Underreamed FoUnderreamed Footings in Jointed Clayotings in Jointed Clay

7/28/2019 Underreamed FoUnderreamed Footings in Jointed Clayotings in Jointed Clay

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Load-Displacement Behavior of Underreamed

Footings Bearing in Jointed Clays

Kenneth E. Tand, P.E., M.ASCE, C. Vipulanandan, Ph.D., P.E., M.ASCE,

And Michael W. ONeill, Ph.D., P.E.

ABSTRACT: Underreamed footings, sometimes referred to as belled shafts or

belled piers, are frequently used in areas where geological conditions are favorable fortheir installation. Underreamed footings are sized to provide resistance to compression,

uplift and lateral loads. In this study, the load-displacement behavior of two full size

underreamed footings load tested to failure in compression and three in uplift wereinvestigated. The measured compression loads ranged from 750 to 1090 kips (3340 to

4850 kN), and the uplift loads ranged from 109 to 220 kips (485 to 980 kN).

Due to the relatively high permeability of the slickensided and fissured clays,

four of the load tests were performed under partially drained conditions. FE analysis

using effective stress parameters allowing for the dissipation of pore pressures betweenloading intervals was required to predict the measured load-displacement behavior. The

at-rest earth pressure coefficient, slickensides and fissures in the clays, and permeabilityof the layers were important parameters to model the behavior of the footings.

Introduction

The bottom of drilled shafts are sometimes enlarged to increase their bearing

capacity to optimize foundation costs by reducing the quantity of concrete andconstruction labor required (ONeill, 1985). These belled shafts are also referred to as

underreamed footings or belled piers. This type of foundation system is commonly

used along the Texas Gulf Coast where favorable subsoil conditions exist for theirinstallation.

In order to better understand the load-displacement behavior of drilled shafts instiff clay, four load tests had been performed on drilled shafts, one underreamed, at the

State Highway 225/Loop 610 interchange in Houston, Texas, (ONeill, et. al., 1972).

These drilled shafts were tested to failure and the skin friction and end bearingdeveloped in the different soil layers were analyzed. Further studies of underreamed

footings performed by ONeill showed that underreamed footings were advantageous

because they could be economically installed below the active depth of the expansiveclays typical of the Texas Gulf Coast area, and that underreamed footings provided

adequate resistance for compression, uplift, and lateral loads.

______________________________________________________________________1Principal Engineer, Kenneth E. Tand & Associates; 2817 Aldine Bender, Houston,Texas 77032; Ph: (281) 590-1711; Fax (281-590-1430); Email: [email protected]

2Chairman and Professor; Department of Civil & Environmental Engineering;

University of Houston; Houston, Texas 77204; Ph. (713) 743-4278; Fax (713) 743-4260; Email: [email protected]

3Cullen Distinguished Professor, Department of Civil & Environmental Engineering,

University of Houston; Houston, Texas 77204 (deceased)


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In another study, an underreamed footing was load tested to failure in

compression at the University of Houston test facility (ONeill, et. al., 1985). The baseof the footing was instrumented with earth pressure cells to determine the pressure

distribution under the footing, and it was observed that a portion of the load was being

carried in suction at the roof (top) of the underream. Later, four more underreamed

footings were load tested in uplift at the same location (Yazdanbod, et. al., 1987), andpiezometers were installed to measure the suction that developed at the base of the

footings.

A recent study investigated the load-displacement behavior of the underreamedfooting load tested to failure in compression at the U of H test facility (Tand and

ONeill, 2003). Finite element analysis was performed to investigate the soil-structure

interaction.

This study included analysis of the two compression load tests, and three uplift

load tests referenced above. The overall objectives were to investigate the effects of the

following factors on the load-displacement behavior of underreamed footings in

compression and uplift:

Suction Fissured and slickensided clay layers. Soil strength parameters and permeability. In situ stress conditions (KO).

Site Geology

The two test sites were located east of downtown Houston, Texas, and they wereabout 4.5 miles apart. Both sites were located on a Pleistocene age deposit known

locally as the Beaumont formation (Mahar and ONeill, 1983). The Beaumont

formation is about 26 feet (8 m) deep at both sites, and it is underlain by an older

Pleistocene deposit known as the Montgomery formation. The subsoils were depositedin shallow coastal river channels and flood plains during the Peorian interglacial stage.

Both formations can be categorized as primarily clay with occasional interbedded seams

and layers of sand and silt. The clays and cohesive silts were overconsolidated to deepdepths due to desiccation that occurred when the water table was lowered during the

Second Wisconsin Ice Age (Bernard, 1962).

After alluvial deposition of the clays, large vertical fissures were formed by

shrinkage due to desiccation. Soil from the surface was washed down into the cracksduring periods of heavy rainfall. The soft sediments in the cracks were then

compressed when the clays swelled leaving locked-in horizontal stress (Al-Layla,

1970). This process was repeated throughout Pleistocene to recent geologic times, and

K0 values of 3.0 and greater have been measured at the U of H test site (ONeill, 2000).The process of desiccation and subsequent rewetting caused cyclic shearing

displacements in the clay mass that produced polished failure planes referred to as

slickensides. The slickensides are widely variable in size and orientation, althoughthere is a slight preference towards inclination at about 45 degrees to the horizontal

(Mahar and ONeill, 1983). The slickensides and fissures are flaws within the clay

mass, and such deposits are often referred to as jointed clays. The system ofslickensides and fissures that occurred during recent drying cycles produced a varying


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degree of desiccation and overconsolidation within the clay mass (ONeill, et. al.,

1995). The state of stress is a function of the distance from the joints.As shown in Fig. 1, desiccation has produced an assemblage of small to large

oddly shaped clusters that lie within the slickensides and fissures (macro structure), and

the soil within the assemblage is somewhat, but not perfectly, intact (micro structure).

The cyclic movements resulted in residual strength parameters along the slickensidesand fissures. The micro and macro structure affects the strength, deformation and

permeability properties of the soil mass. The clays are spatially inhomogeneous, and

exhibit some anisotropic properties due to their stress history.

Soil Samples

Fig. 1 - Conceptual Sketch of Jointed Clay Mass

The jointing structure and loading pattern greatly affects whether macro ormicro behavior will govern overall behavior of the foundation. The macro strength of

jointed clay will have small influence on the bearing strength of clay under a large matwhere the soil is somewhat confined, but the macro strength will have great influence

on the uplift resistance of an underreamed footing with limited embedment.

Compression loading on a footing will tend to squeeze the joints closer and reduce the

mass permeability, but uplift loading will tend to open the joints significantly increasingthe mass permeability. There can be significant differences in the pore pressure

gradient between the clay at the edge of a fissure that has been opened by uplift loading,

and in the center of the clay assemblage. Thus, selection of a constitutive model forjointed clays and soil parameters for that model present a significant challenge to the

geotechnical engineer.The subsoils properties at the U H test site have been documented in the

literature (Mahar and ONeill, 1983, and ONeill, et. al., 1995). A summary of thedatabase can be found on the web at www.uh.edu/nges. Shown in Fig. 2 is a subsoil

profile with a brief summary of key soil parameters. It should be noted that the

database was constructed from a very large number of laboratory tests on smalldiameter samples, and in situ tests.


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Depth

ft. m

0 0ft m

2

2

4

4

6

6

8

8

10

10

12

12

14

16

18

20

22

24

26

28

30

32

34

36

38

40

42

44

StratumW

-%-OCR

Fig. 2 - Generalized Subsoil Profile at U H Test Site

Symbol Key

w - moisture content

LL - liquid limit

Ip - plasticity index

- unit dry densitySu - undrained shear strength

c' - effective stress cohesion

' - internal angle of resistance

qc - cone tip bearingE - soil modulus

KO - at rest earth pressure coefficient

The engineering properties of the subsoils have also been extensively studied at

the state highway interchange (ONeill and Reese, 1972 and Woodward-Clyde, 1982).The site geology is very similar to the U H site, but the water table was at a depth of 15

feet (4.6 m) in 1969. When comparing the subsoil stratigraphy in the two different

Su c' qc ELL Ip - pcf- -ksf- -ksf- ' -ksf- ksf- KO(kN/m3) (kPa) (kPa) (MPa) (MPa)I

20 55 40 115 1.5 .2 25 15 240 3 10(18) (72) (9.6) (0.7) (11.5)

II

_W.T.

III 23 35 20 112 2.0 - - 100 280 2.5 8(17.5) (96) (4.8) (13.4)

IV 28 70 45 110 2.5 .6 20 45 350 2.0 6(17.3) (120) (28.7) (2.2) (16.8)

Stratum I Fill: Mixture of shell fragments and sand with clay binder

II Stiff to very stiff clay

III Very stiff sandy clay, with sand seams and calcareous layers

IV Stiff to very stiff clay

V - Stiff silty clay with silt lenses

VI - Very stiff sandy clay

V

VI

18 24 10 118 2.0 - - 50 480 1.5 5

(18.5) (100) (2.4) (23)

15 30 18 128 4.2 .8 30 70 600 1.2 4

(20) (200) (37) (3.4) (29)


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engineering reports for this site, it should be noted that the site grade was raised about 6

feet (1.8 m) with fill after the load tests were performed in 1969. Cone penetration testsperformed at a later date (ONeill, et. al., 1985) clearly showed that the footing was

bearing in stratum V, but that the strata was 20 feet (6.1 m) below grade at this site.

Finite Element ModelingThe complex subsoil profile and geologic stress history, as well as the variable

stress paths that occur during loading, make conventional methods of analysis

impractical. Hence, the finite element method was selected to investigate the subsoilparameters affecting the load-displacement behavior of the underreamed footings.

The commercially available computer finite element based code PLAXIS (2D

version 8/2002) was used. The following features within the code were used:

Axisymmetric was used due to the symmetry of the geometry and loading. Triangular elements with 15 nodes were used which provided fourth order

displacement field.

Interface elements were used to model differential movement between theconcrete and soil.

Varying KO values were input to model the horizontal locked in stressesoccurring in the different soil layers.

Volumetric contraction was used to model tension stresses induced at thesurface due to shrinkage of the expansive clays.

The tension cut off feature was used so that no effective stress tensionpoints occurred.

A cavitation cut off pressure of 2.1 ksf (100 kPa) was used to limit suctionpressures.

Consolidation was used to model the dissipation of pore pressures.The constitutive model chosen was the hardening soil model (Schanz et. al.,

1999) because it includes the following soil parameters to model real soil behavior:

Mohr-Coulomb failure criteria (c and ) Secant modulus for zero volumetric strains Hyperbolic relationship to model reduction in stiffness due to strain Constrained (odometer) modulus for volumetric strains Both secant and constrained modulus are stress dependent Unloading/reloading modulus Coefficient of permeability to model dissipation of pore pressures Effective stress parameters were input to model undrained and partially

drained conditions. The computer code automatically computes the pore

pressures based on the stiffness of the soil skeleton.

To model the load-displacement behavior of the footings, a detailed iterative

procedure was used whereby soil parameters were changed until good agreement

between measured versus FE predicted load-displacement behavior was obtained for allthe footings tested in compression and uplift at the U H test site. FE analysis was

performed on a desktop PC computer.


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Compression Load Test on F-1

Underreamed footing F-1 was installed at the U H site, and it was bearing at a

depth of 8 feet (2.4 m) in stiff clay. The 2.5 foot (0.8 m) diameter shaft wasunderreamed to a base diameter of 7.9 feet (2.4 m) without the use of drilling slurry.

The as-built details of the footing load test are shown in Fig. 3.

Fig. 3 Footing F-1Load Test Configuration

The reaction system consisted of two underreamed footings placed 9.5 feet (2.9

m) on each side of the test footing (center-to-center spacing). Their diameter was the

same as that of the test footing, but the bearing depth was increased to 18 feet (5.5 m)below grade. A pertinent observation was that groundwater seepage was large enough

that the reaction footings were installed using the slurry displacement method. Thus,

the permeability of the clay mass is much higher than that of intact clays.

Earth pressure cells were installed at the base of the footing to measure the stress

distribution. A 2.5 foot (0.8 m) diameter cardboard casing (sonatube) was set inside theshaft to act as a form to prevent shear from being transferred between the concrete shaft

and clay. However, the concrete rose up 0.75 feet (23 cm) into the annulus between the

cardboard casing and the soil forming a small collar around the shaft at the top of theunderream when concreting the footing.

The load was applied to the top of the shaft in 100 kips (445 kN) increments and

held for sixty minutes to a load of 500 kips (2224 kN) with a hydraulic jack.

Thereafter, the load was applied in 50 kips (222 kN) increments and held for thirty

minutes so that the failure load could be defined clearly. The footing had to beunloaded at the end of the 700 kips (3115 kN) load cycle to insert shims due to the large

displacement. Plunging failure occurred at a load of 750 kips (3334 kN) during thereload cycle. The footing was excavated after completion of the load test, and there was

no indications of structural failure.

Settlement was measured using dial gages attached to an independent reference

beam, and ground heave was measured using optical surveying methods. The earthpressure cells were read at the start and end of each load interval.


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Analysis of F-1 Load Test

During the initial FE study (Tand and ONeill, 2003), no attempt was made to

model the increasing displacement that occurred during each holding period due to theincreased base load resulting from dissipation of suction on the roof of the underream

because it was deduced that most of the suction had dissipated by the end of each

holding period. FE modeling was achieved by placing a small void directly above theunderream during the calculation phase to break any suction.

The initial calculations used soil parameters c', ' (CIU Bar tests), E' and KO

obtained from the U H site database. FE analysis was performed using 1967 triangular

elements, and undrained conditions were assumed. Detailed parameter studies were

performed to investigate the effects of various soil properties on the load-displacementbehavior. The FE computed load-displacement curves are compared to the field loading

test results in Fig. 4a.

(a) Database/Optimized Parameters (b) Effects of KO and E

Fig. 4 Load Displacement for Compression Footing F-1

FE calculations were made using KO = 1 for each layer to evaluate the influenceof the in situ horizontal stresses that occurred due to the geologic weathering process.

As shown in Fig. 4b, there was limited influence when the load is less than one-third

ultimate, but it was significant thereafter. Displacements are significantly increased,and the ultimate bearing capacity was reduced by about 25 percent. Increasing the soil

modulus of each layer by 50 percent indicated that displacements were reduced, but the

ultimate bearing capacity was unchanged. Simple 3D FE analysis indicates that the

influence of uplift from the reaction footings on the test footing was less than 5 percentwhen the loads were less than 50 percent of failure, and less than 10 percent at failure.

Additional FE analyses were performed to determine whether the load-

displacement behavior of F-1 during the loading intervals could be modeled if porepressures were allowed to build up and dissipate on the roof of the underream, and in

the clay mass below the footing during the holding intervals. The boundary conditions

and finite element mesh are shown in Fig. 5. The FE calculations were performed using


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2322 triangular elements. Consolidation calculations were made after each load interval

allowing the pore water pressures to partially dissipate.

Fig. 5 - Boundary Conditions and FE Mesh

As shown in Fig. 6a, the load-displacement behavior of the F-1 footing load test

is compared to the FE predictions. The optimized values of c', ', E' and KO are

compared with the soil parameters in the database in Table 1.

(a) Predicted/MeasuredBehavior (b) Suction above the Underream

Fig. 6 Modeling Field Loading Path for Footing F-1


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Table 1 - Summary of Soil Parameters at U H Test Site

U H Data Base* Optimized Parameters (FEM)

Soil

Layer

Depth

- ft

(m) KO

E

-ksf-(MPa)

c'

-ksf-(kPa)

'

KO

E**

-ksf-(MPa)

c'

-ksf-(kPa)

Kperm.

-ft/day-(m/day)

Ia0-3

(0-0.9)~3.0

230

(11)

0.2

(10)25 1.0

50

(2)

0.2

(10)15 Drained

Ib3-5

(0.9-1.5)~3.0

230

(11)

0.2

(10)25 2.0

100

(5)

0.2

(10)15

50.0

(15)

IIa5-7

(1.5-2.1)~3.0

230

(11)

0.2

(10)25 2.0

400

(19)

0.4

(19)20

1.0

(0.3)

IIb7-9

(2.1-2.7)~2.5

270

(13)

0.2

(10)25 3.0

500

(24)

0.7

(34)25

1 x 10-5

(.3 x 10-5

)

III 9-14(2.7-4.3) ~2.0 300(14) *** *** 2.5 600(29) 1.4(67) 30 1 x 10

-4

(.3 x 10-4

)

IV14-26

(4.3-7.9)~1.5

400

(19)0 25 2.0

500

(24)

0.7

(34)20

1 x 10-5

(.3 x 10-5

)

V

26-50

(7.9-

15.2)

~1.2

710

(34)

0.8

(38) 30 1.5

600

(29)

0.8

(38) 30

1 x 10-4

(.3 x 10-4

)

The following comments pertain to the items marked with asterisks in Table 1.

* c' and ' are from CIU Bar tests. See web database for CKOU Bar and

residual soil parameters.

** E is reported as 2 x the E modulus for comparison purposes only. TheE modulus is the secant modulus in the hardening soil model, and themobilized E is a function of the minor principal stress (3) and power

function (m).

*** Layer 3 is very stiff to hard sandy clay with sand seams and numerouscalcareous nodules. The clay appears to be lightly cemented with

calcium carbonate. This layer was not addressed in the web database.

FE analysis captured the general trend of the loading/holding intervals during

the field loading test. The initial displacements predicted by FE were slightly greater

than measured because the hyperbolic soil model does not consider small strain

stiffness or anisotropy. The permeability of the clays above the underreams, whichcontrols buildup and dissipation of suction, are physically changing during the load testdue to opening of fissures and slickensides and this cannot be modeled.

Analysis indicated that decreasing the permeability of each layer by a factor of

10 had little effect on the load-displacement behavior indicating that the permeabilityused for analysis is resulting in undrained failure of the bearing clays. However,

increasing permeability by a factor of 10 increased the stiffness of the load-

ref

50ref

50


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displacement behavior (Fig. 6a), and increased the bearing capacity by about 15 percent

due to partially drained conditions.The earth pressure cell measurements indicated that about 50 percent of the load

was carried above the base of the footing at about 50 percent of the ultimate bearing

capacity. Since a void had been cast around the footing shaft to break side shear, most

of this load would have been carried in suction on the roof of the underream. Maximumsuction inferred from the measurements was about 4 ksf (192 kPa), which is about 2

times greater than the maximum suction measured during the subsequent uplift load

tests (F-2/3/4) at the U H test site. The most plausible explanation for the high suctioninferred from the earth pressure cells is that the sand layer used to bed the earth pressure

cells was more compressible than the stiff clay surrounding the cells, and that a

disproportionate amount of the bearing stress was being carried by the stiff clay aroundthe cells (arching effect) until the sand was densified.

The earth pressure cell measurements did indicate that the base load increased

during the holding periods thus revealing that suction forces were building up on theroof of the underream when load was applied, and that suction was dissipating with

time. For example, the earth pressure cells indicated an increased base load of 14 kips(62 kN) during the 60 minute holding period at 400 kips (1780 kN). Thus, they infer

that 0.3 ksf (15 kPa) of suction on the roof of the underream dissipated during theholding period, which transferred load to the base.

FE predictions showing buildup and dissipation of suction pressures versus

displacement of the underream are shown on Fig. 6b. The predicted gap above theunderream started to form during the initial 100 kip (445 kN) loading interval. At the

400 kip (1780 kN) load, FE analysis predicts that 0.35 ksf (17 kPa) of suction

developed, and that it completely dissipated during the one hour holding period. TheFE predicted dissipation of suction correlates well with the 0.3 ksf (15 kP a) suction

inferred from the earth pressure cells. It should be noted that the buildup anddissipation of suction is a dynamic process because suction is a function of

displacement, and displacement is a function of base load which is also a function of

suction. FE analysis can model this complex process.FE analysis indicated that suction continued to build up on the roof of the

underream after the 400 kip (1,780 kN) load was applied, but it did not entirely

dissipate thereafter. At the failure load, analysis predicts that about 0.6 ksf (29 kPa) of

suction pressure was still present on the roof of the underream. Thus, the base load atfailure was about 25 kips (110 kN) less than the load applied to the top of the footing

(~3% less). Analysis of the earth pressure cells indicated that the base load was about

715 kips (3,180 kN) at failure which correlates well with the FE prediction of 725 kips(3,230 kN).

The roof of the underream was situated in a weather clay crust where thepermeability was high enough that most of the suction dissipated during the loading

intervals. This is the prime reason that good agreement was obtained when bothneglecting suction on the underream (Fig. 4a) and when including suction (Fig. 6a).

However, if the footing was situated below the water table and the rate of loading was

high, suction effects may be included in the design.


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Two cone penetration tests (CPT) next to the test footing detected a 5 foot (1.5

m) thick layer of very stiff to hard sandy clay about 1.5 feet (0.5 m) below the footing.FE parameter studies were performed to evaluate the effects of subsoil layering:

The thin layer of stiff clay (IIb) directly below the footing was replaced withthe underlying layer of hard sandy clay (III). Qu is computed as 833 kips

(3710 kN) at a displacement of 0.4 feet (S/B = 5%) indicating a 10 percentincrease in Qu.

All the clays below the footing were replaced with the layer of stiff clay(IIb). Qu is computed as 638 kips (2840 kN), indicating a 15 percentreduction in Qu.

All the clays below the footing were replaced with the layer of very stiff tohard sandy clay (III). Qu is computed as 910 kips (4050 kN), indicating an

increase of 20 percent in Qu.

Parameter studies thus indicate that subsoil layering had moderate influence onthe ultimate bearing capacity at this site. However, layering had only minor influence

on the initial slope of the load-displacement curves when loads were less than one-thirdultimate. The prime reason is that there were not large differences in the soil modulusbetween the clay layers.

FE analysis was performed to evaluate the effect of slickensides and fissures on

underreamed footings loaded in compression. As shown in Fig. 7, a random pattern of

slickensides and fissures within the soil mass was modeled using zero thicknessinterface elements. A stiff clay profile [cohesion (c') = 0.75 ksf (36 kPa) and internal

friction angle () = 20] with a constant soil modulus [E = 250 ksf (12 mPa)] and KO

(2.0) was chosen. Residual soil parameters [c' = 0.2 ksf (10 kPa) and ' = 12] were

input for the interface elements. A 6 feet (1.8 m) diameter underreamed footing bearingat depths of 8 feet (2.4 m) and 13 feet (4.0 m) was used. Undrained conditions were

assumed, and suction was allowed to build up on the roof of the underream.

Fig. 7 Use of Interface Element to Model Jointed Clay

ref

50


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As shown in Fig. 8, FE analysis predicted that the effects of jointing on the load-

displacement behavior was minor at loads less than one third of failure. Displacementswere increased thereafter, and the ultimate bearing capacity was reduced by about 6

percent for the shallow footing, and 2 percent for the deeper footing. Also shown in the

figures, there was about a 25 percent reduction in the ultimate bearing capacity when

reducing KO from 2.0 to 1.0 for the jointed clays.Yielding was taking place along numerous slickenside surfaces for the shallow

footing. However, it was only taking place along isolated slickensides in the deeper

footings due to higher confining pressures. FE analysis predicts that a cavity started toform above the roof of the underream for both footings when the initial load was

applied. The maximum suction pressure of 2.1 ksf (100 kPa) developed on the roof of

the underream after a top displacement of 0.6 inches (1.6 cm) for the deep footing, but 3inches (7.6 cm) of displacement was required to develop maximum suction for the

shallow footing. Thus, the base load for both footings was reduced about 53 kips (235

kN) by suction (~ 8%).

(a) Footings at 8 feet (2.4 m) (b) Footings at 13 feet (4.0 m)

Fig. 8 Effects of Jointed Clays for Footings in Compression

The 725 kip (3230 kN) base load at failure computed by FE analysis indicated a

net ultimate bearing pressure of 14.7 ksf (705 kPa). The equivalent undrained shearstrength of the jointed clay mass was computed to be about 1.9 ksf (91 kP a) assuming a

bearing capacity factor Nc = 7.7 (Skempton, 1951) for D/B (depth/diameter) = 1. For

comparison, the average undrained shear strength of the bearing clays below the footing(ONeill, 1983) was about 2.4 ksf (115 kPa). However, the equivalent shear strengthdoes not consider the effects of layering of the clays above or below the bearing

elevation. FE analysis can best be used to analyze geotechnical problems involving

multiple layers, and complex in situ stress conditions.


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Uplift Load Tests on F-2, F-3, and F-4

The load tests on footings F-2, F-3 and F-4 were performed as a series at the U

H test site to study the load-displacement behavior of underreamed footings tested inuplift (ONeill, et. al., 1985). Underreamed footing F-2 had a base diameter of 7.5 ft.

(2.3 m), and it was bearing at a depth of 8 ft. (2.4 m) in stiff clay. Footings F-3 and F-4

had base diameters of 6 ft. (1.8 m), and they were bearing at a depth of 10 ft. (3.0 m) instiff clay.

The reaction system consisted of two W 33 x 220 beams bearing on two timber

mats at the surface. The distance from the center of footing to edge of mat was 11 ft.

(3.4 m). The as built plan of the load test setup for footings F-3/4 is shown in Fig. 9.

Fig. 9 - F-3/4 Load Test Configuration

Two vibrating wire type piezometers were placed near the center of each footing

at the base to measure pore pressures. A pertinent observation is that relief wells had to

be installed to lower the water table so that the piezometers could be placed in the dry.Thus, the permeability of the clay mass is much higher than that of intact clays.

Settlement was measured using dial gages attached to an independent reference beam,

and ground heave by optical surveying methods. The piezometers were read at the startand end of each load interval when the equipment was functioning properly.

Footing F-2 [7.5 ft. (2.3 m) dia.] was load tested in steps of 20 kips (89 kN).

Failure occurred at 109 kips (485 kN) at an upward displacement of about 2.2 inches(5.6 cm). Each load was held for 5 minutes, and measurements were taken at the start

and end of each load cycle.

Footing F-3 [6.0 ft. (1.8 m) dia.] was load tested in steps of 25 kips (111 kN).

Failure occurred at 220 kips (979 kN) at an upward displacement of about 1.8 inches(4.6 cm). Each load was held for 5 minutes, and measurements were taken at the start

and end of each load cycle.


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Footing F-4 [6.0 ft. (1.8 m) dia.] was load tested over a 10 day period at load

steps of 40 kips (178 kN). Failure occurred at 190 kips (846 kN) at a deflection ofabout 0.8 inches (2.0 cm). Each load was held for a 2-day period, and settlement was

monitored periodically.

After completion of the initial loading cycle for each test, additional loading

cycles were then applied. The failure load increased during the additional load cyclesdue to build up of suction pressures that occurred with increasing displacement.

Analysis of F-2, F-3 and F-4 Load Tests

Footings F-3 and F-4 were both bearing at a depth of 10 feet (3 m) below grade,and both had underream diameters of 6 feet (1.8 m). However, F-3 was load tested to

failure in about 1 hour and F-4 was load tested to failure in about 10 days. The ultimate

capacity of F-3 was 220 kips (980 kN) versus 190 kips (845 kN) for F-4. The subsoilconditions indicated by the CPT tests indicate that reasonably uniform conditions

existed, and thus most of the differences are due to short term versus sustained loading.

This difference suggests that about 1.2 ksf (57 kPa) of suction built up on the base of F-

3. This is a simplified assumption because some of the additional 30 kips (134 kN) offorce could have been due to the difference between the partially undrained and drained

strength of the clays.

Both F-2 and F-3 were load tested in about 1 hour, but F-2 carried only half theuplift load of F-3 although the diameter was 25 percent greater. The reasons are due to

2 feet less embedment for F-2, and the fact that the roof of the underream was situated

in the weathered clay crust.

About 0.2 ksf (10 kPa) of suction was measured at an upward displacement ofabout 0.5 inch (1.3 cm) during the first loading cycle in F-3. The maximum measured

suction of 1.9 ksf (90 kPa) occurred during the second loading cycle for F-3 at an

upward displacement of about 7 inches (18 cm). As predicted by Boyles law, suctionis a function of displacement (volumetric expansion).

FE analysis was performed to determine whether the effects of suction on theuplift capacity could be modeled. The boundary conditions and finite element mesh for

typical analysis is shown in Fig. 5. The optimized soil parameters in Table 1 were the

input. The FE calculations were performed using 2214 triangular elements for F-2, and2404 elements for F-3 and F-4. Consolidation calculations were made after each load

interval allowing the pore pressures to dissipate.

The load-displacement behavior for footings F-2, F-3 and F-4 are compared in

Fig. 10a. The measured versus predicted load-settlement behavior for footings F-2, F-3,and F-4 are shown in Figs. 10b, 10c, and 10d, respectively. For comparison purposes,

undrained and drained loading are shown.As shown on Fig. 10a, the load-displacement behavior of footings F-3 (short

term test) and F-4 (sustained test) were very similar up to a load of 190 kips (845 kN).Thereafter, the load-settlement curve of footing F-3 was stiffer, and the ultimate uplift

capacity was about 15 percent greater.


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15

(a) Footings F-2/3/4 (b) Footing F-2

(c) Footing F-3 (d) Footing F-4

Fig. 10 Load-Displacement Behavior for Uplift Tests

As shown in Fig. 10b, the load-displacement behavior of footing F-2 (short term

test) is very similar to FE predictions for drained, not undrained, conditions. The primereason is that the footing is situated in the weathered clay crust where the permeability

is so high that dissipation of pore pressures approaches drained conditions.

As shown in Fig. 10c and 10d, the load-displacement behavior of footings F-3and F-4 is closer to FE predictions for undrained loading up to a load of 160 kips (720

kN). Thereafter, the load-displacement behavior was softer due to dissipation ofsuction. The relationship between undrained and drained conditions is complex for

jointed clays. It was concluded (ONeill, 1986) that air passages formed during theuplift tests, probably due to opening of the slickensides and fissures, which allowed

suction to dissipate rapidly. However, undrained conditions were probably present in

the intact clumps of clay between the slickensides and fissures. For FE modeling, thepermeability of layers IIb and III were input as 1.0 ft./day (0.3 m/day), and 0.5 ft./day

(0.2 m/day), respectively.


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Strength parameters very close to residual properties [c' = 0.2 ksf (10 kPa) and '

= 15] were required to model the weathered clay crust. KO =1.0 was used to computethe in situ stresses for this layer because the tests had been run in the summer when

shrinkage had reduced the horizontal stress. The permeability and the loading rate areimportant parameters because they control dissipation of pore pressures.

FE analysis was performed to evaluate the effect of slickensides and fissuresusing a random pattern of slickensides and fissures modeled using interface elements(Fig. 8a). Shown in Fig. 11a is the load-displacement response for a 6 foot (1.8 m)

diameter footing bearing at a depth of 8 feet (2.4 m), and on Fig. 11b for a footing

bearing at a depth of 13 ft. (4.0 m). FE analysis indicated that there was about a 10percent reduction in the uplift capacity due to the presence of slickensides and fissures

above the underreams for both embedment depths. Also, there was a 10 to 20 percent

reduction in the uplift capacity when KO was decreased from 2.0 to 1.0.

(a) Footings at 8 ft. (2.4 m)(b) Footings at 13 ft. (4.0 m)

Fig. 11 Effects of Jointed Clays on Footings in Uplift

Compression Test on F-5

Underreamed footing F-5 was installed at a test site for the Texas Department of

Transportation at the interchange of Highway 225 and Loop 610 . What is referred to

as shaft S-2 (ONeill et. al., 1972) had been underreamed to a diameter of 7.5 feet (2.4m). The load was applied in 50 kip (222 kN) load steps up to 500 kips (2225 kN) in 2.5

minute intervals, and thereafter the steps were increased to 200 kips (890 kN). Failure

occurred at a load of 1080 kips (4800 kN) at a displacement of 2.8 inches (7.1 cm).Settlement was measured using dial gages attached to an independent reference

beam, and they were read at the start and end of each load interval. The shaft wasinstrumented with 7 levels of Mustran cells (strain gages) to measure the load

distribution with depth, and 4 levels of telltails had been installed to measure

compression of the footing shaft.

The reaction system consisted of two underreamed footings placed 10 feet (3.0m) on each side of the test footing (center-to-center spacing). Their diameter was the


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same as that of the test footing, but the bearing depth was increased to 10 feet (3.0 m)

below the test footing. The as built details of the load test set up are typical to thoseshown in Fig. 3, except for the embedment depths.

Analysis of F-5 Load Test

The strain gages indicated that 190 kips (845 kN) was carried in side frictionalong the shaft. Suction occurring above the roof of the underream was not measured,

and thus the compression loads reaching the base during the load test can only be

estimated using FE analysis.

For initial FE calculations soil parameters c', ' (CIU Bar tests), E' and KO

obtained from the site data base were used. The measured value of 2.4 ksf (105 kPa)side shear from the strain gages was input for the maximum value in the interface

elements along the shaft to reduce the unknowns. Analysis was performed using 2288

triangular elements, and undrained conditions were assumed. In Fig. 12a the load-displacement behavior of F-5 is compared to the FE predictions. The optimized values

of c', ', E', and KO with values reported in the database are summarized in Table 2.

(a) Load-Displacement Behavior (b) Incremented Displacements

Fig. 12 Load-Displacement for Footing F-5 in Compression

It is surmised that undrained conditions existed during this test due to the rapid

rate of loading (~30 minute duration). The fact that the underream was bearing in clayabout 8 feet (2.4 m) below the water table suggests that suction close to the maximum

theoretical value of 2.1 ksf (100 kPa) probably developed. FE analysis predicts that the

maximum suction pressure of 2.1 ksf (100 kPa) build up on the roof of the underreamafter a load of 150 kips (670 kN) had been applied, and that it was present thereafter.

Thus, the base load at failure was reduced about 80 kips (356 kN) due to suction (~9%less).

The CPT test closest to F-5 showed that the strata of stiff silty clay with silt

layers was between 20 and 26 feet (6.1 and 7.9 m) below the 1969 grade. As shown inFig. 12b, FE analysis indicated that this layer is somewhat confined by the stiffer clay

layers above and below, and the silty clay is being squeezed out laterally.


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18

The 810 kip (3604 kN) base load at failure computed by FE analysis indicates an

ultimate bearing pressure of 18.3 ksf (876 kPa). The equivalent undrained shearstrength of the clay mass was computed to be about 2.1 ksf (100 kPa) assuming a

bearing capacity factor Nc = 8.8 (Skempton, 1951) for D/B (depth/diameter) = 3. For

comparison, the average undrained shear strength of the bearing clays below the footing

(ONeill, 1972) was about 2.0 ksf (96 kPa). However, the equivalent shear strengthdoes not consider effects of layering of the clays above or below the bearing elevation.

Table 2 - Summary of Soil Parameters

SH 225 and Loop 610 Interchange

Site Data Base Optimized Parameters (FEM)

Soil Layer

Depth

- ft

(m) KO

E

-ksf-

(MPa)

c'

-ksf-

(kPa)

'KO

E

-ksf-

(MPa)

c'

-ksf-

(kPa)

Ia 0 5(0-1.5)

2.5 375(18) 0.4(19) 23 1.0 50 .2(10) 15

Ib5 10

(1.5-3)1.7

375

(18)

0.4

(19) 23 3.0 400.5

(24) 20

II10 22

(3-6.7)1.3

375

(18)

0.4

(19) 23 1.5 400.6

(29) 20

III*22 26

(6.7-7.9)1.3

35

(2)0.3

(14)30 1.5 200

.5

(24) 30

IVa*26 30

(7.9-9.1)

1.3330

(16)0.5

(24)

30 1.2 6001.5

(72) 25

IVb*30 60

(9.1-18)1.1

360

(17)0.8

(37)30 1.0 600

2.0

(95) 25

The following comments pertain to the items marked with asterisks in Table 2.

* The effective stress parameters for layers III and IV were not reported in

the site database. These parameters were assumed from data at the U ofH site for similar subsoil conditions.

** E is reported as 2 x the E modulus for comparison purposes only. The

E modulus is the secant modulus in the hardening soil model, and the

mobilized E is a function of the minor principal stress (

3) and powerfunction (m).

Conclusions

The load-displacement behavior of two footings load tested in compression andthree in uplift was investigated. The complex subsoil profile and geologic stress history

of the Beaumont clays required numerical analysis to capture all the important features.

ref

50

ref

50ref

50


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19

Very good agreement was obtained between predicted load-displacement behavior and

the field load tests using FE analyses. The following conclusions are advanced fromthis study:

The slickensides and fissures within the clays were modeled usingrandomly distributed interface elements. There was about a 2 to 6 percent

reduction in bearing capacity for footings loaded in compression, but itwas about 10 percent for footings loaded in uplift.

Strength parameters for the weathered clay crust close to residual [c' = 0.2ksf (10 kPa) and ' = 15] were required to model the load-displacement

behavior for footings in uplift.

A cavity was formed above the roof (top) of the underreamed footingsduring the compression load tests which caused a build up of suction. Theultimate bearing capacity was increased by less than 5 percent for the

shallow footing F-1, and it was increased by about 10 percent for the deep

underreamed footing F-5.

Suction built up on the base of the underreamed footings during the upliftload tests. The uplift capacity of footing F-4 tested under sustained

loading was about 15 percent less than the capacity of footing F-3 testedunder short term loading. The magnitude of suction developed was a

function of the displacement, loading rate, and permeability of the clays.

KO was an important parameter in the modeling of the footings. The soilmodulus had a substantial effect on displacements, but it did not effect the

ultimate bearing capacity.


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Appendix - References

Al-Layla, M. T. H. (1970) Study of Certain Geotechnical Properties of Beaumont

Clay,Ph. D. Thesis Texas A & M University.

Bernard, H. A., LeBlanc, R. J., and Major, C. F., Recent and Pleistocene Geology of

Southeast Texas, Geology of the Gulf Coast and Central Texas and Guide Book ofExcursions, Houston Geological Society, 1962, pp. 175-224.

Kim, M. H., ONeill, M. W., and Kramer, S. L. (1999). Performance of Drilled Shafts

with Isolation Tubes in an Expansive Soil Environment, Geotechnical Testing

Journal, ASTM

Mahar, L. J., and M. W. ONeill, Geotechnical Characterization of Desiccated Clay,

Journal of Geotechnical Engineering, ASCE, January, 1983, pp. 56-71.

ONeill, M. W., and S. A. Sheikh, Geotechnical Behavior of Underreams in

Pleistocene Clay,Drilled Piers and Caissons, C. N. Baker, Jr., Editor, ASCE, May,

1985, pp. 57-75.

ONeill, M. W., and Reese, L. C., Behavior of Bored Piles in Beaumont Clay,

Journal of the Soil Mechanics and Foundations Division, ASCE, Vol. 98, No. SM2,

Feb. 1972, pp 195 213.

ONeill, M. W., and Sheikh, S. A. (1985). Geotechnical Behavior of Underreams in

Pleistocene Clay,Drilled Piers and Caissons II, ed. by C. N. Baker, Jr., ASCE,May, pp 57 75.

ONeill, M. W., and Yoon, G. (1995). Some Engineering Properties of Over-

consolidated Pleistocene Soil of the Texas Gulf Coast, Transportation Research

Record No. 1479, Transportation Research Board, Washington, DC, pp 81 88.

Schanz, T., Vermeer, P. A., and Bonnier, P. G. (1999). Formulation and Verification

of the Hardening Soil Model,Beyond 2000 in Computational Geotechnics,

Balkema, Rotterdam, pp 281 290.

Sheikh, S. A., and ONeill, M. W. (1988). Structural Behavior of 45-Degree

Underream Footings, Transportation Research Record 1119. Transportation

Research Board, Washington, DC, May pp 83 90.

Skempton, A. W., The Bearing Capacity of Clays,Proceedings, The Building

Research Congress (U.K.), Division I, 1951.

Tand, K. E., and ONeill, M. W. (2003), Comparison of Computed vs. Measured

Load/Settlement Response of a Footing Bearing on Stiff to Very Stiff Clay, PlaxisBulletin No. 14, pp 10-13.

Woodward-Clyde Consultants (1982), Study to Investigate the Performance of Skin

Friction on the Performance of Drilled Shafts in Cohesive Soils, Technical Report

GL-82-1, U. S. Army Engineer Waterways Experiment Station.

Yazdanbod, A., Sheikh, S. A., and ONeill, M. W. (1987). Uplift of Shallow

Underreams in Jointed Clay,Foundations for Transmission Line Towers, ASCEGeotechnical Special Publication No. 8, ed. by J-L Briaud, April, pp 110 127.

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