Settlement Assessment for the Burj Khalifa, Dubai-HGP-6 (1)

8/10/2019 Settlement Assessment for the Burj Khalifa, Dubai-HGP-6 (1)

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RE-ASSESSMENT OF FOUNDATION SETTLEMENTS FOR THE BURJ KHALIFA,

DUBAI

Gianpiero Russo1

Harry G. Poulos2

John C. Small3

ABSTRACT: This paper deals with the re-assessment of foundation settlements for the Burj

Khalifa Tower in Dubai. The foundation system for the tower is a piled raft, founded on deep

deposits of calcareous rocks.Two computer programs, GARP and NAPRA have been used for the

settlement analyses, and the paper outlines the procedure adopted to re-assess the foundation

settlements, based on a careful interpretation of load tests on trial piles in which the interaction

effects of the pile test setup are allowed for. It then examines the influence of a series of factors on

the computed settlements. In order to obtain reasonable estimates of differential settlements within

the system, it is found desirable to incorporate the effects of the superstructure stiffness which act to

increase the stiffness of the overall foundation system. Values of average and differential

settlements for the piled raft calculated with GARP and NAPRA were found to be in reasonableagreement with measured data on settlements taken near the end of construction of the tower.

Key words: case history, footings and foundations, full-scale tests, piles, rafts, settlement.

s 730

1University of Naples, Italy2Coffey Geotechnics, Sydney, Australia

3Coffey Geotechnics and University of Sydney


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INTRODUCTION

The Burj Khalifa in Dubai was officially opened in January 2010, and at a height of 828m, is

currently the worlds tallest building. The foundation system is a piled raft, a form of foundation

that is being used increasingly to support tall structures where the loads are expected to be

excessively large for a raft alone and where the raft and the piles are able to transfer load to the soil.The foundation design process for this building has been described by Poulos and Bunce (2008).

An important component of the design of a piled raft foundation is the detailed assessment of the

settlement and differential settlement of the foundation system, and their control by optimizing the

size, location and arrangement of the piles, and the raft thickness. Many different methods of

analysis have been devised in order to predict the behaviour of raft and piled raft foundations

(Selvaduri, 1979; Clancy and Randolph, 1993; Poulos, 1994; Ta and Small, 1996; Russo and

Viggiani, 1998; Viggiani, 1998; Hemsley, 1998; Hemsley, 2000), and these range from simple hand

based methods to complex three-dimensional numerical analyses.

In this paper, attention is focussed on two methods that model the raft as an elastic plate and the

piles as interacting non-linear springs. The computer codes implementing these methods aredescribed very briefly and are then applied to the Burj Khalifa, currently the worlds tallest

building, which is founded on a piled raft. The development of the ground modulus values is

described using a combination of field test and laboratory data and the results of pile load tests. The

method of interpreting the pile load test data is discussed, and the importance of allowing for

interaction between the test pile and the surrounding reaction piles in emphasised. The two

programs are then used to compare the computed settlements with available measurements of

foundation settlements, and with the Class A predictions made by the foundation designers and

the peer reviewers.

An important objective of the paper is to explore how pile load test data should be used when

predicting the settlement performance of piled and piled raft foundation systems, and to examine

some factors that may have an important influence on predicted foundation settlements.

COMPUTER ANALYSES

The settlement analyses used in this paper for the Burj Khalifa have employed two computer

programs, GARP and NAPRA, which idealize the piled raft foundation as a plate supported by non-

linear interacting springs. A very brief description of these programs is given below.

Program GARP

The computer program GARP (Geotechnical Analysis of Raft with Piles, Small and Poulos, 2007)

uses a simplified boundary element analysis to compute the behaviour of a piled raft when

subjected to applied vertical loading, moment loading, and free-field vertical soil movements.

The raft is represented by a thin elastic plate and is discretized via the finite element method, using

8- noded elements. The soil is modelled as a layered elastic continuum, and the piles are represented

by elasticplastic or hyperbolic springs, which can interact with each other and with the raft. Pile

pile interactions are incorporated via interaction factors (Poulos and Davis, 1980). Simplifying

approximations are utilized for the raft-pile and pile-raft interactions. Beneath the raft, limiting

values of contact pressure in compression and tension can be specified so that some allowance can

be made for nonlinear raft behaviour. The output of GARP includes the settlement at all nodes of

the raft; the transverse, longitudinal, and torsional bending moments within each element of the raft;the contact pressures below the raft; and the vertical loads on each pile. In its present form, GARP

can consider vertical and moment loadings, but not lateral loadings or torsion.


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Program NAPRAThe computer program NAPRA (Non linear Analysis of Piled Rafts, Russo 1998; Russo &

Viggiani, 1998) computes the behaviour of a raft subjected to any combination of vertical

distributed or concentrated loading and moment loading. The raft is modelled as a two-dimensionalelastic body using the thin plate theory, and utilizing the finite element method, adopting a four or

nine noded rectangular element.

The piles and the soil are modelled by means of interacting linear or non-linear springs. It is

assumed that the interaction between the raft and the soil (the piles) is purely vertical; accordingly,

only the axial stiffness of the springs is required.

The soil is assumed to be a layered elastic continuum. The Boussinesq solution for a point load and

the closed form solution for a rectangular uniformly loaded area at the surface of an elastic half-

space are used to calculate the soil displacements produced by the contact pressure developed at the

interface between the raft and the soil. The layered continuum is solved by means of the

Steinbrenner approximation (Russo, 1998; De Sanctis and Russo, 2002), and as such, invokes thesimple assumption that the stress distribution within an elastic layer is identical with the Boussinesq

distribution for a homogeneous half-space (Russo, 1998).

The interaction factor method is used to model pile to pile interaction and a preliminary boundary

element (BEM) analysis allows calculation of the interaction factors between two piles at various

spacings. Interaction between axially loaded piles beneath the raft and the raft elements is

accounted for via pile-soil interaction factors computed with a preliminary BEM procedure. The

reciprocal theorem is used to maintain that the soil-pile interaction factor is equal to the pile-soil

interaction factor.

A stepwise incremental procedure is used to simulate the non-linear load-settlement relationship of

a single pile, the total load to be applied is subdivided into a number of increments, and the diagonal

terms of the pile-soil flexibility matrix are updated at each step. A computation of the nodal

reactions vector is made at each step, to check for tensile forces between raft and soil and an

iterative procedure is used to make them equal to zero. Basically, this procedure releases the

compatibility of displacements between the raft and the pile-soil system in the node where tensile

forces were detected, although the overall equilibrium is maintained by a re-distribution of forces.

An iterative procedure is needed since after the first run some additional tensile forces may arise in

different nodes. The output of the code is represented by the distribution of the nodal displacements

of the raft and the pile-soil system, the load sharing among the piles and the raft, the bending

moments and the shearin the raft, for each load increment.

Abagnara et al (2011) have compared GARP and NAPRA analyses for a simple case, and have

concluded that both programs give comparable results, but that some of the simplifying

assumptions employed in each program give rise to differences in detail. For example, the

difference in raft settlements may be due to the differences in the details of calculation of the soil

layer stiffness using the Boussinesq/Steinbrenner approach. The difference in plate element types

may also contribute to the differences. For the piled raft, the differences may arise because of

differences in the methods used to compute the single pile stiffness values, the interaction factors

and the pile-raft and raft-pile interactions.

In this paper, attention will be focussed on analyses carried out with NAPRA, although acomparison will also be presented between the GARP and NAPRA analyses.


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SETTLEMENT ASSESSMENT FOR THE BURJ KHALIFA TOWER, DUBAI

Foundation layout

The Burj Khalifa project in Dubai, United Arab Emirates (UAE), comprises a 160 storey high rise

tower, with a podium development around the base of the tower, including a 4-6 storey garage. The

Burj Khalifa is located on a 42000 m2site. The tower is founded on a 3.7m thick raft supported on

194 bored piles, 1.5 m in diameter, extending 47.45m below the base of the raft; podium structures

are founded on a 0.65 m thick raft (increased to 1m at column locations) supported on 750 bored

piles, 0.9 m in diameter, extending 30-35 m below the base of the raft. A plan view of foundation is

shown in Fig. 1.

Figure 1. plan view of the Khalifa Tower foundation system

Ground investigation and site characterization

The investigations involved the drilling of 32 boreholes to a maximum depth of about 90 m below

ground level and 1 borehole to a depth of 140 m under the tower footprint. Standpipe piezometers

were installed to measure the ground water level which was found to be relatively close to the

ground surface, typically at a level of 2.5m DMD. The ranges of measured SPT N values are

summarised in Table 2. There was a tendency for N values to increase with depth, beyond anelevation of about -8m DMD.

Table 2 Summary of Measured SPT Values

Elevation m Range of SPT Values

2.5 to -1 0-40

-1 to -8 50-400

-8 to -14 0-100

-14 to -30 40-200

-30 to -40 100-200-40 to -80 100-400


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The ground conditions at the site comprise a horizontally stratified subsurface profile which is

complex and highly variable in terms of the strata thickness due to the nature of deposition and the

prevalent hot arid climatic conditions. The main strata identified were as follows

1. Very loose to medium dense silty sand (Marine Deposits).

2. Weak to moderately weak calcarenite, generally unweathered with fractures close to mediumspaced interbedded with cemented sand. This material is generally underlain by very weak to

weak sandstone which is generally unweathered with fractures close to medium spaced

interbedded with cemented sand.

3. Very weak to weak calcarenite, calcareous sandstone and sandstone; this formation is slightly tohighly weathered with fractures extremely close to closely spaced and interbedded with

cemented sand. Bands of 1 to 5 m thickness are also present of medium dense to very dense,

cemented, sand with sandstone bands and locally with bands of silt.

4. Very weak to weak gypsiferous sandstone, gypsiferous calcareous sandstone occasionallygypsiferous siltstone. This material is generally unweathered to slightly weathered with

fractures extremely close to closely spaced and interbedded with cemented sand. The formation

is interbedded with dense to very dense, cemented, silty sand and occasionally silt withsandstone bands.

5. Very weak to weak calcisiltite, conglomeritic calcisiltite and calcareous calcisiltite. Thismaterial is generally moderately to highly weathered, occasionally slightly and completely

weathered with fractures extremely close to medium spaced. Calcareous siltstone was

encountered in the majority of the deeper boreholes comprising very weak to weak occasionally

moderately weak calcareous siltstone in bands with a thickness of 0.5 to 14.4 m generally

slightly to moderately weathered occasionally highly to extremely weathered.

6. Very weak to weak and occasionally moderately weak calcareous siltstone, calcareousconglomerate, conglomeritic sandstone and limestone. This material is generally slightly

weathered and occasionally unweathered and moderately weathered to highly weathered.

Occasionally encountered as calcisiltite interbedded with bands of siltstone and conglomerate.

7. Very weak to moderately weak claystone interbedded with siltstone. This material is generallyslightly weathered with close to medium spaced fractures. Between -112.2 m and -128.2 m

occasional bands of up to 500 mm thick gypsum were encountered. Below -128.2 m the stratum

was encountered as weak to moderately weak siltstone with medium to widely spaced fractures.

Table 3 summarizes the stratigraphy adopted for the foundation settlement analyses.

In situ and laboratory test results

A comprehensive series of in situ tests was carried out, including pressuremeter tests, down-holeseismic, cross-hole seismic, and cross-hole tomography to determine compression (P) and shear (S)

wave velocities through the ground profile. The vertical profile of P-wave velocity with depth gave

a useful indication of variations in the nature of the strata between the borelogs.

Conventional laboratory classification tests (moisture content of soil and rock, Atterberg limits,

particle size distribution and hydrometer) and laboratory tests for determining physical properties

(porosity tests, intact dry density, specific gravity, particle density) and chemical properties were

carried out. In addition, unconfined compression tests, point load index tests, and drained direct

shear tests were carried out. A considerable amount of more advanced laboratory testing was

undertaken, including stress path triaxial tests, resonant column testing for small-strain shear

modulus, undrained cyclic triaxial tests, cyclic simple shear, and constant normal stiffness (CNS)direct shear tests.


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Table 3.Stratigraphic model adopted for settlement assessment.

Stratum Description Level at the top

of the stratum

[m DMD]

Thickness

[m]

Adopted

Level at

top of

layer

[m DMD]

UCS

qu

[MPa]1 Marine

deposits

1.15 to 2.96 1.85 to 4.3 2.5

2

Calcarenite/

Calcareous

sandstone

-0.27 to -1.95 2.87 to 10.75 -1.2 2

3a Calcareoussandstone/

Sandstone

-4.13 to -12.06 10.5 to 21.43-7.3 -

3b -13.5 1

4 Gypsiferous

sandstone

-21.54 to -

26.69

1.7 to 7.75 -24 2

5a

Calcisiltite/

Conglomeriticcalcisiltite -27.64 to -

31.15

39.2 to 46.75 -28.5 1.3

5b Calcareous

siltstone

-50 1.7

6

Calcareous/

Conglomeritic

Strata

-67.19 to -

76.04

31 (from 140m

deep BH only) -68.5 2.5

7

Claystone/

Siltstone

interbedded

with gypsumlayers

-98.19 Proved to 39.6

m thickness

-90 -

Some of the relevant findings from the in situ and laboratory testing are as follows:

i. The cemented materials were generally very weak to weak; unconfined compressivestrength (UCS) values ranged mostly between about 0.16 MPa the average values for each

layer being the ones reported in the table 3.

ii. Values of the Youngs Modulus from pressuremeter tests (first and second reload cycle)were found to be in good agreement with values from correlation with shear waves

velocities. From calcarenite (0 m to -7.5 m) to sandstone (-7.5 m to -24 m ) Youngs

Modulus is approximately constant with depth; at greater depths the average values decrease

in the gypsiferous sandstone (-24 to -28.5 m) then they slightly increase in the calcisiltite

(from -28.5 to -68.5 m) and finally decrease in the siltstone (from -68.5 to -91 m).

iii. Triaxial Stress Path Testing (at strain levels of 0.01% and 0.1%) was found to give resultsfor Youngs modulus that were in good agreement with pressuremeter and geophysics

testing results.

iv. Resonant Column Testing was found to give more conservative values for the YoungsModulus when compared with values from pressuremeter tests, geophysics tests and triaxial

stress tests.

v. Constant normal stiffness (CNS) tests were carried out on three samples taken from stratum5a to assess the ultimate skin friction values and the potential for cyclic degradation at the

pile-soil interface. These tests indicated values of peak monotonic shear stress ranging from

360 to 558 kPa, with only a little difference between the peak monotonic and the residual

cyclic shear stress values.


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Geotechnical Model

The key parameters for the assessment of the settlement behaviour of the Khalifa Tower piled raft

foundation system are the values of the Youngs modulus of the strata for both raft and pile

behaviour under static loading. In a non-linear analysis, the values of ultimate skin friction of piles,the ultimate end-bearing resistance of the piles, and the ultimate bearing capacity of the raft would

also be required, but in this paper, only linear elastic analyses have been undertaken using NAPRA

and GARP analyses, having explored the little influence of non linearity up to the maximum

observed load level. Attention has thus been focussed on evaluating relevant values of Youngs

modulus for each stratum.

As a first step in obtaining these values, the relative stiffness of the various soil layers was assessed

considering values of the Youngs Modulus from the following data:

1. Pressuremeter tests (initial loading, first reload, second reload cycles);2. Geophysics tests (correlation with shear wave velocities);

3. Resonant column tests (Initial, 0.0001%, 0.001%, 0.01% strain levels);4. Triaxial Stress Path Tests (0.01% and 0.1% strain levels);

Values of the various Youngs Modulusvalues are plotted in Fig. 2, and although inevitable

scatter exists among the different values, there is a reasonably consistent general pattern of

variation with depth.

Layer 3b (see Table 3) has arbitrarily been chosen as the reference layer, and for each type of

test, values of the Youngs Modulusfor a layer i, Ei, have been related to the value for layer 3b,

E3b. The values of Ei/E3bhave then been averaged, using the following data: reload cycles from

pressuremeter testing; seismic data; resonant column data at a strain level of 0.01%, and the

triaxial stress path tests. Fig. 3 shows the different assessed relative stiffness profiles so

obtained, andTable 4 summarises the average values of relative Youngs modulus that were

adopted for the analyses and the interpretation of the pile load test data. The absolute values of

Youngs modulus for each of the different layers have been then obtained by fitting the load

settlement curves of the single piles obtained from the load tests, and the process of fitting the

load-settlement curves to obtain the Youngs modulus values is described below.

Table 4. Relative Values of Youngs Modulus Used in Pile Load Test Interpretation

Stratum Youngs Modulus, Relative to Value for Layer 3b

2

3a

3b

3c

4

5a

5b

6

2.3

0.6

1.0

1.0

0.8

0.7

0.8

0.7


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Fig. 2. Summary of Youngs modulus values.

Fig. 3. Assessed soil relative stiffness.

Pile Load TestsA program of pile load testing was undertaken which involved the installation of seven test piles in

the podium area near the location of the Khalifa Tower. All the test piles and reaction piles were

bored cast in-situ and constructed under polymer fluid. A permanent casing, 6m long, was installed

from the top of each pile to just above the highest strain gauge level for all the trial piles tested in

compression and tension. Five piles, designated as P1, P2, P3, P4 and P5, were tested in

compression; two, P3 and P5, were shaft grouted. Test pile P6 was tested in tension and test pile P7

was laterally loaded.

Only the compression load tests on trial piles P1, P2 and P4 have been considered for the presentpaper. Table 5 summarizes the main features of these piles. Figure 4a shows the load test

arrangements for piles P1 and P2, which consisted of the test pile and six reaction piles, while

-90

-80

-70

-60

-50

-40

-30

-20

-10

0

0 1000 2000 3000 4000 5000 6000 7000 8000 9000 10000

Young's Modulus [MPa]

Depth[m]

Press Init AVE

Press First AVE

Press Second AVE

Seismic (x 0.2)

TRXL 0.01%

TRXL 0.1%

RC Initial

RC 0.0001%

RC 0.001%

RC 0.01%

adopted profile for drained modulus E'

-90

-80

-70

-60

-50

-40

-30

-20

-10

0

0 1 2 3 4

Relative Stiffness E/E3b[-]

Depth[m]

Press Init AVE

Press First AVE

Press Second AVE

Seismic (x 0.2)

TRXL 0.01%

TRXL 0.1%

RC Initial

RC 0.0001%

RC 0.001%

RC 0.01%

ALL

NO INITIAL PRESS

NO INITIAL PRESS & NO TRXL0.01%

Poulos & Davids 2005


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Figure 4b shows the set-up for pile P4, which consisted of the test pile and four reaction piles. Steel

load distribution plates were grouted to the top of the test piles and hydraulic jacks were placed

between the steel plates and the reaction beams. Steel reaction beams were used to transfer the load

from the hydraulic jacks to the installed reaction piles. Macalloy bars were used as reaction anchors

to transfer the load from the beams to the reaction piles. Six cycles of loading were applied to trial

piles P1 and P2 while nine cycles of loading were applied to trial pile P4, which was the piledesignated to be tested cyclically.

Table 5.Summary of pile load tests.

Trial

Pile

Diam.

[m]

Cut-off

level

[m DMD]

Toe

level

[m DMD]

Length

[m]

Load Test

layout

DWL*

[t]

DML**

[t]

No. of cycles

1 1.5 -4.85 -50 45.15

6 RP

circle with a

4.5 m radius

3000 6000 6

(50%-150% DWL)

2 1.5 -4.85 -60 55.15

6 RP

circle with a4.5m radius

3000 6000 6(50%-150% DWL)

4 0.9 -2.90 -50 47.1

4 RP

square with a 9

m side

1000 3500 9

(100%-150% DWL)

*designated working load;** designated maximum test load.

Four main types of instrumentation were used in the compression test piles:

1. Concrete embedment vibrating wire strain gauges, to allow measurement of axial strains atsix levels along the pile shafts and hence estimation of the axial load distribution;

2. Extensometers, to measure change in length of the piles, and installed at the same levels asthe vibrating wire strain gauges to provide back-up information on axial load distribution

with depth;

3. Displacement transducers at the top of piles, to measure the vertical movement at the pileheads.

4. Load cells, to monitor the load applied to the pile via the jacks.

Back-Analysis interpretation of load tests to obtain Youngs Modulus Values

The computer program NAPRA was used to carry out the back-analyses of compression tests on the

three test piles considered. Since a detailed soil profile at each trial pile location was not available,

the same geotechnical model was adopted for all three piles.

For comparison purposes the three load tests were back-analysed both taking and not taking into

account interaction between test piles and reaction piles. It is now well-recognised that ignoring

interaction between the test pile and the reaction piles can lead to an overestimation of the pile head

stiffness (Poulos & Davis, 1980; Poulos, 2000; Kitiyodom et al., 2004).

Both linear elastic (LE) and non-linear analyses (NL) were carried out. In all analyses, Youngs

modulus for the piles, Ep, was assumed to be 31.8 GPa. For the linear analyses, the theoretical

behaviour was fitted to the observed load-settlement behaviour at pile head displacements of about

0.08% of diameter and 0.2 % of diameter.


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In the non-linear analyses, in order to assess the sensitivity of the back-calculated values of soil

stiffness to the value of ultimate capacity, Qlim, three different values were adopted in the analyses:

1) Qlim was estimated as the asymptote to a hyperbola fitted to the whole measuredload-settlement curve (HYP);

2) Qlimwas based on the load transfer deducedby strain gauges readings (SG);3) Qlimwas based on the load transfer deducedby extensometer readings (EX).

Ultimate skin friction values inferred from the axial load distribution and from the extensometer

readings were employed to assess pile shaft capacity up to depths above -30m, -38m and -30m for

piles P1, P2 and P4 respectively. From pilesoil interface load-strain curves at various depths along

the shafts, these values were found to be representative of the ultimate values in the upper (cased)

part of the shaft. Below these depths, ultimate values of shaft friction were estimated from

correlations with the unconfined compressive strength (UCS) of the rock.

Table 6 summarises the values of Qlimobtained from these three procedures. As might be expected,

the hyperbolic extrapolation procedure gives the largest values, and probably over-estimates thecapacity. There is some difference between the values assessed on the basis of the strain gauge and

extensometer readings, but from the point of view of settlement prediction, such differences are not

very significant.

Figures 5 and 6 show typical fits (for Pile P2) of the computed non-linear load-settlement behaviour

and the observed load-settlement behaviour. Figure 5 is for the interpretation taking account of

interaction, while Figure 6 shows the corresponding fit with interaction between the test pile and

reaction piles being ignored. In both cases, very reasonable fits are obtained with the measured data.

Table 6. Assessed pile capacity with different methods.

Pile

Qlim

[kN]

Hyperbolic Extrapolation

(HYP)

Strain Gauge Readings

(SG)

Extensometer Readings

(EX)

TP1 108,800 93,800 73,200

TP2 115,900 97,300 100,200

TP4 82,600 50,500 59,900

(a)Piles P1 and P2. (b) Pile P4

Fig. 4 Set-up for pile load tests


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Fig. 5Predicted and measured loadsettlement Fig. 6Predicted and measured loadsettlement

for pile P2(interaction considered). for pile P2 (interaction ignored).

Back-calculated values of the Youngs Modulusfor stratum 3b, E3b, are reported in Table 7. In the

linear elastic analyses the first point on the measured load-settlement curve has been considered. Inthis way back-calculated values of soil stiffness in linear analyses are affected by the loading

procedure adopted in the load tests. In the cases of piles P2 and P4, values of back-calculated soil

stiffness are in close agreement with values back-calculated in the non linear analysis (values of

w/D are 0.0008 - 0.0009) while in the case of pile P1, the first point is at a higher displacement

(0.21%), and so the back-calculated value is lower. It should be noted that had the interaction

between the test pile and the reaction piles not been taken into account, the back-calculated values

of pile-soil relative stiffness would have been considerably larger.

Table 7.Youngs Modulusvalues derived from load tests.

TESTPILE

E3b[MPa] with interaction accounted for E3b[MPa] with interaction not accounted for

Linear Analysis(w/D=0.0008)

Linear Analysis(w/D=0.002)

Non-linear Analysis Linear Analysis(w/D=0.0008)

Non-linear Analysis

P1- 350 650 (HYP)-850 (SG) - 900 (HYP)-1100 (SG)

P2 700 6501000 (HYP)-1200 (SG-

EX)1200

1500 (HYP)-1700 (SG-

EX)

P4850 550 650(EX)-850(SG) 1100 850(EX)-1100(SG)

Notes: HYP denotes values derived from hyperbolic extrapolationEX denotes values derived from extensometer readingsSG denotes values derived via strain gauges

From Table 7, the following points can be noted:

1. The consideration of interaction between the test pile and the reaction piles results inbackfigured modulus values which are considerably less than those for which interaction has

been ignored. Thus, there would be a tendency to under-estimate foundation settlements if

interaction effects are ignored.

2. The back-calculated values from the three tests are scattered.

In order to partially overcome the described limitation in the back-analysis of load tests, and to

show its effects on the average settlement assessment, sensitivity analyses have been carried out

with NAPRA by adopting two different values of soil stiffness assessed to be representative oflower and upper bound values.

0

5

10

15

20

0 10000 20000 30000 40000 50000 60000 70000

Q [kN]

w

[mm]

measured

E3b=1000 Mpa (NL)HYP

E3b=1200 Mpa (NL)SG

E3b=1200 Mpa (NL)EX

0

5

10

15

20

0 10000 20000 30000 40000 50000 60000 70000

Q [kN]

w

[mm

]

measured

E3b=1500 Mpa (NL)HYP

E3b=1700 Mpa (NL)SG

E3b=1700 Mpa (NL)EX


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In the GARP and NAPRA analyses described below, for the assessment of the average and

differential settlements, the values of E3bshown in Table 8 were adopted, on the basis of the non-

linear analysis of the load test results.

Table 8. Values of Youngs Modulus (E3b) of Layer 3b Adopted for Foundation Analyses

CaseYoungs Modulus of Layer 3b (E3b) MPa

Best Estimate Upper Bound Value Lower Bound Value

Reaction pile

interaction considered900 1000 650

Reaction pile

interaction ignored1200 1500 900

PROCEDURE FOR FOUNDATION SETTLEMENT RE-ASSESSMENT

The majority of the foundation settlement re-assessment was carried out using linear elasticanalyses with the computer program NAPRA. The mesh adopted for the NAPRA analyses is shown

in Figure 7, and in this mesh, the columns were spaced 1.7 m apart. Preliminary analyses indicated

that using a finer mesh than this produced no change in the results. The actual shape of the raft was

modelled by adopting a piecewise approximation.

Only long-term conditions have been considered, and for the majority of the early analyses, an

average load per pile of 23.21 MN has been used (this is representative of the design dead plus live

loading) and has been applied as a point load on each of the 194 piles. This load corresponds to a

uniformly distributed load on the tower raft of about 1250 kPa.

The majority of analyses were undertaken using the best-estimate modulus value of 900 MPa

derived from the proper interpretation of the load test data (see Table 8).

A series of sensitivity analyses was undertaken to examine the following issues:

1. The influence of non-linear pile response;2. The influence of the range of back-figured values of Youngs modulus;3. The influence of using correct and incorrect back-figured values of Youngs modulus;4. The effect of not considering the raft in the foundation settlement analysis;5. The differences between analyses using NAPRA and GARP;

6. The influence of the assumed loading pattern;7. The effect of incorporating the stiffness of the superstructure;8. The effect of including the podium loading.

The results of these sensitivity studies are described below.


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Fig.7 Model adopted in NAPRA analyses

RESULTS OF SENSITIVITY STUDIES

Influence of Non-Linear Pile Response

For the non-linear NAPRA analyses, the ultimate axial capacity of each pile has been assumed to be

112.5MN. Table 9 compares the computed maximum (S max) and central settlements (S centre),

and the maximum differential settlement (DS max), from the linear and non-linear analyses. There

is very little difference between the two analyses in this case, as it could be expected being the

global safety factor on each pile in the range 3 to 5. Thus the comparison indicates that the

foundation response is essentially elastic under the dead plus live loadings. Accordingly, only linear

analyses have been employed for the remainder of the sensitivity studies.

Table 9. Computed Settlements (mm) from Linear and Non-Linear Analyses

Linear Analysis Non-Linear Analysis

S max S centre DS max S max S centre DS max

52 51 27 53 53 27

The influence of the range of back-figured values of Youngs modulus

Table 10 summarises the computed settlements from the NAPRA analysis, using the range of

values of Youngs modulus back-figured from the correct interpretation of the pile load tests (see

Table 8). As would be expected, the computed settlement for the lower-bound modulus value is

considerably greater than that for the upper-bound modulus value, although the ratio of the

computed settlement values is less than the ratio of the modulus values. This may be explained by

the non linearity source provided by the iterative check on the tensile forces at the raft-soil

interface.


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Table 10. Influence of Using Upper and Lower Bounds of Backfigured Modulus Values

Modulus Value for Layer 3b S max S centre DS max

Lower bound (E3b= 650 MPa) 81 68 50

Upper bound (E3b=1000 MPa) 56 46 35

The influence of using correct and incorrect back-figured values of Youngs modulus

Table 11 shows the influence on the computed settlements of using the best-estimate modulus

values for Layer 3b obtained from the correct interpretation (considering test pile-reaction pile

interaction) and the incorrect interpretation (ignoring this interaction). The settlements computed

using the incorrect modulus interpretation are about 21% less than those using the correctinterpretation, and it is therefore important to properly interpret the test pile load-settlement data to

avoid the tendency to under-estimate the foundation settlements and differential settlements.

Table 11. Influence of Modulus Value on Computed Settlements

Modulus Value Used

Computed Settlements

mm

S max S centre DS max

Correct Interpretation

E3b= 900 MPa

52 51 27

Incorrect Interpretation

E3b=1200 MPa41 40 22

Effect of Not Considering the Raft in the Analysis

NAPRA has been used to analyse the foundation system, both as a piled raft, and as a pile group inwhich there is no raft joining the piles. The correct best-estimate modulus of Layer 3b of 900

MPa has been used. Table 12 shows the computed settlements for both these cases. The difference

between the computed central settlements is negligible, but there is a considerable difference

between the computed maximum settlements and differential settlements. In this case, the

conservatism introduced by ignoring the raft would lead to a 17% increase in the computed

maximum settlement but a 40% increase in the maximum differential settlement. Therefore it is

desirable to incorporate the effect of the raft when computing the settlement distribution within the

foundation system.


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Table 12. Influence of Ignoring Pile Cap on Computed Settlements

Case Analysed Computed Settlements mm

S max S centre DS max

Pile Group (no raft)61 51 38

Piled Raft52 51 27

Analyses Using NAPRA and GARP

Similar meshes have been used for both the NAPRA and GARP analyses and identical analysis

assumptions have been made in both cases. Figure 8 shows the computed profile of settlement along

Wing C from both analyses, and reveals that they are almost identical. Thus, pleasingly, for the

same input, each program is capable of giving very similar results.

Figure 8 Calculated settlements along wing C from NAPRA and GARP

The Influence of the Loading Pattern

The preceding results have all been obtained assuming that the average design load (23.21 MN) has

been applied to each pile location. In reality, the loads will be applied via wall and column

locations, and consequently, NAPRA has been used to examine the influence of the loading pattern

on the computed settlement profile for two cases:

a. Equal loads on all the piles;b. The actual design loadings are applied at the wall and column locations.

The computed settlement profiles along Wing C in Figure 9 show a difference in the computed

settlement patterns, with the equal load assumption giving smaller maximum settlement than the

other case. Thus, it would appear desirable to employ the actual load pattern in design calculations.

0

10

20

30

40

50

60

0 10 20 30 40 50 60 70

Distance along wing cross section [m]

w

[mm]

GARP AV 23210 kN NAPRA AV 23210 kN


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Figure 9 Influence of Assumed Loading Pattern on Computed Settlement Profile

The Influence of the Podium Structure on the Tower Settlement

The podium structure was assumed to be founded at the same depth of the tower raft, but the two

rafts were assumed to be unconnected. The length of the podium piles was taken to be 30m, and the

columns and rows in the NAPRA mesh were spaced at 3m up to a distance of 30 diameters and 4 m

at larger distances. An average load per pile of 23.21 MN was applied as a point load on each of the

194 tower piles while the loads acting on the low-rise area were modelled as point loads of between

2 and 8 MN acting on the podium piles, depending on their location.

The maximum computed settlement was 54mm which was only 2mm larger than the valuecomputed for the tower only. The effect of the 750 piles of the podium was thus very small in this

case, primarily because of the significant distance of many of the podium piles from the tower, and

the relatively small loads that they carried.

The Influence of Superstructure Stiffness

In order to investigate the effect on the computed settlement and differential settlement, and to try

and obtain a more accurate estimate of the pattern of settlement, the stiffening effect exerted by the

superstructure on the raft was taken into account, in various ways, by increasing the bending

stiffness of the raft in each wing (estimated by the structural designers to be equivalent to an

increase of 25200 kNm2per wing). Six alternative methods of incorporating this increased bendingstiffness were adopted:

a. Increasing the thickness of the whole raft to reflect the bending stiffness of the entire tower(Model 1).

b. Increasing the raft thickness over the central part of the wings and on the core, as shown inFigure 10, to reflect the bending stiffness of the entire tower. This is denoted as Model 2.

c. Increasing the raft thickness only below the shear walls (see Figure 11), to reflect thebending stiffness of the entire tower; this case is denoted as Model 3.

d. Model 1, with only 10% of the stiffness of the tower considered (Model 1M).e. Model 2, with only 10% of the stiffness of the tower considered (Model 2M).f. Model 3, with only 10% of the stiffness of the tower considered (Model 3M).

0

10

20

30

40

50

60

70

80

0 10 20 30 40 50 60 70

w

[mm]

d [m]

Averagepile loadsapplied

Loadsapplied atwall &columnlocations


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In each case, the actual pattern of loading via the columns and walls was applied, with only the dead

load component considered.

Figure 12 compares the various computed profiles of settlement across the tower, together with

those in which no account is taken of superstructure stiffness. Also shown is the design profile

developed by Poulos and Bunce (2008), which was for combined dead plus live load, and thereforenot directly comparable. Clearly, there is a considerable difference between the extreme profiles,

those taking all the superstructure stiffness into account, and that in which no account is taken of the

superstructure stiffness. It would appear reasonable to assume that the profiles from Models 1M,

2M and 3M may be more reasonable approximations to reality, and this appears to be borne out by

the comparisons with the measured settlements, as described below.

Fig.10. Raft model 2. Fig.11. Raft Model 3.

Figure 12 Comparison Between Various Calculated Settlement Profiles

Comparisons Between Calculated and Measured Settlements

Detailed settlement measurements were only available up to February 2008, before the end of

construction and well before the commissioning and occupation of the building in January 2010.Nevertheless, anectodal evidence indicated that the additional settlements between February 2008

and January 2010 were relatively small, of the order of 1-2mm.

0

10

20

30

40

50

60

70

80

90

0 20 40 60

Settlement[mm]

Distance along wing [m]

NAPRA Model 1

NAPRA Model 2

NAPRA Model 3

NAPRA Model 1M

NAPRA Model 2M

NAPRA Model 3M

NAPRA-No Structure StiffnessAllowanceDesign Values (Poulos & Bunce(2008)


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Figure 13 shows comparisons between the latest available measured profile of settlement in

February 2008, and the calculated settlement profiles from Models 1M, 2M and 3M. The following

observations are made from an examination of Figures 12 and 13:

1. Without allowance for superstructure stiffness, the calculated maximum final differentialsettlement is about 35mm which is considerably larger than the measured value of about

14mm. The computed maximum settlement is also larger than the measured value, although

some additional settlement would be expected after the building has been in operation for

some years.

2. When allowance is made for the superstructure stiffness, the computed maximum settlementis similar in magnitude to the measured value. However, for Models 1, 2 and 3, in which the

full superstructure stiffness is incorporated (albeit approximately), the computed settlement

pattern differs somewhat from the measured pattern, and the computed differential

settlements are significantly smaller than those measured. It seems clear that it is not

appropriate to allow for the bending stiffness of the entire structure when trying to modify

the foundation stiffness.3. When the allowance for superstructure stiffness is reduced by a factor of 10 (Models 1M,

2M and 3M), there is better agreement between the computed and measured profiles, with a

computed maximum differential settlement ranging between 15 and 21 mm for the three

models, similar to the measured value. In this case, the stiffness of the raft is approximately

53 times its original value, and this latter value is much larger than the value of 10 times

adopted by Hooper (1973) for the Hyde Park Barracks in London and by Sales et al. (2010)

for the Skyper Building in Frankfurt. Interestingly, and almost certainly coincidentally, the

profile for this case is rather similar to that obtained for the case when the average load is

imposed on each pile.

4. There remain some differences between the measured and computed settlement profiles inthe vicinity of the edge of the wing. There may well be scope to refine the method by which

the superstructure is incorporated into the geotechnical foundation analysis.

5. The calculated settlements from the design phase are considerably greater than thoseobtained from the analyses in this paper. The main reason for these larger settlements is that

the settlements were for both dead and live load acting, and in addition, conservative values

of Youngs modulus were used in these analyses, with a somewhat different distribution of

ground stiffness with depth being adopted in those calculations. Once again, this comparison

emphasises the importance of appropriate selection of the ground stiffness values if accurate

foundation settlement predictions are to be made.


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Figure 13 Measured and Computed Settlement Profiles along Wing C

CONCLUSIONS

1. The case history of the Burj Khalifa Tower in Dubai has been re-assessed for the prediction

of average and differential settlement of the piled raft foundation system. Thecomprehensive ground investigation and pile testing program carried out for this project has

enabled the site to be characterized in some detail. The pile load tests is particular have been

an important factor in enabling reasonable settlement prediction to be made.

2. The ground stiffness or modulus is a key factor in the prediction of foundation settlements.If this is to be derived from pile load test data, then the interpretation of the load-settlement

should take into account interaction effects with the reaction system, otherwise the ground

stiffness is likely to be over-estimated and the foundation settlements subsequently under-

predicted.

3. Sensitivity studies have been carried out to explore the effects of a number of factors onpredicted settlement behaviour of the Burj Khalifa tower. In addition to the ground stiffness

or modulus, the consideration of the effects of the raft and superstructure stiffness may be

important factors in influencing both the maximum settlement and the maximum differential

settlement.

4. The assessment of the differential settlementof the Khalifa tower foundation has been byexplored by adopting three different models to account for the stiffening effect of the

superstructure, and they have been found to give reasonably similar results. However, if the

foundation, or parts of it, are stiffened to represent the bending stiffness of the entire

structure, the consequent foundation response is too rigid, and the differential settlements

tend to be under-predicted considerably. For the Burj Khalifa tower, an additional stiffnessequivalent to about 10% of the entire bending stiffness has been found to give improved, but

0

10

20

30

40

50

60

0 20 40 60

Settlement[mm]

Distance along wing [m]

NAPRA Model 1M

NAPRA Model 2M

NAPRA Model 3M

Measured (February 2008)


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20 11 July 2011

by no means perfect, results when compared with measured settlement profiles. This result

suggests that there is a limit to the stiffness that the structure can provide to the foundation

system, and so the full structure should not be taken into account when calculating the

effective increase of stiffness.

5. In this case at least, consideration of the low-rise podium structure leads to only a smallincrease in the settlement under the tower footprint.

6. The method of analysis may be a less significant factor in the prediction of piled raftssettlements, provided the method is sound. For the same input data, the computer programs

GARP and NAPRA produced similar settlementsof the tower.

REFERENCES

Abagnara, V., Poulos, H.G. and Small, J.C. (2012). Comparison of two piled raft analysis programs.

Submitted for 12thAustralia- New Zealand Conf. Geomechanics, Melbourne.

Clancy, P. & Randolph, M.F. 1993. An approximate analysis procedure for piled raft foundations.Int. Journ. For Num. and Anal. Meth. in Geomech, 17(12): 849-869.

De Sanctis, L., Russo, G. and Viggiani, C. 2002. 21.Piled raft on layered soils. Proc. Ninth

International Conference on Piling and Deep Foundations, ___ -___Nice 2002

Hemsley, J.A. 1998.Elastic Analysis of Raft Foundations. Thomas Telford, London.

Hemsley, J.A. (ed) 2000.Design Applications of Raft Foundations. Thomas Telford, London.

Kitiyodom, P., Matsumoto, T. and Kanefusa, N. 2004. Influence of reaction piles on the behaviour

of a test pile in static load testing. Canadian Geotechnical Journal, 41, 408420

Poulos, H.G. & Davis, E.H. 1980.Pile foundation analysis and design. New York, John Wiley.

Poulos, H.G. 1994. An approximate numerical analysis of pile-raft interaction.

Int. Journ. For Num. and Anal. Meth. in Geomech., 18, 73-92.

Poulos, H.G. 2000. Pile testingFrom the designers viewpoint.Statnamic Loading Test 98,

Balkema, Rotterdam, 3-21.

Poulos, H.G. and Bunce, G. (2008). Foundation design for the Burj Dubaithe worlds tallest

building. Proc. 6thInt. Conf. Case Histories in Geot. Eng., Arlington, Virginia, Paper 1.47 CD

volume.

Russo, G. 1998. Numerical analysis of piled rafts.Int. Journ. For Num. and Anal. Meth. In

Geomech., 22, No 6, 477-493.

Russo, G. And Viggiani, C. 1998. 15.Factors controlling soil-structure interaction for piled rafts.

Proc. International Conference on Soil-Structure Interaction in Urban Civil Engineering, Ed. R.

Katzenbach & U. Arslan, Darmstadt,___ - ___.

Sales. M.M., Small, J.C., Poulos, H.G. 2010. Compensated piled rafts in clayey soils: behaviour,measurements, and predictions. Can. Geotech. J.47, 327-345.

Selvaduri, A.P.S. 1979.Elastic Analysis of Soil-Foundation Interaction. Elsevier Publishing Co.,

New York.

Small, J.C. and Poulos, H.G. (2007). Nonlinear analysis of piled raft foundations. Geotech. Special

Publication GSP158, ASCE, CD Volume, GeoDenver 2007.

Ta, L.D. & Small, J.C. 1996. Analysis of piled raft systems in layered soils. Int. Journ. For Num.

and Anal. Meth. inGeomech. 20: 57-72.

Viggiani, C. 1998. Pile groups and piled rafts behaviour.Proc. 3rdInt. Geot. Seminar on Deep

Foundations on Bored and Auger Piles, Ghent, 77-94

Documents

Settlement Assessment for the Burj Khalifa, Dubai-HGP-6 (1)