ADDITIONAL CHEMICAL ENGINEERING INVESTIGATIONS FOR

ARIZONA'S MARIO CPA SSC SITE

Edward A. Nowatzki, Eugene Mulier, Jay S. DeNatale,

German A, Ibarra-Encinas, Abdul-Hakim M. F. AlGhanem (1)

and John Welty (2)

TABLE OF CONTENTS

1.0 DESCRIPTION OF SURFACE STUDIES ............................... 1

1.1 In~Place Testing: Field Slope Testing

2 • 0 IsABORATORY TESTING II.................. II . . . . . " .. II 0 . . . . . . . . . . . . . 3 2.1 Sieve Analyses 2.2 Atteroorg Limits Analyses 2.3 Soil Classifications 2.4 Direct Shear Testing 2.5 Triaxial Testing 2.6 Ar1Ellyses of Collapse Potential

3 . 0 FINDINGS ..................................................... 5

3.1 Results of Drilling (Fanglcrnerate Boreholes Logs) 3.2 Re~llts of Field Slope Testing 3.3 Results of Laboratory Testing 3.4 Canparability of Fru1g1craerate Strength Data

3.4.1 Lab Data Versus Dilatometer Data 3.4.2 Lab Data Versus Field Slope Data 3.4.3 SPT Blow~Count Data Versus Seismic Survey Data

3.5 Open CUt Construction -- Stability Analyses for 90 degree, 60 degree, and 45 degree side slopes

4.0 CONSTRUCTION REQUIREMENTS .................................... 22

5 • 0 RElrEREN'CES ••••••••••••••••••••••••••••••• ell • • • • • • • • • • • • • • • • • • • 23

APPENDIX A: FIELD BORING LOGS

APPENDIX B: FIELD SLOPE TEST FAILURE LOADS AND ~ES

APPENDIX C: LABORATORY DATA SHEETS AND CURVES

1.0 DESCRIPTION OF SURFACE STUDIES

1.1 In-Place Testing: Field Slope Testing

Much of the basin fill and fanglomerate is lightly to moderately cemented, and it is very d.ifficult to acquire undisturbed specimens from standard field sampling techniques. Disturbed and undis·turbed test speclmens will e>d1ibit nearly identical internal friction angles ¢, provIded that the specimens possess similar void ratios. HCJVI1ever, the materIal's apparent cohesion c generally decreases as the degr(~e of sampling·-induced disturbance increases. For this reason, laboratory testing rnay significantly underestimate the shear strength of the material in it's natural, undisturbed state, A novel program of field slope testing was therefore corlducted to .identify the shear strength of the basin fill in it's natural state. A 10-foot deep, 35~foot long benched trench was dug in the material with a ste~n shovel, and several separate slope failures were then illduced by applying a Slwface loading at the crest of the newly created vertical slope.

A one-inch thick, square steel plate, 8 inches or 15 inches in length, was placed near the crest of the slope, as shown in Figure 1 . 1 , 1 . A uniform load of increasing intensity was then applied to the plate by means of a 50--ton hydraulic jack. The rear-end of the steam shovel provided the reaction support for the hydraulic jack. The surface surcharge was gradually increased unt.i1 the slope failed, and the failure load and slip Sl"Lrface geometry were then recorded. Ten separate tests were performed -- five on the upper slope and five more on the lower slope (see Appendix A). The shear strength par~neters of the mldisturbed basin fill were then back-calcu.lated by means of an accepted analytical procedure (described in detail in Section 3.2).

5.0fJ If

Figure 1.1.1

Sketch of Field Slope Testing Set-up, Maricopa Site

TABLE OF CONTENTS

1.0 DESCRIPTION OF SURFACE STUDIES ............................... 1

1.1 In-Place Testing: Field Slope Testing

2 • 0 LABORATORY TESTING ........................................... 3

2.1 2.2 2.3 2.4 2.5 2.6

Sieve Analyses Atterberg Limits Analyses Soil Classificatio~s Direct Shear Testing Triaxial Testing Analyses of Collapse Potential

3 • 0 FIlIDINGS ........•............................................ 5

3.1 Results of Drilling (Fanglomerate Boreholes Logs) 3.2 Results of Field Slope Testing 3.3 ReSlllts of Laboratory Testing 3.4 Canparability of Fanglomerate Strength Data

3.4.1 Lab Data Versus Dilatometer Data 3.4.2 Lab Data Versus Fi~J,d, Slope Data 3.4.3 SPT Blow-Count Data Versus Seismic Survey Data

3.5 Open cut Construction -- Stability Analyses for 90 degree, 60 degree, and 45 degree side slopes

4.0 CONSTRUCTION REQUIREMENTS .................................... 22

6.0 ~................................................... 23

APPENDIX A: FIELD BORING LOGS

APPENDIX B: FIELD SLOPE TEST FAILURE LOADS AND GEam'RIES

APPENDIX C: LABORATORY DATA SHEETS AND CURVES

1.0 DESCRIPTION OF SURFACE STUDIES

1.1 In-Place Testing: Field Slope Testing

Much of the basin fill and fanglomerate is lightly to moderately cemented, and it is very difficult to acquire undisturbed specimens from standard field sampling teclmiques. Disturbed and undisturbed test specimens will exhibit nearly identical internal friction angles ¢, provided that the specimens possess similar void ratios. However, the material's apparent cohesion c generally decreases a.s the degree of sampling-induced distu.rbance increases. For this reason, laboratory testing may significantly underestimate the shear strength of the material in it's natural, undisturbed state. A novel program of field slope testing was therefore conducted to identify the shear strength of the basin fill in it's natural state. A lO-foot deep, 35-foot long benched trench was dug in the material with a steam shovel, and several separate slope failures were then induced by applying a surface loading at the crest of the newly created vertical slope.

A one-inch thick, square steel plate, 8 inches or 15 inches in length, was placed near the crest of the slope, as shown in Figure 1.1.1. A uniform load of increasing intensity was then applied to the plate by means of a 50-ton hydraulic jack. The rear-end of the steam shovel provided the reaction support for the hydraulic jack. The surface surcharge was gradually increased until the slope failed, and the failure load and slip SlU'face geometry were then recorded. Ten separate tests were perfonned -- five on the upper slope and five more on the lemer slope (see Appendix A). The shear strength parameters of the undisturbed basin fill were then back-calclllated by means of an accepted analytical procedure (described in detail in Section 3.2).

Figure 1.1.1

" "

"-"-"-"-:---:: I

~""'" """'."'""' . ...".,....,,, """,',....,,""' .. ",'",.,','.', ""'., ,...;:, '~"-'>------ I ~ ~-r~'M;; ~,

~~ ja.QU ff

t~

Sketch of Field Slope Testing Set-up, Maricopa Site

2.0 LABORATORY TESTING

2.1 Sieve Analyses

Sieve analyses were performed on samples acquired from different depths within the various boreholes. All analyses were conducted in accordance with ASTM D421 and D422 standards for s~nple preparation and testing. A s~nple weighing between 100 and 500 grams was oven~dried and then placed in a sieve stack consisting of #20, #40, #60, #140, and #200 sieves. The stacl< was vibrated in a mechanical sieve shaker for about ten minutes, and the percentage (by weight) of material passing each sieve was recorded. 'I'he s~nple's grain size distribution curve was constructed from the five data points obtained by means of the sieve analysis. The in situ moisture content of each sample was also measured.

2.2 Atterberg Limits Analyses

Atterberg Limits analyses were performed on remolded samples acquired from different depths within the various boreholes. All analyses were conducted in accordance with ASTM 4318 standards.

2.3 Soil Classifications

The borehole samples were classified according to the Unified Soil Classification System (USCS) using the standard ASTM D2487 procedure. The results of the sieve analyses and Atterberg Limits testing provided the necessary nunlerical data relating to grain size and plasticity.

2.4: Direct Shear Testing

Direct shear testing was performed on intact C'ME and diamond drill core samples from boreholes MA-12 and MO-12. All testing was conducted in accordance with ASTI1 D3080 standards. Tests were run at several different normal (vertical) stresses, and continuous measurement was made of both shear stress and shear deformation. Vertical displacement during shear was not recordecl.

2.5 Triaxial Testing

Unconsolidated Undrained (UU) triaxial testing was pel "formed on intact diamond drill core salnples from borehole MO-12 f in accordance wi th ASTM D2850 standards. Each 2-inch (50 llun) dialneter t 4. 5-inch (114 lIDn) high cylindrical sample was tested at it IS natul:'al moisture content. The testing was all strain-controlled, with an axial deformation rate of 0,40 in/hr (10 mm/hr). Each test was carried beyond the point of peak: deviator stress and to an axial strain of at leae;;t 5%.

2.6 Analysis of Collapse Potential

The procedure followed in determining the collapse potential of near-surface soils at the Maricopa Site is described in detail by Nowatzki (1980) and is referred to as the II one-dimensional pseudo"-consolidation test." The test is called a pseudo-consolidation test because the sample in the oedometer is not saturated prior to load application, as is the case in a conventional consoHdation test, Instead, a series of tests is performed in wrlich undisturbed samples, approximately two inches in diameter and one inch thick, are placed at in-situ moisture content in an oedometer. Following application of a 200 psf seating 1~1, each srunple is loaded in a sequence of vertical stress increments until a predetermined rnaximum stress is reached. The appUed stress increments generally double the magnitude of the existing stress. Following application of a given stress increment I displacement readings are taken periodically (usually 15 minutes apart for the first hotu~ and then once every ho'UI") until the difference between two successive readings is less than 0.001 inches. ~len the vertical stress on the sample has been incremented to a level approximately equal to the anticipated allowable foundation pressure, the sarr~le is saturated while still under load, and displacement read1ngs a1:'e taken in the same manner as described above. The anticipated al10wable foundation pressure is predetermined on the basis of design loarJ.s, the foundation soil's strength properties I and the type and size of fou:ndation system that is most economical. In general, the pseudo-consolidation test as described above can be completed within 24 ho'UI"s. The results are usually reported in terms of total applied stress and percent compression (strain). The following loading sequences were used in this study:

1. 200 psf, 1200 psf f saturation. 2. 200 psf, 2400 psf, saturation. 3. 200 psf, 1200 psf, 2400 psf, saturation. 4. 200 psf, 1200 psf, 2400 psf, 4800 psf, saturation. 5. 500 psf, 1000 pst, 2000 psf, 4000 pst, 8000 pst, 16000 psf,

32000 psf, 4000 psf, 32000 ];)Sf I saturation.

3.0 FINDINGS

3.1 Results of Drilling (Fanglomerate Borehole Logs)

Thirteen :new test borings were conducted as part of the most recent geotechnical investigations, The actual field boring logs are contained, for reference, in Appendix A. The borings were taken to a depth of up to 200 feet, A wide range of materials were encountered I ranging from boulders and cobbles, to clayey gravels (GC soils, according to the Unified Soil Cla..",sification System) I to well- and poorly-graded sands, silty sands, and clayey sands (8W, SP, SM and SC soils) I to silts, sandy silts I clays, and sandy clays of low plasticity (ML and cr. soDs). No sil ts or clays of hlgh pla..",tici ty were encountered, Standard penetration testing WOE done at boreholes MA-l, MA-3, MA~3A, MA~4, MA-6, MA-l0, MA-ll, MA-12 , and Mk~13, and all blow count information is contained on the respective boring logs (see AplJendix A) •

3.2 Results of Field Slope Testing

DisDJrbed and unclisttu~~.od test specimens exhibit nearly identical Internal friction angles 'I, provided that the specimens possess similar voId ratlos. Hence, the friction angle of the basin fi11 in its natuxal state may be fixecl at ¢ "'" 32 -- the value that was measured by direct shear laboratory testing on intact samples extracted from borehole MA-12.

A slope failure is associated with a safety factor of F ~ 1. Since the applied surface surcharge I failure suxface geometry, and in s1 tu unit 'Weight were measured for each of the ten individual field slope tests, the cohesive component of shear strength becomes the only relevant uriknow-n, Stability analyses of the slopes may therefore be used to back-calculate the h:'1Sin fill's undisturbed cohesion c. The relatively small loaded surface area led to the development of a roughly wedge-shaped slip surface (Figure 3.2.1), and conventional two-dimensional plane strain slope stabiU ty procedtlres I such as Bishop's Modified Method (Bishop, 1955), Jan.bu's Method (Janbu, 1973) I or the Morgenstern-Price Method (Morgenstern and Price, 1965), would overestimate the cohesion associated with a given safety factor. The method of Hovlanct (1977), on the other hand, aCC01lnts for the "end-effects II associated with a truly three-dimensional failure suxface, and this procedure was therefore used to analyze the field test data,

The cohesion c is calculated from the equation (Hovland, 1977):

where:

C :::::::

rb'sin i [~---------][F

3B 3

tan III ---------] B tan i

2b' B = {[----- sin i]2 + 1}1/2

w

r/J ::::: internal friction angle ::::: 32 i := angle of inclination of the slip surface (measl..lred) b l = maximum depth of the failure surface perpendicular to the

crest of the slope (measured) w := maximum length of the failure surface parallel to the

crest of the slope (measw:'ed) ,. ::::: equivalent tmit weight (which accounts for the unit

weight of the soil plus the applied surface sl.;U'charge) F3= the three-dimensiorv'3l safety factor::::: 1 (at failure)

A complete record of the field slope testing is included, for reference, in Apperu.iix B, An example of the back

3.3 Results of Laboratory Testing

The data obtained from the laboratory sieve and Atterberg Limits analyses, and the direct shear, triaxial, and consolidation testing are included for reference in Appendix C. The results of the sieve and Atterberg Limits analyses are surmnarlzed in Table 3.3.1, and the results of the strength testing program are summarized in Table 3.3.2.

The :i::'esul ts of the pseudo-consolIdation tests are summarIzed in Table 3.3.3 and shown graphically in Figure 3.3.1. Jennings and I{night ( 1975) regard the strain occurring at a saturation stress of 4000 ps:f as an index of collapse called the "Collapse Potential". They define the following critical values for the Collapse Potential (Cp ):

gp.J!l

0-1 1 - 5 5~1O

10 -20 >20

Severity

No problems Moderate problems Problems Severe problems Very severe problems

As can be seen from Table 3.3.3 and Figure 3.3.1, the data obtained from pseudo-consolidation tests performed on sol1s from the Maricopa S1 te suggest a Collapse Potential of approximately 9%. TIlis indicates that there is a potential for settlement problems as a result of collapse. However I the field samples were not retrieved directly into oedometer rings I as is usually the case. Thus, pseudo-consolidation test specimens had to be prepared in the laboratory by extruding them from the field salnpler into the oedometer rings. This dual handling proced:ure undoubtedly resulted in sample disturbances which caused the specimens to have greater values of Cp than conventional "undisturbed" salnples would have had. For this reason, the severity of the collapse problem is expected to be less than that suggested by the laboratory test data. This interpretation is consistent with the evaluation of collapse susceptibHi ty made on the basis of the Gibbs criterion as shown in Fi.gure 3.3.2. Therefore, methods typically recommended by geotechnical engineering consultants in Arizona for stabilizing such soils (e.g. excavation and recanpaction under controlled conditions) are expected to be effective for collapse susceptible soils at the Maricopa site.

The pseudo-consolidation tests in this study were performed on salnples retrieved from depths of from 30 to 60 feet. Ali (1987) has shown that the probability of encmmtering collapse susceptible soils decreases with depth. Therefore, so11s exhibiting a high degree of collapse potential are not expected to occur below a depth of about 30 feet, and protJably more shallow. Since the potential for collapse settlement is such an important consideration in the design of founclations for surface structures, and since the results of laboratory pseudo-consolidation tests are susceptible to sample disturbance I full scale field tests should be performed at the site of the main campus to verify the existence of collapse-susceptible so11s and determine their sever i ty with (:lepth.

3.4 Comparability of Fanglomerate strength Data

~J=?b~~Jl~I§,1~::fLPilatometer Data: The results of the field dilatometer testing program have not yet become available.

3.4.~Lab_Data Versus Field Sl~~: Direct shear testing was performed on intact specimens of basin fill material extracted from borehole MA-12 at depths of 30 to 65 feet. As mentioned previously in Section 3.3, this direct shear testing yielded minimul'll shear strength parruneters of ¢ = 32 and c ::::: 2160 psf. By comparison, the field slope testing on the unweathered lower bench (involving material at depths of 5 to 10 feet) yielded an average cohesion of c = 2430 psf. As a consequence of sampling-induced disturbance, the laboratory specimens should be expected to have a cohesion which is lower than that for the material in its natural (undisturbed) state. Indeed, it appears as though the process of sampling reduces the material's cohesion by at least 10%. The results of the laboratory and field testing are consistent, and minimum strength parameters of ¢ ::::: 32 and c ::::: 2430 IJSf may be assigned, with confidence, to the cemented basin fill material.

3.4.3 SPT_ BIO\l11.::Q9~!~~smi£_Surv~ Dat~: A comparison is made between elastic moduli determined from 8PT blaIJ counts obtained from boreholes on or near the ring alignment with moduli values computed fran velocity data obtained from seismic sturveys performed at or in the vicinity of the borehole locations. The seismic data were correlated to modultlS by the relationship:

where:

E ::::: VC2 [(1+Y)(1-2~}/(1-v)][l/g]

E ::::: modulus of elasticity Vc ::::: compression wave velocity " ::::: Poisson I s ratio ~ ::::: unit weight of soil or rock 9 ::::: acceleration due to gravity

This expression can be rewritten as:

where

E :: V 2K C

K ::::: a material constant

For the materials at the Maricopa site the following values of the material parameters were found to be appropriate:

~ "" 0.33 1" '"' 125 pcf

If these values are used with 9 - 32.2 ft/sec/sec to evaluate K, and Vc is expressed in ft/sec then:

E(tsf) = 0.00131Vc2

The SPT blow count data were correlated to modulus by the following relationship developed for sands and gravels by Wrench and Nowatzki (1987) :

E{tsf) = 22.2 N 0.888 a

Table 3.4.1 shavvs that although the moduli do not appear to correlate well nUJl1erically at shallow depths, there is consistency between the results regarding an increase of modulus with depth. Both sets of data suggest that the material within approximately 20 feet of the surface has a much lower modulus than the alluvial fanglomerate below that depth. At shallow depths the modul~s is generally less than 1000 tsf, whereas at depths greater than approximately 20 feet the modulus is generally greater than 3000 tsf and in some locations greater than 7000 tsf. This agrees with results reported elsewhere (see, for example, DeNatale et aI, 1987 and Beckwith et aI, 1988) .

It should be noted that seismic data give an integrated value for velocl ty and modulus over a range of depths, whereas blow counts are taken at specific depths. With the seismic data, discrete changes of strata dellc'Si ty are recorded by discrete changes in compression wave velocity. SPT blow counts, on the other hand, are non-integrative and. pertain only to the depth at which they are taken, It would not be correct to average blow counts over a range of depths and use the average value as representative of the stratum. In addition, for this study, the SPT blow counts were recorded for three 6--inch increments with the reported N-value (N ) being the sum of the blows over the last two 6-inctl increments. If "the full 6-inch penetration was not reached after 50 hammer blows in any increment, the test was discontinued, and the count was reported as 50 blows for the actual penetration obtained in that increment. In those cases the 8PT blow count was increased by linear proportion to the full 6-inch penetration, so that a value of N could be computed for use in the equation above. This extrapolation &n significantly underestimate the actual 6-inch penetration that would have occurred had the test been continued. Therefore the moduli reported in Table 3.4. 1 for depths greater than 10-20 feet represent lower limits of the in situ moduU that can be expected. In all cases, data were used from the seismic survey conducted closest to the borehole for which 8PT data were available.

3.5 Open-Gut Construction -- Stability Analyses For 90 Degree, 60 Degree, And 45 Degree Side Slopes

The primary component of the sse project is the 52-mile long collider ril19 that will be housed in a 10-foot di~neter concrete tunnel placed 30 feet or more below the ground surface. The most economical way to construct an underground tunnel of this type in Arizona is by the cut-and-cover method. In this approach, a tunnel is formed by excavating downward from the ground surface. Precast cylindrical concrete tm1nel s~nents are placed in the open trench with cranes, and the trench is then backfilled with the excavated soil. At the Maricopa

si te, about half of the underground accelerator ring can be placed by the cut-and-cover lllethod. Since the amount of material to be excavated decreases as the sides of the excavation become more steep, the cut-and-cover method becomes most economical when an l~~lpported vertical excavation profile can be used.

Bishop's Modified Method (Bishop, 1955) provides a conservative relationship between the safety factor F and the excavation depth H in a homogeneous deposit where circ,"Ular sl ip surfaces are expected, If the cohesion and friction angle of the alluvial material are conservatively fixed at c ::::: 2430 psf and ¢ ::::: 32 , a computer-aided stability analysis based on Bishop's Modified Method indicates that it is possible to excavate vertically downward to a depth of 100 feet and still have a safety factor against slope failure in excess of one. Table 3.5.1 presents a SUlIDnary of safety factors for heights of up to 100 feet and slope angles of 45 , 60 I and 90. It should be pointed out that vertical slope faces are actually preferable to inclined ones, since vertical faces are less susceptible to weakening due to water infi! tration, in the event tha.t precipitation occurs before the excavation has b~en backfilled,

TABLE 3.2.1: VALUES OF COHESION AS BACK-GALCULATED FROM FIELD SLOPE TESTING.

Cohesion Test # Bench (in psf)

1 Upper 2 Upper 1393 3 Upper 1203 4 Upper 738 5 Upper 469

6 Lower 7 Lower 2619 8 Lower 2657 9 Lower

10 Lower 2007

Average Cohesion For Upper-Bench Failures 950

Average Cohesion For !.ower-Bench Failures 2430

TABLE 3.3.1: RESULTS OF LABORATORY SIEVE AND A'l'TERBERG LIMITS ANALYSES.

% < #200 Moisture Atterberg Limits Borehole Depth (ft) Sieve Content % LL PL PI

MA~lO 0.5 - 2.0 29 3.4 3.5 - 5.0 26 21 20 1 8.5 - 9.0 14 4.3

13.5 -13.9 17 19.5 -23.5 24 30.0 -30.5 13 23 19 40.0 -40.3 19 55.0 -55.4 22 60,0 -60.5 5.7 65.0 -65.4 30 68.5 -73.0 7 80.0 -80.2 35 85.0 -85.4 7.0 90.0 -90.4 37

MA-11 0.5 - 2.0 31 3.8 10.0 -10.4 14 6.7 20.0 -20.3 35 6.2 30.0 -31.5 14 4.2 19 18 1 40.0 -40.3 23 5.5 55.0 -55.2 21 6.4 65.0 -65.5 19 3.4 73.5 -73.8 11 6.3

MA-12 0.5 - 2.0 12 2.0 22 9 13 15.0 -15.7 22 7.1 25.0 ~25.8 38 7.6 30.0 -32.8 38 33.5 -34.9 16 3.7 48.5 -48.9 41 7.4 53.5 -54.2 34 8.0 61.0 -62.0 27 7.1 37 68.5 -69.1 38 12.6 73.5 -73.9 28 9.2

TABLE 3.3.1: RESULTS OF LABORATORY SIEVE AND ATTERBERG LIMITS ANALYSES. ( Continued)

% < #200 Moisture Atterberg Limits Borehole Depth (ft) Sieve Content % LL PL PI

MR-l 5 - 10 10 2.4 28 19 9 15 - 20 13 7.4 44 22 22 25 - 30 22 2.4 26 22 4 35 ~ 40 14 4.3 31 17 14 45 - 50 14 3.1 55 - 60 5 4.0 65 - 70 6 4.9 75 - 80 11 4.8

MR-2 5 - 10 11 6.2 54 50 4 15 - 20 21 4.3 32 22 10 25 - 30 4 2.9 35 27 8 35 - 40 17 3.6 37 20 17 45 - 50 17 0.9 55 - 60 19 2.5 65 - 70 19 3.6 75 -- 80 13 2.8

TABLE 3.3.2: RESULTS OF LABORATORY STRENGTH TESTING

Source of Type of Laboratory Cohesion Friction Specimens Ma.terial Test c (l2§f} AngIe !

MA--12 Basin Fill Direct Shear 2160 32~

MD~12 Fanglanerate Direct Shear 9000 27·

MD-12 Fanglomerate Trirudal 11400 26-(Peak Value)

MD-12 Fanglomerate Triaxial 2000 39· (at 5% Strain)

TABLE 3.3.3: ~ OF RESULTS OF PSEtJDO-CONSOLlDATION TESTS FOR BOREHOLE MA-12

Applied Stress Strain Sample Depth (psf) (%)

30 ' 3" -- 32' 9" 1200 2.38 1200 S 11.59

C = P 9.21

30'3" - 32'9" 2400 5.87 2400 S 13.01

C = P

7.14-

1200 1.85 2400 4,70 2400 S 11.45

C = P 6.15

30'3" - 32'9" 1200 1. 78 2400 4.38 4800 7.36 4800 S 16.68

C = P 9.32 61'0" - 63'6" 500 0.50

1000 1.42 2000 2.61 4000 3.49 8000 4.91

16000 6.89 32000 8.44

4000 7.20 32000 8.68 32000 S 25.00

Cp = 16.32

TABLE 3.4.1: COMPARISON OF ELASTIC MODULI OBTAINED FROM SPI' BLOW COUNTS WITH MODULI OBTAINED FROM SEISMIC SURVEYS.

SPI' Blow Count Data Seismic SUrvey Data Maximum

Depth Modulus Depth Modulus Boring (ft) (tsf) (ft) (tsf)

MA-l o - 11 < 500 o - 57 5694 11 - 75 >4000

MA-3 o - 11 4000 6 -300 20522

MA-4 o - 11 < 500 o - 5 1500 11 -100 >3000 5 - 36 23784

36 -509 78357

MA-6 o - 5 < 500 o - 15 2046 5 - 70 >5000 15 -260 24481

MA-I0 o - 9 < 500 o - 44 5909 9 - 65 >3000 44 -106 16379

65 -150 >7000 106 -319 56202

MA-ll o - 20 3000 8 - 29 7470

40 - 74 >7000 29 -221 14798

MA~-12 o - 10 < 500 o - 9 247 MA-12 10 - 74 >4000 9 -130 1.0926

MA-13 o - 22 5000 97 -161 16583

TABLE 3.5.1: SAFETY FACTORS AS A FUNCTION OF SLOPE HEIGHT AND SLOPE ANGLE.

Slope ---------- Safety Factor Values -------------Height Slope Angle Slope Angle Slope Angle (feet) of 45 Degrees of 60 Degrees of 90 Degrees

40 4.05 3.35 2.17 50 3.47 2.87 1.81 60 3.08 2.53 1.57 70 2.80 2.28 1.40 80 2.59 2.09 1.27 90 2.41 1.94 1.17

100 2.29 1.82 1.09

• ____________ ~2~.7Le~fut~· ______________ ~7/

- 7 ~ /' ._._/-._._. / /

----------.-4r---

----,---.----.~~-

20.00

-

--

15.00

-

-----

.r--..... -~ -

'--./ 10.00 -

0.. -o ---

--

--

5.00 --

-----

--

0.00

y

V V

V

V /

V V

V

10 Saturation Stress (ksf)

.-*

/ V

Collapse Strain (%) versus Saturation Stress (ksf) Maricopa SSC Site - Borehole MA-12

Collapse Potential = Collapse Strain at Saturation Stress = 4 ksf (Jennings & Knight, 1979)

20.00

15.00

,.-..... I:R

'----'"10.00 Q..

o

5.00

0.00

--------------------

-----------

-.---

1

~

V V V

V V

V V

V

10 Saturation Stress (ksf)

Figure 3.3.1

.--*

L V

Collapse Strain (%) versus Saturation Stress (ksf) Maricopa SSC Site - Borehole MA-12

Collapse Potential = Collapse Strain at Saturation Stress = 4 ksf (Jennings & Knight, 1979)

10 '

-u c.. >-!::: C/)

z w 0

>-c:: 0

50r-----------------~--~----------------~--~----~

60

70

80

90

100

110

COLLAPSING SOILS Specific Gravity

2.6

Specific Gravity

2.7

Depth == 30'3" - 32'-9" A

NON-COLLAPSING SOILS

Depth == 61'0" - 63'6"

•

120~~~----~--------~----~~~--------~--------~ 30 40 50 60 10 20

LIQUID LIMIT (%)

Figure 3.3.2

Collapse-Susceptibility Diagram (After Gibbs,196l) Maricopa SSC Site - Borehole MA-12

/

.--. . ---;/-.

. --/ '--. 1.00 ft.

1.00fl.

'-- .---

Figure 3.2.1

A Typical Field Slope Failure Surface Geometry

20.00 ---------

15.00 -------

,--... -~ -

,---'' 10.00 -a.. o --

------

5.00

----

.---

0.00

1

'y

/ V

V V

V V

V V

10 Saturation Stress (ksf)

Figure 3.3.1

/~ V

Collapse Strain (%) versus Saturation Stress (ksf) Maricopa SSC Site - Borehole MA-12

Collapse Potential = Collapse Strain at Saturation Stress = 4 ksf (Jennings & Knight, 1979)

,

'10'

50~--------------------------------------~----------~

.... -

60

70

u 80 a.

>-t:: C/)

z w o >- 90 c:: o

100

110

COLLAPSING SOILS Specific Gravity

2.6

Depth = 30'3" - 32'-9" A

NON-COLLAPSING SOILS

Depth = 61'0" - 63'6"

•

120~~~----~--------~--------~--------~--------~ 10 20 30 40

LIQUID LIMIT (%)

Figure 3.3.2

Collapse-Susceptibility Diagram (After Gibbs,1961) Maricopa SSC Site - Borehole MA-12

60

4.0 CONSTRUCTION REQUIRmlJENTS

Construction requirements for the injector complex, experimental chambers, and campus building area, as well as analyses for foundation bearing capacity, settlement, and swell/collapse potential are discussed in detail by Beckwith et al (1988). Details of these analyses need not be repeated herein.

5.0 REFERENCES

1. Ali, M.M, (1987), "A Probabilistic Analysis of the Distribution of Collapsing Soil in Tu.cson Using Kriging Method" I PhD. Dissertation, The University of Arizona, Department of Civil Engineering and Engineering Mechanics.

2. Beckwith, G.H., Weeks, R.E., and Cheng, S.S. (1988), "The Prelimina:ey Geotechnical Investigation Main Campus Area, Maricopa SSC Site," Technical Report, SHB Consulting Geotechnical Engineers, Phoenix, Arizona.

3. Bishop, A.W. (1955), "The Use of the S1ip Circle in the Stability Analysis of Slopes," Geotechnique, Vol. 5, pp, 7-17,

4. DeNatale I 3. S . ( 1986) , "Progr'am CSI.IPl : For Slope Stabili ty Analysis by Bishop's Modified Method, with a Search Routine :Based on the Simplex Metb.od of NeIder and Mead, II Technical Report, The University of Arizona, Department of Civil Engineering and Engineering Mechanics.

5. DeNatale, 3,S" Nowatzki, E.A., and Welty, 3.W. (1987), "Geotechnical Engineering Investigations for Arizona's SSC Sites If I Arizona Bureau of Geology and Mineral Technology Open-File Report 87-7.

6. Gibbs, B.3. (1961), "Properties which Divide Loose and Dense Uncemented Soils I" U. S. Bureau of Reclamation Report No. EM-608.

7. Hovland, H .... !. (1977), "Three-Dimensional Slope StabHity Analysis Method, II Journal of Geotechnical Engineering, ASCE, Vol. 103, No. GT-9, pp. 971-986.

8. Janbu, N. (1973) I "Slope Stability Computations," Emba.nkment Dam Engineering, casagrande Volume, pp. 47-86.

9. Jennings, J.E. and Knight, K. (1975), "A Gl'I.ic1e to Construction on or with Materials Exhibiting Addi tiona1 Settlement Due to Collapse of Grain structure," Proceedings of the 6th Regional Conference for Africa on Soil Mechanics and Foundation Engineering I Vol. 1, pp. 99-105.

10. Morgenstern, N,R. and Price, V.E. (1965) I "The Analysis of the Stability of General Slip SUrfaces I" Geotechnique I Vol. 15 I No.1, pp. 79-93.

11. Nowatzki, E.A. (1980), "Settlement Estimates in Collapse-Susceptible Soils I r' Journal of Civil Engineering Design, Vol. 2, No.2, pp. 121-148.

APPENDIX A: FIELD BORING LOGS

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PEPTH HOUR PATE _'~' SERGENT, HAUSKINS & BECKWIT! I" B I CONSUl TOIO OEOTtCHHICA1. ENOINEERS

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_1-js:::11 SERGENT, HAUSKINS & BECKWITH 1~1 CONSUlTING OEottCHNICAl ENGINEERS -- t - PHOENIX' TUCSON • AlBUOUERQUE • SANT 1>. FE • SA!' lAKE CITY • El PASO

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_1-jS~1 SERGENT, HAUSKINS & BECKWITH

1~1 CONSUlTiNG GEOTECHNICAL EtlGlNEERS --- 1 - PHOENIX' TUC~· ALeUOUEAOUE' SAHTA FE· SALT lAKE efTY· EL PASO

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'[g)'" I SERGENT, HAUSKINS & BECKWITH -, '" B I CONSUl TlNO OEO"I'ECHNJCAl ENGINEERS

-- t - PHOENI)( • TUCSOt-I • ALSUOUEAOUE: • SANTA. FE • SALT lAKE CITY • El PASO

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