ML20043A553

Forwards Response to Outstanding Issues & Request for Addl Info from 891128-30 Advanced BWR Seismic Design Audit at GE Ofcs in San Jose,Ca.Info Resolves Sections 2 & 3 to Draft SER & Action Items

ML20043A553
Person / Time
Site:	05000605
Issue date:	05/16/1990
From:	Recasha Mitchell GENERAL ELECTRIC CO.
To:	Chris Miller NRC OFFICE OF INFORMATION RESOURCES MANAGEMENT (IRM), Office of Nuclear Reactor Regulation
References
EEN-9021, MFN-051-90, MFN-51-90, NUDOCS 9005220234
Download: ML20043A553 (90)

Text

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r U. . GE Nuclear Energy sene or tem corwny (3 ' ' '""' '** S* " ** C' *5

l" ) May 16,1990 MFN S 05190 Docket No. STN 50-605 EEN 9021 Document ControlDesk U.S. Nuclear Regulatory Commission Washington,D.C. 20555 Attention: Charles L Miller, Director Standardization and Non Power Reactor Project Directorate

Subject:

Responses to ABWR Selsmic Design Audit Open Issues and Information Requests

Reference:

Dino C. Scaletti to Patrick W. Marriott, ' Meeting Suramary of the ABWR Seismic Design Ardit" May 1990 Encic, sed are thirty four (34 additional information identif)ied at the November 28 30,1989 ABWR Seismie at the GE offices in San Jose, California. Attachment 1 provides the resolution to Section 2 and 3 Draft SER open items and Attachment 2 provides the response to the request for additional L

(q y- information and GE action items.

o, It is intended that, where applicable, GE will amend the SSAR with these responses in a future amendment.

Sincerely, L Q .C. h R. C. Mitchell, Acting Manager i

Regulatory and Analysis Services M/C 382, (408) 925 6948 cc: F. A. Ross (DOE)

D. C. Scaletti (NRC) t, D. R. Wilkins (GE) l J. F. Ouirk (GE)

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1 RESPONSES TO DRAFT SER OPEN ITEMS

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r-DRAFT SER OPEN ITEMS

[\~-} NOT RELATED TO SEISMIC DESIGN OF STRUCTURES 2.5.4.1 The GE staff confirmed that the minimum depth of embedment will be 85 ft for the Reactor Building even if competent rock is met with at ground surface. In view of this, the statement,in the SSAR (Appendix 3A, Section 3A.3.4) that the minimum soil depth of 25.7 m (85 ft) is considered for the HR and EH profiles needs to be amended appropriately, since the HR (hard rock) and EH (extra hard rock) profiles do not have any soil depth at all as seen from Table 3A.3.2.

Resolution Section 3A.3.4 and Table 3A.3-6 is revised accordingly.

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ABWR n4 enc 4s Standard Plant mA where in order to evaluate the effects of water (7 table location variation (assuming no soil V Ko = coefficient of horizontal carth pressure failure) on structural response, three water at rest at depth y table locations are considered, namely, the bi = height of the column of unsaturated soil high, intermediate, and low water tables. The above depth y high water table is the base case which is h2

= height of the column of saturated soil located at 0.61 m (2 ft) below grade. The low above depth y water table is taken to be 25.7 m (85 ft) which is at the base of the reactor building Using 0.55 for Ko, Eq. (3A 4) becomes foundation basemat. The intermediate water table is assumed at 12.2 m (40 ft) below grade

• m = 0.7 ( r h1 + r'h2 ) (3A 5) which is at about the midheight of the reactor building embedment. Since the ground water may Since the shear modulus is a function of the have more pronounced effects on soft sites, the effective confining pressure which is, in turn a UB profile with 45.7 m (150 ft) depth of soil function of the soil unit weight, the location of deposit is investigated for all three water the ground water table affects the shear modulus tables defined above. The average shear wave and the wave propagation velocity in the soil, velocities of the UB profile having the water Since the water is essentially incompressibic and table located at 12.2 m (40 ft) and 25.7 m (85 cannot resist shear, the presence of the water it) below grade are shown in Table 3A.3 5. The has an important influence on the compressional water table level for other site conditions is wave velocity but produces only a minor effect on based on the basic high water table case which ,

the shtar wave velocity, is at 0.61 m (2 ft) below grade.

The effect of ground water on the soli 3A.3.4 Summaryof Site Conditions properties is considered using the following procedure: The above discussions cover the range of site-parameters in terms of soil deposit depth, soil l

O~ (a) Perform one dimensional deconvolution profile and properties, and water table i analysis for the horizontal excitation location. Based on the GESSAR experience (Refs.

component to obtain strain compatible shear 1 and 2) that the shallower soil depth in '

l modulus, G, corresponding to the induced general resulted in higher structural response, strain level. The corresponding shear wave all ujocity profiles, except the HR and EH velocity, V s, is then computed as

'shallo}w soil case, and only a limited soil (pT ,

V3 = / G/r (3A 6) velocity profiles need be considered for other depths. For the deeper deposits of 61 m (200 l (b) Co=pute the corresponding compressional wave ft and 91.5 m (300 f ) in depth, it is ye1ocity, V p, using the ioilowing f su)fficient to consider only the UB profile si equation: its shear wave velocity profile has the largest l variation with depth. Consequently, the effects Vp of variation in soll properties with depth is l

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=

V /g#

s (3A 7) accounted for more representatively. The minimum soll dep@jof 25.7 m (85 ft) is where p is a Poisson's ratio. The lower bound considered for the 7 UB VP4?" trMHyofiles .y of V p or f the submerged soil below water table which adequately cover the range offroilles

. is tot compressional wave velocity of water taken , considered. As mentioned before, the water to be 1463 m/sec (4800 ft/sec). When the l table variation is taken into account by the UB

! computed Vp of soil is sinaller than 1463 in/sec, j profile for the shallow deposit case. A total adjustment of soil Poisson's ratio is required so of 14 site conditions are selected for seismic j that she Vp of soil is equal to 1463 m/sec. ( analysis and are summarized in Table 3A.3 6.

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ABWR nan =xt Standard Plant m'A Table 3A.3 6 -

O Site Conditions Considered (Water Table at .61 m (2 ft) depth except noted)

SOIL PROFIG SHEAR WAVE YE!4 CITY PROFILE DEPTH M.En ill YI2 %I2 iT.i 321 %I5 EB EH w +n 25.7 (85) UB1D85 - - VP4D85 - - HRD85 EHD85 45.7 (150) UB1D150 VP2D150 VP3D150 VP4D150 VPSD150 VP6D150 - -

UB2D150' UB3D150

61 (200) UB1D200 - --- -- -- ---- -- ---

91.5 (300) UB1D300 - - --- - - -- ..

• Water Table at 12.2 m (40 ft)

" Water Table at 25.7 m (85 ft)

• H Namm, Thu n m Ja, >., rse/ s d h an.( L e ]<,e se.sd .Lp(f, g System:

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m: Velocity profile identifier Denn Soil deposit depth in feet

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j. O NOT RELATED TO SEISMIC DESIGN OF STRUCTURES

r, 3,8.1 and 3.8.2 GE will provide-the design details of the containment steel

' liner, j Resolution  !

See resolution to part (a) of item II.3

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REQUEST FOR INFORMATION AND ACTION ITEMS I O ,

I.. REQUEST OF ADDITIONAL INFORMATION~ '

L OR CLARIFICATION II. ACTION ITEMS FOR GE i

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P fr I. REQUEST OF ADDITIONAL INFORMATION OR CLARIFICATION I.1 Section'3.7.1 GE should provide.the justification of doubling the OBE responses to cover'the SSE in the seismic design.

Resolution ,

To demonstrate that the 0.3g SSE floor response spectra can be conservatively taken to be two times the OBE responst spectra (Note: the peak ground acceleration for OBE is assumed to be 0.15g), specific SSE SSI analyses are

. performed for two representative sites covering a range of site conditions.

The soil site UB1D300, which is softest soil profile among all-sites considered, is chosen since it may result in most significant effect on' structural response in terms of resonant frequencies. The rock site HRD85 is chosen since

-it generally governs the structural responses in terms of-

. amplitudes.

1 In order to satisfy the SRP 3.7.1 enveloping requirements of the RG 1.60 design spectrum of-7% damping for concrete l structures unde" SSE excitation, the synthetic input time l

-%) history is.incr21ced by 7% in amplitudes. This results in

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an actual 0.3210 peak ground acceleration for the 0.3g SSE

. definition.

The strain-compatible soil properties under SSE excitation for the'UB1D300 soil profile are calculated using the SHAKE l computer program. The strain-compatible P-wave velocities at layers below the water table, which is at 2 ft below grade-in this case, are kept 4800 ft/sec minimum. For the HRD85 rock site.the strain-compatible properties are essentially unchanged from their initial values.

Therefore, a uniform shear wave velocity of 5000 ft/see is directly used in the SSI analysis.

The SSE-SSI analyses using the SASSI 2D option are performed.for the reactor building situated at these two sites for the horizontal X direction response calculations. The corresponding OBE analysis cases are C13X for UB1D300 and C10X for HRD85.

The calculated SSE response spectra of 2% damping are compared to two times OBE spectra at four key locations in Figs. 1 through 4 for the UB1D300 soil site and Figs. 5 through 8- for the HRD85 rock site. As expected, a minor .

downward shift of the fundamental SSI frequency is observed

-s at superstructures (Figs. 1 through 3) for the UB1D300 soil

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f site under SSE excitation. This has no effect on design since it.is within the broadening range of the 2*OBE spectra. For the HRD85 rock site no frequency shift is I.1-1

/f hs c noted. In examining Figs. 1 through 3 and 5 through 7, the 280BE responses at superstructure locations bound the actual SSE responses at dominant resonant frequencies (including the ZPA). The slightly higher SSE responses at frequencies below the fundamental SSI frequency at these locations-are resulted from the 0.321g input motion as opposed to 0.3g for 2*0BE. This difference in the magnitudes of the input motion is also-the major cause for the slightly higher SSE responses at the basemat (see Figs. .

4 and 8). In no case however is the actual SSE response greater than the broadened and smoothed SSE (i.e., 2*OBE) site-envelope spectra.

In conclusion, it is adequate to use 2*0BE site-envelope

. loads for the SSE design of the ABWR standard plant.

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EIG. 3 COMPARISON OF 2*0BE AND SSE RESPONSE.(2% DAMPING) FOR UB1D300

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FIG. 5 COMPARISON OF 2*OBE AND SSE RESPONSE,(2% DAMPING) FOR HRD85 AT RB TOP-(NODE 95)-

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D. I. REQUEST OF.ADDITIONA!., INFORMATION OR CLARIFICATION  :

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' I.2.'Section 3.7.1-

- GE'should provide results of the seismic. analyses and structural design details-for the control' building and the-radwaste building substructures.

- tSee FAX from J. Fox,.GE to--D. Scaletti, NRC dated 8/31/89)

- Resq1ution To be provided by 6/30/90.

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I. REQUEST OF ADDITIONAL INFORMATION OR CLARIFICATION

[)DI I.3 Section 3.8.1 &~3.8.2 GE should provide details of the containment steel liner design. GE should also provide breakdown of the stress components at the critical locations in both reactor building and containment building.

Resolution- ,

a) Subsection 3.8.1.1.2 is updated and a typical section of I the containment liner-plate and anchor design is shown in Fig.-3H.3-6.Lin Appendix 3H.

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- SECTION 3H.3 c' ILLUSTRATIONS ,

Figure Etig East 3H3-1 Containment Structure Wall Reinforcement 3H3-50 3H3 2 Containment Structure Opening Reinfcrcement 3H3-51 3H3-3 Containment Structure Opening Reinforcement 3H3-52 3H3-4 Containment Structure Top Slab Reinforcement 3H3-53 3H3-5 Reactor Building Foundation Reinforcement 3H3-54 Tyna A suL,, . i ,,,,, pg,,( ,

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3H3-7 Diaphagm Floor Reinforcement 3H3 56 4

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' MM '23A6100AE Standard Plant REV B j 3.8.1.1.2 Containment IJoer Plate Boiler and Pressure Vessel Code, Division 2, y Section III, Subsection CC. I_

H f, j P The internal surface of the containment is l lined with welded steel plate to form a leaktight

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i 3.8.1.23 General Design Criteria, Regulatory l barrier. The liner plate is fabricated from Guides, and industry Standards ]

,, carbon steel except that stainless steel plate or

,; clad is used on wetted surfaces of the suppres. (1) 10 CFR 50, Appendix A ' General Design

.I p ~~ sion chamber. TE: t::::: Of d : ::: b!::; ;;;;! ' Criteria for Nuclear Power Plants' Criteria l'  :: .. % :; L. d.e . 77. ..... d. 6. .. 6 m. 1, 2, 4,16 a nd 50. Conformance is -

E T:g ;; 3M.34 The liner plate is stiffened by discussed in Section 3.1.

use of structural ecctions'and plates to carry the design loads and to anchor the liner plate to (2) U.S. Nuclear Regulatory Commission (NRC)'

as stm '~iFconcrete/The liner plate is thickened 1o. Regulatory Guides. Regulatory Guide 1.136. i Materials, Construction and Testing of q M-i'3gy6-cally majorand additional structural anchorage attachments suchisasprovidcd at penetration Concrete Containment. '

L sleeves, structural beam brackets, the RPV pedes-

-1 tal and the SRV quencher support connection to (3) Industry Standards ,

", the basemat, and the' diaphragm floor connection Nationally recognized industry standards

! to the containment wall. such as thue published by the American Society for testing and Materials (ASTM) and The erection of the liner is performed using the American National Standards Institute:

standard construction procedures. The (ANSI) as referenced by the Applicable ' '

containment wall liner and top slab liner are Codes, Standards, and Regulations are used, used as a form for concrete placement. The liner on the bottom of the suppression chamber and 3A1.2.4 Containment Boundary lower drywell is placed after the foundation mat concrete is in place. The jurisdictional boundary for application -

of Section III, Division 2 of the ASME code to .

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3.8.1.2 Applicable Codes, Standards.and Specifications the concrete containment is~ shown in Figure 3.812. The boundary extends to the:

The design, fabrication, construction, (1) Outside diameter of the containment wall' testing, and inservice inspection of the contain- from the foundation mat to the containment ment conforms to the applicable codes, standards, top slab.

specifications, and regulations listed below, except where specifically stated otherwise. (2).The foundation mat within the outside .

Jiameter of the containment wall.

3AL2.1 Regulations (3) The containment top slab from the drywell (1) Code of Federal Regulations, Title 10, head opening to the outside diameter of the Energy, Part 50, ' Licensing of Production containment wall.

and Utilization Facilities."

(4) The intersection of the RPV pedestal on top

-(2) Code of Federal Regulations (CFR), Title 10 of the basemat.

- Energy, Part 100, Reactor Site Criteria, (10 CFR 100), including Appendix A thereto, (5) The intersection of the diaphragm floor with=

" Seismic and Geologic Siting Criteria for the containment wall. .;

Nuclear Power Plants."

The concrete containment pressure boundary is 3.8.1.2.2 Construction Codes of Practice limited to the cylindrical wall of the drywell and wetwell, and the drywell top slab.

American Society of Mechanical Engineers (ASME)

Amendment 4 3.82 1.3-3

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h b) Stresses

-are not meaningful due to individual loads were for reinforced not calculated concrete since structures, they especially with significant thermal loads.

The stress analysis approach used for the ABWR containment /

reactor building is shown schematically in Figure 1. For individual loads such as dead load (D), live load (L), LOCA pressure loads (Pa), hydrodynamic lords (CO and SRV), seismic loads (OBE and SSE) and LOCA thermal leads (Ta), forces and moments are obtained at selected sections (Figure 3.8 - 14 in Section 3.8) based on elastic analysis using computer program-Stardyne. The hydrodynamic loads are applied as equivalent static pressure loads.

Using "Stardyne Post" which is a post processor program, element forces and moments due to all pertinent loads including hydrodynamic (but excluding seismic and thermal) for each load combination are first combined. Element forces-and moments due to seismic loads and thermal loads are not combined but treated separately. The forces and moments of the sub-combined loads (excluding seismic and thermal) and individual seismic and-thermal loads are included in Appendix 3H at Selected sections in the containment and internal structures. The corresponding data at Section #10 for the reactor building is shown in Table-1.

() To find stresses in rebar and concrete, computer program CECAP-has been used. The input to CECAP consists of rebar ratios, material properties and element geometry at the section under consideration.together with-the forces and moments for a load combination. The program accounts for self relieving of thermal moments due of concrete cracking, and gives stresses and strains in rebar and concrete at that section. The seismic forces and moments are combined lineraly with the others.in the most conservative manner so as to produce maximum stresses in the rebar and maximum compressive stresses in concrete for.the load combination under consideration.

The total stresses at selected sections in the containment and internal structures are included in Appendix 3H for the load combinations considered. For the reactor building at Section

1. 10,. total stresses resulting from the element forces and moments due to contributing loads shown in Table-1, are shown in Table-2, for the load combination (D+L+Pa+Co+SRV+SSE+Ta).

4

3. 3- 6

. I

' STARDYNE sE0 METRY-MAT. PROPERTIES STIPPNESS MATRIX DECOMPOSITION ECOMPO ED

' APPL! CAT 10N OF STARDYE 10 LOAD CASES 4 '

STATIC ANALYSIS TRIx

/ SELECTION OF STARDYNE POST gggng, ELEMENTS FOR 4 COMBINATION OF I 1 PORCES 8.

LOAD COMB. LCADS MOMENTS RESAR AMOUNT CECAP- 1. D+L+P,+c0+sav

. CONC. SECTION  % , 2. E LINER PLATE ETc, CCD881DERIM M I' IA-STRESS RELIEP DUE TO CDCRETE CRACX!M 1P REBAR STRESSES CONCRETE STRESSES LINER PLATE STRAIN FIGURE 1 Analysis Flow Chart 9 ,

T.3-7

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t if 3 76 z/

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O

'INEE 2 REBAR AND OEDGE'IE SIRESSES IN 'ITE R/B EXITRIOR NAIL AT SECTIGES 10 (EUBENT 601)

! 11E 'IO IDADING CIMBINATION (D&IAPa+00&SRV+SSE+Ta)

(-

l

neinforcirn Steel concrete '

g Calculated Stresses (ksi) Allowable Calmlated Allowable Inside axm outside Face Shear Stress Stress Stress Y Vert. Horiz. Vert. Horiz. Ties (ksi)' (ksi)' (ksi) to 16.39 16.96 13.80 21.12 16.76 54.0 -0.26 -3.4 1

f i

't AS4R-S -l s

1

-. ~ .- - -.

1 I.-REQUEST OF ADDITIONAL INFORMATION'OR CLARIFICATION L 1 lL I.4 Section 3A.1 Concerning site condition (7 in Section .I 3A.1 GE needs to address site conditions with soil deposit depths between 0 and 85 ft.

L Resolution For soil sites, the minimum soil depth considered for the L ABWR standard plant seismic design is 85 ft-which is the.

embedment depth of the reactor building. For_ rock sites, zero soil depth is considered, i.e., the~ building embedded to the same depth is situated'at a uniform rock halfspace.

• From the site enveloping standpoint, these two extreme-cases should adequately cover any site with specific-soil depth falling in between.

To demonstrate this point, free-field response analyses using the SHAKE computer program are performed for the UB profile (soft soil) with the intermediate depth of' soil at 20,-40, and 60 ft. The resulting free-field horizontal

motions at depth 20, 40, 60, and 85 ft along the embedment are compared to.those of the EMD85 rock site (zero soil)

L . .

.and UB1D85 soil site (85 ft soil)Lin Figs. 1 through 4,

['] respectively. As'can be seen from these figures, the

(/ free-field responses of the uniform rock case (zero soil deposit) bound those of all soil sites with_ varying soil depths up to 85 ft. ,

The:results of the SSI case C8X for UBlD85 ri'v of 85 ft

' soil depth and case C16X for END85 rock site of zero soil depth are. compared to demonstrate the effect of soil depth on SSI response. The floor response spectra of 2% damping

-for these two cases are shown in Figs. 5 through 8 at four key locations in the reactor building complex. These figures clearly show that the rock site, which has lesser free-field motion attenuation, results in higher structural ,

responses than the 85 ft soil depth case for the same building embedment. -Therefore, the seismic response of the ABWR standard plant located at sites with soil depth within 85 ft is-expected to be bounded by the all-site enveloping response used for design, when all site conditions defined in'Section 3A.1 are satisfied.

In summary, it is not required to add a soil depth limitation in site condition #7 in Section 3A.1. '

I. 4 -l

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FIG. 8 RESPONSE SPECTRA (2% DAMPING) AT RB TCP (NODE 95)

,. .n ...v - , ....a.. -. ,n. - ~ - - . . , . ,- , , . . . , , - - . - - - - _ - - - _ . - - - . _ _ - - - -

I. REQUEST OF ADDITIONAL INFORMATION OR CLARIFICATION O I.5 Section 3A.3 GE should address the effect of the potential separation of side soil from the buildings and the associated dynamic structural response. This should be done for all applicable site conditions.

Resolution To predict potential separation between the foundation and side soil under seismic excitation, the at-rest soil pressures and lateral seismic soil pressures enveloped over a range of soil conditions are compared. It has been found thni the top 10 ft of the embedded reactor building foundation may be separated from the soil due to 0.15g OBE. The separation depth may become 16 ft for the 0.3g SSE. For conservatism a 20 ft separation is assumed to evaluate the effect on structural responses.

To evaluate the soil-foundation separation effect on a generic site basis, a soft side soil site (UB1D85) and a rock site (HRD85) are analyzed for 0.15g OBE with the top 20 ft soil or rock adjacent to the foundation walls removed, using the SASSI 2D analysis option. The corresponding analysis cases without considering the separation are C8X for the soil profile UB1D85 and C10X for O the rock profile HRD85 .

For the soft side soil case, the effects of soil-foundation separation on response spectra are minor as shown in Figs.

1 through 4. For the rock case, however, lower responses are generally predicted as shown in Figs. 5 through 8, when separation is assumed. This is mainly attributed to a less overall toundation motion due to lack of direct application of input excitation at the portion of the foundation being separated from the soil. Reduction in response is not observed in the soft soil case because the coupling effect is not strong to begin with. As a result, the dynamic characteristics of the soft SSI system are not affected significantly.

The same trend is noted for peak force responses as shown in Table. 1.

In summary, the assumption of no separation between the soil and foundation is adequate for design purposes.

O 1.5-I

TABLE l. ETTECT OT SOIL-POUNDATION SEPARATION ON MAXIKUM TORCES DUE TO 0.150 e B E O BEAM ELT LOCATION

RESPONSE

TYPE CBX (UB1DB5)

V/0 SEP V/ SEP C10X (HRD85) SITE W/0 SEP V/ SEP ENVELOPE 2B SHROUD SUPPORT SHEAR 95 98 201 157 264 MOMENT 644 639 1450 1130 1740 69 RPV SKIRT SHEAR 205 219 470 361 780 MOMENT 1450 1460 3280 2080 5322 78 RSV BASE SHEAR 272 281 797 459 1044 MOMENT 1430 1470 40B0 2120 5064 86 PEDESTAL BASE SHEAR 1130 1130 1340 1080 3343 MOMENT 19800 21000 25100 19000 72077 92 RCCV BASE SHEAR 6530 6550 6330 5590 16300 MOMENT 147000 152000 220000 142000 295000 102 R/B BASE SHEAR 10100 10900 6110 6650 21000 MOMENT 290000 294000 400000 262000 564000 UNITS: SHEAR IN TONS; MOMENT IN TON-METERS O

O I,5 ~1

_ ___ _ _ _ _ _ _ _ __- _ _ ___ _ -_ _ _ _ _ -- _ __ _ o

(V i i

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ig 20-FT SEPARATION o

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EFFECT OF SOIL-FOUNDATION SEPARATION FOR C8K (UB1D85)

- 2% DAMPING - AT RB TOP (NODE 95)

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EFFECT OF SOIL-FOUNDATION SEPARATION FOR CSX (UB1D85)

- 2% DAMPING - AT RPV/MS NOZZLE (NODE 33)

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EFFECT OF SOIL-FOUNDATION SEPARATION-FOR C10X (HRD85)

- 2% DAMPING - AT BASEMAT (NODE 88) .

I. REQUEST OF ADDITIONAL INFORMATION OR CLARIFICATION O I.6 Section 3A.8 GE should justify the model of the reactor building wall in SASSI 3-D analysis as shown in Fig. 3A.8-23.

Resolution Dynamic model of the reactor building is a 3D stick model which represents the building stiftness and mass properties along the height. SASSI SSI model of the reactor building incorporates this stien medel with the foundation model.

To properly consider the impedance and scattering effects of the embedded foundation, the perimeter shear walls are modeled by plate elements. The properties of these walls are subtracted from the corresponding 3D stick model properties in the embedded part of the building in the combined SSI model in order to avoid double counting of the perimeter wall properties. The 3D model is subsequently connected to perimeter walls as shown in Fig. 3A.8-23. The connection of stick model to perimeter walls follows the basic assumption used in 3D stick model development in that the walls perpendicular to the direction of shaking (cross walls) do not provide significant stiffness to the box wall -

system in resisting earthqucke loads. Consequently, the 't g

gs stick model is connected to the walls as shown in Fig. 1, and the cross walls are allowed to deform as illustrated in this figure.

In computing shear and moments acting on the perimeter walls, the following loadings commonly used for nuclear structure are considered:

l

1. Thu static soil pressure is computed using the at-rest condition. At-rest pressure is selected since the ,

walls due to rigidity of floor diaphragm may not ,

undergo sufficient displacements and rotations to allow for active pressure to develop under static conditions.

2. The seismic pressure on the walls is computed from elastic solution obtained from SASSI analysis of the reactor building. A column of soil elements modeled by brick elements (see Fig. 1) is used to obtain the seismic pressure along the height.of the walls. This solution is an elastic solution, and can be shown that it is conservatively larger than the' commonly used solutions obtained from Mononobe-Okabe method based on plastic equilibrium of soil-wall system.

L 3. The inertia load (forces and moments) obtained from s SASSI analysis along the building height is used.

1.64

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1 a

(~N In summary, the loading considered-for perimeter walls is

't/ - an absolute sum of the site-envelope loads obtained from 1) static at-rest pressure, 2) seismic soil pressure, and 3) inertia loads.

The combined loading considered is conservative, and adequately includes the static and seismic soil pressures and inertia loads for reactor building walls.

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I. REQUEST OF ADDITIONAL INFORMATION OR CLARIFICATION

1. 7 Section 3G.1 GE has agreed to provide SASSI User Manual and Validation Report (complete set). To establish more confidence on the SASSI code, the staff needs additional application-oriented exercises of the program including SSI response of deeply embedded flexible structure founded on rock.

Resolution a) The following SASSI documents are provided under separate covers (1) SASSIO15 Users Guide, NEDE-31496, GE Proprietary.

(2) SASSI User's Manual, Version 1, Rev. O, Bechtel Proprietary.

(3) SASSI Theoretical Manual, Bechtel Proprietary.

(4) SASSI Computer Program Validation Report, GE SASSIO15 CRAY-XMP Version, Bechtel Proprietary.

Note: Appendices are not included. Should NRC request access to these appendices, they could be made available for inspection in the GE ABWR project office.

(5) Summary of validation problems 13 and 14, GE Proprietary b) A validation problem is designed to address the conditions as stated. The problem configuration is shown in Fig. 1. The structure has the same overall dimensions as the actual reactor building of the ABWR standard plant.

The structure is embedded into a 85 ft deep soft soil-of uniform properties, and'is founded directly on bedrock.

The structure is modeled as a center-lined, lumped-mass stick model as shown in Fig. 2. This model basically is the outer stick of the actual reactor building structural model. The excavated soil model represented by 2D plane strain. elements is shown in Fig. 3. As noted, only half of the soil volume is included because of symmetry. The 2D foundation model for the flexible side wall and basemat is shown in Fig. 4 which also shows the connections to the superstructure stick model.

The control motion is the H2 component of 0.15g peak acceleration considered for the ABWR plant. It is 9 specified at the grade level and is assumed to be a vertically propagated shear wave.  ;

I . 7 -- l

V i L

f e t i

' - l A In order to evaluate the reasonableness of the SSI response, natural frequencies of the structural nodel are first calculated for three different boundary conditions as ,

illustrated in Fig. 5. Condition (a) is the fixed base  !

nodel which neglects the restraints provided by the side o

soil and treats the bedrock as perfectly rigid. The '

resulting structural frequencies are upper bound values. t Condition (b) considers the bedrock flexibility but neglects the side soil. The base springs are frequency independent and their stiffnesses are calculated according  ;

to Ref. 1, in which the embedment effect is neglected. The ,

associated frequencies are, as expected, lower than the  !

fixed base frequencies (Fig. 5). Condition (c) includes  !

both bedrock flexibility and side soil restraints, which -

simulates the actual SSI system more realistically. The  ;

base springs are the same as those considered for condition t (b), and the lateral springs are calculated according to l' the Novak formulation for plane strain case (Ref. 2). The calculated impedances are frequency dependent. However, for natural frequency calculations, a frequency independent stiffness, which is taken to be the peak value of the real  !

part of the frequency dependent impedances, is chosen for i the lateral soil spring. As shown in Fig. 5, the condition r L (c) . frequencies fall in between conditions (a) and (b).

The trend is reasonable. '

t]

\ The SSI response calculation is made using the 2D option of the SASSI code. The calculated response spectrum of 2%

damping at the top of the building is shown in Fig. 6. The .

predominant spectral peak occurs at 4.27 Hz which is close -

to the fundamental SSI frequency (4.16 Hz) predicted by the ,

condition (c) model. The reasonableness of the response '

amplitudes is evaluated by examining the transfer functions. The transfer function (absolute values) of i total acceleration response at the building top.is shown in  !

Fig. 7. According to the half-power method (Ref. 3), the i damping value associated with the fundamental mode is about 11%. The material damping considered for the structure and the surrounding soil is 4%. The difference is due to the radiation damping for the layered system considered. The total damping value of 11% is judged to be reasonable. For  ;

a damping value of 11%, the amplification factor of total acceleration transfer function at the resonant frequency is

l. calculated-to be 4.65 (Fig. 7), which is very close to 4.8 l- calculated by SASSI. I l

E The basemat SSI response is compared to the free-field -

l' response at the elevation corresponding to the basemat base in Fig. 8. The SSI and free-field responses are L essentially the same. This is reasonable since the structure is directly supported by rock of which the SSI

! effect is not significant. The basemat also experiences

\

rocking motion because of variation in free-field motions >

l 1. 7 - 1

- .. - .~

l l

I l

at depth. The calculated vertical response at the edge of

(/ ') . the basemat is shown in Fig. 9 which is an indication of the basemat rocking response. As shown in this figure, the predominant rocking frequency is at 4.27 Hz which coincides i with the frequency of the peak translational response at j the building top (Fig. 6). This indicates that the j fundamental mode of response is a rigid body response. )

In view of the above observations, the SASSI results of l this application-orientated validation problem are believed

)

reasonable. Therefore, the SASSI code is adequate for i application to the ABWR seismic design. -

)

)

Referencer:

1. ASCE 4-86, Seismic Analysis of Safety-Related Nuclear Structures and Commentary on Standard for Seismic  ;

Analysis of Safety Related Nuclear Structures, Table 3300-2, pp 30, September 1986, t

2. Novak, M., Nogami, T., and Aboul-Ella, F., Dynamic Soil Reactions for Plane Strain Case, ASCE, Journal of Engineering Mechanics, Vol. 104, No. EM4, August 1978. .

3. Clough and Penzien, Dynamics of Structures, pp 72, '

McGraw-Hill, 1975

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('N - I. REQUEST OF ADDITIONAL INFORMATION OR CLARIFICATION I.8 Section 3G.2 GE should justify the cut-off frequency equal to 18 Hz for UB sites and to 25 Hz for stiffer sites as shown in response spectra curves.

Resolution ,

The SASSI 2D analysis cases C8X for UB1D85 soil profile as a representative soil site and C10X for HRD85 rock profile as a representative stiffer soil / rock site are reanalyzed up to 33 Hz. The existing soil mesh for case C8X is refined to ensure the 33 Hz frequency transmissibility.

The soil model used in the original C10X case is capable of transmitting 33 Hz and no refinement is made.

The calculated response-spectra are compared with those analyzed to lower frequencies in Figs. 1 through 4 and Figs. 5 through 8 for cases C8X and C10X, respectively, at four key locations. For case C10X rock site, the 33 Hz cutoff essentially has no effect on the results obtained

from a 25 Hz cutoff analysis at superstructure locations (Figs. S through 7). At these locations, the responses of case C8X soil site are relatively sensitive to the cutoff

(~]

(~/

frequency (Figs. 1 through 3); however, the effects are minor except for the response at node 89 (Fig. 2). The response at this location is, nevertheless, governed by the rock site, as can be seen by comparing Figs. 2 and 6. At the basemat both cases, when analyzed to 33 Hz, predict slightly higher response amplitudes at frequencies ranging ,

about 20 to 30 Hz; however, the increased responses are still bounded by the broadened and smoothed site-envelope spectrum shown in Fig. 3G.4-5 in Appendix 3G.

• The peak force responses at key locations are compared in Table. 1 to show the effect of cutoff frequencies. The closeness in results demonstrates that solutions at these locations are sufficiently converged at frequencies lower than 33 Hz. The site-envelope forces are also included in this table for comparison purposes.

In summary, the site-envelope seismic loads developed using the cutoff frequency at 18 Hz for soft soil sites and 25 Hz for stiffer soil / rock sites are adequate.

I. E H W- -*'* w - - - ' -

TABLE 1. ETTECT OF CUTOTT FREQUENCY ON MAXIMUM TORCES DUE TO 0.15G OBE RESPONSE C8X (UB1D85) C10X (HRD85) SITE

-BEAM ELT LOCATION TYPE 18 HZ 33 HZ 25 HZ 33 HZ ENVELOPE 28 SHROUD SUFPORT SHEAR 95 93 201 202 264 MOMENT 644 665 1450 1450 1740 69 RPV SKIRT SHEAR 205 216 470 476 780 MOME!'T 1450 1480 3280 3280 5322 78 RSV BASE SHEAR 272 287 797 797 1044 MOMENT 1430 1470 4080 4080 5064 86 PEDESTAL BASE SHEAR 1130 1130 1340 1340 3343 MOMENT 19800 21400 25100 25200 72077 92 RCCV BASE SHEAR 6530 6630 6330 6340 16300 MCMENT 147000 150000 220000 221000 295000 102 R/B BASE SHEAR 10100 11800 6110 6080 21000 MOMENT 290000 293000 400000 401000 564000 UNITS: SHEAT. IN TONS; MOMENT Ill TON-METERS O -

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---n.-

I. REQUEST OF ADDITIONAL INFORMATION OR CLARIFICATION I.9 Section 3H GE should provide data and methodology that resulted in the maximum bearing pressure shown in SSAR Appendix 3H Section 3H.1.2. .

Resolution

1.

Introduction:

The combined effect of earthquake ground motions (vertical and horizontal) and the response of the structure may cause overturning moment and vertical inertia forces that exceed the stabilizing effect of the dead load. This may be the case when the Dead Load is combined with the vertical seismic component acting upward together with the overturning moment imposed by the horizontal ground motion. This condition is generally accompanied by basemat uplift where it is assumed that the soil does not take tension nor any bond exists:

between the soil and the . foundation. The conventional analyses, in which the vertical force component and the overturning moment.are considered as static loads which are in equilibrium with the soil bearing stresses, may be too conservative.

As a result, the simple method for predicting basemat uplift and maximum soil bearing pressures as presented in Reference 1 was used in the soil bearing stress analysis for the - ABWR 4 Standard Plant.

The ABWR Standard Plant Structure is partially imbedded in the foundation soil. Although the embedment effect of the side soil was included in'the seismic soil structure interaction analysis,_but the resistance of side soil was not included, the analysis is conservative.

2. Relationshin Between Rocking Motion, Overturnina Moment and contact-Area of a Rectancular' Foundation:

The basemat of the ABWR is rectangular, and is considered Figure 1 shows the configuration of a rigid rigid.

rectangular basemat in the uplifted position.

The rocking spring constant of the rectangular basemat shown

. on Figure 1 is:

K, = G/ (1-v) S,(2R) 2 B (Page 3-16 of Reference 2)

I.3 -[

L _ _--____ __ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _

in which R= Half width of the basemat in tho' plano of horizontal groend motion B= Length of basemat perpendicular to . the

'h plane of horizontal ground motion v= Poisson's ratio of foundation medium G= Shear modules of foundation medium

1. , = Constant, a function of the dimensional retio (2R/B)

The above formula is for foundation without embedmont.

The basemat area moment of inertia ist 3

I, = (1/12) B(2R)3 = (2/3) BR The basemat is considered to be supported by a set of uniformly distributed compression (no tension) vertical soilL springs having stiffness k, per unit area, which are determined by the following k, = N/I, (1)

Referring to Figure 1, the foundation vertical force P and the overturning moment M in the uplifted. position can be related

<< ~

to soil pressure p(x) and the foundation contact length D.

D P= p(x)B dx From Figure 1: pod (x) = (D-x)6k in which 6 is the' rocking-rotation of basesEt, and k, is d,etermined following . formula (1) above..

D-Thus: P = k,6B [(D-x) dx o

or P = %k,6BD'D. (2)

D M= p(x) f(R-x)B dx = k,6B (R-x) (D-x) dx 3

M = k,6B(RD8 -\ (R+D) D' + (1/3)D] (3)

Rocking rotation and moment at which basemat uplift initiates; are the following:

Go = PR/ (3K,)

(4)

Mo = K,6o Holding the vertical load as a constant quantity, and varying the contact length D, one can determine the relationship WP01/1220 I . 9 - 2.

between acaent = Mi -'and e i from equations (2) and (3)' for a

. finite number of variations for D.-

j .3 . .Basamat Unlift Prediction and Commutation of Soil Bearina j Strass i The relationship of the' overturning M and rocking rotation 6- ,,

is non-linear when D < 2R, M > Mo and 6 > G. The Energy j o

Balance - Method is based on the consideration that the soil strain energy associatLd with the rocking of basemat is the same for.the linear response ignoring the-uplift effect, and for the nonlinear response considering the uplift effect asj illustrated in Figure 2.

The soil strain energy stored in the soil spring in a linear >

response mentioned above can be expressed as follows:

U,, = \ K,9 8 , (5)

In which 4, = M/K, or U, = %M9, (Sa)

The soil strain nergy in case of uplift can be expressed as:

Ui= ' M de (6) or ,

Ui= Mtoo+ Mi d6, (6a)

In which: Mo and to are quantities which have been determined in equations-- (4) for- moment and rocking rotation at the-Mi and 4 are the moment and 4 initiation of basemat uplift.

rocking rotation at uplift determined in,Section 2 by varying _ ,

the contact length D using_ equations (2) and (3) . Expression

-(6a) can be expressed as numerical summation:

Ui= Mtoo+ (M i + Mi ,) (4, - e i. 3) (6b)

The " Energy Balance Method" stipulates that the above summation for U, be extended until it equals U, calculated in (5) or (Sa). Or the'n'" step of the summation in expression The en as determined in balancing (6b) should be equal to U .

the energy used to calcul, ate the maximum bearing stress:

Fon , = e an k ,* Dn (7) in which Dn is the contact length at n'h step at the time when the energy balance is reached.

According to Reference 1, the above computation tends to WP01/1220

1. 9 - 3

.underpredict the uplift. The amount of under prediction tends -

- to be bigger with larger 6. Therefore, an empirical-L -

correction to the " Energy Bal"ance Method" is developed to improve the prediction. Reference 1 introduces a coefficient a which is the normalized instantaneous secant stiffness of basemat rocking 9, and moment M n or s = (M/Mo)/(6/t o) (8)

The improved " Energy Balance Method" is:

63= 4/K The contact length D, corresponding to 9 can be calculated 3 from:(2) as follows:

D3= P/ (k,4 B) 3 The' upper bound maximum soil stress is:

F 3 = o *g k ,* D3

3. References

1. " Simplified Methed of Predicting Seismic Basemat Uplift-of. Nuclear Power Plant Structures", Transactions of the

9 6th International Conference on SMiRT,- Paris, France, August 1981, by'Tseng, W.S., Liou,_D.D.

'2. " Topical Report Seismic Analyses of Structures. and -

Equipar.nt for Nuclear Power Plants" No BC-TOP-4-A Revision _3, November 1974, by Bechtel Power Corporation, San Francisco, California.

i WP01/1220 T. b 4

f ti

k.

t am 2 o -

y A

W s%. i

_ __ o _

T- em j g .

{ . - a JL -

p(a) ,, a n,

4L Fig. 1- Basemat Configuration and Pressure Distribution i

O WP01/1220 T . 9 -6

.s

3

'e t

JL M'

/

H, b

b

=

o .. .. .. .

Area =. Area y 4 he,M,= ho Mo o + h , (M i + Mi . 3) (#i-#5 1)

• = (M n/Mo)/ (*n/#o) e, = e,,/*

Fig. 2 Concept of energy balance method WP01/1220 T.9-6

_ _ ,;p i

z'}. .

N f' ~ 1x10 .Recent earthquake recordings-in the central and eastern United States have indicated higher' spectral content at frequencies greater than 10 Hz than that found in the R.G. 1.60 spectrum. What impact-does this ground motion have on the-response of the plant structures and components? = Assess its significance. (Note: This is not an audit question)..

. Resolution

- -Seismic response of' structures and components depends not only on the frequency / amplification content of the input earthquake, but also on the dynamic characteristics of the structural system.

The major acceleration amplification region of the R.G.

1.60 horizontal spectra is in the frequency band ranging from 2.5!Hz to 10 Hz approximately. The-amplification factors in this frequency range are generally higher than that of the ground response spectra recommended in NUREG/CR-0098. For structural systems having dominant resonant frequencies in the major amplificaiton region, the total responses predicted using the R.G. 1.60 as the earthquake definition.should be adequate since the conservatism in the modal responses of the more significant modes below 10 Hz-should more than compensate the potential

.- under-prediction of the modal responses of the less significant modes higher than 10 Hz. The primary system of the.ABWR reactor building structures coupled with the RPV and its-major internal components has several modes below 10 Hz and the cumulative modal mass up to 10 Hz is about 86% of the total system mass for the fixed base condition.

More modes.below 10 Hz will be excited.when the SSI effect

'is-considered. .Therefore, the potentially larger ground motion ^ input above 10 Hz should not have adverse impact on the total response of the primary system of the reactor building complex.

With regard to the seismic. response of components using floor response spectra as input, the potentially larger ground motion input above 10 Hz is also expected to have no adverse impact, since

1) the building structures and soil act as low-pass filters which effectively eliminate high frequency components of

• the broad-band input ground motion. This is especially the case for the floor motions at higher elevations in the

. building, and

2) the site-envelope floor response spectra were established in a conservative manner. The inherent design margin should be more than sufficient to bound the '

..8- potentially higher spectral amplitudes above 10 Hz.

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. II. ACTION ITEMS FOR GE

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' II.1. Section 2.0'- In Table'2.0-1 of.the ABWR-SSAR, GE

' should clarify that the minimum soil' shear, wave velocity used will be 1000 ft/sec after the soil property-.

l uncertainties-have been applied.

l

' Resolution

, Table-2.0-1 is-revised accordingly.  !

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ABM 22^6 oo^o REV B Standard Plant TABLE 2.01 ENVELOPE OF ABWR STANDARD PLANT SITE DESIGN PARAMETERS Maximum Ground Water level: Extreme Wind: Basic Wind Speed:

2 feet below grade 110 mph (1)/130 mph (2)

Maximum Flood (or Tsunami) 12 vel:(3) Tornado:(d) 1 foot below grade Maximum tornado wind speed: 260 mph Translationalvelocity: 57 mph Radius: 453 ft Pacipitation (for Roof Design): Maximum atm AP: 1.46 psid

- Maximum rainfallrate: 10in/hr Missile Spectra: Per ANSI /ANS 23

- Maximum snow load: 50 lb/sq. ft.

Design Tempentures: Soll Properties:

- Ambient - Minimum Bearing Capacity (demand): 15ksf

- 1% Exceedance Values Minimum Shear Wave Velocity: 1000 fps c6)

- Maximum: 1000F dry bulb /770F coincident wet 1.iquification Potential: _ ,

bulb None at plant site resulting ~1 from OBE and SSE(7) A Minimum: 100F 0% Exceedance Values (Historical limin Maximum: 115 F dry bulb /82 F coincident wet Seismology:

bulb *

- OBE Peak Ground Acceleration (PGA):

0.10g(5) (6)

O- Minimum:- 400F

- Emergency Cooling Water Inlet: 950F SSE PGA :030g(5)

- Condenser Cooling Water Inlet : 5.1000F SSE Response Spectra: per Reg. Guide 1.60 SSE Time History: Envelope SSE Response Spectra (1) 50-year recurrence interval; value to be utilized for design of non safety related structures only.

(2) 200 year recurrence interval; value to be utilized for design for safety related structures only.

(3) Probable maximum flood level (PhiF), as defined in ANSI /ANS 2.8, ' Determining Design Basis Flooding at Power Reactor Sites."

(4) 1,000,000 year tomado recurrence interval, with associatedparameters based on ANSI /ANS 2.3.

(5) y,ee.peld, atplant grade elevation.

.(6) For conservatism, a value of 0.15g is employed to evaluate structural and component responses in Chapter 3.

5 (7) See item 3 in Section 3A.]for additionalinformation. n d A su s&a< ns (S)7%aa & m wms%rm.aeme after- & s.d yni.edy

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?^s' II. ACTION ITEMS FOR GE '

's-II.2 Section 2.5.4.8 GE should amend the SSAR to specify the characteristics of the earthquake loading (magnitude ,

and-frequency) for which the liquefaction potential of the site will'be investigated-and leave only the site specific soil properties as interface requirements. t Resolution i

The SSE peak ground acceleration considered for the ABWR standard plant is 0.39 This is a definition of the maximum design earthquake at the site.- This local level of-

-ground. shaking can be related to different magnitudes of earthquakes, depending on epicentral distances. Without ,

knowing the exact site location and its seismic hazard .

, characteristics, it'is not possible to specify earthquake 2 magnitude and' associated-frequency (or number of strong  ;

motion cycles) for liquefaction evaluation of a given l site. Such design parameters should be established on a "

site specific. basis. .

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-.fh'j II. ACTION ITEMS FOR GE

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II.3~ Section 3.7.1 For Figures 3.7-24 and 3.7-25, GE will recalculate the power spectral density (PSD) for H1 and H2 time histories based on the smoothing technique and the new equation for target curve as described din Appendix A of  :

SRP 3.7.1 Rev. 2. These curves should be-normalized to 0.15g.

I Resolution -

The power spectral density functions (PSDFa):of the H1 and H2 synthetic ~ time histories normalized to 0.15g peak ground-

-acceleration are recalculated according to SRP 3.7.1, Rev.

2 criteria. ' Subsection 3.7.1.2 is revised accordingly.

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Figure ILLUSTRATIONS (Continued)

Tult East 33 15 Synthetic Time History, Vertical Direction,

' Damping Ratio 0.01 3 3-49  ;

1 Synthetic Time History, Vertical Direction, 3316 "

Damping Ratio 0.02 33 50-33 17 Synthetic Time History, Vertical Direction, Damping Ratio 0.03 3,7 51 3.7 18 Synthetic Time History, Vertical Direction, Damping Ratio 0.04 3.7 52 4 33 19 Synthetic Time History, Vertical Direction, Damping Ratio 0.07 33 53

. 3.7 20 Synthetic Time History, Vertical Direction, <

Damping Ratio 0.10 33 54 ,

33 21 Coherence Function C12 for Earthquake 33 55 .'

Components H1 and H2 '

() _

Coherence Function C 13 for Earthquake

.Q'.

C/ 3.7 22 Components H1 and V 3.7 56 33 23 Coherence Function C 23 for Earthquake Components H2 and V - 33 57 33 24 Power Spectral Density Function of Synthetic ,

H1 Time History 33 58 33 25 Power Spectral Density Function of Synthetic H2 Time History ~ 33 59 33 26 O.s.i.i;"e ;C RJ O ei, T. dond4

-Spdr9 91-TeswHistory-A D e. l e.4 c. ) 3.7 60 33 27. -C-uJ.06" woiS Hi.l b i Tuudeud t i 4

Spi:9 F2 TE: Mete y---e D e.I e.4.c.4 33 61 3.7 28 Seismic System Analytical Model 3.7-62 3.7ix Ameadment 1 E. 3 ~ 2.

21A6100AE RFV B Standard Plant c,

The frequency range used in generating the for tne anwk design.

m: response spectra from synthetic histories is 0.2 ~

aio in furthe: comperiscus of PSpfs, the to 33 Hz. The frequency range intervals used in Sistories, generating those spectra is the same as given in cumu tive PSDFs of the two ti gy contribu.

Table 3.7.11 of SRP Section 3.7.1. which re esent the cumulative e tions as a netion of frequey , are computed The coherence function for the three earthquake as the cumula 've area unde (the PSDF curves.

acceleration time history components Hi, H2, and V The cumulativ PSDF,s'versus frequency as I time histories are shon are generated to check the statistical indepen- computed for H1 an dence among them. The coherence function for H1 in Figures 3.7 26 d .7-27, respectively. For and H2 is given in Figure 3.7 21; for H1 and V in comparison, the trespon 'ng cumulative target Figure 3.7 22; and for H2 and V in Figure 3.7 23. PSDFs are als shown in th e figures. As can All values within the frequency ratge between 0 to be seen f,rpfn these figures, the Iculated time SO Hz are calculated at a frequency ineretcent of historyAmulative PSDFs envelop cumulative

1. PSDF with a wide margin in thetequency 0.1 Hz. The small values of these coherence targy(between 0.2 Hz and 34 Hz. N functions indicate that the three components are range

- sufficiently statistically independent.

3.7.1.3 Critical Damping Values To assess the energy content of the synthetic time history, the power spectral density functions The damping values for OBE and SSE analyses (PSDFs) are generated from the two horizontal are presented in Table 3.7-1 for various components H1 and H2. The PSDFs are computed at a structures and components. They are in i frequency increment of 0.024 Hz, and are smoothed compliance with Regulatory Guides 1.61 and 1.84 using the bmm /aserage method as For seismic system evaluation of the SSE, the recommended ind vm 3. eference

• • • "2 '# larger SSE damping values shown in Table 3.71 The stationary duration used in the calculation are not used. The SSE loads are obtained by a is taken to be 22 seconds which is tbc total doubling the OBE loads that result from the OBE

.M' duration of the synthetic time history. The Seismic System analysis based on the lower OBE calculated PSDFs for the H1 and ri2 time histories damping values (see Subsection 3.7.1.2).

g

,p" e h i Figures 3.7 24 and 3.7 25, respec-For analysis and evaluation of seismic

/ FQp ^are s own ntively, for frequencies d A 7,r/ derpe r M u 4 2r rangingsubsystems from Mto (piping, MHz.and equipment),

components The target PSDFsli; i fx; cf S;iT jisi' the floor response spectra are chtained from the BSDP.a Tr% Reference 3 are also plotted OBE time history response of the seismic system, The floor on t ese figures for comparison., L X.u.a that supports the subsystems.

Tapei-P&DF-rsgtverby < response spectra are computed (see Subsection

1. " 3.7.2.5) for damping values that are applicable uW 1 i s.9).-s M-g.7 :)o to the subsystems under OBE as well as SSE; and o further the OBE spectra are doubled to obtain the SSE floor response spectra for input to the e4 Q Weg)W ~ SSE analysis in design of the subsystems.

/ , , , ,  %

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"" Qw/wgP 3.71 A Supporting Media for Seismic Category hich So = 1,100 in 2 7,ee3 (t h i Xa^lu e I Structures i

corre ds to a peak acceleration ), wg =

f nd (g = 0.9793 s can be seen The following ABWR Standard Plant Seismic 10.66 rad a '

g from Figures 4 and -25, the calculated Category I structures have concrete mat g

PSDFs generally eny ' p target PSDFs in the foundations supported on soil, rock or compacted frequency range row 10hQut they fall below backfill. The maximum value of the embedment the target - s in the frequenevrange above 10 depth below plant grade to the bottom of the t

Hz. However,it should be noted that confortnance base mat is given below for eacE structure.

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to th[ target PSDF is currently not a requirement

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