Pile test at the Shard London Bridge

Transcription

1 technical paper Pile test at the Shard London Bridge David Beadman, Byrne Looby Partners, Mark Pennington, Balfour Beatty Ground Engineering, Matthew Sharratt, WSP Group Introduction The Shard London Bridge, designed by Renzo Piano and developed by Sellar, is adjacent to London Bridge railway station. The tower will become the tallest building in Western Europe standing at 310m and will include 60,000m 2 of office, hotel and residential accommodation. Sellar appointed WSP as the structural consultant for the project and MACE as principal contractor. Stent Foundations (now Balfour Beatty Ground Engineering) undertook the design-and-build contract for the foundation piles and the secant pile retaining wall, with Byrne Looby Partners undertaking the detailed design of the foundations and retaining wall for Stent Foundations. The foundations comprise up to 60m deep 1.8m diameter bored piles founded in the underlying Thanet Sand. This paper reports on a 30MN preliminary pile test carried out on a 53.8m deep 1.2m diameter pile. The pile was instrumented in detail to record settlements and strains at various levels, enabling the mobilisation of shaft friction with settlement to be determined through the various strata. Ground conditions A site investigation for the development was completed in 2007 and the ground conditions encountered in the exploratory holes comprised made ground overlying alluvium, River Terrace Gravel, London Clay, Lambeth Group Beds, Thanet Sand and the Upper Chalk. A fault is present below the site trending in a general north-to-south direction. This fault has a 5m throw and a downthrow to the east of the fault line. Notation mod metres Ordinance Datum msd metres Shard Datum (0mSD=4.3mOD) FOS factor of safety c u undrained shear strength α adhesion factor for undrained soil, where skin friction = α.c u N q * equivalent bearing capacity factor for Thanet Sand σ av average effective stress σ v effective vertical stress K s effective lateral stress coefficient at perimeter of pile, where lateral stress = K s.σ v The presence of such geological features within this area of London is documented in the Ciria report Building Response to Tunnelling (Burland et al, 2001) on the Jubilee Line construction where similar features were noted during the tunnelling works. The fault lies below the footprint of the tower and resulted in displacement of the Table 1: Strata levels Table 2: Soil strength parameters (pile design) Strata Density (kn/m 3 ) Strata Angle of friction (degrees) K o ratio of horizontal effective stress to vertical effective stress under at rest conditions Φ angle of friction of soil δ angle of friction on perimeter of pile ε y average strain reading of the four strain gauges placed at a distance y from ground level A G cross sectional area of the pile E G gross stiffness of the pile E c stiffness of the concrete (varies with strain) London Clay, Lambeth Group Beds, Thanet Sands and the Upper Chalk. The typical ground conditions on either side of the fault are listed in Table 1. Data from either side of the fault was plotted and indicated that a good match was achieved when plotted as metres below the top of the London Clay rather than metres above the Thanet Sands or chalk, Elevation of top of strata msd West of Fault East of Fault Preliminary pile record Made Ground 0 Alluvium Not recorded Terrace Gravels -4.4 to to London Clay -9.3 to to Lambeth Group to to Thanet Sand to to Chalk to to Not recorded Effective cohesion (kn/m 2 ) Undrained cohesion (kn/m 2 ) Maximum shaft friction limit (kn/m 2 ) Maximum base bearing capacity limit (kn/m 2 ) Made Ground 18 n/a n/a n/a n/a n/a Terrace Gravel n/a n/a n/a London Clay 20 n/a /m 140 n/a Lambeth Group 20 n/a n/a Thanet Sand n/a ,000 suggesting that the faulting occurred relatively early in the depositional history of the London Clay. The made ground comprises hardstanding surfaces over soft to firm slightly sandy gravelly clay which in turn overlay soft brown gravelly clay (alluvium) and medium dense to very dense sandy gravel (River Terrace Gravel). The London Clay comprises firm to very stiff brown grey to grey brown or grey fissured or locally thinly laminated clay. The Lambeth Group beds comprise the following typical sequence of strata: n Very stiff multicoloured fissured clay (Upper Mottled Clay); n Very stiff grey clay with lenses/ pockets of silt and sand (Laminated Beds); n Very stiff grey or black shelly clay (Lower Shelly Beds); n Very stiff multicoloured fissured clay (Lower Mottled Beds); n Clayey gravel, green gravelly sand, green shelly gravelly clay (Upnor Formation); The underlying Thanet Sands comprise a very dense silty fine to medium sand. Pressuremeter testing using a Menard pressuremeter was undertaken within Thanet Sands and the results from these tests plotted against depth below the top of the Thanet Sands are presented as Figure 1. The results showed the expected decrease in limit pressure with depth in the Thanet Sands, as described by Nicholson et al (1992). Groundwater monitoring indicated that hydrostatic conditions were generally present through the London Clay and underlying Lambeth Group beds with under drainage in the Thanet Sands and Upper Chalk. Within the River Terrace Gravels, groundwater was present at -4.8mSD. The lower aquifer level, in the Upper Chalk, was recorded at approximately -38mSD. Pile construction The Shard has a three-level basement over the entire footprint of the site. The piles to support the structure were constructed from the existing ground level with approximately 15m of empty bore above the cutoff level of the pile. Many piles were cast with plunge columns to facilitate top-down construction. 24 ground engineering january 2012

2 Elevation mod BH7 BH8 BH9 Figure 1: Pressuremeter limit pressure v elevation (mod) The test pile was constructed west of the fault from the existing ground level and debonded to the proposed basement level to model the actual loading applied to the piles in the permanent condition. The debonding was achieved using a bitumen coated steel tube. Figure 2: Slip liner prior to installation Limit pressure MPa The photograph (Figure 2) shows the bitumen coated liner prior to installation. All the piles including the test pile on the site were constructed using standard rotary boring techniques. Temporary casing was installed through the made ground, River Terrace Gravels and approximately 1m into the London Clay to provide a water seal during pile construction. The piles were excavated in an open bore to a depth of approximately 35m below ground level. Bentonite support fluid was introduced into the pile bore to provide support to the pile during construction to the required toe depth. The majority of piles were founded in the Thanet Sand at a depth of approximately 55m below ground level. For the test pile, a 1,650mm OD temporary casing was installed to a depth of approximately 12m below ground level. The pile was then excavated at 1,500mm diameter through the casing to a depth of 19.5m. The bitumen-coated slip liner was lowered into the bore. The annulus around the liner was then grouted in two stages to secure it into position. The first-stage grout was poured to a level just below the temporary casing. This was allowed to reach initial set prior to the temporary casing being extracted and the second stage of grout was poured. The pile was then excavated through the bitumen-coated liner to the toe depth of 53.8m below ground level. Bentonite support fluid was introduced into the bore during the excavation. It was anticipated that each plunge column pile would take two days to construct, being partially excavated during the first day and then left overnight prior to completing the excavation, placing the reinforcement, concreting and installing the plunge column on the second day. To assess the impact of leaving the pile partially excavated overnight, the test pile was constructed over a similar two-day period. In the event, a problem with the piling rig during the construction of the test pile meant that the pile was left excavated to approximately 1m above the final toe level for two nights. The bentonite support fluid was maintained in the bore to a level inside the permanent casing during this time. A full-length reinforcement cage with four vibrating wire strain gauges at six different levels was installed into the pile bore. Strain gauges were positioned at the top and the bottom of the slip liner to establish the shaft friction along the slip liner and at other levels to identify the mobilised shaft friction in the various strata. Two extensometers were installed into the pile to measure the movement of the pile toe during loading. Following the cage being positioned, the concrete was tremied into the pile displacing the bentonite support fluid. The test pile was constructed between 25 and 27 March 2009 and tested over a four-day period from April Concrete cubes were crushed to ensure the concrete had reached adequate strength to resist the applied stresses. The proposed pile test load meant that the concrete stresses under the jacks applying the load and the concrete stress in the top section of the pile were high. A heavily reinforced pile cap was designed and constructed to spread the loads from the jacks into the pile (Figure 3). Pile design The soil strength parameters adopted for the design are tabulated in Table 2. In the London Clay, the undrained shear strength was assessed based on the Stroud correlation between SPT and undrained shear strength of c u = 4.5N. The skin friction in London Clay and in the cohesive Lambeth Group was based on α.c u where c u is the undrained shear strength. α (alpha) was taken as 0.5 assuming pile construction within a single shift. Where pile construction was predicted to extend beyond 24 hours, the value of α was reduced to The maximum shaft friction was limited to 140kN/ m 2 as recommended by the London District Surveyors Association (1999). The angle of friction for the Thanet Sand was based on a typical value from published data (see references, specifically Chapman et al, 1999). The ultimate end bearing was limited to 20,000kN/m 2 (Suckling and Eager, 2001) and the ultimate shaft friction to 200kN/ m 2 (Chapman et al, 1999). The calculated base capacity was based on N q*.σ av where N q* = 47 and σ av = average effective stress at the pile ground engineering january

3 technical paper Figure 3: Pile cap and jacks toe (Troughton and Platis, 1989), which was based on a K o value (ratio of horizontal to vertical effective stress under at rest conditions) of 1.15 by Troughton et al. The base capacity exceeded the limiting 20,000kN/m 2, particularly when the existing lower ground water pressure was applied (as noted below). The base capacity did not, however, exceed that indicated from the pressuremeter testing, which would be of the order of 32,000kN/m 2. The shaft capacity was based on K s.tanδ.σ v where lateral stress coefficient, K s = 0.7, pile interface friction angle, δ = angle of friction of the soil, Φ and σ v = effective vertical stress. The calculated skin friction at the depth of the Thanet Sand exceeded the limiting 200kN/m 2. The Thanet Sand on this site was deeper than for the majority of the case histories reported to date and therefore the vertical and average effective stresses were much higher in the Thanet Sand. Pile design theory relates the shaft and base capacities directly to the effective stress as noted above and therefore it is important to be cautious when extrapolating pile capacity from existing case history data. This was the reason for applying the conservative limits on the shaft friction and end bearing capacities. It is important to recognise that the theoretical design pile capacity is based on the worst case conditions anticipated during the working life of the pile, specifically ground water conditions. Conditions during the pile test may be less onerous. The ground water level in the Terrace Gravels was measured as 4m below ground level during the pile construction. The water level in the underlying Thanet Sand was recorded in the pile drilling log as 46.5m below existing ground level (-46.5mSD), although the site investigation and the recent deep aquifer monitoring data both indicate the water level to be higher, at approximately -38mSD. This higher figure is considered to be realistic. The results of the analysis Table 3: Strata levels of the test pile using the measured strata levels (Table 1) and the ground water level at -38mSD are included in Table 4, using the same limits on shaft friction and base capacity as adopted in the design calculations, with an alpha value of 0.5. It is noted that the differing ground water level records (-38mSD or -46.5mSD) do not change the results because of the use of the limiting values of shaft and base resistance. The piles were designed for an overall factor of safety of Working pile tests proved very difficult to locate on the site due to the congested nature of the site and the presence of many existing piles. It was therefore agreed that working pile tests would not be carried out. Instead, it was decided to design for a factor of safety of 2.25 instead of the normal 2.0 associated with preliminary pile and working pile tests or 2.5 for working piles tests only, as defined by the London District Surveyors Association (1999). The pile design was also checked for a factor of safety of 1.2 on the shaft friction alone as a crude method of controlling pile settlements. The ultimate pile capacity was calculated as 41,150kN. Applying a factor of safety of 2.25 gave a working load of 18,300kN. However, as noted above, the pile design was also checked for a factor of safety of 1.2 on the shaft and this reduced the pile capacity to 15,450kN. This was taken as the Table 4: Test pile design capacity Capacity Overall ultimate capacity Value kn 41,150 Shaft Capacity 18,530 Base Capacity 22,620 Working load (FOS 15, or 1.2 on shaft) Figure 4: Pile test frame Parameter Nominal Diameter (m) Length (m) Toe Level (msd) Temporary casing toe level (msd) Temporary casing internal diameter (m) Temporary casing external diameter (m) Founding Stratum Penetration into Founding Stratum (m) Value Thanet Sand 8 26 ground engineering january 2012

4 Displacement Displacement (mm) (mm) Displacement Displacement (mm) (mm) Load Load (MN) (MN) Load (MN) Figure 5: Pile load v displacement Stress Stress N/mm N/mm 2 2 Stress Stress N/mm N/mm BS8110 Part Part 1 1 Figure Figure fcu BS8110 fcu = 42N/mm = Part Part Υm Figure Υm = 1.0 = fcu fcu = = 42N/mm Υm 500 Υm = = Strain Strain 800 x10 x Strain x10 x10-6 Figure 7: Measured stress-strain for pile concrete compared to BS 8110 curve Stiffness Stiffness x10 6 x10 kn/m 6 kn/m 2 2 Stiffness Stiffness x10 6 x10 kn/m 6 kn/m 2 2 Depth Depth msd msd Depth Depth msd msd Strain Strain 800 x10 x Strain x10 x10-6 Figure 6: Measured concrete stiffness v strain Load Load 25 kn kn (000s) (000s) Load kn kn (000s) Figure 8: Load distributions with depth working load for the purposes of assessment of the test results. Pile test data The test was carried out in accordance with the ICE Specification for Piling and Embedded Retaining Walls (2007). The pile was loaded (Figure 3) using a load cell and a reaction system consisting of a load frame and four 1,500mm diameter anchors to provide the reaction for the test as illustrated in Figure 4. The test pile was subjected to three cycles of load to maximum loads of 100, 150 and 222% of the originally specified working load of 13,500kN. The pile settlement in relation to load is shown on Figure 5. The maximum settlement at ground level recorded at 222% of the original working load (30,000kN) was 52.18mm. The maximum relative movement of the base recorded by the extensometers for the maximum load was 23.36mm, giving a maximum base movement of 28.82mm. The raw data from the pile test was assessed initially and any unusual or inconsistent readings obtained from the strain gauges were discounted. Errors in the strain gauges could be a result of erroneous or incorrectly installed gauges or damage to the gauges during construction of the pile. The results from three strain gauges, one each at -51mSD, -35mSD and -1.5mSD were discounted at this stage and the average strain at these levels calculated using the remaining three gauges. Temperature was measured for the gauges. However, Table 5: Load from strain gauges Jack load (kn) temperature compensation for the strain gauges was not considered necessary. A fluctuation of approximately 25 micro-strains occurred in the range 0-80 C which was not considered to compromise the reliability of the strain readings. The strain recorded in the strain gauges at -1.5mSD was considered to correspond to the load applied to the pile, due the proximity of Load (kn) at each level (msd) the strain gauges to the head of the pile and point of application of the load (note that the diameter of the concrete through the casing is 1,230mm and this figure has been used for the assessment at -1.5mSD and at -19.5mSD). The stiffness of the concrete in the pile was estimated from the strain gauge at -1.5mSD for each load increment and the varying stiffness was , ,625 2,834 3,115 6, ,131 2,591 3,954 6,383 6,787 10, ,206 4,488 6,455 9,9,189 13,565 1,825 3,723 6,782 9,369 13,644 13,761 16,879 2,519 5,099 9,125 12,059 16,909 17,070 20,254 3,980 6,953 12,098 15,008 20,014 20,239 23,626 6,200 9,456 15,140 18,067 23,185 23,530 27,000 8,586 12,023 18,084 21,059 26,429 26,800 30,005 10,788 14,419 20,731 23,712 29,161 29,638 ground engineering january

5 technical paper used to obtain the loads at all other strain gauges as described below. The stiffness of the concrete during the test period was found to vary with the strain as shown on the graph presented in Figure 6, from a maximum value of 42x10 6 kn/m 2 at 60x10-6 strain to 26x10 6 kn/m 2 at 780x10-6 strain. This is compared to the graph presented in Figure 2.1 of BS 8110 Part 1 as shown on Figure 7. A similar stiffness curve is found if a concrete cube strength of 42N/mm 2 is adopted, setting the concrete material factor to 1.0 for this comparison. The strain gauge data for each load increment at each level was averaged and was converted to a corresponding load, using the varying concrete stiffness as noted above. The formula below was used to obtain the load at each strain gauge. F y = ε ya G E G where: ε y is the average strain reading of the four (or three) strain gauges placed at a distance y from the surface, A G is the cross sectional area of the pile and E G is the gross stiffness of the pile. The pile gross stiffness allows for the pile reinforcement (34 H40s as compression reinforcement) in addition to the pile concrete. The stiffness of the casing has been ignored in this assessment because the strain gauges at -1.5mSD and -19.5mSD are close to the top and bottom of the casing and therefore the casing does not appear to change the pile stiffness. A G E G = (π x / 4 34 x π x / 4) x E c + 34 x π x / 4 x 210 x 10 6 kn/m 2 ie Concrete area + Steel Area Table 6: Measured and design shaft friction in different strata (kn/m 2 ) Jack load (kn) Debonded length London Clay London/Lambeth Clay Lambeth Clay Thanet Sand / msd -1.5 to to to to to , , , , , , , , , Values assumed in design Table 7: Measured pile settlement Total applied load from strain gauge at -1.5m (kn) Shaft friction measured over debonded length (kn) Load applied at pile cut off level (kn) Measured settlement at top of pile (mm) Settlement at pile cut off level (mm) 0 3, , , , , , , , , , , , , , , , , , where: E c is the stiffness of the concrete relative to the strain at each strain gauge. Load transfer along shaft of pile Table 5 summarises the jack load and the load at each level from the strain gauge results. The assessment of the load at each level based on the strain gauge results and the assessment of the concrete stiffness do not give exact figures. The measured loads at -1.5mSD do not correspond precisely to the recorded jack load for the first load increment. This may be because the casing had some effect on the stiffness at this depth for the first load increment. The bottom of the casing was at -19.5mSD and was unlikely to affect the measured strain at -19.5mSD. The comparison between the recorded jack load and the measured results at -1.5mSD are within 1% except for the first load increment. These results are plotted in Figure 8. Average values for the ultimate shaft friction in the various strata at each major load increment are presented in Table 6. The bond stresses in each stratum have been compared with the design values and are discussed as follows: n The bond stress over the debonded length was small, with 7kN/m 2 maximum for the final load increment. n The maximum measured average shaft friction in the London Clay was 161kN/m 2 which was slightly above the 140kN/m 2 limit applied in the design and the delayed concreting of the pile appears to have had a minimal effect. (Alternatively it could be considered that the shaft friction in the London Clay would have been higher if the pile construction had not been delayed.) n The zone between 28 and 35m depth was the lower section of the London Clay and the upper Lambeth Group and the peak shaft friction of 113N/m 2 was slightly less than the 140kN/m 2 limit applied in the design (16% reduction). The lower part of the London Clay often contains sandier bands as noted on the borehole logs and these zones may have deteriorated during the extended construction period. The Upper Lambeth Group was referred to in borehole BH1 as silt which similarly could have reduced the shaft friction. n The shaft friction in the lower Lambeth Clay appeared to be increasing with each load increment and it is not clear if the peak was reached. The maximum recorded value of 176kN/m 2 was more than the 140kN/m 2 limit applied in the design. n The shaft friction in the Thanet Sand appeared to be increasing with each load increment and it was not clear if the peak was been reached. The maximum recorded value of 148kN/m 2 was less than the 200kN/ m 2 limit applied in the design. The base capacity was not mobilised in this test, with 10,788kN recorded as the maximum load applied at some 3m above the toe level of the pile, equivalent to a stress of 9,540kN/m 2. Performance of the Pile The settlements at pile cut off level were calculated based on the average measured strain at 28 ground engineering january 2012

6 Load/ultimate load (%) Settlement/diameter (%) Pile shaft diameter (D S ) = 1.2m Pile base diameter (D B ) = 1.2m Deformation modulus below base (E B ) = 300,000kN/m 2 Young's modulus of concrete (E C ) = 50x10 6 kn/m 2 Friction length coefficient (K E ) = 0.4 Upper pile length carrying no load (L O ) = 0m Pile length transferring load by friction (L F ) = 34.3m Flexibility factor (M S ) = Pile design load (P T ) = 16,680kN Ultimate shaft friction load (U S ) = 18,000kN Ultimate pile base load (U B ) = 45,000kN Figure 9 Fleming analysis of the pile test References 1. Burland J B, Standing J R and Jardine F M: Building response to tunnelling Case studies from construction of the JLE CIRIA Special Publication 200; ICE Specification for Piling and Embedded Retaining Walls (ICE 2007) published by Thomas Telford, London 3. Suckling T P and Eager D: Nonbase Grouted Piled Foundations in Thanet Sand for a Project in East India Dock, London Underground Construction Symposium. London Troughton V M and Platis A: The effects of changes in effective stress on base grouted pile in sand. Proc. Int. Conf. Piling and Deep Foundations, London 1989 pp Chapman T J P, Connolly M L, Nicholson D P, Raison C A and Yeow H C: Advances in understanding of base grouted pile performance in very dense sand. Tunnel Construction and Piling, London 8-10 Sept pp Guidance notes for the design of straight shafted bored piles in London Clay. London District Surveyor s Association Publications No 1, Fleming W G K: A new method for single pile settlement prediction and analysis, Geotechnique 42, No 3, Sept 1992 pp BS 8110 Structural Use of Concrete - Part 1: Code of Practice for Design and Construction British Standards Nicholson D P, Chapman T J P, Morrison P. Pressuremeter proves its worth in London s Docklands. Ground Engineering, March 2002 pp mSD and -19.5mSD, over the debonded length. The values (Table 7) indicated that the pile settlement at cut off level for the design working load of 15,450kN was approximately 6mm, which includes the axial shortening of the pile below pile cut off level. At 1.5 x working load, 23,175kN, the settlement at pile cut off level was approximately 23mm, again including axial shortening of the pile below cut off level. These values were less than the specified maximum values of 12mm and 36mm. It should be noted that these values have been recorded for a pile that was left open for two days between boring and concreting. Ultimate Capacity of the Pile The pile test results were used to carry out load/settlement analysis for the pile to extrapolate to an ultimate capacity at a settlement equivalent to 10% of the pile base diameter, namely 120mm total settlement. The data was analysed in accordance with the Fleming Method (1992) (Figure 9). The input parameters were manipulated to best fit the data from the pile test and a prediction for the ultimate load was estimated from the extrapolated load/settlement curve. The base capacity can only be estimated approximately given the limited load applied to the pile in comparison to the ultimate capacity. For a pile deflection of 10% of the pile diameter (120mm) the ultimate loading on the pile was estimated conservatively to be 43,690kN. Based on a SWL of 15,450kN this provided an overall FOS of 2.8 for the ultimate capacity of the pile. Alternatively, for a FOS of 2.25, the SWL of the pile was 19,400kN. Whilst the base capacity in particular is a very approximate estimate, the figures suggest that the base capacity is about 22,500kN/m 2 at 120mm pile settlement, similar to the 20,000kN/m 2 limit. Conclusions The test confirmed an overall factor of safety in excess of The pile performance was also satisfactory, giving less than 12mm settlement at pile cut off level at working load and less than 36mm settlement at 1.5 x working load. These values were satisfied in spite of leaving the pile open for two days between boring and concreting. The test measured a maximum average shaft friction of 161kN/ m 2 in London Clay. The lower 3m of London Clay and the upper 4m of the Lambeth Beds recorded a reduced shaft friction of 113kN/ m 2. These values are considered to represent peak values. The maximum recorded shaft frictions in the lower Lambeth Group and the Thanet Sand of 176kN/m 2 and 148kN/m 2 respectively are not considered to represent the peak values because the values continued to increase with each load increment. Similarly, the maximum base capacity in Thanet Sand was not mobilised, although extrapolation of the data indicates very approximately that the base capacity is similar to the 20,000 kn/m 2 limit. ground engineering january