1 INTRODUCTION Excavation protected by diaphragm walls may be considered as the default execution method for sta-tions and other rail- and underground structures in urban areas. Diaphragm walls of reinforced concrete are built section by section. The principal construc-tion phases of such a section, called panel, are as fol-lows: (a) construction of guide wall (b) panel trench excavation under bentonite slurry or other support-ing fluid (c) bottom-up concrete filling of the trench using a tube (d) reinforcement installation. Dia-phragm wall panels may reach important dimensions (up to 1.2 m width, 8 m length and 60 m depth are not exceptional).

There are several reasons why the study of the ef-fects due to panel installation is interesting. First (di-rect effect) to assess the possible impact of the panel installation on its environment, as sometimes panel collapse occurs and, quite more frequently, signifi-cant movements in the vicinity of the panel are regis-tered during construction. Second (indirect effect) because even if ground movements induced by panel execution are small, they may lead to stress changes in the environment relevant to later excavation stag-es.

Direct effect of panel installation has more signifi-cance in the case of soft soils. The indirect effect is more important in hard overconsolidated soils. As a first approximation, ignoring the indirect effect tends to leave design on the side of safety, since the broad consequence of panel excavation is horizontal stress relaxation. Ignoring the direct effect instead, is not safe, since the movements induced by the diaphragm

wall installation are added to those produced by the subsequent excavation of the structure.

Given the increasing requirements regarding limi-tation of ground movements in urban areas, devel-opment of tools to evaluate the effect of the installa-tion of diaphragm walls in soft soils is necessary. In the paper, we briefly review some observations on the problem available in the literature and present modeling results relevant to a real case.

2 STATE OF THE ART

2.1 Field observations Although well documented cases of instrumented excavation of diaphragm wall panels are rare, there are enough in the literature (Di Biagio and Myrvoll 1973, Poh and Wong, 1988, Tsai and others 2000, De Wit and Lengkeek , 2003) to obtain some general lessons. Table 1 lists the main geometric features of these cases, i.e. the width (W), length (L) and depth (D) of the studied panel. A description of soil type and a measure of the characteristic resistance (CR) are also included for each case. The correspondent values are also given for the case study presented later in this paper.

Where panel installation does not lead to failure (a) final superficial settlements measured at the edge of the panel vary between 10 and 20 mm (b) those settlements attenuate rapidly with the distance from the panel, being negligible for distances greater than 10W (Fig. 1) (c) horizontal displacements at a cer-tain depth tend to reach values several times higher than the surface settlements. Its attenuation with dis-

Coupled analysis of movements induced by diaphragm wall installation

B. Garitte, M. Arroyo & A. Gens Department of Geotechnical Engineering, UPC, Barcelona, Spain

ABSTRACT: The construction process of diaphragm walls can lead to movements in the surrounding area that are seldom taken into account. However, these movements may be important in situations where soft soils dominate. In this paper, after briefly reviewing the state of the art on this issue, we present results of a hydro-mechanical simulation of the problem in a case located in deltaic soils of Barcelona. The auscultation record of the settlements of a nearby building is employed to validate the computational model. Panel length and bentonite slurry level were found to be the most influent parameters on induced displacements.

Preprint: 7th European Conference on Numerical Methods in Geotechnical Engineering, NUMGE 2010, Trondheim, Benz & Nordal, eds., Taylor & Francis pp 547-552

tance from the panel follows a law similar to the sur-face settlement.

Table 1. Cases of panel installation experiments in the littera-ture.

Case Soil CR W(m) L(m) D(m)

Barcelona (this study)

Sand qc = 5 – 10 MPa 1-1.2 2.8-6 25-35 Clayey

silt Su=40–80 kPa

Oslo (Di Biagio and Myrvoll,

1973)

Soft clay

Su=30-40 kPa 1 1.8- 5 28

Singapur (Poh y Wong,

1998)

Soft clay

Su=14-40 kPa 1.2 2 - 6 55.5

Taiwan (Tsai et al.,

2000)

Silty sand

qc=20–40 MPa 0.9 8 15

Amsterdam (De Wit and Lengkeek,

2002)

Clay and turf

Su=30-40 kPa 0.8 3 35

Sand qc=20 MPa

Figure 1. Effect of the installation of a panel on the superficial settlements (Poh and Wong, 1998).

Some characteristics of the construction process are clearly reflected in the observations. For example, settlements and movements toward the panel that occur during the excavation phase decrease during concreting. Occasionally, the movement direction is even reversed (Fig. 2). Unloading (excavation) and loading (concreting) during the installation of a pan-el also produce pore water pressure changes in less permeable materials. In that case, measurements dur-ing the excavation correspond to a situation of par-tial drainage and more movements may occur during the subsequent consolidation phase.

Other interesting observations are those concern-ing the effect of variations of the basic construction process. In practice, such variations may be due to

work incidents (a) excavation induced movements are very sensitive to the sustaining fluid level in the panel trench. Apparently small decreases (e.g. 1m for a 55m deep panel) can produce significant addi-tional movements and even the collapse of the ex-cavation (b) movements are not increased signifi-cantly during important waiting times (e.g. one day), as long as a constant fluid level is maintained inside the trench. (c) Once the impervious cake has devel-oped satisfactorily on the panel wall, excavation in-duced movements are relatively insensitive to the density of the sustaining fluid (in the range 1 to 1.3 t/m3).

One aspect that is not clear from the cases de-scribed in the literature is the influence of the size of the panel on induced movements. Greater displace-ments are expected to be induced by panels of larger length and depth, but there are few data to quantify this phenomenon.

Figure 2. Settlement at 3m from the panel measured with an hydraulic cell (Di Biagio y Myrvoll, 1973).

Figure 3. Measured fresh concrete pressure in a 45m deep pan-el (Schad et al., 2007)

Finally, observations have repeatedly confirmed (Uriel and Oteo, 1977, Lings et al., 1994; Schad et Al., 2007) that the pressure applied by fresh concrete on the panel walls is not always hydrostatic. In fact,

hydrostatic pressure seemed to be maintained only until a certain depth, referred to as the "critical depth" (Fig. 3). The observed critical depths vary be-tween 1/3 and 1/5 of the panel depth. This character-istic is still poorly understood, but seems to depend on the interplay between concreting rate and harden-ing.

2.2 Analytical solutions There are a number of analytical solutions to the problem of panel stability (e.g. Fox, 2004; Tsai, 2000). Most of them were derived using limit equi-librium and the differences between them lie mainly in the degree of complexity of the alleged failure surface. A major drawback of these analytical solu-tions is their limited applicability for layered pro-files. Moreover, only stability is dealt with and they are not useful for quantifying the settlement induced by panel excavation.

However, it is interesting that these analytic solu-tions also indicate the enormous importance of the level of bentonite on the panel stability. For instance, Fox (2004) predicts a security factor reduction from 2 to 1, by a decrease of bentonite level of 2m in a 15m deep trench of 8m length in sand (φ’ = 34º, γ = 20 kN/m3).

2.3 Numerical models An interesting alternative to obtain quantitative an-swers to the panel problem is to use numerical mod-els. However this approach is relatively costly. For practical reasons numerical modeling of excavations still takes place mostly in 2D. In such models, dia-phragm wall construction is simultaneous for the en-tire wall length. Available examples (Ng and Yan, 1998; Gourvenec and Powrie, 1999, Schafer and Triantafyllidis, 2006) make clear that it is very diffi-cult to obtain approximate results if the three-dimensionality of the problem is ignored. The cited authors also emphasized the importance of the initial earth pressure coefficient K0, because of its influ-ence on stress redistribution.

Several common modeling features can be noticed from the precedents (1) guide wall construction is not considered; (2) excavation under bentonite is re-produced by removing the elements included in the volume of the panel and prescribing the hydrostatic pressure of bentonite on the new contour; (3) fresh concrete pouring is represented by changing the boundary condition of total stresses from the hydro-static bentonite profile to a bilinear profile; (4) final-ly, to represent hardened concrete, the total stress boundary condition is removed and elements in the panel volume are re-activated, with material parame-ters corresponding to those of reinforced concrete. Note that, while this procedure is in accordance with the above mentioned field observations, it does ne-

glect the tangential frictional stresses between soil and fresh concrete that will be necessary for equilib-rium. The critical depth is usually taken equal to one third of panel depth.

3 CASE STUDY

3.1 Background The case that inspired the studies described here

can be considered typical of the excavations in soft soil of deltaic areas near Barcelona. The motivation arose from the observation of significant movements in a building near some excavation works during di-aphragm wall installation.

Construction activities were complex because most diaphragm wall installation was simultaneous to other potentially disturbing activities, like jet-grout treatment and micro-pile installa-tion. However, the construction sequence began with the execution of a diaphragm wall section near the building (Fig. 4). During that period, which preceded all other construction activities, significant building movements were already registered (Fig. 5). Those records made clear that diaphragm wall construction had produced some movement, but they did leave open the magnitude, because settlement occurring after the diaphragm wall section was finished might have been due to consolidation or to other, later ac-tivities. Since similar diaphragm walls needed to be constructed in the vicinity a detailed study of this problem by means of numerical simulations seemed necessary.

Figure 4. Plane view of the study area . The investigated dia-phragm wall (R1-5) and reference measurement points (P1-7) are indicated.

3.2 Geotechnical site characterization The geotechnical profile at the site might be de-scribed by five main levels. Made ground (2m thick), clay (2m), sand (11m), silt (33m) and gravel (unde-fined). The water table is found at the top of the sand layer. The geotechnical site characterization proce-dure cannot be described here other than it was heav-ily reliant on in situ tests. The in situ measurement campaign also provided information on the earth pressure coefficient K0 which was used to prescribe initial stress state.

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Figure 5. Settlements measured during the execution of the R-diaphragm wall. Vertical dotted lines indicate the day in which each panel was built (R5 to R1). The silt package is slightly overconsolidated and has a relatively high deformability. It is likely that much of its deformation takes place under plastic re-gime. Hence, the characterization of this regime has some importance and a modified Cam-clay (MCC) model was chosen. For the other materials a simpler Mohr-Coulomb (MC) model was selected. The most important geomechanical parameters employed in the calculations are shown in Table 2.

Table 2a. Geomechanical parameters (MC)

κ E ν c φ

(m/d) (kPa) (kPa) (º)

Made ground 0.22 3000 0.3 30 25

Clays 0.005 15,000 0.3 50 29

Sand 31.3 20,000 0.3 1 33

Gravel 105 31,000 0.3 50 36

Table 2b. Geomechanical parameters (MCC)

κ φ OCR κ λ (m/d) (º)

Silt 0.003 27 1.15 0.012 0.06

3.3 Characteristics of the numerical models Precedent published analyses of this problem mod-eled clayey soils, for which undrained behaviour might be safely assumed. The geotechnical profile in this case includes layers of very different permeabil-ity and any generic assumption about drainage was not granted. For this reason fully coupled hydrome-chanical computations were performed.

The program employed was Code_Bright, a finite element code developed in the Department of Ge-otechnical Engineering of UPC (Olivella, 1995). The modeling of the construction process of a panel re-quired some modifications to the program, the most important being the implementation of a boundary condition with (bi)linear stress variation with depth. The new implementation was verified by bench-

marking against a case reported in the literature (Gourvenec and Powrie, 1999).

Two types of analysis were performed: a paramet-ric study of the excavation of an isolated panel and a detailed modeling of a particular excavation se-quence of five panels, namely the R-diaphragm wall (Fig. 4). Figure 6 shows the mesh used for modeling the R-diaphragm wall excavation sequence (the mesh used to simulate the installation of one panel is similar). Advantage was taken of the vertical longi-tudinal symmetry plane of the panels.

Figure 6. 3D mesh used for the analysis of the R-diaphragm wall excavation sequence.

Coupled hydro-mechanical computations require the explicit specification of construction times. After some consultation with the site managers and in-spection of construction records, a site-representative construction sequence was established for the base case. Excavation under bentonite sup-port is modeled by removing meter by meter the el-ements of a panel during 3.2 hours in the base case (25 m deep panel). Once a certain volume has been removed in a panel, hydrostatic bentonite pressure is applied as total stress on the new wall. After the ex-cavation sequence finishes, a five hour waiting time represents bottom cleansing and reinforcement placement. The hydrostatic bentonite pressure profile is then replaced by the bilinear profile representing fresh concrete. Concrete hardening time was esti-mated as 12 hours, after which solid elements with concrete properties are placed in the panel. Concrete hardening is thus modeled as an instantaneous pro-cess, which is clearly unrealistic.

4 RESULTS

4.1 Base case Geometric features of the base case (depth D, length L and width W) are shown in Table 3. Bentonite lev-

el within the panel, nb, and the assumed critical depth Dc, are also indicated.

Surface settlement histories at different distances from the panel edge are given in Figure 7. Times at which excavation ends, fresh concrete is poured and hardening is assumed are indicated by vertical dotted lines. Excavation produces settlement, reaching about 3.5mm at 2m from the panel edge. Most set-tlements induced by the excavation occur simultane-ously to it and later consolidation has a moderate in-fluence. Fresh concrete deposition results in heave and after concrete hardening, settlements are re-sumed. Final settlement values are similar to those registered after excavation. It is worth noting the qualitative similarity with the measurements by Di Biagio and Myrvoll (Fig. 2), especially during fresh concrete injection.

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End of excavation

Concrete hardening

Fresh concrete pouring

Figure 7. Time evolution of settlements at different distances from the panel edge (base case). Table 3: Modeling of single panel. Parametric study cases.

Case D(m) L(m) W(m) nb (m) Dc (m) Base (V1) 25 3.6 1.2 0 8 V2 25 3.6 1.2 2 8 V3 25 6 1.2 0 8 V4 35 3.6 1.2 0 8 V5 25 3.6 1.2 0 5 V6 25 3.6 1 0 8

4.2 Parametric study of a single panel The parametric study includes five variations on the base case. As outlined in Table 3, a single parameter was changed from the base case for each variant. In variant 2 (V2) the level of bentonite, nb, was lowered by 2m; in variant 3 (V3) a length of 6m is considered for the panel; in variant 4 (V4) a deeper panel was excavated up to 35m; in variant 5 (V5) the critical depth was modified to 1/5 of the panel depth instead of 1/3 and in variant 6 (V6) the panel was made thinner.

A comparison between the different cases is given in Figure 8, in terms of settlement evolution and hor-izontal displacement of the panel wall. According to the simulation results, the most damaging cases are

V3 and V2, i.e. a longer panel and a lower bentonite level. Differences between the other cases are negli-gible for superficial settlement measurements, alt-hough, as expected, a deeper panel causes higher movements at depth.

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Figure 8. Left graph: evolution of settlements at 3m from the panel edge for the 6 cases of the parametric study. Right graph: horizontal displacements of the panel wall after the excavation phase for the 6 cases of the parametric study.

4.3 R-diaphragm wall construction sequence The R-diaphragm wall panels (Fig. 4) have similar dimensions to the base case panel: all panels have a length of 3.6m, excepted one (R1), which is 4.9m long and the depth of the wall is 24m instead of 25m in the base case. The construction sequence was simulated according to work site infor-mation. Bentonite level and critical depth were taken as in the base case because no measurements were available.

This computation confirmed some of the previous results. For instance, the installation of one 6m long panel produces more settlement than the consecutive installation of two 3.6m long panels (Fig.9). The dif-ference in settlement behaviour appears mostly dur-ing the excavation phase. After concrete hardening, quite similar final settlement levels are predicted for the 6m long and the joint 7.2m long excavation.

Simulated and measured settlements at observa-tion points P1 and P2 are compared in Figure 10. The position of those points was illustrated in Figure 4. Quite a good agreement between measurements and simulation is obtained. One remarkable differ-ence is the smooth evolution of the measurements when compared with simulation results. This is like-ly due to the abrupt modeling approach adopted for concrete hardening (instantaneous).

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Figure 9. Superficial settlement simulated at a distance of 3m from the panel wall for two panels of 3.6m length and one pan-el of 6m length.

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Figure 10. Comparison of measured and simulated settlements at observation points P1 and P2. Measured and simulated settlement profiles after construction are drawn in Figure 11. Two simulation profiles have been plotted: one at the level of R4, corresponding to observation point P1 and one at the height of R1-R2, corresponding with the remaining points. The fit is quite good for measurement points close to the wall, but its quality decreases with dis-tance to the wall. The observed discrepancies are probably due to the fact that the building structure is not taken into account in the computation. The struc-ture may act as a stiffening element, hindering set-tlement recovery during fresh concrete injection. The overall good match between simulation results and measurements can be considered satisfactory and gives more credibility to the results of the parametric study.

CONCLUSIONS Diaphragm wall installation in soft soils may pro-duce settlement in its neighborhood. Numerical models may help to quantify and understand the problem. The presented parametric study allow for isolating two influent parameters: bentonite level and panel length. Other parameters, like panel width, depth and critical depth were found to be less im-portant.

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Measurements @ 12/02/2007 (other plots)

Figure 11. Measured and simulated settlement profiles after construction.

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on environmental impact of diaphragm wall trench installa-tion in Amsterdam – the final results. Proceedings of the in-ternational symposium on geotechnical aspects of under-ground construction in soft ground, Toulouse, France (eds R. Kastner, F. Emeriault, D. Dias and A. Guilloux), pp. 433–440. Lyon

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Fox, P.J. (2004) Analytical solutions for stability of slurry trench, ASCE Journal of Geotechnical and Geoenvironmen-tal Engineering, Vol. 130, No. 7, 749-758

Gourvenec, S. M. & Powrie, W. (1999). Three-dimensional fi-nite-element analysis of diaphragm wall installation. Géotechnique 49, No. 6, 801- 823

Lings M, Ng CWW, Nash DFT. (1994) The lateral pressure of wet concrete in diaphragm wall panels cast under bentonite. Proceedings of the Institution of Civil Engineers: Geotech-nical Engineering; 107:163–172.

Ng CWW, Yan RWM. (1998) Stress transfer and deformation mechanism around a diaphragm wall panel. Journal of Ge-otechnical and Geoenvironmental Engineering; 128(7):638–648.

Olivella, S., 1995. Nonisothermal multiphase flow of brine and gas through saline media. Doctoral Thesis, Technical Uni-versity of Catalonia (UPC), Barcelona, Spain.

Poh TY, Wong IH. (1998) Effects of construction of diaphragm wall panels on adjacent ground: field trial. Journal of Ge-otechnical and Geoenvironmental Engineering; 124(8):749–756

Schad, H., Vermeer, P.A., Lächler, A. (2007) Fresh concrete pressure in diaphragm wall panels and resulting defor-mations. In: Grosse, Ch. U. (Ed.): Advances in Construction Materials, Berlin: Springer Verlag, 2007, pp. 505-512.

Schafer R, Triantafyllidis T. (2006) The influence of the con-struction process on the deformation behaviour of dia-phragm walls in soft clayey ground. International Journal for Numerical and Analytical Methods in Geomechanics; 30:563–576

Tsai, J.S., Jou, L.D., Hsieh, H.S. (2000) A full scale stability experiment on a diaphragm wall trench, Canadian Geotech-nical Journal, 37, 379-392

Uriel S. y Oteo C. S. (1977) Stress and strain beside a circular trench wall. Proc. 9th ICSMFE, Tokyo, 1,781-788.