Plate Load Tests on Cemented Soil Layers Overlaying Weaker Soil

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Plate Load Tests on Cemented Soil Layers OverlayingWeaker Soil

Nilo Cesar Consoli, Ph.D.1; Francisco Dalla Rosa, M.Sc.2; and Anderson Fonini, M.Sc.3

Abstract: This paper addresses the interpretation of plate load tests bearing on double-layered systems formed by an artificially cementedcompacted top soil layer �three different top layers have been studied� overlaying a compressible residual soil stratum. Applied pressure-settlement behavior is observed for tests carried out using circular steel plates ranging from 0.30 to 0.60 m diameter on top of 0.15 to0.60-m-thick artificially cemented layers. The paper also stresses the need to express test results in terms of normalized pressure andsettlement—i.e., as pressure normalized by pressure at 3% settlement �p / p3%� versus settlement-to-diameter �� /D� ratio. In the range ofH /D �where H=thickness of the treated layer and D=diameter of the foundation� studied, up to 2.0, the final failure modes observed in thefield tests always involved punching through the top layer. In addition, the progressive failure processes in the compacted top layer alwaysinitiated by tensile fissures in the bottom of the layer. However, depending on the H /D ratio, the tensile cracking started in differentpositions. The footing bearing capacity analytical solution for layered cohesive-frictional soils appears to be quite adequate up to a H /Dvalue of about 1.0. Finally, for a given project, combining Vésic’s solution with results from one plate-loading test, it is possible �knowingof the demonstrated normalization of p / p3%-� /D, where the pressure-relative settlement curves for different H /D ratios produce a singlecurve for all values of H /D� to estimate the pressure-settlement curves for footings of different sizes on different thicknesses of acemented upper layer.

DOI: 10.1061/�ASCE�GT.1943-5606.0000158

CE Database subject headings: Shallow foundations; Soil cement; Layered soils; In situ tests.

Introduction

In Southern Brazil, urban and industrial developments often takeplace in terrain where the underlying soils are highly drainableweakly bonded residual soils with high void ratios. These residualsoils are usually partly saturated �degree of saturation of about80%� and even when saturated, any generated pore pressure israpidly dissipated. Hence, immediate settlement and not consoli-dation settlement is important for these soils. These soils cansuffer a high reduction in volume when the bonds are broken. Thesolution adopted for supporting building loads is usually a deepfoundation. However, for some projects, a deep foundation maynot be economically feasible. An example is construction of low-rise, low-cost housing or commercial buildings, where pilingcosts can represent an unacceptably large proportion of the totalinvestment. The present research studied an alternative to deepfoundations: shallow spread footings are placed on an upper layer

1Associate Professor, Dept. of Civil Engineering, Federal Univ. of RioGrande do Sul, Av. Osvaldo Aranha, 99, 3 andar, CEP 90035-190, PortoAlegre, Rio Grande do Sul, Brazil �corresponding author�. E-mail:[email protected]

2Ph.D. Student, Dept. of Civil Engineering, Federal Univ. of RioGrande do Sul, Av. Osvaldo Aranha, 99, 3 andar, CEP 90035-190, PortoAlegre, Rio Grande do Sul, Brazil. E-mail: [email protected]

3Ph.D. Student, Dept. of Civil Engineering, Federal Univ. of RioGrande do Sul, Av. Osvaldo Aranha, 99, 3 andar, CEP 90035-190, PortoAlegre, Rio Grande do Sul, Brazil. E-mail: [email protected]

Note. This manuscript was submitted on March 12, 2008; approvedon May 26, 2009; published online on May 27, 2009. Discussion periodopen until May 1, 2010; separate discussions must be submitted for indi-vidual papers. This paper is part of the Journal of Geotechnical andGeoenvironmental Engineering, Vol. 135, No. 12, December 1, 2009.
©ASCE, ISSN 1090-0241/2009/12-1846–1856/$25.00.
1846 / JOURNAL OF GEOTECHNICAL AND GEOENVIRONMENTAL ENGIN

J. Geotech. Geoenviron. Eng.

that has been mechanically improved by incorporation of an en-gineered cementing material. The approach produces a double-layer foundation system in which the upper layer has beenartificially cemented through mixing, compaction, and curing.

There have been many studies of shallow foundations on lay-ered systems, most of them concentrating on cases in whicha sand layer overlies a soft clay stratum �e.g., Tcheng 1957; Mey-erhof 1974; Burd and Frydman 1997; Kenny and Andrawes1997�. Only a few �e.g., Vésic 1975� dealt with cohesive-frictional soils. The accuracy of the Vésic �1975� solution in pre-dicting the ultimate bearing capacities of footings on anartificially cemented upper layer overlying a weakly bonded re-sidual soil layer with a high void ratio is unknown. The stress-strain-strength behavior of artificially cemented soils has beenstudied in the past by several investigators �e.g., Clough et al.1981; Coop and Atkinson 1993; Huang and Airey 1998; Consoliet al. 2000, 2003, 2006, 2007, 2009a,b,c,d; Dalla Rosa et al.2008�.

The results presented in this work are part of a comprehensivetesting program being run at the experimental site of the FederalUniversity of Rio Grande do Sul �UFRGS� in southern Brazil.Three different layers of compacted soil treated with differentcementitious materials were constructed over the weak residualsoil stratum. Shallow footings were placed on the artificially ce-mented layer and tested. The main purpose of the research is toinvestigate:• The effectiveness of using cementitious materials for improv-

ing the strength and stiffness of the top soil layer;• The possibility of normalization �in terms of dimensionless

variables� of the results to develop an overall understanding of
the different mechanisms of load transfer for shallow founda-
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tions bearing on layered systems overlaying residual soilstrata;

• The range of application and accuracy of the footing bearingcapacity analytical solution of Vésic �1975� for layeredcohesive-frictional soils for a range of studied H /D values;and

• The possibility of developing a methodology for assessing thepressure-settlement curve for footings of different sizes on dif-ferent thicknesses of a cemented upper layer.The results of the present research might also help the under-

standing of shallow foundations of offshore structures �conven-tional jack-up rigs and small gravity-based platforms� in stratifiedsoil profiles, consisting of cemented and uncemented sand and siltfound in many parts of the world, such as the calcareous profilesfound on the North-West Shelf of Australia.

Layered System and Site Characterization

The project involved building on the top of a residual soil stratumcompacted layers of three different cemented soil mixtures: sandysoil to which coal bottom ash and carbide lime layers had beenadded, coal bottom ash with Portland cement, and local residualsoil mixed with Portland cement. The following section presentsthe characteristics of the residual soil site and the improved upperlayers.

Table 1. Grain-Size Distribution and Atterberg Limits of the ResidualSoil Layer

Sandy siltred clay

Nonplasticsilt sand

Bottomash

Medium sand �%� 4 9 13

Fine sand �%� 30 47 55

Silt �%� 26 39 30

Clay �%� 40 5 2

wL �%� 43 22 —

wP �%� 22 19 —

Gs 2.67 2.70 2.08

Fig. 1. Grain-size distribution curves

JOURNAL OF GEOTECHNICAL AND GEOE


Residual Soil Stratum

The characteristics of the residual soil at the UFRGS experimen-tal site have been determined through an extensive testing pro-gram that consisted of a combination of laboratory tests, mainlytriaxial tests and in situ tests. The in situ tests included conepenetration tests �CPTs�, standard penetration test, cross-holetests, down-hole tests, dilatometer tests, Ménard pressuremetertests at different depths, and plate load tests using both steel platesand concrete footings. The site investigation revealed a 3.5-m-thick upper layer of homogeneous sandy-silty red clay, which isclassified as low plasticity clay �CL� according to the Unified SoilClassification System. A grain size distribution curve for this re-sidual soil is given in Fig. 1. The grain size distribution, as well asthe Atterberg limits and specific gravity of the solids, are pre-sented in Table 1. The average bulk unit weight was about17.5 kN /m3; the water content was typically 25.0% �suction val-ues were measured in situ and revealed maximum values of about8 kPa�; and the degree of saturation was around 80%. Unliketransported soils with similar grain size distribution, this residualsoil is the product of in situ weathering, which decreased the unitweight, increased the porosity and exhibit parent rock featuressuch as interparticle bonding. Saturated drained conventional tri-axial tests were carried out on cylindrical �50 mm in diameter and100-mm high� specimens at effective confining pressures of 20,60, and 100 kPa �see Fig. 2�a��. The peak friction angle �� forthe in situ residual soil specimens was 26° and peak cohesionintercept �c�� about 17 kPa.

The soil profile to a depth of 4 m of the test site is presented inFig. 3, where measurements of cone tip resistance qc and thevalues of the maximum shear modulus G0 obtained from cross-hole and down-hole tests are presented. It is clearly observed thatboth the CPT and shear moduli are fairly constant, which indi-cates that the homogeneity of the upper soil layer is not restrictedto strength, but can also be extended to deformation behavior.Further information on the laboratory and in situ characterizationof the residual soil layer can be found in previous work by Con-soli et al. �1998�.

Top Cemented Layers

Before work was done on compacting the soil-cemented top lay-ers, a 1.0-m-thick layer of the residual soil was removed through-out the testing area. After removal, several improved soil areas,each 2.5 m�2.5 m in plan, were constructed as test areas. Theprepared test areas consisted of areas with 150-mm-thick layers,300-mm-thick layers, and 600-mm-thick layers. Material charac-teristics, mixture proportions and compaction parameters are de-scribed below.

MaterialsThe sandy soil samples used in the present study were obtainedfrom the region of Porto Alegre, in southern Brazil. The soil isclassified as nonplastic silty sand �SM� according to the UnifiedSoil Classification System. The grain size distribution, as well asthe Atterberg limits and specific gravity of the solids are presentedin Table 1.

The bottom ash selected �type F according to ASTM C618�ASTM 1998a�� for mixing with the sandy soil is a pozzolanicmaterial residue of burning coal obtained in a thermal power sta-tion, located close to Porto Alegre. The grain size distribution and
specific gravity of the solids are presented in Table 1. The mate-
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rial is classified as nonplastic silty sand �SM�. Fig. 1 also includesgrain size curves for both the sandy soil and the bottom ash.

Carbide lime, a by-product of the manufacture of acetylenegas, obtained from White Martins Industrial Gas plant also lo-cated near Porto Alegre, was used in part of this investigation asthe cementing agent. A chemical analysis showed that the carbidelime is 97% CaO and 3% CaCO3. Standard Portland cement wasalso used in part of the investigation as the cementing agent.

Mixture Proportions and Compaction ParametersThe three improved top layers were prepared using a rotatingdrum mixer by mixing air-dried soil and/or bottom ash, cementi-tious material �Portland cement or carbide lime�, and water.

The first mixture contained equal weights of sandy soil andcoal bottom ash, to which was added 5% carbide lime �by weightof dry soil plus bottom ash�. Water was added to obtain an overallwater content of 26%, which corresponds to optimum moisture

Fig. 2. Consolidated drained triaxial test responses for �a� residualsoil; �b� compacted sandy soil-coal bottom ash-carbide lime; �c� coalbottom ash-cement; and �d� local residual soil-cement field specimens

content under standard Proctor energy. The amount of bottom ash



used in this mixture lies above the range suggested by the Na-tional Cooperative Highway Research Program �NCHRP� �1976�for fly ashes and was selected according to regional practice. Theminimum carbide lime content obtained was based on pH changesfollowing the method proposed by Rogers et al. �1997� for initialconsumption of lime.

The layer of this mixture was built over the residual soil inconsecutive sublayers, each 100-mm thick, using tamping equip-ment to reach a dry unit weight of 12.5 kN /m3 �density ratio of52%� at optimum moisture content. This layer was allowed tocure for 28 days before being loaded with model footings. Con-ventional static drained triaxial tests on specimens of this layerretrieved from the field after 28 days of curing were also carriedout under full saturation at effective confining pressures of 20, 50,and 80 kPa �see Fig. 2�b��. The peak friction angle for this mix-ture was 47° with peak cohesion intercept of about 88.5 kPa.

The second mixture consisted of coal bottom ash, to whichwas added 5% by dry weight of Portland cement. Water wasadded to obtain an overall water content of 44%, which corre-sponds to optimum moisture content under standard Proctor en-ergy. The amount of cement was selected according to regionalpractice. The layer of this mixture was built over the residual soilin consecutive sublayers, each 100-mm thick, in the same way asfor the previous mixture, using tamping equipment to reach a dryunit weight of 9.9 kN /m3 �density ratio of 47.5%� at optimummoisture content. This layer was allowed to cure for 28 daysbefore being loaded with model footings. Conventional staticdrained triaxial tests on specimens retrieved from the field after28 days of curing were also carried out under full saturationat effective confining pressures of 20, 60, and 100 kPa �seeFig. 2�c��. The peak friction angle for this material was 44° witha peak cohesion intercept of approximately 125 kPa.

The third mixture consisted of the local residual soil, to whichwas added 5% by dry weight of Portland cement. Water wasadded to obtain an overall water content of 20%, which corre-sponds to optimum moisture content under standard Proctor en-ergy. Initial trials showed that the plasticity of the residual soilmade it difficult to mix uniformly and compact to the requireddensity. The layer of this mixture was built over the residual soilstratum in consecutive layers, each 100-mm thick, in the sameway as for the previous mixtures, using tamping equipment to

3

Fig. 3. Soil profile to depth of 4 m containing �a� CPT tip resistance;�b� cross-hole and down-hole initial shear moduli �G0� values

reach a dry unit weight of 15.9 kN /m �density ratio of 59%� at


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optimum moisture content. This layer was allowed to cure for 28days before being loaded with model footings. Static drained con-ventional triaxial tests on specimens retrieved from the field after28 days of curing were also carried out under full saturation ateffective confining pressures of 20, 40, and 60 kPa �see Fig. 2�d��.The peak friction angle for this material was 47° with a peakcohesion intercept of approximately 29 kPa. Visual and micro-scopic examination of exhumed field specimens of the three toplayers showed that the mixtures were satisfactorily uniform.

Field-Testing Program

The field-testing program was carried out at the experimental sitedescribed previously. Plate load tests were conducted using tworigid circular steel plates, one 300 mm in diameter and 50.8-mmthick, and the other 600-mm diameter and 101.6-mm thick. Thesetup used for conducting the plate load tests was in accordancewith ASTM D1194 �ASTM �1998b��. The load was appliedthrough a system comprising a hydraulic jack, a reaction beam, aloading platform, and a calibrated load cell. Three dial gaugeswith divisions of 0.01 and 50 mm travel were used for settlementmeasurement. Measurements were made directly on the plate, aswell as on the top of the treated layer adjacent to the plate. Settle-ments on the top of the plate were taken at three points, separated120° from each other. The gauges were fixed to a reference beamand supported on external rods. The load was applied in equalincrements of not more than one-tenth of the estimated ultimatebearing capacity. For each load increment, measurement of settle-ments was made at the following fixed times: 0, 30 s, 1, 2, 4, 8,15, 30, 60, and 120 min. In accordance with Brazilian StandardNBR-12131 �Brazilian Standards Association 1991� and ASTMD1194 �ASTM 1998b�, each increment was maintained for aminimum of 30 min, and until the following criterion wasachieved:

Ln − Ln−1 � 0.05 � �Ln − L1� �1�

where Ln=average dial gauge reading at a specified time intervalt; Ln−1=average dial gauge reading immediately previous to Ln;and L1=first reading of the stage of loading taken just after stageloading application.

Results of the plate load tests have been presented using thenotation PLTDxHy to identify each test, where “PLT” stands forcircular steel plate, “D” refers to plate diameter, “x” is the platediameter in centimeters, “H” refers to the thickness of treated toplayer, and “y” is the thickness of the treated layer in centimeters.

Finite-Element Analysis

In order to investigate the failure mechanisms in the plate tests,and to extend the normalization of the results to larger plate sizes,a two-dimensional axisymmetric finite element study of the platetest on layered soil was undertaken using the PLAXIS code. Thisinvolved a double-layer system with a cemented top layer ofthickness H, and plate of diameter D. Studies were carried outfor four different H /D ratios: 0.25, 0.5, 1.0, and 2.0, involvingH=0.30, 0.60, and 1.20 m, and D=0.30, 0.60, 1.20, and 2.40 m.An elastic-plastic constitutive model with the Drucker-Prageryield criterion was used for both soils. All simulations used initialstresses that assumed at-rest coefficient �Ko� equal to unityfor both soil layers. The unit weight was fixed at 17.5 kN /m3 for

3
the local residual soil layer and 15.75 kN /m for the cemented


top layer. The model parameters used in the modeling were basedon the typical triaxial tests presented in Fig. 2, i.e., c�=17 kPa,��=26°, �=0°, E=22.5 MPa, and �=0.25 for the local residualsoil layer and c�=88.5 kPa, ��=47°, �=16°, E=600 MPa, and�=0.25 for the top cemented layer �corresponding to the first-mixture studied�. Note that this gives tensile strengths ofc� / �tan �� equal to 35 and 83 kPa for the residual soil and ce-mented soil, respectively. The plate loading was simulated using aconstant displacement condition for the area loaded by the plate.The finite element �FE� domain was extended to 16D in the hori-zontal direction, and 16D in the vertical direction.

Test Results and Analysis

Applied Pressure-Settlement Response in the PlateLoad Tests

Fig. 4 shows plots of applied pressure versus plate settlement forfive different loading cases involving plates of different sizes onthe surface of the compacted layer of the first mixture �sandysoil-coal bottom ash-carbide lime�. As already explained, thelegend indicates the plate diameter �30 or 60 cm� and the thick-ness of the improved layer �15, 30, or 60 cm�. Fig. 5 shows thecorresponding plots for plate-loading tests on the compacted layerof the second mixture �coal bottom ash-Portland cement�, whileFig. 6 shows the corresponding plots for tests involving the thirdmixture �local residual soil-Portland cement�.

For the tests shown in Fig. 4, failure load was reached in allfive loading tests and thus in each case the maximum load-bearing capacity was well defined. Once the failure load had beenreached, it was necessary to pump the hydraulic jack continuouslyin order to keep the load constant as settlement continued. Thisproduced large further displacements �settlements�. In the case of

Fig. 4. Pressure-settlement curves of loading tests on compactedsandy soil-coal bottom ash-carbide lime top layer overlying residualsoil stratum

loading tests carried out on top of the compacted coal bottom


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ash-Portland cement �Fig. 5� and local residual soil-Portland ce-ment �Fig. 6� failure load was reached in most of the tests.

Normalization of Curves of Applied Pressureversus Settlement

It is often helpful to present relationships in terms of dimension-less variables so that advantage can be taken of the scaling effectsof continuum mechanics �Wroth 1988�. Normalization of the re-sults of plate-loading tests can establish general tendencies inrelation to different plate diameters and the thicknesses of the

Fig. 5. Pressure-settlement curves of loading tests on compacted coalbottom ash-cement mixture overlying residual soil stratum

Fig. 6. Pressure-settlement curves of loading tests on compactedlocal residual soil-cement mixture overlying residual soil stratum



cemented layer. Such an approach has been used with data fromFig. 4. The first step toward normalization involves plotting theapplied footing pressure against the settlement-to-diameter ratio�� /D� for each footing. The resulting plots for the tests on thecompacted layer of sandy soil, coal bottom ash and carbide limeare shown in Fig. 7. In such figure, results for plate bearing testson the unimproved local residual soil �Consoli et al. 1998� areincluded, for comparison. These plots are labeled PLTD30H00

Fig. 7. Pressure versus settlement �� to diameter �d� ratio on com-pacted sandy soil-coal bottom ash-carbide lime top layer overlyingresidual soil stratum

Fig. 8. Pressure versus settlement �� to diameter �d� ratio curvesfrom numerical simulations for varying geometrical parameters Dand H �FE analysis considered parameters of the first-mixturestudied—sandy soil, coal bottom ash and carbide lime—as the com-pacted cemented top layer�


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and PLTD60H00, indicating plate diameters of 30 and 60 cm,respectively, and improved layer thickness of zero in both cases.Fig. 7 suggests that the test results appear to be related strongly tothe H /D ratio, with some amount of scatter: as H /D increases,the behavior becomes stiffer and stronger. The correspondentplots for the tests on the layer of compacted coal bottom ash-Portland cement and tests on the local residual soil-Portland ce-ment present similar trends and are not presented for concisenessreason.

Following a similar normalization argument, Fig. 8 presentsresults from the two-dimensional axisymmetric FE studymentioned earlier. For different thicknesses �H=0.30, 0.60, and1.20 m� of the cemented layer and different values of plate diam-eter �D=0.30, 0.60, 1.20, and 2.40 m�, unique curves are obtainedfor each H /D ratio when the results are plotted as applied pres-sure versus � /D. The results of the numerical simulations confirmtrends shown by the field testing and extend the normalization forlarger plate diameters. Figs. 9�a–d� present realistic agreementbetween calculated and measured pressure—�� /D� curves �FEanalysis considered parameters of the first-mixture studied—

Fig. 9. Calculated and measured pressure versus settlement �� to dcoal bottom ash and carbide lime—as the compacted cemented top l

sandy soil, coal bottom ash and carbide lime—as the compacted



top cemented layer� for H /D=0.25, 0.5, 1.0, and 2.0, re-spectively. The limit pressure �p3%� was satisfactorily predicted�error of about 25%� using the FE method, except for the H /D=0.25 in which bearing capacity was overestimated by about75%. At working loads, however, using a factor of safety of three,there is a reasonable, and probably acceptable, agreement be-tween the estimated and the observed settlements for all studiedcases.

A further normalization of the results from numerical simula-tions was obtained by dividing the pressure �p� at any value ofsettlement by the value of the corresponding limit pressure de-fined as the pressure at � /D=3% �p3%�. Berardi and Lancellotta�1991� analyzed the behavior of more than 200 shallow founda-tions, and found that maximum displacements at working pres-sures were generally of the order of 1%. Thus, the � /D criterion�3%� used here to define the maximum �or limit� pressure is threetimes the value of � /D observed by Berardi and Lancellotta�1991� at working pressure. Values of p / p3% can then be plottedagainst relative settlement �� /D�. When this is done �Fig. 10�, thesimulated pressure-relative settlement curves in Fig. 8 for differ-

r �d� ratio curves �considering the first-mixture studied—sandy soil,or �a� H /D=0.25; �b� H /D=0.5; �c� H /D=1.0; and �d� H /D=2.0

iameteayer� f

ent H /D ratios produce a single curve for all values of H /D. The


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results in Fig. 10 are for the same set of soil parameters as usedfor Fig. 8. Values of p / p3% for the field tests are plotted against� /D on Fig. 11 corresponding to the plots of pressure versus � /Din Fig. 7. Results from field plate-loading tests for the differentH /D ratios plot closely together, with their associated variability,normalizing in the same way as the simulated results. This indi-cates that the normalized curves can be applied to all footingdiameters and thickness of cemented layer within the range ofH /D from 0.25 to 2. This suggests a means of extrapolating theresults of a single plate-loading test to a design process for full-

Fig. 10. Normalization �p / p3% versus settlement �� to diameter �d�ratio� of the numerical simulations for different H /D relations pre-sented in Fig. 8

Fig. 11. Normalization using pressure divided by pressure at 3% ofrelative settlement �� /D� on compacted sandy soil-coal bottom ash-carbide lime top layer overlying residual soil stratum



scale footings on the same cemented upper material and lowercompressible foundation soil. However, what is still needed fordesign purposes are values of the limit pressure for the particulargeometries being considered.

Progressive Failure Mechanism

During the plate load tests, cracks developed on the top of thetreated layer. The cracks were concentric to the plate and firstbegan to appear at about 70% of the final failure load. Whenthe failure load was reached, large settlements of the plate oc-curred, which produced penetration of the plate into the treatedlayers. The failure surfaces started close to the outer edge of thefooting. Vertical boreholes excavated below the plates after test-ing identified similar “punching failure” mechanisms of failure.Figs. 12�a and b� present photographs of vertical surfaces exca-vated just below two separate test plates. The photographs showobserved failure mechanisms for H /D=0.25 and H /D=1.0, re-spectively, representing cracking and fissuring observed in suchtests.

Fig. 12�b�, for H /D=1.0, also shows some cracks below the

Fig. 12. Photos of the failure mechanism observed in the plate loadtests �a� H /D=0.25; �b� H /D=1.0

center of the plate extending upwards from the bottom but not


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reaching the top of the treated layer. No fissures were observedbelow the plate in Fig. 12�a�, in which H /D=0.25, except directlybeneath the foundation perimeter.

In order to determine the patterns of how progressive failurestarts, more results of the aforementioned axisymmetric FEstudy are presented and discussed. Contours of the horizontal ten-sile stress distribution obtained in the numerical analyses forH /D=0.25 and 1.0 are presented in Figs. 13�a and b�, respec-tively. The initial failure process starts with horizontal tensilestresses �and corresponding strains� at the bottom of the treatedlayer �see in Figs. 13�a and b� the development of high horizontaltensile stresses in distinct regions below the plate �depending onH /D value��. Once the tensile strength is reached, the strains

Table 2. Soil Parameters Used for the Three Different Cemented Soil L

Lower layer �weak residual soil�

Cemented top layer 1 �compacted sandy soil-coal bottom ash-carbide lim

Cemented top layer 2 �compacted coal bottom ash-cement mixture�

Cemented top layer 3 �compacted local residual soil-cement mixture�

Fig. 13. Horizontal tensile stresses �results of numerical analysis� inthe top cemented layer for �a� H /D=0.25; �b� H /D=1.0



cause opening of fractures. It should be pointed out that the FEmodel does not develop fractures, and so cannot really model thewhole process. In the FE model, once tensile failure is reached, aconstant tensile stress is maintained in this region, whereas inreality, tensile failure in a brittle material results in 0 tensile stressbeing reached across the fracture, and a stress redistribution oc-curs. Thus, the FE model can really only show where tensilefailure initiates, but cannot then follow through to allow the fullpunching mechanism to be investigated explicitly.

When H /D=0.25 in the FE analysis, the fractures initiate atthe bottom of the cemented layer under the edge of the plate, ascan be seen in Fig. 13�a�. After this initial crack, it is postulatedthat the fractures tend to extend upwards in the cemented layer.As the pressures applied by the plate increase, a second failureprocess development, which depends on shear stresses rather thantensile stress, is proposed. High shear stresses develop in the stifflayer just below the edge of the test plate. Once they reach theshear strength of the layer, a failure shear surface is postulatedto progress downwards. As loading on the test plate increases,the tensile failure at the bottom and the shear failure at the topevolve into a single vertical failure surface, producing a punchingfailure. Such a proposed failure mechanism is coherent with thefailure patterns observed in the photograph of the field test shownin Fig. 12�a�. When H /D=1.0, the suggested processes leadingto localized tensile and shear failure are the same as those pro-posed in the previous case. However, the greatest horizontal ten-sile stress is first observed in the bottom of the cemented layerbelow the center of the plate, as can be seen in the horizontalstresses distribution obtained in the numerical analyses consider-ing H /D=1.0, as presented in Fig. 13�b�. As loads increase fur-ther, new fractures are postulated to initiate outside the initialfracture, where the horizontal tensile stress are also high. Shearstresses at the contact between the outer edge of the test plate andthe stiff top layer are supposed to start inclined shear failure sur-faces directed downwards and outwards. As the loading continuesto increase, it is postulated that the two sets of fractures evolveinto a single punching failure. Such a proposed failure mechanismis consistent with the failure patterns observed in the field loadingtests �Fig. 12�b��.

Bearing Capacity Analysis on StratifiedCohesive-Frictional Soil Profile

The overall response of shallow foundations placed on weak re-sidual soil can be significantly improved by building a top layerof compacted cemented soil. In these cases, existing analyticalsolutions for layered cohesive-frictional soils �e.g., Vésic �1975��might effectively be employed to predict bearing capacity.

Vésic �1975� assumes that the footing failure surface is verti-cal and follows the foundation perimeter. According to his analy-sis, the bearing capacity �q� of a foundation resting on an upper,stronger layer �compacted cemented soil in the present study�

as well as for the Weak Residual Soil Stratum

Cohesion intercept�kPa�

Friction angle�degrees�

Bulk unit weight�kN /m3�

17 26 17.5

ure� 88.5 47 15.75

125 44 14.25

29 47 19.1

ayers,

e mixt


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with strength parameters c1� and �1� and underlain by a lower layer�weak residual soil in the present study� with strength parametersc2� and �2� is

q = �q0 + � 1

K�c1� cot �1��e4K tan �1��H/D� − � 1

K�c1� cot �1� �2�

where q0=bearing capacity of a fictitious footing of the same sizeand shape as the actual footing, but resting on the top of the lowerlayer K=1+sin2 �1� /1+sin2 �1�; H=thickness of the treated layer;and D=diameter of the plate. Following Terzaghi’s recommenda-tions for punching failure, reduced shear strength parameters wereused for the lower layer of weak residual soil, as suggested byConsoli et al. �1998�.

This solution was used to estimate the limit pressure defined asthe pressure at � /D=3% �p3%� for the plate bearing tests on thethree different improved layers already discussed. The parametersused in the present study for the three cemented layers, and theunderlying residual soil stratum are presented in Table 2. Table 3

Table 3. Geometrical Characteristics of Treated Layers and Field and A

Thicknessof treatedlayer H

�m�

Platediameter D

�m� H /D n

Cemented top layer 1�compacted sandysoil-coal bottomash-carbide lime mixture�

0.15 0.30 0.5 P

0.30 0.30 1.0 P

0.60 0.30 2.0 P

0.15 0.60 0.25 P

0.30 0.60 0.5 P

Cemented top layer 2�compacted coalbottom ash-cement mixture�

0.15 0.30 0.5 P

0.30 0.30 1.0 P

0.30 0.60 0.5 P

Cemented top layer 3�compacted local residualsoil-cement mixture�

0.15 0.30 0.5 P

0.30 0.30 1.0 P

0.30 0.60 0.5 P

Fig. 14. Footing pressure versus settlement �� to diameter �d� ratio



summarizes the results of these calculations. The limit pressure�p3%� was satisfactorily predicted by this method for H /D�1.0,with the solution giving bearing capacities ranging from 62.8 to101.6% of the measured field values �limit pressures obtainedaccording to the criterion established in the present research�.However, the method overestimated the bearing capacity for theone field test with H /D�1.0 �i.e., PLTD30H60, with H /D=2.0�by more than 200%. Therefore, there is still a need for furtherresearch in order to produce reliable bearing capacity solutionsfor layered cohesive-frictional soil for design applications forH /D values greater than unity.

Assessment of the Pressure-Settlement Curveof Footings on Layered Soil with a CementedTop Layer

Once the value of the limit pressure has been successfully deter-mined using the Vésic �1975� method �for H /D up to 1.0�, and

al �Vésic 1975� Limit Pressures

testlature

Limit pressureplate test �p3%�

�kPa�

Bearing capacityVésic �1975�

�kPa�

Vésic �1975�to field limit

pressure�%�

0H15 912 693 76.0

0H30 2,015 1,655 82.1

0H60 3,221 7,271 225.7

0H15 432 439 101.6

0H30 825 746 90.4

0H15 1,296 814 62.8

0H30 Did not reach � /D=3% 2,025 —

0H30 Did not reach � /D=3% 863 —

0H15 730 528 72.3

0H30 1,145 1,160 101.3

0H30 Did not reach � /D=3% 575 —

mented silt overlying silty soil stratum �adapted from Finnie �1993��

nalytic

Plateomenc

LTD3

LTD3

LTD3

LTD6

LTD6

LTD3

LTD3

LTD6

LTD3

LTD3

LTD6

on ce


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knowing of the substantiated normalization of p / p3%-� /D, wherethe pressure-relative settlement curves for different H /D ratiosproduce a single curve for all values of H /D, it becomes possibleto estimate the complete pressure �p�—settlement �� curves forfootings of different sizes bearing on different thicknesses ofcemented upper layers. So, for a specific footing size �D� and agiven thickness of cemented upper layer �H�, and the representa-tive curve of p / p3% versus � /D obtained from a single plate-loading test for the particular project being examined, it ispossible to obtain the particular p-� curve determining p3% �as-sessed in the present study using Vésic �1975� method� forthe representative H /D and then plot the curve p-� by substitut-ing the values of p3% and D for their specific values in thep / p3%-� /D curve.

Comparison with Model Tests on Layered Soilsin Centrifuge Tests

Finnie �1993� carried out model footing tests on a geotechnicalcentrifuge, which, at the acceleration level used in the tests,represent prototype circular footings with diameter ranging from3 to 10 m. The soil profile consisted of a medium strength crust,with a prototype thickness of 2 m and a cone tip resistance ofabout 2.6 Mpa, over a much weaker layer of silty calcareous soil.The results of three tests at three different H /D ratios �0.20, 0.50,and 0.67� are shown in Fig. 14, plotted as footing pressure �p�against normalized settlement �� /D�. The legend indicates thediameter D and the thickness of the cemented layer H in cm.Fig. 14 shows the stiffness of the p-� /D relationship increasingwith increasing H /D values from 0.20 to 0.67, similar to theresults observed in the present study.

The results in Fig. 14 are replotted in Fig. 15 consideringnormalized pressure �p / p3%� plotted against � /D. As an outcome,roughly a unique curve was found in Fig. 15 for these three valuesof H /D �H /D=0.20, 0.50, and 0.67�, corroborating the results

Fig. 15. Footings normalization using pressure divided by pressurestratum

obtained in the present study.



Conclusions

Results have been presented from field loading tests carried outon two-layered systems consisting of a compacted cemented layeroverlying weaker residual soil. The cemented layer consisted ofmixtures of:• Sandy soil, coal bottom ash and carbide lime;• Coal bottom ash and Portland cement; and• Local residual soil and Portland cement.

The outcomes from this work can be summarized as follows:• For a given layered profile, expressing the results as normal-

ized pressure �p / p3%� versus settlement-to-diameter �� /D�ratio produces a unique relationship.

• The failure modes observed in the field tests always involvedpunching through the top layer. The progressive failure pro-cesses of the compacted top layer always initiated by tensilefissures in the bottom of the layer, and depending on the ratioH /D value, the tensile cracking started in different positions.

• Good estimates of the bearing capacity could be obtainedusing the method of Vésic �1975� for H /D�1.0. Therefore,this method, when combined with results from one plate-loading test, make it possible �knowing that p / p3% versus � /Dproduces a unique relationship� to estimate the pressure-settlement curve for footings of different sizes on differentthicknesses of a cemented upper layer.

• The normalization suggested here was also found to apply to aseries of model footing tests on the centrifuge carried out byFinnie �1993�. These tests represented large diameter �2- to10-m diameter� footings resting on a cemented crust �H /Dfrom 0.2 to 0.67� overlying weaker silty calcareous soil.

• This suggests that the results of the present research mightalso help the understanding of shallow foundations of bothonshore and offshore structures �conventional jack-up rigs andsmall gravity-based platforms� in upper cemented stratifiedsoil profiles considering a large range of cemented top layer

of relative settlement �� /D� on cemented silt overlying silty soil
at 3%
strengths.


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Acknowledgments

The writers express their gratitude to the Brazilian Agency forElectrical Energy �ANEEL-Project P&D CEEE-GT/UFRGSGrant No. 9936455� and Brazilian Council for Scientific andTechnological Research/Brazilian Ministry of Science and Tech-nology �CNPq/MCT–Projects Produtividade em Pesquisa GrantNo. 301869/2007-3, Edital Universal Grant No. 472851/2008-0,and PNPD Grant No. 558474/2008-0� for the financial support tothe research group. The writers would like to thank the anony-mous reviewers for their insightful comments and suggestionsthat improved the content of this paper.

Notation

The following symbols are used in this paper:c� � effective cohesive intercept;D � plate diameter;E � Young’s modulus of elasticity;

Gs � specific gravity of the solids;H � thickness of treated layer;

Ko � at-rest coefficient;Ln � average dial gauge reading at specified time

interval t;Ln−1 � immediately previous average dial gauge

reading to Ln;L1 � first dial gauge reading of the stage of

loading;p � pressure applied to the plate;

p3% � limit pressure: pressure at � /D=3%;q � bearing capacity;

q0 � bearing capacity of a fictitious footing of thesame size and shape as the actual footing, butresting on the top of the lower layer;

wL and wP � Atterberg limits �liquid limit, plastic limit�;� � settlement;

� /D � relative settlement �settlement to diameterratio�;

a � axial strain;v � volumetric strain;� � Poisson’s ratio;

a � axial stress;r � radial stress;t � tensile stress;� � angle of dilation; and

�� effective friction angle.

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