---gy-=uy--+
wey-py
,w w+
e-v-=rw-+g&-=-*w--=wt-w-*--g--sw---
1 4
?.
)
i due to the hydrostatic pressure of the coolant.
The most probable bowing condition would be a small waving effect along the liner length.
This small waving condition would develop because the required pressure to maintain the plate straight is greater than the hydrostatic pressure of the water.
The j
small bowing effect of the plate will accommodate the 0.19 i
inch thermal expansion.
For analysis purposes, the worst stress condition would occur if the plate were assumed to be straight rather I
than in a bowing configucation.
Under this condition, the i
total expansion would go into compressing the plate,-leading to ma::imum stresses at any location throughout the length of liner.
Additional compressive stresses in the' plate which 3
would occur due to hydrostatic loading of the coolant and rack leg to liner contact loads were included.
Maximum st' rain energy theory was used to determine the resultant liner stresses.
The stress is defined by the following equation:
i T" I6 2+
2+8 2 1/2
- 2p(S Sy2+382 3 + 8 3 Il j
S 31 1
2 3
V l
where i
are the directive stresses, S
S2 ""0 83 y,
S is the total compressive stress.
T 1
l i
,...,,-,---...,,--_,---v.r.
,,----,-..-,n.,
.,,..-r
--,,,w.,
,,.,.-,.,m,,
,,y-
,,m,,,
--,.,,.n,__,.,,,,._,,-,y,
-~.,--,-, ----- - _.-.,~- me,
o
. 1 Based on the relative thermal expansion between the t
concrete and liner of 150*F, stresses in the two horizontal directions become t
1, 2" I"304ss ~"c)
E T
S S
s I
where AT = thermal expansive temperature range
-6
= coefficient of thermal expansion is 5.5 x 10 ac l
in/in*F 304ss = coefficient of thermal expansion a
6
-6
= 9.55 x 10 in/in*F 6
E
= modulus of elasticity = 27.49 x 10 psi.
s The directional compressive stress in the liner becomes 16,725 psi.
Rack leg contact stresses and hydroctAtic compressive stresses are respectively 549 psi and 28 psi, I
making the total through thickness liner directional stress S3
= 677 psi.
Solving the maximum strain energy stress equation for Poission Ratio, p,
equal to 0.29 given a total compressive stress, S
= 19,610 psi.
Minimum yield strength for 304 T
stainless steel at 237'F is S
= 24,125 psi.
A comparison shows that the total compressive stress at a rack leg location is below the yield strength material level.
1 1 POOL RACK STRENGTH CONSIDERATIONS With respect to the spent fuel racks, boiling conditions of the coolant represent an 87'F increase in temperature above the design temperature of the racks, 150*F.
This change in temperature would cause the racks to experience a slight thermal expansion.
Since the racks in the Big Rock spent fuel pool are not restrained, no additional stresses would develop.
Only the change in rack material strength properties requires investigation.
A comparison of the strength properties for the existing aluminum racks and new stainless steel racks at the design and boiling temperature is given below.
Strength properties are for the stainless steel and aluminum racks.
Yield Ultimate Temperature Strength Stretagth Rack Material T( F)
_S (ksi)
S (ksi) u 304 Stainless Steel 150 27,500 73,000 237 24,125 69,150 Aluminum Alloy 6061-T6 150 39,755 43,605 237 38,316 43,360 The yield strength of the stainless steel racks decreased by 12.3% while the yield strength of the aluminum racks decreased by only 3.6%.
. i In summary, the new spent fuel storage racks would
,i withstand the loads at the boiling conditions of the coolant, while maintaining structural integrity.
The reduction in the i
j aluminum racks' yield strength would be only 3.6%.
This reduction in strength is small and, hence, the effect of boiling is insignificant.
4 CONCLUSION I
The foregoing analysis demonstrates that the structural integrity of the spent fuel pool walls, floor, liner, and racks are noc adversely affected by the ef fects of pool boiling.
INCREASED CONCRETE DEGRADATION EVALUATION' SUPPLEMENT Subsequent to the Exhibit A analysis an additional evaluation was performed which accounts for increased degra-dation.
This evaluation, attached as Exhibit B, was performed in order to establish whether concrete degradation in the i
range of 50-60% could be tolerated by the Big Rock Point spent fuel pool.
As explained above, such degradation is not considered possible.
Testing performed by Oak Ridge National Laboratory (References 7 and '.))
indicated a 15% reduction in I
bond strength for limestone aggregate concrete specimens.
(Reference 8) identified bond Testing performed in Japan e
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1 strength reductions of 50 to 60% for temperatures ranging from 212*F to 572*F for silica aggregate.
Based on my review of Walter H.
Flood and Company Test Reports for Bechtel
]
Corporation, cited at page 4-1 of Exhibit B, it is my judgment that limestone aggregate was used for the construction of the Big Rock pool.
Therefore, the Oak Ridge National Laboratory data was considered more representative of the Big Rock spent fuel pool concrete.
Nonetheless, for study purposes the increased degradation of 50-60% was investigated.
It can also be postulated that since bonding represents the strength which assures that the reinforcement bars stay attached to the concrete, one mode of failure could be a shear or tension failure of the concrete in contact with the reinforcement bars.
Consequently, although such a result is not substantiated by any technical paper or test data that I have found, the pool shear strength was also considercd to be degradated by the higher bond strength reduction levels.
l When I performed the initial analysis that forms the basis of Exhibit A, the actual compressive strength of the Big Rock pool concrete was not made available to me.
In per-forming the latest evaluation, Exhibit B, I was able to establish a more realistic starting basis for the initial pool concrete strength.
Actual construction site compressive strength specimens made from the same concrete pours as the i
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pool were used in the Exhibit B analysis to determine a statistical lower bound strength.- The statistical applications were in accordance with recommended American Concrete Institute methods.
A minimum strength increase of 10% due to aging was used in the Exhibit B analysis.
Review of similar one-year concrete curing curves, however, shows j
that strength increases due to aging of 19 to 27% can be anticipated.
The conservative 10% increase due to aging that was used results in an actual lower bound pool concrete compressive strength of 3,472 psi as compared to the 2500 psi value used in the Exhibit A analysis.
This compressive strength, which is one parameter in the expressions for determining the required reinforcement development length (bonding strength) and shear strength capacities was used as the base strength at ambient temperature.
Shear strength capacities for the walls and floor were calculated based on the actual lower bound compressive strength of 3,742 psi.
In order to determine if enough margin existed to accommodate a 50%-60% reduction in strength a comparison was then made between these strength capacities and the maximum shear loads.
An a result of this comparison it was determined that reduction of over 70% could be tolerated for even the maximum shear location for the east and north t
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. walls.
This was not the case for similar locations at the south wall and floor.
However, the American Concrete Institute Standard (Reference 2) does not require considera-tion at these points.
The Standard only requires that sections located less than a distance d (distance from extreme compressive fiber in the concrete to the center of the tension reinforcement) from the face of a support may be designed for the same shear as that computed at a distance d from the face of the support.
The face of the support in the case of the south wall and floor are respectfully 3'-0" and 5'-0" from the wall and floor intersections.
At these points reductions of 60% in shear strength could be tolerated when compared to the original shear strength capacities.
Reinforcement development lengths were also calcu-lated in accordance with the requirements of the Reference 2 standard.
These values were then doubled to account for a 50%
reduction in bonding strength.
Critical locations in the wall where the ends of reinforcement bars overlap each other known as splices, along with wall and floor intersectine corners were checked to assure that the necessary length of reinforcement bar existed beyond the points of applied loads.
A summary of this review (actual development lengths to required development lengths for a 50% reduction in bond strength) is as follows.
. Location Actual l Required l d
d Wall to Floor Intersections 68" 21" Floor Slab Top Reinforcement 36" 32.8" Splices at 10' Above the North 30" 31.4" or South. Walls Horizontal Wall Peinforcement 24" (at least) 19.6" In all cases but one the actual development length is greater than the required length for a 50% reduction in bond strength.
In the case of the spliced reinforcement located 10' above the north and south walls, the actual development length is 1.4" shorter than that required to reach yield strength in the reinforcement bars.
The maximum tension stress in this rebar, however, is only 1,189 psi or 3.2% of yield.
Therefore, only a minimum amount of the 30" development length is needed in order to maintain reinforcement bonding and the ability to resist bending moment loads.
In conclusion, a reduction of 50-60% in bond strength can be tolerated in the Big Rock Point spent fuel i
pool design.
This margin is based on three items:
the originally applied reduction of-50%, the margin still identi-fied between actual and required l values to reach yield and d
the fracticn of yield stress actually reached in the rebar.
I EFFECTS OF POINT LOADING FROM STORAGE RACKS Having addressed the overall structural strength of the pool for boiling conditions, I will now consider local point loading from the storage rack legs.
Section 2.5.5 of the Description and Safety Analysis dated April 1982* addresses local bearing and punching shear point loading conditions at the new storage rack legs in Section 2.5.5.
The bearing stress results from the load applied over the rack leg pad surface in contact with the concrete.
The punching shear stress is simply the local loading condition under the rack leg which could shear or punch a hole through the pool floor.
These conditions were analyzed at a temperature of 150 F, the maximum normal allow-able operating temperature.
I have reviewed the analysis summarized in the licensing submittal for technical adequacy, and I will explain in niore detail the results.
In addition, I have reduced the margins calculated in that analysis in order Consumers Power Company, Big Rock Point Plant, Spent Fuel Rack Addition, Consolidated Environmental Impact Evalua-tion and Description and Safety Analysis, April 1982.
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to account for lower allowables at the higher temperature 1
condition of 237 F.
4 The new spent fuel storage racks rest on the pool j
floor on either four or six rack legs.
Each support leg i-consists of 3.50 inch diameter screws which are threaded into the rack bacc.
These screws are then attached to 5.0 inch j
diameter pads which rest on the pool floor.
The largest bearing load would result from a full rack having only four support legs.
For this maximum bearing condition a margin of 1.61 exists wh'en the actual bearing stress was compared to the limits for local bearing stress cited in Section 2.5.5 of the f
Description and Safety Analysis.
This margin, however, is 4
based on the maximum normal allowable operating pool tempera-t l
ture of 1,50 F.
The bearing stress is a function of compressive strenoth, so a reduction of 20% in the concrete strenth at this temperature was applied.
This reduction would
)
result in a margin of 1.29 between actual bearing stresses and limits at the 237*F boiling temperature.
s The existing spent fuel storage rack design consis-ted of 7 1/2" square pads resting on the pool floor at every other storage cell location.
There are a minimum of 36 bearing pads for each existing rack resulting in essentially a 4
i uniform load distribution.
This distribution of loading, combined with a lower total full weight of the existing racks, i
_- __. _ _ _ _ _ _ _ _ _., _.. _. _ _ _ _.. _ _. _ _ _. -.. ~ _ _ _ _. _
O would make the new rack local or point loading conditions limiting.
In the analysis summarized in the Description and Safety Analysis, two cases were considered for punching shear stress.
The first case considered local punching shear under one rack support leg.
The second case considered punching shear conditions in a region where the nearest three support legs from adjacent racks met.
The second condition was slightly more severe than the first, however, margins were in excess of 10.0 even after considering the reduction to the ACI Standard limits of 25% due to the postulated effects of the 237*F boiling temperature condition.
In the Exhibit A analysis, I did not attempt to calculate point loading ca the floor from the storage racks; rather, I calculated the maximum bending moments and shear stresses resulting at critical locations such as the walls and the floor intersection.
As stated previously, a uniform loading condition equal to the weight density of the new storage racks was considered to exist over the entire pool floor surface.
The actual load distribution consisted of essentially a uniform loading condition for the existing racks and a total of fourteen support leg point loads dispersed over
s half the floor surface for the new storage racks.
The amount of deadweight loading imposed in the Exhibit A analysis is approximately 75% larger than the actual deadweight loads.
This is due to over 20% of the floor surface being. vacant space and over 40% of the floor surface being covered by existing fuel storage, equipment, and fuel channel storage racks.
The uniform loading for these racks is lower in magnitude than that for the new higher density racks.
As a result of the actual loading condition, maximum bending moment and shear stress values summarized for deadweight loading conditions in the Exhibit A analysis exceed actual values.
In summary, the fuel pool concrete is strong enough to withstand all point loading conditions even at the maximum coolant boiling temperature of 237 F.
. - _ ~ -.
e T AliLE la Load Combination - 5trength Comparison Summary Moment in Moment in llorizontal Moment Ver t 6 cal Moment Shear 5twar Reinforcement Capacity Margin R e6nf orcement Capacit y Margin Force
' Capacity Mar gin
_ all- - - -
Location it-lb/f t it-lb/f t lt-lb/f t it-Ib/f t lb/ft Ib/ft W
rast I
-7973*
-46801 8).O
-4340
-26838 83.8 2119 14688 83.4 2
-829)
-46801 82.)
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S t.9 3
I 19.1
-29942
-l.19tal0' 74.9 17136 33080 68.9
-2.206:10l Nor th 1
-46210 a
-2.206:10 83.1 10416 3.34)l mig 78.9 2
-37183
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-33205 4.206x10 85.0 3108) 2.30)n 10 79.6
-1.010l
- 57 69.3 171 %
27340 37.8 2.0x%
76.I
-17384 South 1
-23911 10 St.l 2
-18334
-l.0ml0 31.7 37112 3
4 35437 1.0 10 64.5
-I5323
-37000 72.8 Moment in Moment in East-West Nor th-South Reinforcement Reinforcement it-lb/f t it-lb/f t 76.7
-7 % 26
-2.629x t D' 69.7 35712 48960 27.1
-2.629mulf riour 1
-6121) 5 2
-33004
-2.269s10 79.1 I)6888 2.526x10 45.8 3
3 3
-47957
-2.314 mig 80.9 103913 2.668s10 61.1 3
4 ll190L 1.878ml0 37.2
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REFERENCES 1.
American Concrete Institute Standard 318-71, " Building Code Requirements for Reinforced Concrete," 1971.
2.
American Concrete Institute Standard 318-77, " Building Code Requirements for Reinforced Concrete," 1977.
3.
Technical Report CONF-790408-3, " Strength Properties of Concrete at Elevated Temperatures," by G.
N.
Freskakis, R. C.
- Burrow, E.
B.
Debbas, presented at the American Society of Civil Engineers Specialty Conference on the Structural Design of Nuclear Plant Facilities, Boston, Massachusetts, April 2, 1979.
4.
J.
P.
Callahan and G. C.
Robinson, " Uniaxial Compressive Strengths of Concrete for Temperature Reaching 1033 K,"
Nuclear Engineering and Design, 45 (1978), pp. 439-448.
5.
Oak Ridge National Laboratory Report ORNL-4227, " Pre-stressed Concrete in Nuclear Pressure Vessels, A Critical Review of Current Literature," by Chen Pang Tan, The Franklin Institute Research Laboratories, May 1968.
6.
American Concrete Institute Special Publication SP-34, Concrete for Nuclear Reactors, Copyright 1972, paper SP 34-26, " Influence of High Temperature on Strength and Deformations of Concrete" by H. Weigler and R.
Fisher.
7.
Technical Report CONF-810801-53, " Testing Program for Concrete at Temperatures to 894 K,"
by D.
J. Naus, C.
B.
Oland, and G. C.
Robinson, presented at the International Conference on Structural Mechanics in Reactor Technology, Paris, France, August 17, 1981.
8.
American Concrete Institute Special Publication SP-34, Concrete for Nuclear Reactors, Copyright 1972, paper SP 34-21, " Strength Elasticity and Thermal Properties of Concrete Subjected to Elevated Temperatures," by T.
Harada, J.
- Takeda, S.
Yamane, and F.
Furumura.
9.
Oax Ridge National Laboratory Report, ORNL/TM-6086, " Bond Between Concrete and Steel Reinforcement at Temperatures of 149*C," by C.
B.
Oland and J.
P.
Callahum, April 1978.
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2 10.
American Society for Testing and Materials Standard C 234-71, " Standard Task Method for Comparing Concretes on the Basis of the Bond Developed with Reinforcing Steel."
11.
American Society of Mechanical Engineers, " Boiler and Pressure Vessel Code,"Section III, Division I, Appen-dices, Nuclear Power Plant Components," 1977 Edition.
12.
American Concrete Institute Standard 349-69, " Criteria for Reinforced Concrete for Nuclear Power Plant
_Contain. ment Structures," Appendix A, 1971.
j, e
~
t
- 13.-
Elleuch, " Behavior of Special Shielding Concretes and of
'Their Constituents Under Neutron Irradiation " Fourth
-United Nations' International Conference on the Peaceful l
'Use of Atoni* Energy,llA/CCNF.'49/P1613, July 1971.
c 14.
- Sommers, J.
F.,
"Ga'mma Radiated Damage of Structural Concrete Immersed in Water," Health Physics, Volume 16, April 1969.
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RAYMOND F. SACRAMO EDUCATION University of Pittsburgh, M.S., Mechanical Engineering,1977 Penn State University, B.S., Mechanical Engineering,1974 Drexel University, Civil Engineering,1967 EXPERIENCE NUS CORPORATION,1977 Present Westinghouse Electric Corporation, 1974-1977 United States Army, 1968-1970 Beaumont Birch Company, 1966-1967 NUS - Responsible for directing the analytical efforts required to support Engineering Division design projects. These projects cover a variety of modifications and new system incorporations at more than 25 nuclear power plants resulting from TMI, retrofit, or NRC l.E. Bulletin nuclear-industry-related programs. Associated analytical efforts include various levels of finite element modeling using both static and dynamic analysis techniques, as required, to investigate normal, upset, and emergency loading conditions imposed on mechanical or electrical systems and associated support structures.
Responsible for developing finite elen$ent models and performing structural integri y stress anal-t yses for: ASME B&PV Code Class Il and ill piping systems; safety-related components; and high-density spent-fuel racks and spent-fuel-pool concrete structures. Analyses have employed static, response spectrum and nonlinear displacement time-history analysis techniques. Analyses are performed in accordance with the rules and requ;rements of the appropriate ASME, AISC, or ACI Codes. A!sc responsible for the fabrication followup efforts for high-density spent-fuel racks designed by NUS.
Westinghouse Electric Company - On the Clinch River Breeder Reactor Project, was respon-sible for assessing the mechanicsl reliability and integrity analyses on mechanical components interfacing with the shutdown system. Responsible for the development of design support docu-ments for the reactor upper internal structure, core support structure, bypass flow and lower inlet modules, head heating and cooling systems, and riser assemblies. The documents considered statistical evaluations of thermal, seismic, and accident-condition transients and predicted the failure probabilities of structural-integrity-related failure modes resulting from these transients.
The documents were presented at final design reviews and were responsible for design changes that decreased the probabilities associated with component failure modes, which affected the mechanical shutdown system reliability.
United States Army - Assigned to perform mechanical design and drafting.
Beaumont Birch Company - Performed drafting and mechanical design.
MEMBERSHIP American Society of Mechanical Engineers N
8 s
l3 NtJS COAPORATION
1 RAYMOND F. SACRAMO Page Two PUBUCATIONS i
"Statis.tcal Fatigue Testing and Analysis of 304 Stainless Steel," Thesis, University of Pittsburgh, 1377.
" Reactor Scram Experience for Shutdown System Reliability Analysis" (coaatner) American Wucle ir Society Techr cal Paper. June 1976.
u "Probabilistic Assessment of Primary Piping integrity" (coauthor), American Nuclear Society Technical Paper, April 1976.
1 I
l NUS CORPOAAPON
I'XHIBIT A N'U9f))w $
El S EM D-RFS-013 u
U COAPORATION INTERNAL COPIRESPONDENCE TO:
G. Antenucci, Jr.
CATE:
May 12,1950 FROM:
R. Sacramo f[ff COPlES:
H. Eckert J. Wawr:eniak SU5 JECT:
Structural Integrity cf the Spent Fuel Racks After a Pool Coclant System Failure at Eig Rock Point Nuclear Power Plant RIFERENCES: 1.
USNRC Standard Review Plan 3.8.4 "Other Category I Structures,"
November 24, 1975.
2.
ASME Boiler and Pressu:e Vessel Code;Section III, Divisien 1, Appendices Nuclear Power Plant Components,1977 Edition.
3.
Nuclear Systems Materials Handhcok, TID-26666, Volume 1,
" Design Data" Sock 2.
Property Ccries 2101, 2102 and 2105, Revision 0, September 30, 1976.
4.
Specification #514S-P-103, " Revision 2 " Design Input Require-ments for Big Rock Pcint Spent Fuel Rack Modification Project, May 14,1979.
In the event that the spent fuel pool coolant system were to fail, the temper-ature of the coolant could eventually reach the boiling point. De maximum s
boiling temperan:re of the ecclarit, 23f F, would occur at the top of the active length of the spent fuel assemblies. Bis represents an 8fF increase in temperamre from the design temperature of the racks,150*F, identified in Reference 4.
The change in temperature would cause the racks to experience a slight ther-mal expansion. With the racks not being restrained, and constructed of the same mhterial, no additional stresses will develop. Only the change in rack t
material strength properties requires investigation to assure structural integ-l l
rity of the racks. If design conditions for the racks can be maintained after the temperature rise, rack distortion would be prevented, and the distances to assure criticality will be maintained.
r D.iD-RTS-013 May 12,1980 Page Two A comparison of de streng$ properties for the existing aluminum racks and new stainless steel racks at the design and boiling temperature is given Strength properties were cb:ained from Paferences 2 and 3, respec-below.
tively, fcr Me stainless steel and aluminum racks.
U1timate Temperature Yield Strength Strength Paek Meterial
_T
(* F)
_ Sv Ocsi)
Su (ksi)_
0 27500 73000 304 Stainless Steel 237 24125 69150 150 39755 43605 Aluminum Alloy 6061-T6 237 38316 43360 he yield strength of the stainless steel racks' decreases by 12.3% while the yield strength of the aluminum racks decreases by only 3.6%.
l j
he stainless steel racks were designed in accordance with the elastic Bis criteria permits only the extreme fibre design criteria of Reference 1.
De new racks meet this criteria for the stresses to slightly exceed y'ield.
worst design condition, which also included seismic loads. For the abncrmal loading condition of pool coolant boiling, the plastic design criteria of De Reference 1 allows the entire cross section to reach 90% of yield.
margin realized under this design criteria is well in excess of the 12.3%
reduction in material yield strength. Derefore,the new stainless steel racks would meet the plastic design criteria of Reference 1 for even' the combined loading condition of seismic loading during boiling conditions of the coolant.
In the case of the existing aluminum racks, the same analogy would hold true.
.b j
EM D-F.FS-013 May 12, ISB0 Page Taree In summa./, de new sgant fuel s crage racks would withstand the worst design loads, even at boiling con:iin ns cf ie coolant, while maintaining structural ir.tegrity and criticalief limits. The reductio:i in the aluminum racks' yield streng6 would be cnly 3.6%. D.is reduction in strength is small and, hence, :he effect of boiling is insignificant.
[ Exhibit A also includes Revision 1 to NUS-3567, " Structural Analysis of the Spent Fuel Pool Liner'and Concrete.Due to Coolant System Failure for the Big Rock Point Nuclear Power Plant," dated April 27, 1982.]
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