ML19290D603

Forwards Amend 6 to Rept, Spent Fuel Pool Mod, Re Application for Expansion of Storage Capacity, in Response to NRC Request for Addl Info.Response to Round 1 Question 10 Is Withheld (Ref 10CFR2.790)

ML19290D603
Person / Time
Site:	Hatch
Issue date:	02/18/1980
From:	Widner W GEORGIA POWER CO.
To:	Office of Nuclear Reactor Regulation
References
NUDOCS 8002220327
Download: ML19290D603 (15)

Text

Georgia Power Company

• 230 Peachtree Street

. Post Office Box 4545 k Atlanta, Georgia 30302 Telephone 404 522-6360 b{

m Power Generation Department Georgia Power February 18, 1980 Director of Nuclear Reactor Regulation U. S. Nuclear Regulatory Commission Washington, D. C. 20555 NRC DOCKETS 50-321, 50-366 OPERATING LICENSES DPR-57, NPF-5 EDWIN I. HATCH NUCLEAR PLANT UNITS 1, 2 SPENT FUEL POOL STORAGE EXPANSION Gentlemen:

Georgia Power Company hereby submits Amendment 6 to the report entitled

" Spent Fuel Pool Modification" which was included as part of our July 9, 1979, application for expansion of the spent fuel storage capacity at Plant Hatch Units 1 and 2. This amendme.nt has been prepared in response to a request by your staff for additional information concerning (1) the response to Round 1 Question 10 that was submitted to you in Amendment 2 by our letter dated September 21, 1979, and (2) the response to Round 2 Question 8 that was submitted to you in Amendment 5 by our letter dated December 31, 1979.

Please note that the response to Round 1 Question 10 contains General Electric Company Class III proprietary fuel design information that has been included as an enclosure to this letter and, therefore, is hereby requested to be withheld from public disclosure. An affidavit providing the basis for the request is attached as part of the response.

Also contained in this amendment are revisions to the report pertaining to the friction coefficients associated with the fuel storage module foot pads. To support our installation schedule an alternate manufacturer had to be selected by General Electric Company to supply the material used for fabrication of the foot pads. Consequently, the lower range friction coefficient for this material is below that considered in the original analysis, i.e. ,

0.132 as compa nd to 0.145. Accordingly, a reanalysis was performed without modifying any of the original methodology d.iscussed in Section 4.0 of the subject report to evaluate the effect of such a change. The results demonstrate, as documented in this amendment that a lower foot pad friction coefficient does not reduce the adequacy of the high density fuel storage system design.

8002220 M

Georgia Power d Director of Nuclear Reactor Regulation U. S. Nuclear Regulatory Commission Washington, D. C. 20555 Page Two February 18, 1980 We regret the need for this late notification of a change in our design bases which was dictated by material availability and trust that your review can proceed in a timely fashion to support our installation schedule.

Yours truly, W. A. Widner Vice President and General Manager Nuclear Generation RDB/TMM/mb Enclosures xc: Ruble A. Thomas George F. Trowbridge, Esquire -

R. F. Rogers III

Edwin I. Hatch fluclear Plant Units 1 and 2 Spent Fuel Pool Modification IfiSERTI0fl IllSTRUCTI0flS Page Instruction 4-2 Replace 4-3 Replace 4-5 Repl ace 4-6 Repl ace 4-7 Repl ace 4-8 Replace 4-9 Repl ace Table 4-1 Replace Table 4-2 Replace Table 4-5 Repl ace Table 4-6 Replace Q10-1 (Response to fiRC Replace letter of 8/24/79)

Second, the derived total mass of the module was used to perform dynamic analysis for the OBE and SSE. As seen in Figure 4-9, for a typical 13 x 13 module, when the added-mass terms from the hydro-dynamic mass effect were included, the fixed base frequency decreased.

Third, both finite-element and lumped-mass models of a module were then developed to provide a basis for delecting simplified module models to be used in the module and support system analysis and module sliding analysis. The finite-element model also was used to obtain the distribution of shear forces in the module plate elements.

Fourth, an eleven-node lumped-mass model was then developed by lumping the tributary module mass to the corresponding node point and ini-tially selecting the stiffness properties based on beam theory. The stiffness properties of this model were based on matching the natural frequencies of the finite element model.

The model is represented as a triangle with three masses. This model preserves the overturning and tilting moment of the rectangularly shaped module. A rectangular model with more mass node would not pro- 4 duce higher effects. Thus, there would be no dif ference in results if a rectangular model was used.

In the nonlinear analysis used to calculat.e the amount cf sliding and tilting, a two-node 1.., 9-mass model was found to. adequately re-present the module and suppoi6 system analyses, since the response of the module support system was shown to be primarily first mode and rigid body motion and both the first mode and rigid body dynamic 3 properties could t e simulated. The lumped mass at the top of the two-mass model wat selected so that the base shear force of the first mode was preserved. The height of the model was selected to preserve the overturning m) ment at the base of the module for both the first mode response and rigid body motion. The summation of tr.e two lower masses and the upper mass used in the model equals the total mass of 4 the module. The distance of the two lower masses was selected to pre-serve the mass mcment of inertia of the module. This ensured that the shear force at tte base was preserved for rigid body motion. Finally, 3 the stiffness of the structural element was selected to preserve the fundamental frequency of the module. The effects of the corner sup-ports were added to the model by including base springs and the final model was used in the sliding analysis. The horizontal spring repre-

, sents the stiffnass of the support pad and the vertical spring repre-sents the stifftass of the fuel support plate, the foot pad, and the support pad.

The mechanism for controlling the shear force in each module is the limiting of the coefficient of friction between the module and the support pad by the selection of a non galling, corros, ion-resistant material with a low coefficient of friction to be used as the module foot pads which are in contact with the stainless steel support pads.

The range of friction coefficient for the selected materials was found to be between 0.132 and 0.203. The friction coeffir.ient between the l6 4-2 Amend. 3 10/79 Amend. 4 11/79 Amend. 6 2/80

stainless-steel support pads and the stainless-steel liner is at least 0.349. This difference insures that sliding will occur between the foot pad and the support pad, and not between the support pad and the floor liner (References 8 and 9). l4 The sliding analysis was done using the two-dimensional, non-linear DRAIN-20 and SEISM computer codes. DRAIN-2D was originally developed at the University of California at Berkeley; SEISM was developed by GE. Both computer codes have been design reviewed and meet NRC-QA requirenents. Sliding and overturning of the module were studied for the SSE and OBE conditions. All of the modules were found to be stable under the worst postulated seismic loading conditions, and the minimum 2-inch clearance between modules precludes contact during a seismic event.

4.2 Stress Analysis The HDFSS module has been designed to meet Seismic Category I require-ments. Structural integrity of the rack has been demonstrated for the load combinations below using linear elastic design methods.

Analysis was based upon the criteria and assumptions contained in the following documents:

a. ASME Boiler and Pressure Vessel Code Section III, Subsection NF.

b. USNRC, Regulatory Guide 1.92, Combining Modal Responses and Spa-tial Components in Seismic Response Analysis.

c. Hatch 2 Final Safety Analysis Report, Seismic Design Criteria.

OBE - Operating Basis Earthquake SSE - Safe Shutdown Earthquake

d. Light-Gage Cold-Formed Steel Design Manual, 1961 Edition, American Iron and Steel Institute.

Acceptance criteria were based on:

a. Normal and upset (0BE) Appendix XVII, ASME,Section III.

b. Faulted (SSE) Paragraph F-1370, ASME Section III, Appendix F.

c. Local buckling stresses in the spent fuel storage tubes were calculated according to " Light-Gage Cold-Formed Steel Design

. Manual" of American Iron and Steel Institute in lieu of Appendix WII, ASME,Section III, because of its applicability to these Iight gaoc tubes. Only the strength of the outer wall thickness of 0.090 inch nominal is considered in the stress calculations.

4-3 Amend. 4 11/79

4 Thermal loads were not included in combinations because the design of the rack makes them negligible; i.e., the rack is not attached to the structure and is free to expand or contract under pool temperature changes. Assuming the boundaries of the module are completely fixed and the module is not allowed to expand, the maximum thermal stress between loaded and unloaded cells is less than 6,400 psi. This is well within the allowable compressive stress in the tube wall. Further-more, according to ASME Section III, Subsection NF, Paragraph NF-3230, 4 Appendix XVII Article F-1370, thermal stresses need not be considered in the stress calculation but only in the buckling analysis for the SSE condition. This is consistent with industrial practice for piping stress analysis where thermal stress is treated as secondary stress.

Therefore, under the cooling water flow conditions in the modules, the heat rise in the wall of a loaded storage tube caused by gamma heating is no more than 5 F and the maximum water temperature rise from bottom to top of a storage tube is 19 F. Thus, the maximum tempegature grad- 3 ient between a loaded and an empty cell is no more than 24 F, as is explained in Section 8.5. Temperature-induced stresses are not addi-tive from module to module because each module is independent of the others.

Stress analyses were done for both OBE and SSE conditions, based upon the shears and moments developed in the finite-element dynamic anal-ysis of the seismic response. These values were compared with allow-able stresses referenced in ASME Section III, Subsection NF (Table 4-1). Values given in Table 4-1 are based on the maximum stresses l1 calculated for all module sizes. A dynamic load amplification factor of 1.514 has been apslied to stresses due to the horizontal seismic 6 load to account for the effects of impact between the fuel and the module. A deri/ation of this factor is given in Section 4.3. Addi-tional analyses were then performed to determine the dynamic fre-quencies, earthquake loading reactions, and maximum amount of sliding.

The stability of the modules against overturning was also checked and I they were found to be stable. Those values are summarized in Table 4-2.

The force path in the module caused by a horizontal earthquake is shown schematically in Figure 4-10. This figure shows the path of the horizontally induced earthquake fuel element inertial forces from the fuel element to the module support pads. Part of the fuel bundle inertial forces induced by the motion of the module are transferred either through the water or directly to the tube walls perpendicular to the direction of motion (Point 1 in Figure 4-10). These walls then transfer the forces to the side tube walls, which carry the forces down the walls and into the fuel support plates (Point 2). The por-tion of the fuel bundle load which is not transferred to the fuel tube walls is transferred directly to the fuel support plate at Lhe point where the lower end fitting of the fuel bundle is supported vertically (Point 3). The fuel support plates, acting as a relatively rigid diaphragm, transfer the in plane shear forces to the long casting which then transfers the shear forces to the module base assembly 4-5 Amend. 1 7/79 Amend. 3 10/79 Amend. 4 11/79 Amend. 6 2/80

plate (Point 4). The forces are carried in the module base assembly (Point a) until they are ultimately transferred to the foot pad and to the support pad and the pool slab (Point 6).

The vertical forces caused by earthquake and gravity loads become axial forces in the foot pads. The critical location for the com-pression forces from the foot pads is in the long castings and tubes directly above the foot pads. For stress analysis purpose, these compression forces are considered to be resisted by four fuel tubes sitting directly above the support pad.

Fuel assembly drop accidents were analyzed using analytical methods in accordance with the " Operating Technical Position for Review and Acceptance of Spent Fuel Storage and Handling Applications". In estimating local damages in the module, the maximum strain energy resulting from plastic deformation is equated to the maximum potential energy of the fuel. Energy dissipation attributable to the viscosity 3 of the water and plastic deformation of the fuel bundle was ignored for conservative results. The stainless steel for the module is assumed to exhibit a bi-linear hysteresis relationship, with yield stress and ultimate stress as the two control points. The results are summarized in Table 4-3.

Also evaluated was the damaging effect of a fuel bundle drop through an empty storage position along the outer rows of.the module, impart-ing the base frame. It was determined that the fuel bundle will not possess enough energy to perforate the 1-inch thick base frame. The resulting configuration of the module will be adequate to maintain the fuel in a safe condition. This case is less critical than the cases discussed in Table 4-3.

The loads that may be carried over the spent fuel pool are listed in Table 4-4. A free fall of these loads onto the fuel pool liner plate and storage racks was evaluated. It was determined that a fuel assembly drop causes the most damaging effect due to its weight and geometrical configuration. Also, none of the other loads can be lifted to a posi-tion higher than that of a fuel assembly above the liner plate and storage racks. Regarding the integrity of the liner plate, the evalu- 4 ation demonstrated that the energy developed by a freely falling fuel assembly would not cause perforation (Reference 7). A free falling fuel assembly dropping from a height extending 27 inches above the height of a module with 0 ft/sec initial velocity is calculated to have a final velocity of 26.5 ft/sec when it comes in contact with the slab liner plate after traveling through the water. The re-quired steel plate thickness to just perforate, based on this velocity, is less than the liner plate thickness that is provided for the pool slab. The presence of concrete below the liner plate was conservative-ly neglected in the computation. Regarding the integrity of the fuel and storage racks, the consequences of drapping any of the items listed in Table 4-4 are no more severe than that of the fuel assembly drop accidents summarized in Table 4-3. The provisions employed to prevent movement of heavy objects over the spent fuel pool are discussed in Section 11.0.

4-6 Amend. 3 10/79 Amend. 4 11/79

The HDFSS design does not require any different fuel handling pro- 3 cedures from those discussed in the Unit 1 and Unit 2 FSAR The loads experienced under a stuck fuel assembly condit. ion are less than those calculated for the seismic condition and have therefore not been included as a load combination.

4.3 Fuel Bundle / Module Impact Evaluation An analysis was performed to evaluate the effect of an impact load that is possible because of gaps between the fuel bundle and the fuel storage modul e. In the seismic analysis for the Hatch high density spent fuel storage module (results in Table 4-2), gaps were not considered and the fuel bundle was treated as an integral part of the module in addition to the hydrodynamic mass due to surrounding water.

A gapped element mcdule was prepared to study the effect of impact loads on the module. This model is shown in Figure 4-11. The distinct feature of this model is that the fuel bundle is separated from the. module and is free to vibrate within the confines of the storage position in the module.

The fuel bundle is pinned supported at the base and the entire module is submerged under water and free to slide. For comparison purposes re-garding the impact load effect, a lumped element model was also con-structed. The lumped element model is identical to the gapped element model shown in Figure 4-11 except that the gaps between the fuel bundle and the module are ignored.

The objectives of this evaluation are:

5

a. to assess the difference in maximum internal forces in the module as determined from a gapped element model and a lumped element model, and

b. to assess the effect of impact loads on the maximum sliding dis-placement of the module.

To evaluate gap effects on rigid body displacements, the two models were subjected to a constant 1.0g base acceleration for a period of 0.8 sec-onds. This acceleration was applied for two cases, corresponding to fric. tion coefficients of 0.132 and 0.2. The use of a constant 1.0g base 16 acceleration was mandated by the lack of a definitive time history to use in conjunction with rigid body displacements. Gap effects on internal forces were evaluated by subjecting both models to the Hatch time history.

This was done for three cases: p = 0.132, p = 0.2, and p*= (fixed base). 16 The results of these analyses are presented in Tables 4-5 and 4-6 for rigid body displacements and internal forces, respectively.

Table 4-5 shows the displacement ratio between the gapped and the lumped element model. It indicates that there are no significant differences between the rigid body displacements as determined from the gapped and lumped element models for both p = 0.132 and p = 0.2. Thus, it can be 16 concluded that gap effects on the rigid body motions can be neglected and that the results provided in Table 4-2 are adequate for design purposes.

4-7 Amend. 310/79 Amend. 5 12/79 Amend. 6 2/80

Table 4-6 indicates that the internal forces (or spring loads) in the module determined from the gapped models are significantly less than the corresponding forces in the lumped models for the two cases p = 0.132, and 15 p = 0.2. For the case where p+= this situation is reversed, however, and .

S the internal force in the gapped model exceeds the internal force in the 6 lumped model. Thus it can be concluded that where rigid body motion is -

permitted and friction forces are within the range of interest, the in-ternal forces are conservatively determined from the lumped model. The ratio between the spring forces in the gapped model and the lumped model (fixed base case shown in Table 4-6) is treated as the dynamic load ampli-fication factor and used in the stress analysis comparison in Table 4-1. 6 This approach is conservative since results for the sliding model indicate that there is a reduction in internal stresses for the gapped element model. -

4.4 Effects of Increased Loads on the Fuel Pool Liner and Structures The Unit 1 and Unit 2 spent fuel pool structure and liner plate have adequate capacity to carry the increased loads imposed by the new high 4 density spent fuel storage racks.

The spent fuel pool structure for each unit was evaluated for new loads based on the following criteria:

1. " Code Requirements of Nuclear Safety Related Concrete Structures", The ACI 349-76 Code.

2. USNRC Regulatory Guide 1.142.

3. USHRC Standard Review Plan, Section 3.8.4.11.

4. USNRC Operating Technical Position for Review and Acceptance of Spent 5

Fuel S'.orage and Handling Applications.

3 Based on the above criteria, the following is a listing of the primary loads that were considered in the structural evaluation:

1. The dead weight of the structural elements (D).

2. The live loads acting on the structural elements (L).

3. The hydrostatic load due to the water in the pool (F).

4. A three component OBE seismic load (Eg ).

5. A three component SSE seismic load (Ess)*

6. A thermal loading based on normal operating conditions pogl water temperature of 150 F and ambient air temperature of 90 F (Tg ).

7. A thermal loading based on accident conditions pool wa er tem-perature of 212 F and ambient air temperature of 90 F (T ).

4-8 Amend. 3 10/79 Amend. 4 11/79 Amend. 5 12/79

8. A thermal loading based on normal operating conditions poo water temperature of 150 F and ambient air temperature of 110 F (T ). 3

9. A thermal loading based on accident conditions pool watgr tem-perature of 212 F and ambient air temperature of 110 F (Ta )*

The following seven loading combinations that produce the most severe loading to this type of structure were used in the evaluation:

1. U = 1.4 (D) + 1.7 (L) + 1.4 (F) + 1.9 (Eg )

1 2.

U = (D) + (L) + (F) + (Ess) + (Tg )

2 3.

U = (D) + (L) + (F) + (Ess) + (T )

4. )

U = (D) + (L) + (F) + (Ess) + (

5. )

U = (D) + (L) + (F) + (Ess) +

1

6. U = 0.75 [1.4 (D) + 1.7 (L) + 1.4 (F) + 1.9 (Eg ) + 1.7 (T )3 2

7. U = 0.75 [1.4 (D) + 1.7 (L) + 1.4 (F) + 1.9 (Eg ) + 1.7 (T )3 5

A three-dimensional mathematical model was developed for each spent fuel pool structure. Each mathematical model is composed of plate /shell ele-ments, beam elements, truss elements, and boundary elements to idealize the existing structure. Structural properties for the elements were selected based on insitu conditions.

The analysis was performed using a computer code name "BSAP". This com-puter code is a modified version of SAP IV which is in the public domain.

The analysis was broken into two parts. First, each structure was ana-lyzed for load combination 1 using gross concrete structural properties.

This will verify that each structure will carry the mechanical loading and place an upper bound on the structure's stiffness. Second, each structure was analyzed for all seven loading combinations listed above using cracked concrete and reinforcing steel structural properties. This will verify that each structure will carry the mechanical as well as the thermal loading combinations, placing a lower bound on the structural stiffness.

The results of the analyses, forces, moments and shears for each loading combination and analysis condition were evaluated based on " strength criteria" for each of the fuel pool elements.

The evaluation shows that the fuel pool structure for each unit meets the design criteria for the conditions stated.

4-9 Amend. 3 10/79 Amend. 5 12/79

TABLE 4-1 Comparison of Calculated Stress vs Allowables (psi)

OBE Condition SSE Condition Location / Type Calc Stress A110wables l Calc Stress Allowables l Tube wall shear 6,040 11,000 7,400 22,000 Tube wall compression 7,180 14,880 8,400 29,760 Tube weld throat shear 8,540 11,000 10,400 22,000 6

Angle, weld throat shear 8,540 11,000 10,400 22,000 Casting wall shear 6,240 11,000 9,210 22,000 Casting wall compression 11,600 16,500 12,500 33,000 Casting base weld shear 4,920 11,000 7,250 22,000 l6 Support plate weld throat 3,400 11,000 .5,700 22,000 shear Closure plate compression 6,570 14,880 7,460 29,76'O Closure plate shear 6,840 11,000 8,450 22,00p 6 Closure plate weld shear 9,120 11,000 11,300 22,000 Corner tube local compressive - - -

6,900 17,224 stress check for local buckling 1

Allowable stresses referenced in ASME Section III, Subsection NF Amend. 1 7/79 Amend. 6 2/80

TABLE 4-2 DYNAMIC FREQUENCIES, EARTHQUAKE LOADING REACTIONS, AND MAXIMUM AMOUNT OF SLIDING Module Size Direction Fundamental Frequency (Hz) Max. Reaction (lbs) Max. Sliding (in)

Il x 13 North-South 9.6 95,000 0.62 (Unit 1 Only) East-West 12.0 88,000 0.67 13 x 13 North-South 10.9 98,000 0.79 (Unit 1 Only) East-West 10.9 98,000 0.79 13 x 15 North-South 9.9 109,000 1.02 6 a

(Unit 1) East-West 11.6 110,000 .98 13 x 15 North-South 11.3 110,000 0.87 (Unit 2) East-West 9.4 105,000 1.14 1 13 x 17 North-South 9.5 118,500 1.72 (Unit 1) East-West 12.1 100,500 1.22 6

13 x 17 North-South 11.7 110,500 0.88 (Unit 2) East-West 9.0 110,500 1.75 13 x 19 North-South 12.1 122,500 0.91 (Unit 2 Orly) East-West 8.5 121,000 1.40 15 x 19 North-South 11.2 128,000 1.02 (Unit 1) East-West 9.6 137,000 1.50 l6 15 x 19 North-South 11.3 128,000 1.02 (Unit 2) Eas t-Wes t 8.8 137,000 1.50 l6 Amend. 1 7/79 Amend. 6 2/80

TABLE 4-5 Normalized Rigid Body Displacements Of Lumped And Gapped Models Friction Coefficient Gapped Element Model/

p Lumped Element Model 0.132 1.01 15 0.2 1.02 Amend. 5 12/79 Amend. 6 2/80

TABLE 4-6 Spring Forces In Lumped And Gapped Models Friction Coefficient Gapped Model Lumped Model (p) Force (1bs.) Fort .(1bs.)

0.132 0.551 x 10 5 0.852 x 10 5 l6 0.2 0.592 x 10 5 1.20 x 10 5 p+= 2.09 x 10 6 1.38 x 10 6 (fixed base)

Amand. 5 12/79 Amend. 6 2/80

QUESTION 10 For our evaluation of the difference between the maximum calculated k, of 0.87 given in your submittal and the maximum actual k that might occur in the spent fuel pool, the following information should* N provided:

a. The quantity and distribution of the uranium-235 in the fuel pool storage lattice calculation for this maximum k,;

b. The quantity and distribution of gadolinium-155 and gadolinium-157 in the fuel pool storage lattice calculation for this maximum k,;

c. The quantity and distribution of the fission products and actinides in the pool storage lattice calculation for this maximum k,;

RESPONSE

The response to this question contains General Electric Company Class III proprietary information which was provided by our letter dated 6 February 18, 1980.

Amend. 2 9/79 Amend. 6 2/80 Q10-1