Introduction
From May 2005 to January 2010, a large basement was excavated for a 301.8-m-high landmark skyscraper, the Gate of the Orient, in Suzhou, China. This megastructure would consist of two 68-story main towers, a 7-story podium, and a 5-story basement. The main towers would comprise reinforced concrete core tubes to be connected with perimeter mega columns via steel outriggers; the podium would be typical frame structures. The two main towers would converge at the height of 238 m to form a 301.8-m-high superstructure in an arch shape. Its basement would have dimensions of approximately
to 30.7 m deep (see Fig.
1). The site was bounded by several utility pipelines buried at depths of 0.5–1.0 m below ground surface (BGS) to the east, an artificial river (Xiangmen) to the south, and green lands to the north and west.
The site was located on Yangtze River Delta, approximately 110 km west of Shanghai. Prior to construction, field exploration tests, including 21 standard penetration tests (SPTs) and 41 cone penetration tests (CPTs), had been carried out across the site to characterize its subsurface condition. As shown in Fig.
2, the ground comprised 2-m-thick fill (Layer I), overlying 22-m-thick firm to stiff clayey and silty soils (Layer II). Beneath was stiff to hard silty clay and clayey silt (Layer III) to a depth of 38 m BGS, overlying medium dense to very dense silty sand (Layer IV) to 44 m BGS. Below Layer IV, it was stiff to very hard clay and silty clay with a thickness of approximately 48 m (Layer V), followed by very dense silty fine sand (Layer VI) to the termination depth of field exploration at 130 m BGS. The phreatic water level varied from 1.0 to 2.2 m BGS, fluctuating with weather changes. There were two aquifers: one was a perched aquifer, located at 6–12 m BGS with an elevated water level at 1.3–1.6 m BGS; the other was a confined aquifer located at 38–44 m BGS with an elevated water level at 4.33–4.63 m BGS. This indicated that if the excavation depth was over 20 m, the weight of the soil strata above the confined aquifer would be unable to hold down the underlying aquifer pressure; i.e., rupture or uplift of basal soils would occur (e.g.,
Moore and Longworth 1979;
Chow and Ou 1999;
Tan and Lu 2017a). To derive soil parameters for design, laboratory tests (e.g., consolidation, direct shear, triaxial, and seepage analysis tests) were performed on Shelby tube soil samples, and measured soil parameters are presented in Fig.
2. Cohesion
and friction angle
were derived from consolidated undrained direct shear (CUDS) tests; consolidated undrained cohesion
and friction angle
as well as effective cohesion
and friction angle
were measured by consolidated undrained (CU) triaxial tests; void ratio
and constrained modulus
were derived from one-dimensional compression (consolidation) tests; water content
was measured by the oven-drying method; liquid limit
and plastic limit
were measured by photoelectric liquid-plastic testers; horizontal and vertical permeability coefficients
and
were measured by constant head permeability tests in the laboratory; and total soil permeability
was measured by in-situ injectivity tests.
Design and Construction
To save project cost and shorten construction duration, composite earth-retaining systems were designed for excavation of this basement. The topmost 7.85 m soils along the north and west pit sides were excavated by the sloped open-cut method (Fig.
3); as for the south side, the adjacent artificial river (18 m wide and 5.85 m deep) was drained, followed by demolishing the north river embankment, and then soils were cut to 7.85 m BGS (Fig.
S1). Because of the high phreatic level, light-well points were adopted for open-cut to lower groundwater below the excavation surface. To prevent slope failure, exposed cut surfaces were covered by steel-welded meshes and sprayed shotcrete. To safeguard Xing-Gang Street with buried utilities, a house and a bridge to the east (Fig.
1), upper soil removal along the east pit side was retained by soil nailing wall (SNW), i.e., mixed-in-place (MIP) piles, also known as soil-mixing wall (SMW), anchored by soil nails (Fig.
3).
Regarding excavation of the soils from 7.85 to 21.5 m BGS, it was supported by continuous bored pile (CBP) walls braced by concrete struts at three levels. To further restrain lateral wall deflection, the soils at 14.0–25.5 m BGS on the excavation side against the east CBP wall were reinforced by MIP piles (
). Soil removal below the first level of struts followed a basin-type excavation procedure down to 21.5 m BGS; i.e., excavation was conducted at the pit center, immediately followed by pouring concrete struts normal to CBP walls; once concrete reached 28-day strength, residual earth berms against the perimeter walls were removed, accompanied by casting diagonal struts. Fig.
4 shows the plan layout of struts cast at Levels 1 to 3. To facilitate soil disposal, concrete decking slabs (30 cm thick) were cast on normal struts at Level 1 (the shaded areas in Fig.
4). The small inner pits, were retained by SNWs; meanwhile, soils below excavation bases were jet-grouted (see Fig.
5). Detailed construction schedules refer to Table
1.
To build waterproof curtains along the pit perimeter, MIP piles were constructed along the outer perimeter of the CBP walls and gaps between them were compaction-grouted to achieve satisfactory watertightness; see Figs.
1 and
3. Throughout excavation of soils below 9.25 m BGS, groundwater inside the pit was kept at 0.5–1.0 m below the excavation surface via one hundred and thirty 26–28 m deep pumping wells (Fig.
6). Because the confined aquifer located at 38–44 m BGS could cause uplift or rupture failure of basal soils or floating of basement structures, nine 44-m-deep relief wells were constructed inside the pit to reduce artesian pressure (Fig.
6).
Instrumentation
Compared to excavations in soft clays (e.g.,
Ou et al. 1993;
Koutsoftas et al. 2000;
Hashash et al. 2008;
Wang et al. 2010;
Tan and Li 2011;
Tan and Wei 2012;
Tan and Wang 2013a,
b;
Tan et al. 2014,
2015b,
2016,
2017;
Finno et al. 2015;
Whittle et al. 2015;
Xu et al. 2015;
Tan and Lu 2017b,
2018), excavations in stiff clays have received much less attention (
Tedd et al. 1984;
Ng 1998;
Long 2001;
Liao et al. 2015;
Tan et al. 2015a). With respect to a large deep excavation supported by composite earth-retaining systems, none was reported in the literature, and current methods for predicting pit performance might not be applicable to such cases any more. With the progress of urbanization, more and more large-sized deep excavations (e.g.,
Tan and Wang 2013a,
b;
Tan et al. 2015a) have been, and are to be, constructed in urban areas for using underground space. To save project cost or shorten construction duration, unusual composite earth-retaining systems are frequently used for such large excavations, which inevitably bring uncertainties to design and construction. Therefore, a well-documented case history such as this one would be a useful addition for upgrading the current state of the art and practice in deep excavation.
To ensure project safety and investigate excavation performance, this pit was extensively instrumented (Figs.
1 and
4). The monitored items included: (1) lateral CBP wall displacements at P1 to P14, (2) vertical and lateral displacements at waterproof curtains (Q1 to Q12) along the east side and slope crests along other sides (Q13 to Q23), (3) vertical CBP wall displacements at W1 to W38, (4) vertical column displacements at LZ1 to LZ43, (5) axial strut forces at 19 locations for each level (Zi-1 to Zi-19;
, 2, 3), (6) ground settlements along six survey sections (C1-1–C1-6 to C6-1–C6–5) and lateral ground displacements at T1 to T11, (7) basal heaves at HT1 to HT12 [at each location, heaves were measured at excavation base and 3 and 6 m below, designated as HTi-1 and HTi-3 (
–12)], and (8) variations in both phreatic levels at SW1 to SW20 (20 m deep and 2 m behind CBP wall) and artesian levels at CY1 to CY4 (40 m deep and 25 m behind CBP wall). Moreover, settlements of the adjacent building (F1–F4), bridge abutments (X1–X4), and those utility pipelines buried below Xing-Gang Street were closely monitored throughout construction as well.
Lateral displacements of CBP walls were monitored by deflection measurement systems (Fig. S2), which consisted of plastic inclinometer casings, standard slope indicators with accuracy of 0.1 mm, vibrating wire readout boxes for taking readings, and computer software for data analysis and processing. The used slope indicators (probes) had a resolution of , and the tilt of the walls was measured at 1-m intervals along depth with an accuracy of . As for the lateral ground displacements, they were measured by inclinometer casings placed inside predrilled boreholes in the ground behind the pit. The measured deflection data by slope indicators were the relative displacements to the casing tops. To know the actual lateral displacement caused by excavation, the lateral movements at casing tops were also surveyed. This was done by a theodolite with an accuracy of 2 s. Therefore, the actual lateral wall or ground displacements were the sum of the measured data of inclinometer casings and the surveyed data of casing tops.
Regarding vertical displacements of CBP walls, at each selected survey point, a stainless steel piece was embedded at the top of CBP walls (Fig. S2), and their vertical movements were measured by a level instrument with an accuracy of 0.01 mm. The survey of vertical and lateral displacements of waterproof curtains, slope crest, and interior columns was adopted in the same way for CBP walls. Ground settlements behind the pit were measured via surface markers. They consisted of round-head steel rebars with a diameter of 20 mm and length of 200 mm, which were encased inside capped protection wells (Fig. S3). The protection wells were cast-iron pipes with a diameter of 150 mm and length of 700 mm. In order to secure the surface markers, the soils around the protection wells were compacted with cement mortar. With respect to the settlement survey for buried utility pipelines, at each survey point, the ground was excavated to the pipeline. Then, the pipeline was enclosed with a cast-iron ring, which was connected with an iron rod located vertically on the top of the pipeline. The iron rod was protected by a capped PVC pipe and the gap between them was filled with expansive clay. The iron rod was approximately 0.5 m below the ground surface (Fig. S4). The survey points for settlements of the building and bridge abutments consisted of steel pieces, which were inserted into predrilled holes on the exterior walls of the buildings or bridge piers to be monitored. Gaps between the steel pieces and the walls or piers were filled with cement mortar (Fig. S5).
Prior to commencement of the work, the ground, building, bridge, and utility pipelines within the predefined influence zones were surveyed to establish benchmarks (reference points) for monitoring the influence of the excavation. These benchmarks were constructed at approximately 100–300 m distant from pit, which was far enough not to be disturbed by the excavation. They consisted of concrete piers, which were embedded into undisturbed natural subgrade and protected by concrete wells. The vertical distance from the ground surface to the top of the piers was 0.2 m. Configuration of a typical benchmark is illustrated in Fig. S6.
Axial forces in reinforced concrete struts were measured by vibrating wire steel stress meters. Before concreting, four stress meters were welded on four reinforcing bars along the longitudinal direction of the struts, which were located at the middle points of the four sides of the rectangular steel reinforcement cage (Fig. S7). Cables of the stress meters for taking readings were protected by PVC pipes. Then, the instrumented rectangular steel reinforcement cage, along with the stress meters and PVC pipes, were poured inside the struts. The changes of axial force in the concrete struts would cause variation of tensile force in the steel wires of the stress meters and then change their vibration frequency. By recording vibration frequency with a frequency meter, the axial force of the concrete struts can be calculated using the calibrated relationship curves between frequency and axial force.
Groundwater levels were monitored by standpipes at approximately 2 m (SW1 to SW20) and 25 m (CY1 to CY4) behind the pit. Holes with a diameter of 100 mm were drilled in the ground at the selected locations (20.5 m deep at SW1 to SW20 for phreatic levels and 40.5 m deep at CY1 to CY4 for artesian levels). Then, 20–40-m-long and 50-mm diameter PVC pipes with many tiny holes, which were wrapped up with nylon nets to facilitate seepage of water, were placed into the holes at SW1 to SW20 and CY1 to CY4, respectively. The heads of the PVC pipes were 50 mm above the ground surface and capped for protection. Then, gaps between the PVC pipes and the holes were filled with soils. The fills consisted of expansive clay in the upper 4 m for both SW1 to SW20 and CY1 to CY4, whereas clean sands were used in the lower sections for SW1 to SW20 and in-situ soils for CY1 to CY4; see Fig. S8. Via a portable water level indicator, water level changes inside the standpipes were measured.
Basal heaves were measured by 12 clusters of magnetic extensometers installed inside 12 boreholes (designated as HT1 and HT2 in Fig.
4). In each borehole, three magnetic loops were installed at 0, 3, and 6 m below the excavation base (designated as HT1-1–HT1-3 and HT12-1–HT12-3 along the depth). The magnets were placed on the outside of a PVC pipe inside each borehole, and the gap between the borehole and the PVC pipe was filled with sandy soils. By pulling up the PVC pipe slightly, the barbs of the magnetic loops could pierce into the fill sands and clinch to the natural subgrade tightly (Fig.
S9). By lowering a magnetic probe into the PVC pipe, the vertical displacements of the magnetic loops could be measured.
Ground Settlement
The ground behind the pit developed significant settlements
during excavation, which were up to 168 mm behind the east pit side and 107 mm behind the north and west sides; see Fig.
12. The bar charts in Fig.
13 summarize ground settlement increments
in each construction stage along C1-1 to C1-6, C2-1 to C2-4, and C3-1 to C3-4 behind the pit middle spans, in which
was normalized by the corresponding final measurement
, and distance
between survey point and pit was normalized by
. Throughout excavation,
increased dramatically over time and did not show a sign of stopping until completion of the base slab in Stage S7. Although settlement rates in Stages S7 to S8 were much smaller than those in Stages S3 to S6, the cumulative
developed in 1 year (Stages S7 and S8) were considerable, up to
. Considering that the rapid recovery of artesian water after completion of excavation (Fig.
7) would try to push up its overlying impermeable clayey strata (aquitard) somewhat, the majority of these postexcavation
should result from postexcavation inward lateral wall displacements associated with dismantling rigid concrete struts (Fig.
9) rather than consolidation of the clayey strata over time. This deduction can be also validated by the very limited postexcavation settlements of the shallowly buried utility pipelines and the bridge piers behind the east CBP wall (Fig.
10).
Ground Settlement Profile
Fig.
14 presents distribution of
along distance
from the pit, in which both
and
were normalized by
. Furthermore, data available from Cases 11, 12, 16, and 17 and Moormann (
2004), as well as the empirical envelopes in literature (
Peck 1969;
O’Rourke 1976;
Clough and O’Rourke 1990;
Hashash et al. 2008), were included. Obviously,
behind this multipropped excavation had spandrel-type profiles typical for excavations retained by cantilever wall (e.g.,
Peck 1969;
Clough and O’Rourke 1990) instead of the concave type typical for multipropped excavations (e.g.,
Ou et al. 1993;
Tan and Wang 2013a,
b). This inconsistency should derive from excavation of the uppermost 7.85 m soils being achieved by open cut or retained by SNWs. Evidently, the composite earth-retaining systems adopted in this excavation made its ground settlements distinctively different from those of other multipropped excavations. Within
from the pit, the ground along C1-1 to C1-6 behind the east pit middle span had greater
than the corresponding C2-1 to C2-6 and C3-1 to C3-4 behind the west and north pit middle span. In spite of this discrepancy, they had similar ground settlement influence zones, which can be bounded by envelope (5). Along the same side, C1-1 to C1-6 behind the middle span had apparently greater settlements than both C5-1 to C5-5 and C6-1 to C6-5 distance away, which exhibited clear evidence of spatial corner stiffening behavior like those cases in Tan et al. (
2014).
For this excavation, zone I of Peck (
1969) for sand and soft to hard clay highly underestimated its
and ground settlement zone;
in this case was up to 0.8% within zone II of Peck (
1969) for very soft to soft clay. Envelope (2) of O’Rourke (
1976) for soft to medium stiff Chicago glacial clay would highly overestimate
within
for this case, but underestimate its
beyond
. Regarding envelope (3) of Clough and O’Rourke (
1990) for stiff to hard clay, it could make a reliable estimation on the ground settlement zone, but would highly underestimate
. Envelope (4) of Hashash et al. (
2008) for the CA/T project in medium stiff BBC would highly underestimate
, but overestimate the ground settlement zone. As for the other four Suzhou excavations, envelope (3) of Clough and O’Rourke (
1990) could make a good prediction in terms of both
and the ground settlement zone. With the exception of this case, most
in Suzhou and Moormann (
2004) were within zone I of Peck (
1969). Apparently, the measured ground settlements in this case were much greater than those of Cases 11, 12, 16, and 17 in Suzhou.
It has been recognized in the literature (e.g.,
Galloway et al. 1999;
Xu et al. 2016;
Wu et al. 2017) that discharging of ground water (phreatic and artesian water) could cause pronounced land subsidence. Considering the significant drawdown in artesian water levels behind this excavation (Fig.
7), it was suspected that the exceptionally larger ground settlements of this excavation compared to the other excavation cases in Suzhou might be associated with dewatering. As illustrated in Table
2, Case 11 had
and
; i.e., its DW had penetrated through the confined aquifer at 38–44 m BGS and extended into the underlying impermeable clay and silty clay stratum (aquitard). Because the DW had relatively good watertightness, pumping of aquifer water inside the pit for releasing artesian pressure in Case 11 would not cause substantial drawdown in the water level outside the pit; i.e., dewatering-induced ground settlement would not be significant for Case 11. As for Cases 12, 16, and 17, their
were no more than 20 m, and dewatering for releasing underlying confined aquifer pressure was not required; i.e., as mentioned previously, for excavations with
in Suzhou, the weight of overburden soils was able to resist upward confined aquifer pressure. Like this case, the DWs of Cases 12, 16, and 17 were toed in the upper clayey strata and did not penetrate into the underlying confined aquifer. In light of the good watertightness of the DW and the low permeability of the upper clayey strata (aquitard), pumping of phreatic water inside the pit would not cause a substantial drawdown in the water level outside the pit for Cases 12, 16, and 17. Different from Cases 12, 16, and 17, pumping of confined aquifer water inside the pit was conducted in this case because of its great excavation depth (21.5–30.7 m). However, its waterproof curtain was toed at 27.3 m BGS and did not penetrate deeply through the confined aquifer stratum [see Figs.
3 and
8(a)]. Consequently, significant drawdown in the water level outside the pit was monitored during dewatering (see Fig.
7). On the basis of these analyses, it was reasonable to infer that the dramatically larger ground settlements of this excavation compared to those of Cases 11, 12, 16, and 17 should, to some extent, result from the significant drawdown in the aquifer level.
Discussion on Ground Settlement Caused by Dewatering
The preceding analyses disclose that the drawdown in water level imposed a significant adverse influence on ground settlement. As plotted in Fig.
7(a), the measured drawdown in phreatic level by the 20-m-deep observation wells at 2 m behind CBP walls was up to 2 m along the east pit side and up to 11 m along the other three sides. The drawdown in the underlying confined aquifer level measured by the 40-m-deep observation wells at 25 m behind the CBP walls was up to 14 m along the pit perimeter; see Fig.
7(b). As schematically illustrated in Fig.
8(b), drawdown of the water (phreatic or artesian) level in the ground behind the pit would incur consolidation-induced settlement
for the upper 38 m silty clay and clayey silt strata (aquitard) and compression
of the underlying silty sand layer (confined aquifer). Therefore, dewatering-induced total ground settlement
can be given by
As for the upper aquitard [Fig.
8(b)], dewatering of its phreatic water levels from
to
caused increment of vertical effective soil stress
with a magnitude of
in the aquitard, in which
= density of water,
= phreatic level before dewatering, and
= phreatic level after dewatering. Meanwhile, releasing (dewatering) of artesian water in its underlying confined aquifer led to reduction of
in the upward artesian pressure against the aquitard bottom, which can be treated as a downward pressure
applied to the bottom of the upper aquitard after dewatering, where
= confined aquifer level before dewatering and
= confined aquifer level after dewatering. According to the one-dimensional consolidation theory of Terzaghi (
1925), the change in pressures acting on the boundaries of soil strata would result in consolidation (compression). Although the water level in the upper perched aquifer at 6–12 m BGS was not monitored during excavation, it should rarely affect ground settlement. As illustrated in Fig.
8, the waterproof curtains (CBP wall and MIP) extended to 27.3–34.05 m BGS, and this perched aquifer was intermediate between the two impermeable aquitards; therefore, dewatering conducted inside the excavation hardly influenced the water level of the perched aquifer outside the excavation. Hence, it can be postulated that the upper perched aquifer should rarely affect ground settlement behind the excavation. Based upon the preceding analyses, it can be concluded that for the upper 38-m-thick soil strata overlying the confined aquifer at 38–44 m BGS, its
primarily consisted of
due to drawdown
in its phreatic level and
due to drawdown
in its underlying confined aquifer level. In accordance with the one-dimensional consolidation theory of Terzaghi (
1925),
can be approximately estimated by the following equations:
where,
= consolidation-induced ground settlement at time
for the upper aquitard;
associated with drawdown in phreatic level;
associated with drawdown in the underlying confined aquifer level;
= consolidation ratio;
= final consolidation-induced ground settlement;
associated with drawdown in the phreatic level;
associated with drawdown in the underlying confined aquifer level;
= thickness of the upper aquitard, equal to 38 m;
= compressibility coefficient of the upper aquitard;
= initial void ratio of the upper aquitard;
= constrained modulus of the upper aquitard; and
= permeability coefficient of the upper aquitard.
As shown in Fig.
7, the reductions in the phreatic level and underlying confined aquifer level were up to 2 and 14 m behind the east pit side (
;
) and 11 and 14 m behind the other three sides (
;
). Thereafter, the phreatic levels showed limited variation throughout Stages S3 to S7. According to Eqs. (
2)
–(
10), the roughly estimated
during excavation was approximately 68 and 55 mm for the upper aquitard behind the east and the other three sides, respectively.
Regarding the underlying confined aquifer layer (silty sand), drawdown in the aquifer level would increase its effective stress with a magnitude of
, which would lead to compression settlement
of the silty sand layer. It can be approximately estimated by
where
= thickness of the confined aquifer layer, equal to 6 m; and
= constrained modulus of the silty sand (confined aquifer layer). The estimated
could be up to 60 mm. Based on the previous analysis, the total ground settlement (
) related to dewatering could be up to 128 mm behind the east pit side and 115 mm behind the other three sides.
As illustrated in Fig.
8(a), the 44-m-deep relief wells for releasing confined aquifer pressure were completely penetrating wells; i.e., they penetrated through the confined aquifer and reached the underlying aquitard. The influence zone due to pumping of aquifer water can be roughly delineated with the empirical equation proposed in Liu and Wang (
2009), i.e.,
where
= radius of influence from relief well;
= drawdown in aquifer level; and
= permeability of confined aquifer layer measured by in-situ injectivity test. The estimated
could be up to 45 m distant from this pit, more than
. This in part explains why the ground settlement of this case measured at a distance of
behind the pit was still significantly greater than the other cases in Suzhou; see Fig.
14.
The previous theoretical analyses disclose that the majority of the pronounced ground settlement measured behind this excavation should derive from the significant drawdown in the aquifer level outside the pit, whose influence could reach as far as
behind the pit. The significant drawdown in the water level largely arose from the previously mentioned deficiency in the waterproof design. This deficiency led to the water flow path in the confined aquifer layer not being cut off between the pit and the ground outside [Fig.
8(a)] during discharging underlying confined aquifer water inside the pit; consequently, a significant drawdown in water level was monitored behind the pit. If such a project were constructed in a densely populated urban environment, a safer design would have to be adopted to minimize ground subsidence to protect adjacent structures or facilities, e.g., (1) extending the waterproof curtain deeply enough to penetrate through the underlying confined aquifer despite the risk of increasing project cost dramatically, or (2) recharging the underlying confined aquifer with the water discharged by relief wells inside the pit in the proximity of the structures or facilities to be protected (see the illustrations in Fig.
15).
Vertical Displacements of CBP Walls, Waterproof Curtains, and Interior Columns
Fig.
17 presents development of vertical displacements
over time for CBP walls. Except for the survey points near the northeast (W1 to W3) and southwest corners (W18 to W20), both the east and south CBP walls moved downward during excavation. On the contrary, the north and west CBP walls moved upward. These opposite movements should be related to the different retaining systems used. As for the east CBP wall, upward force on the inner wall face and wall toe resulting from basal rebounding could not overcome downdrag force on the outer wall face caused by ground settlement; consequently, the east wall settled instead of upheaving. Regarding the north and west CBP walls, because of the open cut, downdrag force on the outer wall face because of ground settlement behind the pit could not suppress wall upheaving resulting from basal rebounding; hence, they moved upward rather than settling. As for the south CBP wall, its open cut was not as large as those behind the north and west walls; therefore, its
had an intermediate development tendency. In spite of this inconsistency, following completion of excavation and subsequent construction of underground structures, all CBP walls started to settle gradually over time.
Fig.
18(a) presents typical vertical displacements of waterproof curtains along the east wall (Q1 to Q12). Like their corresponding lateral displacements (the maximum was 92 mm near the pit middle span and the minimum was 53 mm near corners (see Fig.
19), settlement data of the east waterproof curtain exhibited strong corner stiffening behavior as well; i.e., the maximum settlement (130 mm at Q6) occurred near pit midspan and the minimum (52–54 mm at Q1, Q2, and Q12) near corners.
Fig.
18(b) plots development of typical vertical displacements
over time for interior columns within the northeast pit zone. Generally,
data of columns LZ1 to LZ4 near the east CBP wall were much smaller than those of the others. Contrary to expectations,
of LZ27 and LZ28 near the north wall was not smaller than those of LZ29, LZ31, LZ32, and LZ42 near the pit center featuring the largest basal heave. This should be the combined results of the following two factors: (1) columns near the pit center were directly below heavy decking slabs—hence, their upheaving was suppressed to some extent; and (2) settling of the east CBP wall restrained upheaving of adjacent columns.
Axial Strut Force
Figs.
21 and
22 summarize axial force developments for concrete struts cast at Levels 1 and 2. Because the measured axial strut forces at Level 3 (Fig.
S10) had similar magnitudes and development patterns as those of Level 2, they are not presented herein. Distinct from those at Levels 2 and 3, axial forces of normal struts were significantly larger than axial forces of diagonal struts by 10-fold at Level 1. These extremely large axial forces should arise from operation of heavy-duty trucks (Fig.
4), which caused bending of underlying normal struts (i.e., significant bending stresses were generated). Although upper soils outside both the southwest and northwest pit corners were openly cut, axial forces of diagonal struts Z1-15 and Z1-16 against the northwest corner were much larger than those of Z1-6 and Z1-7 against the southwest corner. If soils above 7.85 m BGS behind the CBP wall were treated as surcharge (Fig.
23), much larger additional stresses were generated behind the north and west walls than the south wall. Therefore, Z1-15 and Z1-16 should sustain greater axial forces than Z1-6 and Z1-7, which was evidenced by smaller lateral displacement of the south wall than the north and west walls. As shown in Figs.
3 and
23, the ground behind the east CBP wall sustained a much larger surcharge than the ground behind the other walls. However, axial forces of diagonal struts Z1-1, Z1-2, Z1-3, and Z1-4 against the east wall were not greater. As mentioned previously, the east SNW (Q1 to Q12) moved significantly toward the pit during excavation, and its lateral displacements were up to 53–92 mm (Fig.
19); thus, anchor forces of the soil nails were mobilized to resist lateral SNW displacement. This resulted in the cohesive soil mass anchored by soil nails being pulled upward somewhat (i.e., the vertical load from the anchored soil mass behind the east CBP wall was alleviated to some extent) and meanwhile, the east CBP wall was pushed inward by the displaced SNW. Hence, struts against the east wall did not sustain larger lateral earth pressures than others, whereas the east CBP wall developed greater lateral deflections.
Different from axial strut forces at Level 1, axial forces of diagonal struts at Levels 2 (Z2-6 and Z2-7) and 3 (Z3-6 and Z3-7) against the southwest corner were not smaller than those of diagonal struts at Levels 2 (Z2-15 and Z2-16) and 3 (Z3-15 and Z3-16) against the northwest corner. This indicates that at 13.6 m BGS (Level 2), additional stresses induced by the surcharge (uppermost 7.85 m soils) were already insignificant.
As shown in Figs.
21 and
22, for struts cast at the same levels, their axial force magnitudes varied significantly. Beyond lateral earth pressures, struts might sustain bending stresses. Vertical differential displacements between adjacent columns could be indicative of strut bending. To verify this, vertical column displacements below normal struts at different excavation levels were plotted in Fig.
24. Apparently, there was no differential displacement between the two columns near either Zi-18 or Zi-19. In contrast, the columns adjacent to Zi-9, Zi-10, and Zi-11 had significant differential uplifts. These observations corresponded well with Zi-9, Zi-10, and Zi-11 having the maximum axial forces, whereas Zi-18 and Zi-19 had the minimum.