ML20063P299
| ML20063P299 | |
| Person / Time | |
|---|---|
| Site: | Midland |
| Issue date: | 10/08/1982 |
| From: | Hendron A CONSUMERS ENERGY CO. (FORMERLY CONSUMERS POWER CO.), ILLINOIS, UNIV. OF, URBANA, IL |
| To: | |
| References | |
| ISSUANCES-OL, ISSUANCES-OM, NUDOCS 8210130265 | |
| Download: ML20063P299 (119) | |
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{{#Wiki_filter:- - - . - - - - - - - -
. O 00CMETED USNRC UNITED STATES OF AMERICA 2 OCT 12 P250 NUCLEAR REGULATORY COMMISSION
,;,,.- .r Atomic Safety and Licensing Boar?.
In the Matter of ) Docket Nos.
)
CONSUMERS POWER COMPANY ) 50-329 OM
) 50-330 OM (Midland Plant, Units 1 ) 50-329 OL and 2) ) 50-330 OL
)
I, Alfred J. IIendron, Jr. , being first duly sworn,- state that my accompanying testimony evaluating the bearing capacity of the footings of the Diesel Generator Building is true and correct to the best of my knowledge and belief. Alfred J. liendron, Ur. SUBSCRIBED AND SWORN TO before me this 8 day of October , 1982. tum 8. Y//tdl Notary Public g2 0 30 5 I a b3 J'
r' UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION 4 . ATOMIC SAFETY AND LICENSING BOARD In the Matter of ) Docket Nos. 50-329 OM
) 50-330 OM CONSUMERS POWER COMPANY )
) Docket Nos. 50-329 OL
.(Midland Plant, Units 1 and 2)) 50-330 OL TESTIMONY OF DR. ALFRED J. HENDRON, JR.
This is the testimony of Dr. Alfred J. Hendron, Jr. My testimony is concerned with the evaluation of the bearing capacity-(ultimate load carrying capacity) of the footings of the Diesel Generator Building. The bearing capacity of the Diesel Generator Building footings ~was evaluated for the normal long term static loads, which consisted of the dead weight of the structure and.the design live load. The avail-able bearing capacity was also evaluated for the additional footing loads induced by the Safe Shutdown Earthquake (SSE) . On the basis of the analyses presented in this testimony I have concluded that the Factor of Safety against a bearing
-capacity failure is adequate under normal operating conditions and SSE conditions.
QUALIFICATIONS AND EXPERIENCE My detailed biographical, educational, and professional record is set forth in Attachment 1. The following is a summary. 8 4 .
k I completed the requirements for the degree of Bachelor.of Science-in-Civil Engineering from the University l 3 . of Illinois in June of 1959. In September of 1959 I enrolled in the Graduate College of the University of Illinois where I began full-time study on a University of Illinois Fellowship in tho' Department of Civil Engineering, specializing in Soil Mechanics and Foundation Engineering. I received the degree of Master of Science in Civil Engineering from the University of' Illinois in June of 1960. During the summer of 1960 I was employed by the consulting engineering firm of Shannon and Wilson, Inc. of Seattle, where I worked on foundations for the building complex which housed the World Fair in Seattle, and conducted an analysis of the stability of a 200 acre slope adjacent to the Pacific Ocean to be used in expert testimony by Mr. S. D. Wilson and Dr. A. Casagrande. In the fall of 1960, I began full-time study at the University of Illinois'towards a Ph.D. in Foundation Engineering and full minors in Geology and in Theoretical and Applied Mechanics. During this phase of my education, my academic adviser was Dr. Don U. Deere, under whom I studied engineering geology. I studied foundation engineering under Dr. R. B.. Peck. My studies on the application of shear strength of cohesive soils to prob'ams of slope stability and bearing capacity, problems use r analyzed by limit equilibrium methods, 9 4 -
were in courses taught by:Dr..R..B. Peck, Dr. Alan. Bishop,. and Dr.'D: U. Deere..
) . . . .
In June of 1961, I became a full-time-research associate at the University of Illinois where I conducted-a research project in soil dynamics under Professor N. M. Newmark and I taught an undergraduate course in Soil Mechanics and Foundation Engineering.unber Dr. R. B. Peck. These duties continued until the completion of my Ph.D.. thesis in the summer of 1963, when I.became a 1st Lieutenant in the U.S. Army Corps of Engineers. From September 1963 to September 1965 I was assigned to the-U.S. Army Engineer Waterways Experiment Station-as a. research engineer. At the Waterways Experiment Station I worked-in research in soil dynamics and structural dynamics as applied to problems.in the design of protective structures. I was also-instrumental in starting a rock mechanics research program at the Waterways Experiment Station, and I served as'a special consultant for the office of the Chief of Engineers on construc-tion problems associated with the Minute Man System in North Dakota and as a consultant to NATO in Norway on the construction of a large underground cavity in rock. Since September of 1965 I have been employed in teaching and research at the University of Illinois in the Department of Civil Engineering, as an Assistant Professor from 1965 to 1968, an Associate Professor from 1968-70, and
-9 4 -
as a Professor of Civil Engineering from 1970 until present. 3 During this time I have taught graduate courses in soil and rock dynamics, applied foundation engineering, earth dams, rock mechanics, and rock engineering. I have also taught undergraduate courses ir. Introductory Soil Mechanics and Foundt ~.on Engineering.- I have been active in research on slope stability in rock masses, design of tunnel linings in soil and rock, ground vibrations produced by construction blasting and by nuclear explosions, theoretical studies of inelastic and time-dependent stress distribution around tunnels in soil and rock, and compressibility of large sized granular materials such as rock fill materials. While a staff member at the University of Illinois, I have been engaged by various public agencies, consulting engineering firms, utility. companies, and foreign governments to consult on the design and construction of nuclear power plants, dams, tunnels, and slope stability problems. A number of these consulting assignments are listed in detail in Attach-ment 1. I have previous experience in bearing capacity analysis related to fills placed on soft clays. In addition, I have either checked or made calculations for the bearing capacity of footings under both static and earthquake laodings t 4 -
4 forLmany conventional structures as.well as-for nearly- .. all of the nuclear plants listed in my experience record shown in At'tachment 1. Diesel Generator Building Description and NRC Inquiries Regarding Bearing Capacity The Diesel Generator Building for the Midland Nuclear Station is located to the south of the Turbine Building as shown in Figure 1. The building is founded on a 10 ft wide continu-ous wall footing around the perimeter of the stru=ture; in addition three North-South wall footings, also 10 ft wide, support the walls which separate the four bays of the 4 Diesel Generator Building. A plan view of the footings is shown in Figure 2. The strip footings analyzed and reported in'this testimony are shown by the dark shaded areas in Figure 2. A cross-section taken from west to east through the Diesel Generator Building is shown in Figure 3. The base of the footings are founded at elevation 628 and the finished grade elevation is 634. The static loads including . dead and live loads are 48.54 kips / lineal ft along the l [ perimeter wall which yields a static contact pressure of ( 4.86 ksf. For an SSE of 0.12 g, the dynamic increment of wall loading results in an increase of 1.20 ksf in the footing pressure. An evaluation based on contact stresses due to a i , dynamic increment of 1.5 times the SSE values is also given l l l l- I t 4
,' c , _
, t 5
- . - later_in the testimony. The-evaluation for a dynamic increment of,1.5 the SSE
- value was done to coaform to a commitment made by-
' CPCO-to the NRC.
.Before surcharging in' January of 1979 there_was an-exploratory program conducted in-the plant area by Bechtel from August 1978_to December 1978 whichwincluded borings, Dutch cone soundings, and a test pit. .A' series;of
!! "undrained" triaxialsshear strength tests were performed on=- l-samples'taken from these borings. 'In-addition dry;. densities-l and waterEcontents were determined for many samples taken
. from beneath the Diesel Generator Building. I have been
, involved with evaluation of the foundation. soils for the
- ; Diesel Generator Building since October of 1978, when I-made my/first field inspection of. test pits adjacent to the
$ Diesel ~ Generator building foundations after surveys: detected i settlement-of the structure,before' construction was complete. F In the applicants' response to NRC Question 35 (10-1 CFR 50.54f) , the ~information. cited above was used to estimate - the bearing. capacity-of the foundations under static'and , combined. static and SSE loadings. The analyses ~ presented by the applicant.were rejected as " unsatisfactory" by the Corps l-of Engineers in a letter to Dr. Robert:E. Jackson from James i Simpson:on July 7,_1980,.because "... the basis of selection
~
i j of' shear strength for computing bearing capacity does not p reflect'the' characteristics of soils under the Diesel Generator i i g 14 .] ' .
Building. . A bearing-capacity computation should be submittted 3 . based on test results of samples from new borings which we have requested in a separate memo." Consumers Power has subsequently conducted the additional boring and sampling program requested by the Corps of Engineers. .On some of the samples from the area of the Diesel Generator Building, triaxial tests have been~ con-ducted by Woodward-Clyde Consultants on representative samples. These test results were used to calculate bearing capacity of the Diesel Generator footings under static and earthquake loadings. These triaxial tests were conducted within the consolidation pressure ranges which I recommended. The consolidation pressures utilized are given on page 19 of this testimony. In this testimony I have reviewed both the Goldberg Zoino triaxial tests, conducted on samples obtained before surcharging, and the results from Woodward-Clyde tests, con-ducted on samples which were taken after surcharging. The bearing capacities calculated are based on the Woodward-Clyde test data because the densities and strengths of these materials inherently include the effects of surcharging. It is clear from a comparison with the Goldberg Zoino tests however that a conservative estimate of bearing capacity could have been made on the basis of the Goldberg Zoino P
1 1 I I triaxial-test. data and on measurements of dry density made s by Goldberg Zoino on soil samples from beneath the Diesel Generator Building, even if the Woodward-Clyde tes ts -had not ; been conducted. Fundamentals of Static Bearing Capacity and Definition of Factor of Safety Figure 4 represents a cross-section through a long foo ticg , such as the footing beneath the north or south wall of the Diesel Generator Building.- The footing has a width B and'is founded at a depth D f below the ground surface. The quantity D g is referred to as the surcharge depth. If the footing illustrated in Figure.4 fails, the wedge of soil a'o'bd, as shown on the left side of Figure 4, must be displaced upward and to the left. The. weight of the wedge and the shearing strength of the soil along a'o'bd tend to resist-failure. Thetotalultimatecontactpressuregj (Figure 4) which is developed at failure for gilen shear strength parameters of the soil is computed from accepted analytical techniques, (Terzaghi and Peck, 1948), (Terzaghi and Peck, 1967), and is expressed as: I 1
1-gj = 7 BY Ny + cN + YD gNq .Eq. (1) 3 . where: .B = width of the footing Df = surcharge depth Y = unit weight of soil c = cohesion of soil N,N.Ny q c = bearing capacity factors which are funtions of p, the' angle of shearing resistance of the soil. The first term of Eq. _(1) represents the portion of the bearing capacity due to-the weight of the material in wedge a'o'ba (Figure 4). The second term represents that portion of the bearing capacity due to the cohesive shearing resistance along surface a'o'b (Figure 4), the third term _is that. portion of the bearing capacity due to the surcharge
' pressure, YDf, acting on the plane a-b or a'-b' (Figure 4).
In the analytical procedures it is usually assumed that the
- influence of the soil above the base level of the footing can be replaced by the uniform pressure YDf as shown on planes a-b and a'-b' (Figure 4). This is conservative since
, the shearing resistance of the soil above the base elevation, such as the' shearing resistance along b-d (Fig. 4), is neglected. The value of qf from Eq. (1) is the total ultimate bearing capacity. The " net ultimate bearing capacity", gd' is defined as the pressure that can be supported at the base of the-footing in excess of the pressure at the same level due to the surrounding surcharge; thus, i T 1
9d"9d - YD f
.Eq. (2) 3 ,
and 1 9d= 3 BYNy + cN c +YDf(Nq-1) ~Eq. (3)' The total applied contact pressure at the base of the footing is designated as qt. The net applied pressure at-the base of the footing is the total applied contact pressure minus the adjacent surcharge pressure, YDf. Thus the net applied pressure,-qnet, is given by 9 net " 9t - YDf Eq. (4) The Factor of Safety under static loading is
' defined.as the ratio of'the available net ultimate bearing capacity to the net applied pressure. Thus, F.S. = qd = qA - yDe 9 net 9t - TDf Eq. (5)
For long term static loads the' net ultimate bearing capacity, qd, can be calculated from Eq. (3) using the shear strength parameters c' and.d' in terms of effective stresses. This method of calculation is appropriate for calculating i the static factor of safety for the Diesel Generator Building
-considering the dead load plus live load acting on the footings. Under these circumstances it is usual engineering practice to require the calculated factor of safety to be l greater than 3.0.
f 1 f 1
Factor of Safety Under-Combined Static and Earthquake Loading 3 . For.the general case of considering both static and' earthquake loads on the footing, Eq. (3) cannot;be used because the soil behaves as if it is "undrained" under the dynamic pressure increment induced by the earthquake. Thus in an analysis of combined static and earthquake loadings it is necessary to first consider the initial static " effective" stresses developed along a potential failure' surface because , these " effective" stresses will govern the undrained shear
- strength appropriate-for.use in the analysis. For this i-purpose a potential sliding mass, as shown in Figure 5 was divided into slices as is'normally done in limit equilibrium analyses of slope stability problems. The Method of Slices used in this analysis is that given by Lowe and Karafaith (1959), which is described and illustrated in EM 1110-2-1902 of the Corps of Engineers. The procedure is a trial'and error method in which I have taken c' = 0 and have assumed-different values of the'" effective" drained angle of shearing resistance mobilized, d', at the base of each slice until-the value of pf is found which-is just necessary to maintain
. static equilibrium. The normal " effective" stress, &n, and the shear stress, in mobilized at the base of each slice
.can then be determined.
f 1
. . _ -. _ . - . _ - . _ _ . . . _ . . ._ .~ - _ . . _ _ _ . _ - _ _._ - . _ .
The magnitude of the principal stresses'at the base s . of each slice can then'be calculated by assuming that the base of the slice coincides with the failure plane and that the orientation of the principal planes and failure plane are
. independent of the factor of' safety. For these conditions the principal stresses El ""d 3 can be estimated from 31= Eq. (6) n + *n tan &
B3" n -*n. 1 Eq. D) tan $ where v = 45 + d'/2 Eq. (8) In this procedure it will generally be found that the ratio of E1/33 will be nearly constant for all slices and can be used as a guide in selecting the " effective" stress ratio to-use in anisotropically consolidated undrained shear strength tests. The absolute values of d i and a3 f r all slices will indicate the range of cohfining stresses which should be ' used in the consolidation phase of consolidated-undrained
. tests.
From the analysis of the static conditions described above, consolidated-undrained tests can be conducted at the indicated ratio of 61c/33c and the absolute values of &lc and &3 c. Thus the Mohr circle for the consolidation stresses is shown by the dashed circle in Figure 6. The Mohr circle at failure for this sample tested undrained is shown in f 1
e Figure 6 as the solid circle. Thus T ff shown on Figure 6 3 , - is the undrained shear strength on the failure plane which was consolidated at an effective normal stress U n. For this test it is appropriate to plot T ff vs E n as illustrated 4 by point'A, Figure 6. Thus with a series of tests a diagram is developed of undrained shear strength on the failure plane Tff versus Un , the effective normal consolidation pressure.on-the same plane. This diagram is useful because from the static analysis described above, the effective normal consolidation stress 6 n can be determined at the base of each slice. Then from Figure 6 the undrained shear strength, T ff, available at the base of each slice'can be determined. When the combination of static foundation loads plus additional foundation loads due to earthquake are considered, the undrained shear strength, T gg, at the base of each slice is the shearing strength which can be mobilized along the failure surface. In order to calculate the factor of safety under combined static and earthquake loading, it is necessary to calculate the ultimate load, P , which the footing can sustain, under this undrained loading condition. The main sources of resistance which contribute to the ultimate bearing force, Pu, which the footing can maintain are: I 1
a) the presence of the surcharge 3 , b) the weight of the sliding mass c) the undrained shear strength which can be mobilized along the surface of sliding. The computation of P is illustrated in Appendix A for various failure surfaces. After P u is determined it is then possible to cal-culate the factor of safety. The net bearing capacity available is qd = Pu/B -YD g Eq. (9) and the net applied pressure is gnet " 9t + Oe -YD f Eq. (10) B where qt = total contact pressure under static loading Q e = Additional earthquake force per unit. length of wall Thus the Factor of Safety is given by F.S. = qt + O e/B -Y Df Eq. (11) Shear Strength of Foundation Materials Before Surcharge Shear strength tests and dry density determinations were conducted on a testing program carried out by Goldberg Zoino from August to December of 1978. The samples were f 1
obtained from borings throughout the plant site and in a part.of this study nineteen undrained triaxial tasts were conducted on compacted fill. The pertinent information about these samples regarding boring location, depth, initial dry unit weight Yd, and water content, w%, as well as initial confining pressure are given in Appendix B. Strength paramecers for the materials underneath the Diesel Generator foundations before surcharging were estimated by: 1- Determining representative values of the dry density of materials underneath the Diesel Generator building and 2- Using shear strength parameters obtained from triaxial tests carried out with samples of similar dry densities. Effective strength parameters, c' and p', for the materials under the Diesel Generator building were estimated from triaxial tests carried out on compacted fill samples with dry densities similar to those of the fill materials under the DGB. Pore pressure measurements taken during these tests made it possible to calculate the principal i
" effective" stresses di and 63 at failure. From these results Modified Mohr Coulomb diagrams in terms of " effective" stress were plotted as shown in Figures 7 and 8. These diagrams f 1
I i indicate'that the compacted fill has effective strength 3 , parameters of c' = o and p' = 30*. Undrained shear strengths representative of the fill materials under the Diesel Generator Building before surcharging were-obtained from triaxial, tests. Maximum undrained shear strength values, (al -c3) , measured in 2 max these tests were plotted versus the isotropic " effective" consolidation stress on the sample prior to shearing (Fig. 9). Undrained shear strengths of the materials under the Diesel Building, were estimated from the relationships shown in Figure 9 for possible use in a bearing capacity analysis with full knowledge that these values do not account for any increase in shear strength due to increases in density which may have occured during surcharging. Fig. 10 shows the data from Fig.-9 plotted in the form of i gg, (the undrained shear strength on the failure plane) versus E n, the initial consoli-dation stress on the failure plane, assuming the failure plane is at 45+p'/2 for the major principal plane, where p' was taken as 30*. The values of T gg in Fig. 10 are below values of S y given in Fig. 9, as would be expected and thus would yield more conservative shear strengths to be used in an undrained analysis to evaluate bearing capacity under static and earthquake loadings. I l
Shear-Strength and Density of' Fill Determined After
-Surcharge During the period 13 March 1981 to June 1981, Woodward-Clyde Consultants contucted a soil boring and testing program of the fill and natural soils immediately adjacent to the Diesel Generator Building. The final report containing.the test results on soils sampled at the Diesel Generator Building are given in Woodward-Clyde Consultants' report of 8-July 1981. The boring locations adjacent to the Diesel Generator Building are shown in Fig. 11.
Stratigraphy borings COE-8, 9, 10, 11, 12, and 13 (13R) (Fig. 11) were continuously sampled, below a depth of 5 ft, using " undisturbed" sampling techniques. Typically, undisturbed samples were obtained (and retained) in 3-in.- diameter thin-walled tubes using-Osterberg, Hvorslev, and Pitcher samplers. Stratigraphy borings were advanced a minimum distance of 5 ft into foundation soils underlying fill materials, and Standard Penetration Tests (SPT) using a split-barrel sampler were performed in the bottom 5 ft of the borings. Following processing of the stratigraphy boring tubes in the field laboratory, the sampling intervals in the engineering property borings were selected. Engineering property borings CO2-9, 10A, 11A, 12A (12B), and 13A (13B) were typically drilled within about 3 ft of their corresponding stratigraphy borings. These borings were logged and the I l
cohesive fill materials were continuously sampled in 3-in.-diameter thin-walled tubes using Osterberg and Pitcher samplers. Scheduled. engineering property boring COE-8A was not drilled because cohesive fill materials were not encountered in stratigraphy boring COE-8. Index property testing included processing of tube samples and determinations of density,'ater w content, consistency (pocket penetrometer) , liquid and plastic limits, particle-size distribution, and specific gravity. Strength testing included isotropically consolidated-undrained l triaxial compression tests with pore water pressure measure-ments (CIU) and anisotropically consolidated-undrained triaxial compression tests with pore water pressure measurements (CAU). Compressibility testing consisted of one-dimensional consolidation tests with an unload-reload cycle. Index Property Testing Index property test results and other pertinent data are listed for borings COE-9 through COE-13B in Table 1. In this table the elevation of each sample, the Unified Soil Classification, the total unit weight, the dry unit weight, and the water content are given for each sample. In general the information on the dry density, Y'd, indi ates that the dry densities to a depth of 10 ft below the footing elevations I l
. . - .~ . -- . . . .
T 2 (628 '618) are higher'than those determined-from samples before surcharging by. Goldberg 1and .Zoino. (Appendix LB) .. ' The
, average dry density,y d, f soil' samples between elevations 618.and 628-from Tablell:is'125 lbs/ft3 as determined from forty-four samples.
Strength Testing
- Strength testing' included one series each of'CIU 1and CAU triaxial compression tests performed-on specimens of cohesive fill from the DGB borings.- .In all but one case, test specimens-were selected.from. representative high-
. quality samples-within'the elevation range (el. 628 to el.
618) requested by the writer; one'CAU test,-however, was performed on a test specimen from el. 617.2. [ A series of'five CIU triaxial' tests was performed E
'on cohesive fill specimens consolidated.to pressures ranging from 0.7 to 8.2.ksf. In addition, a series of six CAU tri-1 axial tests was performed on similar cohesive-fill specimens.
These CAU tests were anisotropically consolidated at an effective principal stress ratio (k e
- lc/ 3c) f 1.9 and at maximum " effective" principal consolidation stresses, (o '
lc of 1, 2,~4, 8, 12, and 16 ksf. [ The sample location, dry density, water content, . soil classification, and effective consolidation pressure
'for the CIU tests'are shown in Table 2. The sample location, a.
.e water content, density, soil classification, and the major and minor principal stresses for the consolidation phase are given for samples tested in CAU Tests in Table 3.
The results of both the-isotropically and aniso-tropically consolidated triaxial tests are tabulated in Table 4. For each test the dry density, water content, and the initial effective consolidation stresses for the first phase of the triaxial test are given. 'Also listed for each test are one-half the maximum stress difference (f'l-' 2 03) at failure as deter-mined from the maximum stress difference failure criteria, the strain at failure, and the A coefficient at failure. From the data presented in Table 4 a plot can be made, as-shown in Figure 12, of 1/2 the stress difference at failure as ordinate versus 1/2 the sum of the major and minor principal I . effective stresses at failure as abcissa. As shown on Figure 12, the slope of a straight line through the data points and through the origin gives a slope, a ,oof 27*. As shown on Figure 12 this corresponds to an effective angle of shearing resistance, d', of 30.5 . The information shown in Figure 12 is the basis for my choice of using an effective angle of shearing resistance of 30* to represent the effective angle of shearing resistance for the compacted soils beneath the Diesel Generator Building. It should be noted that the same conclusion was reached by the writer at an earlier date I l
fromLthe Goldberg Zoino triaxial data shown in Figures 7 tnd 8 of-this testimony. According.to the graphical construction shown on Figure 6, the triaxial data shown in Figure 13 for both the isotropically and anisotropically consolidated undrained tests have been plotted in a form which shows the effective normal consolidation stress on the failure plane, En, f r each sample versus the.undrained shear strength,T gg, which was developed on that plane. -Also shown on Table 4 are the maximum values of the undrained strengths, 1 03 ,. 2 which occurred for the same tests. It should be noted that the orientation of the failure plane was taken as 45 + d'/2 with the major principal plane where d' was taken as 30 . It should be noted from Table 4 that the values of T gg are always about 16% lower than the maximum values of the undrained shear strength. In analysis of stability problems involving undrained shear strength it is conservative to use the' values of T gf ratherthan(1- 3] A plot of the relationship between the undrained shear strength, T gg, on a potential failure plane versus the effective normal consolidation stress on that plane from Figure 14 is shown for anisotropically connolidated data on test smmples taken from Table 4. The solid line shown in f I
_.y s e . .
- 1 Figure .14 is ' the re'lationship between undrained strerigth and !*
initial consolidation pressure'which~best epresents the' undreined shear' strength which can'be' developed at the~ base
~
offafpotentialLsliding' surface'beneath,the Diesel Generator a
. Building fotndation.- I feel ~.that this relationship is the .
i best-representation of~the undrained shear strengths of~the materials-which-should beLused in a calculation of the factor ? , of. safety because..itLaccounts for the initial' anisotropic.
,. stresses which exist in the ground.due to the long. term static loads;- and, . the test data shown ;i.n Figure L 14 represent 1 L the properties df the fill after application of the surcharge. '
3 Thus the undrained shear strengths are slightly higher than=
- f, :the'undrained shear strengths'which would have been predicted
.from .the Goldberg-Zoino test; data' shown in Figure 10.
Calculated Factors of Safety Against-a Bearing Capacity Failure o Long Term Factor of Safety Under Static Loads , Under long term static loads the pore pressures i produced'by shearing strains along a potential failure t p surface can.be assumed to be zero because they have had a p .long; time to dissipate and the appropriate shear. strength parameters for assessing bearing capacity-for this situation v
- are the long. term or effectiveishear' strength parameters c',
.+ and p'. The net bearing capacity, qd, fo'r these conditions i-t s' I * .3 I 6 , ~ ~
*T^s Y
s
] ., f
. . , _ .,, _ - _ _ . _ __ _ _ . _ _ _ =. _ m _
is given. by Eq. - (3) . As discussed previously, the maximum. ground. water elevation considered was at elevation 627, and for the first term of Eq.- (3) the submerged unit weight Y ' -was used. The effective angle of shearing resistance, d', has been taken as 30' and in order to evaluate the bearing capacity factors N y, N in Eq. (3). q, and N The values of N y, N g and Nc -corresponding to an angle of shearing resistance of 30* are 22.4, 18.4 and 30.1 respectively, according to Vesic (1975). If the foundation width, B, is considered to be 10 feet and the~ surcharge depth, D f, is considered to be six feet, then the net bearing capacity, qd is given by. gd " hl lb/f t )3 (lo' f t) (22.4) - (135 lb/f t )3 (61f t,) (17.4) = 22,214 psf for c' = 0. /' Thus the long term fact'or of; sEfety'is given by
, F=
F.S. = 9a = 22,214 psf =15.49 qt -YD f 4,044 psf In foundation engineering it is commo,n-to design for a factor of safety of 3 when considering th$ dead load of the structure plus ordinary live loads. Thus the value of 5.49 ilndicated above is considered adequate. 6 . For'the permanent dewatered condition, the unit weight, Y , should be 135 pcf,for all terms'in Eq. (3). Thus 4 J 9 3
. e the net bearing capacity would be given L/
gd = 1/2 (135) pcf (10) ft (22. 4 + - (135) pcf (6) ft (17.4)
= 29,214 psf Thus the long term factor of safety for the dewatered condition is ,
p*3* . 29,214 psf 4,044 psf = 7.22 Summary.of. Calculated Factors of Safety For Combined Static and Earthquake Footing Loads The mechod for computing the factor of safety against a nearing capacity failure for the combined static and earth-quake loads has been .given in an earlier section of this testi t mony. The details of these calculations for all cases f considered are given in Appendix C. For all cases consi[ered the total contact pressure at the base of the footing due to the static loads was considered to be 4.85 ksf. For an SSE of 0.12g, I have been given, by the Bechtel structural engineers, an additional pressure increment of 1.20 ksf at the base of the 10 ft wide footing. As indicated in Appen-
- dix C, I have considered a deep circular surface of sliding, a shallow failure surface, and an intermediate depth failure. ,
surface to ensure that the analysis would include a slipsur-face-with the minimum factor of safety. For each of these cases the water level was considered to be at elevation 627. A summary of the calculations which are included in Appendix C is given in Table 5. f 1
From an inspection of the results in Table 5 and the calculations given in Appendix.C, it is seen that the intermediate depth surface designated as case V yields the minimum factor of safety. For this case a triangular wedge is considered under the footing and a circular surface is considered between the triangular wedge under the footing and a passive triangular wedge out to the side of the footing. The factor of safety for Case V for the anisotropic undrained shear strength given in Fig. 14 is 2.67; and, the factor of safety for case V for the lower isotropic undrained shear strength given in Fig. 13 is 2.37. For the combined effects of dead load, live load, and loads with a low probability of occurrence such as earth-quake loads, it is common engineering practice to design footings to have a factor of safety of about 2 if these loads are considered to act simultaneously. For both shear strength > relationships mentioned above the factor of safety is suffi-ciently above a factor of safety of 2 to be considered adequate for the performance of the Diesel Generator Building under the SSE. In fact, when the permanent dewatering is considered to lower the water table below the slip surfaces assumed for Case V, then the factor of safety of 2.67 above increases to 2.85 and the factor of safety of 2.37 above increases to 2.54. f 1
If the dynamic increment is considered to ima 50% greater i than that for the peak acceleration associated with an SSE of
.12 g, then the . increase in footing loads due to the SSE will be 1.80 ksf. If the earthquake load on the footing is considered to be 1.80 ksf instead of 1.20 ksf, then the factor.of safety of 2.85 referenced above would decrease to 2.55 and the factor of safety of 2.67 referenced above would be reduced to 2.40.
Thus I believe the factors of safety against a bearing capacity failure indicate that this mode of failure is precluded for the dewatered case, even with a more intense SSE than has previously been considered. 4 f 1
CONCLUSIONS In previous sections of this testienny I have presented analyses for calculating the ultimate bearing capacity for the DGB foundations for static loadings and for combined static and earthquake' loadings. The factor of safety against a static bearing capacity failure was found to be greater than 5.0 as compared to an acceptable value of 3.0 in foundation engineering practice. The factor of safety against a bearing capacity failure for combined static and earthquake loads consistent with an SSE of 0.12 g was found to be greater than 2.6 if the effects of dewatering are taken into account. The factor of safety was shown to be equal to 2.4 if an SSE consistent with increasing the dyna-mic forces by 50% is considered. It is acceptable practice for the factor of safety to be greater than 2.0 for-combined static and earthquake loading. It is my opinion that the results of the calculations quoted here indicate that the Factor of Safety against a bearing capacity failure are high enough to preclude a bearing capacity failure under combined static and earthquake loads considered. f 1
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RESUME PROFESSIONAL BACKGROUND AND EXPERIENCE N ame: ALFRED J. HENDRON, JR. Address: 2230c Newmark Civil Engineering Laboratory University of Illinois at Urbana-Champaign Urbana, Illinois 61801 Date of Birth: October 4, 1937 Marital Status: Married with 2 children Citizenship: Natural Born - U. S. EDUCATION Ph.D. 1963 University of Illinois Major: Soil Mechanics Urbana, Illinois Foundations Minors: Geology Theoretical and Applied Mechanics M.S. 1960 University of Illinois Civil Engineering Urbana, Illinois B.S. 1959 University of Illinois Civil Engineering Urbana, Illinois (Bronze Tablet) NATIONAL AND INTERNATIONAL AWARDS OR RECOGNITIONS Received American Society of Civil Engineers Walter L. Huber Research P rize,1974 Invited to give 6th Nabor Carrillo Lecture, by Mexican Society for Soil Mechanics, November 1982 Listed in Who's Who in Engineering, 1980 POSITIONS HELD September 1970 - Present Professor of Civil Engineering, University of Illinois September 1968 - September 1970 Associate Profesor of Civil Enginering, University of Illinois September 1965 - September 1968 Assistant Professor of Civil Engineering, University of Illinois September 1963 - September 1965 1/Lt. U.S. Army Corps of Engineers Research Engineer U.S. Army Engineer Waterways Exp(riment Station
Alfred J. Hendron, Jr. Page 2 June 1961 - September 1963 Research Associate June 1960 - September 1960 Engineer, Shannon & Wilson Soil Mechanics and Foundation Engineers, Seattle, Washington TEACHING EXPERIENCE Undergraduate Courses, University of Illinois 1961 - 1963 Introductory Soil Mechanics 1965 - Present Introductory Soil Mechanics
' Foundation Engineering Civil Engineering Design Course for Senior Honors Students Graduate Courses, University of Illinoit 1965 - Present Earth Dams Rock Mechanics Applied Rock Mechanics Applied Soil Mechanics Soil Dynamics (includtag earthquakes effects)
RESEARCH EXPERIENCE 1961 - 1963 Research Associate, University of Illinois Conducted research on the high pressure compressibility of sands and measurement of the coefficient of earth pressure at rest. 1963 - 1965 U.S. Army Engineer Waterways Experiment Station Conducted research on stress wave propagation in soils, design of structures for dynamic loading, and developed a research program in rock mechanics. 1965 - Present University of Illinois i Presently conducting research on the follnwing specific topics: (1) Ground vibrations produced from blasting tunnels and open cuts in rock. (2) Compressibility of large sized granular materials such as that used in rock fill and rolled earth dams. (3) Theoretical studies of inelastic and time dependent stress distribution around tunnels. (4) Effect of pore pressures on the strength of rock. (5) Three dimensional analysis of slope. (6) Design of tunnel linings in soil and rock. (7) Earth dam design.
---c ,- - - -- --- - -
Alfred O. Henorco, Jr.
- Page 3 OFFICES AND OTHER SERVICES TO PROFESSIONAL SOCIETIES I
(1) Member of the Research Committee of the Soil Mechanics and Foundations Division of the American Society of Civil Engineers, 1967-69. f (2) Member of the Subcommittee 12 of Committee 0-18, ASTM, Properties of Soil and Rock, 1965-1970. (3) Co-chairman of Panel on " Stress Wave Propagation in Soils," International Symposium on Soil Dynamics, Albuquerque, New Mexico, sponsored by ASCE & NSF, August, 1967. (4) Panel member for " Dynamic Loading," session of a National Specialty Conference on Placement and Improvement of Soil to Support Structures," sponsored by the Soil Mechanics and Foundations Division of the American Society of Civil Engineers, M.I.T., August,1968. (5) April,1968 - Gave lectures on rock mechanics to Metropolitan Section ASCE, New York City. (6) April 1969 - Gave lectures on rock mechanics to Metropolitan Section, Washington, D.C. (7) Selected to give a lecture on " Field Instrumentation in the Design of Underground Structures in Rock," Metropolitan Section, ASCE, New York City, May, 1970. (8) Panel mem*aer on " Dynamic Loadings and Deformations," session for ASCE, l Soil Mechanics and Foundations Division Specialty Conference on " Lateral )- Streses in the Ground and the Design of Earth Retaining Structures," Cornell University, June, 1970. (9) Panel member on " Deformation Modulus of Rock Foundations," ASTM Symposium on Deformation Properties of Rock, Denver, February, 1969. I (10) Selected by NSF as one of the U.S. members to Exchange Meeting with Japanese Engineers on the topic of Ground Motions Produced by Earthquakes, U. of California at Berkeley, August, 1969. (11) Member of Committee on Soil Dynamics, Soil Mechanics Division, ASCE, 1970-1972. (12) Member of Publications Committee for Journal of the Soil Mechanics and Foundations Division, ASCE, 1970-1972. (13) Member of Committee on Rock Mechanics, ASCE,1978-preseat. (14) Member Corps of Engineer Advisory Board for Geotechnical Engineering Research, 1978-1981.
Alfrec J. Hendron, Jr. Page 4 EXAMPLES OF FOUNDATION ENGINEERING AND EARTHQUAKE ENGINEERING EXPERIENCE (1) Consultant to Williams Brothers Construction Company on slope stability problems encountered in construction of the Transandean Pipeline in southern Columbia, S.A. (2) Consultant to Woodward-Clyde and Associates on the Foundation Design of Davis-Besse Nuclear Reactor for earthquake loadings.. (3) Consultant, as an associate of Dr. N.M. Newmark, on the foundations for a 40-story building in Vancouver, B.C., design for earthquake loading. (4) Consultant to Waterways Experiment Station on the Earthquake Stability of Dam Slopes. (5) Consultant to H.G. Acres Ltd. on seismic conditions for Nuclear Reactor Foundations as a part of a study for 6 New England States on Projected Power Needs. (6) Consultant, as an associate of Dr. N.M. Newmark, to the Divisions of Reactor Licensing and Reactor Safety of the Atomic Energy Commission, on the adequacy of nuclear reactor foundations to resist earthquake loading, September 1967-1980. The following is a list of the Nuclear Power Station Foundations reviewed during this time: Ft. Calhoun Arnold Cooper Pilgrim #1 Surry Crystal River Shoreham Prairie Island Salem Farley Rancho Seco Calvert Cliffs Diablo Canyon Oconee Sequoyah Indian Point Hatch D.C. Cook Brunswick Zimmer Kewaunee 3 Mile Island Fitzpatrick Russellvile Fermi Easton Turkey Point LaSalle Bell (7) Dynamic stability assessment of 3 TVA dams subjected to design earth-quakes. (8) Consultant on dynamic stability of Jackson Lake Dam to Bureau of Reclamation. (9) Consultant on re-evaluation of foundations for the Olympic Tower and Stadium structures, Montreal, Quebec. (10) Cerron Grande Dam, El Salvador, Consultant to Harza Engineering, lique-faction potential of foundation.
-AlfregJ.Hendron,Jr. ,
Page o (11) Dynamic and static stability of %th soil and rock slopes for Alyeska Pipe Line Service Company, Alaskan Pipe Line. (12) . Dynamic analysis of cooling pond dike for Skadgit Nuclear Station, Consultant to Bechntel Engineering. (13) Assesment of the static and dynamic stability of Chatuge and Nottely Dams, Consultant to TVA. (14) Kapong Hydroelectric Project, Ghana, assessment of dikes, Consultant to TVA. (15) Suarez River Project, Colombia, S.A., Evaluation of four dam sites for dams ranging in heights from 500-600 ft in height. (16) Pond Hill Dam,100-ft high dam, Consultant to Tippetts, Abbett, McCarthy
& Stratton (TAMS).
(17) Susitna Project, Consultant to H.G. Acres on 800-ft high earth dam and 600-ft high arch dam for Alaskan Power Authority. (18) Remedial measures for Tarbela Dam, Service Spillway, Pakiston, Consultant to TAMS; static and dynamic slope stabilt/. (19) Libby Reregulating Dam, member of Board of Consultants, Seattle District, Corps of Engineers. (20) Consultant to TAMS on Department of Interior study to evalute the design procedures of the Bureau of Reclamation for the design of earth dams; static and earthquake. (21) Evaluation of site for Patia II Dam for Hidrostudies Company of Bogota, Colombia; 600-ft high dam on Patia River; static and dynamic stability. EXAMPLES OF ROCK ENGINEERING EXPERIENCE (1) Consultant to the American River Constructors on the stability of 300-ft high rock slopes for the spillway cut at Hell Hole Dam, American River Project. (2) Consulted on rock mechanics problems related to the foundations of 'che World Trade Center Building, New York City (110-story office building). (3) Consultant to New York Port Authority on Controlled Blasting Techniques to reduce damage to adjacent structures for Journal Square Subway Terminal. (4) Consultant to Western Contracting Company on stability of 150-ft high vertical spillway cut, Stockton Dam, Stockton, Mo.
Alfred J. Hendron, sr. Page 6 (5) Consultant to British Columbia Hydro Authority, Canada, on assessing stability to Portage Mountain Underground Powerhouse. (6) Consultant to Fenix and Scisson on the design of a rock cavity and steel casing at a depth of 6,000 ft in weak rock on Amchitka Island. (7) Slope stability problems along the Transandean Pipeline, Colombia, S.A., for Williams Brothers Construction Co. (8) Consultant to Joseph S. Ward, Foundatior. Engineers, on the design of a school to resist blasting vibrations, Manchester, New Jersey. (9) Consultant to Architect's Collaborative, Cambridge, Mass., on controlled blasting technques and blasting vibrations on IBM building complex, Fishkill, N.Y. (10) Stability of rock slopes for Trans-Alaskan Pipeline Terminal, Valdez, Alaska, Alyeska Pipeline Service Company. (11) Consultant to DeLeuw Cather & Co. on blasting specifications for Washington, D.C. Subway. (12) Stability of cpen pit mine slope, Climax Molybdenum Co, Climax, Co. (13) Consultant to British Columbia Hydro on the efects of a new reservoir on the stability of Downie Slied (1 billion cubic meter slide). (14) Consultant to Gibbs & Hill on a slope adjacent to the Ohio River near Pittsburgh for sludge pipeline construction, slope 500-ft high. (15) Consultant on effect of blasting on stability of slopes of Caue Mine, Itibira, Brazil, slope 800-ft high. (16) Consultant to HydroQuebec on underwater rock blasting upstream on Manic
#5 buttress arch dam.
(17) Contract study for Corps of Engineers to re-evaluate Vaient Slide failure. (18) Consultant, R&M Consultants, Pillar Mountain Slide, Kodiak City, Alaska. EXPERIENCE ON DESIGN OF PROTECTIVE STRUCTURES AND NUCLEAR EFFECTS
- (1) Consultant to TRW Systems, Redondo Beach, Ca., on dynamic soil properties pertinent to the hardnesss of the Minuteman System.
(2) Member of a panel in Dept. of Defense to review design of all safeguard structures for vulnerability and hardnesn
A'Ifred J. Hencron, Jr.
. Page 7 (3)~ Consultant to Omaha District Corps of Engineering on the construction of underground protective structures in rock.
;(4). Consultant to Air Force Space and.Missle Systems Organization of
~ Hardness of Minuteman Structures as an associate of Dr. N.M Newmark.
(5) ' Consultant on problems in soil dynamics and rock mechanics to the U.S. Army _ Engineer _ Waterways Experiment Station, Vicksburg, Miss. (6) ' A member of the "Decoupling Advisory Group," formed by the Defense Atomic Support-Agency. Responsibility is to comment on stability problems which might be encountered in building underground cavities 100-360 ft in diameter and to give shear strength properties of rock masses which are importarit in determining the decoupling characteristics of cavities overdriven by the detonation of a nuclear device. I-(7) -Received Army Commendation Medal in 1965 for representing the Chief of the Corps of Engineers as a consultant to the Norwegian Government and NATO on the engineering of large underground facilities. , EXAMPLES OF UNDERGROUND - CONSTRUCTION EXPERIENCE (1) Consultant to British Columbia Hydro on aspects of Portage Mountain - Underground Powerhouse. 1 (2) Consultant to Duke Power Co., design of Bad Creek Underground Powerhouse. (3) Consultant to ENEE, Honduras, El Cajon Underground Powerhouse.
-(4) Consultant to ENEE, H'onduras, El Nispero Pressure Tunnel.
(5) Consultant to Dominican Republic, Pressure Tunnel, Tavera Project. (6) Consultant to Howard-Needles, Tammen & Sergandoff, Mt. Baker Ridge Highway Tunnel, Seattle, Wa. (7) Consultant to Dames and Moore, City of San Francisco Sewer Tunnel, Fisherman's Wharf. (8) Consultant to City of Rockford,11., sewer tunnel in sands below the water table. (9) Consultant to Metcalf & Eddy, design of sewer tunnel lining, Sao Paulo, Brazil. (10) . Consultant to Port of New York Authority, N.Y.C.: Stability of Lincoln Tunnels Stability of PATH Tubes ' 4 1 ,
. +.
Alf" a J. Hendron, Jr. Pc;e 3 _ LIST OF PUBLICATIONS Hendron, A. J., Jr., "The Behavior of Sand in One-Dimensional Compression," Ph.D.1963. July, thesis, University of Illinois, Department of Civil Engineering, Hendron, A. J., Jr. and T. E. Kenedy, "The Dynamic Stress-Strain Relations for a Sand as Deduced by Studying its Shock Wave Propagation Characteristics in a Laboratory Vol. II, West Point, Device," N.Y., Proceedings of the 1964 Army Science Symposium, June, 1964. Hendron, A. J., Jr., and M. T. Davisson, " Static and Dynamic Constrained Moduli of Frenchman Flat Soils," Proceedings of the Symposium on Soil-Structure Interaction, University of Arizona, Tucson, Az, September, 1964. Hendron, A. J., Jr., G. B. Clark, and J. N. Strange, " Damage to Model Tunnels Resulting from an Explosively-Produced Impulse," U.S. Army Engineer Waterways Report 1 May Experiment 1965. Station, Vicksburg, Miss., Research Report No. 1-6, Hendron, A. J. , Jr., D. U. Deere, - F. D. Patton, and E. J. Cording, "The Design of Surface Construction in Rock," Chap. Rock, American Institute of Mining, Metallurgical and PetroleumII in Failure and Engineer, 1967. Hendron, A. J., Jr., and H. E. Auld, "The Effect of Soil Properties on the Attenuation of Air Blast-Induced Ground Motions," pp. 29-47, Proceedings of the International Symposium on Wave Propagation and Dynamic Froperties of Earth Materials, University of New Mexico Press,1968. Hendron, A. J., Jr., " Mechanical Properties of Rock," Chap. 2, pp. 21-53, of Rock Mechanics in Engineering Practice, edited by K. G. Stagg and O. C. Zienkiewicz, John Wiley & Sons, London, 1968, 442 pp. Hendron, A. J., Jr., and N. Ambraseys, " Dynamic Behavior of Rock Masses," Chap. 7, Rock Mechanics in Engineering Practice, edited by K. G. Stagg andpp. 442 0. C. Zienkiewicz, pp. 203-326, John Wiley and Sons, London, 1968, I Hendron, A. J., l Jr., and H. E. Auld, " Discussion of Wave Propagation in Earth i Materials," Proccedings, International Symposium on Wave Propagation and Dynamic Properties of1968. Earth Materials, University of New Mexico Press, Albuquerque, pp. 94-98, l Hendron, A. J. , Jr. , J. C. Gamble, and G. Way, " Foundation Exploration for l Proceedings of the Twentieth Annual Highway Geology S University 1969. of Illinois, Engineering Experiment Station, Urbana, 126 pp.,
Alfred J. Hendron, Jr. Page 9 Hendron, A. J., Jr., and R. E. Heuer, "Geomechanical Model Study of the Behavior of Underground Openings in Rock Subjected to Static Loads: Report 1, Development of Modeling Techniques," WES Contract Report N69-1 Rept. 1, October, 1969. Hendron, A. J., Jr., G. Mesri, J. C. damble and G. Way, " Compressibility Characteristics of Shales Measured by Laboratory and In Situ Tets," pp. 137-153, ASTM Special Technical Publication 477, Determination of the In Situ Modulus of Deformation of Rock, June, 1970. Hendron, A. J., Jr., E. J. Cording, and D. U. Deere, " Rock Engineering for Underground Caverns," Proccedings, ASCE Symposium on the Design of Large Underground Openings, Phoenix, Az, February, 1971. Hendron, A. J., Jr., E. J. Cording, and A. K. Aiyer, " Analytical and Graphical Methods for the Analysis of Slopes in Rock Masses," NCG Technical Report No. 36 prepared for the U.S. Army Engineer Waterways Experiment Station, Vicksburg, Miss. , July,1971. Hendron, A. J., Jr., R. B. Peck, and B. Mohraz, " State of the Art of Soft-Ground Tunneling," Proceedings of the 1st North American Rapid Excavation and Tunneling Conference, Chicago, 11., June 5-7, 1972, AIME, pp. 1585-1610. Hendron, A. J., Jr., and L. L. Oriard, " Specifications for Controlled Blasting in Civil Engineering Projects," Proceedings of the 1st North American Rapid Excavation and Tunneling Conference, Chicago,11., June 5-7,1972, AIME, pp. 1685-1610. Hendron, A. J., Jr., and M. Amin, " Earthquake Resistance of Earth and Rock-fill Dams: Feasibility of Simulating Earthquake Effects on Earth and Rock-Fill Dams Using Underground Nuclear Events, Report 3, Appendix B," Misc. Paper S-71-17, September,1972, Waterways Experiment Station, 258 pp. Hendron, A. J., Jr., and W.M.C. Emerson, " Measurement of Stress and Strain During One-Dimensional Compression of Large Compacted Soil and Rockfill Specimens," Technical Report, Waterways Experiment Station, 241 pp. Hendron, A. J., Jr., and R. E. Heuer, "Geomechanical Model Study of the Behavior of Underground Openings in Rock Subjected to Static Loads: Report 2, Tests on Unlined Openings in Intact Rock," Waterways Experi-ment Station Contract Report N69-1, Rept. 2, February, 1971. Patton, F. D., and A. J. Hendron, Jr., " General Report on ' Mass Movements'," 2nd International Congress of the International Congress of the Inter-national As::ociation of Engineering Geology,18-24 August,1974, Sao Paulo, Brazil. Hendron, A. J., and C. H. Dowding, " Ground and Structural Response Due to Blasting," Advances in Rock Mechanics, Vol. II, Part 8, Proceedings of the Third Congress of the International Society for Rock Mechanics, Denver, Co., September 1-7, 1974, pp. 1359-1364.
Alfred J. Hendron, Jr...
'Page 10 Hall, W. J.,
N. M. Newmark,.and A. J. Hen'dron, Jr., " Classification, Engineer-ing Properties.and Field Exploration of Soils, Intact Rock and In Situ Rock Masses," U.S. Atomic Energy Commission, Washington, D.C., Report WASH 1301, 256 pp, 1974. 1
- Semple, R. M.,
i A. J. Hendron, Jr., and G. Mesri, "The Effect of Time-Dependent Properties of Altered Rock on Tunnel Support Requirements, Proceedings, 2nd Rapid Excavation and Tunneling Conference, Vol. 2, pp.1371-1373, 1974. Paul, S. L., C. E. Kesler, A. J. Hendron, Jr.,,et al., "Research to Improve
-Tunnel Support Systems," Report No. FRA ORD & D 74-51, Department of-Transportation, Federal Railroad Administration, Washington, D.C., 285 pp . , 1974. i 1
Cording, E. J., A. J. Hendron, Jr., H. H. MacPherson, W. H. Hansmire, R. A. . ~ Jones, J. W. Mahar, and T. D. O'Rourke-(1975) " Methods for Geotechnical Observations & Instrumentation in Tunneling," UILU-ENG 75 2022.- Mohraz,.B., A. J. Hendron, Jr., Randall E. Ranken, and Mohammad H. Salem, '
" Liner-Medium Interaction in Tunnels," Journal of the Construction Division, ASCE, Vol. 101, No.' C01, March, 1975, pp. 127-141.
Tarkoy, P. J., and A. J. Hendron, Jr., " Rock Hardness Index Properties and Geotechnical Parameters for Predicting Tunnel Boring Machine Performance," prepared for National Science Foundation under research grant GI-36468, University of Illinois, Urbana, II., September, 1975, 325 pp. Cording, E. J. , A. J. Hendron, Jr. , W. H. . Hansmire, J. W. Mahar, H. H. MacPherson, R. A. Jones, and T. D. O'Rourke, " Methods for Geotechnical Observations and Instrumentation in Tunneling,' University of Illinois i Report No. ULIU-ENG 75 2022, for the National Science Foundation under grant no. GI-3364X, December,1975, - Vols.1 and 2.
. Hendrca, A. J.', Jr., " Dynamic Stability of Rock Slopes," Earthquake Engineering and Landslides, Taipei, Taiwan, Republic of China, 29 Aug.-
Sept. 11, 1977, 56 pp. Hendron, A. J., Jr., " Engineering of Rock Blasting on Civil Projects," Structural and Geotechnical Mechanics, A Volume Honoring Nathen M. New-mark, W. J. Hall, Editor, Prentice-Hall, Inc., Englewood Cliffs, N.J., 1977, pp. 242-277. 4 Ghaboussi, J., R. E. Ranken, and A. J. Hendron, Jr., " Time-Dependent Behavior of Solution Caverns in Salt," Journal of the Geotechnical Engineering Division, ASCE, October, 1981. i t s,- _ _ _ ,,,_ _ _ , __ , , , . . . ,. c,_.._.,_,_r. _ ,,, , , ,_m.. , ,,_ _ - . s. m. . .,m.._,. , .. __m._-
INDEX PROPERTY TEST RESULTS FROM FILL SAMPLES TAKEN AFTER SURCHARGE Table 1 Ytotal Yd Soil Boring Elev. pcf pcf w% Classification i CO E-9 625 135.9 119.5 13.7 CL-623 137.9 122.8 12.3 CL 620 139.3 125.5 11.0- CL 618 136.6 121.8 12.1 CL 617 139.2 124.5 11.7 CL 617 140.9 126.1 11.8 CL COE-9A 625 143.6 128.6 11.7 CL 620 141.5 127.3 11.2 CL 614 137.5 120.8 13.8 CL COE-10 628 133.0 127.8 4.1 627 112.2 108.2 3.7 SP 626 120.7 108.5 11.2 SP 620 128.4 106.4 20.7 620 140.3 124.8 12 - 618 129.0 109.7 i '. . 617 125.1 107.8 16.0 616 128.6 109.5 17.5 615 111.9 94.7 18.1 608 115.0 102.1 12.6 607 123.7 111.8 10.7 SP-SM 606 118.4 104.0 13.9 605.6 119.4 103.1 15.7 SP-SM COE-10A 623 137.2 123.5 11.1 CL-ML 622 141.7 127.5 11.2 CL 620 142.0 127.8 11.1 CL 620.1 133.0 118.4 12.3 CL
Tab.e 1. .(continuation) Ytotal Yd Soil . Boring Elev. pcf pcf w% Classification COE-ll 627 139.1 123.3 12.8 624 137.6 123.2 11.7 620 139.9 124.6 12.2 619 140.8 126.9 10.9 618 139.4- 123.9 12.5 617.4 139.9 126.7 10.4 612 116.2 103.4 12.3 SP 612 121.1 105.9 14.3 608 115.1 100.7 14.4 608 122.3 104.5 17.1 603 120.4 103.8 16.0 SM 602 123.1 103.9 18.5 COE-llA 621 138.6 125.4 10.5 CL 620 141.9 128.2 10.7 CL 619 137.9 123.3 11.8 CL 619 142.1 128.6 10.5 CL
'617 135.1 120.5 12.1 CL
4 Table 1 (continuation) Y total Yd Soil
, Boring Elev. pcf pcf w% Classification- ,
COE-12 627 137.5 123.2 11.6 627 136.5 120.1 13.7 CL 4 626 140.7 126.3 11.4 CL
! 626 140.6 '126.6 11.0 625 129.2 116.0 11.3 624 140.2 126.7 10.6 j 623 139.8 126.6 10.4 620 140.2 125.3 11.9 619 140.0 -
126.8 10.4-618 140.9 128.2 9.9 617 139.2 126.0 10.5 CL 616.8 142.5 129.1 10.4 CL
- 616 141.4 127.8 10.7 615 137.5 121.5 13.2 614 138.9 124.1 11.9 614 144.3 128.2 12.6 612 136.2 119.4 14.1 CL 611 136.5 119.9 13.9 CL 609 137.4 122.8 11.8 608 137.3 120.0 14.3 608 136.5 117.9 15.9 607 134.1 117.2 14.5 606 132.9 112.6 18.0 CL 604 115.4 97.7 18.1 CL 601 118.3 99.1 19.3 SP-SM 600 109.0 99.6 9.4 SP-SM l
i I { l i
Table 1 (continuation) Ytotal Y d. Soil Boring Elev. pcf pcf w% Classification COE-12A 624 139.2 125.2 11.2 .CL 622 139.0 123.9 12.2 CL 621 138.8 124.1 11.S-621 147.2- 135.2 8.9 617 141.5 128.9 9.7 CL 616.6 141.4 129.0 9.5 615 132.3- 112.1 18.0 CL 614.5 130.5 109.8 18.8 613 132.1 112.7. 17.2 CL 611 135.9 118.6 14.6 CL 608 135.0 118.0 14.5-608 120.2 13.7 136.6 606- 131.1 110.4 18.8- CL i. f
, ,r -
Table 1 (continuation) Ytotal Yd soil Boring Elev. pcf pcf w% Classification COE-13A 623 142.8 129.0 10.7 CL 620 142.2 128.9 10.3 CL 618 142.6 128.5 11.0 CL 615 136.5 121.7 12.1 CL 613 140.3 126.2 11.1 612.5 145.4 132.2 10.0 CL 611 144.9 130.6- 10.9 CL 609 141.3 127.6 10.7 CL 607 140.1 126.9 10.4 CL COE-13B 625 137.9 125.5 9.9 CL 625 142.7 128.9 10.7 CL 619.4 145.1 131.9 10.0 CL 619 142.2 129.8 9.6 CL-ML ] i 4 4 i
4 TABLE 2 TRIAXIAL ~ TESTS ON ISOTROPICALLY-CONSOLIDATED PLANT FILL SAMPLES DIESEL GENERATOR BUILDING - WOODWARD AND CLYDE (1981) Elev. Yd c Test Boring Sample DeptE. pcf w% ksf USCS
'9 1 COE-13B S-1-C 128.9 10.7 0.7 CL 2 COE-13A S-2-B 620; 128.9 10.3 1.38 CL 13 COE-13A 622.8
, 3 S-1-B 9 129.0 10.7 2.73 CL 4 COE-llA S-4-C 14.6 128.6 10.5 5.44 CL , 5 COE-13B F-3-C f*41 131.9 10.0 8.18 CL a b t s
-. - , ., . .,-r
..? ,i TABLE 3 Ii TRIAXIAL TESTS ON '
ANISOTROPICALLY CONSOLIDATED PLANT FILL SAMPLES U DIESEL GENERATOR BUILDING - WOODWARD AND CLYDE-(1981)
.q 0
Elev. Yd # 1 "3 2;
- Test Boring Sample Depth pcf w% ksf ksf 3 USCS lA COE-9A S-1-C g 128.6 11.7. 1.00 0.53 1.891 CL 4
2A COE-10A S-2B 0.6 3_ 127.8 11.1 2.05 1.15 1.79 CL 3A COE-10A S-1C 622.2 127.5 11.2 4.1 2.14 1.916 CL 4A COE-12A S-3B yy } . 123.9 12.2 8.04 4.2. 1.912 CL i SA COE-9A G1 .8 S-4B y 2 127.3 11.1 12.12 6.32 1.917 CL 6A COE-9~ 617 S-6D 6$8 126.1 11.8 16.0 8.43 1.90' CL k i
TABLE 4 TRIAXIAL TEST RESULTS DIESEL GENERATOR BUILDING WOODWARD AND CLYDE (1981) Consolidation Stress Y d 31 1~D max c T. b 2 f A ff Test ocf w% ksf ksf ksf % f ksf 1 128.9 10.7 0.7 0.7 1.87 20.1 -0.284 1.6 2 128.9 10.3 1.38 1.38 1.81 19.9 -0.099 1.55 3 129.0 10.7 2.73 2.73 2.44 19.7 0.078 2.1 4 128.6 10.5 5
. 44 5.44 5.84 2 0 .~ 0 -0.012 4.9 5 131.9 10.0 8.18 8.18 7.47 20.0 0.083 6.4 1A 128.6 11.7 1.0 0.53 1.86 20.1 -0.038 1.6 2A' 127.8 11.1 2.05 1.15 2.43 19.9 -0.314 2.1 3A 127.5 11.2 4.1 2.14 3.2 20.0 -0.254 2.8 4A 123.9 12.2 8.04 4.2 3.9' 20.0 0.121 3.3 5A 127.3 11.1 12.12 6.32 7.88 20.0 -0.106 6.8 i
6A 126.1 11.8 16.0 8.43 7.9? 16.1 0.010 6.8 4
, . , . - -- - - ~ . , , - . . - , ,. -
e . TABLE 5 4
SUMMARY
OF CALCULATED FACTORS OF SAFETY FOR COMBINED STATIC AND EARTHQUAKE LOADS Failure Case No. Surface Undrained Shear Strength Factor of' Safety Case I Deep - Circular From Anisotropically Consolidated Tests Only. Fig. 14 in Text 3.25 From Both Anisotropically and Isotropically Consolidated Tests. Fig. 13 in Text 2.86 Case II Shallow Failure From Anisotropically Surface Formed By Consolidated Tests Only A Triangular Wedge & Fig.-14 in Text 2.95 A Circular Section From Both Anisotropically and Isotropically Consolidated Tests. , Fig. 13 in Text 2.59 Case III -Shallow Failure From Anisotropically Surface Formed By Consolidated Tests Only A Triangular Wedge, Fig. 14 in Text 2.81 A Circular Section & A Passive Wedge From Both Anisotropically and Isotropically Consolidated Tests. Fig. 23 in Text 2.49
Table 5, continued Failure Case No. Surface Undrained Shear Strength Factor of Safety Case IV Intermediate Failure From Anisotropically Surface Formed By A Consolidated Tests Only. Triangular Uedge and Fig. 14 in Text 2.83 A Circular Section From Both Anisotropically and Isotropically Consolidated Tests. Fig. 13 in Text 2.48 i I Case V Intermediate Failure From Anisotropically Surface Formed By Consolidated Tests Only. Triangular Wedge, Fig. 14 in Text 2.67 A Circular Section and A Passive Wedge From Both Anisotrpically and Isotropically Consolidated Tests. Fig. 13 in Text 2.37
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COE 13A (13B) ( ) -Indicates A First Attempt To Do A Boring u NORTH Scale: 1in. = 50 ft Figure 11. Location Plan of Borings Taken After Surcharge - Diesel Generator Building a r-yw -----m- - - - - - - -
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, :e .
APPENDIX A EXPLANATION OF CALCULATION METHOD USED TO ' ASSESS BEARING CAPACITY OF FOOTINGS UNDER COMBINED STATIC AND EARTHQUAKE LOADING
' INTRODUCTION The ultimate bearing capacity of the D.G.B. Foundation under combined static and earthquake loading was. evaluated for different potential sliding surfaces; and, the factor o,f safety against failure, was determined for each of the potential failure surfaces considered. The sliding surface with the lowest factor of safety is the most probable failure surface and therefore determines the factor of safety of the footing. Three potential sliding-surfaces were considered in this analysis. These surfaces were a deep sliding surface, a surface with an intermediate-depth, and a shallow slidiny surface. The main characteristics of these sliding surfa?,es are shown in Figures A-1, A-2 and A-4 and described in-the following paragraphs:
a) Deep failure surface - a deep, circular sliding surface shown by the dotted line in Figure A-1 was assumed in this analysis; as indicated in Figure A-1 the radius of the surface, r, is equal to the width of the footing which is 10 ft. The center, 0, of this circular surface is lccated at the lower right-hand corner of the footing.
-A/2 s
b) Intermediate failure surface - an intermediate
, depth sliding; surface, shown byithe heavy line.
;ba Figure lA-2.was assumed,in this analysis.
~
As shown in' Figure A-2, the-intermediate sliding surface'is formed by a combination of a. straight and a circular section.. The center of the circle.is located at the lower right-hand corner of..the footing. A variation-of-this failure. surface was also considered in which the: outer. portion of the circular.section of the sliding surface ~is replaced by a passive
' wedge'as shown in Figure-A-3.
c) Shallow failure surface - a shallow failure surface, formed by a combination of a straight and a circular section,.was assumed in'this analysis. 1 The straight portion of the sliding' surface, as indicated in Figure A-4,.has~an inclination l of 45* with the bottom of the footing and ends ,
~a t~a point below the center of.the footing. From
[ this point on, the sliding. surface becomes cir - C. !'i cular; the center of this circle :Us located at I the lower right-hand corner of the footing. A variation of this sliding surface was also con-sidered in which the outer portion of'the circular i
~
k i
~
A/3-section of the sliding surface is replaced by a passiveLwedge as shown in Figure A-5. FACTOR OF SAFETY UNDER COMBINED STATIC AND-EARTHQUAKE LOADING Calculation of the_ factor of safety against a bearing capaci ty failure under the combined static and earthquake loading included the following steps: 1- Calculation of the long-term effective stresses along the potential sliding surface due to the over-burden and the permanant structural loads on the footing; for this purpose, the potential sliding mass was divided into the slices shown in Figure A-6. Effective normal stresses' E n, and shear stresses, T , acting on the potential sliding surface, were calculated on the base of each slice. 2- After the effective stress, E , n rmal to the base of n each slice was calculated, the corresponding undrained shear strength along the sliding surface at the base of each slice was determined from the relationships shown in Figures 13 and 14 in the main text. 3- Once the undrained shear strength available along each sliding surface was determined, calculations were carried out to estimate the ultimate load, pua that would induce failure along each potential sliding surface. This ultimate load, Pu, corresponds to the
A/4 maximum resistance of the soil along the potential failure surface; which develops from three main sources: a) The weight of the surcharge (the soil outside the footing and above the level of the base of the footing) b) the weight of the sliding mass below the base elevation of the footing, and c) the shear strength developed along the sliding surface. The contribution of each of these factors to the ultimate failure load, Pu'.was estimated separately from the equations establishing the rotational equillibrium of the sliding mass. Equilibrium requires that the summation of moments produced by each one of these forces around a previously selected point be equal to the moment produced by the failure load, Pu, around the same point. The steps above are explained in detail in the following para-graphs. CALCULATION OF LONG-TERM EFFECTIVE STRESS ON POTENTIAL SLIDING SURFACES The method of slices was used to estimate the effective stresses acting along each potential sliding surface. The method is a trial and error procedure in which a uniform
" mobilized" friction angle is assumed to develop at the bottom
~
A/5 of each slice. This " mobilized"'-friction angle is adjusted until equilibrium of all forces ~ acting on each slice is achieved. Side forces between slices were assumed to'act at an angle given by 'the average value of the inclination of.the failure
. surface'and the ground surface at'the contact between the slices. Pore water pressures were accounted for by using bouyant weights for.the portion of slices below the water table in the polygon of forces acting on the slice. A typical trial and error construction of the quilibrium force polygons on each slice is-shown in Figure A-7 As indicated in this figure, equilibrium of all forces acting on each slice resulted in a " mobilized" friction angle of 19 along the potential failure surface analyzed.
4 The average, " effective" normal and shear stresses at the base of each slice, J n andTn,: respectively, were obtained from the equilibrium force polygon, an example of which is shown in Figure A-7. The stresses were calculated as: R cos e o = m
^ Eq. (A-1)
LC l and R sin 0 T n
= m Eq. (A-2) where R is the magnitude of the reaction force at the base of-the slice, obtained from the equilibrium force polygons:
A/6 e, is the mobilized friction angle (along the potential sliding. surface) required to achieve equilibrium of all forces acting on each slice;'and L is the length of the chord subtending the c bottom of the slice. The magnitude of the principal stresses E lc and E 3c along the failure surface was calculated assuming that the angle $ between the plane of. principal stress and the sliding surface is independent of the factor of safety. Under this condition, principal stresses E lc and F 3c an be calculated as: 5
=En+'n tan & Eq. (A-3) 0 3c " A n -T n/ tan $ Eq. (A-4) where $ is defined as:
$ =w +p' Eq. (A-5) 4 2 and d' = effective angle of shearing resistance.
Calculation of Ultimate Load, P u Ultimate failure loads, P u, for the various failure surfaces chosen in this analysis were calculated as follows:
-Deep Failure Surface The. ultimate load, Pu, required to induce a bearing capacity failure along this surface was determined from the equation establishing the rotational equilibrium of the sliding
A/7: mass. Equilibrium requires, in this particular case, that the sumr.iation of mortents about the center, 0, of'the circular sliding surface be equal to zero,. Figure A-1: IM g = 0 Eq. (A-6) Because of the circular geometry of the sliding surface, the moments around point 0 induced by the weight of the slices cancel out,'and_the contribution of the two other factors, shear strength and surcharge, can be estimated in a single equation as: r+ 3 Pu o T Pd ii Eq. (A-7)
.I 1=1 ffl
, g.
1 ,El 1 wher P u. = Ultimate load required to induce a bearing capacity failure, L = length of arm between the center of the g footing and the center of the sliding surface, T ffi = maximum undrained shear strength available along the bottom of slice i, 1 = length of the bottom of slice i, 1 r = radius of the sliding surface, P i
= surcharge load on slice i, di = length of arm between surcharge load P g and the center of the sliding surface.
-Intermediate and Shallow Surfaces with No Passive Wedge The ultimate load, Pu, required to induce a bearing
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.A/8 l
capacity failure along the shallow or the' intermediate F ' sliding; surface with-.no passive-wedge was determined.by dividing- ~
-the. sliding mass at_the. contact between the circular'and-straight portions;of the? sliding surface, line 00' in Figures A-2 and A-4; and establishing the equilibrium of each of these portions. separately. The rotational equilibrium of-the c'ircular portion of the7 sliding mass was established by making the moments around the center, 0, of the circular surface .(Fig.
A-2 and A-4) equal to-zero. The equation representing this ' condition was established'as: IM g
= 0 Eq. (A-8)
The contribution of each one of the resistance factors to the moment around this point was calculated'as: Contribution of shear strength along sliding surface: P' m g =- fT ffi ti ri Eq. (A-9) where P' = resultant normal force (due to shear strength) acting on-the surface 00' that divides the circular portion of the sliding mass.from the wedge underneath the footing (Figures A-2 and A-4), m = length of arm between-the point of application o of force Pf and tne-center of the sliding' surface, in this case one-half the distance between 0 and~0' (Figures ~A-2 and A-4),
. rf f i . = maximum undrained shear strength available-along the bottom of slice i, t
., _~r + i_.
A/9 1 = length of the base of slice i, 1 r1 = radius of the circular. portion of the sliding surface. Contribution of surcharge: Py m g =.if'P1 i di .Eq. (A-10) where PE = resultant normal force (due to surcharge) acting on 00' (Figures A-2 and A-4) mg = as defined'above for contribution of shear strength P = surcharge load cn1 slice i i di = length of arm between surcharge load P i and the center of the sliding surface. Contribution of weight of the sliding mass: m P'"mi = 131 W i ng Eq. (A-ll) where P"' =
-resultant normal force (due to weight of sliding mass) acting on 00' (Figures A-2 and A-4)
=
mi length of arm between point of application of force P"'and the center of the sliding , surface,Ein this case two-thirds the distance 00', measured from the_ center of the sliding surface W. = weight of slice i included in the circualr 1 portion of the sliding mass i I i
. .a.
- A/10-
= length'of arm between weight of slice i-n1 and the center of the' sliding surface.
Once the magnitude of P',.P", and F" are known, the equilibrium of the remaining portion of_the sliding mass (" triangular wedge") below the footing, Figure A-8a) is established by making the summation of the forces acting on.the triangular wedge equal to zero, (force poli gon in Figure A-8b). The force polygon in Figure A-8b also includes-the shear strength developed along the straight portion of the sliding surface, t ff6 1 6, as well as the weight of'the soil wedge beneath the footing, w.
-Intermediate and Shallow Surfaces With Passive Wedges As shown in Figures A-3 and A-5 for the alternative
! slip mechanism considered for the intermediate and shallow sliding surfaces, a portion of the circular section of the !. sliding surface was replTced by a passive wedge. The sliding mass was divided at the contact between the straight section beneath the footing and the subsequent circular section, line 00' in Figures A-2 and A-4; the equilibrium of each of these , portions of the sliding mass was established separately as described above. The rotational equilibrium of the circular portion of the sliding mass was established by making the moments around the~ center, 0, of the circular surface, equal to zero. However,
A/ll due to the presence of the passive wedge, the contribution of each one of the resistance factors was' calculated as follows. Contribution of shear strength along sliding sur-face: As shown in Figures A-3 and A-5, the outer portion of the circular sliding surface at the base of slice 1
. was replaced by the straight section bounding the passive wedge. The contribution of the_ passive wedge was cal-culated as the passive resistance of the fill materials along the vertical plane A-B in Figures'A-3 and A-5. This passive resistance is given by the trapezoidal stress distribution on plane A-B shown in Figure A-9. The passive resistance o gp ando Bp at points A & B respectively is given by:
{ o gp
= 2 (T ffg) =q Eq. (A-12) o 2 (T ffB) +
=
BP ovB Eq. (A-13) where q = vertical affective confining stress at point A, which in this case is equal to the stress Y imposed by the surcharge. - T = undrained shear strength at point A, ffg which is a function of the effective confining stress q, c = vertical effective confining stress at ovB point B, and T ffB
= undrained shear strength at point B, which is a function of the effective confining stress ovB-
f A/12 The resisting force on plane A-B resulting from the
. passive stress distribution described above was divided in two. resultant forces, F1 and F2, shown in Figure A-9, respectively.
~
F1 corresponds to the. rectangular portion of the stress distribution and is located midway between A-B at a distance Hy -from the bottom of the fotting; F2 corresponds to the triangular portion of the stress distribution'and is located at a distance H 2 from the bottom of the footing equal to 2/3 the distance between points A & B. The contribution of' shear strength along the sliding surface in maintaining rotational equilibrium of the cir-cular portion of-the sliding mass is now given by: Pf m0 if2 Tffi.Ei r + F H11+FH22 Eq. (A-14) where the terms, Pf, mo , T fff, ti, and ri remain as pre-viously defined. Note that the contribution of shear strength along-the straight portion of the sliding surface. which bounds the passive wedge is accounted for in Eq. A-14 by the moments F H1 1.and F H22 If n Passive wedge is assumed, the contribution of the shear strength along the outer portion of the sliding surface is given by the moment, developed by the shear strength along the base of slice 1, T ffy 1 1 ri, Eq. A-9. s
- z. . . . .
. t A/13 contribution of-surcharge: The contribution of the surcharge' located outside of the vertical plance A-B in Figures A-2 and A-3 has been accounted-for'in the estimate of the passive resistance along this plane.
.Ther2 fore the direct contribution of the surcharge to-
,the rotational equilibrium of the circular portion of the
~
sliding mass reduced to P" mg = i2Pdii Eq. (A-15) where P", m, a Pi, and di remain as defined-above. Contribution of weight: The contribution of the weight of the sliding mass located on the right-hand of plane A-B was accounted for in the estimate of the passive resistance along this plane (the effective con-fining stress at point B is determined by the weight of this mass). Therefore the resisting moment produced by the weight of the sliding mass is given'by: P'" m i = if2 i "i Eq. (A-16) where f' and m i remain as defined above for Eq. A-ll; however, the term Wi ni includes only the weight of the-circular portion of the sliding mass. Numerical calculations of the ultimate load, P u' required to induce a bearing capacity failure for all the poten-tial failure surfaces considered in this analysis are given in Appendix C. p y = g <
i e c il di Sdn e Pi hu t o r 3 yA B p
+
r dU i e f cd _ i ue u dc S on ro 6 Pl o sB 0 : t nt 0
= a eh t o f mign M
oe io u MWP I P ui S P U N N r
,I ill ! I i
P 2 N d ,i U
' s I ,ll li I N
d 2 3_ d_ P 3 U f f f O % C\_ c f r e a u
,l .I l S
= -
g n i A h1 d i l S r a _ l _ u u _ P ;- c r i C p e e D s A . 1 A e r u g
/ i F
1 I
4 I I l l Pi 2 P e P 3 c 3 - f a r a u S
- l g n
[ 4 i d o e /
% i l
S S e P o / t a P u .' [ i d
- e U m
'O r e
t n I 2 A e r u g i F 8 7 6 5 4 3 2 1 2 2 2 2 2 2 2 2 6 0 6 6 6 6 6 6
o I l* l { 4 4 Pi 4 e 4 g d 4 e 4 W e 4 I v i I s A O s u a 2 f P P h { 2 t i 4 w o e c P 3
=
{ 4 1 0
/
ef S a r u g n b /
! 4
- i d
i
/ i S
o e / w S
! t e
'P / a
/ i d
P u
=
! / e m
W 'O
/
I r e t n 3 A e _ r _ u g i F 8 7 6 5 4 3 2 1 2 2 2 2 2 2 2 2 6 6 6 6 6 6 6 6 l ,
s ' ' 0 v d
- I l l li " ui S
P b
,g 9 ,'
_- A h
* . Il g l e2 di i.l
,l i l l i j
/
2
- 3 d w d
3
}
3 n
,l #Ja e
c I! g _~ i-!' o ; a f
'l 4 r u
/ lf S
.,/w
/
/
5 A% i d g n P i
/ l P
u y \
,P
/
'O
\ S l
l w o a h S 4 A _ e r u g i F . 8 2 6
\ 11 { lt i l l ill llI _
[ Pi e g A A C 0 d
.l l ll g I l
' e W
P 2 N e di >l
=
l} 2 \ v
\ i s
ll ! i l l [ Il l l N s a
/
2 3 \ P d l{ P g h
" 3 3 t d
O g i W ll I rt ~ i
/ e h / 4 c
a A P o
/ ' tfS r
u g
/5 i n
; 'p / d P
u / i l S p\ w o l l a h S 5 g A e r u g i
- - - ~ - _ - - - F 4 8 3 3 2 2 6 6 6 \ll}l
oy
< 1 w
l s e c i
, l 2 4 S
- n 4 i d
e
. 3 / d i
i v O D s s 4 a
\
M g n i d i 5 l S v l a i t n 6 e _ t o _ P a 6 A e y r
, u
_ g i _ F 4 8 3 3 0 8 2 2 2 1 _ 6 6 6 6 6
Ri I R2 2 o R3 3 4 Force Between Slices 3 and 4 ] Total Weight On Top of Slice 4 Reaction Force At Base 4 R of Slice 4 u 5 R5 v I, Jb 5 kips ( 6 l t i Rs 6m 53 i I p Figure A-7. Force Polygon - Shallow Failure Surface
s w f 4 Pu l l'
,0 a /
/ Pu Tff6 6 UW6w / Tff6 A 6
, /
/ Po.
/ a "
PE p p, p pn. Ww Susf s n.
- y '
N'o a) Forces Acting on Triangular b) Equilibrium Force Polygon Wedge Below Footing for Forces on Triangular j Wedge Below Footing i Figure A-8
v 2 H u P Ha g_ e y u g
=
j_ 2 P d p - B e e, +{ # W e v A i i1 ; l ! ll l l s
\ s 2 \ a P P d' ,
y 2 \
\ h t
il ; l I l ! l l 's i w d 2 3 \ e
. I!!
= _
l d l} 3_
~
l P O n
\3 sfS c
a r u M wi g 4 n
/ i
/ ,
d l P / S
/5 w P / o u / l P
l U a
\ h S
9 A e r u g i F 4 8 3 3 2 2 6 6 6
,: V ~
APPENDIX B SHEAR STRENGTH OF FOUNDATION MATERIALS BEFORE SURCHARGING Shear strength parameters for the foundation materials under the Diesel Generator Building were evaluated from results obtained from nineteen (isotropically consolidated, undrained) triaxial tests carried out on fill samples taken at various plant locations before application of the surcharge. Results of these tests are contained in Response to NRC Request Regarding Plant Fill, 50.54 (f) . The locations of the borings from which the triaxial samples were obtained are shown in Figure B-1. Pertinent sample information regarding boring location, depth, initial dry unit weight, Yd, water content, w, and initial consolidation pressures used in the triaxial tests are given in Table B-l. Test results, including maximum principal stress difference (0 -o 1 3) max at failure, undrained shear strength, su" ( 1- 3 max and strain at failure, eg, are given in Table B-2. The pore pressure coefficient, Xf , at failure, is also given in Table B-2 where A f is defined' as: Ag = "s Eq. (B-1) I 1- 3) max where Ap hange in pore water pressure at failure. s
g .: J
., . s .
~
' B/2 ..
" ~
'ShearLstrengthl parameters for the: materials'under-
- . ' neath the' Diesel Generator Building foundations were estimated r -
~
by': il)L Determining representative values ofLthe dry density, of;the materials underneath:the Diesel Generator i; 4- Building, and 1-t 2)~ Using-shear strength parameters, both drained and undrained, obtained from triaxial tests-carried x out on samples of similar dry densities. b i
' Dry Density of Fill Materials and Corr'esponding Triaxial Tests I
[
-Representative values of the'in-situ dry density-of r
f the materials underneath the DieselIGenerator Building were-obtained from samples taken.at various locations and depths within the fill underneath the building. The' location of j the different borings where these samples were taken.is-shown I 4 1 in Figure B-2. Pertinent sample information regarding depth, total and dry unit-weights, water content, Atterberg: Limits, grain size and soil classifications are given in Table B-3.
~
j; . Representative values of-the in-situ dry density and corresponding. i 4 water content within the ten feet of fill immediately beneath ^ 1 the Diesel Generator Building foundations were obtained'from i .
- the information given in-Table B-3. It was determined that the.
1 i
. in-situ dry-density of the fill decreased from an average value of about 122 pcf between elevations 628 and 623 to an average value of about 116 pcf between elevations 623 to 618.
h 4 I i ~.,.,- ,
.- 4 ..-.a---._.,__.-___. , . _ ~ _ _ , . , , , . . - . - _ . _ . . . . . _ _ , - . _ . . . . . , . _ _ , . _ . - . _ ..L..
, ..=-
-D/3.
Corresponding water contents increased from an average of about 10% to 12% in the upper portions of the fill between elevations 628 and 623,.up to about 17% to 19% at elevation. 618. Triaxial test samples from various plant locations-with dry unit' weights similar.to those measured under the Diesel Generator Building were considered to be representa -
'tive of fill under this building. These samples were divided into two groups:
Group I - corresponds to those samples with average dry densities of 122 pcf considered to be representative. of the five' feet thick fill layer immediately underneath the footings (between elevations 628 and -623) . Seven samples from three different borings, T-12, TR-4 and CT-6, are included in this group. Pertinent sample characteris tics (depth, in-situ dry density, water content, etc.) as well as measured shear strength values are given in Table B-4. Group II - corresponds to those samples with average dry densities of 116 pcf representative of the five feet of fill between elevations 623 and 618. Five samples from five different borings, T-16, T-9, TR-2, T-15 and TR-5, are included in this croup. Pertinent sample character-istics as well as measured shear strength values are given in Table B-5. 1 1
~ . a B/4 Effective Shear Strength Parameters Effective shear strength parameters, c' and d', for the materials under the Diesel Generator Building were obtained from the results of undrained triaxial tests carried out on the fill samples-in Groups-I and II. Pore pressure measure-ments taken during these tests.made it possible to calculate the principal effective stresses, 3 1
and E 3'" "II"#** Values of q and p, where q = 71 - 3 2 and p = 1+ 3 2 at failure (Tables B-4 and B-5) were plotted separately for each group of samples on the modified Mohr-Coulomb diagrams shown in Figures B-3 and B-4. These diagrams indicate that the samples from both groups exihibit aln.ost identical effec-tive-strength parameters of c' = 0 and p' = 30*. Undrained Shear Strength Undrained shear strengths representative of the fill materials under the Diesel Generator Building were also obtained from the triaxial test results of the samples in Groups I and II. Maximum undrained shear strength values, 1 (o -o
- 3) max, 2 1 measured in these tests have been plotted versus the corresponding 4
isotropic effective confining stress in the sample prior to s y_ -
- e.
- B/5 shearing. Two distinct relationships, shown in Figure B-5,.
resulted.from these. plots. Scaples with the higher dry densities of '22 pcf, (dotted line in Figure B-5) showed, for all values af effective confining stress, undrained shear strengths substantially higher than those for the samples with the lower, 116 pcf, dry densities. Undrained shear strengths of the materials under .the Diesel Generator Building, during an earthquake event, were obtained from the relationships in Figure B-5. Undrained shear' strengths along the failure. surface, T ff, as defined by Lowe and Karafiath (1959), obtained from these triaxial test results were also plotted versus the corresponding isotropic effective confining stress prior to shearing; these relationships are shown'in Figure B-6.
* \ g y
i
.,oo ..f f 1
J I L - u, ... .. g . tu
-.. - . ;wr. (.S/ ..
. ,M."'
U .:. r.. a '*' . g.. ".t* , _'T, * - * "* ;; .,,, E '
- I", ,. w *?
s.us =
~
,.G s.. . ~ E: _.. -
somcs w#eee ,.4
~f 4' ,..
Tgf u gy)L res7* !- *m "M.. .,
*"* CD e.
,,.,.) ,,,p;, .~. '#, , , , , . . .
- . . . . g. ,,
SAr1fLES 4082E *" S ' ' ' '
*b., ..-.
A r e'd . '..
.. j :: ***Y '.: ":'. t *" w '"*
- h. "~
. m., ~
** e ..?. .,."
Q m
.~-
s g
, m f -
.. . - - ~. .-.
,r, f. ..
,. ,, 7
.- 5, . ..... ) * * . A . ..
n v... ,. .- .. v - . . . . , .... . , ... r -..*zCAW. 8C.
,. r* ~.~' @'r,M
.* ,. ." , *,., gm,.. .=... . -. 4;- ,
=
f,.. ,4g * =.c=
.- . .d
.-"",.g*.-- , .,
e" s' , n (~ ,. r s a. r* (* <r* .** <
-. c . + . , - * . , . . . . .t
. . .- . ,. g w-
,,oo _ . . . ..~ . g ...~ T 1 ,
~. . l
.. g. ,,,,-**-- ,.. . p.Y ,
3""~' h
-...'"*#~ .h.
- a
% *.9 $..: ' g J....
- r.m. t
,. _pp %X ~,+"' -.- ..
y* * . . L., 3
.. 7
**J.- .- ::[ a c'
.3,oo .
r .n ... r,53,, m -- xg - 1- . . . .< ..
" . . m..
' i
. . - ~ . ,
M IIIIII IIIlIlIII I I I I I I I I II'N T A I N Ar EN TI
$ O PD !
5 SWL-4 6 ( . Ij 4 15 0 A X- 9 AX- 18 TW-5 AX-12 ~ AX-5 AX~4 dOTW-4
/
% A X-15 M6 AX-19 o O 3x_30 O ow-5 16 t.r n . ggg AX-an tow-s Y TEW-3" 'I4900 0 A
"TEw-7 -o D o TEW-I C - 233
,! TEW-0 9 TP-8 TEW-2
' Ow-14 G
/ TEW-5 g4} O LOW-7
- LOW - 13 i TURBINE BLOG, D-19
$ D-16 I$ D~EO h Q - 10 D-37 LOW-9.g 4
S-1 6 C O-i2 b g-g S-2 P-13 P -14 OPD P -il D-2 i 4 -o- %p-12 0 4 A -17 8 17 , A-17d M..C-23 4 m 9 y D-47
$ fr D-54 P -19 -20 C-P -17 DG-28 P-18 ODG-29 0G-{y.Q-
,, DG-573 0 D G-31 % D-3 9 TEST 355
~
D )l ' l CG-Ikhc
'fl'l4 edD,-4gG-9 8 eDG-20 eDG-8e e DG-14 e O
^
v TR-2 O P-22-n TR-5
- DG-IO DG-13 TR-l . .TR-3 Oq ,g
,,CH-IS DG-21 DG-234 DG-22 D-45 q!D-43 p-56 D- e D-44 OPD-13 PD-38 5 r DG-15 ;$9 4 ,42DG-17 0 -12 th UNIT 2 4 PDlBO a TRANSFORMERS DG-26g DG-t6
'D-43 DG-il4DG-24D CH-16 N *DG-7
~
C H-l7'C H-18 0-52 I # 0-58 -- O Cl, D-50 0-51 t$- D-59 0-48 e D-60 80 0 I 6 PD-16O DG 29 DG-3 0-2 DG-5 00-30 DG-6 DG-4 O
. S WL'I CL b Fig tre B-2. Plan view DG g5 0 CHLORINATION PD-19 showing location of O D PD-20C BUILDING CT c- borings at the PD-208 hPD-20A n. hIr- G Diesel Generator Buildin9 PD-20 [ g.T - 6 .. -
O f ,
}
n 6000 - Samples With Average yd ss l16 pcf 5000 - c' = 0 psf 4' = sin-'(tang a) = 30* T-15 / TR-5 / + 4000 - Mo. ro a = 27' b 3000 -
+_
N b
+ TR-2
! g 2000 - T-16
- +
T-9[ 1000 - I I I I I ! I l O - O 1000 2000 3000 4000 5000 6000 7000 8000
~
el+a3 , psf P= 2 Figure B-3. Modified Mohr-Coulomb Diagram - Samples with In-Situ y d 116 pcf
d 4 n 12,000 - , Samp;es With 1 Average yd a 122 pcf
'U # -
c' = 0 psf 4'= sin-I(tanga) = 30* EQ.
~
- CT-6 TR-4
^
m a = 27* l6 i N 6000 - i gg CT-6 1 Il CT-6 cr 4000 - TR-4 , 2000 - T-12j + TR-4 l l l I I t O I I I I o 2000 4000 6000 8000 10,000 p 12,000 14[XX) lis/XX) P* WI2+05 ~ ' E8I n
) Figure B-4.
Modified Mohr-Coulomb Diagram - Samples with In-situ y d 122 pcf
s TR- 4+ + CT-6 / l 7000 - y i
/
yd a 122 pcf / j
/
6000 -
/
/ yd a ll6 pcf i CT-6 + CT-6 /
+ /
/
5000 T-15 y/
/'
/e TR-5
/
4000 - . / u, f a- + 7 3 y 2 3 cog _ y
/
TR-2/ + Samples With Average y
/ yd a 122 pcf
- 2000 +TR-4 y T-12+ e Samples With Average
- e T-16 /
j / yd a l16 pcf +- / e T-9 1 /
/
1000 - p
/
/
/
, g i I I I I I I I I 2000 3000 4000 O 1000 5000 6000 7000 8000 9000 l Isotropic, Consolidation Stress, c-j Figure B-5. Maximum Undrained Shear Strength Versus Effective Consolidation Stress from Triaxial Test on Fill Samples Taken Before Surcharge
d n 7000 - yd a l22 pcf
+ ,
/
f sooo - / f {
/
/
- a. / .
/
. 5000 -
/
V + / -
- , + f yd a l16 pcf i .c /
a 5 /
]
E 4000 - - j/ 5 /*
\ u /
o l
. ,/ .. .
m 3000 -
+
/
.n /
8 /
'6 /
-ti
! $ 2000 -
+ l
>/
+ e /
.. /
/
- 1000 /
p
/
/
/
0 I I I I I l l 'l 1 - 0 1000 2000 3000 4000 5000 sooo 7000 sooo sooo ' Normal Consolidation Stress On Failure Plane , Fn , psf. j Figure B-6 Undrained Shear Strength on Failure Plane Versus Consolidation Effective , Stress - Triaxial Tests on Fill Samples Taken Before Surcharge ! a
Table B-1
~
Test Boring- Depth Location . yd w% o c 225.1.1 T16 6 . 2 ' ' -- 6.8' Tank Farm Aiea 114.4 16.9 -1440 South East of 139.1.1 TR2 12.6' - 13.2' Turbine' 128.8 9.7 1440 Building South East of 140.1.1 TR2 16.6' --17.1' Turbine 114.6 14.6 2880 Building 223.1.2 T15 16.0' - 16.5' Tank-Area West 130.8 9.3 5760 223.1.1 T15 15.2' - 16.0' Tank Area West 131.3 10.9 2160 - 222.1.1 T15 6.9' - 7.5' Tank Area West 118.6 14.2 5860 218.1.1 T12 11.4' - 11.9' Tank Area South- 124.1 13.1 2880 - 217.1.1 T12 3.9' - 4.4' Tank Area South 122.9 12.9 -720
" 2880 214.1.1 T10 7.6' -
8.1' 126 11.9 Co n r 213.1.1 T9 21.6' - 22.l' Tank Area South- 117.9 14.4 2880 8 bine 122.7 13.7 1440 142.1.1 TR4 13.6' - 14.1' Bu d 143.1.1 TR4 15.8' - 16.3 1 * ' u ld ng - b 145.1.1 TR4 26.6' - 27.1' Bu d ng b 147.1.1 TR5 16.l' - 16.6' *
- Bu d ng Condensate 164.1.1 CT4 22.7' - 23.2' Storage Tank 126 13.9 8640 SE Condensate
.167.1.1 CE6 15.2' - 15.7' Storage Tank 125.2 13 2160 North
7,..
.=
Table B-1 Continued Test Boring Depth Location. Y
- c d
Condentate 167.1.2 Cr6 15.9' .16.4' Storage _ Tank- 121.5 13.7- 4320 North condensate 168.1.1 CT6- 17.3' - 17.8' Storage Tank 122.7 13 . 6 8640 North
-Condensate 169.1.1 CT6 -19.7' - 20.2' Storage _ Tank 122.9 14.6 2160 North b
Table B-2 TCU TRIAXIAL TEST RESULTS
*f
~
Test c- l 3 max' u - No. Y w% psf psf iksf % u d f 164.1.1 126 13.9~ 8640 15,700 7.85 14 0 167.1.1 125.2 13 2160 7,600 3.80 20 - 0.16 167.1.2 121.5 13.7 4320 10,600 5.3 18 - 0.05 168.1.1 122.7 13.6 8640 14,500 7.25 20 0.1 169.1.1 122.9 14.6 2160 10,900 5.45 20 - 0.3 213.1.1 117.9 14.4 2880 2,800 1.40 20 0.55-214.1.1 126 11.9 2880 5,800 2.90 16 0 217.1.1 122.9 12.9 720 3,650 1.83 20 - 0.3 218.1.1 124.1 13.1 2880 4,500 2.25 20 0.17 222.1.1 118.6 14.2 5860 9,400 4.70 17 0.1 223.1.1 131.3 10.9 2160 5,900 2.95 20 - 0.1 223.1.2 130.8 9.3 5760 17,000 8.50 20 .0.1 I 225.1.1 114.4 16.9 1440 3,650 1.83 20 - 0.15 139.1.1 128.8 9.7 1440 9,000 4.5 23 - 0.37 140.1.1 114.6 14.6 2880 4,800 2.4 20 0.05 142.1.1 122.7 13.7 1440 4,000 2.0 15 - 0.3 143.1.1 120.3 16.2 2880 6,750 3.38 17 - 0.r3 145.1.1 121.9 13.9 5760 14,500 7.25 20 - 0.1 147.1.1 117.9 14.1 5760 8,750 4.38 20 + 0.2 o
Table B-3 SOIL SAMPLES TAKEN UNDER DIESEL C US:RATOR BUILDING Y J - Steve Sleve Depth Y Y total d Average l' 3 e Torvane LL PL 40 200 Y. d Coil Borine Suple Elev. ft per per per v% psf g tsf J J T 7 Averare C1mssification LG 1 614.4 19.7-20.2 127 7 10 5 5230 0 20 12 67 CL (63k.k) 615.h 19.1 141.8 127.2 11.5 0.T 614.8 12T.T 19.6 141.8 128.3 127.8 10.5 1.1 614.7 19.7 141.8 128.7 10.2 1.3 61k.1 20.3 141.8 127 1 11.5 2.3 DC 2 617.0 16.7-17.2 117.4 13.7 163 0 21 12 45 CL (634.0) 614.6 19.1-19.6 115.1 15.8 3163 0 25 14 84 610.1 CL 23.6-24.1 116.0 14.7 2125 0 29 14 68 CL 617.4 16.6 137.1 116.9 17.3 0.04 617 3 16.7 137.1 - 0.05 614.9 19.1 130.6 104.3 I 25.2 0.10 614.4 19.6 11 I 130.6 113.0 15 5 1.20 116.35 610.4 23.6 134.0 115.1 16.4 0.80 609.8 24.2 13k.) 116.T - 14.8 1.40 608.4 25.6 131.4 110.2 19.2 607.9 0.3 26.1 131.4 112.0 17 3 0.7 DG 3 613.2 21.2 135.0 118.0 14.k 0.28 (634.4) 612.7 21.7 135.0 118.1 1h.3 19 610.8 23.6 132.3 108.17 22.3 610.6 116.h0 0.3 23.8 132.3 102.0 29.6 0.05 619.3 24.1 132.3 112.h 17 7 1 75 609.7 24.T 132.3 117 2 12.9 1.62 DG k 616.9 17 5 127.3 114.8 1 10.9 2.0 rL (634.4) 615 2 19 2 134.4 110.2 1 22.0 09 614.6 19.8 134.4 117.0 116.75 14.8 0.69 614.0 20.4 134.4 118.6 l 13.3 0.60 DG 7 614.3 16.4-16.9 121.9 12 7 1418 0 22 12 59 CL-(631.0) 610.6 20.1-20.6 121.4 13.0 1402 0 20 12 58 CL 597.1 33.6-34.1 109.2 19.4 7292 o M 17 100 ct 613.8 17.0-17.3 125., 11.8 20 12 58 CL 596.6 34.1-34.6 112.8 18.3 kk 17 100 CL 614.8 16.2 133.8 99.6 3k.3 .02 118.16 614.6 16.h 133.8 118.8 12.6 .49 614.1 16.9 133.8 120.0 11.6 78 610.9 20.1 135.8 116.40 119.1 14.0 .k2 610.4 20.6 135.8 120.3 12.9 0.50 597.h 33.6 130.1 108.6 19.8 >2 5 596.9 34.1 130.1 109.2 19.1 >2.5 596.4 34.6 130.1 109.9 18.h >2 5 DG S 617.k 10.3-10.8 121.6 l 12.k 1890 0 20 11 54 (628.0) 618.0 10.0 136.5 122.4 121.0 11.5 1.4 121.6 617.2 10.0 136.5 118.9 l 14.8 1.4
w .
./
Table B3 (continued) SOIL S/JELES UNDER DIESEL GENERATOR BUILDING Yd - Sieve Sieve Depth Y total Id Average D3 e Torvane LL PL LO 200 Y d Aversae soil Enrir.g Cample Elev. ft _ ref pc f per w% psf. M tsf % % % % Classification DC 11 618.6 9.1-9.6 120.7 10.6 1204 0 21 12 49 609 9 17.8-18.3 119.9 13.3 1494 0 21 12 69 (62S.0) 618.3 9.6-9.8 127.T 122.4 9.9 21 12 36 122.76 618.9 91 134.3 121.3 10.7 0.4 Clo.k 17.6 132.6 105.0 l 26.5 0.2 62L.8 122.2 11.5 2155 0 19 12 53 CL DC 12 3.0-3.k CL (628.0) 617.0 10.1-11.2 110.7 13.8 504 0 22 12 55 13.8 813 0 21 12 56 CL 616.5 11.2-11.7 117.3 625.2 2.8 111.2 98.5 12.9 0.3 62L.6 3.k 111.2 1.3 621.k 6.6 138.8 123.0 12.8 0.6 11 'll 120.1 115 35 15.6 1.6 620 9 7.1 138.8 620.3 7.7 138.8 123.3 12.5 1.5 C 7.4 10.6 127.5 112.0
- 13.8 0.26 616.8 11.2 127 5 108.3 17 7 0.36 115.7 10.2 1.75 CL 616.3 11.7 127 5 613.9 1L.1 134.7 117.8 1k.3 0.35 CL 13.2 2667 0 25 12 62 CL DG 13 619.0 8.7-9.2 121.9 (628.0) 615.9 11.8-12.3 104.6 16.2 613.0 1L.7-15 2 103.8 17.9 3 627.L 0.6 116.78 15 5 0.35 110.1 626.9 1.1 13.4 0.52 62L.k 3.6 134.7 119.6 12.6 0.k3 623.8 L.2 13k.7 122.2 10.2 0.t6 619.3 8.7 138.5 122.8 12.8 1.3 618.8 9.2 138.5 122.6 13.0 1.2 to 14 626.9 1.1 142.5 126.7 12.5 0.88 (628.0) 625.1 2.9 142.7 129.4 10.3 0.63 62L.4 3.6 142.7 128.3 11.2 1.30 620.8 7.2 135.T 120.3 12.8 0.65 620.1 79 135 7 121.4 123.43 11.7 1.20 618.9 9.1 141.8 122.3 15.9 1.50 618.4 9.6 1kl.8 127.6 11.1 1.80 615.1 12.9 130.4 111.45 17 0 0.35 LG 15 625 1 2.7-3.1 126.1 l 10.h 3656 0 21 12 51 l (628.0) 622.5 5.5 139.8 120.8 124.06 15.7 1.75 126.1 622.1 59 139.8 125 3 l 11.6 1.60 l
4 TableB3 (continued) SOIL SAMPLES U' DER DIESEL GENERATOR BUILDING Yd - Sleve Sieve Y Y T Depth total d Average l' 3 c Torvane IL PL k0 200 d Soil Borira Sample Elev. ft pe r per per v% psf g tsf % % % % Averare Classification i DG 16 626.5 1.2-1.7 122.9 11.3 19 12 63 (628.0) 625 9 1.8-2.3 129.9 10.0 5180 0 20 12 53 626.9 1.1 140.8 122.8 Ik.7 0.37 626.3 1.7 1k0.8 12'i.5 .125.kk 10.4 1.55 126.k 623.5 4.5 131.7 121.0 8.8 1.0 N2.8 52 131.7 122.k 7.6 >2.5 625.0 3.0 144.8 131.6 10.0 1.5 DG 17 610 5 17.2-17.7 120 5 13.4 h166 0 26 12 61 l CL (628.u) 609 9 17.8-18.3 125.3 l 12.2 23 13 69 cL 121.6 121.85 1.k 122.9 610.9 17.1 138.1 13.6 610.3 17.7 138.1 120.0 15.0 I l 1.25 DG 19 626.3 1.3-2.0 126.9 11.9 3120 0 22 12 Sk CL (628.0) 614.8 12.9-13.k 107.8 17.3 250 0 27 13 64 614.2 13.5-1k.0 11k.2 15.3 26 13 65 626.8 1.2 1hl.1 125.8 12.5 1.13 615 9 12.1 11k.78 21.9 0.17 116.3 615.1 12.9 21.0 0.55 614.6 13.4 16.9 0.35 606.9 21.1 131.2 111.8 17.h 0.36 6C6.2 21.8 131.2 113.2 15 9 0.40 DC 20 614.0 14.0 128.6 109.0 l 17.9 0.25 (628.0) 613.6 1k.k 128.6 106.0 107 5 21.4 c.15 607.k 20.6 l 13.6 c.35 DG 21 616.5 11.2-11 7 123.0 11.8 900 0 20 12 61 CL (628.0) 615.6 12.1-12.6 116.5 13.5 300 0 20 12 56 617 0 10.8-11.1 110.7 17.T 22 13 60 615.0 12.7-13.2 119.8 13.1 20 12 53 626.9 1.1 132.9 121 5 9.k 0.88 626.1 1.9 132.9 120.6 10.2 1.0 622 9 51 137.9 124.k 10.8 1.63 617.k 10.6 1c6.8 118.8 0.8 117.5 135 3 26.7 616.9 11.1 135.3 115.3 17.3 0.88 616.3 11 7 135.3 117.2 15.k 0.40 615.8 12.2 135.3 120.4 12.4 0.40 615.9 12.1 135.1 114.1 18.4 0.10 615.k 12.( 135.1 119.0 13.5 0.65 613.1 14.9 13k.7 120.9 11.h 0.65 il -
Tatle B3(continued) SOIL V SA.JLC L'.CER GENERATCR BUITDINO Yd Y a Depth I Y total d Average 1 3 c Tcrvane LL FL k0 200 Average Soil Forinz Cample Elev. ft ref per per v5 rst M tsf T T % 5 ref Classification LG 22 616.1 11.6-12.2 118.7 12.8 13k0 0 21 12 51 (6:8.0) 626.5 1.5 141.7 124.4 13.9 1.h 625.A 2.6 136.6 122.6 11.k 0.7 62k.6 3.k 136.6 I?k.4 9.8 1.4 623.8 4.2 138.9 12k.0 12.0 0.5 623.3 4.7 133.9 12k.5 11.6 0.6 621.2 6.8 127.3 115.7 10.0 1.7 620.5 7.h 127 3 113.8 121.10 ll 0'I3 8.6 13e.9 122.3 11.'8 9 0.63 118.7 619.k 618.6 9.4 136.9 120.0 1k.1 0.55= 617.4 10.6 141.7 124.2 1k.1 0.85 617.4 'O.6 141.7 12k.2 Ik.1 1.50 616.4 11.6 130.5 114.9 13.6 0.45 7 615.8 12.2 130.5 113.7 14.8 0.50 12 606.h 21.6 Ik3.T 127.0 13.2 c.40 605.8 22.2 1h3 7 124.6 15.3 0.63 EG 23 617.0 10 7-11.2 109.2 13.5 365 0 20 12 60 CL (623.0) 616.6 11.3-11.5 117.7 13 3 19 11 55 CL 626.9 1.1 8.7 2.0 626.7 1.3 8.4 1.75 625.3 27 134.0 122.h 95 2.1 523.0 5.0 128.6 117.6 115.45 9.3 1.37 113 k5 617 9 10.1 128.8 110.8 16.2 0.2 617.3 10.7 128.8 112.7 14.3 0.3 616.S 11.2 128.8 113.0 13.9 0.18 616.1 11.9 123.8 11k.2 12.8 0.63 615.3 12.2 137.6 126.5 13.2 0.45 IX: 2k 626.8 1.0-1.4 127.0 10.1 36G C 21 12 32 (629.0) 1.k 127!0 11.2 127!0 626.6 i 1 75 l D3 26 613.6 1k.1-14.6 118.h 13.1 750 0 2k 12 57 (628.0) 613.3 14.6-14.8 93.2 30.0 39 14 91 625.8 2.2 1k3.8 129.2 11.3 1.13 625 5 2.5 l' .k 131.1 10.1 2.0 623.9 k.1 129 9 117.2 10'8 l'0 US.k 4.5 129.9 118.0 US 53 10.1 1.25 623.5 623.2 L.8 129.9 118.5 9.6 1.0 613.9 14.1 136.9 1c5.8 29.h 0.1 613.h 14.6 136.9 110.4 2k.0 0.33 612.4 15.6 134.0 113.5 18.1 0.95
Table B-4 Triaxial Test Results and Sample Characteristics-Sample's-Average Dry Densities, yd: 122 pcf ' Triaxial Test p' = U} 'oy -o3' _ Y # "
. Depth d Initial 2 ,
max 1+_3
. Boring Location ft pcf w% . confinement psf 2 fail -
T-12 Tank Area (south) 3.9-4.4 122.9 12.9 720. 1840 3500
~
TR-4 Turbine - Building 13.6-14.1 122.7* 13.7 1440 2000 4300 TR-4 (South 15.8-16.3 120.3 16.2 2880 3400 6600 . West) TR-4 26.6-27.1 121.9 13.9 5760 7250 14400 CT-6 Condensate 15.9-16.4 121.5 13.7 4320 5300 10500 CT-6 Storage . Tank 17.3-17.8~ 122.7 13.6 8640 7250 14100 CT-6 (North) 19.7-20.2 122.9 14.6 2160 5450 11000 e 9
.* w
s Table B-5 Triaxial Test Results and Sample Characteristics , Sample's average dry densities, yd : 116 pcf Triaxial Test p= p' =
"l "3 lY 3
~
c initial T Boring Location Depth ft d pcf w% confinement psf 2 psf max [2 psf,failad T-16 Tank Farm Area 6.2-6.8 114.4 16.9 1440 1825 3410 T-9 Tank Area , (south) 21.6-22.1 117.9 14.4 2880 1400 2730 TR-2 Southeast of Turbine Bldg. 16.6-17.1 114.6 14.6 2880 2400 5130 T-15 Tank Area (west) 6.9-7.5 118.6 14.2 5760 4700 9200 TR-5 Turbine Bldg. (s. west) 16.1-16.6 117.9 14.1 5760 4375- 8600 e t i
o ca. --
-APPENDIX C DETAILED-CALCULATIONS FOR DdTERMINATION OF FACTOR:OF SAFETY.FOR COMBINED STATIC AND EARTHQUAKE LOADINGS INTRODUCTION_
This appendix contains the calculations carried out-to assess the. factor of safety against a bearing capacity failure under the combined static e.nd earthquake leading considered in this analysis. These calculations are shown separately for the three failure surfaces (deep, intermediate and shallow) considered as potential sliding surfaces in case of a bearing capacity failure. DEEP' FAILURE SURFACE 5 The mass within the deep sliding surface considered in this analysis was divided into six different slices as shown in Figure C-1. The magnitude and inclination of the various forces , acting on each slice.are given in Table C-1. Equilibrium 1 force polygons for each one of these slices as well as for the entire sliding mass invcived in-th,is deep failure surface are given in Figure C-2. As indicated in Figure C-2, equi ' ibrium of all~ forces acting on each-slice resulted 2 i a " mobilized" friction angle of 16 along the sliding surface. " Effective" r
~
normal and shear' stresses at the base of each slice o and T n respectively, obtained from.the equilibrium force polygon i I I' u '
C/2 shown in Figure C-1, are given in Table C-2 Principal consolida-tion stresses B lc and J 3c at the base of each slice are also given in Table C-2. The magnitude of the resisting moments against sliding, originated by the presence of the~ surcharge as well as by the undrained shear strength that can be developed along the sliding surface are given in Table C-3. The undrained shear-strength values, T ff, at the base of each slice given in Table C-3 were obtained from the relationship between shear strength, Tff, and effective confining stress, E n, shown in Figure 14 of the main text; as previously indicated this relationship was obtained from the results of triaxial tests carried out on fill samples consolidated under anisotropic consolidation stresses similar to those calculated along the sliding surface (Table C-2). Calculations of the facter of safety against bearing capacity failure along this sliding curface are given in Table C-4. As indicated in Table C-4, the factor of safety was equal to 3.25. A corresponding factor of safety was also calculated assuming that the undrained shear strength along the deep sliding surface was determined from a different shear strength,. T ff, versus normal confining stress, 6n, relationship given in Figure 13 of the main text; as indicated in the main text,
C/3 this relationship includes triaxial test results obtained from anisotropically as well as isotropically consolidated fill samples. Results from triaxial tests carried out with isotropically consolidated-samples resulted in lower undrained shear strengths along the sliding surface. The magnitude of the undrained shear.strengch along the base of each slice as well as the magnitude of the resisting moment produced by the shearing strength along the failure surface are given in Table C-5. The magnitude of the resisting moment against sliding due to the surcharge is.also given in Table C-5. Calcu-lations of the factor of safety for the deep surface against a. bearing capacity failure under static and earthquake loadings are shown in Table C-6. As indicated in this table, the calculated factor of safety for the conditions assumed was equal to 2.86. INTERMEDIATE SLIDING SURFACE The mass within the intermediate sliding surface con-sidered in this analysis was. divided into six different slices as shown in Figure C-3. The magnitude and inclination of the various forces acting on each slice are given in Table C-7. Equilibrium force polygons for each one of. the slices as well
- as for the entire sliding mass included in the intermediate failure surface are shcwn in Figure C-4. As indicated in Figure C-4, equilibrium of all forces acting on each slice resulted in a " mobilized" friction angle of 18.5* along the sliding surface. " Effective". normal and shear stresses at e -
'C/4
- the: base of each. slice in.and tg respectively,'obtained fr'om-the equilibrium force polygon shown.-in Figure C-4 are givenin; Table'C-8.. Principal consolidation stresses'61c ""0 3c at the base of'each slice are-also given in Table C-8.
Factors of safety against a bearing capacity ' failure along the intermediate sliding' surface shown in Figure C-3 or aus alternate sliding surface in which the outer portion of the circular section was replaced by a passive wedge, Figure C-5, were calculated in this analysis. Factors of safety were cal-
^
culated for four different conditions as follows: Case.1) A factor of safety was calculated assuming the-- sliding surface shown in Figure C-3, and deter - mining the undrained' shear strength T gg, along. this surface from the shear strength, T gf, versus
-normal confining stress E n, relationship shown ,
in Figure 14-of'the main text. .As'previously indicated this relationship was obtained from the results of' triaxial tests carried out on , fill samples consolidated under anisotropic stresses similar'to those developed along the potential sliding surface.lar the static loads on the footing. Case 2)- A second factor of safety was calculated, assuming the same sliding surface of Case 1 above (Fi~ure g C-3) but determining the undrained-
C/5 shear strength along this surface, T , from the relationship shown in Figure 13 of the main text; as previously indicated this relationship between undrained shear strength, T ff, and normal confining stress, B , was obtained from triaxial test results carried out with anisotropically as well as isotropically consolidated fill samples. Case 3) A third factor of safety was calculated assuming the sliding surface shown in Figure C-5, in which the outer portion of the circular section was replaced by a passive wedge; the undrained shear strength, T ff, along this surface was determined from the relationship shown in Figure 14 of the main text; and Case 4) A fourth factor of safety was calculated assuming a the s'me sliding surface as in Case 3 above, Figure C-5, but determining the undrained shaar strength, rff, along this surface from the relationship shown in Figure 13 of the main text. Values of the undrained shear strength at the base of each slice of the sliding surface considered in Case 1 are shown in Table C-9. The magnitude of the resisting moments
, . , - . , . . . , - --e ~
-C/6 against. sliding developed by the weight of the sliding mass, the surcharge materials between the base of the footing and the ground surface, and the undrained shear strength available along this sliding surface are'also given in Table C-9.
Calculations of the factor of safety against a bearing capacity failure for Case 1 are given in Table C-10. As indicated in this table, the calculated factor of safety under these condi-tions (Case 1) was equal to 2.83. Values of the undrained shear strength at the base of each slice for conditions considered in Case 2 are shown in Table C-ll. The magnitude of the resisting moments against sliding developed in this case are also given in Table C-11. Calculations of the factor of safety against bearing capacity failure for these conditions are given'in Table C-12; these - calculations resulted in a factor of safety of 2.48. Values of the undrained shear. strength at the base of each slice of the sliding surface considered in Case 3 are shown in Table C-13. The magnitude of the resisting moments against sliding under these conditions, including the passive wedge resistance, are also given in Table C-13. Calculations of the factor of safety against bearing capacity failure under the conditions given in Case 3 are shown in Table C-14. As indicated in Table C-14 a factor of safety f of 2.67 was calculated for these conditions. l -
. C/7?
s Valuss ofithe undrainedishear strength alongfthe slidingisurface for! the~ conditions considered--in Case 4 are shown in'_ Table C-15.. The' magnitude of the res'istin'g moments- ~
- against1 sliding developed.under'these conditions,Jincluding-the passive _ wedge ~ resistance, are also'given in' Table'C-15.
Calculations-of the. factor of' safety.against a bearing capacity' failure under the' conditions given in' Case 4 are~shown'in Table C-16. As indicated-in this table, a factor'of: safety of: . n 1 2.37 was calculated. t SHALLOW FAILURE SURFACE The mass within~the shallow' sliding surface considered in this. analysis was divided =into six different slices as shown in Figure _C-6. -The magnitude a'nd inclination of the various forces acting on each slice are given in Table C-17. Equilibrium force polygons for each one of'the slices, as well [ as.for the entire. sliding mass, are shown in Figure C-7. As indicated in Figure C-7, equilibrium of all forces acting on each slice resulted in a " mobilize' d " friction angle of 19* along the' sliding surface. Effective normal and shear stresses at the base of each slice, U n and.T respectively, obtained n
'from the equilibrium force poly og n in Figure C-7 are given in Table C-18. Principal consolidation stresses a lc and E 3c at the base of each slice are also given in Table C-18.
. Factors of-safety against a bearing capacity failure i ( 9
'C/8 along-the shallow sliding surface shown in Figure C-6 or along zul alternate. sliding surface ~as shown in Figure C .were_ calculated in this-analysis.
.Four factors'of safety, corresponding to four condi-tions similar to those mentioned above were calculated as follows:
Case 1A) A factor of safety was calculated assuming the , sliding surface shown in Figure C-6; the undrained shear strength, T ff, along this surface was determined from the shear strength, T ff, versus normal confining stress, En, relationship-shown in Figure 14 of the main text. Case 2A) A second factor of safety was calculated assuming the same sliding surface as Case 1A) above, but the undrained shear strength along this surface, T ff, was obtained from the shear strength relationship shown in Figure 13 of the main text. Case 3A) A third factor of safety was calculated for the j sliding surface shown in Figure C-8. The undrained shear strength,-r ff, along this surface was determined from the relationship shown in Figure 14 of the main text.
7
~C/9 Case.4A) A fourth factor of safety was calculated assuming the same sliding' surface as in case 3A.
The undrained shear strength, Tff, along this surface was determined from the relationship shown in Figure 13 of the main text. Values of the undrained shear strength at the base of each slice of the sliding surface considered in Case 1A are given in Table C-19. The magnitude of the resisting moments against sliding along this surface are also given in Table C-19. Calculations of the factor of safety against a bearing capacity 4 failure under the conditions in Case 1A, are given in Table C-20. As indicated in Table C-20, the calculated factor of_ safety for these conditions was equal to 2.95. Values of the undrained shear strength at the base of each slice under the conditions considered in case 2A are given in Table C-21. The magnitude of the resisting moments against sliding along this surface are also given in Table C-21. Calculations of the factor of safety against a bearing capacity failure under the conditions in Case 2A are given in Table C-22. The calculated factor of safety for Case 2A was 2.59. l Values of the undrained shear strength at the base of each slice under the conditions considered in Case 3A are ! given in Table C-23. The magnitude of the resisting moments l against sliding along this' surface are also given in Table C-23. l l i
E C/10
-Calculations of the factor of safety against a bearing capacity failure under_the conditions in Case 3A are given in Table C-24.
As_ indicated in Table C-24, the calculated factor of safety for Case 3A was 2.81.- Values of the undrained shear strength at the base of each slice under the condition considered in Case 4A are given
'in Table C-25. The magnitude of the resisting moments against sliding along this surface are also given in Table C-25.
Calculations of the factor of' safety against a bearing capacity failure under the conditions in Case 4A are given in Table C-26. As indicated in this table the calculated factor of safety was 2.49. DEWATERED CONDITION An additional analysis was carried out to evaluate the' factor of safety against bearing capacity failure under-combined static and earthquake loading taking into account the effects of dewatering. In this case the water level, which had been previously assumed at an elevation of 627 was lowered to an elevation of 603 or below. Since the intermediate failure surface with a passive wedge resulted in the lower factor of safety against bearing capacity failure this surface, Figure C-5, was chosen to estimate the effects of dewatering on the bearing capacity of the footing. The magnitude and inclination of the various forces acting on each slice are
.given in Table C-27. Equilibrium force polygons for each one
. - - , m,.
C/ll of these slices as well as for the entire sliding mass are given-in t.gure C-9. As' indicated ir Figure C-9 equilibrium of all forces acting on each slice resulted in a mobilized friction angle of 15.5* along the sliding. surface. Effective normal and shear' stresses at the base of each slice, E and T n respectively, obtained from the equilibrium force polygon shown in Figure C-9 are given in Table C-28. Principal con-solidation stresses ele and 6 3c at the base of each slice-are also given in Table C-28. Two factors of safety against bearing capacity failure were calculated under these conditions. In one case, Case 1B, the undrained shear strength along-the failure surface was determined from the shear strength shown in Figure 14 of the main text. In the other case, Case 2B, the undrained shear strength along the sliding surface was determined from the shear strength relationship shown in Figure 13 of the main text. Values of the undrained shear strength at the base of each slice under the' conditions considered in Case 1B are given in Table C-29. The magnitude of the resisting _ moments against sliding under these conditions are also given in Table C-29. Calculations of the factor of safety against a bearing capacity failure for Case 1B are given in Table C-30. As indicated in Table C-30, permanent dewatering increased the factor of safety under these conditions to a value of 2.85. Values of the undrained shear strength at the base
i
' C/12-l of each slice'for the condition considered'in Case 2B are given in Table C-31.- The magnitude of the resisting moments against sliding developed'under these conditions are also given in Table C-31.- Calculations of the factor of safety against bearing. capacity for Case'2B are given in Table C-32. As
- indicated in this table, permanent dewatering increased the factor of safety to 2.54 .
i l I f
., ,,m..,.-----
E m, 4i UNITED STATES OF AMERICA NUCLEAR REGULATORY COMMISSION BEFORE THE ATOMIC SAFETY AND LICENSING BOARD In the Matter of: ) Docket Nos. 50-329 OM
) . 50-330 OM CONSUMERS POWER COMPANY ) Docket Nos. 50-329 OL (Midland Plant, Units 1 & 2) ) 50-330 OL CERTIFICATE OF SERVICE I, Michael I. Miller, hereby certify that. copies of the following in the above-captioned proceeding have been
. served upon all persons shown in the attached service list by deposit in the United States mail, first-class postage prepaid and by Federal Express, this lith day of October, 1982:
- 1. Testimony of Donald F. Lewis on Behalf of the Applicant Regarding Underground Piping at the Midland Plant
- 2. Testimony of Dr. Alfred J. Hendron, Jr.
Concerning Bearing Capacity of the Footings of the Diesel Generator Building
- 3. . Testimony of Dr. Alfred J. Hendron, Jr.
Concerning Seismic Shakedown
/
1s (d Michael I.' Miller Isham, Lincoln & Beale Three First National Plaza Suite 5200 Chicago, Illinois 60602 (312) 558-7500
-s .
[, l' , 4 SERVICE LIST
-Frank J. Kelley, Esq. St' eve Gadler Attorney General of the 2120-Carter Avenue State of Michigan- St. Paul, Minnesota 55100 Caro 3a Steinberg, Esq. .
Assist . Attorney General Atomic Safety ~&-Licensing Envite .antal Protection Div. Appeal Panel 720 Law Building U. S. Nuclear Regulatory Comm.
. Lansing, Michigan 48913 Washington, D.C. 20555 Myron M.-Cherry, Esq. **Mr. C. R. Stephens One IBM Plaza Chief, Docketing & Services Suite 4501 U.S. Nuclear Regulatory Comm.
Chicago, Illinois 60611 Office of the Secretary
' Washington, D.C. 20555
** Mr. Wendell H. Marshall 4625 S. Saginaw Road **Ms. Mary Sinclair Midland, Michigan 48640 5711 Summerset Street Midland, Michigan 48640
** Charles Bechhoefer, Esq. ** William D. Paton, Esq.
Atomic Safety & Licensing Counsel for the NRC Staff Board Panel U.S. . Nuclear Regulatory Comm. U.S. Nuclear Regulatory Comm. Washington, D. C. 20555 Washington, D.C. 20555
** Dr. Frederick P. Cowan Atomic Safety & Licensing 6152 N. Verde Trail Board Panel Apt..B-125 U.S. Nuclear Regulatory Comm.
Boca Raton, Florida 33433 Washington, D.C. 20555
** James E. Brunner, Esq. ** Jerry Harbour Consumers Power Company Atomic Safety & Licensing 212 West Michigan Avenue Board Panel Jackson, Michigan 49201 U.S. Nuclear Regulatory Comm.
Washington,-D.C. 20555 Mr. D. F. Judd ** Lee L. Bishop Babcock & Wilcox Harmon & Weiss P. O. Box 1260 1725 "I" Street, N.W. #506 Lynchburg, Virginia 24505 Washington, D.C. 20006
** Barbara St iiris 5795 North River Road f Route 3 Freeland, Michigan 48623
** Federal Express
.}}