1. 1. Vertical stress distribution in soil bentonite slurry walls Rahul V Mukherjee MDH Engineered Solutions, Member of SNC Lavalin Group, Saskatoon, Saskatchewan, Canada. Moir D Haug Department of Civil Engineering, University of Saskatchewan, Saskatoon, Saskatchewan, Canada. ABSTRACT Backfill arching or wall hang-up of backfilled slurry wall trenches has the potential to significantly reduce the effectiveness of these walls. Soil-bentonite (SB) slurry walls are one of the most popular techniques to minimize the horizontal migration of contaminants. This technique involves the excavation of a narrow trench in the presence of bentonitic slurry down through permeable strata and backfilling with a low permeability material. This paper presents the results of a laboratory testing program and modelling study to investigate the vertical stress distribution in slurry walls. The program was conducted to supplement the design and installation of an 11,000 m long slurry wall. This slurry wall is being installed through low hydraulic conductivity glacial till containing intermediate permeable zones. The depth of this 1 m wide slurry wall is approximately 50 m. Backfill hang-up in these deep walls is a particular concern. In order to calculate hang-up it is necessary to determine friction along the wall of the trench. The testing program involved large strain consolidation testing of selected field backfill samples. A Marriott bottle system was also employed to continuously measure hydraulic conductivity and lateral stress during testing. Additional backfill samples were tested to provide data on sensitivity of horizontal and vertical stress to variations in backfill fines content. An analytical model and a coupled seepage stress-strain finite element model (FEM) were used to predict vertical stress changes with time and depth for the different backfill materials. The results of the research study found that trench width was the most important variable in determining wall hang-up. It also showed that the upper portions of the backfilled trench may take years to consolidate and thus provide the desired level of impermeability. RSUM La vote ou la paroi moule de tranches remblayes a le potentiel de rduire significativement lefficacit de ces murs. Lutilisation de parois moules constitues dun mlange de sol-bentonite (SB) est lune des techniques les plus populaires pour limiter la migration horizontale des contaminants. Cette technique implique lexcavation dune tranche troite, en prsence dune suspension de bentonite, travers des strates permables et le remblayage avec un matriau faible permabilit. Cet article prsente les rsultats dun programme dessais en laboratoire ainsi que dune tude de modlisation de la distribution des contraintes verticales dans les parois moules. Le programme a t ralis dans le but de complter la conception et linstallation dune paroi moule de 11 000 m de long. Cette paroi est en cours dinstallation dans un till faible conductivit hydraulique contenant des zones permables intermdiaires. La profondeur de cette paroi moule de 1 m de large est denviron 50 m. Les conditions daccrochage du remblai le long de ces parois profondes est une importante proccupation. Afin de calculer laccrochage, il est ncessaire de dterminer le niveau de friction le long de la paroi de la tranche. Le programme dessais comprend des essais de consolidation grande dformation dchantillons slectionns de remblai. Un systme de bouteille Marriott a galement t employ pour mesurer en continu la conductivit hydraulique et la contrainte latrale durant les essais. Des chantillons de remblai supplmentaires ont t tests afin de fournir des donnes sur la sensibilit des contraintes horizontales et verticales en fonction de la fraction de particules fines contenues dans le matriel de remblai. Un modle analytique et un modle par lments finis (FEM) ont t utiliss pour prdire les changements de contraintes verticales en en fonction du temps et de la profondeur de chacun des diffrents matriaux de remblai. Les rsultats de ltude ont montr que la largeur de la tranche tait la variable la plus importante pour dterminer laccrochage aux parois. Il a galement t dmontr que les parties suprieures de la tranche peuvent prendre des annes avant dtre consolides et donc fournir le niveau dimpermabilit dsir. 1. INTRODUCTION The potential for slurry wall backfill arching and hang-up was examined in a laboratory testing program. The permeability of clay till slurry backfill is highly dependent on the load placed on the backfill. As a result permeability usually decreases with depth as the backfill consolidates slowly under increasingly high loads. If the width of the backfill trench is narrow and the depth great, there is a potential for arching and backfill hang-up. Laboratory testing was carried out to characterize changes in hydraulic conductivity and the coefficient of lateral earth pressure with increase in effective stress. Soil-bentonite (SB) slurry walls are one of the most popular techniques to minimize the
2. 2. horizontal migration of contaminants. Backfill arching or wall hang-up of the backfilled trenches has the potential to significantly reduce the effectiveness of slurry walls. A 45 m plus deep, 11,000 m long slurry wall is currently being constructed at a Potash mine in eastern Saskatchewan. This wall is being constructed through high clay content drift. The slurry wall is being keyed into low permeability upper Cretaceous clay-shale. The purpose of this deep slurry wall is to cutoff seepage of brine through stacked aquifers which exists between the glacial till under the site. These aquifers represent pathways for the migration of chloride impacted fluids. 2. BACKGROUND Soil bentonite (SB) slurry walls are vertical barriers constructed by excavating a narrow trench through high hydraulic conductivity material or zones and backfilling with a mixture of soil, bentonite and water to achieve a low hydraulic conductivity barrier (Xanthakos 1979, Haug 1983, Evans 1995). The backfill soil either comes from the excavated trench or from a source nearby, if the soil is deemed to be unsuitable for use. The slurry in the SB wall is a mixture of water and dry bentonite (4% - 7%) by weight (Evans 1994). Before placement in the trench the backfill is mixed with desired amount of slurry from the trench to provide a desired consistency. Since SB walls are mainly used as low hydraulic conductivity barriers and constructed as a part of polluted site remediation process, hydraulic conductivity value plays an important role. In general, they range from 10 -7 m/s to 10 -11 m/s (DAppolonia 1981, Haug 1987, Evans 1995, Yeo et al 2005). The hydraulic conductivity for the backfill mixture depends upon many factors such as (i) the amount of dry bentonite added to the mix (Ryan 1987, Evan 1994, Yeo et al 2005), (ii) the amount of fines in the mixture (Xanthakos 1979, DAppolonia 1981, Yeo et al 2005), (iii) resistance of the mix against the contaminant being contained and (iv) the state of stress in the slurry wall (Evans 1995, Filz 1999). The performance assessment of these walls in terms of hydraulic behaviour is generally done through laboratory testing of the backfill mixture. Laboratory models and field studies on the slurry walls have shown that the stress in the wall does not increase linearly with depth as assumed earlier (McCandless and Bodocsi 1987, Evans 1995).The friction at the trench/backfill interface governs the consolidation behaviour of the backfill- slurry mix, which is fairly compressible. This phenomenon can be explained by arching. 3. FOCUS OF THE STUDY Backfill hang-up and arching is of particular concern for deep slurry walls. This potential is a function of the width of the trench, the depth of the trench, shear strength of the backfill and potentially the sequence and spacing of permeable zones cut-off by the slurry wall. In 2009, construction was started on an 11,000 m long, 45 m plus deep slurry wall through glacial till at a Saskatchewan potash mine. This slurry wall is intended to control the migration of chloride impacted water. The hydraulic conductivity and the potential for osmotic consolidation is a function of the stress environment of the trench backfill. As a result, the potential for arching or wall hang-up is a concern. Further complicating the situation is the presence of stacked aquifers which will result in more rapid backfill consolidation at various depths in the trench. Figure 1 shows the general situation. Figure 1 Schematic diagram of complex site geology 4. EXPERIMENTAL PROGRAM The experimental program involves slump cone tests, direct shear test, large strain consolidation testing along with hydraulic conductivity measurement of the backfill mixes. A Marriott bottle system was also used to enable continuous monitoring of hydraulic conductivity during loading. Miniature load cells were installed on the walls of the mold at different depths to measure the stress profile. Figure 2 shows a sketch of the large strain modified consolidation mold used in the present study. 4.1 Test Materials A glacial till brought from the site was used as a host soil for preparing the backfill mixes. The natural water content of till was about 6% 8%. Air dried and pulverized till passing through 4.75 mm was used for preparing the trial backfill mixes. The hydraulic conductivity of the till was about 1 x10 -8 m/s. The preconsolidation pressure of the host soil ranged from 1200 1500 kPa. A total of four different backfill mixes were examined in the testing program. Three backfill trial mixes (TM1, TM2 and TM3) were prepared in the laboratory by varying the fines content (particle size < 75 m) and the fourth mix was collected in the field (FB) from the site of the slurry wall construction. The percentage of fines in the backfill mix made in the laboratory (TM1, TM2 & TM3) was varied as 10%, 25% and 50% respectively. Table 1 shows the description of the backfill mixes used in this study. Bentonite slurry was prepared in the laboratory in the same proportion as that in the field to simulate the field conditions. The bentonite was mixed with tap water using a high-speed colloidal shear mixer. Then the bentonite-water
3. 3. slurry was allowed to hydrate for 48 hrs 72 hrs in order to properly hydrate the bentonite powder. The density of the slurry was 0.86 g/cc and the Marsh cone viscosity was 39 s (API 1990). The pH of the slurry was about 9.20. Table 1 Description of backfill mixes Mix Percent of fines without bentonite Percent of fines with bentonite Amount of bentonite in the mix Trial Mix1 (TM1) 10 12.9 2.9 Trial Mix2 (TM2) 25 28.9 3.9 Trial Mix3 (TM3) 50 52.8 2.8 Field Backfill (FB) N.A 28 N.A 4.2 Experimental Setup A test apparatus for measuring the co-efficient of lateral earth pressure (K) during large strain consolidation was developed. This apparatus was based on original design by Gan et al; 2011 (Nov 2011). The Proctor-type stainless steel mold shown in Figure 2 (150 mm diameter) was modified to measure the co-efficient of lateral earth pressure and consolidation characteristics of the backfill mixtures under increasingly large strains. Hydraulic conductivity (k) of the test mixes were also measured simultaneously. Button type load cells were installed in the mold to measure the lateral pressure exerted on the wall of the mold by the backfill mixes during consolidation. The mold was modified to contain ports for measuring hydraulic heads at 8 different depths. The bottom outlet of the mold was connected to a Marriott bottle setup to continuously run constant head hydraulic conductivity test along with consolidation. The modified mold (Figure 2) was loaded by a Conbel up to 900 kPa. Two load cells and the linear potentiometer were connected to a computer by a data logging system. Lab view software was used to communicate with the data logging system to record the raw data automatically on to the computer. A Marriott bottle was connected to the bottom outlet of the mold to run the constant head hydraulic conductivity (k) test simultaneously along with consolidation. All the eight ports in the mold were connected to a manometer to enable reading heads at different heights in the sample. 4.3 Test Procedure Preliminary tests such as determination of Atterberg limit were performed on all the soils used in the test program, namely the glacial till (host soil) obtained from the field, the field backfill (FB), TM1 (having 10 % fraction passing 75 m), TM2 (having 25 % of fraction passing 75 m) and TM3 (having 50% of fraction passing 75 m). Slump cone tests were conducted on the samples. This helped in determining the consistency of the SB mix. This is an indirect way of measuring the flowability of the mix. In the present work, each backfill mix was mixed with 5% bentonite-water slurry in various proportions to evaluate the relationship between the gravimetric water content of the resulting backfill-slurry mixture and the resulting slump. The slump was measured according to ASTM C 143, and the slump tests performed for each specimen at a given water content were repeated two to three times to account for variability in the measured values of slump. Figure 2 Modified mold Consolidated drained (CD) shear box tests were conducted to study the shear strength parameters of the backfill mixes. All the samples were sheared under increasing normal pressures of 54.7 kPa, 163.2 kPa and 316 kPa. In total 12 samples were tested. Samples from different mixes were allowed to consolidate in large consolidation molds at different specified normal loads. After, consolidation the samples were extruded and trimmed to fit into a 150 mm x 150 mm diameter direct shear box. The samples were left for 1 - 3 days depending upon the normal load. The shear box test was then carried out according to ASTM D-3080-04. The saturated samples in the molds were compacted at a water content corresponding to a slump of 120 mm and at a constant dry density. The depths in the field were simulated corresponding to the load applied onto the sample. The samples once compacted were left to consolidate under the weight of the loading plate, till the load cell and potentiometer reading stabilized. The whole setup was then loaded using Karol-Warner Conbel. Load cell and deflection readings were recorded and monitored continuously. A constant head hydraulic conductivity test was run simultaneously, with use of the Marriott bottle arrangement (Figure 2). Once the equilibrium was attained, the next incremental load was applied to the sample. 5. DISCUSSION OF LABORATORY TEST RESULTS Figure 3 shows, the failure envelopes of TM1, TM2, TM3 and FB obtained from the shear box test data. The effective angles of internal friction () measured from the failure envelope were 32 , 29 , 26 o and 23 o for TM1, FB, TM2 and TM3, respectively. This was because the sand contents in TM1, TM2 and TM3 were 87%, 71% and 47.2% respectively. The amount of fines present in the mix also justifies the values. These results were used as an input, for modelling consolidation of the SB wall. These values are
4. 4. considerably lower than 35 degrees previously reported in the literature. Figure 3 Failure envelope for different types of backfill mixes Figure 4 shows the e - log p curves for all the test materials tested for compressibility in a large proctor mold, the diameter of which was 150 mm. The sample thickness was 160 mm. Consolidation test was performed by applying air pressure on the sample and by monitoring settlement of the sample. The equilibrium void ratio corresponding to the applied pressure was plotted against log of applied pressure as shown in Figure 4. TM3 showed highest compressibility because it had highest fines content and the other mixes, namely TM2, FB and TM1 which had lower fines content resulted in reduced compressibility. Figure 5 shows the variation of measured hydraulic conductivity, k (m/s) with the applied vertical stress. These tests were run simultaneously with consolidation. This was done using Marriott bottle arrangement as described in previous section. All the tests were run 3 times to account for variability in results. The data shows that hydraulic conductivity (k) decreased with increasing in vertical stress (i.e. decrease in equilibrium void ratio). The data also show that the hydraulic conductivity value of TM3 which had highest fines content of all the materials was lower than other materials. The applied pressure hydraulic conductivity relation for FB and TM2 was close because the fines content (silt and clay) in these materials was nearly equal. TM1, which had the least fines content and highest coarse content resulted in the highest value of hydraulic conductivity corresponding to the equilibrium void ratio. Figure 4 e log p curves from large strain consolidation test Figure 5 Hydraulic conductivity (k) plots for various backfill mixes The hydraulic conductivity for TM1 decreased from 1x10 - 4 m/s to 2.5x10 -7 m/s with the increase in stressfrom 9 kPa to 900 kPa. Whereas the k for TM3 decreased from 8x10 -9 m/s to 4x10 -10 m/s with the same increase in stress. Hence, the amount of fines present in the mix plays an important role in controlling k. Similar results have been reported in literature by Yeo et al, 2005 where hydraulic conductivity for their test mixes also decreased with the increase in fines content. 6. MODELLING A numerical method based on principles of statics (discrete model) and secondly the finite element method was used for this study. This method does not consider time dependent behaviour such as secondary consolidation, creep etc. The results of modeling using both methods have been discussed and comparisons between both have been drawn. 6.1 Discrete model Discrete model is based on the principle of statics. A body is said to be in a state of static equilibrium, if and only if all the forces acting on it are balanced. Assumptions in statics model include; 1. The co-efficient of lateral earth pressure, K = h /v where h and v are horizontal and vertical effective stresses (Figure 6). These are not principal stresses. 2. For a given trench width, at any depth,v is uniform across the trench. 3. Co-efficient of lateral earth pressure (K) is constant all throughout the wall. 4. Friction is fully mobilized along the two vertical interfaces of the slurry wall. Note that this solution does not take into account stress- strain relations of the material.
5. 5. Figure 6(a) shows a trench of B unit wide which has been discritised into small elements of thickness, h. Consider an element of thickness h, at a depth of h from the ground surface (Figure 6a). vn is the stress acting on the top of the element which is the cumulative stress being transferred by the elements above it. Figure 6(c) shows the free body diagram for the shaded element under consideration. Pn (force acting downwards) = Force being transferred due to the normal stress to the top of the element under consideration by the elements above it. w is the self weight of the element. F/2 is the frictional force acting upwards on each side of the block. Pn+1 is the net force acting upward on the segment to keep it in static equilibrium. From equilibrium, Figure 6(c) Po + W P1- 2*F/2 = 0 P1 = Po + W 2*F/2 Writing the forces in terms of stresses, vn+1 * B * 1= (vn * B * 1) + ( * B *1 * h) 2*F/2 Where, F/2 = (tan ) h h= (tan ) Kv h vn+1 = (vn) + ( * h) 2*(tan Kvn/B) h [1] Hence, from the above equation, vertical stress on any element is the function of the co-efficient of lateral earth pressure K, width of the wall B and angle of internal friction, 6.1 Parametric study using discrete model The variation of vertical effective stress in the wall, with depth, for different width of the wall and for K = 0.35 and = 29 is shown in Figure 7. The width (w) of the wall is varied as 0.5, 1, 2, and 4 m. The dashed line (w = 1m) represents the stress distribution for the SB wall. The vertical stress at any depth moves towards the geostatic stress (h) distribution with an increase in the width of the wall. Thus it can be concluded that arching reduces as the width of the wall increases. Figure 7 shows that for all trench widths the stress up to 2 m depth is approximately equal to geostatic stress. Below that depth, the vertical stress deviates significantly from geostatic stress. Figure 8 shows the variation of vertical effective stress (v) with depth of the wall for different angles of internal friction (for K = 0.35 and w =1 mBased on Rankine active and passive earth pressure theory, K should lie between 0.35 and 3.25. This is the lowest and highest values of K calculated using the value shown in the figure. The angle of internal friction, 'for the SB mixes varied from 2 to 2 . he dashed line in Figure 8 represents the properties of backfill from the slurry wall construction site. The width of this wall was 1m. Figure 6. Discrete Model Figure 7 Variation of vertical effective stress (v) with depth for different widths of the wall Figure 8 Variation of v with depth for different angle of internal friction () of the SB mix
6. 6. Stress distribution in the wall was found to move further away from the geostatic stress with increase in '. For all the values of ', v in top 3 m, is approximately equal to geostatic stress and thereafter starts to move away (less than geostatic stress). As ' increases, the vertical effective stress decreases. This is because friction along the side walls acts upward and is function of '. Hence, with increase in ', the backfill has increased resistance to settlement. It also resists consolidation and strength gain. Figure 9 Variation of v with depth for different values of K The vertical effective stress in top 2 m for all values of K is approximately same. The stress at D = 8 m, 20 m and 50 m decreases as K increases from 0.25 to 0.75. Based on Rankine earth pressure theory for 29; K should lie between 0.34 and 2.88. The percentage difference between the geostatic stress and v at 50 m depth for K = 0.35 and 0.75 is 95.4% and 98%. This difference is not overly significant. Hence, K has little impact on arching in the SB walls. Figure 10 Vertical stress distribution for different SB mix used in the present study It can be concluded that arching in slurry wall can be reduced by selectively choosing the mix parameters. The width of the wall plays a significant role in the performance of SB walls. If possible, the wall should be wider if the trench is deep (Figure 7). The stress distribution using trial mix 3 (TM3) would result in the least arching (Figure 10). The vertical effective stress at 50 m depth is 22 kPa, 25 kPa, 29 kPa and 36 kPa for TM1, FB, TM2 and TM3 respectively. Stresses are about 90% of the geostatic stress (h) for a given depth. 6.2 Finite Element Model A finite element model (FEM) of Soil Bentonite cutoff wall was developed for this research program. The model simulates the consolidation of the backfill material. Finite element software Geo-studio 2007 was used for analysis. Seep and Sigma packages in Geostudio were coupled for this work to simulate fully coupled behaviour. The coupled analysis simultaneously solves two groups of nodal equations, equilibrium (stress-deformation) and continuity (flow) equations, across the finite element mesh (GEOSLOPE 2007). The basic material-parameters required for such analysis on fully saturated systems are: Youngs Modulus, E; Poissons ratio, ; Unit weight of the material, ; and saturated hydraulic conductivity, k. However, the software does not take in to account secondary consolidation and the deformation due to creep. The model-domain was delineated as a 1 m wide soil bentonite column from the top of the ground surface and keyed at a depth of 50 m into a stiff material such as heavily consolidated clay or bedrock (which is shale in this case). Figure 11 shows the simplified representation of the complex site geology. Only fully saturated SB wall was modelled, with water table at the top. The hydraulic and stress boundary conditions used in the model are laid out in the Table 2. Figure 11 Cross-section of aquifer used for the present study 4.3 Test Procedure Preliminary tests such as determination of Atterberg limit were performed on all the soils used in the test program, namely the glacial till (host soil) obtained from the field, the
7. 7. field backfill (FB), TM1 (having 10 % fraction passing 75 m), TM2 (having 25 % of fraction passing 75 m) and TM3 (having 50% of fraction passing 75 m). Slump cone tests were conducted on the samples. This helped in determining the consistency of the SB mix. This is an indirect way of measuring the flowability of the mix. In the present work, each backfill mix was mixed with 5% bentonite-water slurry in various proportions to evaluate the relationship between the gravimetric water content of the resulting backfill-slurry mixture and the resulting slump. The slump was measured according to ASTM C 143, and the slump tests performed for each specimen at a given water content were repeated two to three times to account for variability in the measured values of slump. Table 2 Boundary conditions used in FEM analysis Type of analysis Flow boundary Displacement boundary In-situ Water table at top Sides - x direction fixed (x = 0) Bottom x & y direction fixed (x and y 0) Transient Water table at top (Inherited from In- situ analysis) Zero pressure at top Flow into the side walls (total head = 50 m throughout the depth) Sides and bottom fixed (x and y 0) 6.2.1 Discussion of results from finite element model It was important to wait for excess pore pressure to dissipate completely as it reflects the end of consolidation process (primary consolidation). In the present case (Figure 12), it takes 1.75 years to dissipate more than 95 % of the excess pore pressure. With the increase in fines content of the mix , the time taken for more than 95% of excess pore pressure to dissipate increases. This is because the amount of void in the soil matrix decreases with the increase in fines content of the mix. Construction of soil bentonite (SB) slurry wall is a site remediation technique to contain waste at the contaminated site. In the present study, 11 km long SB wall is being constructed around a potash mine tailing management area (TMA) to retard and prevent the flow of brine into the nearby aquifers and water channels. Since the top few meters of the wall is prone to attack by brine which can cause osmotic consolidation, it is important to understand the vertical effective stress distribution in SB walls with time and depth of wall. The variation of v with time and depth of the wall for FB is shown in Figure 13. Similar curves for other trial mixes used in this study were also plotted. The time taken for effective stress to build up at any given time and depth, increases with the increase in fines content in the mix. For example, the stress at time, T = 0.5 years at 20 m depth for TM1 and TM3 are 6 kPa and 2 kPa. This is about 65% and 95% less than the effective stress at the same depth at the end of consolidation. The stress distribution shown in Figure 14 compares the Final v at the end of consolidation i.e. when all the excess pore pressures have dissipated. In the case of above mix, it took 1.75 years (Figure 12) for the excess pore pressure to dissipate completely and reach the hydrostatic state (steady state in this case).The dashed line in Figure 14 shows the vertical stress distribution as predicted by the discrete model. It shows that after a depth of about 5 m from top, the stress becomes constant with depth. The solid line representing FEM prediction shows an increasing trend of stress from top to bottom of the wall with the increase in depth. Figure 12 Dissipation of excess pore pressure with time Figure 13 Variation of vertical effective stress (v) with time The FEM and discrete model prediction seems to match over the entire depth of the wall. The stress at a depth of 5 m from the ground surface is 65% and 62% less than the geostatic stress whereas the stress at 50 m depth (at the bottom of the trench) is 96.5% and 95.5% less than geostatic stress as predicted by FEM and discrete model respectively. Hence, it can be concluded that the vertical effective stress in SB walls never reaches geostatic stress condition.
8. 8. Figure 14 Comparison of final vertical effective stress, v with depth of the wall as predicted by the two models 7. CONCLUSIONS The results of this study show that there is a significant loss of effective vertical stress in initial slurry wall backfill with depth, and that this loss increases for narrower a trench. The following conclusions can be drawn from the laboratory test results. i. Grain size distribution has significant impact on but minimal impact on K. The value for the SB mix ranged from 2 for M (maximum amount of fines) to 2 for M1 (minimum amount of fines) ii. Hydraulic conductivity (k) of the SB mix was a function of amount of fines. At a given consolidating pressure, k decreased with increase in fines content in the backfill mix. Similar results have been reported by Baxter, 2001 and Yeo et al, 2005. iii. Hydraulic conductivity for TM3 and TM1 at consolidating pressure of 900 kPa was 2.80 x 10 -10 m/s and 5.28 x 10 -8 m/s respectively. TM2 and FB which had nearly the same fines content have similar k at any given consolidating pressure. iv. Compressibility of the mix was proportional to the amount of fines. TM3 which had the maximum amount of fines (50%) was the most compressible, while TM1 with the least amount of fines (10%) was the least compressible. Vertical stress distribution in SB walls was predicted using analytical and finite element methods. Both models show arching in the SB wall. Arching is defined as a phenomenon of stress transfer from a yielding mass to an adjacent stationary solid body. In the present study, arching is observed by stress distribution pattern in long narrow SB slurry walls. Arching in SB walls decreases with the increase in the width of the wall i.e. the stresses move more towards geostatic stress condition. This was due to decrease in frictional contribution which opposes the consolidation of the SB mixes. i. Co-efficient of effective angle of internal friction (') of the material influences the consolidation of materials in narrow trenches. The greater the amount of fines in the mix, the smaller is the arching. Hence, arching in deep narrow walls is largely proportional to '. ii. Arching in the SB walls decreases with an increase in trench width. iii. Although the final stress in the wall may be reached in a few years, it never reaches the geostatic stress condition. 8. REFERENCES ASTM D 3080 04. Standard test method for direct shear test of soils under consolidated drained conditions DAppolonia, D. (1980) Soil-bentonite slurry trench cut-offs. Journal of Geotechnical Engineering Division, 106(4), 399- 417. Evans, J.C., Costa, M.J. and Cooley, B. (1995) The state-of- stress in soil-bentonite slurry trench cut-off walls, Proc., Geoenvironment 2000, ASCE Geotechnical Special Publication No.46, Y.B. Acar and D.E. Daniel, eds., New Orleans, Louisiana,1173-1191 Filz, G.M. (1996) Consolidation stresses in soil-bentonite backfilled trenches, Proc., Environmental Geotechnics, Kamon(Ed), Balkema, Rotterdam, 497-502. Haug, M.D. (1983) Selection criteria for slurry trench cut- offs. Canadian Journal of Civil Engg, 10(3), 527-537. Marston, A. and Anderson, A.O. (1913) The theory of loads on pipes in ditches and tests of cement and clay drain tie and sewer pipe. Iowa Engineering Experimental station Bulletin (31), Ames, Iowa. McCandless, R.M. and Bodocsi, A. (1988) Hydraulic characteristics of model soil-bentonite slurry cut-off walls. In Proceedings of the 5 th National Conference on Hazardous Wastes and Hazardous Materials. Evans, J.C. 1994. Hydraulic conductivity of vertical cut-off walls. Hydraulic conductivity and waster contaminant transport in soil, ASTM STP 1142, D.E. Daniel and S.J Trautwein, eds., Chapman and Hall, London, 430-454. Gan, J, Nader, I.B and Haug, M.D. (2011) Consolidation and hydraulic conductivity testing of Slurries. TAILINGS AND MINE WAS E 11, Proceedings of the 15 th International Conference on Talings and Mine Waste, Vancouver, BC, November 6 to 9, 2011,259-265. Ryan, C.R. 1987. Soil bentonite cut-off walls, Geotechnical practice for waste disposal87, R.D Woods, ed., ASCE, New York, 182-204.
9. 9. Terzaghi, K. (1945) Stability and stiffness of cellular cofferdams. Transactions, ASCE, 110(2253), 1083-1119. Xanthakos, P.P. (1979) Slurry walls, McGraw Hill, New York, NY. Yeo, S.S., Shackelford, C.D. and Evans, J.C (2005) Consolidation and hydraulic conductivity of nine model soil- bentonite backfills. Journal of Geotechnical and Geoenvironmental Engineering, 131(10), 1189-1198.