NORTHWESTERN UNIVERSITY Automated …iti.northwestern.edu/publications/finno/Langousis-2007-Automated...inclinometer data. In addition, it is illustrated how soil stress changes due

NORTHWESTERN UNIVERSITY

Automated Monitoring and Inverse Analysis of a Deep Excavation in Seattle

A THESIS

SUBMITTED TO THE GRADUATE SCHOOL

IN PARTIAL FULLFILLMENT OF THE REQUIREMENTS

For the Degree

MASTER OF SCIENCE

Field of Civil and Environmental Engineering

By

Miltiadis Langousis

EVANSTON, ILLINOIS

March 2007

ii

ABSTRACT

Automated Monitoring and Inverse Analysis of a Deep Excavation in Seattle

Miltiadis Langousis

Performance monitoring of deep excavations typically includes slope inclinometers,

optical surveying of soil deformation, tiltmeters and strain gages. Current monitoring data

collection and processing requires time consuming site visits and manual data reduction

by project engineers. Development of robotic and remote access geotechnical

instrumentation conceptually allows processed data to be made available to project

engineers and contractors in “real time.”

Deep excavation design methods usually employ empirical methods and 2-

dimensional (plane strain) finite element analysis, based on soil characteristics

determined by in-situ and laboratory tests. Inaccurate predictions are often produced

when the soil input parameters are not correctly evaluated and when changes in soil

stress, due to ancillary activities, are not taken into account. A number of case histories

and numerical analysis have also demonstrated that deep excavation deflections are

greatly influenced by excavation sequence and 3-dimensional corner restraint (Finno et

al. 2007).

Inverse analysis methods can be implemented in finite element analyses for deep

excavations support design to help minimize uncertainties associated with the soil

constitutive model parameters. The inverse problem employs iterative algorithms to

update selected soil parameters based on the observed soil response and support system

performance in early excavation stages. The updated soil parameters are applied to

iii

simulations of future excavation stages to provide better soil response predictions.

Effective use of inverse analysis in finite element models requires timely collection of

soil response and excavation support system data to ensure that support design may be

updated, if needed, without causing construction delays.

This thesis focuses on the installation and performance of an automated surveying

system, a relatively new and still developing monitoring technique, at the construction of

the Olive 8 Towers in Seattle, WA. It summarizes the philosophy behind the real time

instrumentation and shows how the total station data can complement the conventional

inclinometer data. In addition, it is illustrated how soil stress changes due to ancillary

activities prior to the excavation and careful design can be used to great effect in finite

element simulations. Lastly, the observed responses from the inclinometers were used

for inverse analysis to find parameters that resulted in a good agreement with observed

data and to evaluate whether the soil stiffness input parameters had been accurately

defined in the design stage of the project.

iv

ACKNOWLEDGEMENTS

First and foremost I would like to thank my advisor Professor Richard Finno for

giving me the chance to study at Northwestern. I want to express him my sincere

gratitude for providing hours of advice, guidance and encouragement during this process.

His patience and support especially at the beginning of my academic career in the States

was remarkable and I have to thank him for that. I still believe though, that the Cubs are

more fun than the White Sox.

I also want to thank the members of my committee, Professor Charles Dowding

and Professor Jose Andrade, for their effort and their valuable suggestions and comments.

The help provided by Dave, Mat and David, from the ITI department, during the

automated monitoring set-up and the processing of the monitoring data was very valuable

and has to be acknowledged.

First of all, I would like to say I really big thank you, or as we would say in Greece

“ΕΥΧΑΡΙΣΤΩ” to all the people I know or I met here at Northwestern, either close or

distant friends of mine. Every interaction with you was really valuable for me, helped me

grow and made me the person that I am today. I truly thank you.

Thank you graduate students Izzo, Taesik (Dda), Cecilia, Young-Hoon (“doc”),

Xuxin, Sandi and the fellow “Aquitards” later known as “Earthworms” for making our

geotech group experience entertaining outside our basement (lab). I have to thank also the

people from café Ambrosia; many litters of coffee were consumed there especially when

I was writing this thesis.

I also would like to specifically thank some people that were really close to me

during these two years and we supported each other during happy or moody days. My

v

friend Wan Jei or “Evil doctor Cho” for the numerous hours we spend together in and out

of the lab; any interaction with him was priceless and made me a better person; and

Francesca or “Fra” who tolerated me the most and supported me especially during the

completion of this thesis. I also want to thank my friends in the States: Sotos, Tassos,

Kostas, Stavros, Passant, Derin, Rennaud and my friends in Greece: Dimitris, Rena,

Leonidas, Christoforos and Spyros. And last but not least I would like to thank Greg

Andrianis (with gold and capital letters); his distant help is greatly appreciated.

Finally, I want to thank the people that helped above all others in this effort, my

father Spyros, my mother Kiki and my brother Andreas. Thank you Andrea for all these

hours we spent talking on the phone solving all of my questions and listening to my

problems. Thank you mom and dad for your help and the sacrifices you made to see your

sons studying at a place so far away from you. I know you are proud for both of us.

vi

Table of Contents

Acknowledgements……………………………………………………………………...iv

Table of Contents………………………………………………………………………..vi

List of Figures…………………………………………………………………………….x

List of Tables……………………………………………………………………….......xiv

1. Introduction…………………………………………………………………………...1

2. 8th and Olive Excavation……………………………………………………...………4

2.1. Olive 8 Project Overview…………………………………………………………...5

2.1.1. Soil Geology and Stratigraphy………………………………………………6

2.1.2. Adjacent Structure Qwest Building………………………………….………8

2.1.3. Excavation Support System………………………………………………...10

2.1.3.1. Pile #6………………………………………………………………...14

2.1.3.2. Pile #13……………………………………………………………….14

2.1.3.3. Pile #18……………………………………………………………….15

2.2. GeoEngineers Finite Element Model……………………………………………...16

2.2.1. Constitutive Model ………………………………………………………...16

2.2.2. Finite Element Mesh………………………………………………………..17

2.2.3. Excavation Sequence Simulation…………………………………………..19

2.2.4. Numerical Model Predictions…………………………………....…………20

2.3. Project Instrumentation……………………………………………………………21

2.3.1. Slope Inclinometers………………………………………………………...22

vii

2.3.2. Automated Optical Surveying……………………………………………...23

2.3.2.1. Prism Location………………………………………………………..26

2.3.2.2. Leica TPS 1100 Professional Total Surveying Station……………….28

2.3.2.3. Instrument Errors……………………………………………………..30

2.3.2.4. Rigid Body Translation and Rotation Corrections……………………36

2.3.2.5. Data Acquisition Software……………………………………………41

2.3.2.6. Remote Communication Methods……………………………………42

2.4. Summary…………………………………………………………………………..44

3. Field Performance…………………………………………………………………...46

3.1. Construction Sequence…………………………………………………………….46

3.2. Wall Deformation………………………………………………………………….49

3.2.1. Slope Inclinometers………………………………………………………...49

3.2.2. Automated Optical Surveying……………………………………………...53

3.2.3. Comparison of Inclinometer and Optical Survey Data…………………….56

3.3. Summary of Performance………………………………………………………….61

3.3.1. Total Station Difficulties During Construction…………………………….63

3.3.2. The Problem of the Rigid Body Rotation Correction………………………63

3.3.3. Summary…………………………….………………………………...……64

3.4. Predicted and Observed Responses.……………………………………………….64

3.5. Summary…………………………………………………………………………..65

4. Alternate Finite Element Model And Inverse Analysis…………………………...68

4.1. Finite Element Model ……………………………………………………………..68

viii

4.1.1. Finite Element Mesh………………………………………………………..69

4.1.2. Constitutive Model and Input Parameters………………………………….72

4.1.3. Excavation Sequence……………………………………………………….77

4.2. Finite Element Results and Analyses……………………………………...………79

4.2.1. Drained Analysis…………………………………………………………...79

4.2.2. Undrained Analysis………………………………………………………...85

4.3. Analysis of Results………………………………………………………………...88

4.3.1. 3-D effects………………………………………………………………….88

4.3.2. Numerical Simulation of Tiebacks…………………………………………90

4.3.3. Summary……………………………………………………………………94

4.4. Inverse Analysis…………………………………………………………………...95

4.4.1. Procedures………………………………………………………………….95

4.4.2. Selection of Field Observations for Inverse Analysis……………………...98

4.4.3. Finite Element Simulations for Optimization……………………………..104

4.4.4. Optimization for Fully Drained Analysis…………………………………105

4.4.4.1. Optimized Parameters……………………………………………….105

4.4.4.2. Results……………………………………………………………….107

4.4.5. Optimization for Undrained Clayey Silt Analysis………………………...119

4.4.5.1. Optimized Parameters……………………………………………….119

4.4.5.2. Results……………………………………………………………….121

4.4.6. Discussion…………………………………………………………………134

4.5. Conclusions………………………………………………………………………135

ix

5. Summary and Conclusions………………………………………………………...138

References……………………………….……………………………………………..142

x

LIST OF FIGURES

Figure 2.1 Picture of the site taken by Google Earth……………………………………...4

Figure 2.2 The construction site at the beginning of the project………………………….5

Figure 2.3 Soil profile defined by boring logs…………………………………………….7

Figure 2.4 Proximity of Qwest building…………………………………………………..9

Figure 2.5 Plan view of the construction with the design sections………………………10

Figure 2.6 Design Sections………………………………………………………………13

Figure 2.7 Finite Element mesh for a typical design section…………………………….18

Figure 2.8 Predictions for the wall deflection given by Plaxis…………………………..21

Figure 2.9 Inclinometer on Pile #6………………………………………………………23

Figure 2.10 Total station on parapet wall………………………..………………………24

Figure 2.11 Location of monitoring prisms on site………………………………………25

Figure 2.12 Points screwed and glued on the wall………………………………………26

Figure 2.13 Location of monitoring points and fixed point #7…………………………..27

Figure 2.14 Location of fixed point #8…………………………………………………..28

Figure 2.15 Image of Leica TPS 1101…………………………………………………...29

Figure 2.16 Instrument axes……………………………………………………………..30

Figure 2.17 Line of sight error…………………………………………………………..31

Figure 2.18 Titling axis error…………………………………………………………….32

Figure 2.19 Vertical axis error…………………………………………………………...33

Figure 2.20 Compensator error…………………………………………………………..34

Figure 2.21 V-Index error………………………………………………………………..35

xi

Figure 2.22 Collimination error………………………………………………………….35

Figure 2.23 Rigid body translation and rotation correction……………………………...38

Figure 2.24 Hardware used for remote communication…………………………………44

Figure 3.1 Deflection for Pile #6………………………………………………………...50

Figure 3.2 Deflection for Pile #13……………………………………………………….51

Figure 3.3 Deflection for Pile #18……………………………………………………….52

Figure 3.4 Data acquired by automated surveying at Pile #6……………………………54

Figure 3.5 Displacement along the wall of the monitored points………………………..55

Figure 3.6 Comparison of inclinometer and optical survey movement towards and along

the excavation for Pile #6………………………………………………………………..58


the excavation for Pile #13………………………………………………………………59


the excavation for Pile #18………………………………………………………………60

Figure 3.9 Data lost when the instrument went out of balance………………………….62

Figure 3.11 Error in calculating the settlement when the vertical compensator was

OFF………………………………………………………………………………………64

Figure 3.12 Comparison of predicted versus observed wall deflection………………….66

Figure 4.1 Finite element mesh for typical design section………………………………71

Figure 4.2 Relative shear stress before applying the changes in soil stress history……...80

Figure 4.3 Shear stress levels (psf) after applying the changes in soil stress history……81

Figure 4-4 This study, compared to the GeoEngineers predictions and the observed

xii

deflections for Pile #6……………………………………………………………………82


deflections for Pile #13…………………………………………………………………..83


deflections for Pile #18…………………………………………………………………..84

Figure 4.7 Comparison of drained and undrained finite element analysis in contrast with

the observed and predicted displacement profiles……………………………………….87

Figure 4.8 Effects of plan dimensions on PSR (after Finno et al, 2007)………………...89

Figure 4.9. Effects of plan dimensions and depth of excavation on PSR………………..90

Figure 4.10 Deflection profiles for tiebacks and equivalent struts………………………93

Figure 4.11 Observations vs. instrument error for conventional and in-place inclinometer

at Pile #6. ………………………………………………………………………………100

Figure 4-12. Comparison of inclinometer errors……………………………………….101

Figure 4.13 Observed and stated by manufacturer instrument error…………………...102

Figure 4.14 Observed versus the calculated and optimized deflection for three optimized

parameters………………………………………………………………………………107

Figure 4.15 Objective function, RFI (%) and stiffness parameters versus the number of

iterations………………………………………………………………………………...109

Figure 4.16 Observed versus the calculated and optimized deflection for a three

optimized parameters…………………………………………………………………...111

Figure 4.17 Objective function and stiffness parameters versus the number of

iterations….……………………………………………………………………………..113

xiii

Figure 4.18 Optimization results and wall deflection versus number of optimized

parameters………………………………………………………………………………114

Figure 4.19 Shear strains (%) in function with depth and distance from the West

excavation wall………………………….……………………………………………...118


parameters………………………………………………………………………………121

Figure 4.21 Objective function RFI (%) and stiffness parameters versus the number of

iterations………………………………………………………………………………...124




iterations………………………………………………………………………………...127

Figure 4.24 Observed versus calibrated displacement deflection profiles for two and three


Figure 4.25 two parameter optimization results for drained and undrained Clayey Silt

analysis………………………………………………………………………………….131

xiv

LIST OF TABLES

Table 3.1 Excavation sequence during construction for Pile #6…………………………46

Table 3.2 Excavation sequence during construction for Pile #13………………………..47

Table 3.3 Excavation sequence during construction for Pile #18………………………..48

Table 3.4. Final movement recorded for monitored points……………………………...56

Table 4.1 Soil Parameters used for design……………………………………………….74

Table 4-2 Comparison of K0 values from GeoEngineers and this analysis……………..74

Table 4.3 Properties of concrete elements……………………………………………….75

Table 4.4 Stiffness of structural elements………………………………………………..76

Table 4-5. Pre-stress load for West wall tiebacks………………………………………..78

Table 4-6. Pre-stress load for East wall tiebacks………………………………………...78

Table 4-7 Representation of Tiebacks by equivalent struts……………………………...92

Table 4.8 Observations used for inverse analysis………………………………………104

Table 4-9. Results of sensitivity analysis……………………………………………….105

Table 4.10 Optimization for three parameters………………………………………….110

Table 4.11 Optimization for two parameters…………………………………………...113

Table 4.12 Overview of Stiffness Input Parameters……………………………………115

Table 4-13. Results of sensitivity analysis……………………………………………...119

Table 4.14 Optimization for three parameters………………………………………….122

Table 4.15 Optimization for two parameters…………………………………………...126

Table 4.16 Overview of Stiffness Input Parameters……………………………………129

Table 4.17 Overview of Stiffness Parameters…………………………………………..133

Chapter 1. Introduction

1

Performance monitoring of deep excavations typically includes slope inclinometers,

optical surveying of soil deformation, tiltmeters and strain gages. Current monitoring data

collection and processing requires time consuming site visits and manual data reduction by

project engineers. Development of robotic and remote access geotechnical instrumentation

conceptually allows processed data to be made available to project engineers and

contractors in “real time.”

Deep excavation design methods usually employ empirical methods and 2-

dimensional (plane strain) finite element analysis, based on soil characteristics determined

by in-situ and laboratory tests. Inaccurate predictions are often produced when the soil

input parameters are not correctly evaluated and when changes in soil stress, due to

ancillary activities, are not taken into account. A number of case histories and numerical

analysis have also demonstrated that deep excavation deflections are greatly influenced by

excavation sequence and 3-dimensional corner restraint (Finno et al. 2007). Inverse

analysis methods can be implemented in finite element analyses for deep excavations

support design to help minimize uncertainties associated with the soil constitutive model

parameters. The inverse problem employs iterative algorithms to update selected soil

parameters based on the observed soil response and support system performance in early

excavation stages. The updated soil parameters are applied to simulations of future

excavation stages to provide better soil response predictions. Effective use of inverse

analysis in finite element models requires timely collection of soil response and excavation

2

support system data to ensure that support design may be updated, if needed, without

causing construction delays.

The purpose of this study was to test the performance of an automated survey system,

a relatively new and still developing monitoring technique, summarize the philosophy

behind the real time instrumentation and show how the total station data can complement

the conventional inclinometer data. In addition, it will be illustrated how soil stress

changes due to ancillary activities and careful design can be used to great effect in finite

element simulations.

Construction of the Olive 8 Towers in Seattle, WA, provided the opportunity to

install both traditional and developmental geotechnical instrumentation. Traditional

monitoring instrumentation included slope inclinometers for the shoring wall. To

supplement the traditional data, an automated remote-access total survey station was

deployed to monitor the 3-d wall movements.

Several sets of finite element analyses were performed to evaluate the effects of soil

stresses, tieback representations and 3-d corner restraint on the predicted displacements.

Also the observed responses from the inclinometers were used for inverse analysis to find

parameters that resulted in good agreement between computed and observed data, and to

evaluate how well the soil stiffness input parameters had been defined during the design

stage of the project.

Chapter 2 presents an overview of the Olive 8 project excavation, with particular

attention to shoring wall design for the west side of the excavation. It includes the lateral

wall deflection predictions made by GeoEngineers Inc., the geotechnical consultant of the

3

project, based on numerical simulations of the excavation sequence. The project

instrumentation program is summarized, including slope inclinometers and automated

optical surveying for the shoring wall and the adjacent structures.

Chapter 3 presents the performance monitoring data collected from the

instrumentation program during the excavation. The observed west shoring wall

deflections are summarized and compared to the predictions given by the design finite

element simulation employed in design. The automated monitoring data is compared with

the inclinometer observations at the top of the west shoring wall. Finally, the difficulties

with application of the automated surveying are discussed and the corrections which was

used to process the survey data are presented.

In Chapter 4, a finite element simulation is presented that explicitly accounts for the

changes in stresses history caused by construction of the Qwest building. Stiffening at the

corners of the excavation due to 3-D effects and representation of tieback restraint in the

finite element simulation are also discussed. Results of inverse analysis of this numerical

model are presented, as are the optimized parameters that results in the best fit between the

calculated and observed deflection profiles.

Chapter 5 summarizes this thesis and presents its conclusions.

4 Chapter 2. 8th and Olive Excavation

The 8th and Olive development is a project in downtown Seattle that includes a deep

excavation to accommodate the underground parking facilities. Figure 2.1 shows the

location of the project site from Google Earth. Finite element analyses were conducted in

the design phase of this project because of the complexity of the shoring system. The city

of Seattle imposed a design constraint that the soil would not move more than 1 inch

during the excavation. The movement limitations were checked during the excavation by

instrumentation which included conventional inclinometers and optical survey points and

a robot total station, which is not widely used in practice.

Figure 2.1 Picture of the site taken by Google Earth

5

2.1. Olive 8 Project Overview

The Olive 8 tower development occupies one half of a city block bordered by Eight

Avenue on East, Olive Way on the north and an alley on the west in downtown Seattle,

Washington. The development consists of 33 stories above grade and 5 parking levels

below grade. The excavation for the underground facilities extended as much to 71 ft.

below the ground level. The urban location of the construction and the surrounding

existing buildings required a deep supported excavation. In addition, the proximity of the

Qwest building west of the excavation demanded a rather unique design for the adjacent

shoring wall. Figure 2.2 shows the construction site at the beginning of the project during

the installation of the soldier pile wall

North Wall

West Wall

East Wall

Figure 2.2 The construction site at the beginning of the project

6

For the North, South, and East shoring walls the retaining structure consisted of a

temporary soldier pile wall with wooden lagging and 5 or 6 rows of tiebacks. This design

was done by HartCrowser Inc. Their October 2000 report also included a proposed design

for the West Shoring wall.

For the West shoring wall, GeoEngineers Inc, of Redmond, Washington, proposed an

alternative design for the retaining structure on August 2005. The west wall was designed

and constructed according to their recommendations. In addition, to predict more

accurate the wall deflections they performed finite element model analyses with PLAXIS

V8.

In their design, the West shoring wall consists of a soldier pile wall with shotcrete, rather

than wood lagging, 9 to 10 rows of soil nails and 4 or 5 rows of tiebacks.

2.1.1. Site Geology and Stratigraphy

Based on boring logs presented in the HartCrowser Inc. geotechnical report (2000),

the soil generally consisted of silty sand over hard silt and clay underlain by very dense

sand. The soil samples from the boring logs were visually classified in the field and then

taken to the laboratory for further testing and classification under a relatively controlled

environment. The soil stratigraphy as defined in the boring logs is shown in Figure 2.3

and is described below.

Fill. The fill consists of loose to dense sand with variable silt and gravel content. It was

anticipated to be as thick as 10 ft. The design unit weight was γ=125 pcf, the cohesion

7

was c=200 pcf and the friction angle is φ= 38o. The SPT resistance varied from 18 to 25

blows per foot.

Silty-Sand (SM). Below the fill, a layer of dense gravelly silty sand was encountered;

horizontal slickenslides were observed in the soil samples. The density of the sand

increased with depth. The thickness of this layer ranged from 30 to 35 feet. The design

unit weight was γ=130 pcf, the cohesion was c=200 pcf with a friction angle of φ= 34o.

These fractured soils are found also in similar sites in downtown Seattle. The SPT

resistance varied from 30 to 60 blows per foot and the water content from 10 to 18 %.

Figure 2.3 Soil profile defined by boring logs (GeoEngineers Inc., 2005)

8

Clayey Silt (ML). A hard, moist clayey silt layer underlay the silty sand layer. The

thickness of the clayey silt ranged from 25 to 30 ft. The layer contained some hard lenses

and some occasional horizontal slickensides. The design unit weight was γ=125 pcf, the

cohesion was c=200 pcf and the friction angle was φ= 38o. The SPT resistance varied

from 25 to 30 blows per foot and the water content from 18 to 30 %.

Very Dense Sand (SP). The bottom of the boring showed very dense wet, gravely sand

underlying the clayey silt. The design unit weight was γ=140 pcf, the cohesion was c=0

pcf and the friction angle was φ= 40o. The SPT resistance was at least 60 blows per foot

and the water content was less than 10 %.

At the time of the drilling, ground water was not encountered in the boreholes.

2.1.2. Adjacent Structure-Qwest Building

The existing Qwest building, as shown in Figure 2.4, is located to the west of the

Olive 8 excavation. A 16 foot wide alley separates the west temporary support wall from

the Qwest building. The Qwest building has several below grade levels that extend to

approximately 52 feet below the ground surface. The exterior walls are supported on a

perimeter 7-foot-thick and 10-foot-wide strip footing. The core of the building rests on a

7-foot-thick, 90 ft x 90 ft wide mat foundation. The design bearing pressure for the Qwest

building perimeter footing and mat foundation is 10 ksf.

9

Qwest Building

Alley

Utility Vault

Figure 2.4 Proximity of Qwest building

In the alley between the Qwest building and the excavation, there are buried power and

communication utilities with 5 utility vaults. The utility vaults are located at the central

part of the west excavation wall and extend 22 feet below the ground surface.

Consequently, 3 separate sections were designed to provide lateral support for the west

wall. Figure 2.5 presents the plan view of the construction including the three design

sections.

10

Figure 2.5 Plan view of the construction with the design sections

Due to the proximity of the Qwest building, the rather conventional soldier pile and

tieback walls used for the North, South and East shoring walls were not feasible. The

shoring system for the West Wall designed by GeoEngineers consisted of soldier piles

with shotcrete lagging and soil nails in its upper reaches, with tiebacks with wood lagging

in it lower reaches.

2.1.3. Excavation Support System

This thesis focuses on the design and performance of the West shoring wall. The

three design sections for this wall are shown in Figure 2.6.

11

Soldier Piles. The soldier piles consist of W27x146 sections set in a drilled shaft which

was filled with concrete following the placement of the beam. The spacing of the soldier

piles was 8 feet, the same as that of the soldier piles used for the Qwest excavation.

Soil Nails. The soil between the Qwest building and the excavation was retained with

several rows of soil nails angled 15 degrees below the horizontal (the number of rows

depend on the wall section). The soil nails consisted of #10 rebar grouted in an 8-inch

diameter hole and are spaced at 3 feet vertically and 4 feet horizontally. Depending on the

wall section, an additional soil nail was added at an elevation approximately 30 feet

below the ground surface at the center of the 3x4 feet pattern. The length of the soil nails

was the maximum possible. The drilling stopped when the drill hit the Qwest basement

wall. The drilling for the soil nail was performed with a rotary auger, and with

subsequent grouting resulted in a load transfer of 3 kips/foot without the use of post

grouting.

Tiebacks. Below the elevation of the soil nails, tieback anchors were used to retain the

wall. The tiebacks were placed as high as possible, while maintaining a minimum 3 foot

clearance below the perimeter strip footing of the Qwest building. The maximum

declination of the tiebacks was 45 degrees below the horizontal. The tieback horizontal

spacing is 8 feet and 4 to 5 rows of tiebacks were deployed, depending on the wall

section. The bonded and unbonded length of the tiebacks is approximately 30 and 25 feet,

respectively. The maximum allowable load transfer design load was 4.5 kips/ft. The

12

drilling was performed with a rotary drill and the anchors were grouted under high

pressure with no regrouting.

As previously mentioned, in the alley between the excavation and the Qwest

building there are buried power and communication utilities in 5 utility vaults. To reduce

the movements and make the soil around the larger utility vault stiffer, grout was injected

at the sides and between of the utility vault and excavation wall after the soldier pile wall

was installed, as shown by the cross hatched area in section Pile #13.

13

Soil Nails

Tiebacks

Figure 2.6 Design Sections

13

14

2.1.3.1. Pile #6 (Design section 1)

The design section 1 refers to the wall section around pile #6 and it is shown in

Figure 2.6a. The top elevation of the fill is at +133.5 ft from the sea level. The beams

used for the soldier pile wall were 83-feet-long, W27x146 sections. Because of presence

of utility vaults, the first 19 feet of the wall are cantilever. Below the cantilevered part,

the wall is retained by 9 rows of soil nails spaced with a 3x4 feet pattern (vertical x

horizontal). After the 4th row of soil nails an extra soil nail was inserted at the center of

the 3x4 foot pattern. Tieback anchors were placed below the soil nails. For section 1, 4

rows of tiebacks are used, spaced 8 feet horizontally and 5 feet vertically. The typical un-

bonded length was 25 feet with a grouted length of approximately 30 feet. The bottom of

the excavation is at +71.0 feet. The wall extends about 20 feet below the bottom of the

excavation. Hereafter, Section 1 will be referred to as Pile # 6.


Design section 2 refers to the wall sections around pile #13 and includes the large

utility vault as shown in Figure 2.6b. The top elevation of the fill is at +131.5 ft over the

sea level. The beams used for the soldier pile wall were 86-feet-long, W27x146 sections.

Due to the size of the utility vault, the first 25 feet of the wall are cantilevered. Below the

cantilevered part, the wall is retained by 5 rows of soil nails spaced with a 3x4 feet

pattern (vertical x horizontal), with an additional soil nail at the center of the 3x4 foot

pattern. Tieback anchors were placed below the soil nails. For section 2, 5 rows of

15

tiebacks were used, spaced 8 feet horizontally and approximately 5 feet vertically. The

typical un-bonded length was 25 feet and the grouted length was approximately 30 feet.

The bottom of the excavation is at elevation +60.0 feet, with the wall extending 15 feet

below this elevation. It needs to be mentioned; that the strength of the soil surrounding

the utility vault in the alley, was improved by grout injections. Hereafter, Section 2 will

be called as Pile # 13.


The design section 3 refers to the wall section around pile #18 and is shown in

Figure 2.6c. The top elevation of the fill is at elevation +129.5 ft over the sea level. The

beams used for the soldier pile wall were 83 feet long W27x146 sections. Because of the

presence of the utility vault the first 18 feet of the wall was cantilever. Below the

cantilevered part, the wall is retained by 9 rows of soil nails spaced with a 3x4 feet

pattern (vertical x horizontal). After the 4th row of soil nails an extra soil nail was inserted

at the center of the 3x4 foot pattern. Tieback anchors follow the soil nails. For section 3,

5 rows of tiebacks are used, spaced 8 feet horizontally and 5 feet vertically. The typical

un-bonded length was 25 feet with a grouted length of approximately 30 feet. The bottom

of the excavation is at +60.0 feet. The wall extends about 15 feet below the bottom of the

excavation. Hereafter, Section 1 will be referred to as Pile # 18.

16

2.2. GeoEngineers Finite Element Model

GeoEngineers numerically simulated the excavation with the finite element

program PLAXIS V8. This program is a two dimensional (2D) finite element code

written for soil and rock mechanics analysis. One can represent the soil stratigraphy and

the structural elements as well as responses of soil and structural elements, while

numerically simulating the construction process. Since PLAXIS V8 is a 2D program, the

shoring system is modeled in plane strain conditions, where an equivalent 1-foot wide

section of the wall is analyzed with corresponding structural properties scaled by the

horizontal spacing of the structural elements.

2.2.1. Constitutive Model

The stress-strain behavior of the soil layers described in section 2.1.1 was

represented in PLAXIS V8 by the hardening-soil model (PLAXIS Manual, 2002). The

Hardening-Soil includes soil dilatancy and a volumetric yield surface that isotropically

expands due to plastic straining. Shear hardening is also included to model irreversible

plastic strain due to primary deviatoric loading.

Some features of the Hardening-Soil model are:

• Stress dependent stiffness according to a power law, defined by the power of

stress-level dependency of stiffness, m.

• Plastic straining due to primary deviatoric loading, defined by the secant stiffness

in standard drained triaxial test, E50ref.

17

• Plastic straining due to primary compression, defined by the tangent stiffness for

primary oedometer loading, Eoed,ref.

• Elastic unloading reloading, defined by the unloading and reloading stiffness

Eur,ref and the Poisson’s ratio for unloading and reloading, νur.

• Failure according to the Mohr-Coulomb model, defined by the cohesion c, the

friction angle φ and the dilatancy angle ψ.

2.2.2. Finite Element Mesh

The finite element mesh (Figure 2.7) used by GeoEngineers to model the

excavation was 134 feet high and 235 feet long. The mesh encompasses the adjacent

Qwest building and almost the entire excavation width. The left (West) and the right

(East) boundaries extend 134 feet and 100 feet away from the west excavation wall,

respectively. It has to be mentioned that the East boundary is not the symmetry axis of

the excavation since the entire width of the excavation is almost 110 ft. This mesh did not

include the west wall of the Qwest building and the east excavation wall. The lower

bound of the mesh was set at the elevation of +0.0 feet. For the mesh, 15-node triangular

elements were used. The mesh was more refined near the wall, the soil nails and the

tiebacks. The fixities specified for the mesh boundaries were zero horizontal and vertical

movements for the lower boundary of the mesh and zero horizontal movements for the

left and right boundaries. In addition, zero horizontal movements were used at the East

Qwest rigid wall.

18

Fill

Sand

Clay

Dense Sand

135 ft 100 ft

+134.0

+119.5

+84.0

+59.5

+0.0

West East

Qwest Building

Olive 8 Excavation

Figure 2.7 Finite Element mesh for a typical design section

19

2.2.3. Excavation Sequence Simulation

The construction sequence was modeled by GeoEngineers in PLAXIS V8 in a way

that approximately represented the effects of the Qwest building, the main excavation

stages, and the soil nail and tieback installations. The Qwest building was inserted in the

geometry after the initial geostatic conditions. The east rigid wall was fixed in the

horizontal direction so as to allow only vertical movement. The displacements were set to

zero after establishing the initial geostatic conditions, applying traffic loads in the alley

and installing the Qwest building.

The modeling sequence in GeoEngineers numerical simulation is:

1) Set up model geometry and calculate initial stress conditions.

2) Excavate the Qwest building and install Qwest basement wall.

3) Apply Qwest foundation loading and install Qwest basement wall

4) Install soldier pile wall

5) Excavate 2 feet below soil nail or tieback level

6) Install soil nail or tieback (including post tensioning of tiebacks)

7) Repeat steps 5 and 6 until the bottom of the excavation is reached

During the numerical simulation of the Olive 8 excavation, a rather uncommon

technique was adopted to promote the rigid block failure of the soil and the soil nails in

the alley. This technique was used only for the Upper Sand and Clayey Silt layers and is

described below.

20

Rankine theory implies that the failure plane of the soil behind a retaining wall is

oriented 45 + φ/2 from the current excavation level. As the Olive 8 excavation proceeded,

the interface friction between the soil and the soil nails was reduced by 60% at locations

higher than the point where the failure plane from each excavation level intersected the

Qwest wall. The use of this technique reduces the contribution of the soil nails to the

shoring wall stiffness and forces the retained soil to separate from the interface of the

Qwest wall. However, this technique is considered conservative because the soil nails

slide easier through the retained soil and this increases the predicted wall deflections.

2.2.4. Numerical model predictions

The results of the finite element analysis, for the 3 design sections are shown in

Figure 2.8. For Pile #6 the simulation predicted 0.8 inches of movement for the top of

the wall and 0.8 inches of deep seated movements at the bottom of the excavation. For

pile #13 the result indicate a deflection of 0.5 inches for the top of the wall and deep sited

movements of 1.2 inches. Finally for pile #18, results suggest 0.5 and 1.5 inches of

deflection for the top and the bottom of the excavation respectively. These predicted

movements exceed the limiting movement of 1 inch, which would result in a work

stoppage while the movements and required support system were revaluated. However,

GeoEngineers were satisfied with this design, and were confident that the 1 inch limits

will not be exceeded because they considered their finite element model very

conservative in terms of the wall’s stiffness, the strength of the retained soil and stress

history induced by Qwest building.

21

Pile #6

Predicted Displacement (in.)

0.0 0.4 0.8 1.2 1.6

Ele

vatio

n (ft

.)

60

70

80

90

100

110

120

130

Pile #13


0.0 0.4 0.8 1.2 1.6

Ele

vatio

n (ft

.)

40

50

60

70

80

90

100

110

120

130

Pile #18


0.0 0.4 0.8 1.2 1.6

Ele

vatio

n (ft

.)

40

50

60

70

80

90

100

110

120

130

Figure 2.8 Predictions for the wall deflection given by Plaxis

2.3. Project Instrumentation

An instrumentation program was established to monitor the performance of the

shoring system and to provide early detection of deflections that could potentially

damage the nearby structures. This program included installing three inclinometers along

the west wall to measure lateral movements, optical surveying of the shoring wall and

nearby structures before and during the construction, and strain gages on the soil nails

22

and load cells on the tiebacks to monitor the load transfer from the ground to the

structural components of the wall.

If the lateral movements were observed to be in excess of ½ inch between two

successive readings or if the total wall movements exceeded 1 inch, the construction of

the shoring wall would be stopped to evaluate the cause of the movement and to establish

the type and extent of remedial measures required. Based on past performance data of

excavations through similar soil conditions, typical deflections for excavations of this

height were expected to vary from 0.001H to 0.003H (¾ inch to about 2.5 inches).

Clough and O’Rourke (1991).

2.3.1. Slope Inclinometers

Three inclinometers were installed along the west wall attached to the back flange

of piles #6, #13 and #18, as shown in Figure 2.9. The casing was securely attached on

the pile every 10 feet and it was filled with water prior placing the concrete into the shaft.

All inclinometers extended to the bottom of the soldier piles. Inclinometer readings were

performed approximately once a week by GeoEngineers. The first reading was completed

prior to drilling adjacent piles.

It has to be mentioned that at this location, the inclinometers observations are

affected by the stiffness of the Soldier Pile and in common practice they should be

installed 2 to 3 feet behind the shoring wall.

23

Figure 2.9 Inclinometer on Pile #6

2.3.2. Automated Optical Surveying

Northwestern University installed an optical surveying system to monitor

automatically 3D deflections of the top of west wall at the positions where the

inclinometers were located. These same prisms were also placed in the alley behind the

wall. A Leica TPS 1101 Total Surveying Station was placed atop the parapet wall of the

rooftop at the Paramount Hotel, which was adjacent to the south excavation wall. To

access the total station, one had to have access to the Hotel rooms that connected to the

rooftop. To mount the instrument on the wall steel brackets were used. These brackets

were “π” shaped and they were tightened around the front and the back of the parapet

24

wall with the use of rubber pads. The brackets are shown in Figure 2.10. Although the

instrument was fixed on the brackets, it could not be considered as fixed since the

brackets were not screwed into the wall, so they were free to slightly translate or rotate.

A sketch of the prism locations is shown in Figure 2.11. Since this total station

could not be considered fixed in space, the prisms P-7 and P-8 were established at

locations far away from the excavation to serve as a reference baseline for the

measurements.

Figure 2.10 Total station on parapet wall

25

P7

P8

Figure 2.11 Location of monitoring prisms on site.

Readings were obtained for the monitored points 8 times a day. The first reading was at 4

a.m., about 3 hours before work started on site. Readings were taken at 7 a.m., 9 a.m.,

11a.m., 1 p.m., 3 p.m. and 5 p.m., while construction activity occurred. A final reading

was taken at 11 p.m. This cycle would be repeated for every day until the excavation was

completed. For each point, the instrument shoot at a prism, then rotated automatically

180o and the lens rotated also 180o to re-shoot the same prism. Here and on, this will be

referred as two face measurement. This technique eliminates the deviation between two

measurements on the horizontal plane.

26

2.3.2.1. Prism location

To monitor the displacements on the west wall six (6) points were used. Figure 2.12

illustrates the typical mounting of the prisms. Points P1, P3 and P5 were screwed into the

shoring wall adjacent to the piles with the inclinometers. The others were located (Figure

2-12) behind the piles at the other side of the alley and were glued on a small brick wall.

Figure 2.12 Points screwed and glued on the wall

27

Because the total station’s position could translate and rotate, displacements of the

points were calculated with respect to two fixed points that were far away from the

excavation. The first point (Point 7) was placed across of the excavation inside an alley at

a place where it could not easily seen (or maybe stolen by pedestrians). Point 8 was glued

on a building east of the excavation. The exact positions of points 1 through 7 are shown

in Figure 2-13. Point 8 is in Figure 2.14.

Point 1

Point 5

Point 2

Point 4

Point 6

Point 7

Point 3

Figure 2.13 Location of monitoring points and fixed point #7

28

Point 8

Figure 2.14. Location of fixed point #8

2.3.2.2. Leica TPS 1101 Professional Total Surveying Station

A Leica TPS 1101 Professional total surveying station, shown in Figure 2.15, was

used to monitor the displacement of eight optical prisms around the Olive 8 project. The

instrument includes monitoring firmware and robotic, automatic target recognition

capabilities which enabled repeated, automated position measurement of predetermined

points marked with prisms.

Measurement data can be accessed from storage on the instrument flash memory or

via remote communication employing the RS232 serial port of the instrument.

29

According to the manufacturer, the Leica TPS1101 is operational in temperatures

ranging from -20o C to 50o C and up to 95% humidity. Automated point measurements

can be taken within a range of 1000 meters. The stated angular accuracy of the instrument

is 1.5 seconds. The stated accuracy of the instrument is 3 mm at distances up to 300

meters.

Figure 2.15 Image of Leica TPS 1101

30

2.3.2.3. Instrument Errors

To understand the instrument errors one must to understand the instrument axes

(Leica TPS 1100 User Manual). The three axes of a total station are the vertical axis V

(standing axis), the titling axis T and the line of sight L and are shown in Figure 2.16.

Figure 2.16 Instrument axes

Instrument errors occur if the instrument differs from the following conditions:

• The line of sight L is perpendicular to the titling axis T

• The titling axis T is perpendicular to the vertical axis V

• The vertical axis is perfectly vertical

31

Line of sight error

As illustrated in Figure 2-16, the line of sight error is cause by the deviation “c”

between the optical line of sight and the perpendicular to the titling axis. This error

affects all horizontal readings and increases with steep sighting. The effect on the vertical

angle is very small and normally can be ignored. This effect can be eliminated by

precisely defining the line of sight error with the instrument’s on board calibration

function or two face measurements.

Figure 2.17 Line of sight error

Titling axis error

As shown in Figure 2.18, the titling axis error is caused by the deviation “a” of the

mechanical titling axis from the line perpendicular to the vertical axis. The titling axis

error can be observed when the telescope is moved vertically along a vertical line and the

32

crosshair moves away from the vertical line although the instrument has not turned in the

horizontal. This error increases with steep sightings, but has no effect on horizontal

sightings. The effect on the vertical angle is small and can be ignored. This effect can be

eliminated as well by precisely defining the titling axis error with the on board calibration

function, or with two face measurements.

Figure 2.18 Titling axis error

Vertical axis error

The vertical axis error (also called the standing axis error), as shown in Figure 2.19

is not an instrument error, but a set up error. It occurs if the vertical axis is not truly

plumb. The vertical axis error affects both the horizontal and the vertical angle readings

and cannot be eliminated with two face measurements. It can be avoided by carefully

leveling the instrument. The Leica TPS 1101 has two built-in axis compensators.

33

Leveling the instrument can be roughly done manually using the circular bubble and the

fine leveling to the plump line is automatically performed by the vertical compensator.

Figure 2.19 Vertical axis error

Compensator index error

The principle of the electronic compensator is quite similar to that of a circular

bubble. It consists of a small container filled with a special fluid. Due to gravity the

surface of this fluid is always horizontal. Sensors interact with the surface of this fluid to

determine the longitudinal and lateral tilt of the instrument. The horizontal and vertical

angles are then automatically corrected. As indicated in Figure 2.20, a compensator

index error occurs when the zero point of the compensator is not in the plump line. With

a dual axis compensator the index error of the compensator is divided into two

components, one alongside (longitudinal l) and the crosswise (transversal t) to the

telescope. The longitudinal index error “l” is similar to the V-index error and affects only

34

vertical angle readings. On the other hand the transversal compensator index error “t” is

similar to the titling axis error and mainly impacts the horizontal angle. The effect of both

errors can be eliminated by running the compensator calibration or by two face

measurements.

Figure 2.20 Compensator error

Vertical index error

A vertical index error “I”, as shown in Figure 2.21 exists if the 0o mark for the vertical

reading does not coincide with the instrument’s mechanical axis. The vertical index error

that affects all vertical angle readings is independent from the steepness of the shooting

point. The V-index error has no impact to the horizontal angle. The effect of the V-index

error can be avoided by precisely defining this error with the instrument’s on-board

calibration function or with two face measurements.

35

Figure 2.21 V-Index error

Collimination error

The collimination error, shown in Figure 2.22 is the angular divergence between the line

of sight and the ATR (Automated Target Recognition) camera axis. The horizontal

component of the collimination error affects the horizontal angle where as the V

component affects the vertical angle. The calibration routine allows one to set the

alignment of the centre of the camera with the optical line of sight.

Figure 2.22 Collimination error

36

2.3.2.4. Rigid Body Translation and Rotation Corrections

To evaluate the displacement of the monitored points a 3D rigid body translation

and rotations corrections were applied (L. Mcmillan, Comp136, Lecture 15 at USC,

1996). The correction procedure accounts the fact that the total station is not fixed on the

parapet wall. This transformation is only valid when all coordinate axes (x, y, z) remain

mutually perpendicular.

When the instrument was installed on the parapet, initial readings for all the

monitored points were taken. For this initial reading the total station was considered to be

at the origin of the coordinate system, i.e. at point (0, 0, 0). Thereafter, all data recorded

may have a slightly different origin because the surveying instrument slightly rotated and

translated.

The vectors that the monitored points form will be used to explain the correction

mathematically.

In section 2.3.2 it was mentioned that points 7 and 8 are far away from the

excavation and they did not move during the excavation. This means that the vectors

and should always be the same. However, in practice, due to rotations, translations

and instrument errors, a vector that is formed between two points will be different from

the one formed after the initial reading, and from all subsequent vectors. To correct these

slight changes, the vectors and have to be adjusted by a rotation and translation

matrix so that the vectors are exactly the same as that measured during the initial reading.

78V

87V

78V 87V

37

If point 7 is selected as the reference point, then the initial vector, , and the

vector at a reading i, , are:

oV78

iV78

⎥⎥⎥

⎦

⎤

⎢⎢⎢

⎣

⎡

−

−

−

=07

08

07

08

07

08

078

zzyyxx

V (2-1)

⎥⎥⎥

⎦

⎤

⎢⎢⎢

⎣

⎡

−

−

−

=ii

ii

ii

i

zz

yy

xx

V

78

78

78

78 (2-2)

In the same way the vectors from Point 7 to all the monitored points can be calculated

( and ……. and ). 071V iV71

076V iV76

The rigid body rotations and translation are shown in Figure 2-23.

38

38

Figure 2-23. Rigid Body Rotation and Translation correction

39

The unit vectors of and are: 078V iV78

078

078

0

78

^

VV

V = , where 078V is the magnitude of the initial vector. (2-3)

i

ii

VV

V78

7878

^= , where iV78 is the magnitude of the vector at a reading i. (2-4)

Let α be a new vector which is given by the cross product of:

0

78

^

78

^

0

78

^

78

^

VV

VVi

i

×

×=α (2-5)

α is a new vector which is perpendicular to the plane formed by vectors and ; it is

the vector that determines how much the current vertical axis z’ must be rotated to be

aligned with the initial vertical axis z.

078V iV78

If we take the dot product of and we get the cosine of angle θ that rotates the

current x’-y’ plane to the original x-y plane.

078V iV78

( ) 07878cos VV i •=θ (2-6)

and, ( )07878

1cos VV i •−=θ (2-7)

One must compute a rotation matrix, R, given as:

( ) ( ) ( )θθαθα cossin)()cos1( ∗∗∗ ++−= IskewsymR (2-8)

where symα is a symmetric matrix given by the product:

40

Tsym ααα *= (2-9)

and, skew(α ) is a skew matrix which is given by:

⎥⎥⎥

⎦

⎤

⎢⎢⎢

⎣

⎡

−

−=

00

0)(

xy

xz

yz

skewαα

αααα

α (2-10)

and I is the identity matrix.

Once R has been determined for each set of readings, the transformed vectors are given

by:

ij

Transj VRV 77 *= , where j is the number of the monitoring point (j=1, …, 6) (2-11)

To take into account the change in distance between and one must proportion the

components by:

078V iV78

iV

VScale

78

078

= (2-12)

So the final corrected vector will be,

Transj

Corrj VScaleV 77 *= , where j is the number of the monitoring point (j=1, …, 6). (2-13)

The monitored points P1 through P6 are expected to move, so any changes in the

corrected vector’s coordinates reflect the point’s displacement.

077 j

Corrj

j

j

j

VVzyx

−=⎥⎥⎥

⎦

⎤

⎢⎢⎢

⎣

⎡

ΔΔΔ

(2-14)

41

The same procedure can be performed starting from Point 8. For this case the vectors are

and , with: 087V iV87

⎥⎥⎥

⎦

⎤

⎢⎢⎢

⎣

⎡

−

−

−

=08

07

08

07

08

07

087

zz

yy

xx

V (2-15)

⎥⎥⎥

⎦

⎤

⎢⎢⎢

⎣

⎡

−

−

−

=ii

ii

ii

i

zz

yy

xx

V

87

87

87

87 (2-16)

For perfect measurements the transformed and scaled matrix should be identical,

whether we use either Point 7 or Point 8 as a reference. However, there is instrument

error in each measurement. For this reason we use the average of the displacements given

by points 7 and 8.

2.3.2.5. Data Acquisition Software

Infrastructure Technology Institute (ITI) staff wrote a program code in Java to

evaluate the rotation matrix based on the theory expounded in section 2.3.2.4. The

program structure that they followed was based on work by Blackburn (2005) related to

the excavation of the Ford Center in Evanston, IL. The Java code could detect if the data

collection was successful and, if it was unsuccessful, what type of error occurred. Data

collection and remote operation errors can occur due to power loss, communication

errors, obscured prisms, and instrument disturbance.

42

In the same script, there was the ability to allow the user to alter the frequencies of

the data collection.

2.3.2.6. Remote communication Methods

In section 2.3.2.1, it was mentioned that the monitoring data could be accessed

either from storage on the instrument’s flash memory or via remote communication by

employing the RS232 serial port of the total station. In this project, the second option was

used because it provided data without human intervention. From the time the instrument

takes the readings until it is displayed on a project website, a hardware communication

system intervenes to transfer the data. Figure 2.24 shows the setup that was used on site

to obtain the data. The total station was connected through a RS232 port to a computer

that stored the monitoring data in a hard drive. A wireless and a dial up modem were

connected on this computer to provide access to the internet. Power for these instruments

was provided by an electrical line after it was transformed to the proper voltage.

The sequence of data acquisition and transfer was:

1) A reading was taken and stored in the computer via the RS232 port. During every

measurement cycle, 2 sets of measurements were taken at each target point. Each set of

measurements contained a two-face reading. Thus, after each shot to a target, the

instrument rotated 180o, the lens rotated also by 180o and re-shot at the same point. In

each measuring cycle there were 4 data points for each reported measurement.

43

2) The computer connected to the internet at preset time intervals and transferred the

recorded data to a computer at Northwestern University. The first connection attempt was

through the wireless modem.

3) If the wireless modem failed to connect, the computer tried to connect to the internet

through the dial up model.

4) If both modems failed to connect, then the computer waited until the next reading was

taken and attempt to connect again.

5) At Northwestern University, the transferred data were inserted into a software where

the 4 readings for each point were averaged, and then the displacements were computed

using the correction procedures discussed in Section 2.3.2.4.

To protect the hardware from the weather, a waterproof box was used which was

mounted on steel brackets as shown in Figure 2-24.

44

Computer

Dial up modem Wireless modem

Power supply

Figure 2.24 Hardware used for remote communication

2.4. Summary

Olive 8 project included an 70 feet deep excavation to provide space for 5

underground parking levels. The shoring system is rather unique, along the West shoring

wall, due to the proximity of the Qwest building. The maximum allowed displacement,

imposed by the city of Seattle, was 1 inch. The subsurface investigation pointed that the

soil consisted of fill, dense gravelly silty sand, hard moist clayey Silt and very dense

gravely sand (going downwards). GeoEngineers Inc. designed the West wall shoring

system, which consisted of a soldier pile wall with shotcrete lagging, 9 to 10 rows of soil

45

nails and 4 or 5 rows of tiebacks. Three separate sections were designed to provide lateral

support for the West wall, due to the presence of a large utility vault between the

excavation and the Qwest building.

Numerical analysis performed by GeoEngineers showed that the maximum

predicted wall deflections vary from 0.8 to 1.5 inches which are higher than the 1 inch

limiting displacement. However they proceeded with the proposed design because

shoring wall and the soil layers were modeled conservatively and the actual behavior was

expected to be stiffer than the numerically predicted behavior.


West shoring wall. It included three conventional inclinometers to measure the lateral

wall movements, strain gages and load cells on the soil nails and tiebacks respectively.

An automated surveying system was also installed to monitor the 3D deflections of the

top of the West wall, at the positions where the inclinometers were located. The wall

movements were calculated relative to reference baseline points that were established far

away from the excavation. Readings were taken every day throughout construction until

the excavation was completed. Rigid body translation and translation corrections were

performed to account for the fact that the total station was not rigidly attached to the

parapet wall where it was mounted.

46 Chapter 3. Field Performance

3.1. Construction sequence

Tables 3.1, 3.2 and 3.3 present the construction procedure at the three design

sections along the west wall, as recorded for daily construction reports made by

GeoEngineers site personnel.

Construction Stages at Pile # 6 Stage

Number Activity Description 1 Sheet Pile Installation. 2 Excavation to +117.3 3 Soil Nails Row 2 Installation 4 Excavation to +113.4 5 Soil Nails Row 3 Installation 7 Excavation to +106.4 8 Soil Nails Rows 4, 5 and 6 Installation 9 Excavation to +101.4

10 Soil Nails Row 7 Installation 11 Excavation to +97.4 12 Soil Nails Rows 8 and 9 Installation 13 Excavation to +93.4 14 Soil Nails Row 10 Installation 15 Tieback Row A Installation 16 Excavation to +86.9 17 Tieback Row B and C Installation 18 Excavation to +75.5 19 Tieback Row D Installation 20 Excavation to +71.5

Table 3.1 Excavation sequence during construction for Pile #6

47


Number Activity Description 1 Sheet Pile Installation. 2 Excavation to +103.2 3 Soil Nails Rows 6 and 7 Installation 4 Excavation to +99.2 5 Soil Nails Row 8 Installation 6 Excavation to + 97.7 7 Soil Nails Row 9 Installation 8 Excavation to +93.2 9 Soil Nails Row 10 Installation

10 Tieback Row A Installation 11 Soil Excavated to +86.2 12 Tieback Row B Installation 13 Excavation to +74.2 14 Tiebacks Rows C and D Installation 15 Excavation to +65.2 16 Tieback Row E Installation 17 Excavation to +61.2


48


Number Activity Description 1 Sheet Pile Installation. 2 Excavation to +114.7 3 Soil Nails Rows 2 and 3 Installation 4 Excavation to + 110.5 5 Soil Nails Row 4 Installation 6 Excavation to +103 7 Soil Nails Rows 5 and 6 Installation 8 Excavation to +98 9 Soil Nails Rows 7 and 8 Installation

10 Excavation to +94.5 11 Soil Nails Row 9 Installation 12 Excavation to +92.5 13 Soil Nails Row 9 Installation 14 Tieback Row A Installation 15 Excavation to +84.5 16 Tieback Row B Installation 17 Excavation to +75.5 18 Tiebacks Rows C and D Installation 19 Excavation to +64.5 20 Tieback Row E Installation 21 Excavation to +60.5


The excavation sequence assumed in the numerical analysis made during the design

portion of the project by GeoEngineers, was described in Chapter 2.2.3. During

construction, the excavation sequence differed from that assumed in design. For example,

the specifications dictated that before installing any soil nail or tieback, the soil was not

to be excavated more than 2 feet below the installation level. However, the contractor

excavated soil below this specified depth, and occasionally two or even three rows of soil

nails were installed at the same time. In addition, the shotcrete lagging was to be placed

within 24 hours after the soil was exposed, but during construction the soldier pile wall

remained without lagging for as much as three days. While the former could be easily

49

incorporated into a finite element analysis of the excavation, the latter could only be

accounted for if a material-rate dependent model was used to represent the soil stress-

strain responses.

3.2. Wall Deformation

Slope inclinometers and automated surveying instruments were installed to monitor

the response of the wall due to the excavation process.

3.2.1. Slope Inclinometers

Pile #6: A cantilever movement profile was observed as the soil was excavated to

elev. 97.1 ft., as summarized in Figure 3-1. The maximum movement at the top of the

soldier pile wall was 0.42 inches. The reinforcing effect of soil nails is clearly in the

small lateral deflection that was observed at the level of the soil nails. As the excavation

was made to elevation +71.5 ft., small lateral movements were observed on the order of

0.1 inch. Most of the cantilever movement was obtained by the time that the excavation

level reached +97.1 ft, or about ½ the full depth of the case. Thereafter small deep seated

movements developed.

50

Pile #6

Deflection (in.)-1.0 -0.5 0.0 0.5 1.0

Elev

atio

n (ft

.)

60

70

80

90

100

110

120

130

May-25 July-20 September-7

+71.5

+97.1

+121.5

Qwest

Figure 3.1 Deflection for Pile #6

Pile #13: Pile #13 was the wall section where the large utility vault was located. In

Figure 3.2 it is observed that most of the lateral movements occurred after the excavation

depth exceeded elevation +97.0 ft. Due to the presence of the vault and the jet grouted

soil, the displacement at the top of the wall was 3 times less that the one observed at pile

#6; or approximately 0.1 inches. At this wall section the deep-seated movements were

51

slightly larger that the ones recorded at the cantilever section (0.16 versus 0.13 inches),

and were larger at a location about 5 feet above the bottom of the excavation.

Pile #13

Deflection (in.)-1.0 -0.5 0.0 0.5 1.0

Ele

vatio

n (ft

.)

50

60

70

80

90

100

110

120

130

May-25 July-20 September-7

+119.0

+97.0

+61.2

Jet-Grout

Utility Vault

Qwest


52

Pile #18

Deflection (in.)

-1.0 -0.5 0.0 0.5 1.0

Elev

atio

n (ft

.)

50

60

70

80

90

100

110

120

130

May-25June-20September-7

+117.0

+94.5

+60.5

Qwest


Pile #18: Response at Pile #18 is similar to that at Pile #6, as indicated in Figure

3.3. The maximum movement at the top of the soldier pile wall was 0.34 inches. At

about 10 feet above the final end of the excavation, deep-seated movements of about 0.08

inches were observed. Most of the cantilever movement developed by the time that the

excavation reached elevation +94.5 ft. After this point, small deep-seated movements

53

occurred. Clearly, in all cases the lateral movements were less than the 1 inch maximum

value targeted in design.

3.2.2. Automated Optical Surveying

As mentioned in Chapter 2, the survey prisms were placed on top of the soldier pile

wall and on the brick wall behind the alley at the sections monitored by the inclinometers.

Recorded data for Pile #6 are shown in Figure 3.4 where the displacements towards and

along the excavation, as well as the settlements are plotted versus time. The movements

parallel to the excavation are negligible. The “gaps” between the collected data are

attributed to failure of the total station to perform the point measurement which will be

discussed in section 3.3.1. The settlements recorded after July 10th do not represent the

actual response, as it will be discussed in Section 3.3.2. Figure 3.5 presents the deflection

towards the excavation for all monitored points. For Point #2, there are no data after the

2nd of June because the prism was lost due to construction activities. It is seen that the

displacements of the prisms on the brick wall (Points #2, #4 and #6) were negligible.

Table 3.4 summarizes the lateral movement recorded for the monitored points at the end

of the construction.

54

Point #1 - Pile #6

Displacement towards the excavation

4/24 5/15 6/5 6/26 7/17 8/7 8/28 9/18

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5

1.0

Displacement along the excavation

4/24 5/15 6/5 6/26 7/17 8/7 8/28 9/18

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5

1.0

Vertical Settlement

4/24 5/15 6/5 6/26 7/17 8/7 8/28 9/18

Settl

emen

t (in

)

-2.0

-1.0

0.0

1.0

2.0

Figure 3.4 Data acquired by automated surveying at Pile #6

55

Point #54/24 5/15 6/5 6/26 7/17 8/7 8/28 9/18

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5

1.0

Point #34/24 5/15 6/5 6/26 7/17 8/7 8/28 9/18

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5

1.0

Point #14/24 5/15 6/5 6/26 7/17 8/7 8/28 9/18

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5

1.0

Point #24/24 5/15 6/5 6/26 7/17 8/7 8/28 9/18

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5

1.0

Point #44/24 5/15 6/5 6/26 7/17 8/7 8/28 9/18

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5

1.0

Point #64/24 5/15 6/5 6/26 7/17 8/7 8/28 9/18

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5

1.0

Figure 3.5 Displacement along the wall of the monitored points

56

Deflection (inches)

Point No

Towards the

excavation Along the

excavation

#1 0.47 0.05 #2 - - #3 0.17 0.02 #4 -0.01 0.08 #5 0.19 -0.01 #6 0.06 0.03

Table 3.4. Final movement recorded for monitored points

3.2.3. Comparison of Inclinometer and Optical Survey Data

The location of points #1, #3 and #5 in regard with the inclinometers at piles #6,

#13 and #18, respectively, provided the opportunity to compare the lateral movement

data from automated surveys to that based on the inclinometer data.

Figures 3.6 and 3.7 compare lateral movements based on total station data at the

location of Pile #6 and Pile #13, respectively. It can be seen that the results agree quite

well throughout the entire construction period and the maximum difference is 10%.

For the section at Pile #18, Figure 3.8 shows that the total station recorded smaller

deflections than the inclinometers throughout construction. As the excavation reached

final depth, the data converged somewhat with the inclinometer data, showing slightly

larger displacements than the optical survey data. The maximum difference between the

optical survey and the inclinometer is 35% (or approximately 0.12 inches).

The total station’s accuracy is mentioned in Section 2.3.2.2 and is stated to be

±3mm (0.1 inch) at shooting distances up to 300 meters. Point 5 (at Pile #18) is located

approximately 45 meters away from the total station. Assuming that the error increases

57

linearly with distance, the error at point 5 is approximately 0.45 mm or 0.015 inches. The

inclinometer’s measurement error at the top of the shoring wall is 0.21 inches (as it will

be discussed in Section 4.2.2). This value is larger than the maximum difference between

the inclinometer and the survey data. Judging from Figures 3-6 and 3-7 and knowing the

accuracy of the Total Station the difference between the monitoring and the survey data,

at pile 18, can be caused by the inclinometer’s measurement error.

58


4/24/06 5/15/06 6/5/06 6/26/06 7/17/06 8/7/06 8/28/06 9/18/06D

ispl

acem

ent (

in)

-1.0

-0.5

0.0

0.5

1.0Total Station DataInclinometer Data


4/24/06 5/15/06 6/5/06 6/26/06 7/17/06 8/7/06 8/28/06 9/18/06

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5



the excavation for Pile #6

59


4/24/06 5/15/06 6/5/06 6/26/06 7/17/06 8/7/06 8/28/06 9/18/06D

ispl

acem

ent (

in)

-1.0

-0.5

0.0

0.5



4/24/06 5/15/06 6/5/06 6/26/06 7/17/06 8/7/06 8/28/06 9/18/06

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5




60


4/24/06 5/15/06 6/5/06 6/26/06 7/17/06 8/7/06 8/28/06 9/18/06D

ispl

acem

ent (

in)

-1.0

-0.5

0.0

0.5



4/24/06 5/15/06 6/5/06 6/26/06 7/17/06 8/7/06 8/28/06 9/18/06

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5




61

3.3. Summary of performance

In terms of accuracy, the automated surveying recordings agree with the recordings

taken from inclinometers. However, problems were observed are related to the practical

application of the automated surveying and with the correction which was used to process

the survey data. Both of these issues are discussed in the following sections.

3.3.1. Total Station Difficulties During Construction

As described in Chapter 2 the Leica TPS 1101 has two built-in axis compensators.

The first is the horizontal compensator, which corrects the line of sight error, the tilt error

and the standing axis error (only when the vertical compensator is turned on). The second

is the vertical compensator, which measures the longitudinal and transverse tilts of the

vertical axis, and ultimately brings the instrument back to the plumb line. The working

range of this compensator is ±0.1o from the vertical axis; if the instrument diverges more

than ±0.1o, then the total station shuts down and no readings are made. During the

excavation, the working range of the vertical compensator often was exceeded due to

weather and wind conditions. To reactivate the Total Station, one had to have access to

the total station to manually re-level the vertical axis back to the plumb line.

As previously mentioned in Chapter 2, the Total Station was mounted on top of the

parapet wall on the rooftop of Paramount Hotel. Access to the instrument was limited to a

hotel room window that led to a roof top. When the instrument was out of commission,

one could only access the instrument when the room was not rented. This less that ideal

situation resulted in access being prevented for periods between three days and two

62

weeks, depending on the demand of the room. These access restrictions were the cause of

the gaps in the total station data in Figures 3.5 though 3.8

Displacement towards the excavation for pile #6

4/24/06 5/15/06 6/5/06 6/26/06 7/17/06 8/7/06 8/28/06 9/18/06

Dis

plac

emen

t (in

)

-1.0

-0.5

0.0

0.5

1.0

9 days

16 days 9 days 18 days

7 days

Figure 3.9 Data lost when the instrument went out of balance

In Figure 3.9, the discontinuities between the collected data show that the Total Station

shut down numerous times, during the excavation monitoring. However, even with these

gaps, the trends in the data were defined. As shown in Figure 3.9, these shutdown

periods did not affect much the ability to define trends in responses. However, those

delays are clearly unacceptable in ground applications and real time monitoring, where

total station data are necessary.

63

3.3.2. The Problem of the Rigid Body Rotation Correction

The frequency of the total station shutdown led to the decision to turn off the

vertical compensator in an attempt to avoid losing data. Turning off the vertical

compensator does not affect the horizontal measurements. However, the vertical angles

are affected because they are measured relative to the standing axis which differs now

from the plumb line.

As mentioned in section 2.3.2.4, the rigid body rotation correction is applicable

only when the axes of the coordinate system are perpendicular to each other. So when the

vertical compensator is disengaged, the standing axis is no longer in plumb and the

calculated settlements of the monitored points are not correct. However, the horizontal

deflections along and towards the excavation are close to the inclinometer readings

because the horizontal compensator remained active.

Change of Elevation for pile #6

4/24/06 5/15/06 6/5/06 6/26/06 7/17/06 8/7/06 8/28/06 9/18/06

Dis

plac

emen

t (m

m)

-40

-20

0

20

40

Compensator ON Compensator OFF

Figure 3.10 Error in calculating the settlement when the vertical compensator was OFF

64

In Figure 3.10 the settlements using the rigid body rotation correction (Chapter 2.3.2.4)

are presented. It is obvious that after the compensator was turned off, the scatter in the

data is large unacceptably.

3.3.3. Summary

The lateral displacements based on the automated surveying data of prisms at the

top of the wall agree well with the inclinometer data. The less than rigid connection

between the total station and the parapet throughout the construction caused the total

station to shut down numerous times during the excavation. However, in an attempt to

provide continuous readings the vertical compensator was shut off. The settlement data,

collected during this shutdown, did not reflect the actual response because the standing

axis was no longer vertical to the horizontal plane. However, the horizontal

measurements were correct since the instrument’s horizontal correction was ON.

3.4. Predicted and Observed Responses

The recorded deflection from the inclinometers is compared with the deflections

predicted by GeoEngineers in Figure 3.11 for the three design sections. The deflections

were over predicted, especially for the Upper Sand and Clay layers. However, the

patterns of observed and predicted deflection profiles are similar implying that the

shoring wall and the soil responded more stiffly than expected.

The 1 inch movement limitation specified by the city of Seattle was not exceeded at

any time or at any section of the wall. In particular, the maximum observed lateral

movement was half of this limitation. In addition, the observed deflections at the Upper

65

Sand and Clayey Silt layers were, at certain depths, even 10 times smaller that the

predictions.

During the shoring wall design, the only laboratory data available were the index

properties of the Upper Sand and the Atterberg limits of the Clayey Silt. In addition, the

only triaxial tests available were performed back in 1972 when the adjacent Qwest

building was constructed. The difference between the wall deflections, at the Upper Sand

and the Clayey Silt layers is likely to be the result of conservative selection of soil

strength and stiffness parameters. In the Chapter 4, a finite element model made after the

excavation was completed will be presented. This model accounts for changes in soil

stress caused by construction of the Qwest building. Inverse analysis was performed to

define soil parameters that result in computed lateral wall deflections that are similar to

the observed values.

66

Pile #6

Deflection (in.)

0.0 0.4 0.8 1.2 1.6

Elev

atio

n (ft

.)

60

70

80

90

100

110

120

130

Pile #13

Deflection (in.)

0.0 0.4 0.8 1.2 1.6

50

60

70

80

90

100

110

120

130

Pile #18

Deflection (in.)

0.0 0.4 0.8 1.2 1.6

50

60

70

80

90

100

110

120

Observed DeflectionPredicted Deflection

Fill

Silty Sand

Clayey Silt

Clayey Silt Clayey Silt

Silty Sand Silty Sand

Fill Fill

Dense Sand Dense Sand

Figure 3.12 Comparison of predicted versus observed wall deflection

3.5. Summary

1) The construction sequence followed differed from the sequence used in the

design.

67

2) The inclinometer observations showed the 1 inch movement limitation specified

by the city of Seattle was not exceeded. In particular, the maximum observed

movement was half of this limitation.

3) The horizontal wall deflections towards and along the excavation based on the

total station agreed well with the movement with the movement of the top of the

wall based on inclinometer data.

4) The Total Station shut down because the working range exceeded the vertical

compensator’s capacity due to less than rigid connection with the hotel’s parapet

wall.

5) The rigid body rotation correction requires all axes to be perpendicular to each

other. When the vertical compensator was turned off, the vertical axis was not

perpendicular to the horizontal plane, resulting in unacceptably large scatter in the

settlement data.

6) The design finite element predictions of the magnitude of wall deflection were

larger than that observed. But the observed and predicted displacements, along the

length of the wall, have the same pattern, implying that the shoring wall and the

soil responded more stiffly than expected.

68

Chapter 4. Finite Element Simulation and Inverse Analysis

Chapter 3 showed that the numerically-predicted deflections made during design

were larger than the observed deflections. In this chapter a numerical model is presented

that explicitly accounts for the changes in stresses history of the soil caused by

construction of the Qwest building. This model was subjected to inverse analysis where

the lateral observed wall deflections were compared to the computed values. Because the

inclinometer data were collected manually at this project, these analyses were not

performed in real time, but after the excavation was completed.

4.1. Finite Element Model

The west shoring wall design was too complex to simulate in 3-D without substantial

simplifications. Therefore the excavation was simulated in 2-D assuming that the

complexity of the support system could be reasonably modeled.

The limitations of the pre-construction numerical simulations included:

1) The stress history induced by the construction of the Qwest building was not

considered explicitly. The building was simply inserted into the finite element

mesh with horizontal fixities, approximating the rigidity of the basement wall, and

values of K0 were assigned to the soil that were less than the at-rest values.

2) The soil strength and stiffness parameters used for the analysis apparently were

conservative because the computed movements were significantly larger than

those observed at the West wall.

69

To represent the conditions at the start of the Olive 8 excavation, the finite

element simulation presented in this chapter included a staged excavation for the Qwest

building retained by a soldier pile wall and four levels of struts. After the excavation was

completed, the soldier pile wall was replaced with a rigid wall and the struts with floor

slabs. The soil strength and stiffness parameters were the same initially in both designs

with the exception of the K0 which reflected the values expected for heavily

overconsolidated soil (The K0 values are shown in section 4.1.2). Numerical simulations

were performed for all three design sections. The purpose was to compare the computed

wall deflections with the design predictions, when the effects of the Qwest building

installation are taken into account. Moreover, inverse analysis was performed to calibrate

the soil parameters that provided the best fit between the numerical and the observed wall

deflections.

It is important to note that the only soil data available when selecting the design

values were the index properties of the Upper Sand and the Atterberg limits of the Clayey

Silt, SPT values for the entire soil profile, and triaxial tests performed in 1972 when the

Qwest building was constructed. So the initial choice of the soil strength and stiffness

parameters was primarily a case of applying engineering judgment to a rather scant set of

data.

4.1.1. Finite Element Mesh

Figure 4.1 presents the finite element mesh of a typical design section. The

70

dimensions of the mesh were 900 ft x 135 ft. The left and the right mesh boundaries

extend to a distance of 5H (H defined as the excavation depth) from both Olive 8 shoring

walls. At this distance the boundaries have no influence on the movements near the wall.

15-node triangular elements were used in the mesh to represent the soil and the

Qwest mat foundation. The mesh was more refined around the wall, the soil nails and the

tiebacks. The structural elements (shoring wall, struts, tiebacks, soil nails) supporting the

Qwest building excavation were modeled as elastic materials. In addition, frictional

interface elements were placed between soil and the shoring walls, the soil nails, and the

grouted length of the tiebacks to account for relative slip between the soil and the

structural element. The concrete elements (foundations, slabs) were modeled as elastic

materials. For the bottom of the mesh, zero horizontal and vertical movement was

specified. For the left and right boundaries, only zero horizontal movement was

precluded; the soil was free to move vertically.

71

71

+134.0

+119.5

+84.0

+59.5

+ 0.0

265 ft. 135 ft. 120 ft. 380 ft.

Fill

Silty Sand

Clayey Silt

Dense Sand

Figure 4.1 Finite element mesh for typical design section

72

4.1.2. Constitutive Models and Input Parameters

The numerical analysis adopted the same soil profile as GeoEngineers used for the

design predictions. This profile is described in Section 2.1.1 and was based on the

subsurface explorations performed by HartCrowser Inc. at their report regarding the

Olive 8 excavation in 2000. While the water table was not encountered in any of the

boring logs, both drained and undrained numerical analyses were used to simulate the

behavior of the Clayey Silt. In all cases, drained conditions were assumed for the Fill and

the Silty Sand.

Drained analysis does not generate pore pressures and the finite element model

accounts for volumetric changes that occur due to compression of the voids in the soil

skeleton. This is the case for dry or fully drained soils due to high permeability or low

rate of loading.

Undrained analysis simulates constant volume condition, and thus the excess pore

water pressures. This option is used to model soils when one is interested in short term

stability and displacement estimates.

For this excavation, the subsurface investigation showed that the soil layers were not

dry but moist (10-18% water content for the Silty Sand and 18-30 % for the Clayey Silt

layer). For the case of the Silty Sand, drained analysis can represent relatively accurately

the actual material behavior. The behavior of the Clayey Silt, with its low hydraulic

conductivity, however, cannot be modeled accurately with drained analysis. The actual

73

behavior of a partially saturated, partially drained clay falls between the limits of

drained and undrained analysis. The finite element simulation was made assuming both

drained and undrained analysis for the Clayey Silt layer to compare the computed

deflections based on the two approaches.

The constitutive soil models in the analysis were either the Hardening–Soil, or the

Mohr-Coulomb models. The hardening-soil model represents the response of all soil

layers except for the jet grout around the utility vault in design section Pile #13, which is

assumed to respond as a Mohr-Coulomb material. Table 4.1 summarizes the soil

parameters used in this analysis. All the soil parameters used for the simulation were the

same with the ones used for the GeoEngineers predictions, with the exception of the

lateral stress coefficient, K0. K0 was increased by 40% for the overconsolidated Upper

Sand and the Dense Sand layers more accurately represented in in-situ conditions prior to

the excavation of the Qwest building basement. The K0 values from GeoEngineers and

this finite element analysis are compared in Table 4-2. The dilation angle, ψ, is assumed

to be zero for all the soil layers. The dilation angle is a value that measures the dilatancy

of granular soils during shearing. When ψ is equal to zero the soil expansion due to

shearing is zero.

74

Soil Parameters

Parameter Fill Silty Sand

Clayey Silt

Dense Sand Jet-Grout

Constitutive Law S/H S/H S/H S/H M/C Soil Unit Weight, γ (pcf) 125 130 125 130 145

Friction Angle φ (o) 32 38 34 40 0 Cohession,c (pcf) 100 200 200 0 7000

Lateral Stress Coefficient, K0 0.47 0.6 0.7 0.6 - Poisson's Ratio, ν 0.3 0.3 0.2 0.3 -

Dilation Angle, ψ (o) 0 0 0 0 - Soil Stiffness (ksf) refE50 600 1000 500 1500 -

Unload/Reload Stiffness (ksf) refurE 2600 3000 1600 4500 -

Oedometer Stiffness, (ksf) refoedE 500 1000 700 1500 -

Interface Reduction Factor, Rinter 0.67 0.67 0.67 1 - Reduced Interface Factor, Rinter - 0.2 0.2 - - Reference Pressure pref (atm) 1 1 1 1 1

S/H stands for the Soil Hardening Model M/C stands for the Mohr-Coulomb Model

Table 4.1 Soil Parameters used for design

Lateral Stress Coefficient, K0

Layer GeoEngineers Design

This Analysis

Fill 0.47 0.47 Upper Sand 0.38 0.66

Glaciolacustrine Silt/Clay 0.7 0.7 Lower Sand 0.36 0.6

Table 4-2 Comparison of K0 values from GeoEngineers and this analysis

Tables 4.3 and 4-4 illustrate the input parameters of the concrete and structural

75

elements, respectively. All the elements are assumed to behave as elastic materials.

Similarly to the soil parameters, all concrete elements, soil nails, tiebacks, the West

soldier pile wall and the Qwest rigid wall, have the same properties as did in

GeoEngineers design. The Qwest foundation and basement slab is represented by 15-

node elements. The Qwest shoring wall was assumed to have identical properties with the

Olive 8 West shoring wall. In addition, the struts used to retain the Qwest excavation

were defined as circular steel sections, 24 inches in diameter and 0.5 inches thick. The

vertical and horizontal spacing of the struts was 12 and 30 feet, respectively. The floor

slabs, have the same normal stiffness (EA) as the Qwest rigid wall. The East Olive 8

shoring wall consists of W24x131 sections spaced every 8 ft. The tiebacks used for the

retaining system, have the same properties with the ones used at the West Olive 8 wall.

Parameter Mat Footing

Basement Slab

Element Type 15-node element

15-node element

Unit Weight, γ (pcf) 145 145 Poisson's Ratio, ν 0.15 0.15

Young’s Modulus E (ksf) 7.2 x 105 4.32 x 105

Table 4.3 Properties of concrete elements

76

nt g

I) t)

al

t) t

e Structural ElemeBendin

Stiffness, (E(kip-ft2/f

NormStiffness, (EA)

(kips/fElemen

Typ

West Wall of Olive 8 (W27x146) 05 06 PLATE 1.42 x 1 1.24 x 1East Wall of Olive 8 (W24 x 131) 05 05 PLATE 1.1 x 1 3.1 x 1

Qwest Temporary Wall 05 05 PLATE 1.1 x 1 3.1 x 1Qwest Internal Bracing n/a 05 ANCHORS1.4 x 1

Qwest Rigid Wall 08 08 PLATE 1.89 x 1 3.5 x 1Qwest Slabs /a 08 PLATE n 3.5 x 1Utility Vault 03 05 PLATE 6.1 x 1 2.9 x 1

Soil Nails (4' by 3' spacing) /a 04 D n 6.4 x 1 GEOGRISoil Nails (4' by 3' spacing wi

additional Nail at center patter

th of

n) /a 04 D n 9.6 x 1 GEOGRI

Tieback Anchors (No Load Zone) /a 03 ANCHORSn 8.2 x 1Tieback Anchors (Grouted Body) /a 04 D n 3.5 x 1 GEOGRI

Table 4.4 Stiffness of structural elements

In PLAXIS, Plate elements are defined using a flexural rigidity, an axial stiffness,

EA, and an ultimate bending moment, EI.

Geogrids are flexible elastic elements that represent a grid or a sheet of fabric, they are

used to model soil nails, the grouted part of tiebacks and geomembranes. Geogrids cannot

sustain compressible forces, the only property is the elastic axial stiffness, EA.

Anchors are elastoplastic spring elements that are used to model struts and the non-

grouted portion of tiebacks.

For tiebacks (node-to-node anchors) it is usual to neglect any shear stresses mobilized

between the soil and the free anchor length, so the spring is connected to the wall at one

end and to the soil and fixed anchor length at the other end. A node-to-node anchor

requires specifying the axial unit stiffness, EA.

77

For struts (fixed-end anchors) the anchors are springs that are used to model a

reaction at a single point. The length is defined as the distance between the anchor

connection point and the fictitious point in the axial direction where the displacement is

assumed to be zero. A fixed end anchor requires one to specify the axial stiffness, EA,

and the out of plane spacing.

4.1.3. Excavation Sequence

The excavation sequence for each of the three design sections consists of two parts:

A) The Qwest building construction, which is the same for all three design sections.

B) The Olive 8 excavation which varies for every design section.

The Qwest building construction is simulated by following the steps:

1) Soldier pile installation, with no change in soil stresses

2) Soil excavation to a level 2 feet below each strut level

3) Strut installation

4) Repeat 2, 3 until the lowest strut is installed

5) Excavation to final depth

6) Mat foundation construction by activating the concrete elements representing the

slab and applying the design load 10 ksf for the Qwest foundation and 0.3 ksf for

the traffic loads in the alley

7) Placement of permanent Qwest structure by replacing the soldier pile wall with a

rigid concrete wall and the struts by floor slabs

78

8) Reset displacements to zero before the start of the Olive 8 excavation

The Olive 8 excavation simulation remained almost the same as described in section

3.1. In addition, it included the simulation of the east soldier pile wall and the East wall

tieback installations. The East shoring wall was retained by 6 rows of tiebacks that were

installed in a pattern of 8 ft horizontal and 10 ft vertical spacing. It was assumed that the

East shoring wall was installed at the same time as the West shoring wall. The

excavation grade was uniformly lowered across the site and the east wall tiebacks were

installed at the same time as the lateral support on the West wall that was located at the

same elevation. Tables 4.5 and 4.6 present the pre-stress load applied to each row of the

tiebacks for the West and East walls, respectively.

Pre-stress Load at West Wall Tiebacks

Tieback Row No.

Maximum Pre-stress Load

(kips)


(kips/ft) 1 113 14.13 2 118 14.75 3 119 14.88 4 127 15.88 5 113 14.13

Table 4-5. Pre-stress load for West wall tiebacks

Pre-stress Load at East Wall Tiebacks

Tieback Row No.


(kips)


(kips/ft) 1 62 7.75 2 150 18.75 3 115 14.38 4 125 15.63 5 120 15.00 6 65 8.13

Table 4-6. Pre-stress load for East wall tiebacks

79

4.2. Finite Element Results and Analyses

4.2.1. Drained Analysis

Figures 4.2 and 4.3 show the computed relative shear stresses that occurred as a

result of simulating the Qwest building excavation. They represent the results of the finite

element simulation at the stage right before installing the soldier pile for the Olive 8

excavation. Figure 4.2 presents the relative shear stress as a result of wishing the Qwest

building in place and adding the foundation loads. Figure 4.3 presents the relative shear

stress after simulating the Qwest building excavation. The relative shear stress gives an

indication of the proximity of the stress point to the failure envelope and is defined as:

maxτττ =rel (4-1)

where τ is the maximum value of shear stress (radius in Mohr stress circle) and τmax is

the maximum value of shear stress for the case where Mohr circle is expanded to touch

the Coulomb failure envelope keeping the intermediate principal stress constant. A value

of zero indicates isotropic stress and a value of 1 indicates failure.

In GeoEngineers design (Figure 4-2), the low K0 values at the Silty Sand and Dense

Sand layers generally produced higher relative shear stresses. The soil in the alley, next to

the Qwest building, was not greatly affected by the Qwest building installation. This is

because the entire building was simply inserted in the mesh and the horizontal fixities at

the stiff basement wall did not allow any lateral deflection. Changes in shear stress levels

80

are also observed next to the upper portion of the Qwest basement wall and below

the Qwest foundation as a result of applying foundation loads.

As a result of simulating the Qwest excavation, the relative shear stresses decreased

at the Silty Sand and Dense Sand layers (Figure 4-3). However, in the alley next to the

Qwest building relative shear stresses near failure developed. This happened because the

Qwest shoring wall was allowed to deflect laterally. The maximum relative shear stresses

are located at the lower portion of the Qwest basement wall due to deep seated

movements that developed during the excavation, and below the Qwest foundation as a

result of applying foundation loads.

Figure 4.2 Relative shear stress: No simulation of Qwest excavation

81

Figure 4.3 Relative shear stress: With simulation of Qwest excavation.

Figures 4.4 through 4.6 compare the results of the finite element model simulations

including the Qwest excavation for all design sections with the results from

GeoEngineers (no Qwest excavation simulation) finite element predictions and the

recorded deflections from the inclinometers. In general, taking into account the

installation of the Qwest Excavation in the finite element analysis produced smaller wall

deflections than the simulation without it; however, these computed lateral deflections

were still larger than the observed soil responses. The numerically-calculated deflections

in the fill, for Pile #6 and Pile #18 reasonably agree with the observed at Figures 4.4 and

4.6. At all sections, the largest difference between the observed and the numerically-

calculated deflections occurred at the Upper Sand and the Clayey Silt layers. These layers

82

apparently responded in the field more stiffly than assumed in the finite element

model calculations; the maximum difference in responses occurred at the Clayey Silt

layer. However, the patterns of the observed wall deflections have similar shapes with the

ones calculated by the finite element analysis.

Pile #6

Displacement (in.)

-1.0 -0.5 0.0 0.5 1.0

Dep

th (f

t.)

60

70

80

90

100

110

120

130

Observed Defl.GeoEngineers This Study

Qwest

Figure 4-4. This study, compared to the GeoEngineers predictions and the observed

deflections for Pile #6

83

Pile #13

Displacement (in.)

-0.5 0.0 0.5 1.0 1.5

Dep

th (f

t.)

50

60

70

80

90

100

110

120

130

Observed Defl. GeoEngineers This Study

Jet Grout

Qwest

Utility Vault



84

Pile #18

Displacement (in.)

-0.5 0.0 0.5 1.0 1.5

Dep

th (f

t.)

50

60

70

80

90

100

110

120

Observed Disp. This StudyGeoEngineers

Qwest



The difference between the observed and the calculated displacement profiles can be

caused from several factors. First, the inclinometers were located on the piles and the

85

observed profiles are influenced by the stiffness of the soldier piles. In a plane strain

finite element analysis, the stiffness of the soldier pile wall and soil system is smoothed

when the axial stiffness (EA) and the bending moment (EI) of the piles are divided by

their out-of-plane spacing of the soldier piles. Thus, if the inclinometers are not far

enough behind the piles, the results will be influenced by the locally stiff pile section, in

contrast to the average plane strain value.

Other factors that influence the difference between the calculated and the observed

displacement profiles are, the use of drained instead of undrained analysis to model the

Clayey Silt layer, conservative choices of strength and stiffness parameters, 3-D effects,

and the way that the west wall tiebacks were modeled as anchors. All these factors will

be examined further in the following sections.

4.2.2. Undrained Analysis

To provide a lower limit at the computed response in the Clayey Silt layer, finite

element simulations were performed for all design sections with Clayey Silt assumed to

behave as an undrained material. Note that when simulating the Qwest building

excavation, the Clayey Silt was assumed to behave as a drained material because Qwest

was constructed 25 years ago and any excess pore pressures that developed would have

dissipated. While this approach is clearly an approximation, it is intended to serve as an

indicator of the stiffest response one could expect using the selected soil parameters.

Figure 4.7 compares the computed lateral deflections for drained and undrained

86

analysis with the observed and GeoEngineers predicted displacement profiles (no

Qwest excavation). Smaller deformations are seen in the Clayey Silt layer at all the

design sections when undrained analysis is used. The displacements are closer to the

observed displacement profile; however, the computed displacements still exceed the

observations. Other factors contribute to the differences in responses, as will

subsequently discussed.

87

Pile #6

Displacement (in.)

-1.0 -0.5 0.0 0.5 1.0

Dep

th (f

t.)

60

70

80

90

100

110

120

130

Pile #13

Displacement (in.)

-0.5 0.0 0.5 1.0 1.5

50

60

70

80

90

100

110

120

130

Observed Deflection Drained AnalysisGeoEngineersUndrained Analysis

Pile #18

Displacement (in.)

-0.5 0.0 0.5 1.0 1.5

50

60

70

80

90

100

110

120

130

Figure 4.7 Comparison of drained and undrained finite element analysis in contrast with

the observed and predicted displacement profiles

88

4.3. Analysis of Results

4.3.1. 3-D Effects

Figures 4.8 and 4.9 (Finno et al. 2007) show the 3-D effects in deep excavations ias

a function of the excavation depth and the dimensions of the excavation. The results of

the analyses are represented by the plane strain ratio, PSR, defined herein as the

maximum movement in the center of an excavation wall computed by 3-D analyses

divided by that computed by a plane strain simulation. L/He is the ratio of the length of

wall to the excavation depth and L/B is the ratio of the plan dimensions of the excavation,

with L being the side where movements are computed.

Figure 4.8 shows the PSR plotted versus L/B ratio. When the L/B ratio is less than

4, L/HE must be taken into consideration for determining the PSR. The Olive 8 west

shoring wall was 240 feet long and the width of the excavation was 120 feet. This defines

a ratio L/B equal to 2; the PSR varies from 0.75 to 1.0 and depends on the L/HE ratio at

any stage of the excavation.

89

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

1.1

PSR,

δ3D

/δ2D

0 1 2 3 4L/B

Figure 4.8 Effects of plan dimensions on PSR (after Finno et al. 2007)

The trends at Figure 4.9 indicate that L/HE ratios greater than 6 result in an

excavation response which has a PSR approximately equal to 1, thus suggesting that

results of plane strain and 3-D analyses will yield the same maximum wall displacement

in the center of the excavation for this geometry. Large differences between plane strain

and 3-D responses are apparent when L/He is less than 2, implying that as the excavation

gets deeper relative to its length, more restraint is provided by the sides of the excavation.

The Silty Sand Layer can be located from 13 to 50 feet below the ground surface.

This defines a ratio L/He which varies from 5 to 18. When the excavation is less than 50

ft deep, or when Silty Sand layer is being excavated, the excavation can be represented

adequately by plane strain conditions. The Clayey Silt layer is located 50 to 75 feet below

the ground surface. When excavating through the Clayey Silt layer at these depths, the

L/He varies from 3 to 5, so the excavation cannot be accurately represented by plane

90

strain conditions, especially when the excavation reaches the full depth.

All other things being equal, this means that the movements that occur when

excavating between 50 and 75 ft. are over-predicted when plane strain finite element

analysis is used to simulate the west shoring wall excavation. The stiffening effects of the

corners of the excavation are not considered in plane strain analysis which accounts for

some of the difference between the computed and observed results.

0 2 4 6 8 10L/HE

120.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

1.1

PSR,

δ3D

/δ2D

Silty SandClayey Silt

1 1

Figure 4.9. Effects of plan dimensions and depth of excavation on PSR

4.3.2. Numerical Simulation of Tiebacks

Another parameter that has to be evaluated in order to validate the difference

between the observed and the numerically calculated displacement profiles is the

numerical representation of the tiebacks in the finite element analysis.

91

Tiebacks have a three-dimensional geometry, and their representation in plane

strain analysis involves significant approximations. A question in this analysis was

whether the 2-D representation of the tiebacks brings them so close in the resulting mesh

that the stresses transmitted to the soil overlap, reducing the tieback load-bearing capacity

and producing excessive displacements.

A representation of the tiebacks as (elastic) struts approximates the field conditions

in plane strain without concentrating stresses in the soil near the anchors.

Struts that are equivalent to the tiebacks were modeled as springs that transmit axial

force. The major anchor property is the axial stiffness, EA, entered in the unit of force.

Assuming an elastic behavior, this implies the equivalence:

StrutEquivalentTieback LEA

LEA

⎟⎠⎞

⎜⎝⎛=⎟

⎠⎞

⎜⎝⎛ (4-1)

Therefore the equivalent stiffness of a strut is:

( ) StrutEquivalentTieback

StrutEquivalent LL

EAEA ⋅⎟⎠⎞

⎜⎝⎛= (4-2)

Table 4-7 illustrates the calculations for the axial stiffness of the equivalent struts. For

this calculation, the axial stiffness, EA, of the tiebacks was that used at the finite element

analysis. The spacing was assumed 8 ft, the same with the one used for the tieback

simulation. For a strut length of 60 ft, the EA for the equivalent strut, that is appropriate

for a plane strain analysis, is shown in the last column of Table 4-7.

92

Representation of Tiebacks by Equivalent Struts

Normal Tieback Stiffness

(EA)

Normal Tieback Stiffness

(EA)

Tieback Length EA/L

Length of Equal

Strut

Normal Equal Strut

Stiffness (EA)

Normal Equal Strut

Stiffness (EA)

(kips/ft) (kips) (feet) (kips/ft) (feet) (kips) (kips/ft) Row 1 8200 65600 21.1 3109 60 186540 23318 Row 2 8200 65600 16.9 3882 60 232899 29112 Row 3 8200 65600 15.1 4356 60 261355 32669 Row 4 8200 65600 15.0 4373 60 262400 32800

Stiffness and Length for the tiebacks refer to the non grouted tieback body The normalized values refer for 8 feet spacing

Table 4-7 Representation of Tiebacks by equivalent struts

Figure 4.10 presents the computed deflections at the location of Pile #6 for both

representations of the tiebacks. These simulations were performed when the Clayey Silt

was modeled as a drained material. The results show that modeling the tiebacks as

equivalent struts has little effect on the calculated deflections.

93

Pile #6

Displacement (in.)

-0.5 0.0 0.5 1.0

Dep

th (f

t.)

60

70

80

90

100

110

120

130

Equal Struts Tiebacks

Fill

Silty Sand

Clayey Silt

Figure 4.10 Deflection profiles for tiebacks and equivalent struts

4.3.3. Summary

The finite element numerical analysis that included the Qwest excavation produced a

displacement profile that was closer to the movements observed in field than that based

94

on a similar analysis without an explicit simulation of the Qwest excavation.

However, the computed deflections in the Silty Sand and Clayey Silt layers were still

larger than observed.

The stiffness of the soldier piles influenced the observed deflection profile because

the inclinometers are located on the soldier piles. In a plane strain analysis this stiffness is

smeared because the stiffness of the soldier pile is divided by the out-of-plane spacing of

the soldier pile.

Modeling the Clayey Silt layer as an undrained material reduced the calculated

deflections. However, the deflections at the bottom of the excavation still were larger

than observed. This difference can also be related to the three dimensional stiffening

effects at the corners of the excavation. For the shallower excavation through the Silty

Sand layer, the 3-D effects were insignificant and the shoring wall can be represented by

plane strain conditions. For the deeper excavation through the Clayey Silt layer, the 3-D

effects were significant, especially as the excavation reaches full depth..

Representing the tiebacks as either tiebacks or equivalent struts, had little effect on

the computed lateral deflections.

95

4.4. Inverse Analysis

Because of the differences between the observed and the computed lateral

deflections, inverse analyses were conducted to find the soil parameters that provided the

best fit to the observed lateral deflections. A significant question for this excavation case

history was whether the small observed displacements at the Silty Sand layer, were large

enough to be used for the optimization procedure. For these small lateral displacements to

be useful in inverse analysis, they must be larger than the instrument errors associated

with the inclinometer data (Rechea 2006). Two optimizations were performed for the Pile

#6 design section. Wherein the Clayey Silt was assumed to act as either drained or

undrained.

4.4.1. Procedures

Finno and Calvello (2005) and Finno and Rechea (2006) presented an inverse

analysis procedure that used construction monitoring data to update predictions of

deformations for supported excavation systems. The field observations were obtained

from inclinometers that measured the lateral movements of the soil approximately 6 ft

behind the supporting walls of the excavation throughout the construction. The

constitutive soil responses were represented by the Hardening-Soil model. Of the six

basic input parameters of that model, up to three parameters per layer were optimized

( , , ), while the other parameters were kept constant or related to the

updated value of the optimized value. This methodology was effectively used to calibrate

refE50refurE ref

oedE

96

the model of the excavation based on responses measured at the beginning stages of

construction, so that good predictions can be made of the soil lateral deformation close to

the wall at later stages. UCODE (Poeter and Hill, 1998) was used to optimize the finite

element models for several deep excavation case histories. UCODE is a universal inverse

code that can be used with any application model; it performs gradient-based inverse

modeling and estimates the parameters by calculating the values that minimize the

objective function using non-linear regression.

The difference between the calculated and the monitored quantities is evaluated

through an objective function, also known as error function; the optimum values are the

ones that minimize it. The expression for the weighted least-squares objective

( )bS ' function is:

( ) ( )[ ] ( )[ ] eebyybyybS Tωω =−−= ••

−

'' (4-3)

where:

b : is a vector containing values of the parameters to be estimated,

y : is the vector of the observations being matched by the regression,

( )by ' : is the vector of the computed values which correspond to observations,

ω : is the weight matrix reflecting the error of the measurements,

e : is the vector of residuals, y- y’(b).

This function represents a quantitative measure of the accuracy of the predictions.

97

A quantity called Relative Fit Improvement, RFI, can be employed to indicate

the percent improvement at each stage, i, compared to that based on the initial

predictions:

( ) ( )( )initial

iinitiali bS

bSbSRFI −= (4-4)

where ( )initialbS is the objective function based on the initial estimate of parameters and

( )ibS is the objective function at stage i.

Key factors for successful calibration are a reasonable representation of constitutive

responses of stress history of the soil, wall installation effects and the construction

sequence.

A sensitivity analysis was used to identify the soil parameters most likely to be

successfully calibrated. The sensitivity analysis calculates the relative importance of the

parameters being simultaneously estimated. A useful statistic parameter UCODE is the

composite scaled sensitivity, cssj, which indicates the total amount of information

provided by the observations for the estimation of parameter j, and is defined as:

⎥⎥⎥

⎦

⎤

⎢⎢⎢

⎣

⎡

⎟⎟⎠

⎞⎜⎜⎝

⎛⎟⎟⎠

⎞⎜⎜⎝

⎛

∂∂

∗= ∑=

b

ND

jijj

j

ij b

by

NDcss

1

2

2/1'1 ω (4-5)

where

y´i is the ith computed value,

bj is the jth estimated parameter,

98

j

ib

y∂

∂ is the sensitivity of the ith computed value with respect to the jth parameter,

and

ωij is the weight of the ith observation with respect to the jth parameter

The inverse analysis presented herein uses the inclinometer observations and

adjusts the selected parameters until the difference between the observed and calculated

quantities falls below a pre-established threshold. Herein, this threshold was either 10%

change of the objective function, or stabilized RFI over 90% for three consecutive

iterations. One factor that affects the accuracy of the inverse problem is the measurement

error, which depends on the instrument accuracy and error, field conditions and

operator’s ability. The instrument error is the only factor that can be quantified, the

others can not be defined easily but they can be minimized with good field technique.

4.4.2. Selection of Field Observations for Inverse Analysis

The manufacturer specified inclinometer measurement error is ±0.25 mm/m. The

accuracy of the measurement (in meters) versus the measurement depth, d, (in meters

also) is given by:

derror ⋅±=1000

25.0 (4-6)

The 95% confidence interval for the accuracy of a measurement is:

96.1100025.0 d

⋅±=σ (4-7)

99

The weight of each observation depends on the measurement depth and is given by:

)0001.0(11

22 dweight

⋅==

σ (4-8)

Generally, the observations employed in the analysis should to be larger than the

measurement error, and at least one order of magnitude bigger than the 95 % confidence

interval of the accuracy of the measurement (Rechea, 2006).

If all inclinometer measurements were used for the inverse analysis regardless the

measurement error, then each observed displacement is calibrated relative to one

calculated with the predetermined weight given by equation (4-6). Hence, there would be

uncertainty that enters into the parameters optimization because, when selecting the

observations for the inverse analysis, the accuracy of the observations is not considered.

Inclinometer data from Pile 6 was used for the inverse analysis. The data are the

readings taken at the end of the excavation. Recordings from earlier excavation stages

could not be used to optimize the Clayey Silt because reliable deep-seated movements

started to develop only after the excavation level reached an elevation about 2/3 (two

thirds) of the full excavation depth. Moreover, reliable observations in for Silty Sand

layer started to develop after the excavation level reached ½ of the full excavation depth.

Figure 4.11 shows the inclinometer deflection at Pile 6 versus the 95% σ accuracy of the

measurement.

It can be observed that only three readings at the Upper Sand layer are higher than

the 95% σ accuracy of the measurement. The thickness of the Upper Sand layer was 25

100

feet and the observations used were located at the first 5 feet. Within the clay layer,

the readings are larger than the instrument error and most of them can be used for the

optimization. Thus at excavation stages earlier than the final stage, not enough

observations were available for a meaningful calibration.

Pile # 6: Deflection and Instrument Error vs Elevation

Deflection And Error (in.)-0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 0.2 0.3 0.4 0.5

Ele

vatio

n (ft

.)

60

70

80

90

100

110

120

130

Deflection95% Confidence

Silty Sand

Clayey Silt

Fill

Figure 4.11 Observations vs. instrument error for conventional and in-place inclinometer

at Pile #6

A way to reduce the instrument error is to use inclinometers with accuracy greater

than the conventional inclinometers or to obtain a site specific measure accuracy. This

can be accomplished by evaluating reading taken at the start of the project before any

101

construction related to ground movement develops. Figure 4.12 compares, at Pile

6, the 95% σ accuracy of a conventional inclinometer versus a more accurate sensor. If a

more accurate sensor was used, then the number of reliable observations could be larger.

Pile # 6: Deflection and Instrument Error vs Elevation

Deflection And Error (in.)

-0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 0.2 0.3 0.4 0.5

Ele

vatio

n (ft

.)

60

70

80

90

100

110

120

130

Observed DeflectionConventional Incl. errorIn-place Incl. error

Silty Sand

Clayey Silt

Fill

Figure 4-12. Comparison of inclinometer errors

Figure 4.13 presents the 95% σ accuracy of the measurement stated by the

manufacturer versus the observed error the inclinometer at Pile 6, based on data collected

prior to the start of excavation when the wall has not started yet to deflect. The observed

error can be roughly approximated by a straight line. The equation of the line is:

error (inches) = 8.53•10-4•h (in feet) (4-9)

102

or

error (meters) = 6.38•10-5•h (in meters) (4-10)

where:

h is the height for from the deeper measurement, where the error is assumed to be equal

to zero. The measured error is about 70% less than the manufacturer’s stated value.

Pile #6: Deflection and Instrument Error vs Elevation

Observed and Stated error by Manufacturer (in)

-0.5 -0.4 -0.3 -0.2 -0.1 0.0 0.1 0.2 0.3 0.4 0.5

Dep

th (f

eet)

60

70

80

90

100

110

120

130Fill

SIlty Sand

Clayey Silt

95% Confidence (Manufacturer)

Observed Error

Observed Deflectionerror (in)= + 8.13*10-4 * h (ft)

Figure 4.13 Observed and stated by manufacturer instrument error

103

Table 4.8 lists the observations in the calibration. Observations are named

after the soil layer where they were measured: F for the Fill, S for the Silty Sand layer, C

for the Clayey Silt layer and DS for the Dense Sand layer. The 95% σ measurement

accuracy, stated by the manufacturer, is calculated with equation (4-5). The 95% σ

observed measurement accuracy is calculated by (4-8). The observations used for the

optimization procedure for fully drained analysis are in bold. For these observations the

stated accuracy is used. The observations used for the optimization procedure when the

Clayey Silt is modeled as an undrained material are in bold and italic bold.

Observations marked in italic are excluded from the inverse analysis because their

magnitudes are either lower than the 95% σ measurement accuracy (stated or observed)

or the observed deflection is negative (movement away from the wall). Negative

deflections were not used because the inclinometer was located on the soldier pile. If the

inclinometer was installed 4-6 feet behind the soldier pile, negative deflections likely

would not have been observed. The observations for the Fill and the Dense Sand layers

were not used for the calibration.

104

Table 4.8 Observations used for inverse analysis

4.4.3. Finite Element Simulations for Optimizations

Two sets of inverse analysis were performed. The first section presents the inverse

analysis when the Clayey Silt is modeled as a drained material and the instrument

accuracy is the one stated by the manufacturer. The second section presents the inverse

analysis when the Clayey Silt is modeled as an undrained material and the instrument

error is the observed error at the beginning of the Olive 8 project.

105

4.4.4. Optimization for Fully Drained Analysis

4.4.4.1. Optimized Parameters

From the parameters that define the responses of the Clayey Silt and Upper Sand

layer, only two were chosen for the optimization. These are the reference values for

primary deviatoric loading , and for elastic unloading and reloading , because

they are parameters that most influence the behavior of an excavation when the

movements are “small” (Finno and Calvello 2005). A sensitivity analysis was performed,

before the optimization, to define which of the previously mentioned parameters were

more susceptible to calibrate Table 4-9 presents the results of the sensitivity analysis.

refE50refurE

Composed Scaled Sensitivities

Parameter Silty Sand

Clayey Silt

ref50E 0.748 0.94 refurE 0.343 9.92

Table 4-9. Results of sensitivity analysis

Based on the composed scaled sensitivities, of the Clayey Silt layer is the most

important parameter to be calibrated from the inverse analysis. The of the Silty Sand

and Clayey Silt layer are one order of magnitude smaller the for the Clayey Silt.

The of the Silty Sand layer has the least influence on the computed results and

therefore it was not included in the optimization procedure because the composite scaled

refurE

refE50

refurE

refurE

106

sensitivity was relatively close to zero and this parameter would likely not be

optimized.

Two optimizations were performed using inverse techniques:

1) An analysis where three parameters were calibrated simultaneously, refE50 for the

Silty Sand layer, refE50 and refurE for the Clayey Silt layer.

2) An analysis where parameters were calibrated simultaneously, refE50 for the Upper

Sand layer and refE50 for the Clay layer.

The parameters that were not optimized did not change and remained at their input

values, or were related to the updated value of the optimized value. Specifically, when

not being explicitly optimized, the unload reload stiffness was calculated in

function of as:

refurE

refE50

refurE =3* (4-11) refE50

and the and the oedometer stiffness was, in all cases, computed as: refoedE

refoedE =0.7* (4-12) refE50

For the second optimization the most important parameter, of the Clayey Silt

layer, while not optimized, it was directly related to by a factor of 3.

refurE

refE50

107

4.4.4.2. Results

Three Optimized Parameters

Wall Deflection (in.)

-1.0 -0.5 0.0 0.5 1.0

Ele

vatio

n (fe

et)

60

70

80

90

100

110

120

130Observed Deflection Before Optimization 3 Par. Optimization

SIlty Sand

Clayey Silt

Very Dense Sand


parameters

Figure 4.14 shows for Pile #6, the calibrated displacement profile for the Three

108

Parameter Optimization versus the inclinometer observations and the numerically

calculated profile before the optimization. The calibrated displacement profile agrees

well with the observed profile at the Fill and Clayey Silt layers. At the Silty Sand Layer,

due to lack of reliable observations at greater depths, the fit is not as good compared with

the fit at the previous two soil layers.

Table 4.10 shows the results of the inverse analysis for three optimized parameters.

The initial values of (for the Upper Sand and Clayey Silt layer) and (for the

Clayey Silt layer) were the ones used in the GeoEngineers finite element analysis, given

in Table 4.1. The initial objective function value was 5132 and the final was 138 after 24

iterations. It must be mentioned that the optimization was terminated because for the

Clayey Silt layer overcame the finite element code limitation of ≤ 20* . The

results show that the reference stiffness increased by 38.5% for the sand layer and

161.2% for the Clay layer; the unload-reload stiffness for the Clay layer increased by

1700%. was increased with a rate much higher than the rate of and eventually,

after 24 iterations, became 20 times larger than so the optimization stopped

due to the finite element code limitation.

refE50refurE

refurE

refurE refE50

refurE refE50

refurE refE50

Figure 4.15 shows the objective function and the optimized stiffness parameters

versus the number of iterations. It can be seen that when the optimization was terminated,

by the finite element code limitation, the relative fit improvement (RFI) was 97% and the

objective function was stabilized, so the optimization was close to the optimum solution.

109

The for the Silty Sand and Clayey Silt layers were not greatly affected by the

optimization in contrast to of the Clayey Silt layer, which is the most important

optimization parameter.

refE50

refurE

Iteration Number

0 5 10 15 20

RFI

(%)

0

20

40

60

80

100Iteration Number

0 5 10 15 20

Obj

ectiv

e Fu

nctio

n

0

1000

2000

3000

4000

5000

6000

Iteration Number

0 5 10 15 20

kips

/ft2

0

5000

10000

15000

20000

25000

30000E50,SandE50,ClayEur,Clayref

ref

ref

Figure 4.15 Objective function, RFI (%) and stiffness parameters versus the number of

iterations

The optimization for three parameters showed that the importance of at the

Clayey Silt layer is significant in comparison with the rest of the optimized parameters.

refurE

110

However its very large value at the end is not representative of the Clayey Silt, and

likely reflects the 3-D stiffening effect of the corner of the excavation. This kind of

response was also noted by Finno and Calvello (2005).

3 Optimized Parameters Iteration 1 Iteration 24

Layer Parameter Estimated Parameter (kips/ft2)

Objective Function

Estimated Parameter (kips/ft2)

Objective Function

Change (%)

GeoEngineers Parameters

(kips/ft2)

Silty Sand

refE50 1000 1385.3 38.5 1000

refE50 500 1306.2 161.2 500 Clayey Silt ref

urE 1600

5132

28808

138

1700.5 1600

Table 4.10 Optimization for three parameters

Two Optimized Parameters

Figure 4.16 shows the calibrated displacement profile for the Two Parameter

Optimization, the inclinometer observations and the computed profile before the

optimization for Pile 6. Again, the calibrated displacement profile agrees with the

observed profile in the Fill and Clayey Silt. In the Silty Sand Layer, due to lack of

reliable observations at greater depths, the fit is not as good compared with the fit in the

other two soil layers.

111

Wall Deflection (in.)-1.0 -0.5 0.0 0.5 1.0

Elev

atio

n (fe

et)

60

70

80

90

100

110

120

130Observed Deflection 2 Par. Optimization Before Optimization

Silty Sand

Clayey Silt

Very Dense Sand


optimized parameters

Table 4.11 summarizes the results of the inverse analysis for two optimized

parameters. The initial values for the Upper Sand and the Clayey Silt layer were refE50

112

taken from the three parameter optimization. The rest of the parameters were related

to the updated value of the optimized value by equations (4-11) and (4-12). The initial

objective function value was 2690 and the final was 154 after 13 more iterations.

Figure 4.17 shows the objective function, the RFI (%) and the optimized parameters

versus the number of iterations. This optimization was stopped when the objective

function changed less than 8% for 3 consecutive iterations. The RFI stabilized at about

93% over these three last iterations. It has to be mentioned, that when the objective

function started to stabilize, the calculated deflections profile was slightly affected by the

extra number of iterations performed. The for the Silty Sand layer was not affected

greatly by the optimization. The optimization for two parameters showed that the

importance of at the Clayey Silt layer is significant. In this optimization though,

was not optimized directly but through according to the monotonic

relationship (4-11). Both parameters were changing during the optimization and this

explains the similar calculated soil profiles at the two and the three parameter

optimizations. However the very large values at the end are not representative of the

Clayey Silt, and likely reflect the 3-D stiffening effect of the corner of the excavation. It

must be noted that for the Clayey silt is one order of magnitude higher in the two

parameter optimization and this slightly affected the calibrated deflection profile.

refE50

refurE

refurE refE50

refE50

113

Iteration Number

0 2 4 6 8 10 12 14

Obj

ectiv

e Fu

nctio

n

0

500

1000

1500

2000

2500

3000

Iteration Number

0 2 4 6 8 10 12 14

kips

/ft2

0

3000

6000

9000

12000

15000E50,SandE50,Clayref

ref

Iteration Number

0 2 4 6 8 10 12 14

RFI

(%)

0

20

40

60

80

100

Figure 4.17 Objective function and stiffness parameters versus the number of iterations



Observed Function


Observed Function

Change (%)


(kips/ft2)

Silty Sand

refE50 1385 1150 -17.0 1000

Clayey Silt

refE50 1306 2690

12437 154

852.3 500

Table 4.11 Optimization for two parameters

114

Wall Deflection (in.)

-1.0 -0.5 0.0 0.5 1.0

Elev

atio

n (fe

et)

60

70

80

90

100

110

120

130Observed Deflection 3 parameter optimization 2 parameter optimization

Silty Sand

Clayey Silt

Very Dense Sand

Figure 4.18 Optimization results and wall deflection versus number of optimized

parameters

Figure 4.18 compares the inclinometer data with the computed deflections after the

two optimizations. The deflections in the Clayey Silt are almost identical for the two

115

optimizations and they are similar to the observed values. For the upper portion of

the wall the three parameter optimization gives slightly better results than the two

parameter optimization. This better fit is reflected in the lower final objective function

and the higher RFI for the Three Parameter Optimization. Hence, the three parameter

optimization gives a better solution, even though the simulation was terminated by the

finite element code.

Table 4.12 compares the optimized stiffness parameters for the two analyses with

their initial values.

Initial Stiffness Input Values versus the Values after the Three and Two Parameter Optimizations

3-par Optimization 2-par Optimization

Layer Parameter (kips/ft2)

Initial Value

Final Value

Initial Value

Final Value

refE50 1000 1385 1385 1150 refoedE 700 970 970 805 Silty

Sand refurE 3000 4156 4156 3450 refE50 500 1306 1306 12437 refoedE 350 914 914 8706 Clayey

Silt refurE 1600 28808 3919 37311

Table 4.12 Overview of Stiffness Input Parameters

The value in the Silty Sand increased by 38 and 11.5 % after the three and the

two parameter optimization, respectively, a reasonable result if ones considers that the

drained conditions represent the behavior of the Silty Sand layer and the 3-D cornering

refE50

116

effects are negligible. However, the stiffness parameters were optimized from

observations given by an inclinometer attached to the soldier pile.

The value in the Clayey Silt is the same order of magnitude in both

optimizations, but much larger than could be reasonably expected for this soil. The

is one order of magnitude higher in the 2 parameter optimization, than in the 3 parameter

optimization, because the most important parameter, , while optimized was directly

related to by a factor of 3.

refurE

refE50

refurE

refE50

An important factor is that the plane strain analysis conducted herein does not take

into account the 3-D stiffening effects near the corners of a deep excavation. This effect

implies that the optimized parameters in a plane strain analysis will be larger than those

that actually exist. This result was also noted by Finno and Calvello (2005) for the

Chicago-State excavation.

Finally, the stiffness parameters were optimized from observations given by an

inclinometer located right on the soldier pile. So the calibrated parameters are expected to

be greater than the actual, because they are affected by the stiffness of the soldier pile,

and not the smeared stiffness in a plane strain simulation. The slight kickback within the

Silty Sand layer would not have been measured in a inclinometer located about 6 ft

behind the soldier pile.

Figure 4.19 shows the computed shear strains (%) in the ground for displacements

induced by the excavation at pile 6. Herein, shear strain εq, refers to the invariant

117

deviatoric shear strains and is defined as:

( ) ( ) ( ) ([ ])222222 621

32

zxyzxyxxzzzzyyyyxxq εεεεεεεεεε +++−+−+−= (4-13)

In a plain strain simulation, this shear strain can be written as:

( ) ( ) ( ) ( )[ ]2222 621

32

xyxxyyyyxxq εεεεεε +++−= (4-14)

The strain components were obtained from the results of finite element calculations,

as those values that developed during excavation. Recall that displacements were zeroed

just before excavation begun. With the aid of Golden Software Surfer®, a contour plot of

the shear strain levels was generated.

The shear strains behind the Olive 8 shoring wall due to excavation are less than or

equal to 0.1%, with the exception of some localized shear strains (0.2 %) around the

bottom of the shoring wall and the lower portion of the Qwest basement wall. For this

excavation, the optimized stiffness parameters represent strains on the order of 0.1%.

Conventional laboratory soil testing without internal instrumentation allows one to

measure strains as low as 0.1%. If one does not consider the effect of the excavation of

the Qwest building, the secant modulus for the Silty Sand and the Clayey Silt layers only

can be determined only by lab tests that involve the use of research-type equipment

which is generally not available in commercial laboratories. Because no data were

available to compare computed with observed results for the Qwest building, only these

incremental strains can be evaluated. Successful calibrations thus would depend on

118

accurately defining the stresses in the ground by other means before excavating for

the Olive 8 building. (e.g. pressuremeter data to obtain “K0”)

Figure 4.19 Shear strains (%) in function with depth and distance from the West

excavation wall

119

4.4.5. Optimization for Undrained Responses of Clayey Silt

4.4.5.1. Optimized Parameters

As in the case of the drained analysis, two parameters that define the responses of

the Clayey Silt and Silty Sand were chosen for the optimization. These again were the

reference values for primary deviatoric loading , and for elastic unloading and

reloading . In this case the response of the Clayey Silt was assumed undrained. The

parameters are the same, but a penalty formulation is employed to enforce a condition of

no volume change in the layer. A sensitivity analysis was performed before the

optimization to define which of these parameters had the greatest impact on the

computed results. Table 4-8 presents the results of the sensitivity analysis.

refE50

refurE

Composed Scaled Sensitivities

Parameter Silty Sand

Clayey Silt

ref50E 0.380 3.18 refurE 0.802 9.85

Table 4-13. Results of sensitivity analysis

Based on the composed scaled sensitivities, of the Clayey Silt layer is the most

important parameter to be calibrated from the inverse analysis and of the same layer

is the second most important parameter. The of the Silty Sand layer is one order of

magnitude smaller than this value. The of the Silty Sand layer was the least influent

refurE

refE50

refurE

refE50

120

in the computed results, and was not included in the optimization procedure

because the composite scaled sensitivity was relatively close to zero and this parameter

would likely not be optimized.

Two optimizations were performed using inverse techniques:

3) An analysis where three parameters were calibrated simultaneously, refurE for the

Upper Sand layer, refE50 and refurE for the Clayey Silt.

4) An analysis where two parameters were calibrated simultaneously, refurE for the Upper

Sand layer and refurE for the Clay layer.

The parameters that were not optimized did not change and remained at their input

values, or were related to the updated value of the optimized value. Specifically, when

not being explicitly optimized, the reference stiffness was calculated in function of

as:

refE50

refurE

refur

ref EE ∗= 31

50 (4-15)

and the oedometer stiffness was, in all cases, computed as: refoedE

refrefoed EE 507.0 ∗= (4-16)

121

4.4.5.2. Results

Three Optimized Parameters

Displacement (in)

-0.5 0.0 0.5 1.0 1.5

Ele

vatio

n (ft

)

60

70

80

90

100

110

120

130

Observed DeflectionBefore OptimizationAfter Optimization

Fill

Silty Sand

Clayey Silt


parameters

122

Figure 4.20 shows the calibrated displacement profile for the Three Parameter


optimization for Pile 6. By visual inspection of the optimization results one can see that

the optimized profile does not fit with the observed.

Table 4.14 shows the results of the inverse analysis for three optimized parameters.

The initial values of (for the Upper Sand and Clayey Silt layer) and (for the

Clayey Silt layer) were the ones used in the GeoEngineers finite element analysis, given

in Table 4.1. The initial objective function value was 58317 and the last was 25273 after

9 iterations. The previous objective function values are higher than the ones found at the

optimizations for fully drained analysis. This is due to the fact that for this optimization

the observed error was used instead of the stated error. Since the observed error was

smaller then the weight of the observations derived from equation (4-9) increases and

consequently the objective function given by equation (4-3) increases also.

refE50refurE



Objective Function


Objective Function

Change (%)

Design Parameters

(kips/ft2)

Silty Sand

refurE 3000 1803 -39.9 3000

refE50 500 277 -44.6 500 Clayey Silt ref

urE 1600

58317

5794

25273

262.1 1600

Table 4.14 Optimization for three parameters

123

The optimization was terminated because for the Clayey Silt was larger

than that imposed by the finite element code limitation of ≤ 20* . Consequently,

this optimization was terminated at an early stage and this is the reason that the calibrated

displacement profile does agree well with the observed profile. The results show that the

decreased by 40% for the Silty Sand layer and increased 262% for the Clayey Silt;

the for the Clayey Silt layer decreased by 45%. increased rapidly while

decreased and eventually, after 9 iterations, became 20 times larger than so the

optimization stopped due to the finite element code limitation.

refurE

refurE refE50

refurE

refE50refurE refE50

refurE refE50

Figure 4.21 shows the objective function, the RFI (%) and the optimized stiffness

parameters versus the number of iterations. It can be seen that when the optimization was

stopped by the finite element code limitation, so the inverse analysis had not reached an

optimum solution defined by the convergence parameters.

124

Iteration Number

0 2 4 6 8 10

Obj

ectiv

e Fu

nctio

n

0

10000

20000

30000

40000

50000

60000

Iteration Number

0 2 4 6 8 10

kips

/ft2

0

2000

4000

6000

8000

10000Eur, SandE50, ClayEur, Clayref

ref

ref

Iteration Number

0 2 4 6 8 10

RFI

(%)

0

20

40

60

80

100


iterations

125

Two Optimized Parameters

Displacement (in)

-1.0 -0.5 0.0 0.5 1.0

Ele

vatio

n (ft

.)

60

70

80

90

100

110

120

130

Observed DeflectionBefore OptimizationAfter Optimization

Fill

Silty Sand

Clayey Silt


optimized parameters

. Figure 4.22 shows the calibrated displacement profile for the Two Parameter


126

optimization for Pile 6. The calibrated displacement profile agrees well with that

observed in the Fill and Clayey Silt layers. In the Silty Sand layer, the fit is not as good

compared with the fit at the other two soil layers.

Table 4.15 shows the results of the inverse analysis for two optimized parameters.

The initial values for the Silty Sand and the Clayey Silt layer were those in Table

4.1, used in the GeoEngineers finite element analysis. The rest of the stiffness parameters

were related to the updated value of the optimized value by equations (4-15) and (4-16).

The value of the objective function decreased from an initial value of 65612 to 4467 after

15 iterations.

refurE



Observed Function


Observed Function

Change (%)


(kips/ft2)

Silty Sand

refurE 3000 5319 77.3 3000

Clayey Silt

refurE 1600

65612 18840

4667 1077 1600

Table 4.15 Optimization for two parameters

Figure 4.23 shows the objective function, the RFI (%) and the optimized parameters

versus the number of iterations. This optimization was stopped when the objective

function changed less than 8% for three consecutive iterations. The RFI stabilized at

about 92% over these three last iterations. When the RFI stabilized, the calculated

deflection profile was affected little by the additional iterations performed. The value of

127

refurE for the Silty Sand layer gradually changed as the number of iterations

increased. The value of of the Clayey Silt layer increased rapidly as the number of

iterations increased.

refurE

Iteration Number

0 2 4 6 8 10 12 14

Obj

ectiv

e Fu

nctio

n

0

10000

20000

30000

40000

50000

60000

Iteration Number

0 2 4 6 8 10 12 14ki

ps/ft

20

5000

10000

15000

20000

25000

30000Eur, SandEur, Clayref

ref

Iteration Number

0 2 4 6 8 10 12 14

RFI

(%)

0

20

40

60

80

100


iterations

Figure 4.24 illustrates the updated profiles for two and three optimized parameters

versus the observed deflections. The profile given by the three parameter optimization

does agree particularly well with the observations because the calibration ended before an

128

optimum solution was reached. In the upper portions of the wall and the Clayey

Silt, the two parameter optimization provides a very good fit with the observed data in

contrast to that in the Silty Sand layer where not enough reliable observations existed to

perform the optimization.

Displacement (in)

-1.0 -0.5 0.0 0.5 1.0

Ele

vatio

n (ft

)

60

70

80

90

100

110

120

130

Observed Deflection2 par. Optimization3-par Optimization

Fill

Silty Sand

Clayey Silt

Figure 4.24 Observed versus calibrated displacement deflection profiles for two and

three optimized parameters

129

Table 4.16 presents the calibrated parameters from the two and the three

parameter optimizations assuming undrained conditions in the Clayey Silt and their initial

values. It can be seen that the results from the 3 parameter optimization differ greatly

from the results of the 2 parameter since the optimization could not finish. Because the fit

was much better in the later, hereafter the only the results of the two parameter

optimization will be discussed.

Comparison of 2 and 3 Parameter Optimization for Undrained Clayey Silt Analysis

3-par Opt.

2-par Opt.


Initial Value

Final Value

Final Value

refE50 1000 601 1713 refoedE 700 421 1199 Silty

Sand refurE 3000 1803 5139 refE50 500 277 6280 refoedE 350 194 4396 Clayey

Silt refurE 1600 5794 18840

Table 4.16 Overview of Stiffness Input Parameters

The value in the Silty Sand increased by 71% after the two parameter

optimization, a reasonable result if ones considers that the drained conditions represent

the behavior of the Silty Sand layer and the 3-D cornering effects are negligible.

However, the stiffness parameters were optimized from observations given by an

inclinometer attached to the soldier pile.

refE50

130

The value in the Clayey Silt increased by an order of magnitude but is

much larger than could be reasonably expected for this soil. The plane strain analysis

conducted herein does not take into account the 3-D stiffening effects near the corners of

a deep excavation. This effect implies that the optimized parameters in a plane strain

analysis will be larger than those that actually exist. This result was also noted by Finno

and Calvello (2005) for the Chicago-State excavation.

refurE

Finally, the stiffness parameters were optimized from observations given by an

inclinometer attached at the soldier pile. So the calibrated parameters are expected to be

greater than the actual, because they are affected by the stiffness of the soldier pile, and

not the smeared stiffness in a plane strain simulation. The slight kickback within the Silty

Sand layer would not have been measured in an inclinometer located about 6 ft behind

the soldier pile.

131

Displacement (in)

-1.0 -0.5 0.0 0.5 1.0

Ele

vatio

n (ft

)

60

70

80

90

100

110

120

130

Observed DeflectionUndrained Clayey SiltDrained Clayey Silt

Fill

Silty Sand

Clayey Silt

Figure 4.25 two parameter optimization results for drained and undrained Clayey Silt

analysis

Figure 4.25 presents the observed deflections and the calculated displacement

profiles for both two parameter optimizations. It can be seen that the calibrated profiles

132

are similar in the fill but differ slightly in the Silty Sand and Clayey Silt. In the Silty

Sand layer, when the observed error was used instead of the stated error (undrained

Clayey Silt analysis) the calculated displacement profile fitted better to the observed data.

The later is due to the fact that the number of reliable observations increased.

At the Clayey Silt layer the observed profile lies between the profiles given from the

drained and undrained analysis. For the undrained Clayey Silt response, the difference

from the observed profile is smaller. Since the number of observations used for both

optimizations is almost the same (15 instead of 14 when using the stated error), the better

fit is caused due to the undrained material response and the smaller measurement error

that produced a higher weight for each observation.

Table 4.17 presents an overview of the stiffness parameters for the two optimized

parameter calibrations when the Clayey Silt is modeled as a drained and undrained

material. For the Silty Sand layer the stiffness parameters were increased by 50% for the

later optimization. Since the observed error was used instead of the stated error, the

number of reliable observations increased and this resulted to a better fit of the optimized

to the observed profile. As mentioned in Section 4.3.1, the 3-D effects are negligible for

the Silty Sand layer; however the stiffness parameters were optimized from observations

given by an inclinometer attached on the soldier pile. So the calibrated parameters are

expected to be greater than the actual, because they encounter also for the stiffness of the

soldier which is smeared in a finite element simulation.

133

Comparison of 2 Parameter Optimizations for Drained and Undrained Clayey Silt Analysis

Drained Undrained


Initial Value

Final Value

Final Value

refE50 1000 1150 1713 refoedE 700 805 1199 Silty

Sand refurE 3000 3450 5139 refE50 500 12437 6280 refoedE 350 8706 4396 Clayey

Silt refurE 1500 37311 18840

Table 4.17 Overview of Stiffness Parameters

For the Clayey Silt layer, it can be seen that after using undrained analysis the

optimized stiffness parameters for the Clayey Silt reduced by 50%. Drained analysis

resulted long term wall deflections and the soil stiffness parameters had to increase more

to fit the calculated with the observed deep-seated displacements. The numerical

simulation of the behavior of this layer affects greatly the optimization results. However,

these calibrated stiffness parameters for the Clayey Silt do not reflect the actual

parameters that one would expect after high quality laboratory testing. The plane strain

analysis conducted herein does not take into account the apparent 3-D effects that

increase the stiffness at the corners of a deep excavation. This means that the optimized

parameters in the plane strain analysis will be larger than those that actually exist. This

result was also noted by Finno and Calvello (2005) for the Chicago-State excavation.

Finally, as previously mentioned, the optimized stiffness parameters are also affected by

134

the location of the inclinometer.

4.4.6. Discussion

The selection of observations for the inverse analyses showed that when the stated

error is used, most of the observed displacements at the Silty Sand layer are lower than

95 % of the standard deviation of the instrument error. Measurements like these are not

reliable to be used in inverse analysis because they created uncertainties in the

optimization results. If one compares the optimized and the observed profile in the Silty

Sand layer, then the lack of reliable observations created the difference between these

two displacement profiles. The manufacturer measurement error is relatively high for

conventional inclinometers and at projects where the predicted movements are small,

more accurate surveying instruments should be used, or site specific measurements

should form the basis of error estimates.

When the observed instead of the stated error was used, the number of reliable

observations increased at the Silty Sand layer and the optimization produced a better fit

for the calculated displacement profile. At the Clayey Silt layer the number of

observations used remained almost the same; however the calibrated profile slightly

improved since the observed error at this depth was very small and the weight of these

observations was increased.

Simulating the Clayey Silt as an undrained instead of drained material produced, as

expected, smaller values for the optimized parameters at this layer. However, still these

135

parameters do not reflect the actual parameters that one would expect after

extensive laboratory testing and are influenced by 3-D stiffening effects and by the

stiffness of the soldier pile which is smeared in plane strain analysis.

The shear strains behind the Olive 8 shoring wall are less than or equal to 0.1%, thus

the secant modulus can be determined only by lab tests that involve the use of research-

type equipment which is not available in commercial laboratories.

4.5. Conclusions

1) With the same set of soil parameters, the finite element numerical analysis that

included the Qwest excavation produced a displacement profile that was closer to the

movements observed in field, than a similar analysis without an explicit simulation of

the Qwest excavation.

2) When the changes in soil stresses were taken into account, the relative shear stresses,

before the Olive 8 excavation, increased behind the West excavation wall. The

calculated movements were smaller due to the soil stress relief after the Qwest

excavation.

3) Using undrained instead of drained analysis for the Clayey Silt layer resulted in

smaller computed deflection profiles.

4) The 3-D stiffening effects at the corners of the excavation are negligible in the Silty

Sand layer. In the Clayey Silt layer the 3-D effects become important especially when

the excavation reaches full depth.

136

5) The deep-seated deflections in the Clayey Silt layer are over-predicted in plane

strain analysis due to 3-D effects that increase the stiffness at the corners of a deep

excavation.

6) The observed profiles based on inclinometers attached to the soldier piles are

influenced by the stiffness of the soldier piles. In plane strain finite element analysis,

the stiffness of the soldier pile wall is smeared when dividing with the out-of-plane

spacing and behavior at the wall will be different than behind the wall.

7) Little difference was observed in computed displacement profiles when the tiebacks

were represented as equivalent struts or equivalent anchors.

8) The selection of observations for the inverse analysis showed that most of the

observed displacements at the Silty Sand layer are lower than 95 % of the standard

deviation of the manufacturer instrument error.

9) The manufacturer measurement error is relatively high for conventional inclinometers

and at projects where the predicted movements are small, more accurate surveying

instruments should be used or should be calculated from site specific measurements.

10) When optimizing the Clayey Silt layer as an undrained material the calibrated

parameters were smaller than the parameters produced when this layer was modeled

as a drained material.

11) The low shear strains behind the Olive 8 shoring wall indicate that the secant modulus

only can be determined by lab tests that involve the use of research-type equipment

137

with on specimen instrumentation which is generally not available in

commercial laboratories.

12) 3-D effects were negligible in for the Silty Sand layer and the optimized parameters

are affected by the location of the inclinometer which was next to the soldier pile. The

stiffness of the soldier pile affects the observations.

13) The calibrated stiffness input parameters for the Clayey Silt layer do not reflect the

actual parameters that one would expect after extensive laboratory testing because of

the 3-D stiffening effects.

138 Chapter 5 5. Summary and Conclusions

The Olive 8 development in Seattle, WA, included a 70 feet deep excavation to

provide space for 5 underground parking levels. The shoring system was rather unique

along the West shoring wall, due to the proximity of the Qwest building. The maximum

allowed displacement imposed by the city of Seattle was 1 inch.

GeoEngineers Inc., the project consultants, designed the West wall shoring system,

which consisted of a soldier pile wall with shotcrete lagging, 9 to 10 rows of soil nails

and 4 or 5 rows of tiebacks. Three separate sections were designed to provide lateral

support for the West wall, due to the presence of a large utility vault between the

excavation and the Qwest building. GeoEngineers performed numerical analysis to

predict the wall deflections which were found to be higher than the 1 inch limiting

displacement. However they proceeded with the proposed design because shoring wall

and the soil layers were expected to behave stiffer than the numerically-derived

predictions.


West shoring wall. It included three conventional inclinometers to measure the lateral

wall movements and an automated surveying system to monitor the 3D deflections of the

top of the West wall at the positions where the inclinometers were located. The wall

movements were calculated relative to reference baseline points that were established far

away from the excavation. Rigid body translation and translation corrections were

performed to account for the fact that the total station was not rigidly attached to the

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parapet wall where it was mounted. The lateral displacements based on the automated

surveying data of prisms agreed well with the inclinometer data. The less than rigid

connection between the total station and the parapet throughout the construction caused

the total station to shut down numerous times during the excavation.

To supplement GeoEngineers finite element analysis, a numerical model was

presented that explicitly accounted for the changes in stresses history caused by

construction of the Qwest building. Separate simulations were performed with the Clayey

Silt modeled as an undrained material and the tiebacks represented as equivalent struts.

The 3-D stiffening effects at the corners of deep excavations were evaluated for this

specific case history.

The finite element model that simulated the Qwest excavation was used in an

inverse analysis wherein the observed and the computed deflections are compared to find

the soil parameters that provided a best fit. A significant question for this excavation case

history was whether the small observed displacements in the Silty Sand layer were large

enough to be used for the optimization procedure.

The conclusions drawn from the excavation monitoring performance are:

1) The inclinometer observations showed the 1 inch movement limitation specified

by the city of Seattle was not exceeded. In particular, the maximum observed

movement was half of this limitation.

2) The horizontal wall deflections towards and along the excavation based on the

total station agreed well with the movement with the movement of the top of the

wall based on inclinometer data.

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3) The Total Station shut down because the working range exceeded the vertical

compensator’s angular limit due to the less than rigid connection with the hotel’s

parapet wall.

4) The rigid body rotation correction requires all axes to be perpendicular to each

other. When the vertical compensator was turned off, the vertical axis was not

perpendicular to the horizontal plane, resulting in unacceptably large scatter in the

settlement data.

5) The design finite element predictions of the magnitude of wall deflection were

larger than that observed. But the observed and predicted displacements, along the

length of the wall, have the same pattern, implying that the shoring wall and the

soil responded more stiffly than expected.

The conclusions drawn from the finite element analysis conducted herein are:

1) With the same set of soil parameters, the finite element numerical analysis that

included the Qwest excavation produced a displacement profile that was closer to the

movements observed in field compared to those from a similar analysis without an

explicit simulation of the Qwest excavation.

2) Using undrained instead of drained analysis for the Clayey Silt layer resulted in

smaller computed deflection profiles.

3) The 3-D stiffening effects at the corners of the excavation are negligible in the Silty

Sand layer. In the Clayey Silt layer, the 3-D effects become important especially

when the excavation reaches full depth. The computed “deep-seated” deflections in

the Clayey Silt layer are larger than observed due to 3-D effects that increase the

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stiffness at the corners of a deep excavation.

4) The observed profiles based on inclinometers attached to the soldier piles are

influenced by the stiffness of the soldier piles. In a plane strain finite element

analysis, the stiffness of the soldier pile wall is smeared when dividing with the out-

of-plane spacing and lateral movements of the soldier pile wall differ than these

between the piles at the wall. For inverse analysis purposes of this type wall system,

the inclinometer should be located several feet behind the wall so the localized effects

do not impact these results.

5) Little difference was observed in computed displacement profiles when the tiebacks

were represented as equivalent struts or equivalent anchors.

The conclusions drawn from the inverse analysis results are:

1) The selection of observations for the inverse analysis showed that most of the

observed displacements at the Silty Sand layer are lower than 95 % of the standard

deviation of the manufacturer instrument error.

2) The manufacturer measurement error is relatively high for conventional inclinometers

and at projects where the predicted movements are small, more accurate surveying

instruments should be used or should be calculated from site specific measurements.

3) When optimizing the Clayey Silt layer as an undrained material, the calibrated

parameters were smaller than the parameters produced when this layer was modeled

as a drained material.

4) The low shear strains behind the Olive 8 shoring wall indicate that the secant modulus

only can be determined by lab tests that involve the use of research-type equipment

142

with on-specimen instrumentation which is generally not available in commercial

laboratories.

5) The calibrated stiffness input parameters for the Clayey Silt layer should not reflect

the actual parameters that one would expect after extensive laboratory testing because

of the 3-D stiffening effects.

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REFERENCES Blackburn, J.T. (2005). “Automated Sensing and Three-dimensional Analysis of

Internally Braced Excavations”, Ph.D. thesis, Civil and Environmental Engineering Department, Northwestern University, Evanston, Illinois.

Calvello, M. (2002). “Inverse analysis of a supported excavation through Chicago glacial

clays”, Ph.D. thesis, Civil and Environmental Engineering Department, Northwestern University, Evanston, Illinois.

Calvello, M. and Finno, R.J. (2004). “Selecting Parameters to Optimize in Model

Calibration by Inverse Analysis,” Computers and Geotechnics, Elsevier, v 31, n 5, p 411-425.

Calvello, M. and Finno, R.J. (2003). “Modeling Excavations in Urban Areas: Effects of

Past Activities,” Italian Geotechnical Journal, v 37, n 4, p. 9-23. Clough G.W. and O’Rourke T.D. (1990). “Construction Induced Movements of in-situ

Walls”. Design and Performance of Earth Retaining Structures, Proceedings of a Specialty Conference at Cornell University, ASCE, New York, 439-470.

Clough, G.W., Smith, E.M. and Sweeney, B.P. (1989). “Movement Control of

Excavation Support Systems by Iterative Design”, Current Principles and Practices, Foundation Engineering Congress, Vol. 2, ASCE, p. 869-884.

Dames And Moore (1972). “Supplemental Analysis Proposed Multi-Story Office

Building, Seventh Avenue and Pine Street, Seattle Washington”. Finno, R.J., Bryson, S. and Calvello, M. (2002). “Performance of a stiff support system in

soft clay”, Journal of Geotechnical and Geoenvironmental Engineering, v 128, n 8, p 660-671.

Finno, R.J., Blackburn, J.T. and Roboski, J.F.(2007), “Three-dimensional Effects for

Supported Excavations in Clay,” Journal of Geotechnical and Geoenvironmental Engineering, ASCE, v 133, n 1, p 30-36

Finno R.J. and Calvello, M. (2005). “Supported excavations: Observational method and

inverse modeling”, Journal of Geotechnical and Geoenvironmental Engineering, v 131, n 7, p 826-836.

Finno, R.J. and Roboski J.F. (2005), “Three-Dimensional Responses of a Tied-back

Excavation Through Clay”, Journal of Geotechnical and Geoenvironmental Engineering, v 131, n 3, p 273-282.

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GeoEngineers Inc. (2005), “Geotechnical Engineering Shoring Design Report for the West Wall, Office Tower 8th Avenue and Olive Way, Seattle, Washington”.

HartCrowser Inc. (2000), “Geotechnical Engineering Design Study, Proposed Office Tower 8th Avenue and Olive Way, Seattle, Washington”.

Leica Geosystems (2002), User Application Manual v2.2, Leica Geosystems AG,

Heerbrugg, Switzerland. McMillan, L. (1996), “Rigid Body Transformations”, Class at UNC, Introduction to

computer graphics (Comp136) , Lecture 15. Plaxis (2002). 2D-Version 8 User Manual, Editor R.B.J. Brinkgreve, A.a. Balkems

Publidhers, Lisse. Poeter E. and Hill M. (1998). “UCODE a Computer Code for Universal Inverse

Modeling”, U.S. Geological Survey. Rechea, C.B. (2005), “Inverse Analysis of Excavations in Urban Environments” Ph.D.

thesis, Civil and Envir. Engrg. Dept., Northwestern University, Evanston, Illinois. Roboski, J.F. (2004), “Three-dimensional Performance and Analyses of Deep

Excavations” Ph.D. thesis, Civil and Envir. Engrg. Dept., Northwestern University, Evanston, Illinois.

Roboski, J.F. and Finno, R.J., (2006) “Distributions of Ground Movements Parallel to

Deep Excavations,” Canadian Geotechnical Journal, v 43 n 1, p 43-58.

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NORTHWESTERN UNIVERSITY Automated …iti.northwestern.edu/publications/finno/Langousis-2007-Automated...inclinometer data. In addition, it is illustrated how soil stress changes due