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Numerical analysis of lateral movements and strut forces in deep cement mixing walls with top-down construction in soft clay
Computers and Geotechnics 88 (2017) 174–181
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Numerical analysis of lateral movements and strut forces in deep cement mixing walls with top-down construction in soft clay Pitthaya Jamsawang a,⇑, Sittisak Jamnam a, Pornkasem Jongpradist b, Pornpot Tanseng c, Suksun Horpibulsuk c a b c
Soil Engineering Research Center, Department of Civil Engineering, King Mongkut’s University of Technology North Bangkok, Thailand Department of Civil Engineering, Faculty of Engineering, King Mongkut’s University of Technology Thonburi, Bangkok, Thailand School of Civil Engineering and Center of Excellence in Innovation for Sustainable Infrastructure Development, Suranaree University of Technology, Thailand
a r t i c l e
i n f o
Article history: Received 18 October 2016 Received in revised form 22 March 2017 Accepted 23 March 2017
Keywords: Deep excavation Deep mixing Finite element Simulation Top-down construction in three dimensions Wall
a b s t r a c t This article presents the observed and simulated lateral movements and strut forces induced in deep cement mixing walls under deep excavation using top-down construction techniques in soft Bangkok clay. The walls are supported laterally by permanent concrete slabs and temporary struts. A threedimensional numerical model is first calibrated with observed data from a case study. Then, a parametric study is performed to compare this construction method with the bottom-up method and investigate the influence of the DCM wall thickness on lateral movements and strut forces of the wall. Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction Deep cement mixing (DCM) walls have been used for deep excavation works in soft clays to protect adjacent properties in many countries [1]. In a DCM wall, columns are formed by mixing in situ soil with cement. The DCM wall cross section is necessarily relatively thick due to its low tensile strength, and it is typically excavated without struts using the bottom-up (BU) construction technique. The top-down (TD) construction method is used for deep excavations in urban areas when there are extremely strict environmental protections, insufficient working spaces and extremely short construction times. One advantage of this method is that a basement excavation and a support from the critical path of the project can be removed after the walls and pile foundations are constructed and the first slab is cast in place. The slabs act as permanent lateral braces for the wall, which are considerably stiffer than cross-lot braces that should minimize adjacent ground movements typically encountered in BU construction [2]. Long [3], Moormann [4] and Wang et al. [5] presented databases of a large number of case histories of deep excavations through soft soils. The walls constructed using the TD method included ⇑ Corresponding author. E-mail address: [email protected] (P. Jamsawang). http://dx.doi.org/10.1016/j.compgeo.2017.03.018 0266-352X/Ó 2017 Elsevier Ltd. All rights reserved.
contiguous pile walls and diaphragm walls, whereas the sheet pile walls and DCM walls were constructed using the BU method. They concluded that TD methods generally resulted in smaller lateral wall movement. A large amount of research has been conducted using three-dimensional (3D) finite element analysis to investigate the lateral movements of diaphragm walls and contiguous pile walls with the TD and BU methods in the context of deep excavation in soft soils [1,6–13]. However, research on the lateral movements and strut forces of DCM walls with TD construction for deep excavations in soft clays has been limited. This paper focuses on the numerical analysis of a field case study of a DCM wall for a deep excavation with TD construction in soft clay in Bangkok, Thailand. In the field study, lateral movement profiles and strut forces were observed during the final stages of excavation. The 3D finite element analysis incorporated in the commercial software program PLAXIS 3D version 2013 was used for the numerical analysis. The numerical analysis simulated the lateral movement behavior and axial forces of the struts. In addition, the numerical analysis was used to investigate the influence of the thickness of the DCM walls on lateral movements and the strut forces for the DCM wall, and to compare the results of the construction method with those of the BU method.
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2. Subsoil conditions and project details The project construction site was located on Sukumvit Road in central Bangkok, Thailand. The soil profile at this site was a 2-m-thick weathered crust underlain by an 11-m-thick soft clay, a 3-m-thick medium stiff clay, a 1-m-thick stiff clay, and an underlying thick dense sand layer, as shown in Fig. 1a. The average water table was approximately 4 m below the ground surface. The basic and geotechnical engineering properties of the soils are summarized in the soil profile presented in Fig. 1b–k, including the wet unit weight, natural water content, liquid limit and plastic limit, soil specific gravity, initial void ratio, and undrained shear strength obtained from unconfined compression tests. Conventional oedometer tests and conventional triaxial tests based on consolidated undrained tests were performed on foundation soil specimens taken from the project site at depths of 1.5, 7, 14 and 18 m for the weathered crust, soft clay, medium stiff clay and stiff clay, respectively, to determine the soil parameters for the numerical simulations. The results of the triaxial tests indicate that the effective friction angle varied from 23° to 28°, whereas the effective cohesion varied from 2 to 30 kPa. These shear strength parameters are consistent with the values for the numerical modeling of Bangkok clays reported by Jamsawang et al. [1]. The ratios of swelling index to compression index obtained from the oedometer tests were 0.13–0.25 for the stiff and soft clays, respectively, which falls within the range of ratios of swelling index to compression index for the Bangkok sub soils reported by Bergado et al. [14]. The over consolidation ratio profiles determined by the oedometer tests show that the weathered crust, medium stiff clay and stiff clay were heavily over-consolidated and that the soft clay was slightly over-consolidated. The project was an 18-story condominium project with two underground car parks and a maximum excavation depth of 7.90 m. The excavation was performed in the soft clay layer only. With insufficient space to construct a thick gravity DCM wall, a DCM wall with temporary bracing systems was designed to reduce the wall thickness. The TD construction technique was used to minimize the construction time, employing a permanent basement slab for lateral support. Fig. 2a shows the layout of the DCM wall and pile foundation and the location of the temporary struts. Fig. 2b presents the cross section of the DCM wall. The maximum
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excavation depth was 7.90 m on the western side of the excavation area, which was close to a public road. On the eastern side of the excavation area, the excavation depth was 6.30 m. There were two basements, B1 and B2. The levels of the two basement slabs were different; the B1-A slab was at a level of 2.90 m, and the B1-B slab was at a level of 4.50 m. The excavation for B2 was unequal; for the B2-A side, the excavation level was 6.3 m, whereas the excavation level for the B2-B side was 7.90 m. Because the construction employed the TD method, the B1-A and B1-B basement slabs were used as the lateral supports for the DCM wall, and they were installed before the construction of the mat foundation. Therefore, the temporary stanchions, which were embedded into the bored piles prior to the excavation work, were required to support the basement slabs. The DCM wall used at this site comprised four rows of 0.7 m diameter DCM columns with 0.1 m of overlap. The entire thickness of the DCM walls was 2.5 m, and the tip was 14 m from the ground surface. The tip of the DCM was embedded 1 m into the medium stiff clay layer. The DCM walls were installed using a low-pressure mechanical mixing method. The excavation construction sequences versus elapsed time are listed in Table 1. The construction project was started on June 16, 2012, and completed on November 21, 2012. Four inclinometer casings were installed up to the stiff clay layer at a depth of 19 m. They were located at the middle of the walls on four sides of the excavation boundary to monitor the wall lateral movement, as shown in Fig. 2a. Electrical strain gauges were attached to the neutral axes of the struts to avoid a bending stress component to measure the forces in the temporary struts installed between B1-A and B1-B. A dummy strain gauge was used to eliminate any effect of temperature.
3. Laboratory tests on cement-admixed clay samples The required unconfined compressive strength (qu) of the DCM columns used for excavation work in this project was 1 MPa. Thus, a trial mix was prepared to determine the optimum cement content to be used in in situ mixing before construction initiated. Portland cement type I was used as an admixture according to the standard of the Department of Highways of Thailand. The chemical compositions and classifications of cement used in this study are
Fig. 1. Soil profiles and soil properties.
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P. Jamsawang et al. / Computers and Geotechnics 88 (2017) 174–181 Table 1 Excavation construction sequences versus elapsed time. Stage
Detail
Elapsed time (day)
1
Construct 0.7- and 0.8-m-diameter bored piles to support the structural load Install a temporary column (stanchion) usingH400 197 kg/m steel embedded into the bored piles to transfer the weight from the basement floor during construction to the bored piles Install DCM column walls by deep wet mixing around the excavated area to be used as a temporary retaining structure for the construction of the basement Excavate to 3.20 m for floor zone B1-A and install concrete slab zone B1-A Excavate to 4.80 m for floor zone B1-B and install concrete slab zone B1-A Install temporary struts between slab zones B1-A and B1-B using H 300 94 kg/m steel to transfer the lateral load due to the excavation Excavate to 6.30 m for space between floor zones B1B and B1-A Excavate to 6.30 m for floor zone B2-A Excavate to 7.90 m for floor zone B2-B
–
2
3
4 5 6
7 8 9
–
0
10 30 37
65 88
Table 2 Summary of the chemical composition of cement used in this study. Composition
Content (%)
Cao (Lime) SiO2 Al2O3 (Alumina) Fe2O3 (Iron) MgO (Magnesia) Na2O and K2O (Alkali) SO3 (Sulfuric anhydrite) Specific gravity Loss of ignition (%) Classification
62.81 21.20 4.95 2.82 4.00 0.30 2.63 3.15 1.23 Ordinary Portland cement Type1-Grade 53
ranged from 15 to 16.5 kN/m3, and the water contents varied from 35% to 70%. Unconfined compression tests were performed on samples 50 mm in diameter and 100 mm in height to determine the field qu. The values of the field qu ranged from 1.4 to 2.1 MPa, with an average value of 2.0 MPa, which was twice as high as the required qu. The undrained Young’s modulus (Eu) ranged from 120 to 290 MPa, with an average value of 200 MPa, indicating an empirical relationship Eu = 100qu, which corresponded to the test results of Jongpradist et al. [18] and Jamsawang et al. [1,19,20–22]. Fig. 2. (a) Plan view of the excavation area and (b) cross-sectional view A-A of the excavation.
4. Numerical analysis of the field case study 4.1. Finite element mesh and boundary condition
presented in Table 2. A series of unconfined compression tests was performed according to ASTM D2166-00 [15]. Flexural strength and splitting tensile tests in the laboratory followed the standard procedures of ASTM D1635-00 [16] and ASTM D3967-95a [17], respectively. The cement contained between 140 and 260 kg/m3 of wet soil; the water-cement ratio was fixed at 1.1. The results show that the required cement content was 200 kg/m3. However, the cement content used for field mixing was increased to 250 kg/m3 due to the non-uniformity of in situ mixing. The empirical relationships indicate that the flexural and the splitting tensile strengths were estimated as 0.3 and 0.16 time the qu, respectively. After the DCM wall construction was completed, core samples were extracted from the DCM columns at various depths from three locations, BHC-1, BHC-2 and BHC-3, for laboratory tests. Fig. 3 presents the test results of the core samples. The unit weights
A finite element simulation using the PLAXIS 3D version 2013 software was used to describe the performance of the DCM wall constructed using the TD method. The 3D finite element model comprised the DCM columns, bored piles and foundation soils. The soil volume was modeled using ten-node tetrahedral volume elements. The stanchions and struts (Fig. 2a and b) were modeled using beam elements, whereas the basement slabs (Fig. 2a and b) were modeled using plate elements. Fig. 4a and b illustrates the 140,000-element 3D finite element mesh used in the analysis. At the bottom of the finite element, the displacements were set to zero in the three directions, x, y, and z. The vertical model boundaries were fixed in the x- and y-directions and free in the zdirection. To avoid boundary effects, the length and width of the model were 160 and 140 m, respectively, and its depth was 30 m.
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Fig. 3. Properties of the cored DCM columns.
Fig. 4. (a) 3D finite element mesh of the wall and (b) enlargement of the modeling excavation area.
4.2. Constitutive model and model parameters The stiffness of the DCM column was much higher than that of the foundation soils, and the final stage of the excavation was performed under working stress. Therefore, a bilinear elastic-perfectly plastic Mohr–Coulomb model was used to model the DCM wall [1,20–27]. The tensile strength of the DCM columns was 0.16–0.23qu based on the measured correlation, which was considered using the tension cutoff in the model. The undrained type C was employed to simulate the undrained behavior of the DCM columns using a total stress analysis. The parameters of the Mohr–Coulomb model are listed in Table 3. Since the elapsed time for each stage of construction was relatively short (Table 1), the creep effect was much smaller than the stress relaxation from the excavation. The lateral movement induced due to excavation was much larger than the movement arising from the creep of the soft clay layer. The hardening soil model is a non-linear advanced model that is frequently used for simulating the behavior of soft soils and stiff soils under deep excavation work [1,20–22,28–30]. A value of v ur ¼ 0:2 is typically used in this model
Table 3 Parameters used in Mohr–Coulomb model. DCM column Unit weight (kN/m3) Undrained Young’s modulus (MPa) Undrained Poisson’s ratio Undrained cohesion (MPa) Friction angle (°) Material behavior
15 200 0.49 1 0 Undrained type C
[30–32]. The hardening soil model was used to model the behavior of the weathered crust, soft clay, medium stiff clay and stiff clay in this study. The soft soil creep model was not used to model the behavior of the soft clay because the elapsed time for each stage of construction was shorter than the stress relaxation time for the excavation. Thus, lateral movement induced due to excavation was much more pronounced than that due to the creep effect of the soft clay layer. The linear elastic model was employed to model the behaviors of the concrete slabs, stanchions, struts and bored piles
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[3]. The parameters of the hardening soil model and linear elastic model are listed in Tables 4 and 5, respectively. 4.3. Soil parameter calibration for the foundation soils The laboratory test results and hardening soil model were calibrated using the PLAXIS ‘‘soil test” facility to obtain reasonable soil parameters for simulating the field behaviors of the foundation soils. The test results of the triaxial tests are presented in the form of deviator stress versus axial strain, whereas those of the oedometer tests are shown as plots of the logarithm of the vertical effective stress versus vertical strain curves, as shown in Fig. 5a and b, respectively, for soft clay. The input shear strength parameters /0 and c0 for the foundation soils were obtained from the triaxial test ref ref results, as shown in Fig. 1j and k, respectively. Eref 50 , Eoed , Eur and m are independent input parameters in the hardening soil model. These parameters were adjusted to obtain suitable values to provide the best-fit results of the stress-strain curves. The example of calibration results for soft clay shown in Fig. 5a and b reveals good agreements for the stress-strain curves. The similar fitting curves were obtained for other soil layers. Therefore, suitable parameters of the foundation soils used for the 3D finite element analysis for this study are presented in Table 4.
4.4. Numerical results and comparisons with measurements 4.4.1. Lateral movement profile The observed lateral movement profiles of the DCM wall at the middle of each side of the excavation area were obtained from the four inclinometers (shown in Fig. 2a), I1, I2, I3 and I4, at the west, east, south and north parts of the excavation, respectively. The lateral wall movement profiles at the final stage of the excavation (stage 10) are shown in Fig. 6a–d. In the figures, the computed lateral movement profiles from the 3D finite element analysis are also included for comparison purposes. The trends of the lateral movement profiles were reasonably well captured, and the computed magnitudes were generally in good agreement with the observed data for I1, I2, I3 and I4. The 20% maximum overestimation of the computed dhm at a depth of 8 m (20 mm) for I4 was 5 mm, whereas the 20% maximum underestimation of the calculated dhm at a depth of 9 m (15 mm) for I3 was 5 mm. For the I1 side, the wall movement showed a small curvature, which means that the wall was tilted like a block. The wall was permitted to deflect as a cantilever beam. The maximum lateral wall movement (dhm) located at the top of the wall (near the ground surface) was 58 mm, and the movement at the tip of the wall (at a depth of 14 m) was 15 mm, which implies that the movement pattern was a combination of slight overturning and sliding. For inclinometer I2, the amount of wall movement was less than I1 because the excavation depth was smaller. The dhm was 32 mm at the top of the wall, and the movement at the tip of the wall was 10 mm. The lateral movement profiles for the I3 and I4 sides are presented in Fig 6c and d, respectively. The magnitudes of the lateral movements were considerably smaller than those of the I1 and I2 sides because of the smaller wall length and sufficient lateral support from the concrete slab bracings B1-B and B1-A. Because the final excavation depths were the same (He), there were no significantly different lateral wall deflections on the two sides of I3 and I4. The dhm values were 0.73, 0.51, 0.23 and 0.33% He for walls located at I1, I2, I3 and I4, respectively. The lateral movement profiles developed into a bulged profile inward of the excavation area, indicating that the walls of the two sides were well propped near the surface. Thus, the location of the maximum lateral movement occurred at a deeper depth.dhm was 22 and 26 mm at distances
4.5 and 7 m below the ground surface for the walls located at I3 and I4, respectively. The tip movements of the retaining wall were only 5 mm for I3 and I4. The location of the maximum lateral movement was at the ground surface for I1 and I2, whereas it was 0.63 and 1.0 He below the ground surface for I3 and I4, respectively. Ou et al. [33] found that the location of the maximum lateral movement of eight case histories in Taipei soft soil were often observed near the excavation surface. The analysis of Moormann [4] showed that the dhm for most deep excavations in a soft soil was observed at depths of 0.5–1.5He under the ground surface. Wang et al. [5] reported that the location of the maximum lateral movement was observed at depths of 0.5–1.0 He under the ground surface in 53% of the case histories. For 43% of the case histories, the location of the maximum lateral movement was observed at depths of 1.0–1.4He under the ground surface. The location of the maximum lateral movement was observed at the top of the wall for only approximately 4% of the case histories. The dhm for the DCM walls without internal struts occurred at the top of the walls. Deep-seated wall displacements were observed when internal struts were used to support the DCM walls. The results of Ou et al. [33], Moormann [4] and Wang et al. [5] have been broadly confirmed by this study. The tip movement of the retaining wall was found to occur at I1 and I2. Wang et al. [34] reported that the embedded depth ratio of the wall, which is defined as the embedded length of the wall to final excavation depth, may contribute to the tip movement. Here, the embedded depth ratio was 0.77 for the DCM wall mainly varied between 0.8 and 1.4 and was 1.08 on average [34]. Thus, embedded depth ratio was the smallest in this study. A larger embedded depth ratio could suppress toe movement because more soils under the excavation surface are strengthened. 4.4.2. Strut force Fig. 7 shows comparisons of the measured and computed strut forces induced by the 7.9 m deep excavation for all struts, as shown in Fig. 2. The measured values of the strut forces were 310, 150, 330, 370, 440, 350 and 450 kN, whereas the computed values were 280, 100, 350, 360, 390, 320 and 510 kN for struts S1–S7, respectively. Strut S7 was farthest from the edge of the slab and was exposed to a larger axial force than the other struts. The yield strength of the steel used in the struts was 250 MPa, resulting in a yield axial force of 3000 kN. Thus, a minimum factor of safety against structural failure of 5.8 was obtained for this project. A comparison of the observed and computed data indicated a maximum underestimation of 33% and an overestimation of 13% for struts S2 and S7, respectively. However, the average error in the comparison was only 12%. Therefore, the computed magnitudes of the strut forces were generally in good agreement with the observed data. The calculated force in strut S7 was approximately twice the measured and calculated forces in strut S1 because the spacing of S7 was twice the spacing of S1. This shows that the strut forces determined by computational aid provide reliable results. 4.5. Parametric study on lateral movements and strut forces To investigate the effectiveness of the DCM wall with TD construction (DCM-TD) used for deep excavation work, the performances of DCM walls without a permanent concrete slab bracing or built with a BU construction method (DCM-BU) were characterized. The effect of the construction method was simulated by omitting the concrete slabs and the temporary struts from the excavation area. As shown in Fig. 8, the DCM wall simulation with BU construction resulted in considerably larger lateral movements. The dhm/He ratios were 1.28, 1.42, 0.83 and 0.85% He for DCM-BU
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P. Jamsawang et al. / Computers and Geotechnics 88 (2017) 174–181 Table 4 Parameters used in hardening soil model.
3
Unit weight (kN/m ) Secant stiffness, Eref 50 (kPa) (kPa) Tangential stiffness, Eref oed Eref ur
(kPa) Unloading and reloading stiffness, Power of the stress-level dependency of the stiffness, m Poisson’s ratio for unloading-reloading, v ur Effective cohesion (kPa) Effective friction angle (°) Angle of dilatancy (°) Over consolidation ratio Material behavior
Weathered crust
Soft clay
Medium stiff clay
Stiff clay
16 10,000
14 1200
16 10,000
20 20,000
12,000
960
12,000
25,000
35,000
4000
45,000
95,000
1.0 0.2 15 27 0 2.0 Undrained
1.0 0.2 2 23 0 1.5 Undrained
0.9 0.2 9 26 0 2.0 Undrained
0.9 0.2 30 28 0 2.4 Undrained
Table 5 Parameters used in linear elastic model.
Unit weight (kN/m3) Moment of inertia (m4) Cross-sectional area (m2) Young’s modulus (kPa) Poisson’s ratio Material behavior
Concrete slab
Bored pile
Temporary strut
Stanchion
24 1.3 103 m4/m 0.25 m2/m 2.8 107 0.15 Plate element
24 0.012–0.020 0.38–0.50 2.0 107 0.15 Non-porous
78 2.0 104 0.012 2.1 108 – Beam element
78 7.1 104 0.025 2.1 108 – Beam element
Fig. 5. An example of soil parameter calibrations with (a) consolidated undrained triaxial and (b) oedometer test results for soft clay layer.
and 0.81, 0.63, 0.28 and 0.30% He for DCM-TD for walls located at I1, I2, I3 and I4, respectively. These results indicate that the DCM-TD can reduce lateral movement with improvement ratios of 1.6, 2.4, 3.0 and 2.8 for walls located at I1, I2, I3 and I4, respec-
tively, yielding an average improvement ratio of 2.4. The high improvement ratio reflects the strong influence of concrete slab bracing. The shapes of the computed lateral movement profiles for the walls located at I3 and I4 in Fig. 8 tended to move farther toward the excavation, unlike those for the case study. This finding confirms that the presence of a slab affected the type of wall lateral movement profiles for walls located I3 and I4, as noted in Section 4.4.1. The DCM wall thickness for this project was limited to 2.5 m, or four rows of a DCM wall, due to insufficient space. Thus, the effect of DCM wall thickness on lateral movement and strut forces was investigated by varying the number of rows, using 2, 4 (case study), 6, 8 and 10, which correspond to wall thicknesses of 1.3, 2.5, 3.7, 4.9 and 6.1 m, respectively. Fig. 9a presents the relationship between wall thickness and mean values of dhm/He for walls located at I1, I2, I3 and I4. The numerical results show that the mean dhm/He decreases with increasing wall thickness due to increasing wall rigidity [1] for both DCM-TD and DCM-BU. In addition, the DCM-TD for the field case study requires half the wall thickness of the DCM-BU to obtain same mean value of dhm/He. In addition, the difference in the mean value of dhm/He between DCM-TD and DCM-BU decreases with increasing wall thickness, which reflects a decrease in the improvement ratio. The improvement ratios were 2.4, 2.2, 2.0, 1.8 and 1.5 for wall thicknesses of 1.3, 2.5, 3.7, 4.9 and 6.1 m, respectively. Based on past experiences in engineering practice, the improvement ratio should be greater than 2. Thus, the use of a wall thickness greater than 3.7 m or 6 rows in this study was economically undesirable for DCM-TD. Fig. 9b shows the effect of wall thickness on magnitude of all strut forces S1–S7. The figure shows that the wall thickness had a substantial effect on the strut forces. As expected, the results show that the overall tendency is the reduction in strut forces for all locations with increasing wall rigidity, resulting in the possibility of reducing the number of struts. However, this reduction is only significant when the wall thickness varied from 1.3 to 2.5 m. The influence of the wall thickness was not easily detectable when wall thickness exceeded 2.5 m, indicating that the influence of wall thickness is only significant when the wall has a low rigidity.
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Fig. 6. Comparison of measured and computed lateral movements.
1200 1100
Slab B1-B
Strut force (kN)
1000 900
8.0
2.5
S1
800 700 600
2.5
S2
3.1
S3
5.1
S4
3.3
S5
16.0
4.8
S6
S7
Measured Simulated
500 400 300 200 100 0
S1
S2
S3
S4
S5
S6
S7
Fig. 7. Comparison of measured and calculated strut loads.
5. Summary and conclusions A field case study of a DCM wall applied using a TD construction method for a deep excavation in soft Bangkok clay is reported. The lateral movements of the wall system and the strut forces at the proposed excavation depth were observed. A calibration of laboratory test results with the HSM was performed to obtain the optimal parameters for simulating the behavior of the foundation soils. The software program PLAXIS 3D version 2013 was used to simulate the behavior of the wall. Finally, the results of the construction method in comparison to the BU method were investigated, and the influence of wall thickness on wall lateral movement and strut forces was determined to evaluate the effectiveness of TD construction method and to identify an appropriate wall thickness. The following conclusions can be drawn based on the observed and simulated results: 1. The observed data show that dhm were 0.73 and 0.51, 0.23 and 0.33% He for the long and short sides of the walls, respectively. The long sides of the walls were permitted to deflect as a cantilever beam. However, the short sides of the walls were well supported near the surface due to sufficient lateral support of the concrete slab.
Fig. 8. Effect of BU construction on computed lateral movement profiles.
2. The observed strut forces varied from 150 to 450 kN, with an average of 340 kN. The minimum factor of safety against structural failure of 5.8 was obtained for this project, and it has been confirmed by a performance-based design. 3. The numerical results obtained from the 3D model were consistent with the field data for the wall lateral movements and strut forces. Comparisons of the observed and computed data revealed 20% maximum and 12% average errors for the computed dhm and strut forces, respectively. 4. Based on the numerical results, the concrete slab used in the TD construction had a significant effect on the lateral movement of the wall. When the concrete slab was introduced, the lateral
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Fig. 9. Effect of wall thickness on (a) mean value of dh/He and (b) strut force.
movements were sharply reduced, with an average improvement ratio of 2.4. The DCM-TD in the field case study required half the wall thickness of the DCM-BU to obtain the same mean value of dhm/He. A wall thickness greater than 3.7 m was considered an economically impractical design for DCM-TD based on the improvement ratio of 2.0. 5. From the numerical analysis, all strut forces decreased with increasing wall thickness. Thus, the number of struts can be reduced for the thick wall. The effect of the DCM wall thickness appears to be insignificant when the wall thickness is greater than 2.5 m.
Acknowledgments The authors gratefully acknowledge the financial support of the Faculty of Engineering, King Mongkut’s University of Technology North Bangkok Contract no. ENG-59-01 and the Thailand Research Fund under the TRF Senior Research Scholar program Grant No. RTA5980005. References [1] Jamsawang P, Voottipruex P, Jongpradist P, Bergado DT. Parameters affecting the lateral movements of compound deep cement mixing walls by numerical simulations and parametric analyses. Acta Geotech 2015;10:797–812. [2] Arboleda-Monsalve LG, Finno RJ. Influence of concrete time-dependent effects on the performance of top-down construction. J Geotech Geoenviron Eng 2015;141:985–94. [3] Long M. Database for retaining wall and ground movements due to deep excavations. J Geotech Geoenviron Eng 2001;127:203–24. [4] Moormann C. Analysis of wall and ground movements due to deep excavations in soft soil based on a new worldwide database. Soils Found 2004;44:87–98.
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Computers and Geotechnics journal homepage: www.elsevier.com/locate/compgeo
Technical Communication
Numerical analysis of lateral movements and strut forces in deep cement mixing walls with top-down construction in soft clay Pitthaya Jamsawang a,⇑, Sittisak Jamnam a, Pornkasem Jongpradist b, Pornpot Tanseng c, Suksun Horpibulsuk c a b c
Soil Engineering Research Center, Department of Civil Engineering, King Mongkut’s University of Technology North Bangkok, Thailand Department of Civil Engineering, Faculty of Engineering, King Mongkut’s University of Technology Thonburi, Bangkok, Thailand School of Civil Engineering and Center of Excellence in Innovation for Sustainable Infrastructure Development, Suranaree University of Technology, Thailand
a r t i c l e
i n f o
Article history: Received 18 October 2016 Received in revised form 22 March 2017 Accepted 23 March 2017
Keywords: Deep excavation Deep mixing Finite element Simulation Top-down construction in three dimensions Wall
a b s t r a c t This article presents the observed and simulated lateral movements and strut forces induced in deep cement mixing walls under deep excavation using top-down construction techniques in soft Bangkok clay. The walls are supported laterally by permanent concrete slabs and temporary struts. A threedimensional numerical model is first calibrated with observed data from a case study. Then, a parametric study is performed to compare this construction method with the bottom-up method and investigate the influence of the DCM wall thickness on lateral movements and strut forces of the wall. Ó 2017 Elsevier Ltd. All rights reserved.
1. Introduction Deep cement mixing (DCM) walls have been used for deep excavation works in soft clays to protect adjacent properties in many countries [1]. In a DCM wall, columns are formed by mixing in situ soil with cement. The DCM wall cross section is necessarily relatively thick due to its low tensile strength, and it is typically excavated without struts using the bottom-up (BU) construction technique. The top-down (TD) construction method is used for deep excavations in urban areas when there are extremely strict environmental protections, insufficient working spaces and extremely short construction times. One advantage of this method is that a basement excavation and a support from the critical path of the project can be removed after the walls and pile foundations are constructed and the first slab is cast in place. The slabs act as permanent lateral braces for the wall, which are considerably stiffer than cross-lot braces that should minimize adjacent ground movements typically encountered in BU construction [2]. Long [3], Moormann [4] and Wang et al. [5] presented databases of a large number of case histories of deep excavations through soft soils. The walls constructed using the TD method included ⇑ Corresponding author. E-mail address: [email protected] (P. Jamsawang). http://dx.doi.org/10.1016/j.compgeo.2017.03.018 0266-352X/Ó 2017 Elsevier Ltd. All rights reserved.
contiguous pile walls and diaphragm walls, whereas the sheet pile walls and DCM walls were constructed using the BU method. They concluded that TD methods generally resulted in smaller lateral wall movement. A large amount of research has been conducted using three-dimensional (3D) finite element analysis to investigate the lateral movements of diaphragm walls and contiguous pile walls with the TD and BU methods in the context of deep excavation in soft soils [1,6–13]. However, research on the lateral movements and strut forces of DCM walls with TD construction for deep excavations in soft clays has been limited. This paper focuses on the numerical analysis of a field case study of a DCM wall for a deep excavation with TD construction in soft clay in Bangkok, Thailand. In the field study, lateral movement profiles and strut forces were observed during the final stages of excavation. The 3D finite element analysis incorporated in the commercial software program PLAXIS 3D version 2013 was used for the numerical analysis. The numerical analysis simulated the lateral movement behavior and axial forces of the struts. In addition, the numerical analysis was used to investigate the influence of the thickness of the DCM walls on lateral movements and the strut forces for the DCM wall, and to compare the results of the construction method with those of the BU method.
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2. Subsoil conditions and project details The project construction site was located on Sukumvit Road in central Bangkok, Thailand. The soil profile at this site was a 2-m-thick weathered crust underlain by an 11-m-thick soft clay, a 3-m-thick medium stiff clay, a 1-m-thick stiff clay, and an underlying thick dense sand layer, as shown in Fig. 1a. The average water table was approximately 4 m below the ground surface. The basic and geotechnical engineering properties of the soils are summarized in the soil profile presented in Fig. 1b–k, including the wet unit weight, natural water content, liquid limit and plastic limit, soil specific gravity, initial void ratio, and undrained shear strength obtained from unconfined compression tests. Conventional oedometer tests and conventional triaxial tests based on consolidated undrained tests were performed on foundation soil specimens taken from the project site at depths of 1.5, 7, 14 and 18 m for the weathered crust, soft clay, medium stiff clay and stiff clay, respectively, to determine the soil parameters for the numerical simulations. The results of the triaxial tests indicate that the effective friction angle varied from 23° to 28°, whereas the effective cohesion varied from 2 to 30 kPa. These shear strength parameters are consistent with the values for the numerical modeling of Bangkok clays reported by Jamsawang et al. [1]. The ratios of swelling index to compression index obtained from the oedometer tests were 0.13–0.25 for the stiff and soft clays, respectively, which falls within the range of ratios of swelling index to compression index for the Bangkok sub soils reported by Bergado et al. [14]. The over consolidation ratio profiles determined by the oedometer tests show that the weathered crust, medium stiff clay and stiff clay were heavily over-consolidated and that the soft clay was slightly over-consolidated. The project was an 18-story condominium project with two underground car parks and a maximum excavation depth of 7.90 m. The excavation was performed in the soft clay layer only. With insufficient space to construct a thick gravity DCM wall, a DCM wall with temporary bracing systems was designed to reduce the wall thickness. The TD construction technique was used to minimize the construction time, employing a permanent basement slab for lateral support. Fig. 2a shows the layout of the DCM wall and pile foundation and the location of the temporary struts. Fig. 2b presents the cross section of the DCM wall. The maximum
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excavation depth was 7.90 m on the western side of the excavation area, which was close to a public road. On the eastern side of the excavation area, the excavation depth was 6.30 m. There were two basements, B1 and B2. The levels of the two basement slabs were different; the B1-A slab was at a level of 2.90 m, and the B1-B slab was at a level of 4.50 m. The excavation for B2 was unequal; for the B2-A side, the excavation level was 6.3 m, whereas the excavation level for the B2-B side was 7.90 m. Because the construction employed the TD method, the B1-A and B1-B basement slabs were used as the lateral supports for the DCM wall, and they were installed before the construction of the mat foundation. Therefore, the temporary stanchions, which were embedded into the bored piles prior to the excavation work, were required to support the basement slabs. The DCM wall used at this site comprised four rows of 0.7 m diameter DCM columns with 0.1 m of overlap. The entire thickness of the DCM walls was 2.5 m, and the tip was 14 m from the ground surface. The tip of the DCM was embedded 1 m into the medium stiff clay layer. The DCM walls were installed using a low-pressure mechanical mixing method. The excavation construction sequences versus elapsed time are listed in Table 1. The construction project was started on June 16, 2012, and completed on November 21, 2012. Four inclinometer casings were installed up to the stiff clay layer at a depth of 19 m. They were located at the middle of the walls on four sides of the excavation boundary to monitor the wall lateral movement, as shown in Fig. 2a. Electrical strain gauges were attached to the neutral axes of the struts to avoid a bending stress component to measure the forces in the temporary struts installed between B1-A and B1-B. A dummy strain gauge was used to eliminate any effect of temperature.
3. Laboratory tests on cement-admixed clay samples The required unconfined compressive strength (qu) of the DCM columns used for excavation work in this project was 1 MPa. Thus, a trial mix was prepared to determine the optimum cement content to be used in in situ mixing before construction initiated. Portland cement type I was used as an admixture according to the standard of the Department of Highways of Thailand. The chemical compositions and classifications of cement used in this study are
Fig. 1. Soil profiles and soil properties.
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P. Jamsawang et al. / Computers and Geotechnics 88 (2017) 174–181 Table 1 Excavation construction sequences versus elapsed time. Stage
Detail
Elapsed time (day)
1
Construct 0.7- and 0.8-m-diameter bored piles to support the structural load Install a temporary column (stanchion) usingH400 197 kg/m steel embedded into the bored piles to transfer the weight from the basement floor during construction to the bored piles Install DCM column walls by deep wet mixing around the excavated area to be used as a temporary retaining structure for the construction of the basement Excavate to 3.20 m for floor zone B1-A and install concrete slab zone B1-A Excavate to 4.80 m for floor zone B1-B and install concrete slab zone B1-A Install temporary struts between slab zones B1-A and B1-B using H 300 94 kg/m steel to transfer the lateral load due to the excavation Excavate to 6.30 m for space between floor zones B1B and B1-A Excavate to 6.30 m for floor zone B2-A Excavate to 7.90 m for floor zone B2-B
–
2
3
4 5 6
7 8 9
–
0
10 30 37
65 88
Table 2 Summary of the chemical composition of cement used in this study. Composition
Content (%)
Cao (Lime) SiO2 Al2O3 (Alumina) Fe2O3 (Iron) MgO (Magnesia) Na2O and K2O (Alkali) SO3 (Sulfuric anhydrite) Specific gravity Loss of ignition (%) Classification
62.81 21.20 4.95 2.82 4.00 0.30 2.63 3.15 1.23 Ordinary Portland cement Type1-Grade 53
ranged from 15 to 16.5 kN/m3, and the water contents varied from 35% to 70%. Unconfined compression tests were performed on samples 50 mm in diameter and 100 mm in height to determine the field qu. The values of the field qu ranged from 1.4 to 2.1 MPa, with an average value of 2.0 MPa, which was twice as high as the required qu. The undrained Young’s modulus (Eu) ranged from 120 to 290 MPa, with an average value of 200 MPa, indicating an empirical relationship Eu = 100qu, which corresponded to the test results of Jongpradist et al. [18] and Jamsawang et al. [1,19,20–22]. Fig. 2. (a) Plan view of the excavation area and (b) cross-sectional view A-A of the excavation.
4. Numerical analysis of the field case study 4.1. Finite element mesh and boundary condition
presented in Table 2. A series of unconfined compression tests was performed according to ASTM D2166-00 [15]. Flexural strength and splitting tensile tests in the laboratory followed the standard procedures of ASTM D1635-00 [16] and ASTM D3967-95a [17], respectively. The cement contained between 140 and 260 kg/m3 of wet soil; the water-cement ratio was fixed at 1.1. The results show that the required cement content was 200 kg/m3. However, the cement content used for field mixing was increased to 250 kg/m3 due to the non-uniformity of in situ mixing. The empirical relationships indicate that the flexural and the splitting tensile strengths were estimated as 0.3 and 0.16 time the qu, respectively. After the DCM wall construction was completed, core samples were extracted from the DCM columns at various depths from three locations, BHC-1, BHC-2 and BHC-3, for laboratory tests. Fig. 3 presents the test results of the core samples. The unit weights
A finite element simulation using the PLAXIS 3D version 2013 software was used to describe the performance of the DCM wall constructed using the TD method. The 3D finite element model comprised the DCM columns, bored piles and foundation soils. The soil volume was modeled using ten-node tetrahedral volume elements. The stanchions and struts (Fig. 2a and b) were modeled using beam elements, whereas the basement slabs (Fig. 2a and b) were modeled using plate elements. Fig. 4a and b illustrates the 140,000-element 3D finite element mesh used in the analysis. At the bottom of the finite element, the displacements were set to zero in the three directions, x, y, and z. The vertical model boundaries were fixed in the x- and y-directions and free in the zdirection. To avoid boundary effects, the length and width of the model were 160 and 140 m, respectively, and its depth was 30 m.
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Fig. 3. Properties of the cored DCM columns.
Fig. 4. (a) 3D finite element mesh of the wall and (b) enlargement of the modeling excavation area.
4.2. Constitutive model and model parameters The stiffness of the DCM column was much higher than that of the foundation soils, and the final stage of the excavation was performed under working stress. Therefore, a bilinear elastic-perfectly plastic Mohr–Coulomb model was used to model the DCM wall [1,20–27]. The tensile strength of the DCM columns was 0.16–0.23qu based on the measured correlation, which was considered using the tension cutoff in the model. The undrained type C was employed to simulate the undrained behavior of the DCM columns using a total stress analysis. The parameters of the Mohr–Coulomb model are listed in Table 3. Since the elapsed time for each stage of construction was relatively short (Table 1), the creep effect was much smaller than the stress relaxation from the excavation. The lateral movement induced due to excavation was much larger than the movement arising from the creep of the soft clay layer. The hardening soil model is a non-linear advanced model that is frequently used for simulating the behavior of soft soils and stiff soils under deep excavation work [1,20–22,28–30]. A value of v ur ¼ 0:2 is typically used in this model
Table 3 Parameters used in Mohr–Coulomb model. DCM column Unit weight (kN/m3) Undrained Young’s modulus (MPa) Undrained Poisson’s ratio Undrained cohesion (MPa) Friction angle (°) Material behavior
15 200 0.49 1 0 Undrained type C
[30–32]. The hardening soil model was used to model the behavior of the weathered crust, soft clay, medium stiff clay and stiff clay in this study. The soft soil creep model was not used to model the behavior of the soft clay because the elapsed time for each stage of construction was shorter than the stress relaxation time for the excavation. Thus, lateral movement induced due to excavation was much more pronounced than that due to the creep effect of the soft clay layer. The linear elastic model was employed to model the behaviors of the concrete slabs, stanchions, struts and bored piles
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[3]. The parameters of the hardening soil model and linear elastic model are listed in Tables 4 and 5, respectively. 4.3. Soil parameter calibration for the foundation soils The laboratory test results and hardening soil model were calibrated using the PLAXIS ‘‘soil test” facility to obtain reasonable soil parameters for simulating the field behaviors of the foundation soils. The test results of the triaxial tests are presented in the form of deviator stress versus axial strain, whereas those of the oedometer tests are shown as plots of the logarithm of the vertical effective stress versus vertical strain curves, as shown in Fig. 5a and b, respectively, for soft clay. The input shear strength parameters /0 and c0 for the foundation soils were obtained from the triaxial test ref ref results, as shown in Fig. 1j and k, respectively. Eref 50 , Eoed , Eur and m are independent input parameters in the hardening soil model. These parameters were adjusted to obtain suitable values to provide the best-fit results of the stress-strain curves. The example of calibration results for soft clay shown in Fig. 5a and b reveals good agreements for the stress-strain curves. The similar fitting curves were obtained for other soil layers. Therefore, suitable parameters of the foundation soils used for the 3D finite element analysis for this study are presented in Table 4.
4.4. Numerical results and comparisons with measurements 4.4.1. Lateral movement profile The observed lateral movement profiles of the DCM wall at the middle of each side of the excavation area were obtained from the four inclinometers (shown in Fig. 2a), I1, I2, I3 and I4, at the west, east, south and north parts of the excavation, respectively. The lateral wall movement profiles at the final stage of the excavation (stage 10) are shown in Fig. 6a–d. In the figures, the computed lateral movement profiles from the 3D finite element analysis are also included for comparison purposes. The trends of the lateral movement profiles were reasonably well captured, and the computed magnitudes were generally in good agreement with the observed data for I1, I2, I3 and I4. The 20% maximum overestimation of the computed dhm at a depth of 8 m (20 mm) for I4 was 5 mm, whereas the 20% maximum underestimation of the calculated dhm at a depth of 9 m (15 mm) for I3 was 5 mm. For the I1 side, the wall movement showed a small curvature, which means that the wall was tilted like a block. The wall was permitted to deflect as a cantilever beam. The maximum lateral wall movement (dhm) located at the top of the wall (near the ground surface) was 58 mm, and the movement at the tip of the wall (at a depth of 14 m) was 15 mm, which implies that the movement pattern was a combination of slight overturning and sliding. For inclinometer I2, the amount of wall movement was less than I1 because the excavation depth was smaller. The dhm was 32 mm at the top of the wall, and the movement at the tip of the wall was 10 mm. The lateral movement profiles for the I3 and I4 sides are presented in Fig 6c and d, respectively. The magnitudes of the lateral movements were considerably smaller than those of the I1 and I2 sides because of the smaller wall length and sufficient lateral support from the concrete slab bracings B1-B and B1-A. Because the final excavation depths were the same (He), there were no significantly different lateral wall deflections on the two sides of I3 and I4. The dhm values were 0.73, 0.51, 0.23 and 0.33% He for walls located at I1, I2, I3 and I4, respectively. The lateral movement profiles developed into a bulged profile inward of the excavation area, indicating that the walls of the two sides were well propped near the surface. Thus, the location of the maximum lateral movement occurred at a deeper depth.dhm was 22 and 26 mm at distances
4.5 and 7 m below the ground surface for the walls located at I3 and I4, respectively. The tip movements of the retaining wall were only 5 mm for I3 and I4. The location of the maximum lateral movement was at the ground surface for I1 and I2, whereas it was 0.63 and 1.0 He below the ground surface for I3 and I4, respectively. Ou et al. [33] found that the location of the maximum lateral movement of eight case histories in Taipei soft soil were often observed near the excavation surface. The analysis of Moormann [4] showed that the dhm for most deep excavations in a soft soil was observed at depths of 0.5–1.5He under the ground surface. Wang et al. [5] reported that the location of the maximum lateral movement was observed at depths of 0.5–1.0 He under the ground surface in 53% of the case histories. For 43% of the case histories, the location of the maximum lateral movement was observed at depths of 1.0–1.4He under the ground surface. The location of the maximum lateral movement was observed at the top of the wall for only approximately 4% of the case histories. The dhm for the DCM walls without internal struts occurred at the top of the walls. Deep-seated wall displacements were observed when internal struts were used to support the DCM walls. The results of Ou et al. [33], Moormann [4] and Wang et al. [5] have been broadly confirmed by this study. The tip movement of the retaining wall was found to occur at I1 and I2. Wang et al. [34] reported that the embedded depth ratio of the wall, which is defined as the embedded length of the wall to final excavation depth, may contribute to the tip movement. Here, the embedded depth ratio was 0.77 for the DCM wall mainly varied between 0.8 and 1.4 and was 1.08 on average [34]. Thus, embedded depth ratio was the smallest in this study. A larger embedded depth ratio could suppress toe movement because more soils under the excavation surface are strengthened. 4.4.2. Strut force Fig. 7 shows comparisons of the measured and computed strut forces induced by the 7.9 m deep excavation for all struts, as shown in Fig. 2. The measured values of the strut forces were 310, 150, 330, 370, 440, 350 and 450 kN, whereas the computed values were 280, 100, 350, 360, 390, 320 and 510 kN for struts S1–S7, respectively. Strut S7 was farthest from the edge of the slab and was exposed to a larger axial force than the other struts. The yield strength of the steel used in the struts was 250 MPa, resulting in a yield axial force of 3000 kN. Thus, a minimum factor of safety against structural failure of 5.8 was obtained for this project. A comparison of the observed and computed data indicated a maximum underestimation of 33% and an overestimation of 13% for struts S2 and S7, respectively. However, the average error in the comparison was only 12%. Therefore, the computed magnitudes of the strut forces were generally in good agreement with the observed data. The calculated force in strut S7 was approximately twice the measured and calculated forces in strut S1 because the spacing of S7 was twice the spacing of S1. This shows that the strut forces determined by computational aid provide reliable results. 4.5. Parametric study on lateral movements and strut forces To investigate the effectiveness of the DCM wall with TD construction (DCM-TD) used for deep excavation work, the performances of DCM walls without a permanent concrete slab bracing or built with a BU construction method (DCM-BU) were characterized. The effect of the construction method was simulated by omitting the concrete slabs and the temporary struts from the excavation area. As shown in Fig. 8, the DCM wall simulation with BU construction resulted in considerably larger lateral movements. The dhm/He ratios were 1.28, 1.42, 0.83 and 0.85% He for DCM-BU
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P. Jamsawang et al. / Computers and Geotechnics 88 (2017) 174–181 Table 4 Parameters used in hardening soil model.
3
Unit weight (kN/m ) Secant stiffness, Eref 50 (kPa) (kPa) Tangential stiffness, Eref oed Eref ur
(kPa) Unloading and reloading stiffness, Power of the stress-level dependency of the stiffness, m Poisson’s ratio for unloading-reloading, v ur Effective cohesion (kPa) Effective friction angle (°) Angle of dilatancy (°) Over consolidation ratio Material behavior
Weathered crust
Soft clay
Medium stiff clay
Stiff clay
16 10,000
14 1200
16 10,000
20 20,000
12,000
960
12,000
25,000
35,000
4000
45,000
95,000
1.0 0.2 15 27 0 2.0 Undrained
1.0 0.2 2 23 0 1.5 Undrained
0.9 0.2 9 26 0 2.0 Undrained
0.9 0.2 30 28 0 2.4 Undrained
Table 5 Parameters used in linear elastic model.
Unit weight (kN/m3) Moment of inertia (m4) Cross-sectional area (m2) Young’s modulus (kPa) Poisson’s ratio Material behavior
Concrete slab
Bored pile
Temporary strut
Stanchion
24 1.3 103 m4/m 0.25 m2/m 2.8 107 0.15 Plate element
24 0.012–0.020 0.38–0.50 2.0 107 0.15 Non-porous
78 2.0 104 0.012 2.1 108 – Beam element
78 7.1 104 0.025 2.1 108 – Beam element
Fig. 5. An example of soil parameter calibrations with (a) consolidated undrained triaxial and (b) oedometer test results for soft clay layer.
and 0.81, 0.63, 0.28 and 0.30% He for DCM-TD for walls located at I1, I2, I3 and I4, respectively. These results indicate that the DCM-TD can reduce lateral movement with improvement ratios of 1.6, 2.4, 3.0 and 2.8 for walls located at I1, I2, I3 and I4, respec-
tively, yielding an average improvement ratio of 2.4. The high improvement ratio reflects the strong influence of concrete slab bracing. The shapes of the computed lateral movement profiles for the walls located at I3 and I4 in Fig. 8 tended to move farther toward the excavation, unlike those for the case study. This finding confirms that the presence of a slab affected the type of wall lateral movement profiles for walls located I3 and I4, as noted in Section 4.4.1. The DCM wall thickness for this project was limited to 2.5 m, or four rows of a DCM wall, due to insufficient space. Thus, the effect of DCM wall thickness on lateral movement and strut forces was investigated by varying the number of rows, using 2, 4 (case study), 6, 8 and 10, which correspond to wall thicknesses of 1.3, 2.5, 3.7, 4.9 and 6.1 m, respectively. Fig. 9a presents the relationship between wall thickness and mean values of dhm/He for walls located at I1, I2, I3 and I4. The numerical results show that the mean dhm/He decreases with increasing wall thickness due to increasing wall rigidity [1] for both DCM-TD and DCM-BU. In addition, the DCM-TD for the field case study requires half the wall thickness of the DCM-BU to obtain same mean value of dhm/He. In addition, the difference in the mean value of dhm/He between DCM-TD and DCM-BU decreases with increasing wall thickness, which reflects a decrease in the improvement ratio. The improvement ratios were 2.4, 2.2, 2.0, 1.8 and 1.5 for wall thicknesses of 1.3, 2.5, 3.7, 4.9 and 6.1 m, respectively. Based on past experiences in engineering practice, the improvement ratio should be greater than 2. Thus, the use of a wall thickness greater than 3.7 m or 6 rows in this study was economically undesirable for DCM-TD. Fig. 9b shows the effect of wall thickness on magnitude of all strut forces S1–S7. The figure shows that the wall thickness had a substantial effect on the strut forces. As expected, the results show that the overall tendency is the reduction in strut forces for all locations with increasing wall rigidity, resulting in the possibility of reducing the number of struts. However, this reduction is only significant when the wall thickness varied from 1.3 to 2.5 m. The influence of the wall thickness was not easily detectable when wall thickness exceeded 2.5 m, indicating that the influence of wall thickness is only significant when the wall has a low rigidity.
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Fig. 6. Comparison of measured and computed lateral movements.
1200 1100
Slab B1-B
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S1
800 700 600
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S2
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500 400 300 200 100 0
S1
S2
S3
S4
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Fig. 7. Comparison of measured and calculated strut loads.
5. Summary and conclusions A field case study of a DCM wall applied using a TD construction method for a deep excavation in soft Bangkok clay is reported. The lateral movements of the wall system and the strut forces at the proposed excavation depth were observed. A calibration of laboratory test results with the HSM was performed to obtain the optimal parameters for simulating the behavior of the foundation soils. The software program PLAXIS 3D version 2013 was used to simulate the behavior of the wall. Finally, the results of the construction method in comparison to the BU method were investigated, and the influence of wall thickness on wall lateral movement and strut forces was determined to evaluate the effectiveness of TD construction method and to identify an appropriate wall thickness. The following conclusions can be drawn based on the observed and simulated results: 1. The observed data show that dhm were 0.73 and 0.51, 0.23 and 0.33% He for the long and short sides of the walls, respectively. The long sides of the walls were permitted to deflect as a cantilever beam. However, the short sides of the walls were well supported near the surface due to sufficient lateral support of the concrete slab.
Fig. 8. Effect of BU construction on computed lateral movement profiles.
2. The observed strut forces varied from 150 to 450 kN, with an average of 340 kN. The minimum factor of safety against structural failure of 5.8 was obtained for this project, and it has been confirmed by a performance-based design. 3. The numerical results obtained from the 3D model were consistent with the field data for the wall lateral movements and strut forces. Comparisons of the observed and computed data revealed 20% maximum and 12% average errors for the computed dhm and strut forces, respectively. 4. Based on the numerical results, the concrete slab used in the TD construction had a significant effect on the lateral movement of the wall. When the concrete slab was introduced, the lateral
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Fig. 9. Effect of wall thickness on (a) mean value of dh/He and (b) strut force.
movements were sharply reduced, with an average improvement ratio of 2.4. The DCM-TD in the field case study required half the wall thickness of the DCM-BU to obtain the same mean value of dhm/He. A wall thickness greater than 3.7 m was considered an economically impractical design for DCM-TD based on the improvement ratio of 2.0. 5. From the numerical analysis, all strut forces decreased with increasing wall thickness. Thus, the number of struts can be reduced for the thick wall. The effect of the DCM wall thickness appears to be insignificant when the wall thickness is greater than 2.5 m.
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