ML20031D589
| ML20031D589 | |
| Person / Time | |
|---|---|
| Site: | Big Rock Point File:Consumers Energy icon.png |
| Issue date: | 09/28/1981 |
| From: | Sacramo R CONSUMERS ENERGY CO. (FORMERLY CONSUMERS POWER CO.), NUS CORP. |
| To: | |
| Shared Package | |
| ML20031D553 | List: |
| References | |
| ISSUANCES-OLA, NUDOCS 8110130505 | |
| Download: ML20031D589 (20) | |
Text
UNITED STATES OF AMERICA
(
NUCLEAR REGULATORY COMMISSION EEFORE THE ATOMIC SAFETY AND LICENSING BOARD IN THE MATTER OF
)
)
Docket No. 50-155-OLA Pent u ! Pool CONSUMERS POWER COMPANY Exp,n, 9n)
)
(Big Rock Point Nuclear Power Plant)
)
STATEMENT OF RAYMOND F. SACRAMO col <CERNING CHRISTA-MARIA CONTENTION 8 AND O'NEILL CONTENTION IIIE-2 My name is Raymond F. Sacramo. I reside at 17204 Chiswell Road, Poolesville, Maryland. I have been employed with NUS Corporation, an engi-neering consulting firm in Rockville, Maryland, since October 1,1977. Prior to my current position, I was employed with the Advanced Reactor Division of Westinghouse Electric Company in Madison, Pennsylvania.
My position at Westinghouse Electric Company was that of a Reliability Engineer on the Clinch River Breeder Reactor Project. My resume is attached to this statement. Based on my eductional background and work experience, I believe I am qualified to provide the structural integrity analysis of the effects of pool bolling on the Big Rock Spent Fuel Pool which is required to answer Christa-Maria Contention 8 and O'Neill Contention III E-2.
TE$' Board consolidated and rewrote the contentions to read as follows:
The occurrence of an accident similar to TMI-2 which would prevent ingress to the containment building for an extended period of time would render it impossible to O
l Oh o$d[ofj5 PDR'
maintain the expanded spent fuel pool in a safe condition and would result in a significantly greater risk to the
,Q public health and safety than would be the case if the V
increased storage were not allowed.
The structural integrity of the spent fuel pool concrete, Jiner and storage racks under the loading condition 'vhich would develop during the accident scenario defined in Mr. Blanchard's affidavit was investigated. This loading condition being the pool filled to its proposed capacity of 441 assemb!!es during a boiling condition of the coolant which wi!! be maintained above the level required to cover the spent fuel assemblies. This loading condition is considered to exist for an extended period. In order to determine the affects of' this accident scenario the following areas were addressed:
o Affects of the accident envirenment on degrading the pool concrete strength.
Strength load-carrying capacity definition of the pool geometry.
o Thermal gradient, hydrostatic and deadweight loading conditions o
across the various walls and floor during the accident.
Accident condition bending moment and shear loads imposed on the o
walls and floor.
Design margins under the resultant accident loading conditions, o
Pool liner strength and stress conditions during the accident.
o Fuel rack strength considerations during coolant boiling.
o The spent fuel pool being investigated was placed in service during 1962. Over the years, the major parameters which would have affected the concrete strength of the pool are temperature, aging and irradiation. These o -
parameters were investigated considering a variety of aggregate materials by O
reviewing the pertinent technical literature.
A summary of published test data nn concrete strength properties at elevated temperatures was presented at the American Society of Civil Engineers Specialty Conference on the Structural Design of Nuclear Plant Faci!!tles, held in Boston, Massachusetts on April 2,1979. The technical paper, CONF-790408-3, entitled " Strength Properties of Concrete at Elevated Temperatures" was prepared by Messrs. George N. Freskakis, Richard C. Burrow, and Elias B.
Debbas of Burns and Roe, Inc. The literature study, sed on over fourteen different aggregate 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 estab!!shed as a standard by the American Concrete Institute. The literature study considers a sufficient number of aggregates and mixtures and contains a sufficient number of tests on the effects of temperature on 28-day conclude to conclude that it applies generically to concrete, including the concrete in the spent fuel pool at Big Rock Point.
Dhia was extracted from the literature study for the maximum boiling temperature condition of 237 F as defined in Dr. Perelewicz's affidavit.
For all of the aggregates tested, the data shows a range from no reduction in I
strength to a maximum reduction of 25% at 237 F.
Concrete compressive strength does, however, increase during aging beyond 28 days. A test program carried out at Oak Ridge National Laboratory and also summarized in CINF-790408-3 considered such aging conditions. Eight O -
to nineteen month old concrete cylinders of a !!mestone aggregate mix were O
tested. A limestone aggregate mix would lie approximately midway between the strength bounded reduction levels defined above and is therefore a representa-tive example. The cylinders were increased to their tested temperature at a rate 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 sho'ved that the compressive strength at 237 F was approximately 120%-130% of the 28-day-old strcngth at normal temperature. Consequently, concrete aging compensates for the effects of temperature degradation at this temperature level. The remaining temperature consideration is the prolonged effect of temperature on concrete strength. Even though strength, reduces as the temperature condition prolongs, it stabilizes under steady-state conditions.
Steady-state conditions will occur approximately eight days af ter the coolant bolling condition begins. The test results are therefore reasonable for the pool acd#..t environment as postulated by Mr. Blanchard.
Irradiation effects tend to degrade concrete compressive strength.
At the Big Rock Point Plant, spent fuel is randomly placed in the fuel pool racks.
The following analysis is based on placing new spent fuel in one location, b, the center of the pool floor. This assumption is conservative when compared to the actual method of spent fuel storage described above for the Big Rock Point Plant.
The irradiation would be of two types: that produced by neutrons and that produced by gamma rays. Neutron fluxes eminating from spent fuel is
- O,
l normally low. The fluxes are further reduced due to the neutron absorbing O
capability of the water. A conservative total neutron exposure estimate for this center pool floor area after a service life of 40 years would be on the order of II 2
10 n/cm. Test results contained in ORNL-Report-4227, " Prestressed Concrete in Nuclear Pressure Vessels; A critical Review of Current Literature,"May 1968, showed that the compresslu strength properties of a reactor shield made from Barytes-Haydite concrete did not degrade at such irradiation levels. Serpentine concrete specimen test t esur.s summarized by Elleuch, " Behavior of Special Shielding Concretes and of Their Constituents Under Neutron Irradiation,
" Fourth United Nations International Conference on the Peaceful Use of Atomic I9 Energy, A/ CONF.49/P1613, July 1971, showed irradiation dosage levels of 10 2
n/cm produced only s!!ght changes in concrete compressive strength.
An upper-bound estimate of gamma irradiation for a service life of 10 40 years is 10 Rads. Test results summarized by Sommers, 3. F., " Gamma Radiated Damage of Structural Concrete Immersed in Water," Health Physics, Volume 16, April 1969, showed that specimens receiving such gamma ray irradiation levels on the average sustained no degradation. The upper-bound loss in compressive strength was only 10% No appreciable strength reduction in concrete is, therefore, expected to occur from the affects of either neutron or gamma irradiation over a 40-year service 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 tempera-l l
ture (approximately 70 F). Nevertheless, for the purpose of conservatism, a 0
lO
l lower-bound 80%, 28-day compressive strength level at 237 F was used as the design criterion.
The next step in the analysis requires the discussion of the load-carrying capacities which will resist the resultant bending moment and shear forces that can develop during the accident postulated by Mr. Blanchard.
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 reinforce-ment or rebar. The spent fuel pool walls and floor reinforcement which has a yield strength of S = 40,000 psi at 70 F vould be reduced to 37,175 psi at 237 y
F. Steel yield strength and other material properties at temperature were taken from the American Society of Mechanical Engineers "Boller and Pressure Vessel Code;" Section 111, Division I, appeno:ces Nuclear Power Plant Components,1977 Edition.
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 use'd by American Concrete Institute, allowed the stress of the concrete reinforcement, under teaslie loading, to reach 90% of the reinforcement yield strength.
Under bending moment loading conditions, reinforced concrete would fail when a load slightly larger than that which causes the tension reinforcement to exceed the yield j
strength is app!!ed. The failure condition of reinforced concrete is explained in the technical book, " Reinforced Concrete Fundamentals," Third Edition, John lO
Wiley & Sons, Inc.,1973. A summary of the condition of failure is as follows.
When a slight additional load is applied beyond the load causing yield in the reinforced steel, the steel stretches a considerable amount. The increased steel deformation in turn causes the neutral axis of the concrete cross-section to shif t toward the compressive side. The shift in neutral axis reduces the area under compression and thereby increases the unit corapressive stress. This process of additional steel deformation, shift in neutral axis, and compressive area reduc-tion continues until the concrete falls in compression as a secondary effect.
When steel does not have a sharp yield point, the load required for failure will increase, making calculations based on 90% of yleid strength conservative. It should also be noted that the ultimate strength design method assumes that the tenslori 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 surfaces. 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 locatior were developed. Two load-carrying capacities were developed for both the horizontal and vertical reinforcement. The first set of two load-carrying capacities were based on the final resultant moments causing tension in the pool side reinforcement. The second set considered tension in the outside wall reinforcement.
The shear stress capacities were based on the area of concrete and has one value for a given wall location. Concrete is weaker in tension than compression with real shear strength intermediate between the two. Based on O 1
I the large difference between concrete compression and tension strength, molt concrete shear failures that would be termed shear failures are really diagonal tension failures.
The 28-day compressive strength of the spent fuel pool concrete is i = 2500 psi, while the tensile strength, being approximately 10% of c
the compressive strength, is 250 psi. Minimum shear strength values based on American Concrete Institt.te Standards were used in development of shear-I strength allowables. As an example, the typical shear strength allowable of a I
wall section is 1.2[. Based on a compressive strength of 2500 psi, the shear strength allowable would be 60 psi. As can be seen, the selected shear strength is well below what would represent real shear strength.
Having established the load-carrying capacities of the Big Rock spent i
fuel pool, I determined the loading conditions imposed on the pool. Four critical i
j or bounding cross-sections were investigated. The topmost portion of the east i
wall which is 2'-0" thich 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 the coolant boiling i
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 l
coolant. The 6'-9" thick north wall would develop the largest bending moments due to the thermal transient. 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 transient. Again, 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 remalps constant at the boiling temperature of the coolant. The south wall, below elevation 624'-0", varies from j
5'-9" thick on tne 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 O
3-
thickness down to the ground floor elevation. This section of wall would provide O
the least load carrying capacity for 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.
i Thermal loading resulting from these temperature gradients were developed based on the input discussed in Dr. Prelewicz's affidavit and they are defined by the following equations:
2'-0" thick east wall AT = AT (z/t)2.7 = 128 (z/t)2.7 z
6'-9" thick north wall AT = AT (z/t)2.4 = 150 (z/t) z 3'-6" thick south wall M = M (z/t)2.0 =142 (z/t)2.0 z
6'-0" thick floor g, g (,7g)2.2 = 150 (z/t)2.2 z
where AT = the difference in temperature between the poolside and outside surfaces of the wall or floor
- z = the difference in temperature between the outside 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.
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A concrete surface subjected to a curvilinear g Mient, as those i O defined by the above equations, will develop a moment due to the gradient.
Also, a resisting moment caused by forces (P ) acting over a distance (L) results T
from the restraints at the corners. The restraints are the other walls and floor of the pool. A futher effect is the possible cracking of the concrete in tension.
This cracking results from an increase in the stresses in the tension side 4
reinforcement due to a shif t in location (L) of the resisting forces P.g..
i The tension stresses la the reinforcing steel which would develop due to the thermal gradients and the equivalent moments required to develop the stress were based on enalytical methods contained in the American Concrete Institute Standard 349-69 "Cr!teria for Reinforced Concrete for Nuclear Power Plant Containment Structurer," Appendix A,1971.
For the mechanical loading conditions (deadweight and hydrostatic),
the analysis of fuel pool walls and floor was performed using thin plate theory.
Thin plate theory la based on sheat deflections being small compared to the flexural deflections.
This condition is normally maintained by limiting the 8
thickness of the plate to about one-quarter of the least transverse dimension.
For reinforced concrete, the flexutal 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 Jurfaces of the wall, is within one-quarter of the least transverse dimension. Also, since the reinforce-ment is extremely ductile as compared to the concrete, one can assume that the j
shear deflections are small compared to the flexural deflections. The walls and l
l floor will also act as integral members with each other in distributing the loads.
!O _
Loads would be distributad in a propoitional manner to the moments of inertia i
of the various walls and floor comprising the pool. A simplifying and conserva-tive assumption was made that all walls and floor would. ct 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.
}
j The pool walls were analyzed based on three sides fixed and the top side simply supported. 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 pcol filled to the i
maximum elevation when boiling begins. This is conservative since under the scenario conditions defined coolant is allowed to boil off to a level which wl!!
keep the fuei assemblies covered. The applied hydrostatic loads on the walls i
2 were linear, increasing from zero at the top surface to approximately 1935 lb/f t i
at the bottom pool elevation. Deadweight condition on the pool floor was based 4
on a uniform loading of approximately 3,830 lb/ft. The resultant moments and shear forces were developed considering the uniform load being app!!ed 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 2
concrete floor. The weight of the coolant was 1,935 lb/ft. It was assumed that 4
the weight per cell for the higher density stainless steel racks, approximately 2
1,030 lb/ft would be applied across the total surface. This is conservative, since the weight of the aluminum rack per square foot is less, and approximately 0
- 1 l
15% of the pool floor is not covered by a rack. A ualform load of approximately 2
fM5lb/ft represents the weight of the pool floor concrete.
Resultant bending moments and shear forces due to the above loading conditions which develop during the hypothesized accident scenario are sum-marized in Table 1 for locations of maximum bending and shear. Locations defined in Table I can be correlated to wall and floor locations by the number index shown below:
7bP OF WALL.,
af eI 4
t$ (
'O f
(FLoon M.AA/ VIEW)
O h
2 Y
BomoM or wdLL
.}4.-,
h
=
=
b
=
These resultant bending moments were algebraically combined to define the direction of bending. A conservative absolute summation was then O
l l
performed and compared with the load-carrying capacity in the defined direc-tion.
i Minimum margins in bending and shear for each wall and the floor are given below:
Moment Margin Shear Margin East Wall 61.0 %
89.6%
North Wall 60.1%
63.3 %
South Wall 45.8 %
32.8 %
Floor 15.2 %
23.0 %
In summary, the spent feel pool concrete will withstand the loads resulting from the accident scenario postulated by Mr. Blanchard.
The spent fuel pool stainless steel liner acts as a secondary seal to protect the concrete from the pool water. The liner was analyzed for (1) thermal loads due to relative thermal expansion between the liner and concrete walls,(ii) hydrostatic loading of the water and (111) contact loads between the liner and rack iegs assuming the racks were 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 tNekness to prevent the liner from tearing around the anchors. In all cases, the anchors are located above the spent fuel pool t
O l !l -
coolant elevation. This means that shear failure of the anchors would not break O
the water seai provided by the iiaer. structurai iatearity of the nchors is therefore 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, are at the wall edges and rack leg locations. These locations 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 approximately 0.19 inch more than the concrete at the coolant bolling 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 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 small bowing effect of the plate will accommodate the 0.19 inch thermal expansion.
For analysis purposes, the worst stress condition would occur if the plate were assumed to be straight rather than in a bowing configuration. Under this conditiori, the total expansion would go into compressing the plate, leading to maximum stresses at any location throughout the length of liner. Additional compressive stresses in the plate which would occur due to h'/ rostatic loading of d
the coolant and rack leg to liner contact loads were included. Maximum straia energy theory was used to determine the resultant liner stresses. The stress is defined by the following equation:
f O
1 l,
(
l 3 - 2 y (S S + 3 3 + S S [
T * [3 S
+3
+3 g2 23 3g 1
2 where 5,5 and S are the directive stresses 4
3 2
3 i
S is the total compressive stress T
Based on the relative thermal expansion between the concrete and liner from 70' F to 237 F stresses in the two horizontal directions becomes S,5 = (d304ss a AT g
2 c
s where AT = 237 - 70 = 167 F dc - coefficient of thermal expansion is 5.5 x 10-3 in/in' F
- 304ss - coefficient of thermal expansion going from 70' F to 237 F = 9.366 x 10-6 in/in F 6 si P
E - modulus of elasticity = 27.49 x 10 3
The directional compressive stress in the liner becomes 17748 psi.
Rack leg contact stresses and hydrostatic compressive stresses are respectively 649 psi and 28 psi making the total through thickness liner directional stress 5 = 677 psi. Solving the maximum strain energy stress equation for Polssion 3
Ratio, p, equal to 0.29 given a total compressive stress, ST = 20828 psi.
Minimem yield strength for 304 stainless steel at 237' F is S = 24,125 psi. A y
comparison shows that the total compressive stress at a rack leg location is below the yield strength material level.
i O '
With respect to the spent fuel racks, bolling conditions of the coolant represents an 87 F increase in temperature abwe 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.
Ultimate Tempgrature Yle!d Strength Strength Rack Material T( F)
S (ksi)
Su (ksi) y 304 Stainless Steel 150 27500 73000 237 24125 69150 Aluralnum Alloy 6061-T6 150 39755 43605 237 38316 43360 The yield strength of the stainless steel racks decreased by 12.3%
while the y!cid strength of the aluminum racks decreased by only 3.6%.
In summary, the new spent fuel storage racks would withstand the loads at the bolling conditions of the coolant, while maintaining structura!
integrity. The reduction in the aluminum racks' yield strength would be only 3.6%. This reduction in strength is small and, hence, the effect of bolling is insignificant.
10 -
i i
t Conclusion lO The foregoing analysis demonstrates that the structural integrity of a
the spent fuel pool walls, floor, liner and racks are not adversely affected by the 4
effects of pool boiling.
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UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION BEFORE THE ATOMIC SAFETY AND LICENSING BOARD l
In the Matter of
)
)
CONSUMERS POWER COMPANY
) Docket No. 50-155-OLA
)
(Spent Fuel Pool (Big Rock Point Nuclear Power Plant)
)
Modification) l AFFIDAVIT OF DAVID P. BLANCHARD County of Charlevoix)
)
My name is David P. Blanchard. I am employed by Consumers Power Company as a Technical Engineer et the Big Rock Point Plant located near 4
Charlevoix, Michigan.
I have a Bachelor of Science degree in Nuclear j
Engineering from the University of Missouri - Rolla. Upon graduation from engineering school in 1971, I joined Consumers Power Company as a Graduate c
Engineer, and I was assigned to the engineering staff at the Big Rock Point plant.
In 1972, I transferred to the Company's corporate headquarters in Jackson, Michigan.
I was assigned to the Reactor Physics Group as an Associate Engineer. I was responsible for nuclear and thermal hydraulic analyses and design of the reactor cores for the Big Rock Point plant.
In early 1976, I was temporarily assigned to Babcock and Wilcox's facilicy in Lynchberg, i
My work included the performance of reactor Tirginia,asaGeneralEngineer.
f physics analyses for input to the FSAR of the Midland nuclear plant.
I returned to the Big Rock Point plant at the end of 1976, and I was assigned i
the position of Reactor Engineer. - In this position, I was responsible for the. performance of analyses of the Big Rock Point reactor to insure conformance 3
with nuclear fuel therma 1' hydraulic limits associated with transient and
/
accident conditions, as well as steady state operation. I was also responsible for implementation of a program for the handling and accounta-bility of nuclent fuel and reactor internals hardware at the Big Rock Point bite. Theseresponsibilitiesincludeddevelopmentofproceduredforhandling nuclear fuel from the time of its arrival at Big Rock, through its receipt inspection and acceptance, insertion into the reactor for power production, transfer to spent fuel storage on depletion of its useful fissile material and ultimately to its shipment offsite to a fuel reprocessing facility.
Development of controls and procedures for full handling during refueling activities and irradiated fuel examination activities by the nuclear fuel vendors were also among my responsibilities. Accounting for the handling and transfer of fuel within the plant is required by 10 C.F.R. Part 70, and as Reactor Engineer, I was responsible in implementing fuel accountability methods which complied with these regulations.
I have been plant Project Engineer for several plant modifications, including replacement of the emergency core cooling sparger in the reactor vessel. In 1980, I was the plant individual assigned the responsibility of working with Science Applications, Inc., in the development of a Probablistic Risk Assessment for the Big Rock Point plant. In March 1981, I moved to the position of Senior Engineer, and I was promoted subsequently to the position of Technical Engineer.
I have been assigned to the spent fuel expansion project since early 1978, and I was involved in developing a method to pecomplish remote actuation of water addition to the spent fuel pool.
Based on my educationa] background and vol. experience, I yelieve I am qualified to answer Christa-Marja Contention 8 and O'Neill Contention IIIE-2, O'Neill Contention IIG(b), and Licensing Board Question 2.
}
I am the author of the testimony addressing Christa4faria Contention 8 and O'Neill Contention IIIE-2, O'Neill Contention IIG(b), and Licensing Bcard
1 Question 2.
I swear that the statements and information contained in this affidavit, the referenced testimony, and in the figures and exhibits attached thereto are true and correct to the best of uy knowledge and belief.
i Q.
I Executed at Charlevoix, Michigan.
MEA k h aMC Y k
Subscribed and sworn to before me this 2nd day of October,1981.
4,-
- W, L j
Ndfary Public M anddfor the State of Michigan and County of Charlevoix q
Eugene A Dziedzic
% commission expires March 6, 1983.
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