numerical analysis of cylindrical diaphragm wall excavation at coastal area

Instrumentation and Numerical Analysis ofCylindrical Diaphragm Wall Movement During

Deep Excavation at Coastal Area

DONG SOO KIM

Department of Civil & Environmental Engineering,Korea Advanced Institute of Science and Technology,Taejon, Korea

BYOUNG CHUL LEE

Department of Civil Engineering,Shingu University, Sungnam, Korea

The lateral deflection of a cylindrical diaphragm wall and the associated groundmovement induced by deep excavation are analyzed by performing site instrumenta-tions and numerical analyses in the coastal area of Korea. Wall lateral deflection,rebar stress, and pore water pressure were measured and analyzed in eight direc-tions. Variations of soil properties with the decrease of confining pressure are com-pared by performing various in situ tests before ad after excavation. To calculate thewall lateral deflection accurately, the effects of small strain nonlinearity, confiningpressure, and the hysteresis loading=unloading loop developed during excavation areconsidered in the proposed numerical analysis. By comparing numerical results withmeasured ones, the importances of considering small strain nonlinearity and confin-ing pressure reduction in the nonlinear (FEM) are emphasized. Also, the effects ofwall stiffness on the performance of cylindrical diaphragm walls are studied forfuture similar excavation in the coastal area.

Keywords cylindrical diaphragm wall, lateral deflection, instrumentation, FEM,confining pressure, small-strain nonlinearity

Although the reliable evaluation of ground movement has become essential for theeconomy and safety of major offshore construction projects, the current techniquesfor evaluation of marine soil properties and design procedure show a considerablelack of accuracy when compared with the instrumented results. This is usuallyattributed to a misunderstanding of marine soil behavior under working loadconditions.

The strain level experienced in the soil medium under working load conditionusually ranges below about 0.5%, and soil behaves nonlinearly from the small strainranges of about 103% (Jardine et al. 1986; Burland 1989). The stiffness of granularsoils and weathered rock is considerably affected by the in situ confinement, and the

Address correspondence to Dong Soo Kim, Department of Civil & EnvironmentalEngineering, Korea Advanced Institute of Science and Technology, Taejon, Korea.E-mail: [email protected]

Marine Georesources and Geotechnology, 23:117136, 2005Copyright # Taylor & Francis Inc.ISSN: 1064-119X print/1521-0618 onlineDOI: 10.1080/10641190590953728

117

modulus values based on (SPT) obtained by site investigations before major exca-vation would be quite different from those values after excavation. In current designpractice, subgrade reaction method and=or a linear finite element method are oftenemployed for the analysis of deformational behavior of geotechnical offshore struc-tures using soil properties determined by SPT N-values and=or a conventional tri-axial tests which cannot properly consider the effects of confinement, small-strainnonlinearity and hysteresis loading=unloading loop. Therefore, it is necessary todevelop the refined site investigation and numerical analysis methods which can con-sider the aforementioned effects for the reliable ground movement analysis duringdeep excavation at coastal area.

In this article, the case study of the deep excavation for a 56m-depth cylindricalinground LNG storage tank at coastal area was investigated. The 1.7m thick, 80mdiameter, and 75m deep cylindrical diaphragm wall was utilized as a retaining struc-ture for the excavation. Detailed site investigations were performed both before andafter excavations to determine the variation in deformational characteristics due tothe decrease of confinement. The small-strain nonlinearity of the site was evaluatedby effectively combining the maximum shear moduli determined by a downhole testwith the modulus reduction curves determined by a resonant column test. A numberof instruments were installed at the diaphragm wall and the adjacent ground, and thedistributions of wall deflection, pore water pressure, and rebar stress were monitoredat various locations during the excavation (Samsung Corporation 1998). The resultof series of numerical analyses was compared with the monitored wall deflection,and the importance of considering the effects of small-strain nonlinearity and confin-ing pressure was assessed. Finally, the effects of wall stiffness on the performance ofcylindrical diaphragm walls were evaluated for the future similar construction in thecoastal area.

Outline of Diaphragm Wall and Excavation Work

Inground LNG storage tanks were constructed at Inchon in the west coast of Korea.The capacity of each storage tank is 200,000 kl; its diameter is 80m; and its innerexcavation depth is 56m. The 1.7m thick and 75m deep cylindrical diaphragm wallwas constructed as a retaining structure to withstand the earth and water pressuresduring excavation without any prop system. The horizontal support is achieved byvirtue of hoop-compression stresses on the circular wall. The wall was installed intothe soft rock by 3 m and cast by 26 panels. The earth retaining wall was constructedby the slurry wall construction method. The cross section of the LNG storage tank isshown in Figure 1.

The excavation was carried out at eight stages to prevent the occurrence ofnonuniform lateral pressures. The ground water table inside the retaining wall waslowered by deep wells and maintained its level about 1m below the bottom of exca-vation at each stage. The excavation works proceeded smoothly up to the depth ofabout 50m when a thin impermeable clay layer was encountered. Once the clay layerwas excavated, a significant amount of seepage occurred into the excavated area andfurther excavation could not be continued. To stabilize the inward seepage, injectiongrouting was performed into the weathered and soft rocks through the holes installedalong the diaphragm wall. Detailed site investigation was performed again and thedeformational characteristics of the inside soil can be evaluated before and afterexcavation.

118 D. S. Kim and B. C. Lee

Site Profile and Soil Properties

Geological Profile

The site consisted of a layer of reclaimed soil of about 10m, a loose alluvial silty sandof about 37m, an alluvial clay of about 3m, a sandy gravel of 5m, a weathered rockof about 17m and soft rock (Figure 1). The degree of weathering of rock was severeat the upper part, and its intensity was mitigated at the lower part. The alluvial claylayer deposited at a depth of about 50m was of interest in this study. The water tablewas located at a depth of about 7m. Both in situ and laboratory tests including SPT,downhole, pressuremeter, and resonant column tests were performed to obtain thedeformational characteristics of the site (Table. 1).

Comparison of Soil Properties before and after Excavation

In most excavation work, site investigation is usually performed before the exca-vation. If the stiffness of soil is affected by confinement, the deformational charac-teristics of the soil determined after the excavation would be quite different fromthose determined before excavation. In this study, the stiffness of the weatheredrock, which played the major role in resisting the movement of the wall, can beevaluated before and after excavations.

In SPT, the safety type hammer of energy ratio of 60% was used. SPT N valuesdetermined before and after excavations were compared in Figure 2(a), and it isobserved that N values of the weathered rock were significantly reduced as a resultof seepage and decrease of confinement due to the excavation. Menard-type

Figure 1. Cross section of inground LNG storage tank.

Analysis of Wall Movement During Deep Excavation 119

pressuremeter tests were also performed before and after excavations and elasticmodulus of weathered rock obtained after excavation was found to be considerablylower than the modulus before excavation as shown in Figure 2(b).

Downhole test was performed before excavation to obtain the shear wave velo-city (Vs) profile of the site as shown in Figure 3. At depths of 24m, 34.5m, and 40mduring excavation, Spectral Analysis of Surface Wave (SASW) tests were performedto evaluate the variation of vs with confinement. Because of the limited testing space,the receiver spacing in the SASW tests was restricted. It was interesting to note thatthe shear wave velocities measured after excavation were lower than the velocitiesobtained before excavation, and the importance of considering the confinementeffect in the evaluation of deformational characteristics can be explained. Unfortu-nately, it was not possible to perform an SASW test at the bottom of the excavation(at a depth of 55m) because of construction problems. The shear wave velocity afterexcavation was estimated by correcting the downhole test result based on confiningpressure influence factors of weathered and soft rocks determined by resonantcolumn tests as shown in Figure 3. The elastic modulus values determined bySPT, PMT, and shear wave velocity measurements before and after excavation are

Table 1. Comparison of soil properties before and after excavation

Depth EL,mSoil

classification

N-Value E 28N (kPa) Vs (m=s)

before after before after before after

7.42.6 Reclaimed full 17 44667 160 2.616.3 Silty sand 24 65883 200 16.330.0 Silty sand 42 115295 260 30.039.9 Silty sand 70 192158 310 39.942.3 Clay 73 204420 310 42.347.3 Sandy gravel 127 348630 310 47.363.8 Weathered rock 173 85 474906 233335 450 24663.872.1 Soft rock 510 1400011 750 40072.1100.0 Hard rock 1500

Depth (EL, m)Soil

classification

Unit weightc (kN=m3)(before)

Max. shearmodulus

Gmax (Pa)

(before)

ElasticmodulusEmax (Pa)

(before)

7.42.6 Reclaimed fill 18.14 4.736E7 1.260E82.616.3 Silty sand 18.14 7.400E7 1.968E8

17.339.9 Silty sand 18.63 1.284E8 3.417E830.039.9 Silty sand 20.10 1.970E8 5.240E839.942.3 Clay 18.63 1.826E8 4.857E842.347.3 Sandy gravel 20.59 1.922E8 5.113E847.363.8 Weathered rock 23.04 4.759E8 1.266E963.872.1 Soft rock 23.04 1.434E9 3.815E972.1100 Hard rock 25.00 5.738E9 1.526E10

Gmax and Emax are obtained from shear wave velocity Vs.


compared in Figure 4. In the Korean Specification of Highway Bridges (KoreanHighway Corp. 1996), the elastic modulus can be estimated as 28N based on SPTN-values. The maximum elastic modulus at small strains was calculated using theshear wave velocity. Modulus values determined by wave velocities were larger thanthose determined by 28N, and this difference can be explained by the difference in

Figure 2. Comparison of modulus obtained by SPT and by PMT between before and afterexcavation.


strain amplitude as discussed later. The effects of confining pressure reduction onmodulus values before and after excavation are embodied in the proposed numericalanalyses by using the Janbu (1963) formula.

Figure 4. Comparison of elastic moduli between before and after excavation according tovarious testing methods.

Figure 3. Comparison of shear wave velocity profiles with excavation steps.


In the numerical analysis, the layer of weathered rock plays a major role in thewall lateral deflection. The modulus values of weathered rock were significantlyreduced due to the disturbance and the decrease of confinement during excavation.

Figure 5. Normalized modulus reduced curve by RC tests at average effective confining press-ure of each layer.

Figure 6. Comparison between elastic moduli determined byN-value, downhole test, andRC test.


If one used the modulus values determined before excavation, the stiffness of thelayer would be overestimated in the analysis.

Variation in Elastic Modulus with Strain

Because the soil exhibits the nonlinear behavior even at small strains and the strainlevel experienced in the soil media is small (usually less than about 0.5%) underworking stress conditions, the evaluation of small strain nonlinearity is importantfor the reliable analysis of wall deformation (Burland 1989). Samples were obtainedat eight depths for resonant column tests and the test results are shown in Figure 5.Test results were fitted using Ramberg-Osgood model and fitting parameters wereemployed in the proposed model discussed later. Then, the nonlinear stress-strainrelations were determined at each layer by effectively combining the maximummodulus at small strain determined by the downhole test with the modulus reductioncurves determined by resonant column tests (Kim et al. 1997). The variations in elas-tic modulus values with strain levels are presented in Figure 6. The modulus valuesdetermined by 7N and 28N based on Korean Specification of Highway Bridges(Korea Highway Corp. 1996) were also plotted for comparison purposes. Modulus

Figure 7. Monitoring points and devices.


values at each layer were decreased as strain level increased, and it was interesting tonote that modulus value estimated by 7N was considerably low, but values estimatedby 28N were close to the value at a strain level of about 0.1%. Considering the factthat the strain level in the working stress condition is usually less than 0.11.0%, itmay be appropriate to use 28N as a linear elastic modulus.

Instrumentation

An extensive instrumentation program was implemented, which included (1) inclin-ometers in eight directions to measure the wall lateral deflection, (2) pore waterpressure transducers in four directions to measure the variation in pore pressuresduring excavation at four different depths, (3) rebar stress meters in the radial andvertical directions, and so on (Figure 7). The measuring section was divided intoeight directions (31, 73, 128, 170, 211, 253, 308, 350). In the directions of 73

and 253 (MAIN LINE 1) all instruments were performed. In the directions of170 and 350 (MAIN LINE 2) which were almost at a right angle to MAIN LINE1, all instruments excluding pore water pressure meter in the surrounding groundwere installed. In other directions, minimum required instruments were installed.

Figure 8. Lateral wall deflections and average value in final excavation step.


The variations of wall deflections with depth were monitored at eight directionsby inclinometers installed in diaphragm wall. Circumferential rebar stress meterswere installed at five depths in eight directions. Pore water pressure transducers wereinstalled to measure the applied water pressure to the diaphragm wall and to confirmsecurity against permeation through the bottom. Pore pressure transducers wereinstalled at three depths in four directions outside of the diaphragm wall and atone depth in four directions inside of the diaphragm wall.

Observed Performance during Excavation

Lateral Deflection of Diaphragm Wall with Excavation Steps

Figure 8 shows the measured wall deflections at eight different inclinometer locationsafter final excavation to the depth of 56m. The inclinometers were installed insidethe wall and the monitored deflections were corrected based on the absolute deflec-tions monitored at the top of the wall accurately using a transit. The difference inwall deflection at various locations can be explained by the occurrence of eccentricforce derived from surrounding ground.

Figure 9. Comparison of lateral deflection by using rebar stress and inclinometer value at 4th,6th and final excavation steps.


The average value of measured deflections is plotted as a solid line in Figure 8.The excavation was carried out at eight steps during two months, and the diaphragmwall was supported only by the hoop stress of cylindrical diaphragm. The averagevalues of lateral wall deflections with excavation steps of 4th, 6th, and final are shownin Figure 9. Since the shape of the diaphragm wall in cylindrical, the wall lateraldeflection can also be computed indirectly by measuring the rebar stress installedin the wall. The rebar stresses are also plotted in Figure 9 for the comparison pur-poses. The calculated wall lateral deflection shape based on the rebar stresses withexcavation steps matched well the average deflection shape monitored by inclin-ometers, indicating the reliability of using the average deflection shape in the analysis.

Variation in Pore Water Pressure

Pore water pressures were monitored in four directions at depths of 18.4m (EL-11m)and 52.7m (EL-45.3m) and outside of wall and at a depth of 70.4m (EL-63m) bothinside outside of the wall. The variations in measured pore pressures with time areshown in Figure 10. On the 7th of October, pore pressures measured at EL-63m beganto decrease sharply and then stabilized. At thatmoment, the impermeable clay layer thatexisted at a depth of 47m was excavated and the abrupt seepae started to occur due tothe pressure head difference between the outside and inside of the wall. In the designstage, the clay layer was not considered as impermeable and the pore water pressurewas considered as hydrostatic, because it was expected that the wall embedded into softrock which would constrain the seepage. Judging from the implementation results, theprecise expectation of the permeability of the soft rock and the weathered rock wereimportant for estimating the real behavior of the surrounding ground and the wall.

Numerical Analysis

Proposed Nonlinear Elastic Model

In this study, the improved numerical modeling method was proposed to estimatethe behavior of the cylindrical diaphragm wall and surrounding ground movement

Figure 10. Results of pore water pressure measurement.


during deep excavation in coastal area. During excavation, soil strains in the work-ing stress ranges usually experienced were below about 0.5% (Burland, 1989). At thissmall strain range, stress-strain relationships of soil show nonlinearity and hysteresisloop. The backbone curve is very important for the evaluation of loading=unloadingbehavior because it defines the initial stiffness of the soil at small strains and consti-tutes the basis for characterizing the stress-strain behavior of soils for nonlinearanalysis. Ramberg-Osgood stress-strain relationship generally fits the nonlinearityof the experimental data at strain range below 0.1% very well (Anderson andRichart 1975). In this article, hysteresis loop can be represented by the backbonecurve, which is described by Ramberg-Osgood Parameters, and by assumingMasing rule.

One form of the Ramberg-Osgood stress-strain equation for the inintial back-bone curve can be written as (Eq. 1):

e1 r1 r3

Ei

C r1 r3

Ei

R; 1

where e1 is principal strain, r1 r3 is deviatoric stress, Ei Emax is initial modu-lus, C is dimensionless coefficient, and R is dimensioless exponenet.

Since the tangential modulus is defined as the instantaneous tangential slope ofthe stress-strain curve, the tangential modulus can be calculated from (Eq. 1) by dif-ferentiating with respect to the major principal strain as shown in (Eq. 2).

Et @r1 r3

@e1 Ei

1 CR r1r3Ei R1 : 2

According to the excavation steps, the effective confinement decreases due to theremoval of surcharge load and seepage flow, and soil stiffness inside the wall afterexcavation is reduced compared to that before excavation. Based on the site investi-gations, soil stiffness after excavation is estimated about 30% of the value beforeexcavation. Therefore, for the precise estimation of deformation behavior, it is neces-sary to consider a variation of soil stiffness due to confining pressure reduction byexcavation. The initial tangential modulus of stress-strain curve has been observedto be dependent on confining pressure. Janbu (1963) proposed the following equa-tion (Eq. 3) to consider the dependency of confining pressure for initial tangentialmodulus Ei.

Ei Emax KPar3Pa

n; 3

where, Pa is atmospheric pressure, initial tangential modulus factor K and stressinfluence exponent n are nondimensional material parameters.

If (Eq. 3) is substituted into (Eq. 2), the tangential modulus of backbone curveconsidering the effect of confining pressure at working load condition is representedby (Eq. 4).

Et KPa

r3Pa

n

1 CR r1r3KPar3=Pan

R1 : 4


Then, the tangential modulus for the unloading condition is represented byconsidering Masing behavior.

Eur KPa

r3Pa

n

1 CR2R1

r1r3KPar3=Pan

R1 ; 5

where C and R are Ramberg-Osgood fitting curve parameters, and r1 and r3 aremajor and minor principal stresses, respectively. These parameters are determinedfrom the in situ and laboratory tests and listed in Table 2. The advantages of usingthe proposed model is that the soil behavior during excavation can be calculatedautomatically and accurately by considering (1) small strain nonlinearity, (2) effectof confining pressure reduction, and (3) hysteresis loop for unloading and reloading.

The proposed model was implemented in the computer program of ABAQUS(1995). The validity and accuracy of the proposed model was verified by comparingthe results with hyperbolic model developed by Duncan et al. (1980). Because of thelimited space, the details cannot be included in this paper.

Finite Element Analysis

A series of 2D finite element analyses was performed on a deep excavation site usingthe program ABAQUS (1995), where the proposed model is embodied by usingUMAT Subroutine. The parameters of the proposed model (K, n, C, R) weredefined at each layer and listed in Table 2.

For simplicity, an idealized axisymmetric plane strain excavation geometry withan excavation depth H and a width of 40m was considered as shown in Figure 11The site was assumed to be composed of layers of fill, alluvial soil deposit and weath-ered rock overlying a soft rock stratum. A refined mesh, which consists of approxi-mately 5828 nodes and 5673 elements as shown in Figure 12, was adopted tominimize the effect of mesh dependency on the finite element modeling. Becauseof the symmetry about the excavation centerline, only one-half of the excavationwas considered in the finite element model. The finite element mesh extends to adepth of 1.0H below the final excavation platform and laterally to a distance of3.0H from the excavation centerline. The soil and wall was discretized using 8-nodedisoparametric axisymmetric plane strain element. In addition, the interface betweenthe wall and retained soil was modeled using contact pair elements. The nonlinearbehaviors of the fill, alluvial soil deposit, and weathered rock were modeled by usingthe proposed model. Soft rock and wall were treated as a linear elastic material. Inthis study, excavation was simulated numerically by removing the elements rowwisefrom the front of the wall. At each time of such excavation, the excavated surface

Table 2. Values of parameter used in the proposed numerical analysis

Parameters K n C R

Fill 230 0.68 31622 2.57Alluvial soil 420 0.60 79433 2.97Weathered rock 1000 0.62 3981 2.21


was made stress free by calculating the equivalent nodal forces from the removed ele-ments and applying them on the excavated boundary (Ishihara 1970). The impliedload due to excavation was found too large to be applied in a single increment forthe nonlinear soil. Hence the load was spread over a number of increments until

Figure 12. Finite element mesh.

Figure 11. Excavation geometry.


numerical stability was established and convergence was achieved. An excavationstep is repeated until the excavation reaches the final depth EL-49m.

Analysis Results

Comparison Between the Measured and Analyzed Results

Measured lateral wall deflections are compared to nonlinear FEM analysis resultswith excavation steps in Figure 13. Soil parameters listed in Table 2 reflect the effectof confining pressure reduction derived from ground excavation and small-strainnonlinearity of soil modulus. Because the stiffness of the cylindrical diaphragm wallwith thickness of 1.7m is high, lateral wall deflection is relatively small. FEM analy-sis results are well consistent with measured values at each excavation steps. On theother hand, wall deflection obtained from using parameters before excavation showsmuch smaller deflections at depths between 40m and 55m, compared to the mea-sured data at final excavation step as shown in Figure 14. It is explained that para-meters before excavation cannot consider the effect of confining pressure reductionduring excavation.

Figure 13. Comparison between measured and numerical analysis results with excavationsteps.


Comparison Between Linear and Nonlinear Finite Element Analysis

Though nonlinear analysis predicts the actual soil behavior well, it can be somewhatdifficult to use routinely. As an alternative to nonlinear analysis, linear analysis canbe chosen with a constant value of modulus that corresponds to the strain range ofsoil under working load. The deviatoric strain contour is shown in Figure 15 and themaximum strain level is about 0.1%, which is consistent with the fact demonstratedby Burland (1989). Wall deflections obtained by linear analysis using soil modulusvalues at strains of 0.01% and 0.1% are compared with those by nonlinear analysisin Figure 16. It has been found that the magnitude of soil modulus at strains of0.01% and 0.1% are about 80% and 40% of maximum soil modulus, respectively.Deflection curve obtained by nonlinear analysis is close to the curve obtained bylinear analysis with modulus value at strain of 0.01%, and both results matchreasonably well with measured curve. But the result obtained by linear analysis withmodulus value at strain of 0.1% overestimates the deflection because the modulusvalues were underestimated. As shown Figure 15, the average strain level aroundthe diaphragm wall is about 3.5 102% with the maximum strain less than0.12%, and the difference in deflection curves at each analysis can be explained bythe strain distribution. Even though the linear analysis with modulus values at strainof 0.01% provides the better result compared to the nonlinear analysis, it is difficult

Figure 14. Comparison of results with soil properties before and after excavation.


Figure 16. Comparison of results between linear and nonlinear analysis with strain levels.

Figure 15. Contour of maximum deviatoric strain.


to choose the proper strain level in the linear analysis, because it depends on variousfactors such as soil and wall stiffnesses.

In the proposed case, the wall stiffness of the cylindrical diaphragm was so largethat lateral wall deflection appeared very small, but the effect of the nonlinearity ofsoils would influence significantly lateral wall deflection in case of small wall stiff-ness. Using the numerical analyses performed with altering wall stiffness, the differ-ence between linear and nonlinear analysis can be noticed clearly. For similar deepexcavation in the coastal area, the effects of wall stiffness on the lateral wall deflec-tion have been investigated by varying the wall stiffness of EI 13510, 4752, 2750,1408, and 594MNm2, which corresponds to 1.7, 1.2, 1.0, 0.8, and 0.6m thick cylin-drical diaphragm walls, respectively. As shown in Figure 17, the lateral wall deflec-tion decreases with increasing wall stiffness. The shape of lateral wall deflection withwall stiffnesses shows similar pattern. The maximum lateral wall deflection at the0.6m, 0.8m, 1.0m and 1.2m cylindrical diaphragm walls is increased by 93%,64%, 43% and 27%, respectively, as compared to the maximum lateral wall deflec-tion of the 1.7m thick wall.

As shown in Figure 5, nonlinearity of soil varies soil modulus with strains. Aswall stiffness decreases, wall deflection increases as shown in Figure 17. The increaseof wall deflection comes from both the decrease of wall stiffness and the decrease of

Figure 17. Curves of lateral wall deflections with various wall stiffnesses.


soil modulus due to increase of soil strain. By comparing wall deflections obtainedby linear analysis with constant modulus values and nonlinear analysis for all casesof wall stiffness, respectively, the effect of soil nonlinearity can be explained. In thisanalysis, the soil modulus used in linear analysis is 80% of maximum value, whichcorresponds the values at strains of about 0.01%, and the effect of soil nonlinearitycan be defined as the difference of maximum wall lateral difference between linearand nonlinear analysis. It is found that the effect of soil nonlinearity increases withdecreasing wall stiffness. This means that wall lateral deflection is governed by notonly wall stiffness but also soil nonlinearity with decrease of wall stiffness. Therefore,to evaluate exactly wall deflection and surrounding ground deformation, numericalanalyses must be performed considering the small-strain nonlinearity of soil as wellas the effect of confining pressure reduction during excavation.

Conclusions

The lateral deflections of a cylindrical diaphragm wall induced by deep excavationare analyzed by performing site instrumentations and numerical analyses in the cos-tal area of Korea. Comparing the elastic modulus values before and after excavation,it was found that the elastic modulus is reduced significantly after the excavation dueto the effect of confining pressure reduction and the permeation. Therefore, theeffect of confining pressure reduction should be considered appropriately to deter-mine the elastic modulus for an accurate evaluation of the lateral deflection ofdiaphragm wall.

The stress-strain model, which enables consideration of the effects of soil non-linearity and the confinement reduction, was proposed. The propsed model is imple-mented in ABAQUS, and its applicability to analysis of deep excavation works hasbeen found to be satisfactory through the verification with in situ measurementresults. To evaluate wall lateral deflection properly, numerical analyses should beperformed considering the effect of small-strain nonlinearity of soil and confiningpressure reduction during excavation.

References

ABAQUS. Users and theory manuals, V. 5.5. Hibbit, Karlson & Sorensen Inc.Anderson, D. G. and F. E. Jr. Richart. 1975. Effects of shearing of shear modulus of clays.

Journal of the Geotechnical Eng. Div., ASCE 102(9): 975987.Burland, J. B. 1989. Ninth Lauritis Bjerrum Memorial Lecture: Small is beautiful-The

stiffness of soils at small strains. Canadian Geotechnical Journal 26: 5265.Duncan, J. M., P. Bryne, K. S. Wong, and P. Marbry. 1980. Strength, stress-strain and bulk

modulus parameters for finite element analyses of stresses and movements in soil masses.Geotechnical Engineering Research Report No UCB/GT/80-01, University ofCalifornia, Berkeley, California.

Ishihara, K. 1970. Relations between process of cutting and uniqueness of solution. Soil andFoundations 10(3): 5065.

Janbu, N. 1963. Soil compressibility as determined by oedometer and triaxial tests. ProcEurope Conference on soil Mechanics and Foundation Engineering 1: 1925.

Jardine, R. J., D. M. Potts, A. B. Fourie, and J. B. Burland. 1986. Studies of the influence ofnon-linear stress-strain characteristics in soil-structure interaction. Geotechnique 36(3):377396.


Kim, D. S., K. C. Kweon, S. Y. Jeong, and J. Y. Park. 1997. Evaluation of nonlinear defor-mation characteristic of soil using laboratory tests and site tests. Journal of the KoreanGeotechnical Society 13(5): 89100 (in Korean).

Korea Highway Corporation. 1996. Korea specification of highway and bridges. (in Korean).Samsung Corporation. 1998. Subsoil investigation of incheon LNG storage tank, Rpth to

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numerical analysis of cylindrical diaphragm wall excavation at coastal area

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