ML20150A644
| ML20150A644 | |
| Person / Time | |
|---|---|
| Site: | Trojan File:Portland General Electric icon.png |
| Issue date: | 09/20/1978 |
| From: | Broehl D PORTLAND GENERAL ELECTRIC CO. |
| To: | Schwencer A Office of Nuclear Reactor Regulation |
| References | |
| TAC-07551, TAC-08348, TAC-11299, TAC-7551, TAC-8348, NUDOCS 7809270083 | |
| Download: ML20150A644 (186) | |
Text
{{#Wiki_filter:_ _ _ . . . 7 L. - VAM PORTLAND GENERAL ELECTRIC CourAxr las S. W. SAL MON STREET PORTLAND, OREGON 972o4 LOrestam wet passtosur September 20, 1978 Trojan Nuclear Plant Docket 50-344 License NPF-1 Control Building Proceeding Director of Nuclear Reactor Regulation ATTN: Mr. A. Schwencer, Chief Operating Reactors Branch #1 Division of Operating Reactors U.S. Nuclear Regulatory Commission Washington, D.C. 20555 Dear Sirst Enclosed, as promised in my letter of September 1, 1978 (as supplemented by Maurice Axelrad's letter of September 11, 1978 to the Licensing Board), are the final results of the review and evaluation of the recent STARDYNE finite element analysis perfomed for the existing Control Building structure of the Trojan Nuclear Plant. Attachment 1 is "Trojan Control Building Supplemental Struc-tural Evaluation" (September 19, 1978) prepared by Bechtel, which is supplementary to and supportive of previous analyses and evaluations submitted to the Staff on May 5 and May 24, 1978. This report consists of Sections 1 - 8 and Appendices A-E. l Attachment 2 is "Response to Questions from the Nuclear Regu-latory Commission dated August 30, 1978" (September 20, 1978) prepared by Bechtel. Attachment 3 is "Response of Trojan Nuclear Plant Control Building to Specified SSE Event" (September 20, 1978) prepared by Professors Myle J. Holley and Boris Bresler. Sincerely, cc Mr. R. H. Engelken, Director U.S. Nuclear Regulatory Commission i Region V I&E
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's b TROJAN CONTROL BUILDING SUPPLEMENTAL STRUCTURAL EVALUATION SEPTEMBER 19, 1978 5, J -
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r . i -: TABLE OF CONTENTS j 7-. . qf
- 1. INTRODUCTION
- 2.
SUMMARY
AND CONCLUSIONS
- 3. SEISMIC ANALYSIS AND LOAD DETERMINATION
' 4. CAPACITY DETERMINATION ,
- 5. COMPARISON OF LOADS AND CAPACITIES
- 6. M0 MENT AND VERTICAL SHEAR TRANSFER F
- 7. REDISTRIBUTION OF FORCES
, 8. TRANSFER OF SHEAR FORCES T0 ROCK FOUNDATION Appendix A - Comparison Of Seismic Analyses Appendix B - Shear Capacity Criteria Appendix C - Strain Compatibility
() Appendix 0 - Displacement Determination Appendix E - Inelastic Behavior , o e C i
LIST OF FIGURES ( . FIGURE NUMBER TITLE 3-1 "STARDYNE" Model 3-2 Wall Key Plan For Elevation 45'-61' 3-3 Wall Key Plan For Elevation 61'-77' o 3-4 Fundameatal Mode Shape At Top Of Structure. North-South s Direction. Plan At Elevation 117' (Not To Scale) 3-5 Fundamental Mode Shape At Top Of Structure, East-West Direction. Plan At Elevation 117' (Not To Scale) 4-1 Shear Wall Capacity Criteria 8-1 Foundation System v o e N> 11
4 I
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LIST OF TABLES n U ^ TABLE NUMBER- TITLE 3 Shear ~ Forces, N-S Motion, Elev. 45', 8/17/78, SSE = 0.259 , 8 = 5%, Flexible Base 3-2 Shear Forces, E-W Motion. Elev. 45', 8/17/78, SSE = 0.259, 8 = 5%, Flexible Base 3-3 Shear Forces, N-S Motics Elev. 61*, 8/17/78 SSE = 0.259 ,
$- 8 = 55, Flexible Base 3-4 Shear Forces, E-W Motion Elev. 61', 8/17/78, SSE = 0.259, 8 = 5%, Flexible Base 3-5 Shear Forces, N-S Motion, Elev. 45', 8/24/78 SSE = 0.259, B = 5%, Fixed Base 3-6 Shear Forces, E-W Motion, Elev. 45', 8/24/78 SSE = 0.25g, s = 5%, Fixed Base O 3-7 Shear Forces, N-S Motion, Elev. 61', 8/24/78, SSE = 0.25g, 3 = 5%, Fixed Base 3-8 Shear Forces, E-W Motion, Elev. 61', 8/24/78 SSE = 0.259, s = 5%, Fixed Base 3-9 Oynamic Characteristics - Fixed Base e 4-1 Capacities In N-S And E.W Directions (Elevation 45'-61')
4-2 Capacities In N-S And E-W Directions (Elevation 61'-77')
- O iii 1
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( ; TABLE NUMBER TITLE 5-1 Force-Capacity Comparison, N-S Motion, Elevation 45'-61', Fixed Base, SSE = 0.25g, s = 5% 5-2 Force-Capacity Comparison, E-W Motion, Elevation 45'-61', Fixed Base, SSE = 0.25g, 8 = 5% o 5-3 Force-Capacity Comparison, N-S Motion, Elevation 61'-77', Fixed Base, SSE = 0.259 , s = 5% ( 5-4 Force-Capacity Comparison, E-W Motion, Elevation 61'-77', Fixed Base, SSE = 0.25g, 8 = 5% 6-1 Gross Moment Tension Forces For N-S, SSE = 0.259 , s = 5%, Elevation 45' And 61' 6-2 Vertical Shear Forces And Capacity For N-5 SSE = 0.25g, 8 = 5% O LJ 7-1 Base Case Shear Forces, N-S Motion, Elevation 45' 7-2 Base Case Shear Forces, N-S Motion, Elevation 61' 7-3 Case 1 Shear Forces, N-S Motion, Elevation 45', Based On Limited Capacity Of Wall 1 From Elevation 45' To 61' 7-4 Case 1 Shear Forces, N-S Motion, Elevation 61', Based On l Limited Capacity Of Wall 1 From Elevation 45' To 61' 7-5 Case 2 Shear Forces, N-S Motion, Elevation 45', Based On l Limited Capacity Of Wall 3 From Elevation 61' To 77' 7-6 Case 2 Shear Forces N-S Motion. Elevation 61', Based On Limited Capacity Of Wall 3 From Elevation 61' To 77' (_/ iv
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TABLE NUMBER TITLE C) - Case 4 Shear Forces,-N-S Motion, Elevation 45', Based On 7-7' Limited Capac4 ties From Table 4-1 7-8 Case 4 Shear Forces, N-S Motion, Elevation 61', Based On Capacities From Table 4-2 7 Case 5 Shear Forces, N-S Motion Elevation 45', Based On Limited Capacity Of Wall 1 From Elevation 45'-61'
- 7-10 Case 5 Shear Forces N-S Motion Elevation 61', Based On Limited Capacity Of Walls From Elevation 45'-61' 7-11 Summary Of Cases, Elevation 45'-61' 7-12 Suninary Of Cases. Elevation 61'-77' 7-13 Elastic Displacements 8- l ' ' Sliding Resistance And Base Shear 4
O v
9/19/78
- 1. INTRODUCTION The original seismic analysis of the Trojan Control-Auxiliary-Fuel Building complex was completed by Bechtel in May 1971. This analysis was updated by Bechtel in April'1978, when the weight of the Control Building was re-estimated based on the as-built condition and responses were re-calculated using the square-root-of-the-sum-of-the-squares (SRSS) technique. In June 1978, an analysis performed by an independent consultant using the TABS program
- verified that the prior Bechtel analyses were conservative. More recently, in August 1978, additional analyses were perfomed by Bechtel using the finite element STARDYNE program to predict seismic loads. This report provides additional information, analyses results and conclusions associated with these finite element analyses.
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9/19/78 (x._) 2.
SUMMARY
AND CONCLUSIONS Four seismic analyses have been performed on the Trojan Control-Auxiliary-Fuel Building complex: 1) the original spectral response analysis using a stick model, in 1971 (re-evaluated and modified by application of SRSS and use of estimated as-built weights, in April 1978); 2) an analysis by an independent
, consultant utilizing the TABS program, in June 1978; 3) an analysis utilizing a finite element model and the STARDYNE program with flexible-base, in August 1978; and 4) an analysis utilizing a finite element nudel and the STARDYNE program with fixed-base, in August 1978. These analyses are described and compared in Appendix A to this report.
Section 3 of the report provides information on loads predicted by the STAROYNE analysis. This analysis is highly sophisticated and comprehensive; walls and (3 slabs are modeled by finite eier.ents. In some walls, the loads predicted by LJ STARDYNE are higher than previous analyses predicted. Due to the use of elastic finite element analysis, the loads predicted by STAROYNE are upper limits. Section 4 provides criteria for the shear wall capacities used to evaluate the capability of the Control Building to resist loads predicted by the STARDYNE a nalysi s. These capacities are higher than those used in the previous evaluations. Justification for use of higher capacities is provided in Section 4 and Appendix B, along with the developments of the criteria employed to determine the capacities and their correlation with tests. Strain cmpatibility of the $ bear walls is discussed in Appendix C. 2-1
O Sectio" 5 o' '"' r*a "* c a' '"' S'^" '"'-ar'd'cd ' 'd' "'*" '"' '"r wall capacities. Only in a limited number of small walls does the load demand exceed the capacity. The contribution of those walls to the capability of the total structure to resist seismic forces is very small. . Section 6 d. ascribes the capability of the Control Building to resist gross bending moments and the transfer of shear fran side walls to end walls. The gross bending moments are primarily resisted by the end walls perpendicular to the direction of earthquake motion. Sufficient vertical shear capacities exist te permit transfer of load from side walls to end walls. Section 7 documents a detailed investigation which distributes seismic loads
,as a function of both stiffness and capacity. The investigation demonstrates that the Control Building can accomodate the seiwie loads, even with small walls exceeding yield, due to multiple shear walls which have a total capacity higher than the total load.
Section 8 describes the load transfer mechanism from the Control Building to the foundation rock and demonstrates that a sufficient factor of safety exists against sliding. On the basis of the investigation and analyses described in this report, it is concluded that the Trojan Control Building can resist, with a margin of safety, the forces due to a 0.25g SSE with 5% damping. Since the seismic forces resulting from a 0.259 SSE with 5% damping and a 0.11g OBE with 2*. damping are the same, it follows that the 0.119 OBE for the Trojan Control Building continues to be appropriate. O 2-2
] 3. SEISMIC ANALYSIS AND LOAD DETERMINATION Additional seismic analyses have been perfonned by Bechtel using the STARDYNE finite element computer program. The first analysis simulated foundation flexibility and the second has used a fixed base condition. The analyses were based on the following:
- 1) The entire co@ lex of Control, Auxiliary, and Fuel Buildings was modeled.
The model considered all significant wall and floor slabs which were si.nulated by plate finite elements. Access openings were modeled but not pipe penetrations and other small openings. The stiffness of the elements was based on elastic properties. The model is shown in Figure 3-1.
- 2) The stiffness of the foundation for the flexible base case was based on the techniques presented in BC-TOP-4 A (Ref.1). A shear wave velocity of 5500 fps, as defined in the FSAR war used for the foundation rock.
The foundation flexibility was only considered under the Control Building. Radiation damping conservatively was rot considered.
- 3) The .259 SSE response spectra with 5". damping presented in the Trojan FSAR was used in the analyses.
- 4) The seismic analyses were performed by a linearly elastic modal analysis spectrum response technique, utilizing the comercially available STARDYNE computer program. The modal responses were combined using the square root of the sum of the squares technique (SRSS).
O 3-1
" A Key Plan for elevation 45'-61' is shown on Figure 3-2 and the Key Plan for O elevation 61'-77' is shown on Figure 3-3. These Key Plans are for reference j and do not show all openings.
Tables 3-1 to 3-4 show the shear forces at elevations 45' and 61' for both the N-S and E W directions for the flexible base case. Tables 3-5 to 3-8 show the shear forces at elevations 45' and 61' for both the N-S and E-W directions
, for the fixed base case.
, A comparison of these two sets of tables indicates only a 3% difference in a gross response between the flexible base and fixed base cases. Since the foundation i , shear wave velocity is high, the fixed base case is an appropriate representation of the response, and evaluations given in this report are based on this case. The mode shapes for the fundamental and most dominant modes are shown in Figures 3-4 and 3-5 for the top of the structural complex. The most significant modes are documented in Table 3-9 which shows frequency, participation factor and modal effective weights. REFERENCE [1] Bechtel Topical Report BC-TOP-4A. "Seismic Analysis of Structures and
. Equipment for Nuclear Power Plants". Bechtel Power Corporation. San Francisco, CA, Revision 3. November 1974.
O 32
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t . O Figure 3-4 Fundamental Mode Shape At Top Of Structure, North-South Direction Plan at Elevation 117' (Not To Scale)
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Figure 3-5 Fundamental Mode Shape At Top Of Structure. East- West Direction Plan At Elevation 117' (Not To Scale) 4
(~ t L) Table 3-1 Shear Forces, N-S Motion, Elev. 45', 8/17/78, SSE = 0.25g, s = 57, Flexible Base SHEAR FORCE WALL NUMBER (XIPS) LOCATION 1 4380 N-S WALLS , 2 51 0 3 440 4 2020 5 3550 6 390 7 600 8 320 O r=i22io . i 9 1670 E-W WALLS 10 1280 11 140 12 130 13 1610 14 40 15 290 0
/~m I'~J Table 3-2 Shear Forces F-W Motion, Elev. 45',
8/17/78 SSE = .25g, 8 = 5% Flexible Base ! SHEAR TORCES t
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WALL NUMBER (KIPS) LOCATION 1 920 N S WALL I 2 180 3 190 4 400 5 650
- 6 60 l
7 140 8 60 9 1870 E-W WALL 10 1820 11 440 12 260 13 4700 14 480 15 930 [=10500 t O
c Table 3-3 Shear Forces N-; Motion Elev. 61', (s) 8/17/78, SSE = 0.2Sg, s = 5% Flexible Base SHEAR FORCE WALL NUMBER (KIPS) LOCATION 1 4170 N-S WALLS 2 330 3 3000 4 2130 5 510 6 660 [=10800 7 1530 E-W WALLS O 8 1570 9 830 10 130 11 490 12 300 e f O
q s Table 3-4 Shear Forces, E-W Motior., Elev. 61', 8/17/78, SSE = .259, 8 = 5% Flexible Base SHEAR FORCE WALL NUMBER (K!PS) LOCAT10f4 I 1110 N-S WALLS 2 160 3 870 4 370 5 170 6 200 - 7 1810 E-W WALLS l 8 3560 9 850 10 38 0
, 11 1000 12 1400
[=9000 f O l
(m L J Table 3-5 Shear Forces, N-S Motion Elev. 45', 8/24/78, SSE = 0.259, s = 5% Fixed Base SHEAR FORCE WALL NUMBER (KIPS) LOCATION 1 4110 N-S WALLS 2 780
? 560 4 2240 5 3050 6 340 7 540 8 290
[=11910 9 1540 E-W WALLS 10 970 11 110
. 12 130 13 1260 14 30 l 15 240 l
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s*' V Table 3-6 Shear Forces E-W Motion, Elev. 45', 8/24/78, SSE = 0.259, 8 = 5% Fixed Base SHEAR FORCE WALL NUMBER (KIPS) LOCATION 1 91 0 N-S WALLS
. 2 220 i 3 180 4 420 5 570 6 70 -
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. 12 320 ;
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C/ Table 3-7 Shear Forces, N-S Motion. Elev. 61', 8/24/78. SSE = 0,25g, 8 = 57. Fixed Base SHEAR FORCE , WALL NUMBER (K!PS) LOCATION 1 3910 N-S WALLS 2 560 3 3140 4 1910 5 470 , 6 600 [=10590 h (Y 7 1480 E-W WALLS 8 1450 9 650 10 130 11 440 12 220
- O 5
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8/24/78, SSL = 0.259, s = 5% Fixed Base . SHEAR FORCE
'iALL NUMBEA (KIPS) LOCATION
. I 1080 N-S WALLS 2 170 3 750 4 31 0 5 1 50 6 220 (m (
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fs v Table 3-9 Oynamic Characteristics - Fixed Base T EFFECTIVE FREQUENCY PARTICIPATION WEIGHT (CPS) FACTOR (KIPS) O!RECTION 6.81 1.68 31.100 N-S 9.40 1.11 12,500 N-S 11.80 .92 11,800 N-S 8.53 1.79 50,300 E-W 12.47 2.07 9.900 E-W O O
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4 CAPACITY DETERNINATION The allowable capacities in the codes are usually set anticipating a certain level of sophistication when determining the applied loads. Both the ACI and VBC codes have not significantly changed their shear provisions for several years. It appears that these codes have not considered that the user would be applying techniques as sophisticated as an extensive finite element analysis.
. Only recently have computer programs come into use which can consider the flexibility of an entire cornplex of walls and floor slabs and mathematically distribute the loads throughout the elastic system.
The code provisions for determiring shear capacity in walls are based on wall.s which have e height sufficiently large when compared to the base dimension so that 45* type diagonal tension cracks can develop, which essentially run to both outer edges of the wall. In the case of the Trojan Control Building, many of the walls are quite short in height compared to their length and this situation cannot develop. Therefore, particularly as to these walls, the code provisions are extremely conservative. Recognizing the considerations described above, a set of criteria was developed
. to evaluate the capacities of the shear walls. The detailed developnent of these criteria is provided in Appendix B, dnd the strain compatibility of the masonry walls and the concrete core is illustrated in Appendix C. These criteria are given in Figure 4-1. The wall capacities based on these criteria are given in Tables 4-1 and 4-2. The tables also include ccceents as to which criteria governed the capacity. Af ter evaluating the capacities, it was found that none of the walls that were evalucted by criterion (b) were governed by it. They all had lower shear capacities limited by bending. The Key Plans given in 1 4-1
.. . _ . . _ . . _ . . . . - = .
Figures 3-2 and 3-3 are applicable to these tables. The values given are the h lowest values obtained by the criteria between the elevations sta ad. 4 O 9 O 4-2
Figure 4-1 Shear Wall Capacity Criteria b Basic Criteria: Vu = [261-84 + ]WT .2 5h3 10 V
= (218-41 h + 1WT 1.0 3h2 3.0 . V u = (150-19 h + 1WT 3.0 2h$ 4.0 0
V u
=(75+j]WT 4.0 3h
# *# 1 002; oy 3 0007 and oh " .0013 h v forhj.2useh=.2 where:
V = allowable shear force (lbs) o = horizontal reinforcement ratio H = wall height (in) o y
=
vertical reinforcement ratio
-) W = wall width (in) o" = axial stress (psi)
(+ for ccepression; for tension) T = gross wall thickness (in) In order to qualify for the basic criteria, the walls must be subjected to dead load and _have core reinforcing steel. Walls which have structural steel columns completely interruptin9 the core shall use an ra tio
- which limits W to the distance between columns, i
Additional Criteria: a) Walls which are subjected to dead load, but for which no core reinforcing steel is assumed, may be evaluated by the basic criteria; howver, the stress is limited to 150 psi. 7
./
) Figure 4-1 Continued b) Walls which are not subjected to dead load, such as interior walls between floor slabs, shall be evaluated using the following shear criteria used in the May 5, 1978 submittal:
Af V u =>[2.yt,ff+ y)(.8W) where: a = .85 capacity reduction factor for shear f' = concrete compressive strength (6000 psi) teff = total wall thickness -8" or t gross -8a A3 = minimum of horizontal or vertical reinfon:ement (in?/ft) f y= yield strength of reinforcement (psi)
] Other terms have same definition as given under the basic criteria, c) All walls must be checked for combined bending moment and dead load , effects.
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,a Table 4-1 Capacities In N-S And E-W Directions (Elevation 45'-61')
CAPACITY WALL (KIPS) CONTROLLING CRITERIA 1 5390 Basic criteria 2 470 Bending Moment
~
3 490 5 g 4 3810 Basic criteria 8 " " 5 5 5970 a T 6 110 Bending Moment .
=
7 420 8 60 l=16720 ~ 0 9 4730 Bas 4c :r4ter4e 10 5560 Basic criteria 11 240 Bending liament 5 " "
- p 12 420 S
5 13 9350 Basic criteria o 3 14 170 Bending Moment w 15 760 [*21230 O l l l
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Table 4-2 Capacities In N-S And E-W Directions (Elevation 6i'-77') CAPACITY
. WALL I, KI PS ) CONTROLLING CONDITION 1 5100 150 psi limit 2 1 90 Bending Moment o ,
g 3 4520 Basic criteria
.E 4 2240 Bending Moment T 5 650
=
6 750 [
- 13450
(]' 7 4440 150 psi Ifmit 8 9340 Basic criteria e S 9 2820 t3 W 10 1380 Bending Moment 5 * "
, 11 2390 12 2390
[*22760 ' 4 [ O
- 5. COMPARISON OF LOADS AND CAPACITIES This section compares the fixed base case loads, as determined by the STARDYNE cottputer analysis described in Section 3, with the wall capacities determined employing the criteria described in Section 4. The comparisons are given l in Tables 5-1 inrough 5-4 for both directions at Elevations 45'-61' and 61'-77'.
The loads were developed based on elastic analyvs which assumes that each wall l in the system has a yield strength higher than the load. When the load is : higher than the capacity, then the load is fictitious and cannot develcp further in that particular wall. The most important consideration is that the sum of all the wall capacities is greater than the sum of the applied loads, f i Table 5-1, for elevation 45'-61', and the North-South directico, shows a total capacity 40*. higher than the total applied load. Walls with a.s elastic load , higher than capacity will yield and the excess load will be carred by other i members. As indicated in Table 5-1, only the small walls (Walls 2, 3, 6, 7 and I
- 8) have an elastic load greater than capacity and the combined small wall capacity is only 9% of the total capacity.
t l i Table 5-3 for elevation 61'-77' and the North-South direction shtws a tota l i capacity 27!, higher than the total applied load. In this case, only one small wall has an elastic load greater than capacity. [ t for the Ent-West direction, as shown in Tables 5-2 and 5-4, the total capacity l is very high compared to the total load. There are some small walls with the elastic load greater than capacity, but again, they .:01 only yield at their capacity and their effect on the ijstem is essentially negligible. Based on the large excess capacity in the East-West direction, thi; direction will not be 5-1
1 i O considered further in this evaluation. l ! Sectior,7 (Case 4) exanines a more realistic distribution of forces in the l l North-South direction, when the capacities are considered.
. 1 l
l l.. I l l I O l 9 I i 1
!O 3
4 5-2
v Table 5-1 Force-Capacity Comparison, N-S Motion, Elevation 45'-61' Fixed Base, SSE = 0.259, s = 5% SHEAR FORCE CAPACITY OAPACITY WALL NUMBER (KIPS) (KIPS) LOAC 1 4110 5390 1.31 2 780 470 .60* 3 560 490 . E8
- 4 2240 3810 1.70
$ 5 3050 5970 1. 96
- 6 340 110 .32*
A 7 540 420 . 78
- 8 290 60 .21*
[=11910 [=16720 1.40 C, 9 1540 4730 10 970 5560 11 11 0 240 g 12 130 420 N 13 '260
, 9350 m
e 14 30 170 15 240 760
- Ratios less than 1.0 indicate the load is fictitious since the load cannot exceed the capacity.
O
n v Table 5-2 Force-Capacity Comparison, E-W Motion. Elevation 45'-61', Fixed Base. SSE = 0.259, s = 57, SHEAR FORCE CAPACITY CAPACITY WALL NUMBER (KIPS) (KIDS) LOAD 1 910 5390 2 220 470 3 180 490 0 4 420 3810
$ 5970 g 5 57 0 6 70 110 7 180 420 8 60 60 9 1700 4730 2.78 10 1680 5560 3.31 11 510 240 47
- 12 320 420 1.31 m
d y 13 4620 9350 2.02 a 1 2 14 450 170 .38* w l 15 870 760 .87* ,
, [=10150 [=21230 2.09
' Ratios less than 1.0 indicate that the load is fictitious since the load cannot exceed the capacity.
l l l O
/
v Table 5-3 Force-Capacity Comparison. N-S Motion, Elevation 61'-77', Fixed Base SSE = 0.259 8 = St SHEAR FORCE CAPACITY CAPACITY WALL NUMBER (KIPS) (KIPS) LOAD 1 3910 5100 1.30
- 2 560 190 . 34
- m 3 3140 4520 1.44 d
5 4 1910 2240 1.17 m a 5 470 650 1,38 6 600 750 1.25 {=10590 {.13450 1.27 4440 F] 7 1480 8 1450 9340
$ 9 650 2820 9
- 10 130 1380 11 440 2390 12 220 2390
*
- Ratios less than 1.0 indicate the load is fictitious since the load cannot exceed the capacity.
i O
(_-) Table 5-4 Force-Capacity Comparison, E-W Motion. Elevation 61'-77', Fixed Base SSE = 0.259, s = 55 SHEAR FORCE CAPACITY CAPACITY WALL NUMBER (KIPS) (KIPS) LOAD ; i 1 1080 5100 2 170 1 90 l g 3 750 4520
!! 4 510 2240 m
ak 5 150 650 6 220 750 i 7 1670 4440 2.66 () 8 9 3560 790 9340 2820 2.62 3.57 jl 10 350 1380 3.94 [f, 11 950 2390 2.52 12 1310 2390 L82 {=8630 { = 22760 2. 64 4 9 O
N k ~ J'
- 6. M0 MENT AND VERTICAL SHEAR TRANSFER The gross bending moment is primarily resisted by the end walls (perperdicular to the earthquake motion direction). Table 6-1 shows a comparison of the tension in the end walls due to a North-South earthquake of .259 SSE at 5t damping for the fixed base case and the total dead load. The values shown are the totals for the walls frcm Column Line N to R. As indicated in the table.
. there is no gross tension predicted in the end walls.
Sufficient vertical shear capacities are required so that the side walls can transmit load to the end walls and the end walls can resist the major portion of the gross earthquake moment. Thc. vertical shear capacities have been determined by considering the combined strength of the structural steel beam-to-column connectiors and the dowel strength of the horizontal reinforcing steel. The connection capacity was based on twice the AISC, Part I allowable capacity. The reinforcing dewel strength was detemined using 90!, of the dowel ultimate stress times ths cross-sectional area. This method of detemining capacity is conservative since it neglects any transfer of shear by the concrete. A cotparison of total capacity and total applied load from the fixed base North-South earthquake of .25g SSE at 5% darping is given in Table 6 2. As shown on the table, the minimm margin is 1.55.
\ V 6-1
\ -
C)
\-
Table 6-1 Gross Moment Tension Forces For N-S. SSE = .259, s = 5%. Elevation 45' And 61' EARTHQUAKE TENSION DEAD LOAD WALL NUMBER FORCE (K!PS) (K!PS) m 9 3000 3450 j 13 4310 4560 3 7 2440 2470 f 8 2410 3330 O
'9 9
4 9 O
Table 6-2 Vertical Shear Forces And Capacity For N-S. SSE = .259 . 8 = 5% CAPACITY VERTICAL SHEAR REINFORCING CONNECTION l TOTAL LOCATION FORCE (K!PS) (KIPS) (K!PS) (KIPS) MARGIN , R & 41 1910 1430 1530 2960 1.55 N & 41 1030 1850 1080 2930 2.84 N & SS 2070 2980 2330 5310 2.57 R & 55 2660 3650 1575 5225 1.96 ! O I = l r f i
- 7. REDISTRIBUTION OF FORCES
( [) i 7.1 General As shown in Section 5. (Tables 5-1 and 5-3), the STARDYNE analysis predicts load demands in some walls higher than their capacities. These loads are fictitious since the load developed cannot exceed the capacity. This section will examine in detail how dependent the overall system is on the capacity of large walls and it will also examine the fort:e distribution in the system after yielding of small walls. The analyses presented in this section are very conservative since the applied upper limit loads based on linear elastic response are applied as static loads, when yielding occurs in a system it becomes nonlinear and the response decreases. This is illustrated in more detail in Appendix E. O In order to study the redistribution ability of the Control Building, the following evaluations were made:
- 1) The effect of limiting individual large wall capacity and allowing other walls to develop the required capacity without limits, illustrating which other members would be rest affected (Cases I and 2).
. 2) The effect of limiting vertical shear transfer (Case 3).
O '-l
- 3) The redistribution of loads within the entire system with limited O cea cities (c se 4)- Ta4s 4s com 4eered the est re itstic cese studied. l
- 4) A final analysis, limiting the capacity of the West wall (Wall 1), to illustrate that the Control Building's ability to resist the applied loads is primarily dependent upon having sufficient gross capacity i
and is not even significantly affected by single major walls (Case 5). All analyses presented in this section are for the North-South direction with loads that simulate the .25g SSE at Si, damping. 7.2 Analytical Techniques The redistribution analyses were perfonned by applying" an equivalent set of static loads to the existing structural complex and determining the corresponding forces in the structural members. The equivalent set of static loads was obtained by applying a system of loads proportional to the first N-S mode inertial loads. The scale factor for these inertial loads was obtained by equating the base shear in the Western one-half of the structural complex (from Column Line H to R) to the corresponding base shear from the SRSS forces, which is 11,910 k1ps as shown in Table . 3-5. The force in the various members due to the equivalent static loads is shown in Tables 71 and 7-2. These results are to be compared with those shown in Tables 3-5 and 3-7. The ccmparison shows excellent agreement for all N-S walls at both levels. For the E-W walls, the comparison is not as good, but this is to be expected :,ince, for these walls, the first N-S mode is not the only irrportant mde as it is for the N S walls. In all the redistribution analyses, an equivalent static load system of 1.05 7-2
times the first mode inertial loads was used. In assessing the effects of bp redistribution, the comparisons should be trade with Tables 7-1 and 7-2 which are the base cases for redistribution. The analysis technique used a bilinear capacity or resistance relationship. !
, Reference values will be given for capacities; this can be considered as !
the value where the elastic limit is reached. Af ter exceeding this j value, there is a slight increase in force developed since a finite slope following the yield is used in the analysis, t 7.3 Detailed Description Of Cases l The following cases were analyzed using the technique described in Section 7.2. [ a) Case 1: The capacity of Wall I was limited to 2800 kips between elevation 45' and 61'. This low value was chosen to investigate whether single major walls have a significant effect on the overall system I of walls. All other walls had no limit on capacity. The results of this case are shewn in Tables 7-3 and 7-4 and can be compared to Tabies 7-1 an6 7-2. b) Case 2: The capacity of Wall 3 was limited to 2000 kips between elevations 61'and 77'. This low value was chosen to investigate if l- single major walls have a significant effect on the overall system of walls. All other walls had no limit on capacity. The results of this case are shown in Tables 7-5 and 7-6 and can be compared to l Tables 7-1 and 7-2, 7-3
O c) c se 3: '** "ec t of va rt ' ce ' > ^ r t r> "a r **' >= '"ed - ' " - l intersection of Walls 1 and 13 from elevation 45' to 61' was ! separated so that vertical shear could not be transferred. The results are discussed in Section 7.4 below. d) Case 4: This case considered all walls limited to the capacities
. presented in Tables 4 1 and 4-2. The results of this case are presented in Tables 7-7 and 7-8 together with the base case loads l and the capacities.
e) Case 5: This case considered all walls limited to the capacities presented in Tat,les 4-1 and 4 2 except for Wall I between elevations ' 45'and (1' which was limited to 3350 kips. This case assumes a local f i lack of transfer to the foundation for the North pier in Wall 1. This condition is highly unlikely since the pier is Ottar',ed to the grade beam rhich can, in turn, transfer force to 'the comoression zone. [ The results of this case are presented in Tables 7 9 and 7-10 together with the base case loads and the capacities used in this analysis. ; i 7.4 Results I For Case 1, (Tables 7-3 and 7-4), which initially limited the shear capacity f l at elevation 45' 61' for Wall 1 to 2S00 kips, increases occurred in other !
, shear walls oriented in the same direction frcm elevation 45' to 61'. The
{ l total shear force was reduced indicating that it is resisted in the Eastern l i portion of the Auxiliary Building with a slight increase at the Fuel Building; but increases in the fuel Building walls are insignificant compared to their capacities. The shear forces increased in the trajor E-W wils. l i Tne effect at elevation 61' 77' was basically a decrease in the force en Wall 1 at that leni and an increase in forces on other walls. The general j l ; i , l 74 l l;
{) conclusion is t. hat, as the capi, city of one wall is decreased, others will resist the load. For Case 2, (Tables 7-5 and 7-6), the structure behaved similarly to Case I with other walls increasing slightly to resist the ' plied loads. For Case 3, the total vertical shear throughout the height of the stnJcture decreased from 1680 kips to 1280 kips, and the membrane force in Wall 13 ' increased by the difference in shear. The overall moment in Wall 1 increased from 101.000 k-f t to 112,000 k-ft. As illustrated, the systen is not i particularly sensitive to local shear transfer. Case 4. (Tables 7 7 and 7-8), illustrates the effect of smsll wall yielding. : The major change occurred in Wall I at both elevations 45' and 61'. The O- force increased from 4170 kips to 4320 kips at elevation 45' and from 3990 to 4100 kips at elevation 61'. This results primarily from the limited . t capacity of Walls 2 and 3 at elevation 45' to 61'. The slight increase ' in capacity in the small walls resulted from the fact that some stif fness was used in the computer analyses for these walls even though the initial capacity was set at the values shown in the table. Slight decreases in ; the total shear for the sum of the N S walls indicates that the increment l of shear is being resisted by the E-W a; */or N-S walls towan2 the Fuel f Sullding end of the complex, i Case 5 (Tables 7 9 and 7-10), wnich limited outer Wall 1 to 3350 kips l l at elevatien 45' to 61', showed that the increment of force previously ; resisted by this wall is now being resisted by other parallel shear walls ' and also the perpendicular walls which sid in resisting torsion. l t 7- 5
(' Tables 7-11 and 712 sumarize the results of the Base Case and Cases 4 and 5 for ease of comparison. The general trends in load redistribution can be seen from these tables. Capacities are not given for the East walls. ! since with the exception of scue smaller walls, tne large walls are lightly loaded relative to their respective capacities. 7.5 Displacements
. Table 7-13 shows the elastic displacements for the various load cases. The f values at Column L.ine 55 are for the South. West corner with motion in the North-South direction. The values for Walls 2 and 3 are the relativa f values between elevations 45' and 61'. The values are primarily presented for comparison purposes, since they do not include the cracked stiffness 2
i
. of the structure. Displacement predictions considering concrete and '
masonry nonlinear behavior are determined in Appendix 0.
- O !
F Comooring the fixed base value to Case 4 at elevation 117' indicated an i increase by a factor of 1.08. From these values, the amount of yielding can also be estimated for the smil walls with capacities in the range of 400 kips, t
- Wall 2 had a capacity / lead ratio of .60 (Taole 5.') for the fixed tase ;
I case at a displacement of .026" (Table 7-13). Therefore, the
. elastic displacement when the capacity would equal the load is
(.60)(.026) = .016". For Case 4. the displacement is .030" or ( ) = 1.88 times the value when capacity is reached. Based on the hending capacity criteria, the cuter reinforcing steel is allewed to te I at two times yield. Therefore, the maximum strain is estimted to be ! 2(1.88) = 3.76 times yield. Doing the same tyce of calculations for ! Case 5 leads to 6.4 times yield for the outer fiber reinforcing steel. . O ! I
. 76 F l
I
u 4
,m 7.6 Summary Q) Avarietydfcaseshavebeenconsidered. The structure demonstrates en excellent ability to redistribute loads, in the event that some walls yield. The analyses have shown that even after redist'ribution, the
. displacements only increased by about 20%. Finally, the analyses show that the small walls, governed by moment, will react afacity first and then a-yield. However, the outer reinforcing steel will only be abou 6 times
. yield, and the compression stresses in those walls will be quite low relative to the ultimate strength.of the concrete..
.o S
7-7
a s Table 7-l' Base Case Shear Forces, N-S Motion. Elevation 45' SHEAR F0RCE WALL NUMBER (KIPS) LOCATION 1 4170 N-S Walls 2 800 3 5% 4 2270 5 3030 6 330 7 470 8 260 _[ = 11920
-O 9 95o 5-w we " '
10 880 11 70 12 110 13 660 14 0 15 210 .
- . i f
l* v . , .1 i h i O l r p .
I Table-7 8:5e Case Shear Forces, N-S Motion - :
. Elevation 61' SHEAR FORCE WALL NUMBER _ (KIPS) LOCATION 1 3990 N-S Walls [
2 58 0 3 3130 4 1940 '-
, 5 430 6 550
[=10620 '!, I I 7 1030 E-W Walls 8 '060
.O 5 53o i
k 10 80 : 11 420 i t 12 230 O l' _ m
o Table 7-3 ' Case 1 Shear Forces, N-S Motion,' Elevation 45',
- Based On Limited Capacity Of Wall 1 From
~ (]a Elevation 45' To 61' SHEAR FCRCE WALL NUMBER (KIPS) LOCATION
, 1 2840- N-S Walls 2 1080 3 700
. 4 2450
~
5 31 90 6 340 7 480 8 270 [=11350 4 9 1340 E-W Walls 10 1040 11 60 12 210 13 809 14 20 15 210 O .
Table 7-4 Case 1 Shear Forces, N-S Motion, Elevation 61',
,, Based On Limited Capacity Of Wall 1 From
() Elevation 45' To 61' SHEAR FORCE WALL NUMBER (KIPS) LOCATION 1 3630 N-S Walls 2 M0 , 3 3300
. 4 2040 5 440 6 560
[ = 10650 7 1190 E-W Walls
', 8 1130
\
(d 9 680 10 90 11 430 12 280 l t G S
q, I-Table 7-!L:: Case 2 Shear Forces, N-S Motion. Elev. 45', Based On Limited Capacity Of Wall 3 i '
,. From Elevation 6i' To 77
b SHEAR FCRCE WALL' NUMBER (KIPS) LOCATION 1 -4230 N-S Walls
. 2 810 3 580.
4 2050 a , 5 3080 6 330 7 490 8 270 [=11840 9 940 E-W Walls 10 890 11 260 12 110 13 670 14 10 15 220 e i e i d O 1
l Table 7-6 Case 2 Shear Forces N-S Motion, Elev. 61' . Gased On Limited Capacity Of Wall 3
,-y From Elevation 61' To 7 7'
\_)
SHEAR FORCE WALL NUMBER (KIPS) LOCATION 1 4050 N-S Walls
, 2 610 3 2MO 4 2240 5 470 6 580
[.9990 7 1020 E-W Walls 8 1080 9 680 l] 10 140 11 460 12 280 4 a ,
'(_)
s
r-- I. $ Table 7-7 Case 4 Shear Forces, N-S Motion, Elevation 45' Based On Capacities From Table 4-1 WALL BASE CASE SHEAR NEW SHEAR CAPACITY NUMBER FORCE (KIPS) FORCE (KIPS) CONSIDEREG (KIPS) LOCATION 1 4170 4320 5390 N-S Walls 2 800 51 0 470 3 590 500 490 4 2270 2410 3810 5 3030 3250 5970 6 330 120 110 7 470 440 420 () 8 260 70 60 [=11920 ['11620 [ = 16720 9 960 1200 4730 E-W Walls 10 880 680 5560 11 70 90 240 12 110 20 420 13 660 590 9350 14 0 5 170 15 210 200 760 O
Table 7-8 Case 4 Shear Forces, N-S Motion Elevation 61' Based On Capacities From Table 4-2 WALL BASE CASE SHEAR NEW SHEAR CAPACITY NUMBF.R FORCE (XIPS) FORCE (KIPS) CONSIDERED (K!PS) LOCATION 1 3990 4100 51 00 N-S Walls 2 580 260 190 3 3130 3190 4520 4 1940 1820 2240 5 430 400 650 6 550 520 750 [=10620 [
- 10290 {=13450 p 7 1030 820 4440 E-W Walls LJ 8 1060 830 9340 9 630 690 2820 10 80 120 1380 11 420 420 2390 12 230 180 2390
{=22760 O
r - _ ,) Table 7-9 Case 5 Shear Forces, N-S Motion, Elevation 45' Based On Limited Capacity Of Wall 1 From Elevation 45'-61' WALL BASE CASE SHEAR NEW SHEAR CAPACITY NUMBER FORCE (XIPS) FORCE (KIPS) CONSIDERF.D (XIPS) LOCATION 1 4170 3410 3350 N-S Walls
, 2 800 480 470 3 590 490 490 4 2270 2720 3810 5 3030 3500 5970 6 330 110 110 7 470 430 420 8 260 60 60 c
js
'q) [=11920 [=11200 [ = 14680 9 960 1390 4730 E-W Walls 10 880 1080 5560 11 70 90 240 12 110 80 420 13 660 890 9350 14 0 20 170 15 210 170 760
i
'O Table 7-10 Case 5, Shear Forces, N-S Motion, Elevation 61' Based On Limited Capacity Of Walls From .
Elevation 45'-61' ' t
-WALL BASE CASE SHEAR NEW SHEAR CAPACITY NUMBER FORCE (KIPS) FORCE (XIPS) CONSIDERED (KIPS) -LOCATION 1 3990 4200 5100 N-S Wall's !
, 2 580 210 190 3 3130 3300 4520 4 ~1940 1850 2240 5 430 -430 650 6 550 520 750
{=10620 [=10500 [ = 13450
.7 1030 1150 4440 E-W Walls ;
'8 1060 960 9340 9 630 750 2820 10 80 130 1380 i1 420 400 '4390 ,
, 12 230 210 2390 1 4 O ! i
N,) Tat,.e .-Il Sumary Of Cases, Elevation 45'-61' (All Values In Kips)
,WALL BASE CASE CASE 4 C ASE 5 NUMBER LOAD LOAD / CAPACITY LOAD / CAPACITY 1 4170 4320/5390 3410/3350 2 800 510/470 480/470 3 590 500/490 490/490 4 2270 2410/3810 2720/3810 d 5 3030 3250/5970 3500/5970 9
=
6 330 120/110 110/110 7 470 440/420 430/420 8 260 70/60 60/60 () 11920 11620/16720 11200/14680 9 960 1200 1390 10 880 680 1080 11 70 90 90 e 12 110 20 80 d g 13 660 590 890
, $ 14 0 5 20 15 210 200 1 90 (s)
I ) Table 7-12 Sumary Of Cases, Elevation 61'-77' (All Values In Kips) WALL BASE CASE CASE 4 CASE 5 NUMBER LOAD LOAD / CAPACITY LOAD / CAPACITY I - 1 3990 4100/5100 4200/5100 2 580 260/190 210/190 g 3 3130 3190/4520 3300/4520 N 4 1940 1820/2240 1850/2240 m s 5 430 400/650 430/650 6 550 520/750 _ 520/750 10620 10290/13450 10S00/13450 ,
'~
7 1030 820 1150 8 1060 830 960 m 9 630 690 750 d 9 10 80 120 130
$ 11 420 420 400 12 230 180 210 o
k v
1
,/--
Table 7-13 Elastic Displacements t FIXED BASE BASE CASE CASE 4 CASE 5 South-West Corner . Elevation 61' .036 .038 .039 .0 90 77' .072 .076 .078 .118 117' .145 .156 .156 .187 Wall 2 At Elevation 59' .026 .027 .030 .0 51 Wall 3 At Elevation 59' .022 .023 .027 .0 41 All values in inches.
,, )
4
x
. 8. TRANSFER OF SHEAR FORCES TO ROCK FOUNDATION P .The Control Building, like all other Category I structures in the Trojan i plant, is founded on rock. . The columns of the steel framing system for the Control Building are supported by spread footings and the shear vsalls by 4 . reinforced concrete grace beams. The grade beams are interconnected by the reinfcrced concrete ground floor slab. The foundation systen is shown in Figure 8-1. The dead load of the floors is carried mainly by the steel 9-: .franing system and the spread footings. However, since the shear walls were built from the ground floor up, it is apppropriate to assune that their weight is carried by the grade beams. Floor loads imposed after the shear walls l were built are carried partially by the shear walls and partially by the
'~ steel columns. 1 In the Control Building, lateral forces are resisted by the shear walls and 1 O are transferred to the rock foundation through the following means: i ! 1. Grade Beam To Rock Shear Resistance, The greatest resistance to sliding is provided by shear resistance
- between the bottom of the grade beams and the rock. Shear resistance to sliding (S )rcan be expressed using Coulomb's formula as:
Sp = C+u0 l where: e C = cohesion at zero confinement (psi) a = coefficient of friction 0 = dead load (psi)
- O 8-1 4
The above formula is used by the U.S. Corps of Engineers and U.S. Department of the Interior, Bureau of Reclamation to determine the resistance to sliding of concrete dams (Ref.11 It is also used by other organizations in concrete gravity dam design. To develop the value of C, it was conservatively assumed that the foun , dation under the Control Building is composed entirely of flat-lying
- vulcanic tuff having an unconfined compressive strength of 1225 psi, the average of the actual laboratory test results. (See FSAR, Section 2.5.1.5, "Foundation Evaluation".) Based on the average compressive strength of 1225 psi, the design cohesion value of the material is estimated to be in excess of 130 psi. This conservative assessment is based on laboratory results of the unconfined compressive tests of volcanic tuff materials and on triaxial test results on rocks of similar type and strengths. This approach is very conservative for the following reasons:
- 1) The in-situ rock at the plant site is much stronger than the volcanic tuff raaterials used in the laboratory tests.
- 2) No credit was taken for favorable attitude of the rock units underlying the foundation.
- 3) The 130 psi value would be used for normal design conditions.
Usually, higher values are accepted for the more extreme conditions of a large earthquake (SSE). That is, these are normal design working stress levels, not ultitute. 8-2
L
- l. ,,
s_/ The coefficient of friction, u, for this foundation material is estimated to be at least 0.7. This value is a. reasonable value for this rock with jointing and laminations. The dead load, O, used in con,iunction with the coefficient of friction.
~
includes the weight of the shear walls and the grade beams, but includes no contribution from the dead loads and live loads of the floors. It should be noted that Winterkorn and Fang (Ref. 21 indicate allowable cohesion stress values between 0.03f{ and 0.05f'. By using the average of 0.04f', a value of 120 psi is obtained. This compares well with the 130 psi value used in tM S analysis. O ^ve'i bie res'ste"c to s'4d'"9 <ro= =" r res'=t>"ce is s've" 'a 7 6'e 8-1 under the heading "shear".
- 2. Column To Spread Footing Friction Friction between the steel columns and the co.'. crete spread footings pmvides additional resistance to sliding. A coefficient of f riction equal .to 0.7
, between the steel column base plates and concrete was used in calculating the available resistance to sliding. This 1* given in Table 8-1 under the heading "footing".
The results of the analysis of the Control Building presented in Table 8-1 reconfirm that the seismic forces can adequately be transferred to the rock foundation. O 8-3
u/ REFERENCES [1] Design Standards No. 2 Treatise on Dams, Chapter 9. "Gravity Dams", U.S. Department of the Interior, Bureau of Reclamation. [2] Winterkorn and Fang, Foundation Engineering Handbook, pp. 610-611. O i I > e 1 t O l 84 i i l
m a 1'
,s-m) 41
=
31- 0, 4fo
- 91- 0o.
Si 3/-O,
- SS t '\
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. h; ,
i l
' I ;l I !!
l - ~ ~l li : Q I i
+
l i ;l . I !l rc=:=:==::======i l l
.il i
- .. ,. __ ___.---- u4 ,i Oy ,
a., io C_._.______.;-._____i, . m ,,. ] m i I ,I i j . ! ' g m i . l> l I ' l t
.__) i I
l pd ' l 1 1-, m(.. I
! i il e - .
,w.___.
i 1 i
. i at QM
, 'O ,_, C " ~ "' j ..i ' IIf+-l M, I.
't j. <
t l 4 w, ~
!! l l lg .
, I>
4' ! idJ"I '-l T- -
, i 1 ,
,_ . J. ' cj! i L;;, a f. ; . ----i. .
li, I
' I ;,
l' I I i
'3 l[ 1 i I ,
l t
. tf l 1l l M;l "- '
i : i i! .,!- i
!,J!l.L l1 i l T l t ici l"j i i l
( [; , I i I l if l b I i I g- i . ilJi v . I i-----
-- i --
_-- - -r i l _. _ p _._ __ . i _ l l 4- ' ( <} O mi i
=
f Figure 8-1 Foundation System
,7 )
Table 8-1 Sliding Resistance And Bose Shear - North-South! WALL IDENTITY RESISTANCE (KIPS) BASE PER PER SHFAR2 FACTOR OF
, FIGURE 8-1 FIGURE 3-2 SHEAR FOOTING TOTAL (KIPS) SAFETY R 1 449G 860 53 50 4320 1.24 0 2 1090 225 1315 51 0 2.58 N' 3 1010 225 1235 500 2.47 N 4 4680 1350 6030 2410 2.50l i
Between L&N 5,6,7,8 7790 ---- 7790 3880 2.00 Between K&L3 TOTALS: 19060 ., 2660 21720 11620 1.87 O Notes: I Values for the E-W direction are not given because the factors of safety are higher. 2See Case 4, Table 7-17. 3These two walls were modeled in the STARDYNE analysis as a single wall designated No. 8 and rnoved to Column Line L. 4 O
..)
APPENDIX A COMPARISON OF SEISMIC ANALYSES
- 1. INTRODUCTION
. Seismic forces due to the horizontal SSE were determined for the Trojan Control-Auxiliary-Fuel Building Complex based on four separate seisnic analyses and one re-evaluation of the original seismic analysis. Due to the differences in thc analyses, model assumptions and methodologies, different analyses resulted in different respoase forces in the structures. This appendix explains the differences in the analyses and discusses the differences in the responses.
The four seismic analyses and the re-evaluation are designated for the purpose of
,m discussion as the original, the re-evaluation study, the TABS, the STARDYNE
(" ) flexible-base and the STARDYNE fixed-base. The original analysis was completed in 1971; the re-evaluation study in April 1978, TABS in June 1978, and the STAR 0YNE flexible-base and the STARDYNE fixed-base analyses in August 1978. All analyses were linear elastic analyses based on the uncracked structural stiffnesses and used the modal spectraf response technique. The differences , among the analyses are described in the following sections and summarized in Table A-1.
- 2. DESCRIPTION OF SEISMIC ANALYSES 2.7 w iginal Analysis The Control, Auxiliary, and Fuel Buildings form an L-shaped structural l complex. These three structures are tied together by common floor slabs l
at elevations 61', 77', and 93'. The original seismic analysis was perfomed based on a 30 fixed-base beam stick model consisting of A-1
(D
) essentially four sticks representing the Control Building, the Auxiliary Building, the fuel Building Hold-Up Tank, and the Fuc! BuiM;ng Spent Fuel Pool. These sticks were tied together laterally through beam elements representing the common floor slabs.
The lateral and torsional stiffnesses of each stick were detennined based on the uncracked structural stiffnesses of each respective structura. except for the Auxiliary Building stick, which was assumed '.o have no lateral resistance. Therefore, the lateral resistance of the entirt buildin'f complex is shared between the Control Building and the Fuel be11 ding. The masses of the structures were lumped at each floor level for each stick. The total weight of the lumped masses included in the model was O 58 79o xias-The modal analysis of the model indicated the fundamental node frequencies of the structure at 6.2 cps for the N-S direction and 6.9 cps for the E-W direction. The modal effective weights of the fundamental modes account for about 60% of the total weight of the structure. Spectral responte analysis was used to ebtain the modal responses. The total responses were determined by combining the modal responses by the sum-of-absolute-values (ABS) method. O A-2
k The total base shears for the N-S 0.25g SSE with 5% structural damping
~
were 14.210 kips for the Control Building and 11,930 kips for the Fuel Building, giving a total base shear of the entire structure' of 26,140 kips which is 38% of the total weight. The output of the original seismic analysis was converted to force and moment in walls and floor slab by hand calculations. Floor shear forces were distributed according to the relative rigidity of the individual walls; the center of the mass and center of rigidity were considered. 2.2 Re-Evaluation Study The original analysis was reviewed in April 1978. The weight of the Control Building was re-estimated based on as-built conditions and found to be 13% lower than'the values used in the original analysis. The O origin 1 an lysis could have used the square-root-of-the-sun-of-the-squares (SRSS) method, but used the ABS method. In the re-evaluation study, the SRSS method was used. The total response of the Control Building was reduced from the original by 20% to account for the difference between the ABS and the SRSS modal response combination techniques. Applying these two reduction factors to the original analysis base shear resulted in (.87)(.80)(14.210) = 9890 kips. The Fuel Building had a 28% reduction due to SRSS and the base shear was (.72)(11,930) = 8590 kips. O A-3
C' 2.3 TABS Analysis V) This analysis was perfonned using the computer program TABS (Three-Dimensional Analysis of Building Systems) developed by Dovey and Wilson of U.C. , Berkeley (Ref. 1). . The model used for this analysis was in accordance with the basic assurrptions of the TABS program; that is, the building system is idealized as an assemblage of a system of independent plane frame and shear wall elements intercornected by floor diaphragms which are rigid in their own plane. This simplification reduces the global degrees-of-freedom of the building system to three per floor: two horizontal translations and one twisting rotation. The output of the TABS program are moments and shears in the walls, rh C Since shear walls are treated as independent plane elements, the flange effect of cross walls or the beari-like behavior of box-type shear wall systems such as the Fuel Building Hold-Up Tank and Spent Fuel Pool cannot be modeled completely. . The masses of the structures were lumped for each floor at the center of mass of that floor. The total weight of the lumped masses included in the model was independently estimated as 70,200 kips. The modal analysis of the model indicated four dominant modes, having frequencies at 5.2. 7.2, 12.1 and 13.5 cps, respectively. The first two &niinant modes have nearly equal participation in the N-S and E-W directions. The third mode is predominantly an E-W mode while the fourth mode is a N.S mode. The modal effective weights A-4
~
(~ of the four modes are 21, 29, ; 26%.of the total weight, respectively, for the N-S direction; they are 24, 23, 33, and 0.3% of the total weight, respectively, for the E W direction. The seismic modal responses were obtained based on the spectral response
, technique. The total responses were obtained by combining the modal responses using the SRSS method. The total base shears for the N-S, 0.25g SSE, 5% damping were 6830 kips for the Control Building and 8260 kips for the Fuel Building, giving the total N-S SSE base shear for the entire structure complex at 15,090 kips which is 22% of the total structural weight. When the results were reviewed, it was concluded that the original and the re-evaluatic.n study forces were conservative.
O 2.4 STAROYNE Flexible-Base Analysis This analysis was performed based on a 30 finite element model of the entire building complex. All the shear walls and floor slabs were modeled by STAROYNE plate elements. Thus, the floor flexibility effects, in-plane and out-of-plane, were included in the model. To account for the effect of rocking due to the foundation flexibility.
, vertical foundation springs were used under the Control Building to connect the base nodes of the finite element model to a fixed boundary. The stiffnesses of these springs were derived based on the BC-TOP 4A [Ref. 2]
formula. The entire model consists of 6ES plate elements and 56 beam elements. The masses of the structure were re-estimted in detail and were lumped at 347 mass points distributed throughout the structure. The total weight of the lumped masses in the model was 71,160 kips. A-5
r-The modal analysis of this model indicated five dominant modes having (v] frequencies at 6.4, 8.4, 9.3,11.7 and 12.4 cps, respectively. The modal effective weights of these five modes are 41. 5, 0.3,19,17 and 0% of the total weight, respectively, for the N-S direction; and are 1.2, 69, 0, 0, and 15% of the total weight for the E-W direction. Therefore, the first,
, third and fourth dominant modes were N-S modes while the second and fifth dominant modes were E-W modes. The first N-S mode was also a twisting mode - pivoting about the more rigid Fuel Building with major modal deflection at the Control Building end in the N-S direction.
The seismic response was determined based on the spectral response technique using the SRSS modal response combination method. The total base shears for the N-S, 0.259, SSE, 5% danping were 12,200 kips for the Control Building and 7870 kips for the Fuel Building, giving the total N-S, SSE V base shear for the entire structural complex at 20,070 kips, which is 28% of the total structural weight. 2.5 STAR 0YNE Fixed-Base Analysis This analysis was based on the same model as the STARDYNE flexible base analysis described previously, except that the fixed-base condition was prescribed and the vertical foundation springs under the Control Building were eliminated. U A6
-3 The modal analysis in this case indicated essentially the same modal b behavior as that for the flexible base. The total base shears for the N-S, 0.259 SSE, 5% damping were 11,910 kips for the Control Building and 7680 kips for the Fuel Building giving a total of 19.590 kips for the entire structural complex which is 27% of the total weight. The total shear obtained from the fixed-base analysis differs from that of the e
flexible-base analysis by only 3%. indicating that the foundation flexibility effect was negligible.
- 3. DISCUSSION 3.1 Comparison Of Model Behaviors All analytical models used predicted a relatively consistent fundamental mode frequency which is between 6 and 7 cps, except for the TABS model O whicn aredicted the freavenci et s.2 cas. This 4neicetes thet the 1^85 model is softer in lateral stiffness. This is due in part to the TABS model assumption that shear wall elements are plane elements independent of each other, i.e., the "flange" effect of cross walls and the composite behavior of box-type wall systems cannot be representatively modeled.
, All models used also showed a consistent trend that the horizontal seisaic response (N-S or E-W) was governed by one predominant (N S or E-W) mode.
except for the TABS nodel which showed three dcminant modes, two of them having nearly equal participations in both the E-W and N-S directions. O A-7
C; 1 The STAROYNE flexible-base and fixed-base models are the most mathematically O- refined models among all models considered. The models can consider not
- only the gross structural behavior but also the local element behavior of individual walls, frame memb-s, and wall systens. The STARDYNE model alto f aclude the floor flexibi'ity effect with more mathematical preciseness !
that the other models and directly predicts the elastic distribation of i s t
%rponse forces. !
L j 3.2 Weight Determination 7 The original seismic analysis used a total weight of the model for the , entire complex of 68.790 kips. A weight calculation of the as-built l 1 , j condit ton for the re-evaluation study was made in April 1078 for the , r l Control Building area. This calculation indicated a weight 13 lower than , 4 : j that used originally; The weight determination for the TABS analysis was e ) made independently based on the drawings and live load requiremmts. This j detemination was considered as approximate since it did not consider [ j actual as-built conditions, such as the actual equipment weight. Since a i a STARDYNE analysis was going to be used to predict the interaction of j the total complex (existing Control, Auxiliary, and Fuel Buildings), and l 1 the new structural support system, a new weight calculation was made for [ ] i . the entire complex, and the information was used to determine the mass ; t ! in the STAROYNE model. A comparison of the we'j ' used in the STAROYNE ! analyses with that used in the criginal analysis is as follows: i ? i ! i i ! 1 t i !O i A.8 i 4 l t
._._.__--____--______-----,N
q
' I
- 1) The Control Building weight in the STAROYNE analyses,is 8.2% lower than~
that used in the original analys'.. ; i
- 2) The Auxiliary Suilding weight in the STAROYNE analysss is 11.3% i l higher than that used in the original analysis, i; 7
- 3) The Fuel Building weight in the STAROYNE analyses is 4.6% higher than i
, . , that used in the original analysis. !
+
- 4) The total weight of the complex in the STAR 0YNE analyses is 3.5%
higher than that used in the original analysis. ; I j 3.3 Foundation Flexibility Effect 3 The comparison of the STAROYNE flexible-base and fixed-base results shows O t8at the foeeeatioe fiexibiiit, effect is insion4ficant and that a fixed-base analysis is appropriate considering the fact that the foundation f a rigidity of the Trojan Plant is very high (the shear wave velocity of the (r site rock is 5500 ft/sec). L i I 3.4 Comparison Of Results
. The original analysis predicted the highest total N-S SSE base shear
- (26,130 kips) which is mainly due to the sum-of-absolute value (ABS)
( modal response combination technique. The re-evaluation resulted in a total base shear of 18.480 kips. The total base shear of 18,480 kips is close j to the 19.590 kips predicted by the STAR 0YNE fixed-base analysis. The TABS analysis resulted in the lowest total N-S SSE base shear (15,090 kips), f ! All analyses predicted a higher proportion of the total base shear to be O esisted bx the Ceetrol 8uiidins then bx the Feei 8efieies, excent 18e Tass ! A-9 j. I
-,_____..._--______,,_.____._______._______,,____m_.__._.,
(~w analysis whiu, predicted a higher proportion of base shear to be resisted d by the Fuel Building. This appears to be directly related to the particular model behavior of the TASS model as explained previously. Despite <his higher proporttin. the absolute value of the base shear resisted by the Fuel Building as identified by TABS is still lower than base shears predicted
, by the Jther analyses.
The STAR 0YNE analyses, both the flexible- and fixed-base cases, resulted in a higher proportion, i.e., 61%, of the total N-S base shear to be resisted by the Control Building, as compared to 54% frcen the original analysis and 45% from the TABS analysis. This is mainly due te the increased torsional flexibility inherent in the STARDYNE model as reflected in the mode shape of the fundamental N-S mode described previously. This torsional flexibility appears to be due to the combined effects of floor flexibility and increased relative lateral rigidity of the Fuel Building as a result of the finite element modeling of the box-type Hold-Up Tank and Spent Fuel Pool shear wall systems.
- 4. SumARY The inertial forces are summarized in Table A-1. The STARDYNE flexible base analysis had undue conservatisr.: due to the lack of radiation damping. The TABS analysis was used to gain confidence that the re-evaluation was conservative.
The base shear predicted by the original analysis is higher than that predicted by other analyses. O v A-10
The re-evaluation study results are within 6% of the STARDYNE fixed-base analyses for the total base shear. This is an excellent agreement considering the difference in mathematical sophistication that has evolved from 1971 to the present. The major differences result from elastic load distribution which has led to a load increase in the Control Building end of the structure, as reflected
. in Table A-1.
Concrete structures only behave elastically at very low stress levels; beyond such levels, cracking develops and load redistribution occurs. The effects of load redistribution are exanined in detail in Section 7.
, O t
lO l l A-11 h_
r - _ - _ _ _ _ _ _ - _ _ . _ _ _ _ _ _ _ _ i ) REFERENCES [1] Wilson, E.L., and Dovey, R.H. , "Static and Earthquake Analysis of Three Dimenstoral Frame and Shear Wall Buildings". EERC Report 72-1, U.C. Berkeley, 1972. (2) BC-TOP-4A, "Seismic Analysis of Structures and Equipment for Nuclear Power Plants". Topical Report, Bechtel Power Corporation, San Francisco, CA, Revision 3. November 1974. m x G O A-12
i (^') w/ Table A-1 Summary Of Seismic Analyses Total Inertial Load N-S. SSE .259, 8 = 5% ANALYSIS CONTROL BUILDING FUEL BUILDING TOTAL DESCRIPTION SHEAR (KIPS) SHEAR (KIPS) SHEAR (KIPS) Original 14.210 11,930 26,140 Re-Evaluation 9,890 8,590 18,480 TABS 6,830 8,260 15,090 STAROYNE O Flexible-Base (y 8/17/78 12,200 7,870 20.070 STAR 0YNE Fixed-Base 8/24/78 11,910 7,680 19.590 9 O
f l
] APPENDIX B SHEAR CAPACITY CRITERIA
- 1. INTRODUCTION This appendix describes the evaluation of fully grouted hollow concrete block shear wall test results and, based on these test results, the development of shear wall criteria used to evaluate the Trojan Control Building. Two sets of test data are considered applicable to the capacity evaluation of the grouted concrete block shear walls (Refs. I and 21. Although each of these two test series has different test conditions and parameters, the results of both series can be used to derive realistic criteria for the determination of concrete block shear wall capacity.
Empirical relationships arrived at by Schneider (Ref.1], and recent concrete f')
'~
block shear wall test results obtained by the University of California, Berkeley (Ref. 2], are used herein for the basis of criteria. Cyclic degradation is included in the c-iteria by cor irison with cyclic tests. In addition, dowel capacity criteria are diset sed and comparisons are made. Finally, the criteria are given for horizontal and vertical shear capacity. l 2. DESCRIPTION OF TEST SET-UPS AND CONCLUSIONS 2.1 Schneider's Tests l i Tests were performed in 1969 at California Polytechnic College, Pccona, l l California, by Prof. R.R. Schneider (Ref.1) . Vurious parameters were l investigated in these series. These parameters can be cited as aspect ratio l (H/W), vertical and horizontal reinforcement ratios (cyand ch), and grouting. The testing scheme is shown in Figure B-1. B-1 l l [
(] Based on all of the test results, Schneider developed the following empirical relationships: H/W ULTIMATE SHEAR STRESS (PSI) e
-l 0.2 < H/W < l .0 V/TW = 310-175H/W 3v g$$ 1.0 < H/W < 3.0 V/TW = 152.5-17. 5H/W 5'*
g j 'o 3.0 < H/W V/TW = 100
- 0. 2 < H/ W < l . 0 V/TW = 347.5-112.5H/W
-{ l .0 < H/W < 3.0 V/TW = 290-55H/W 3.0 < H/W < 4.0 V/TW = 200-25H/W x-a j j '= 4.0 < H/W V/TW = 100 where:
) H = height of wall (in)
W = length of wall (in) T = gros 2 thickness of wall (in) V = capacity of wall (lbs) Schneider's test results used herein are tabulated in Table B-1.
. For those tests with N = 0 (Figure B-1), the test report (Ref.1) infers some axial compression. Although the induced axial force into the pier is not documented [Ref.1), several simplified linear analyses (with variable specimen modulus) were made by Bechtel with a finite element computer pro-gram. These showed that 25% to 40% of the vertical component of the diagonally applied load would be resisted by the two struts in compression, depending on the diameter and length of the struts as well as the stiffness assigned to the masonry portion of the test specimen.
1 B-2 L
19 L' Having determined that the axial stress present in the pier contributed to the ultimate shear stress capacity, the following expression of the general relationship is applicable: Y test *Yshear # ompression (Eq.1) Comparing to the ACI 318-71 equation: 0 0 , vj=3.3 +j=v e
+ (Eq. 2) leads to the conclusion:
a vc'Vc' 9' If e n is assumed to be equal to vj (maximun possible compres:fon in the pier), then: v (Eq. 4) c =0.75vf Therefore, a conservative expression for the ultimate shear stress is: v = 0.75v' + where: (Eq. 5) e = axial stress; + compression - tension n and v' is the ultimate shear stress to be detemined by Schneider's , empirical relationship. This expression is referred to in this supplemental evaluation as the "modified Schneider criteria". The results presented in Table B-1 are retabulated according to specific parameters in Table B-2 where the test results are compared against the modified Schneider criteria. Applying the total axial load (equal to the total shear load) in evaluation of the term en aM tadng E of De shear O stress hx schmeieer s or49 1eei ro -vietioe is cc.eservet4 e. B-3
x 2.2 University Of California (Berkeley) Tests The University of California (Berkeley) has been engaged in masonry struc-tural research since 1972. Over the years several series of tests. including a series on hollow concrete block, have been perfonned. T he hollow concrete block series have been tested with the test set-up shown in Figure B-2. The test results are sumarized in Table B-3. All the HCBL-ll series had cyclic loading with disp'acement controlled rams. Therefore stiffness and strength degradation influences are incorpJrated. The actual height-to-length ratio (H/W) used in this series o is 1.17. The blocks used are nominally 8"=8"=16". The actual width of the ' V wall is 7-5/8". The net-to-gross sectional area of the block is 587,. The block compressive strength is 3000 psi, while the grout used is 4000 psi. The series is piso of interest since smaller reinforcement ratios are represented. The axial compressive stresses am known in each case. The modified Schneider criteria were compared with the results of the U.C. (Berkeley) tosts. The correlation is given in Table B-4. 2.3 Correlation Of Block Wall Test Results The test results tabulated in Tables B-2 and B-4 and thet? corresponding predicted values by the modified Schneider criteria are plotted in Figure B-3. As seen in this figure, there exists a very good correlativa. O B-4
,r'T 3. OTHER CONSIDERATIONS V
3.1 PCA Tests On Reinforced Concrete Shear Walls (Ref. 3) The results of some reinforced concrete shear wall tests with transverse members at both ends acting as flanges perfomed at PCA indicate results which are , relevant to the criteria used here. High shear strengths were obtained i' rom the PCA tests even when horizontal or vertical steel was not present. , The empirical relationship obtained from this series is: v =8 - 2.Srq (h) + ( )+o ny I 9' 0) where: on = vertical reinforcement ratio f = reinforcing yield strength f' = concrete compressive strength O which can be rewritten as: v
- V +V u c s +'n (Eq. 7) where v , v,, and of refer to contribution to the ultimate Jhear stress by concrete, steel and axial forces, respectively. It is recogni:ed in this relationship that the smaller the aspect ratio, the higher the strength developed regardless of the reinforcement.
Using the PCA relationship with a concrete strength of 3000 psi and dis-regarding the reinforcement results in the following: v u =438-137h+ (Eq. 8) O B-5
I
,m
() 9 The modified Schneider criteria for .2,< g < l.0 and horizontal reinforce-ment areas follows: a y u = ./5(347.5-112.5 h) 4+ 2 (Eq. 9) v u =261-84h+h Comparing the ratios of the two coefficients:
= .60 and h = .61 indicates that reduction of 40% has been made relative to the PCA relation-ship. Thus, application of the modified Schneider criteria results in a capacity 40% less than the capacity derived from the PCA relationship.
3.2 UBC 1976 g Although strength design considerations are not incorporated in Chapter 24 of UBC. the allowable stresses for masonry shear walls (Table 24-H of Section 2418) are expressed in terms of M/Vd which is related to H/W, the aspect ratio. The recognition of the increase in strength with the aspect ratio is thus incorporated into the 1976 UBC. This requires only that minimum reinforcement ratios of oy of 0.0007 and h of 0.0013 be provided. The importance of aspect ratios is further discussed in [Ref. 41 3.3 Comparison Of Block Wall Test Specimens With Trojan Walls The Trojan walls essentially consist of high strength concrete core sandwiched between two wythes of fully grouted concrete blocks. The grout in the cells and the core concrete strength is >6000 psi. In each wall, the reinforcement of the concrete blocks when considered over the gross area meets the minimum reinforcement requirement: of the 1957 USC. If only single wythe concrete block is considered, the reinforcerent ratios in B-6 -
m the Trojan walls for 7-5/8" thicx block wythe ranges from 0.0024 (#6/24") to 0.0048 (2#6/24") for horizontal steel and from 0.0017 (d5/24") to 0.0024(#6/24") for the vertical steel. These are within the ranges of the Berkeley HCBL test specimens. Comparison of important material strengths of the Trojan walls with those of Berkeley (HCBL) series and Schneider's test specimens is given in Table B-6. As shown in this table, the actual strength of Trojan wall blocks, mortar, cell grout and core fill concrete exceed those of the test specimens. The Berkeley single wythe specimens used blocks with net block area equal to 53% of the gross. Schneider's specimens had 50%. The t, lock to gross area percentage of the blocks in Trojan walls is 50%. For the thinnest composite walls at Trojan which are 20" thick, the block area over the gross wall area is 40%. For composite walls in Trojan thicker than 20", the ratio of block to gross area decreases further. Therefore, the composite strength is higher in all the composite walls at Trojan than the strength of the blocks used in the Berkeley and Schneider tests, if the core reinforcement is considered as additional reinforcement other than at the discontinuities, the reinforcement ratios then are higher than most of
, the test specimens considered for establishing the criteria.
t 4 HORIZONTAL SHEAR CAPATITY CRITERIA Comparison with both the Sch.' eider and Berkeley tests has established that the modified Schneider criteria correlate very well in tenns of resistance capacity. All specimens had vertical restraint; this primarily was to ensure the specimens failed in shear and were not limited by a bending failure. Bending failures are O arevemtee ie saecimems its io espect reties (1 t"e r=#9e or 2 o or iess) er i either confining them by upper and lower stories or the application of dead load. B-7 l t .
O The major Trojan Control Building shear walls have significant dead load or can mobilize it by small anvements that would lead to unloading of the steel columns and a transfer of load to the shear wall. Therefore, for these types of walls, the modified Schneider criteria are classified as the Msic criteria. Since all the Trojan walls have more horizontal block reinforcement than vertical . block reinforcement, the Schneider criteria with horizontal reinforcement are used. The basic criteria are obtained by reducing the criteria shown on page B-2 5 by 25% and adding the axial load effects. Basic Criteria: V u
- I20 0 #
- I 1
- V =
[218-41 h + ]WT 1.0 5,h5 3.0 O 3.0 4.0 V = (150-19h+j,]WT 5h5 V = [75 + ]WT 4. 0 5h a h * #v 1 002: o y 1 0007 and ph 5 0013 for h 5 2 use f = .2 where: Vy = allowable shear force (lbs) H = height of wall (in) W = width of wall (in) T = gross wall thickness (in)
#n = axial stress (psi) (+ for compression; - for tension) oh = horizontal reinforcement ratio
#v = vertical reinforcement ratio B-8
(-
\s In order to qualify for the basic criteria, the walls rnust be subjected to dead load and have core reinforcing steel. Walls which have structural steel columns completely interrupting the core shall use an h ratio which limits W to the distance between columns.
Additional Criteria: a) Walls which are subjected to dead load, but no core reinforcing steel is assumed, may be evaluated by the basic criteria; however, the stress is limited to 150 psi. b) Walls which are not subjected to dead Icsd, such as interior walls between floor slabs, shall be evaluated using the following shear criteria: (3 . C/ A Vu = $ [2 v' t,ff + [,f (.8W) where:
? = capacity reduction factor (0.85) f' = concrete compressive strength (6000 psi) t,ff= total wall thickness -8(in) f, = yield strength of reinforcement (psi)
. A g = minimum of horizontal or vertical reinforcement (in2/ft)
Other tems have same definition as given under the basic criteria. c) All walls must be checked for combined bending rmeent and dead load effects. O B-9 [
These foregoing criteria are based on an empirical formulation which correlates very well with recent testing. They are conservative in that:
- 1) The criteria are based on masonry testing and the actual structure is a composite of masonry and reinforced concrete.
- 2) The concrete core and cell fill strength of the Control Building walls are much higher than the test specimens.
- 3) The added strength of the structural steel is not included.
- 5. VERTICAL SHEAR CAPACITY CRITERIA Vertical shear capaci ty is considered as resisted by the dowel action of the horizontal reinforcing steel and the steel beam-to-column connections.
4 O B-10
p 5.1 Reinforcing Dowel Capacity Table B-5 shows a comparison of Schneider test results and dowel strength, ! assuming the dowel could develop a strength equal to its tensile strength of A f . The Schneider test specimen had some induced compression and the 3u
.75 factor applied to the test data is a conservative estinate of what the ,
maximum shear strength would be without the compression.
. Based on this information, the dowel strength ranges from .65A I to over su 1.0A,fu, However, as shown in Figure B-4, the test specimen reinforcing was located at the outer edges of the specimen and one of the dowels would !
l pull out as shown in the figure before attaining fOli strength. l Test 84-3 from the PCA tests [Ref. 31 illustrates a situation where the web of the test specimen carried the entire load since there was no horizontal . reinforc ement. In this test, the ultimate shear strength was v
- 810 psi. The test had: :
h=.5 o y
= .005 c
h
=0 l f = 78 ksi , f u= 120 ksi I
The strength based on dowel action at ultimate is: v =of = (.005)(120,000) = 600 psi O l B-Il t
l Since the specimen reached 810 psi there is an increase of = 1.35. Thus, the test exhibited an ultimate shear strength 1.35 times greater than i the calculated ultimate dowel action strength. The cpparent reason
- for this strength increase is the fact that the reinforced concrete specimen had uniformly distributed reinforcing and there is
, most likely some shear friction. !
Nelson stud tests (k . 51 have shown shear strengths of 95% of the tensile ultimate strength for studs in the same size range as the reinforcing , steel in the Control Building walls. [ t l Tests documented in (Ref. 6] have shown for steel embedments welded to l plate that the shear strength for small steel areas in the range of 1 in2 is 70% of tensile ultimate strength. These tests are very conservative since i O they only had one-sided e,nbedment and they also had axial movement. The one-sided embedment results in higher bending moments than two-sided embedment l and a slightly reduced capacity. [ Based on the preceding, it can be concluded that the dewel shear strength of reinforcing bars in the range of 1 in: and less are at least 90*. of tensile ultimate strength. 5.2 Beam Connection Capacity For the strength of beam-to-column connections, twice the AISC - Part I allowables are used. This value is considered as a lower limit of the ultimate strength that can be obtained. O B-12
O 5.3 Vertical Shear Capacity Criteria The capacity of the vertical planes is detemined by combining both the reinforcing steel dowel shear strength and the beam-to-colum connections shear strength. These capacities are compared to the total shear load in the vertical section under consideration. These capacities ; conservatively neglect any capacity of tha concrete and masonry block. i Total capacity = 2(A!SC Connection Allowables)+ (,9)fjA, I where: f = horizontal reinforcing ultimate strength {A, = sumation of horizontal reinforcing steel areas O : I i i i t I B-13 i I i o
~
l 1 ) l
.i N)
REFERENCES [1] Schneider, R.R., "Shear in Concrete Masonry Piers", California State Polytechnic College, Pomona, CA, 1969.
. (2) Mayes, R.L., Clough, R.W., Hidalgo, P. A. , and McNiven, H.D. , "Seismic Research on Multistory Masonry Buildings", North American Masonry Conference, Boulder, CO, August 14-16, 1978.
[3] Barda. F. , Hanson, J.M. , and Corley, W.G. , "Shear Strength of low-Rise Walls with Boundary Elements" ACI, SP53-8. (4) Froerer, 0.0. , "The Benefits of High-Strength Masonry". North Arerican Kisonry Conference, Boulder, CO, August 14-16, 1978. [5] "Design Data, Nelson Concrete Anchor Studs for Securing Steel to Concrete", Gregory Industries, Inc., Lorain OH. l (6) "Containment Liner Plate Anchors and Steel Embedment Test Results", P.L. Chang-Lo, et al., 4th International Conference on Structural Mechanics in j Reactor Technology San Francisco, CA, August 1977. l 4 O B 14
(oJ . P - b
/> . .
i
> I s/
YT P oiAoONAtlOAD PRAME tman onAlt3 VAAT) Figure B-1 Schneider's Test Scheme (Ref.1]
- y. '
da CJ ig! ,3I- 4 ul
'% ty L .i , r'ea m= .cru.,ou
\
4 1 Eio>i * -5 i, i i' w 4043CtLkt
, - e nv.. :
i
,' i.ui o 4
i
' ; .a < =u s dL ' _
- i LA hf,c!
; 3 Figure S-2 Single Pier Test Set-Up (Ref. 2) l O .
1
a s d b Criteria: 300 - y = 0.75v' + ,
- Schneider's Tests
, + Berkeley (HCBL) Tests .
e
+
+
3 *
-, a 200-lv) ,=
5 - e E
+#
g .
. +tNo rein fo rc emen t 100- ,
o - 0 i i 100 200
/' Calculated (Criteria) - v (psi) 6, j u Figure B-3 Test Results vs. Criteria
O 4 Shear Force [
- G 7
%. Crack ,
., JL b W W Figure 8-4 Typical Reinforcing Bar Location At Cross Section (Schneider's Tests)
O O s O
O V Table B-1 Schneider's Parameters And Test Results (Ref.1] TEST y TEST
- u A
H W u c CAPACITY D P PANEL (FT) ( FT) H/W (PSI) (IN2 ) y h ( KIPS) 4 1 1.33 3.33 0.4 260 32 0 .0038 ---- 83 2 5.33 2.00 2.7 131 1 92 .0063 - --- 25 6 5.33 2.67 2.0 124 256 .0048 -. -- 32 7 4.00 4.00 1.0 140 384 .0032 -- - 54 9 1.33 4.00 0.33 2 90 38 4 .0032 - - - - 11 1 12 4 00 4.00 1.0 243 384 .0031 .0048 93 13 5.33 2.67 2. 0 178 256 .0048 .0 048 46 15 6.00 2.CJ 3.0 128 192 .0063 .0048 25 O 28 s 33 2 57 2o 137 256 oo47 ooio 25 31 4.00 4.00 1.0 158 384 .0047 ----- 61 33 5.33 2.67 2.0 97 256 .0070 ----- 25 l 8 2.67 5.33 0.5 203 512 .0024 ----- 104 i I l O
(G) Table B-2 Schneider's Test Results Compared With His Criteria And The Basic Criteria i SHEAR STRESS (PSI)! PANEL H/W o TEST v* C0lHENTS ch u c 1 40 .0038 - --- 260 245 lio horizontal reinforcing H/W < .5 9 .33 .0032 ----- 290 262 8 .50 .0024 ----- 203 216 7 1.0 .0032 ----- 140 136 No5rizontal reinforcing H/W = 1.0 31 1.0 .0047 ----- 158 141 12 1.0 .0031 .0048 243 237 Horizontal reinforcing H/W = 1.0 6 !2.0 .0048 ----- 124 11 9 No horizontal
] reinforcing H/W = 2.0 ,
33 2.0 .0070 ----- 97 112 13 2.0 .0048 .0048 178 180 Horizontal reinforcing H/W = 2.0 167/ . 28 2.0 .0047 .0010 137 121 " 2 2.7 .0063 ----- 131 112 No horizontal i reinforcing ; 15 3.0 l.0063 .0048 127.5 126 Horizontal reinforcing Notes: *v y = 0.75vSchneider
- I'n/4) *h'" #n
- Vtest /4 !
"Second value considers the specimen has no horizontal reinforcement.
i O i i
s . . o b O O O Table B-3 Single Pier Test Results [Ref. 2] VERTICAL REBAR IIORIZONTAL REBAR Mm E L "o ?* Q C O O G #a 3E S
= En o_
o w o bh
= a . a r . sem _5 ne as PE ry 5 SPEClKN L y go. gm gy oE * '"
gp oE **. gg W o$
.g ~mgg {gg x3g
> "mO "m2 Mb2 gag DESIGNATION g gg mg pg 7 ,
> gg gg ,e y rx g g g ,c ggx 6 g og i:CBL.11-1 1.5 F 1330 None ---- - - - - None ---- ----- ---- ----- 45.0 46.6 44.0 11CBL-11-3 1.5 F 1833 2-#5 70.8 .0016 None - - - - ----- - - - -
.0015 45.0 49.1 25.1 I!CBL-11-4 1.5 F 1833 2-#5 70.8 .0016 1-#5 47.9 .0008 :14.7 .0024 53.1 54.9 39.1 itBL-11-6 1.5 F 1833 2-#5 70.8 .0016 4-#5 47.9 .0032 .58.8 .0048 80.0 82.7 52.7
!!CUL-11-7 1.5 F 1905 2-#8 69.2 .0041 None ---- ----- ----
.0041 52.5 65.8 33.3 ItCBL-11-9 1.5 F 1905 2-#8 69.2 .0041 2-#5 47.9 .0016 29.4 .0057 53.8 56.9 41.9 liCBL 11 1.5 F 1330 2-#8 69.2 .0041 4-#6 73.9 .0046 130.6 .0086 84.4 87.7 50.8 l
l i i _ _ _ . _ . . - , _ _ _ - _ . _ , _ . , _ _ _ _ _ _ _ . . _ _ _ _ , . , _ _ _, _ ~ . , , . . _ _ _ _ _ _ _ _ , . - , . _ , , _ _ _ _ _ -
O t 61e B-4 serkeier test aesuits vs criterie BERKELEY M00!FIED SCHNE10ER N , SPECIMEN Yu(TEST) M CRITERI A DESIGNATION P(%) y P (%) (PSI) .75v' (PSI)1 y (PSI) o HCBL-ll-1 ---- --- 123 99 28 127
. HCBL-ll-3 .16 --- 123 99 16 11 5 HCBL-ll-4 .16 .08 145 99 25 124 HCBL-ll-6 .16 .32 21 9 170 34 204 :
HCBL-ll-7 41 --- 143 99 22 121 HCBL-ll-9 .41 .16 147 9 92 27 126 HCBL-ll ll 41 .46 230 170 33 203 Notes: IThe contribution of axial compressive load. 2 This case has been interpreted as having no horizontal i reinforcement. l l 1 l O l
e p. V Table B-5 Comparison Of Schneider's Test Results And Dowel Capacity (No Horizontal Reinforcement. H/W < .05) SHEAR STRESS (PSI)
.75 TEST TEST PANEL H/W oy TEST .75 TEST 00WEL 00WEL DWEC ;
1 40 .0038 260 195 l 302 .65 .86 ! 9 .33 .0032 290 218 254 .36 1.14 8 .50 .0024 203 152 202 .75 1.00 i i i h o ! f E I i i l i l O 1
~ '
* =
fist rict St9cogths s T:ble B-6 i ! TRWAN WALLS SPECII LED AVEI E E BEhrELEY SolHEIDER'S STREldGTH MIfilPUM VALUE ILLL TEST ITEM PARAMETER NOTATION VALUE (IN-PLANE) TESTS SPECIMENS lleavyweight Ita se .ry f* 20008 4100 3000 3000 (Net) concrete bloc 6s compress ive I 1500 (Cross) strength (; si) Light weight Knonry f* 20008 2700 ---- - - - - concrete blocks couvressive stre ngth (psi) Mor'er Martar f 2000 3500 2500 3000 (Avg.) conpressive ASTM strength Type M (28 day psi) Cell Grout Concrete f' 50002 -60003 4000 3000 ( Avg. ) C conpressive Grout: streregth (psi) 1C:35:20 Core fill Concrete f' 50002 ,6000' ---- ---- C concrete congressive strength (psi) Reinforr.ing Yield stress f 40 >45 48 to 71 55 steel Y (ksi) d flotes: 8 Based ori tiet cry;ss-sectional area. 128-day couvressive strength. 3 Based on 90-slay compressive strength. 1
i\ 4 ,, APPENDIX C I STRAIN COMPATIBILITY r
.This appendix addresses the strain compatibility of the various components in the Trojan Control Building walls. The composite behavior requires that shear strains!'n the masonry and concrete are the same. The related stresses, however, will'be _quite different due to the different material properties. The walls are composed of masonry blocks and a concrete core. The cells in the blocks are filled with concrete. For.a typical wall about 50% of the area is masonry and the other 50% is the concrete core.
Both the Berkeley masonry tests (Reference 2 Appendix B) and the PCA reinforced i concrete tests (Reference 3. Append.x B) had shear strains in about the sa.ne O re se w8e# t8e meximem cea cities were ec84eved. !
- The PCA test showed that cracking occurs at a shear strain in the range of 350 u in/in to 400 u in/in for specimens without large amounts of reinforcement and for aspect ratios between .25 and 1.0. The shear strains at maximum capacity rsnged between 5300 u in/in and 8500 u in/in.
The Berkeley tests for specimens with horizontal reinforcing equal to or greater than the vertical reinforcing had shear strains of about 1250 u in/in at the start of nonlinearity and in the range of 6500 u in/in at maximum capacity. Results of some representative test specimens are plotted in Figure C-1 for comparison. From this plot it is illustrated that maximum capacities are reached O et e8out 18e seme s8eer streia eme t8 t for e 9 4vee streia s8eer stresses 4a concrete are substantially greater than the stresses in masonry. The re fo re , C-1
I l l 0 calculations of shear capacities of composite walls based on masonry strength values for both materials are conservative. G 0 O i 4 l 4 f t 3 l O l C-2
r, i , i t t
> e 800 -
PCA(84-3)o = .5%, o = 0% h f' \ , f PCA(B6-4)p y = .25%, o = .50% h 600 _ /
/
/
/ /'\ \
~~,*N s N s y .
N s,
- s '
^O
? .
[' , f N,N, s s '
~.
O ,
/ \, [
/
j d o 400 -
/ \. ,
y //
/
N,N. (
- E */
j,/
* ~ '
200 - f Berkeley (Average 6 and 11)
- Test 6.p y = .16%, oh ' '32%
! l
{ Test 11,py = .41%, o = 46% j h
, , i i
; i i i i i i i 5000 10,000 Shear Strain u in/in ;
, Figure C-1 Comparison Of She:r Stress Vs. Shear Strain i Of Test Specimer s l N i h i m , - . . _ - - . _ _ _ . _ _ _ _ . . _ , , . _ _ . . _ . _ _ _ _ _ _ _ . . . _ _ _ , _ _ _ _ _ . _ _ _ _ , . , . _ . . _ _ . . _ , . . . _ _ _ . _ _ - -
APPENDIX 0 V DISPLACEMENT DETERMINATION In t'is appendix techniques are developed for determination of the displacement of the structure based on.information presented in Appendix C. Figure C-1 is pre-sented again as Figure D-1 with the shear moduli shown for the Berkeley and the . PCA test specimens. Also shown on Figure D-1 is another curve which is considered representative for the composite Trojan Control Building walls. Since the walls consist of about equal portions of masonry and concrete, the composite curve for the system can be expressed as an average of both portions as shown in Figure 0-1. The composite curve was constructed as follows: The average of the initial slopes based on the lower PCA and the Berkeley tests is: Ov G - 1 = ( 8+.09)x106 s .45x10e psi 2 From the Berkeley shear stress-lateral displacement curves it is estimated that the nonlinear behavior of these specimens started at about 100 psi. Tnis was used as the limit for the initial shear modulus. For the second slopa, again the average was used: G2 = (.07+.02)x106 = .045x106 psi 2 The following relationships determine the shear strain (y) as a function of the shear stress (v): D-1
1 y = (.45x105) for 0 j y j 100 ' psi for y >,_ 100 psi y = (.04h 6) + .00022
. Based on the average shear stress per story obtained 'from results of Case 4, Secti)n 7 and the story height the interstory displacements and the total building , displacement can be calcualted.
The following calculations are for the West wall (Wril 1) for the N S .25g SSE at 8 = 5%. The displacement can be obtained by multiplying the stor height
.by the shear strain.
Elevation 93'-117' H = 288"; y = 60 psi
= .04"
- 6) = [ 45 106 } 288 =
Elevation 77'-93' H = 192"; Y = 115 psi 4 = .11" 2
- I(. ) + .000221192
= Elevation 61'-77' H = 192"; Y = 135 psi l
6 3=(p j, 3) + .000221192 = .19" Elevation 45'-61' H = 192"; Y = 215 psi 2 e,=9;s;'g,.00022i192 = .53" O Total Displacement Elev.117' N-S [ = 87" s .9" D-2
,~; Applying the same technique to the North wall using the stresses obtained from
\' /
Section 3 for the E-W .25g SSE at S = 5% results in the following: Elevation 93'-117' H = 288"; v = 25 psi
= .016"
, 6) *[(45 06))288
. Elevation 77'-93' H = 192"; y = 48 psi 6 = .020" 2
- I(.4 10*)]l92 Elevation 61'-77' H = 192"; v = 56 psi 5
6 3 = [(,45x107)]l92
= .024" i
u Elevation 45'-61' H = 192"; v = 80 psi 6 = .034" 4 = [( .45 106 /1192 Total displacement Elev.117' E-W { = .094"s .09" The predicted displacements at the Control Building roof are 9" for the North-South direction and .09" for the East-West direction. D-3 O
I!
.O l 1
r 800 -
. PCA (86-4)
/
/
* ~~~ [ '* * %
6@
~
/ %s
.O=E / '~%s %.,
~
. /
{ / L 3 / G = .07x106
- o /
- s. 400 - / !
?u /
5 /
/
,/ Composite s l i l G = .8=10s G2 = .045x106 [
200 - C ,#**'-. ! l f \; = 45=105
- G = .02=106 *** ,
1 .- , I
.#. 8erkeley (Average 6 and 11) ;
lY./ G = .09x106 0 , , . . . . . . , . . , l 5000 10,000 Shear Strain u in/in 4 Figure 0-1 Shear Stress vs. Shear Strain Of The Test Specirnens Q , i i l
APPdNDIX E. n INELASTIC BEHAVIOR
..Q Section .7,0 of the report describes how the Control Building structure distributes maximum elastic loads when some of the members yield. Such yielding results in inelastic behavior and the seismic response forces based on elastic analysis' cannot develop. This.section presents a brief discussion of structural inelastic N
behavior. The ability of a structure to resist the dynamic loadings due to an earthquake depends not only on the structural capacity but also on the inelastic deforma-bility and energy absorption capacity of the structure. Seismic forces computed on the basis of a conventional linear elastic design response spectrum are the forces which would develop in the structure if the structure responds to the earthquake linear-elastically. If the yield capacity of the structure is less than that required to resist the earthquake elastically, then inelastic defor-mation occurs. When inelas' tic deformation occurs, a load-limiting mechanism is developed, since there is energy absorption associated with the inelastic defor- '. ma tion. Thus, the inelastic response of the structure is reduced as compared to the elastic response, the amount of the reduction being directly p rel 4ted to the amount of inelastic deformation and thus energy absorption that
. curs.
If ;he structure is idealized, based on its fundamental mode, by an elasto- ., (perfectly)-plastic single degree-of-freedom system in which the "yield" point corresponds to the structural capacity, then the inelastic seismic response of the structure to an earthquake can be expressed as a function of the "ductility factor" of the elasto-plastic system. The ductility factor designated as u is e defined to be the ratio of the maximum displacement (including the elastic and plastic components' to the displacement of the system at the yield point. E-1
/(
The allowable ductility factor of the system represents the maximum inelastic deformation permitted in the structure, which depends on the material,- the structural . type and the type of construction in the structure. The Uniform Building Code (UBC) relies on inelastic behavior to resist earthquake response forces. This is apparent when comparing the specified VBC earthquake forces with elastic response forces as determined from a response spectra, such as that used for the Trojan Power Plant. Based on the results of analytical studies on the seismic response of elasto-plastic systems, Newmark and Hall (Ref. l} developed the inelastic (elasto-- plastic) seismic response spectrum similar to the elastic response spectrum of a linear elastic system. The inelastic response spectrum values represent the seismic responses of an elastc-plastic system to an earthquake for a O specific demo 4as veiue end duct 4iity fector es e functioa of the (eiestic) frequency of the system. For seismic design purposes, the studies concluded that the ratios of the elasto-plastic to the elastic response spectrum values, expressed in terms of the ductility factor u, are as shwn in the following table (Ref. 21: RATIO OF ELASTO-PLASTIC TO FREQUENCY RANGE CORRESPONDING ELASTIC RESPONSE SPECTRUM VALUES o TO THE QUANTITY CONSERVED BETWEEN THE ELASTO-PLASTIC TOTAL RELATIVE ABSOLUTE AND ELASTIC SYSTEMS DISPLACEMENT ACCELERATION Displacenent f < 2.0 cps
- I I f
Energy and velocity " I 2.0 cps < f < 9. 0 c p>* v'2u - l G O Force and Acceleration u 1 f ut 9.0 cps
- j
- Frequency values are approximate.
E-2
L.Y The Trojan Control Building has a frequency of about 7.0 cps in the North-South direction based on elastic uncracked pro;erties. Af ter concrete cracking the frequency will be somewhat reduced, but still remain in the frequency range where energy and velocity are conserved. The required resistance will be (1/vE) of the elastic value and the resulting displacement will be (u//2a-1) times the elastic values. 4 The following table illustrates the resistance required and the displacement for various low ductility factors. DUCTILITY INELASTIC DISPLACEMENT INELASTIC CAPACITY FACTOR (u ) ELASTIC DISPLACEMENT ELAST IC. CAPACI TY 1.25 1.02 .82 1.50 1.06 .71 2.00 1.15 .58 3.00 1.34 45 5.00 1.67 .33 The Trojan Control Building is not a single oscillator, but it is dominated primarily by one mode of response. An indication of its behavior in the elastic range can be seen by making the extreme assumption that the structure had a yield a capacity (or begins to go inelastic) only one-half of that required to resist a
.25g earthquake in the linear elastic range. Based on the values given in the previous table, the required ductility is 2.5 and the displacement is only 25t greater than it would be to resist the earthquake in the linear-elastic range.
9 (v E-3
f ?' h REFERENCES (1). Newnark, N.M. , and Hall, W.J. , "Procedure and Criteria for. Earthquake Resistant Design", National Bureau of Standards, Report No. NBS BS5 46, Building Practices for Disaster Mitigation, Building Science Series,
- No. 46, Vol. 1, pp. 209-236 February 1973.
R [2] Newmark, N.M. , "Current Trends in the Seismic Analysis and Design of High Rise Structures", Chapter 16, Earthquake Engineering, edited by R. Wiegel, Prentice-Hall, NJ, pp. 403-424, 1970. O d J O E-4
1 de,.J
, lYh VY)
- o. ..
RESPONSE TO QUESTIONS FROM THE NUCLEAR RSGULATORY COMMISSION DATED AUGUST 30, 1978 l I September 20,1978 i i i L j
h o ~C Rosponno to NRC Questions QUESTION 1 "Provide a detailed comparison of the original and tne refined seismic analyses of the Turbine Building and a discuss.on of why the results of the new analysis do not alter any previous conclu-sions regarding the integrity of the structure. Identil'y any safety-related components, equipment and systems located in the Turbine Building, and discuss the impacts of the results of this refined seismic analysis on each of these. items and why the impacts of these results are acceptable."
RESPONSE
As stated in the FSAR, the Turbine Building is a seismic Category II structure. It was designed to withstand wind, tornado and UBC seismic loads. The design was evaluated for SSE seismic loads to ensure that it would not fail and af fect the adjacent Category I structures or internal seismic Category I areas. Three inch gaps were provided between the Turbine Building and the adjacent Category I structures to provide for relative building displacements during an SSE. As-built separation of the Turbine Building and Control Building has been checked. It was confirmed that the 3 inch gap exists in all locations except at two columns on the west side where a 1 inch plate is welded to the column flange leaving a 2-inch gap. This plate extends only to about 6 feet above the Turbine Building operating floor (elev. 93 ft.). Supplemental analyses have been performed to confirm the adequacy of the original design. The supplemental analyses considered the as-built masses and structural elements of the building. The original Turbine Building displacement analysis did not consider the as-built structural elements. The displacements in the original analysis were 2.9 inches at the Control Building roof elevation in both N-S and E-W directions based on the "sum of absolute values method".
- Subsequently, the building design was strengthened and stiffened for stress considerations. Since the displacements were within allowable limits before stif fening and the stif fened building would have lower displacements, it was not necessary to recalculate displacements. However, additional analyses were performed in August 1978 to predict Turbine Building displacements.
The predicted maximum E-W displacement of the Turbine Building at the Control Building roof elevation is 2.4" for 5% damping and 1.8" for 7% damping. The predicted maximum N-S displacement of the Turbine Building at the elevation of the control Building roo f ( el . 117 ' ) is 1.5" using 5% damping and 1.0" using 7% damping. The predicted extreme maximum displacement of the Control Building is 0.9" in the N-S direction as presented in Appendix D, "Trojan Control Building Supplemental Structural Evaluation". The predicted extreme displacement of the Control Building in the E-W direction. based on the same considerations as for the N-S direction, is 0.09". The combined maximum Turbine l 1-1 i
e 4 Response to NRC Questions Response to Question 1 cont'd, and Control Building E-W displacements based on the "sum of absolute values" method are 2.49" and 1.89" for 5% and 7% damping in the Turbine Building, respectively. Table 1.1 summarizes these results. Therefore, the 3" gap is adequate. The 2" gap in the E-W direction at the location of the plate (El. 99') notad above is also adequate, since sum of the absolute values method combined displacements at that location are 0.97" and 0.87" for 5% and 7% damping in the Turbine Building, respectively. The sum of the absolute values method combines the displacements in a very conservative manner since it considers that the two structures will attain their maximum displacements simultaneously, as well as in opposite directions. This is very unlikely to occur. The more realistic values are the SRSS combined values, which are also tabulated in Table 1.1. The following safety-related equipment is located within , the Turbine Building :
- 1) Emergency diesel-generator units
- 2) Engineered Safety Features (ESP) Switchgear and Remote Shutdown Panel
- 3) Auxiliary feedwater pumps
- 4) Cable trays for ESF system The emergency diesel-generator units, the ESF switchgear , the remote shutdown panel, and the auxiliary feedwater . pumps are housed in structures which are separated from the Turbine Building structural elements by a 3 inch gap. Therefore, the seismic response of the Turbine Building does not af fect these structures nor the equipment
, located within them. Category I cable trays for the ESF system t are routed from the Diesel-Generator Building to the Control Building and are supported by the Turbine Building f raming system. These supports have been reviewed and are adequate i to withstand SSE seismic loads, as well as the relative displace-ments between the Turbine Building and Control Building during an l SSE. \ 1-2 l
e % TABLE 1.1 Itaximum Displacement (inch) SSE 0.25g N-S Direction E-W Direction El. 117' El. 117' ? El. 99' Turbine Building (T3) 1.5 2.4 0.9 0 5% damping TurbineBuilding(TB) 1.0 1.8 0.8 9 7% damping Control Building (CB) 0.9 0.09 0.07 0 5% ABS Combination 2.4 2.49 0.97 CB 0 5% + TB 0 5% ABS Combination 1.9 1.89 0.87 CB 0 5% + TB 0 7% SRSS Combination 1.75 2.4 0.9 CB 0 5% + TB 0 5% SRSS Combination 1.35 1.8 0.8 CB 0 5% + TB 0 7% , Gap 3 3 2 1-3
. 's Rosponse to NRC Questions QUESTION 2 "Discuss in detail the discrepancies between the original Bechtel analyses, the TABS analysis an2 the STARDYNE analysis of the Control-Auxiliary-Fuel Building complex. Provide a comparison of the results and the basis for any conclusions regarding the acceptability of these inconsistencies. Substantiate (1) that there is no detrimental impact en pt evious statements and conclusions regarding the integrity of the Fuel Building, and (2) that the statements in the FST,R regarding the conservatism of the fixed-base seismic analyses performed previously for all Seismic Category I structures are not invalidated."
RESPONSE
GENERAL: The differences in the various analyses that were performed for the Trojan Control Building are discussed in detail in Appendix A of the "Trojan Control Building Supplemental Structural Evaluation" dated September 1978. Appendix A also contains a comparison of the results and discussion of the acceptability of the inconsistencies. RESPONSE TO ITEM 1: Both the stick model used in the original seismic analysis and the finite element model used in the recent supplemental STARDYNE analyses included the Fuel Building. In the original analysis the hold up tank enclosure structure, the fuel pool and the connecting slabs were represented by beam elements and concen-trated masses. In the supplemental finite element analysis, the walls of the hold up tank enclosure and the fuel pool structure and the connecting floor slabs, all conventional reinforced concrete structures, were represented by plate elements. Distribution of lateral forces among the individual walls within the structure and torsional effects (membrane shear in N-S walls due to E-W earthquakes and in the E-W walls due to N-S earthquakes) are more precisely represented by the results of the supplemental analysis. Table 2.1 shows the membrane shear forces based on the original and supplemental analysec as well as the corresponding shear capacities for each wall of the fuel pool and hold up tank enclosure structures. The forces are the square roots of the sum of the squares of the corresponding responses in the significant modes. Generally, membrane shear forces decrease in the supplemental analysis. Also, the supplemental analysis shows decreased shear forces in all interconnecting slabs. In a few walls, the supplemental analysis shows somewhat increased membrane shear forces. The dif ferences between the original and supplemental analyses can be attributed to the capability of the finite element analysis to define responses more precisely and to the contribution of some walls not considered in the original analyses. 2-1
< i , e R'ispanno to NRC Questione Response-2, continued The tabulated shear capacities of the walls of the Spent Fuel Pool and Hold-Up Tank Enclosure structures are computed in accordance with the following:
V u I (Y c + Y s ); where V e = l2(f )1/ hd and V, = A y y d/s. f Notation: V u = factored shear force at section. Vc = n minal shear strength provided by concrete V,= nominal shear strength from reinforcement
/ = 0.85, strength reduction factor for shear h = overall thicknes of member , in.
d= 6.8 1,, in. 1, = horizontal length of wall, in. Ay = area of shear reinforcement within a distance of sj sq.in. ! s = spacing of horizontal reinforcement in wall, in, f = 3000 psi, specified compressive strength of I concrete, psi. y = 40,000 pai, specified yield strength of f nonprestressed reinforcement, psi. The above expressions are taken from the ACI Standard 318-77; the streng th of materials f rom the PSAR. The reinforced concrete walls and slabs in the Fuel Building have substantial extra capacicy to resist seismic forces. The supplemental analysis fully Confirmed the conclusion of the original analysis that the structural elements of the Fuel Building can resist SSE and the factored OBE loads within the PSAR criteria. 2-2
, o -
Response to NRC Questions Response,2, continued RESPONSE TO ITEM 2: FSAR Section 3.7.2.3 compares results of analyses for flexible and fixed base models of the containment structure in order to substantiate that fixed base seismic analyses are conservative. As documented in Section 3 of the "Trojan Control Building Supplemental Structural Evalua tion," the base shear in the N-S direction for the flexible base case is 12,210 kips and the base shear for the fixed base case is 11,910 kips. The dif ference in these values is only 34 which is negligible. Since the flexible base case neglected radiation damping, the comparison of the results of the two STARDYNE analyses does not invalidate the statements in the PSAR regarding the conservatism of the fixed base seismic analyses performed previously for all Seismic Category I structures. 2-3
o O TABLE 2.1 LOAD VS. CAPACITY OF EACH WALL OF THE HOLD-UP TANK ENCLOSURE AND THE SPENT FUEL POOL (KIPS) s 1.4 OBE (,15 9)
- SHEAR a
g CAPACITY 3 rig. Stick Model Supplemental
- N-S E-W N-S E-W 26 7363 1320 2130 349 27 7363 1320 1764 878 g 35 2718 '
1360 956 1890
- a. m
?S 36 1616 1360 193 729 918
@" 37 1616 1360 60 671 38 2718 1360 140 596 24 11124 4500 1703 823 g 25 13792 4500 1465 515 wa Ny 33 14352
\ 4340 1304 2034 34 16561 2980 555 1346 s
- STARDYNE finite element analysis 27 35 36 37 38 25 26 Hold-Up Tank Enclosure SPENT m u 33 FUEL 34 POOL 24 Key to wall numbers
/
2-4
. mr
Rosponso to NRC Questions QUESTION 3(a)
"Given the results of the supplemental STARDYNE analysis and considering all of tne nonlinearities in the response of the structure ' indicated by these results:
(a) Identify each of the expected nonlinaarities and quantify the ef fects of each on the behavior of the structure."
RESPONSE
Nonlinearities are identified and their effects are addressed in Sections 5, 7, and Appendix E of the repor t "Trojan Control Building Supplemental Structural Evaluation", September 1978, and in the response to Question 6 (infra) . Section 5 demonstrates that the capacities of the major shear walls are suf ficiently large such that major nonlinear behavior is not expected. Section 7 illustrates how the redistribution of forces can be accommodated by the structure. Appendix E discusses the structural response of systems which have significant nonlinear type behavior. l l l I l l l l l i 3a-1 ] f i l
Response to NRC Questions QUESTION 3(b)
"Given the results of the supplemental STARDYNE analysis and considering all of the nonlinearities in the response of the structure indicated by these results:
(b) Provide a detailed discussion of the impact of these results on the floor response spectra for the Control Building.
-Identify all safety related components, equipment and systems in the Control Building. Discuss the impacts that any changes in the response spectra and the increased displacements of the structure would have on each of these items and why these impacts are acceptable."
RESPONSE
Response Spectra An evaluation was made of expccted effects on Control Building floor response spectra of the STARDYNE elastic analysis results (0.25 g SSE fixed-base model, 5% damping ratio) and the possible e.f fects on the spectra of nonlinear seismic response of the Con-trol Building that could occur due to localized inelastic struc-tural behavior. The horizontal floor response accelerations resulting from the STARDYNE analysis are less than those obtained from the 1971 time history analysis which generated the Control Building's initial floor response spectra. The lower values predicted by the STAR-DYNE analysis are primarily due to the very conservative syn-thetic time history used in 1971. As shown in the PSAR, Figures 3.7-6a, -6b, and -6c, the response spectrum of the synthetic time history exceeds the corresponding design ground response spectrum by large amounts in most frequency ranges. This is particularly true in the 6 to 8 Hz frequency range of the Control Building funda-mental horizontal frequencies. A time history analysis using the STARDYNE model and an input time history whleh closely matches the design ground spectrum would produce floor response spectra having floor response accelerations very close to those obtained from the STARDYNE (spectral response) analysis. Since floor response spectra are generated on the basis of linear elastic analysis, the floor response spectra which could be ex-pected from a STARDYNE time history analysis can be obtained by linearly scaling the original floor response spectra amplitude in accordance with the ratios of the horizontal floor accelerations, and by shif ting, while maintaining the original i 10 % broadening criteria, the spectral peaks to correspond to the fundamental 3b-1
Ronponse to NRC Questions Response to 3b continued f requencies calculated by the STARDYNE analycic (6.8 Hz, N-S; 8.5 Hz, E-W). The floor spectra calculated in this manner are conservative because the spectral peak amplification factors resulting from the original analysis are preserved. To evaluate the effect of inelastic behavior on the Control Building floor response spectra, a simplified technique based on a pseudo-elastic system and inelastic response spectrum was util-ized. The stif fness of the pseudo-elastic system was taken to be the elastic stiffness divided by the ductility ratio. The accel-eration response of the pseudo-elastic system was calculated based on the inelastic response spectrum. (Ref: "Trojan Control Building Supplemental Structural Evaluation", Appendix E, Reference 1). A ductility f actor of 1.5 was selected to represent a conservative bound on the Control Building inelastic response. Using the ductility factor 1.5, the floor spectral peak freque to the lower frequency side by a factor of 1/(/A)g,cies shif 82 (i.e., t %), and the amplitude of the floor spectral peaks would decrease f rom the STARDYNE elastic spectral peak amplitudes by a ragio response accelerations which is 1/( Su-1) of horizontal floor (i.e., 71 %). The resulting floor response spectra as constructed have spectral peak amplitudes lower than the amplitudes of both the original and the STARDYNE floor spectra. Also, their spectral peak frequencies lie in between the original floor spectral peak f requencies and the STARDYNE f req-uencies which are higher than the 1971 frequencies. Figures 3b-1 through 3b-8 are 0.5 % damped r esponse spectra at each principal elevation of the Control Building for N-S and E-W d ir ec tions . These figures show the original, the STARDYNE elas-tic, and the pseudo-elastic (ductility ratio /4 = 1.5) curves for comparison. As can be seen in the figures, tae STARDYNE elastic curveu have lower peak amplitude and are located on the high frequency side of the original spectra. The pseudo-elastic spectra do not have an appreciable ef fect relative to the original spectra. A ductility ratio greater than 2.0 would be required to shift the pseudo-elastic spectra beyond the spectral peak frequency range l of the original 1971 spectra. It is therefore concluded that because of the energy absorption associated with the inelastic de-formations, the ef fect of the Control Building's inelastic behav-ior on the original floor response spectra would be insignificant. Seismic qualification documents for safety-related components, equipment, and systems located in the Control Building were re-viewed to determine if possible changes to the floor response spectra, as described above, could produce significant rpectra t 3b-2 l l l
,. a R7rponsa to NRC Qu9stiona Response 3(b), continued amplitude or frequency shif ts beyond the range of the original seismie qualification envelopes. s Tables 3b-1 and 3b-2 summarize the mechanical and electrical safety-related components, equipment, and systems located in the Control Building, their floor elevation locations and the method used for seismic qualification. Various specific methods by which the equipment was seismically qualified are:
- 1. Generic qualification to acceleration levels which envelope the required qualification floor response spectra in the resonant frequency range of the equipment.
- 2. Qualification to required floor response spectra by vibration testing of the actual equipment.
- 3. Analysis to determine equipment natural frequencies, demon-stration that the equipment and supports are effectively rigid, and qualification by stress analysis or testing to show that the equipment can withstand the required acceleration levels.
- 4. Qualification by analysis using the peak acceleration of the floor response spectra.
- 5. Analysis to find resonant f requencies, determination of spec-tral accelerations from the floor response spectra, and stress analysis to demonstrate seismic qualification.
- 6. Qualification by testing to find resonant frequencies and testing at these f requencies to acceleration levels greater than those from the response spectra.
- 7. Qualification by stress analysis by using either the peak acceleration of the floor response spectra or the acceleration at the system frequency.
For equipment qualified by testing (who se reports include test response spectra data), direct visual comparisons can be made of the spectra amplitude and frequency relationships between the test response spectra and the envelope of the original floor response spectra, the equivalent STARDYNE elastic spectra, and conservative pseudo-elastic spectra. Examples of these spectra compar isons are shown in Figures 3b-9 through 3b-13. 3b-3
R7sponso to NRC Questions Response 3(b), continued The detailed review of seismic qualification documentation for Control Building safety-related components, equipment, and systems has shown that the STARDYNE results and possible nonlinear Control Building response characteristics will have no significant ef fect on the previously qualified safety-relcted equipment. A reanalysis of the Control Building safety related cable trays and their supports has shown that they can withstand the shif t in responses predicted both by the STARDYNE analysis and the possible nonlinear response of the Control Building. Service Water System piping to equipment room coolers is the only safety related piping in the Control Building. Seismic reanalysis of this piping identified one 3" section of Service Water System Train " A" whose design adequacy could be affected by the shift in response spectra suggested by the STARDYNE analyses. The addition of two small seismic restraints will provide sufficient rigidity to resolve this. With this exception, it is concluded that original seismic qualification of the safety-related components, equipment, and systems in the Control Building are not af fected by the revised floor response spectra corresponding to the STARDYNE elastic analysis or the pseudo-elastic analysis. Increased Displacements A survey was conducted to determine the capability of safety-related equipment, piping, and electrical cabling to withstand large displacements in the Conrol Building, Auxiliary Building, and Fuel Building. As shown in Appendix D of the "Trojan Control Building Supplemental Stru:tural Evaluation," September 1978, displacements in these buildings during the design-basis earthquakes will be very small, about 0.5" maximum, lateral displacement between the floors, and less than 0.9 in, total displacement at the top of the Control Building. By inspection, the surveyed equipment, piping, and electrical runs will easily accept such small displace-ments. For additional conservatism, the survey investigated the sensitivity of these safety-related components to much larger ( e.g . g rea ter than 1") displacements to confirm that, even if larger displacements were to take place, the function of such components will not be affected. The effect of differential displacement between adjacent buildings (Control vs Turbine, and Auxiliary / Fuel vs Contain-ment) was also evaluated. It should be emphasized that these larger displacements are not expected to occur, and that they could only occur under conditions that are entirely unfereseen at this time. However, to be conservative in the evaluation of the Trojan Control Building capability to withstand the safe shutdown earthquake, it was considered reasonable to investigate such unexpected conditions. 3b-4
i L Response to NRC Questions [ Response to 3(b), continued f Por completeness, the survey also investigated the vulnerability to pieces of falling debris of equipment located near the walls. Again, this is not expected, but was investigated to provide a conservative point of reference. The survey concluded that the safety-related equipment (mechani- ! cal, electrical, and piping) located in the Control Building, l Auxiliary Building, and Fuel Building can withstand large i i displacements (greater than 1 in. between floors, and 3 to 6 in. between adjacent buildings) without loss of f unction. This con-clusion is based on the physical location of equipment, piping, and electrical runs, their associated connections and routing, and . inherent strength of materials used. In addition, it was deter- ! mined that an assumed piece (a 10-lb block) falling from the wall , would no t cause equipment , piping, or electrical components to fail to perform their intended functions. ; l 3b-5
. .- TABLE 3b-1 CONTDOL BUILDING
, MECHANICAL EQUIPMENT SEISMIC QUALIFICATION Floor Elevation Qualification Equipment Number Equipment Description Location Method
- CB-1 Control Building Emergency Ventilation:
VC-142A&B Fans 105' 3 ) 4 5 VF-147A&B Filters 105' 3 VF-148A&B VF-149A&B VF-150A&B VE-159A&B Duct Heaters- 105' 6 VE-160A&B Cooling Coil 105' 3 VE-161A&B Duct Heater 105' 6 C254 (C255 Control Panel & Identical) Instruments 105' 2 C259 Control Panel VE-161A&B 105' 6 C260 Control Panel VE-159A&B 105' 6 Duct and Duct Supports 93' & 105' 3 CB-7 Battery Room: VC-146A&B CB-7 Battery Room Fans 61' 3 VW-187 & 188 Unit Heaters 61' 6 CB-8A&B Elec Aux Room / Cable Spreading Room V-145A,B,C,0 Self-Contained Air 61' 5 (E-SCII) Condition Units Elec-tric Auxiliary Roor. V-143A,B,C,D, Self-Contained Air 77' S 3,F Condition Units Cable Spreading Room 3b-6
TABLE 3b-1 l (Continued) CONTROL BUILDING MECHANICAL EQUIPMENT SEISMIC QUALIFICATION Floor ; Elevation Qualification ' Equipment !1 umber- Equipment Description Location Method ' Service Water Piping, Valves and 61' & 10$' 3 , System Supports in Control Room
- CB-2 & CB-12 Control Room Normal Venti- e lation and Special Exhaust:
DM-10501, Damper Actuators 93' & 105' 3 DM-10502 ,- DM-10503 ' DM-10504
- Qualification Methods See Table 2 Notes 3b-7
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TABLE 3b-2 CONTROL BUILDING ELECTRICAL EQUIPMENT SEISMIC QUALIFICATION Floor Number Elevation Qualification Equipment Equipment Description Location Method
- C17/C27 Duplex Bench Control 93' 4 Board C31A,B,C,D Nuclear Instrumentation 93' 1 C44B,C45A Auxiliary safeguard 93' 1 Cabinet C19/C29 Duplex Bench Control 93' 4 Board C36Al,2,3,4,5 Process I&C Rack 93' 1 Protection, Set 1 C36A6,7,8,9 Process IEC Rack Control 93' 1 Group 1 C36B1,2,3,4,5 Process I&C Rack 93' 1 l
Protection Set 2 C3686,7,8,9 Process IEC Rack Control 93' 1 j Group 2 l C36C1,2,3 Process I&C Rack 93' 1 Protection Set 3 C36C4,5,6 Process I&C Rack Control 93' 1 Group 3 C48A SSPS Train A Logic Rack 93' 1 C49A SSPS Train A Output Rack 93' 1 C47A,B,C,D SSPS Train B Input Rack 93' 1 C48B SSPS Train B Logic Rack 93' 1 C49B SSPS Train B Output Rack 93' 1 C121/122 Rx Trip Switchgear 65' 1 C40A,C40B Engineered Safeguards 93' 1 Test Cabinets 3b-21
4
. TABLE 3b-2 (Continued)
CONTROL BUILDING ELECTRICAL EQUIPMENT SEISMIC QUALIFICATION Floor Number Elevation Qualification Equipment Equipment Description Location Method
- C41 Radiation Monitoring Panels 93' 1 C07A/C07B Hydrogen Recombiner Controls 93' 1 134-1/134-2/ Shutdown & DBA Sequencers 93' 3 134-3/134-4
> C11/21 Duplex Vertical Control 93' 4 Board C12/C22 Duplex Bench Control 93' 3 Board C13/C23 Duplex Bench Control 93' 3 Board C18/C28 Duplex Bench Control 93' 3 Board C14/C24 Duplex Bench Control 93' 4 Board C15/C25 Duplex Bench Control 93' 4
. Bcard C16/C26 Duplex Bench Control 93' 4 Board C36D1,2,3 Process I&C Rack 93' 1 Protection Set 4 C3CD4,5,6, Process I&C Rack Control 93' 1 7,8 Gtoup 4 C46A,B,C,D Solid-State Protection 93' 1 System Train A Input Rack A-2 4.16 kV Switchgear 65' 6 B02/B04 480V Load Center 65' 3 B22/B24/B26 480V Motor Control Center 65' 1 3b-22
[ l L- =.
TABLE 3b-2 (Continued) CONTROL BUILDING ELECTRICAL EQUIPMENT SSEISMIC QUALIFICATION Floor Number Elevation Qualification Equipment Equipment Description Location Method
- Y-ll Vital AC Panel 65' 1 Y-13 Vital AC Panel 65' 1 Y-22 Vital AC Panel 65' 1 Y-24 Vital AC Panel 65' 1 Y-15 Static Inverter il 65' 1 Y-26 Static Inverter #2 65' 1 Y-17 Static Inverter #3 65' A Y-28 Static Inverter 64 65' 1 Q35,36,37,38 Solotron Voltage Regulators 65' 1 011/D12 125v de Station 65' 5 Battery and Racks D21/D22/ Battery Charger 125V de 65' 3 D23/D24 D10./D20/ 125V de Control Power Panels 65' 1 D30/D40 CO2 Main Control Console 93' 3 C04 Main Control Console 93' 3 C05 Main Control Console 93' 3 Class IE -------
Various 7 Cable Tray 3b-23 l l L.
TABLE 3b-2 (Continued) CONTROL BUILDING ELECTRICAL EQUIPMENT SEISMIC QUALIFICATION i l Floor l Number Elevation Qualifiction ; Equipment Equipment Description Location _ Method l NOTE: The items listed below are installed as part of the various panels documented above. Equipment numbers are not applicable since much of this equipment was tested and supplied directly to the panel manufacturer. Indicating Lights Various 6 Control Switches 1 i G.E. Type SBM Various i Westinghouse various , Type W-2 & OT-2 1 ! G.E. Relays Various 6 Type 12HFA51 ' 12HEA62 12HMA25 j 12IAC66K 1 12IAC51A 12HMAll Fischer-Porter Various 6 Type 50AS3000 50ER3000 55EL3000 55ES3300 55FG3000 55GL 6 53G3000 53D3000 l 50EK1000 6 ET215 6 53EL3000 53EG3000 6 Westinghouse Bistables 93' 6 l in SSPS Magnetics Status Lamps 93' 2 3b-24 l
TABLE 3b-2 (Continued) CONTROL BUILDING ELECTRICAL EQUIPMENT SEISMIC QUALIFICATION gu,alification Method:
- 1. Qualified by generic testing.
- 2. Qualified by testing to specific floor response spectra.
- 3. Qualified by analysis showing system is rigid.
- 4. Qualified by analysis using peak acceleration of floor ;
response spectra. '
- 5. Qualified by analysis using actual resonant frequencies and corresponding spectral accelerations.
- 6. Qualification by testing to find resonant frequencies and test motion at identified resonant frequencies. l
- 7. Qualification by a combination of Methods 4 and 5.
l J 1
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l I 3b-25
Posp*nse to NRC Quqstions QUESTION 3(c) _
"Given the results of the supplemental STARDYNE analysis and considering all of the nonlinearities in the response of the structure indicated by these results:
(c) Provide a summary of the stress levels in the most critical wall elements (i.e. consider thore walls which are the most highly stressed, those walls with the highest and lowest he ig ht to length ratios, those wa.11s with the highest and lowest reinforcement ratios, etc.), and the corresponding strength ratios and acceptance criteria. Discuss the impact of cyclic degradation on wall strengths at these stress levels and provide a basis for all conclusions regarding these effects. In addition, discuss the strain compatibility between the concrete core and the masonry at these stress levels to substantiate that each of these materials will reach the assumed stress levels at about the same overall wall displacement such that significant degradation of either one of these materials does not occur before the other has reached its required stress level."
^ ' '
RESPONSE
The most critical wall elements are the walls in the N-S direction at both elevations 45' and 61'. For completeness, the stress levels at all of the major walls in the N-S direction at these two elevations were evaluated. Tables 3(c)-1 and 3(c)-2 provide the proper ties (length, height-to-length ratios, reinforcement percentages) of these major walls as well as their shear stress capacities (based on the criteria presented in "Trojan Control Building Supplemental Stt uctural Evaluation" of September 1978) . The piers are shown in Figures 3(c)-1 and 3(c)-2. The wall designa-tions referred to in Tables 3(c)-1 and3(c)-2 as well as rigures 3(c)-1 and 3 t':'-2 are consistent with Figure 3-2 and 3-3 of the structural evaluation. The shear stresses for the Cixed base 0.25g SSE with 5% damping are also tabulated. Strength ratios (capacity / load) are tabulated in Section 5 of the "Trojan Control Building Supplemental Structural Evaluation". The basis of acceptance is demonstration of grose capacity greater than gross load demand since loads redistribute within the structure as described in Section 7. l Appendix B of the "Trojan Control Building Supplemental Structural Evaluation", dated September 1978, discusses how cyclic degradation is considered in the criteria, while Appendix D evaluates the upper limit displacements, Appendix C of the same document I discusses strain compatibility of the wall systems. i 3c-1
Table 3(c)-1 Shear Stress Load And Capacity. Elevation 45'-61', N-S Direction REINFORCEENT CAPACITY LOAD W H/W T BLOCK CORE SHEAR SHEAR STRESS STRESS
- WALL (FT) (IN) ph(%) py(%) p h( } 'v( } (' } ( }
1 31. 0.26 23 .16 .11 . 38 .08 283 21 5
- 24. 0,33 41 .13 .09 .42 . 09 252 210 4
2 13. 1.08 48 .15 .08 .22 .04 63 68 3 15. 0.93 39 .13 .09 .17 .05 70 72 4 12.08 0.66 26 .20 .14 . 50 .07 297 21 2 8.0 1.0 27 .19 .14 .48 .07 260 142 9.5 0.84 27 .19 .14 . .48 .07 270 142 4.67 1.70 34 .15 .11 .38 .05 211 141 9.17 0.87 27 .19 .14 .48 .07 260 183 4 5 23. 0.61 20 .18 .13 .43 .09 3 54 158
- 23. 0.35 34 .16 .11 .76 .05 231 158 i
- 32. 0.25 20 .18 .13 .65 .09' 240 117 i
where: Note: All reinforement ratios are evaluated from the uniformly W = length of wall distributed reinforements over the gross section. 8 T = gross wall thickness ph = horizor.tal reinforcement ratio oy = ver tical reinforcement ratio 3c-2 i
l . . I Table 3(c)-2 Shear Stress Load And Capacity. Elevation 61'-77', N-S Direction REINFORCEMENT CAPACITY LOAD W H/W T BLOCK CORE SHEAR SHEAR S STRESS
- WALL (FT) (IN) p h(%)
py(%) ph( } 'v( } ( I} 1 31 0.45 41 .13 .09 -- ** - - *
- 1 50 136
.15 .10 -- ** - - *
- 103
- 31. 0.45 35 1 50 0.32 .23 .34 -- ** - - *
- 150 11 6
- 31. 16 2 13 1.08 48 .15 .08 --** - - *
- 25 35 3 12. 0.67 26 .14 .10 .31 .07 248 1 61
- 26. 0.31 32 .16 .12 .30 .06 271 183 9.5 0.84 28 .18 .13 .26 .06 231 1 83
** ** 23 13 0.62 16 .23 .'34 --- --- 20
** ** 23 12 0.67 16 .23 .34 --- --- 20 where: Note: All reinforcement ratios are evaluated from the W = length of wall uniformly distributed reinforcements over the H = height of wall gross section.
T = gross wall thickness
- Loads from Case 4. Section 7.
= horizontal reinforcement ratio **No core reinforcing steel ph is assumed, py = vertical reinforcement ratio 3c-3 l
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5, Wall 5 1' Figure 3(c)-1 Piers Of Pajor Walls (N-S, Elevation 45'-61') 3c-4
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-, I f i i r 8' I El. 65' g, 8' l
* ~ l '
E l . 61 ' Q 0' 1G.0I 12.0'.1 5.0'm 26.0' -. t_ 9 3 ' ,, 9E't 12. 0 ' .- Wall 3 5 Figure 3(c)-2 Piers Of Major Walls (N-S Elevation 61'-75') j ' . l ', 3c-5
e . R9sp7nse to NRC Qu7stions QUESTION 3(d)
"Given the results of the supplemental STARDYNE analysis and considering all of the nonlinearities in the response of the structure indicated by these results:
(d) Discuss in detail the ef fects of the "Basic Shear Wall Capacity Criteria" utilized in the supplemental analysis on the shear transfer capability at the wall-slab inter-faces."
Response
The shear forces developed in the slab are conservatively considered to be transferred to the supporting shear walls by shear friction. Based on shear-friction the capacity is > given by: R = /,M A s fw y Where: R = shear capacity (KIPS) A s = area of slab horizontal reinforcing steel embedded in the wall (square inches / foot) fy = yield tensile strength of reinforcement (45 ksi)
/ = capacity reduction factor in shear (0.85)
,d= fr iction f actor , conservatively assumed as 1.0 W = slab leng th ( f t. )
The capacity of the slabs at the slab-wall interface at elevations 61', 65', 77', 93' and 117' in the Control Building and the western half of the Auxiliary Building were evaluated in accordance with the above equation. The capacities were compared with the shear forces developed in the "Case 4" analysis reported in Section 7.0 of the "Trojan Control Building Supplemental Structural Evaluation" dated September 1978. All slabs had suf ficient capacity to transfer the loads into the shear walls. The following addresses the shear transfer through the walls from higher to lower elevations. I The concrete in the core of the walls has a design strength of 5000 psi (6000 pai, 90 day strength) and the slabs are constructed , from concrete with a design strength of either 3000 psi (4000 psi, 90 day strength) or 5000 psi,(6000 psi, 90 day strength). These 90 day ; strength values are based on actual field cylinder test records. . The tests upon which the "Basic Shear Wall Capacity Criteria", ' documented in Section 4 and Appendix B of the "Trojan Control Building Supplemental Structural Evaluation" dated September i 1978, are based on specimens made f rom concrete having strength in the range of 3000 to 4000 psi. This shear wall criteria does l not reflect the higher strength of the Trojan walls, therefore, the capacities of the walls predicted by this criteria are compatible with the strength of the slabs.
, 3d-l ;
l t
7 e O Response to NRC Questions Question 3(d) Cont'd In transferring the load f rom the upper wall-slab connection to the lower wall, the following load transfer mechanisms are present:
$. . The dowel capacity of the reinforcing steel in the concrete block is based on 0.9 A,f uwhere; A, is the area of the vertical steel, and '
u is the ultimate tensile strength of the f reinforcing steel
- 2. The dowel capacity of the encased columns.
- 3. At Elevations 61' and 65', the encased beams have Nelson studs on the top and bottom flanges. At other elevations, the beams have Nelson studs on the top flange only and load is transfered from the bottom flange into the lower wall by bond.
- 4. Compression across this interface coming from dead load and bending will generate additional load transfer.
At all Elevations, there is adequate load transfer by using the above capacities where applicable. l l 3d-2 i
r-e O Response to NRC Questions QUESTION 3(e)
"Given the results of the supplemental STARDYNE analysis and
< considering all of the nonlinearities in the response of the structurr indicated by these results: (e) Provide a detailed discussion of shear transfer into trans-verse walls and the tensile and compression stresses in these transverse walls to substantiate the assumption that the gross overturning moments in the structure will be resisted by the transverse walls. Include a basis for any conclusions regarding the adequacy of these walls to transmit the required forces."
RESPONSE
The subject of moment and vertical shear transfer is discussed in Sections 6 & 7 of the "Trojan Control Building Supplemental Structural Evaluation", dated September 1978. 3e-1
, Response to NRC Ouestions QUESTION 3(f) *
"Given the results of the supplemental STARDYNE analysis and considering all of the r.onlinearities in the response of the structure indicated bi these results:
(f) Discuss the adequacy of the shear transfer mechanism at the foundation level from the walls into the rock. Include the bases for any conclusions regarding the adequacy of ,
> these mechanisms."
RESPONSE
Analysis has been conducted to evaluate the lateral force transfer mechanism from the Control Building and the Fuel Building structures to the rock foundation based on the loads obtained from the STARDYNE elastic analysis for 0.25 g at 5% damping. As discussed in Section 8 of the "Trojan Control Building Supplemental Structural Evaluation", dated September 1978, the analysis has reconfirmed that the Control Building can adequately transfer the seismic shear forces to the tock foundation. In the Fuel Building, the spent fuel pool and holdup tank enclosures have reinforced concrete mat foundations. The mat foundations were constructed first, followed by the reinforced concrete walls. Fuel Building lateral forces are resisted by the hold-up tank enclosure and the spent fuel pool and transferred to the rock foundation. A conservative coef ficient of friction between the rock and the mat foundations equal to 0.7 was used in calculating the resistance to sliding. Results for the holdup tank enclosure and spent fuel pc01 are given in Table No. 3(f)-1. Since seismic forces for the CBE (0.159) at 24 damping are very close to the SSE (0.259) at 5% damping, values for the CBE are not shown. It has been reconfirmed that the holdup tanks enclosure and fuel pool foundations can safely transfer the seismic shear forces to the rock foundation. 3f-1 i
F TABLE NO. 3 (f) -1 FUEL BUILDING HOLD-UP TANK ENCLOSURE AND SPENT FUEL POOL SAFE SHUTDOWN EARTHOUAKE (0.25 g 9 5 % Damping) DIRECTION OF EARTHQUAKE N-S E-W , i. HORI2ONTAL BASE FACTOR BASE FACTOR STRUCTURE RESISTANCE SHEAR OF SHEAR OF 0.7W SAFETY SAFETY (kip) (Kip) (Kip) HOLD-UP TANK 5950 2612 2.28 2721 2.19 ENCLOSURE SPENT FUEL POOL 10150 2184 4.65 3031 3.35 (WIT!! WATER) SPENT FUEL POOL 8260 2184* 3.78 3031* 2.72 (WITHOUT WATER )
- seismic analysis was not performed for empty pools conservatively the base shear forces corresponding to pool filled with water was used.
1 3f-2 i i i
Responso to NRC Questions QUESTION 5 "Discuss in detail the impacts of the supplemental STARDYNE analysis results on all analysis results submitted previously. Quanttfy the differences between the supplemental and previous l basic shear wall acceptance criteria for the various extremes ( of shear wall behavior (i.e., consider the walls with the highest and i lowest reinforcement ratios for walls in the various height to length ' ratio categories, the walls which were both moment controlled , and shear controlled under the previous criteria, etc.). Include l a basis for using a wall's dowel capacity, deterained by con-sidering that the steel reiaforcement will reach its ultimate : strength, in the supplew atal analysis as the capacity of a wall ; if the dowel capacity is greater than the shear capacity calculated from the "Basic Shear Wall Acceptance Criteria" for walls with height to length ratios less than or equal to 0.5." RESPONSE ; i F I As discussed in Appendix A of the "Trojan Control Building Supplemental Structural Evaluation" dated September 1978, four seismic anelyses j
; have been performed on the Trojan Control Building. These were:
- 1) The original analysis using a stick model (1971), and a
; re-evaluation of the original using SRSS and estimated i as-built floor loads ( April 1978)
[ ] 2) An analysis using the TABS program [ l 3) An analysis using the STARDYNE program with flexible- [ based model
- 4) An analysis using one STARDYNE program with fixed-based model ;
I l While dif ferences between the analyses exist in both the loads predicted and the wall capacities considered, the STARDYNE analyses does not invalidate the conclusions of the results previously submitted. As . eflected in "Trojan Control Euilding Supplemental ! Structural Evaluation", dated September 1978, Appendix A, Table A-1, , the loads predicted by STARDYNE are higher than past analyses ' have shown. On the other hand, higher capacities are appropriate [ for the walls when evaluating their capability to resist STARDYNE predicted loads. The reason that higher capacities are appropriate in the case of the STARDYNE analyses lies in the mathematical sophistication of this analytical technique. 5-1 l \
r ce . .p . Response to NRC Questions Question 5, continued The more conventional analytica. v Mniquen do not have the capability to predict the respon f individual elements in a atructure. Recognizing th chey are employed with conservatisms inherent in tae A(A and UBC codes. With the STARDYNE analysis, these limitatione are eliminated and it is possiole to predict more accurately the loads seen by elements c2 a stru:ture. Accordingly, the need for the code conserv-atisms no longer exists, and it would be inappropriate to ap-ply these conservatisms in predicting capacities. Thus, the use of more realistic capacities is justified. The Control Bui5 ding is capable of resisting the 0.25 g 3SE ear thquake with considerable margin. This is especially so because the damping value 'a a lower limit and the forces based cn elastic analysis are upper limits. Diffe ences between the present and the previous basic shear wall criteria are documented in Section 4 and Appendix B of the "Trojan Control Building Supplemental Structural Evalua-tion", dated September 1978. These shear criteria have been refined from the preliminary criteria presented in the Sep-tember 1, 1978 submittal. The September 1, 1978 submittal used the Schneider criteria but linearly reduced the shear capacity for walls with reinforcing ratios less than 0.0025. In addition, when the Schneider criteria resulted in lesser capacities for walls with an H/W ratio less than 0.5, the dowel strength for those walls was used. Also, it allowed the column dowel action to be included for the determination of total shear capacity. The p: rsent basic criteria uses the Schneider criteria as conservatively modified to account for upper limit compres-sion ef fects due to the method of testing. These modified Schneider criteria were then compared to test results which included low reinforcing ratios and cyclic ef fects. As mod-ified, the criteria predict capacities with excellent corre-lation to these test results. (Refer to Appendix B of the "Trojan Control Building Supplemental Structural Evaluation", September 1978.) The modified Schneider criteria are used only for walls that have confinement due to dead load and also have reinforcing steel in the concrete core. For the few walls that have re-inforcing in the block but for which no reinforcing steel is assumed to exist in the concrete core, the shear stress Fas been limited to a value of 150 psi. For horizontal sheet, the final criteria do not allow the strength to be increased if dowel strength is greater than basic criteria strength. The column shear strength has conservativsly been disregarded. 5-2 i
, , i R0sponsn'to NRC Questions question 5, continued The following compares the shear criteria used in the re-evaluation study to those presented in this submittal. The criteria used in the re-evaluation study relied on the ACI 318-71 provisions for shear walls but 'imited the concrete stress to 2 (fi)'" due to the low reinforcement ratio. These criteria were based on the following: V =4(2]t,ff+ 0.8W (4. SW where: V u
= ultimate shear force capacity of wall (lbs) t,ff = effective thickness of wall = total wall thickness -8 (in)
A s
= minimum f horizontal or vertical reinforcement (in 2
/ft)
W = length of wall (in) t, = yield strength of reinforcing steel (psi) ff = cc:rpressive strength of concrete (psi) 4 = capacity reduction factor for shear (0.85) T = gross wall t' tess (in) The above expression satisfies the ACI Standard 318-71 requirunents except that effective thickness is used instead of the gross thickness (T). Rewriting in terms of shear stress capacity: 0.68[2yt,ff+ ) , u T A 1.etting no
- 127 and using f =45,000 psi;f;=6000ps' y 105.4(T-8) + 30.600on u T l
M
#c s 5-3
Response to NRC Questions Question 5, continued-In the case of the Trojan Control Building, the typical rein-forcement percentages and the corresponding shear stresses evaluated according to the above relationship are given in Table 5-1. The reinforcement ratios do not include any contribution of the core reinforcement. In establishing rebar density, the block reinforcement is taken over the gross area. The total shear stresses shown in Table 5-1 are not representative of the potential shear stresses the walls can resist. The range is from 109 to 119 psi with an average of 1* 4 psi. Table 5-1 indicates that as the thickness of the wall increases, the sheat strength .(in terms of stress) decreases. This trend is not realistic since the thicker the wall, the higher will be the percentage of 6000 psi concrete that exists in its cross-section with a commensurate increase in shear strength. Moreover, these criteria conservatively neglect the aspect ratio, which is recognized in recent formulations in the ACI standards and the UBC. l TABLE 5-1 PSI (From Eq.5-1)
! t(in) fC)(%) (%) v v v
> ~ h c s u 20 .194 .183 63 56 119 27 .140 .190 74 43 117 30 .122 .172 77 37 114 34 .108 .152 81 33 114 40 .090 .130 84 28 112 48 .070 .160 88 21 109 5-4 6 - -
- _ _ _ _ _ _ _ _ _ _ _ _ l
-s: .
I Response to NRC Questions Question 5, continued In Table 5-2, the additional shear frc.a compression is not shown since it varies throughout the major walls. The primary ef fect of the dead load is to prevent the capacity from being limited by the bending moment. Applying both criteria to the West wall of the Control Building at elevations 45'-61' results in the following: TABLE 5-2
- ipecific Comparison Between original and New Criterla WEST WALL Af
- 0.68(2 r{ tg f + y) v' =0.75(348-113H/W) u T
- t H/W (PSI) (PSI) 41" .33 112 233 23" .26 118 239 5-5
Response to Question 5 Con' td Figure 5-1 shows a more general comparison of ' the . original and new criteria based on. the arpect ratio. 250 Walls with dead load and core reinforcing
- 200, 4
- Walls with major dead load U 150 without core reinforcing y --
3 . . . . . . . . - . - 3 C 0 g _ _ _ _ _ _ _ _ . _ _ _ _ _ _ _ , _ . _ . _ _ _ ___
- Re-evaluation study (average) /
100 - 1 2 m 50, Note:
- For comparison, the contribution of dead load is not shewn in this curve.
1,'O 2'. 0 3'. 0 Height (H) 710th 7
- Figure 5-1 Co@arison Of Shear Stresses 5-6 L_
/
Response to NRC Questions QUESTION 6 , 4
" At what OBE level below 0.15 "g" does the supplemental analysis indicate that nonlinear behavior of the wall systems will begin."
RESPONSE
Only concrete structures that have a cracking capacity higher than the applied loads would remain fully linearly-elcstic. Since concrete structures are not designed in this manner, all concrete structures experience some minor cracking and behave somewhat nonlinearly. For purposes of responding to this question, it is appropriate to consider two classes of walls; large walls and small walls. The shear force (loads) and capacities tabulated below were l obtained f rom Tables 5-1 and 5-3, of the "Trojan Control Building l Supplemental Structural Evaluation" dated September 1978. s As indicated by comparing loads and capacities, the small walls governed by moment have the lowest capacity compared to the load that would develop if they remained elastic. Therefore, the loads are fictitious. f The following is a tabulation for small walls for the N-S Direction: ELEVATION 45'-61' CAPACITY LOAD WALL (KIPS) (KIPS) CAPACITY (LOAD (Reft Figure 3-2) 2 470 780 .60 3 490 560 .88 6 110 340 .32 7 420 540 .78 8 60 290 .21 ELEVATION 61'-77' CAPACITY LOAD WALL (KIPS) (KIPS) CAPACITY / LOAD (Ref: Figure 3-3) 2 190 560 .34 3(Partial) 430 470 .91 6-1
\ - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _
a + Response to NRC Questions Question 6, continued The capacity of ell the small walls is only a small portion of the total capacit.y. Since Walls 6 and 8 at elevation 45'-61' and Wall 2 at elevation 61'-77' are for all practical purposes in signi fican t , they are disregarded. Considering the remaining walls, the capacity to. load ratio varies between .60 and .91. Since these walls have about the same strength, the average ratio (0.79) is used. An unfactored OBE of .159 at # = 2% is essentially equal to the .25g SSE at$ = 5%. Therefore, the factors determined for the SSE may be used. Assuming that nonlinearity starts when the average small wall yields, the OBE level at which nonlinear behavior begins can be computed as 0.12g. (0.15g)(.79) = .12g OBE. For large walls, the nonlinear behavior starts when cracking occurs. Cracking is assumed to occur at one-half of capacity, which is a stress level in the range of 130 psi. The following is a tabulation for the large walls in the N-c direction: ELEVATION 45'-61' CAPACITY C RAC KING LOAD WALL (!IPS) (KIPS) (KIPS) CRACKING / LOAD 1 5390 2695 4110 .66 4 3810 1905 2240 .85 5 5970 2985 3050 .98 ELEVATION 61'-77' CAPACITY C RAC KING LOAD WALL _(KIPS) (KIPS) (KIPS) CRACKING / LOAD 1 3100 2550 3910 .65 3 4520 2260 3140 .72 4 2240 1120 1910 .59 P 23 sed on Wall 4 a t elevation 61'-7"' , the OBE value at cracking or star t of nonlinearity would be ( .59)( .159) = .0899 However, Wall 4 is controlled by bending, and is not as important ac Walls 1 and 3. 6-2
- ~d +
Response to NRC Questiorg Question 6, continued l Wall 1 on both elevations is much more indicative of the overall structure. Moreover, they are outside walls. Using Wall 1 results ins (.66)(.15g) = .099 g (approx. . 10g) , In conclusion, it has been estimated that the structure still start to behave in a nonlinear manner between .10g and .12g. However, the total capacity is suf ficient to resist the SSE of .259 at 5% damping. e t 6-3 l
. w
_ Response to NRC Questions QUESTION 7 ,.
"Considering tu a nonlinear behavior of the Control Building indicated by thr supplemental STARDYNE analysis, provide a detailed discus ion of the expected impact of this behavior on the Fuel Buildi g to substantiate that there will be no detri-mental effects af this behavior on the integrity of the Fuel Building."
RESPONSE
In response to this question, the "Trojan Control Building Supplemental Structural Evaluation", dated September 1978, applies the STARDYNZ analysis to two cases. The first case is presented in Section 3.0 and involves a linear elastic analysis with fixed base. The second case evaluates all walls with load carrying capacities limited per "Case 4" of Section 7.0. In the second (Case 4) analysis, the concept of nonlinear behavior of walls for the N-S earthquake was introduced. In both analyses, only the responses in the first mode were evaluated because of the dominant participation of the Control Building in that mode; thus the differences in the effects of the decreased load carrying capacity of the Control Building on the Fuel Building were identified. Comparative shear forces are shown in Table 7.1. It was found that the assumed nonelastic vs. elastic behavior of the Conttol Building results in insignificant increases in membrane shears in the fuel pool and hold-up tank enclosure structures. To f ur ther emphasize the insensitivity of the Fuel Building to nonlinear behav..or of the Control Building, a separate investigation was made to identify the maximum forces that can be transferred ti the Fuel Building through the floor slab system. The tensile strenyth of the floor slab reinforcement sets an upper limit to the flexural stresses, and therefore to the moment that can be transferred from the Control Building to the Fuel Building. The moment associated with twice the yield strain of rebar in the extreme fiber of the slab and the corresponding membrane shear forces in the walls of the fuel pool and hold-up tank structure were com-puted. With these additional forces tr ansferred to the Fuel Building, the total membrane shear forces in the walls will be well below their capacities. No overturning exists under these conditions. 7-1 L
W i TABLE 7.1 SHEAR FORCES ASSOCIATED WITH N-S EARTHQUAXE, FIRST MODE (DOMINANT FOR THE CONTROL BLOG.) E 1.4 GBC (0.159), N-S aW 5"" FIRST SECOND CASE CASE 26 329 335 27 -252 -252 u T{gL 77 83
$ 35 577 567 58 6d 36 148 142 Ei5 37 29 24 38 -87 -97 TOTAL 667 636 E-W 24 714 719 25 265 272 a
Eg T L 979 99) w 33 888 914 34 125 124 A 1013 1038 y 27 5 35 36 37 38 26 Hold-Up Tank Enclosure Spent 33 Fuel 34 Pol Key to wall numbers 7-2 L
7_.
- g. .e-Response to NRL' Questions QUESTION 8 "Discuss the consistency of the "Basic Shear Wall Acceptance Criteria" with the results of the Schneider, Mayes and Clough, PCA, and other appropriate test data, especially with regard to cyclic loadings. Consider and discuss the dif ferences among the various test specimen and how the r esults for the test specimen are related to the Control Building walls.
Include a quantification of any differences."
RESPONSE
Appendix B of the "Trojan Control Building Supplemental Structural Evaluation," dated September 1978, provides the complete development of the shear wall criteria. It includes a discussion of the con-sistency of the shear wall criteria with the results of the Schneider, Berkeley, and PCA tests. As is further discussed in Appendix B, the Schneider criteria were conservatively modified to account for any induced compression due to the test set-up. The modified criteria were then compared to the recent Berkeley test results which involved cyclic ef fects. The correlation of the criteria with the Berkeley test results dernonstrates that the criteria account for cyclic degradation. 8-1
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