混凝土单撑着墙在性能挖掘和冻结期间保留土壤

Performance of a single-propped wall duringexcavation and during freezing of theretained soilPhillip S.K.Ooi a,*,Michael P.Walker b ,Jeffrey D.Smith c a Department of Civil Engineering,Holmes Hall 383,University of Hawaii at Manoa,2540Dole Street,Honolulu,HI96822,USAb GEIConsultants,I nc.,1021Main Street,Winchester,MA 01890,USAc Slattery,Interbeton,J.F.White,Perini,A Joint Venture,210South Street-8th Floor,Boston,MA 02111,USA Received 4January 2001;received in revised form 5June 2001;accepted 12November 2001AbstractA single-propped soldier-pile-tremie-concrete wall was constructed as part of the excavation support system for a tunnel jacking project.This wall supported a 14.3-m-high cut with a sig-nificant unbraced height (13.4m)above the excavation subgrade.A limited area of the exca-vation subgrade next to the wall was jet grouted to provide stability against base heave.The jet grout also served as a ‘‘brace’’for the wall below subgrade.The wall and prop were designed to support lateral loads during excavation as well as during ground freezing behind the wall.A ground freezing operation was implemented and carefully controlled to prevent wall yielding.Its success was facilitated by monitoring of wall movements,by drilling relief holes adjacent to the wall to reduce lateral heave pressures due to the soil freezing operation,and by limiting the exposure time of the wall to frost expansion.Measured wall behavior during excavation and during ground freezing is compared to those from finite element analysis for this very unique case history.Published by Elsevier Science Ltd.Keywords:Displacement;Excavation support;Finite element analysis;Freezing;Inclinometer;Jet grout;Moments1.IntroductionThe analyses and performance of a single-propped soldier-pile-tremie-concrete wall during excavation and during ground freezing behind the wall is presented.The *Corresponding author.Fax:+1-808-956-5014.E-mail address:ooi@ (P.S.K.Ooi).388P.S.K.Ooi et al./Computers and Geotechnics29(2002)387–409wall was constructed as part of the excavation support system for the I-90East-bound(I-90EB)tunnel jacking pit,built for Contract C09A4of the Central Artery/ Tunnel Project in Boston,MA.This tunnel will form part of the Interstate I-90 extension to East Boston.In Contract C09A4,construction of the I-90EB tunnel involves jacking of23.8m wide,11.3m high,tunnel box sections of various lengths underneath railway tracks, which lead to and from South Station,downtown Boston’s train hub.After casting the tunnels in the I-90EB jacking pit,they are jacked through a‘‘window’’in the headwall on the east end(Fig.1).Prior to demolishing the window,the soil behind the headwall is stabilized via ground freezing.In the original contract design,the headwall is supported by four levels of internal bracing with the bottom three levels removed prior to tunnel jacking.The contractor value-engineered the headwall and eliminated the bottom three braces.Thus,lateral support of the headwall is provided only by the upper brace.With this approach,the subgrade soil stabilized by jet grouting in the immediate wall vicinity must act as a strut(Fig.2).Elimination of the bottom three braces resulted in an unbraced height of 13.4m above the jet grout subgrade.The large clearance was necessary for construction of the tunnel boxes within the jacking pit.In addition to supporting a14.3-m-high cut, the wall and bracing system were also designed to withstand soil stabilization loads due to ground freezing behind the headwall.2.Subsurface profile and soil parametersThe soil profile in the vicinity of the headwall consists of an approximately4.6-m-thickfill layer overlaying the following soil strata in order of increasing depth:about 3.7m of organics,about30m of marine clay,about4m of glacial till and the argillite bedrock(Fig.2).The design ground water table is at El.32.3m based on project datum.El.0m of the National Geodetic Vertical Datum equals El.30.5m project datum.The wall was analyzed using thefinite element computer program SOILSTRUCT [1],which models the soil response utilizing the Duncan and Chang[2]nonlinear elastic hyperbolic model.The analysis was performed assuming drained behavior in thefill and glacial till,undrained behavior in the organics,marine clay and jet grout, and linear elastic behavior in the bedrock.Hyperbolic parameters,obtained from the results of triaxial testing,are summarized in Table1.Parameters that dominate wall behavior are discussed below.Being the thickest deposit,properties of the marine clay have a significant influ-ence on wall boratory strength testing performed on the marine clay include direct simple shear tests,and unconsolidated undrained and consolidated undrained triaxial compression tests.The majority of the undrained shear strengths were measured in direct simple shear,a test that yields an average strength between triaxial extension and triaxial compression of the marine clay,and therefore,pro-vides a simple and effective means of considering stress system-induced anisotropy [3].Variation of the shear strength of the marine clay with depth[4]and the shearFig.1.Plan of C09A4jacking pits and controlled ground freezing plan.P.S.K. Ooi et al. / Computers and Geotechnics 29 (2002) 387–409 389strength profile adopted for design are shown in Fig.3.With the exception of the top crust,the marine clay was lightly overconsolidated with OCRs varying mostly between1and2.There was considerable scatter in the plasticity index,which aver-aged about22%.The initial tangent undrained modulus(E i)of the marine clay isTable1Unfrozen soil parametersSoil type Soil unitweight(kNmÀ3)Frictionangle(degrees)K0Undrainedshear strengthS u(kPa)Young’s modulus a Bulkmodulus bInitialtangentmoduluscoef.c KInitialtangentmodulusexponent nStressratio atfailure R fBulkmoduluscoef.K bBulkmodulusexponent mFill20.4320.503000.40.752800.5 Organics17.300.6385500.0110.991700Marine clay(El.26.5m–El.6.1m)18.900.75677940.0110.913,2000Marine clay(El.6.1m–El.-3.0m)18.900.75728510.0110.914,2000Glacial till21.238 1.0010000.50.758330.5 Bedrock d23.6Linear elastic:E=1.92Â106kPa; =0.25eClay jet grout19.600.5184036,0000.0110.92600,0000a Initial tangent modulus,E i=Kp a( 3/p a)n where p a=atmospheric pressure and 3=minor principal stress[2].Tangent modulus,E t=(1-R f SL)2E i where SL=stress level=( 1À 3)/( 1À 3)f, 1=major principal stress and the subscript f denotes the stress at failure[2].b Tangent bulkmodulus,B t=K b p a( 3/p a)m.Bulkmodulus parameters for organics,marine clay and jet grout were selected to give a constant value of Poisson’s ratio=0.49.c For all soils,the unload-reload modulus,K ur,was taken to be equal to K except for glacial till where K ur=2000[5].d Bedrock=severely to moderately weathered argillite.e The bedrockis modeled as a linear elastic material with a Young’s modulus,E,and a Poisson’s ratio, .P.S.K. Ooi et al. / Computers and Geotechnics 29 (2002) 387–409 391based on E i=1200S u,which corresponds to a secant Young’s modulus of660S u at 50%stress level,where S u is the undrained shear strength.For a soil volume that is100%treated by jet grout,the contract documents spe-cified a minimum unconfined compressive strength(q u)of1030kPa.During the design stage,an unconfined compressive strength of824kPa or80%of the specified minimum value was used.However,based on laboratory testing of the jet grout in the jacking pit,the unconfined compressive strength ranged from620to10,100kPa with an average of3690kPa and a standard deviation of1340kPa.Measured jet grout parameters were used in the analysis discussed herein.The initial tangent modulus of the jet grout E i is approximately1000q u,which corresponds to a secant modulus of about1000S u at50%stress level.3.Wall and bracing detailsThe headwall is 25.9m wide,0.91m thickand 22.6m deep,a cross-section of which is shown in Fig.2.The minimum 28-day design compressive strength of the concrete is 27.6MPa.It is reinforced with W27Â217steel sections,with a yield strength of at least 345MPa,spaced 1.37m apart (EI=0.54Â106kNm 2/m and EA=2.80Â107kN/m).The top half of the soldier piles are reinforced with 38mm thick,432mm wide steel cover plates welded on the outside of both flanges (EI=1.24Â106kNm 2/m and EA=3.22Â107kN/m).The wall is considered impermeable and was designed for both soil and ground water pressures.The bracing system for the headwall consists of a horizontal steel truss (Fig.4).The truss,with centerline at El.33.8m,is framed into an end post attached to the south sidewall.The northwest corner of the truss is connected to two cross lot gir-ders and a pin teral support for the headwall is provided by a series of jacks spaced approximately 2.74m on center along the length of the brace.The jacks allow the Contractor to control wall movements into the excavation during ground freezing behind the wall.The stiffness of the horizontal steel truss was determined using finite element models developed to design the brace itself.A known load was applied to the brace and the deflection associated with that load was used to determine the brace stiffness (=51,800kN/m/m ofwall).Fig.4.Jacking pits with horizontal truss bracing at the I-90EB headwall.P.S.K.Ooi et al./Computers and Geotechnics 29(2002)387–4093934.AnalysisThe wall was analyzed assuming plane strain conditions.An incremental analysis procedure was used to model isotropic nonlinear stress-strain response of the soil.The soil was discretized using two-dimensional,sub parametric,quadrilateral (QM5)elements.The QM5quadrilateral element degrades to a triangular element by making the coordinates of two adjacent nodes of the quadrilateral the same [1].The wall is modeled using Euler-Bernoulli beam elements,the brace is represented by a linear spring and interface elements [6]are used to model slippage between the wall and the soil.Both the beam and interface elements have no thickness.The finite element mesh is shown in Fig.5.Bilinear elasto-plastic stress–strain characteristics were used for the interface ele-ments with the ultimate adhesion equal to one half the shear strength of the adjacent soil.The shear strengths for cohesive soils are summarized in Table 1.For cohe-sionless soils,the shear strength is estimated as K 0 v 0tan ,where K 0is the at-rest lateral earth pressure coefficient, v 0is the vertical effective stress at the mid-depth of the granular soil layer and is the friction angle of the soil.The shear stiffness of the interface elements was based on full mobilization of the ultimate adhesion at a relative displacement of 2.5mm [1].Trainloads were included in the analysis in accordance with the recommendations of the American Railway Engineering Association (AREA)manual [7].The wall was designed for 100,100,50and 25%of the train loads on the first,second,thirdFig.5.Finite element mesh.394P.S.K.Ooi et al./Computers and Geotechnics 29(2002)387–409P.S.K.Ooi et al./Computers and Geotechnics29(2002)387–409395 and fourth closest tracks,respectively.Trainloads from thefifth closest track and beyond were deemed to be far away enough that they were considered negligible.5.Modeling of the construction sequencingModeling of the construction sequencing is discussed in two parts:(1)excavation and brace installation,and(2)ground freezing.5.1.Excavation and brace installationThe excavation and brace installation sequencing was modeled as follows:Step1:apply surcharge loads and excavate to El.32.6m(Fig.6a).Step2:install brace at El.33.8m,prestress brace,and excavate to El.20.4m (Fig.6b).Modeling of the excavation from El.32.6m to El.20.4m in Step2was performed in several sub-steps to minimize errors due to non-linear unloading of the soil.To minimize wall movement during the initial cantilever stage,the depth to the brace was limited to0.9m.In the analysis,it was assumed that lateral resistance from side shear between the sidewalls and the jet grout is developed.The side shear is distributed uniformly along the width of the headwall.5.2.Ground freezingAs water freezes,it expands by about8.5%in volume.This implies that uncon-trolled frost expansion of the pore water can result in significant lateral heave pres-sures on the headwall.Therefore,a carefully planned ground freezing scheme was performed in two phases to minimize wall movements(Fig.1).First,the ground was frozen as a single unit,except for a3-m-wide‘‘buffer’’zone immediately behind the wall.In this zone,holes were drilled to relieve lateral pressure during the main soil freeze.Second,the soil in thefinal3-m zone was frozen.In both zones,the frozen soil extended from El.33.2m(1.5m below ground surface)to El.21.9m.In the analysis,the ground freezing stages are applied as a continuation of the excavation and brace installation sequencing as follows:Step3:Construct base slab in front of headwall leaving a0.3-m-wide gap.Install hydraulic jacks in this gap(Fig.6c).In subsequent stages,the loads in thehydraulic jacks are allowed to build up to a maximum of1460kN/m,afterwhich,they are held e of constant-force jacks duringconstruction allows the wall to move laterally during frost expansion,which in turn helps reduce large wall moments that would be generatedhad the wall been restrained by the slab.Step4:Main soil freeze—freeze ground except for a3-m-wide gap immediately behind the headwall.During this stage,a‘‘yield pressure’’increment isapplied to the wall as a lateral traction.It accounts for the lateralpressure increase necessary to cause the unfrozen soil in the final3-m-freeze zone to yield into the relief holes as the soil in the main freezezone expands (Fig.6d).Step 5:Final 3-m freeze—freeze the remaining unfrozen ground behind thee jacks to maintain brace loads constant at values existingafter main soil freeze.Volumetric expansion of this zone is modeled byperforming a thermal expansion analysis (Fig.6e).5.2.1.Main soil freeze (step 4)The yield pressure applied during main soil freeze varies depending on the soiltype and on the activities in the relief holes.Cohesionless soils (fill)easily ravel anddisplace into the relief holes under their own weight or with minimal assistancefromFig.6(continued ).P.S.K.Ooi et al./Computers and Geotechnics 29(2002)387–409397pumping or jetting.Therefore,the effects of lateral pressure relief in thefill are expected to result in an insignificant increase in lateral pressure on the headwall. Correspondingly,no yield pressure is applied to this portion of the wall.On the other hand,cohesive soils will not ravel and displace so readily into the relief holes. To relieve lateral heave pressures in the organics and marine clay,it is necessary to induce plastic yield of the soil between the relief holes,which in turn increases lateral pressures on the headwall.Two rows of relief holes were drilled in thefinal3-m-zone for pressure relief with the following dimensions and layout.In the organics/clay:150-mm-diameter holes at460mm on center,cased to the topof the organics.In thefill:150-mm-diameter holes at460mm on center,cased throughthe top1.5m.A limit equilibrium analysis was performed to estimate the increase in lateral pressures on the headwall in the clay and organic silt.In Fig.7,d,S and P w represent the diameter of the relief hole,the hole spacing and thefluid pressure in a drilled hole, respectively.From geometry,ab¼2S2sin ¼S sin ð1Þtan ¼dð2ÞAssuming that the failure wedge is bounded by points a,b and c in Fig.7a,the free-body diagram for this soil wedge is shown in Fig.7b,where N is the normal force acting on the plane bc,and S u is the undrained shear strength.The force polygon for the free body diagram is shown in Fig.7c.Force equilibrium along two axes parallel and perpendicular to the wall yields Eqs.(3)and(4),respectively: N sin2 ¼P w S sin þS u S cosðÞcos ð3ÞP y S2¼N cos2 þP w S sin þS u S cosðÞsin ð4Þwhere P y is the lateral yield pressure acting on the bining(3)and(4)to eliminate N and simplifying results in the following expression for P y:P y¼P wþS utanð5ÞCombining(2)and(5)yields:P y¼P wþS u Sdð6ÞA number of options were investigated to facilitate yield of the soils,including drilling holes at closer spacing,enlarging holes,lowering thefluid level inside the 398P.S.K.Ooi et al./Computers and Geotechnics29(2002)387–409holes and jetting with water.Two options reflecting the as-built relief hole layoutwere considered when estimating the yield pressure.Option 1:S /d =3[0.15-m-diameter relief holes at 0.45m on center]with water inthe holes up to the ground surface.Option 2:S /d =3with the water removed in the upper half of the holes,i.e.,waterat El.31.7m or approximately 1.5m above the top oforganics.Fig.7.Yield pressure derivation (a)soil wedge failing into relief hole;(b)free body diagram of soil wedge;(c)force polygon.P.S.K.Ooi et al./Computers and Geotechnics 29(2002)387–409399More aggressive options of pressure relief will result in less conservative values of yield pressure.Options1and2are summarized in Table3and Fig.8.From Fig.8,the selected design yield pressures were550and940kPa/m in the organics and marine clay, respectively.These values represent a reasonable average traction for the two pressure relief options and are assumed to act uniformly on the headwall.A lateral traction applied to the wall represents the difference between the yield pressure and the wall pressure when the bottom of excavation is atfinal subgrade.Because the wall pressure undergoes relaxation as the wall displaces,the applied traction has to be iterated until thefinal pressure on the wall equals the yield pressure.In this stage,hyperbolic parameters for the frozen soil in the main freeze zone were used in thefill,organics and marine clay(Table2)while hyperbolic parameters for the unfrozen soil(Table1)were used for thefinal3-m freeze zone.Parameters for the fro-zen soil are based on a literature review([8]through[16]).5.2.2.Final3-m freeze(step5)Thermal expansion analysis was not performed during the main freeze stage as it was modeled by the application of a yield pressure behind the wall.However,the final3-m freeze stage is simulated using thermal expansion analysis.Thermal expansion analysis is not required in thefill.This is because volumetric expansion of Table2Hyperbolic parameters for frozen soils used during main freeze stageSoil type Shearstrength a S u(kPa)Tensilestrength a(kPa)Initial Young’smodulus a E i(kPa)Stress ratioat failure a R fPoisson’sratio aFill b530024008.0Â1060.90.49 Organics24002400 5.0Â1050.90.49 Marine clay38002400 1.0Â1060.90.49a Frozen soil parameters were determined from references[8]through[16].b For thefill,the tabulated soil parameters were also used during the third sub-step of freezingfinal3-m zone.During sub-steps1and2of thefinal3-m freeze,hyperbolic parameters of the unfrozenfill were used. Table3Yield pressure estimates for main freeze stageElevation (m)Soil type Undrained shearstrength S u(kPa)Hydrostaticpressure P w(kPa)Yield pressure aP y b(kPa mÀ1)Option1c Option2d Option1c Option2d30Organics384515520430 27Organics388151640540 27Marine clay678151920830 22Marine clay67125961070970a S/d=3(0.15-m-diameter holes at0.45m on center).b P y=P w+S u S/d.c Option1=water inside relief hole at ground surface(El.34.7m).d Option2=water inside relief hole at El.31.7m.400P.S.K.Ooi et al./Computers and Geotechnics29(2002)387–409free-draining cohesionless soils is relatively insignificant due to their high perme-ability[17].When chilled,water is forced out of the void spaces at the same rate as freezing progresses resulting in a lower frozen water content.Thermal expansion analysis was modeled in three sub-steps in the cohesive soils. During thefirst two sub-steps,which represent91%of the total volumetric strain (Fig.9),adjacent columns of the frozen soil are not in contact,and the stiffness of the partially frozen soil is dominated by the stiffness of the unfrozen soil.In fact,modulus values of the‘‘yielded’’unfrozen soil were used because the organics andmarine clay in the final 3-m freeze zone have already reached plastic yield during mainsoil freeze.Therefore,the soil stiffness for sub-steps 1and 2were set equal to the secantmodulus of the unfrozen soil at about 95and 75%stress levels,respectively.During sub-step 3(representing 9%of the total volumetric expansion),the frozensoil columns come in contact,and the stiffness of the soil was set equal to the fullyfrozen value.Thermal expansion of the clay and organic silt was simulated by applying nodalloads,which are proportional to the product of the thermal expansion coefficient ( )and the temperature change (ÁT ).Appropriate coefficients of thermal expansionwere derived by setting the volumetric expansion of the pore water equal to 8.5%.Hence,the volume change,d V ,is:d V ¼0:085nV ð7Þwhere V is the total soil volume and n is the porosity.The volumetric strain,"v is:"v ¼d V =V ¼0:085n ð8ÞFig.9.Frozen soil volume when ice columns are in contact.402P.S.K.Ooi et al./Computers and Geotechnics 29(2002)387–409For a saturated soil,the porosity is related to the specific gravity,G s,and water content,w,as:n¼e1þe¼G s w1þG s wð9Þwhere e is the void bining(8)and(9),the volumetric strain becomes:"v¼0:085G s w1þG sð10ÞAfter freezing a cube of soil of unit volume,the volume of the expanded soil is as follows:1þ"v¼1þd xðÞ1þd yðÞ1þd zðÞð11Þwhere d x,d y and d z are the increase in length of the soil cube in the x,y and z directions,respectively.If the soil freezes uniformly,the change in length can be expressed as follows:d x¼d y¼d z¼1þ"vðÞ1=3À1ð12ÞThe coefficient of thermal expansion, ,is the change in length per unit temperature change.Therefore,for a unit temperature change, =d x=d y=d z.To account for expansion along the plane of the wall in a plane strain analysis, is reduced as follows:¼1þ"vðÞ1=3À11þð13Þwhere is Poisson’s ratio of the frozen soil and"v is calculated using Eq.(10). Values of are proportioned at each sub-step so that9%of the total volumetric expansion occurs in sub-step3with the remaining91%divided into two equal increments for sub-steps1and2.A unit temperature change was specified for each sub-step.The coefficients of thermal expansion for the organics and marine clay are summarized in Table4.Thermal expansion analysis can only be performed using linear elastic properties of the soil also summarized in Table4.For thefill,unfrozen drained hyperbolic parameters(Table1)were used for thefirst two freeze sub-steps and fully frozen hyperbolic parameters(Table2)were used for the third freeze sub-step.parison of predicted and measured behaviorThree inclinometers that extend below the bottom of the headwall were installed to monitor lateral displacements.Wall movements for a centrally located inclin-ometer are compared to the predicted movements in Fig.10at the end of excavation and after soil freezing.P.S.K.Ooi et al./Computers and Geotechnics29(2002)387–409403Table4Linear elastic properties of cohesive soils duringfinal3-m freeze stageSoil type Average watercontent w(%)Specificgravity,G sVolumetricstrain,"v(%)Coef.of linear thermalexpansion, aYoung’s modulus,E(MPa)Poisson’sratio, Sub-step1Sub-step2Sub-step3bTotal c Sub-step1(Â10À3)Sub-step2(Â10À3)Sub-step3(Â10À3)Sub-step1Sub-step2Sub-step3Sub-steps1and2Sub-step3Organics80 2.55 2.6 2.60.5 5.7 5.77 5.77 1.116144800.49940.49 Marine clay35 2.70 1.9 1.90.4 4.2 4.22 4.220.8912299600.49940.49 a is calculated using Eq.(13).b9%of the total volumetric expansion was applied in sub-step3when the frozen soil columns are in contact.The remaining91%volumetric expansion was divided into two equal increments in sub-steps1and2.c The total volumetric strain is calculated using Eq.(10).404 P.S.K.Ooietal. / Computers and Geotechnics 29 (2002) 387–409406P.S.K.Ooi et al./Computers and Geotechnics29(2002)387–409At the end of excavation,the predicted wall movements are in excellent agreement with the measured values as shown in Fig.10a.In Fig.10b,the predicted move-ments after main soil freeze and afterfinal3-m freeze are plotted with the measured movements prior to wall demolition.The measured wall movements fall close to the predicted values after main soil freeze but are significantly less than those predicted afterfinal3-m freeze.This can be explained with the aid of the following list of key construction milestones and dates.Construction milestone DateStart of main soil freeze August1999 Start offinal3-m freeze October1999 Start of wall demolition December1999 The movements shown in Fig.10b were measured in November1999,one month after the start offinal3-m freeze.The measured movements agree more with the predicted values after main soil freeze rather than afterfinal3-m freeze due to one or a combination of the following:(1)the rate of frost expansion is very slow in the marine clay and organics,due to the low permeability[17].The maximum movement was recorded just one month after the start offinal3-m freeze.Such a short period may not be sufficient for full frost expansion to occur;(2)the coefficient of thermal expansion in thefinal3-m freeze zone may be overestimated because of compression of the soil fabric during the two-stage freezing process.During main soil freeze,ice buds form in the free water within the clay macro pores.Due to differences in local temperature and due to matric potential,both adsorbed water and free water from adjacent macro pores in thefinal3-m freeze zone migrate towards the ice buds in the main soil freeze zone.Subsequent ice growth during thefinal3-m-freeze stage then causes compression against the water-depleted soil fabric,bond breakage,and a breakdown of the soil structure.As a result,the soil particles are brought closer together,thereby reducing the effective thermal expansion coefficient in thefinal3-m freeze zone;and(3)the plane strain analysis ignores relief holes that are also provided between the main freeze zone and the south wall of the adjacent jacking pit.This line of relief holes orthogonal to the headwall‘‘attract’’ice growth,resulting in an overestimation of lateral heave pressures on the headwall.The criterion for success was for the wall movements to not exceed the150-mm-travel of the series of jacks located at the brace level and at the base slab.The maximum measured wall movements at the end of excavation and end of freezing were7and31 mm,respectively,both of which were well below the limit of the jacktravel.Wall bending moments are related to the wall curvature as follows: M¼EIð14Þwhere M=bending moment in the wall,=wall curvature,E=modulus of elasticity of the wall and I=moment of inertia of the wall.The wall curvature was obtained by numerical differentiation of the slope measured by the inclinometer.Prior to cracking,I is taken to be the gross moment of inertia of the concrete neglecting the。