EXPERIMENTAL AND ANALYTICAL STUDY OF · PDF fileEXPERIMENTAL AND ANALYTICAL STUDY OF REINFORCED CONCRETE COLUMN AND CORE SHORTENING IN A TALL BUILDING. By SADUDEE BOONLUALOAH A Thesis

EXPERIMENTAL AND ANALYTICAL STUDY OF REINFORCED CONCRETE

COLUMN AND CORE SHORTENING IN A TALL BUILDING.

By

SADUDEE BOONLUALOAH

A Thesis submitted in fulfilment of

the requirements of the degree of Master of Philosophy

Griffith School of Engineering

Griffith University, Gold Coast, Australia

March, 2010

Abstract

In tall concrete buildings, columns and core walls experience axial shortening due to

combined effect of heavy loading, shrinkage and creep. If the stress distribution

across the floor plan is disproportionate, columns and core walls within a single storey

may shorten by different amounts. This differential can introduce extra loads and

moments within elements and also induce distortion of attached service elements such

as claddings, pipes, partitions, that leads to serviceability problems.

A number of columns and core walls in a 80-storey building (Q1 Tower, Gold Coast,

Australia) were instrumented during its construction. Axial shortening data of these

vertical elements were collected during 3 year construction period. Regular

measurement of shortening of columns/core walls was undertaken using a manual

DEMEC measuring device in a systematic manner over pre-determined time periods.

Laboratory tests were also conducted on the concrete delivered to the construction

site. The properties examined (using standard test methods) include strength, elastic

modulus, shrinkage and creep of the concrete.

A comprehensive study on the prediction models of concrete properties related to

axial shortening was conducted. The performance of major code models for prediction

of the tested concrete properties was evaluated. The laboratory tested concrete results

were compared against predicted values yielded by each model and the most accurate

and reliable model was identified. In some cases, these identified methods are

modified or regression lines are derived from test data and used instead to ensure

optimum prediction accuracy.

An axial shortening estimation method was developed. The method which

incorporates a construction sequence simulation based on actual construction history

of the Q1 Tower and several material models for the estimation of time dependent

properties of concrete including a varity of code material models and a combination

of modified/fitted models was used in the well accepted Age Adjusted Effective

Modulus Method (AEMM) formulation to produce predicted value of axial

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shortening. Comparisons of measured and predicted axial shortenings of the building

were then made. Performance of the adopted axial shortening prediction method when

incorporated with different material models was also assessed. The best model

combination was then identified. Sensitivity analysis of factors affecting axial

shortening prediction was also conducted.

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Statement of Originality

I hereby certify that the work embodied in this thesis is the result of original research

and has not been submitted for a higher degree to any other University or Institution,

and to the best of my knowledge contains no material previously published or written

by other person except where acknowledgment is given and the relevant author cited.

Signed

Sadudee Boonlualoah

iv

Acknowledgements

This research was a cooperative research project between CEMEX Australia and

School of Engineering, Griffith University, Gold Coast Campus, Gold Coast,

Australia under R&D Project RD814. The findings, discussions and conclusion

expressed in this thesis are those of the author and not necessarily of sponsoring

agency.

First of all, I would like to express my thanks to both of my Principle Supervisors,

Professor Yew Chaye Loo (Foundation Chair of Civil Engineering, Director,

Internationalisation and Professional Liaison Science, Environment, Engineering and

Technology Group, Griffith University) and Associated Professor Salvatore (Sam)

Fragomeni (former Head of Griffith School of Engineering, now Head of School of

Architecture, Civil and Mechanical Engineering at Victoria University), without their

guidance and assistance this research would not have been possible. I also would like

to express thanks to Dr.Daksh Baweja, Principal Engineer of CEMEX Australia for

his sponsorships and kind support throughout this project. Thanks also have been

made to my external supervisor, Associated Professor Prasert Suwanvitaya of

Kasetsart University, Bangkok, Thailand and my co-supervisor, Dr. Sanual

Chowdhury (Senior Lecturer, Griffith School of Engineering) for their support during

the duration of my remote candidature.

Particular thanks are due to staff members at CEMEX Technical Services (Beenleigh)

and CEMEX concrete plant at Southport for their assistance on sampling and testing

of concrete specimens. Grateful acknowledgement is also made to the followings

Griffith University’s student research team, Ben Clark, Tim Lancastor, Christ

Dolinko, Aaron Pianta, for their contribution in data collection on site. Thanks are

also due to Sunland Group for site access permission during the duration of in-situ

data collection at the Q1 Tower. Information provided by Mr.Narayan Baida (a master

graduate with the University of Melbourne) and Ms Sonia Bursle (a master graduate

with the University of New South Wale) are also acknowledged. Special thanks are

made to my colleague at School of Engineering research group.

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Last but not least, thanks are due to my family for their endless support. Their

unconditional support which covers a much larger scope than this project has been

most valuable and inspiring.

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List of Publications

The technical papers that have been composed based on this research are list as

follow:

1. S. Boonlualoah, S. Fragomeni, Y.C. Loo and D. Baweja, “Monitoring

Deformations of Vertical Concrete Members in Tall Building”. Presented at the

fourth International Conference on Concrete under Severe Conditions:

Environment & Loading 27th

-30th

June 2004 in Seoul, CONSEC04 Proceedings,

Seoul National University, Korea, (2004).

2. S. Boonlualoah, S. Fragomeni, Y.C. Loo and D. Baweja, “Aspects of Axial

Shortening of High Strength Concrete Vertical Elements in a Tall Concrete

Building”, ACMSM-18 Proceedings, University of Western Australia, Perth,

(2004).

3. S. Boonlualoah, S. Fragomeni, Y.C. Loo and D. Baweja, “Differential

Deformations of Columns and Cores in Tall Buildings – Assessment, Monitoring

and Correction Measures”, ICTB-VI 6th International Conference on Tall

Buildings Proceedings, University of Hong Kong, Hong Kong, China. (2005).

vii

Notations

a………………material specific constants.Used in creep equation of Thomas (1930).

a/c……………..ratio of aggregate to cement. Used in B3 model.

A ……………..material constant. Used in creep equation of Ross (1937).

A1……………..material specific constants, Used in creep equation of Straub (1930).

A2……………..material specific constants.Used in creep equation of Thomas (1930).

A3……………..material specific constants. Used in shrinkage equation of Bazant and

Panula (1978).

A(t0)…………..specific coefficient of concrete and age at loading. Used in creep

equation of the US Bureau of Reclamation (1956).

B ……………...material constant. Used in creep equation of Ross (1937).

B1……………..material specific constants, Used in creep equation of Straub (1930).

B2……………..material specific constants. Used in shrinkage equation of Bazant and

Panula (1978).

BC …………...basic creep (x10-5

cm2/kgf). Used in creep equation of Sakata (1993).

c……………….constant. Used in creep equation of Bazant et al (1976).

c(t,s)…………..creep strain at time t of a concrete member loaded first time at time s.

C………………cement content, in kg/m3. Used in creep equation of Sakata (1993).

Cb……………..basic creep strain, in .

Cd……………..drying creep strain, in .

Ct……………...total creep strain, in .

Cu……………..ultimate creep coefficient.

C2……………..material specific constants.Used in creep equation of Thomas (1930).

C3……………..material specific constants. Used in shrinkage equation of

Aroutiounian (1959).

C(t,s)………….specific creep, in /MPa.

C(t,t0)…………specific creep. Used in creep equation of GL2000 model.

C(t,t ,t0)……….specific creep (x10-5

cm2/kgf). Used in creep equation of Sakata

(1993).

Cd(t,t ,t0)………drying creep. Used in creep equation of B3 model.

C0(t,t )…………basic creep. Used in creep equation of B3 model.

viii

d……………….constant. Used in creep equation of Bazant et al (1976).

DC ……………drying creep (x10-5

cm2/kgf). Used in creep equation of Sakata

(1993).

E………………..elastic modulus, in MPa.

Ec……………….elastic modulus of concrete.

Ecm28……………elastic modulus at 28 days, in MPa. Used in creep equation of

GL2000 model.

Ecmto……………elastic modulus at time of loading, in MPa. Used in creep equation

of GL2000 model.

Ee………………effective modulus. Used in EM method.

E(t0)……………elastic modulus at the application of the loading. Used in EM

method.

Ec(s)……………elastic modulus of concrete at the initial application of a load.

Ec(t)……………elastic modulus of concrete at time t.

Ec(s)……………elastic modulus of concrete at first loading. Used in AEMM method.

fc……………….actual compressive strength at 28days, in MPa. fck28……………28-days characteristic strength. Used in GL2000 model.

fcm………………mean compressive strength, in MPa.

fcm28……………28-days mean compressive strength, in MPa.

fcu………………characteristic concrete cube strength, in MPa.

f’c………………characteristic compressive strength, in MPa.

f’ca……………..adjusted 28-day compressive strength of concretes where 50 MPa≤

f’c ≤ 80 MPa.

f’c28……………28-days characteristic compressive strength, in MPa.

f’c (t)…………...compressive strength at time t, in days. G1………………applied load at a strain of 50 microstrain divided by the cross-

sectional area of the unloaded specimen, in MPa. Used in AS1012-

17.

G2……………….test load divided by the cross-sectional area of the unloaded

specimen, in MPa. Used in AS1012-17.

G2……………….applied stress, in MPa.

h………………....humidity expressed as a decimal (h≤1). Used in shrinkage equation

of GL2000 model.

h…………………effective thickness of the concrete member. Used in shrinkage

ix

equation of CEB-FIP.

h…………………notational size of member, in mm. Used in creep equation of CEB-

FIP.

h…………………percent relative humidity. Used in shrinkage equation of B3 model.

h…………………relative humidity (0≤h≤1). Used in creep equation of B3 model.

Η…………………correction factor accounts for the “additional strength”.

J(t,t0)…………….a compliance function. Used in creep equation of GL2000 model.

kh………………..humidity dependence. Used in shrinkage equation of B3 model.

k1………………..environmental factor. Used in shrinkage equation of AS3600-2001.

k1………………..size factor. Used in shrinkage equation of AS3600-2009.

k2………………..creep-time curve. Used in creep equation of AS3600-2001.

k2………………..factor accounts for the influences of the hypothetical thickness, the

environment and the time after loading to creep. Used in creep

equation of AS3600-2001.

k2………………..factor of the creep-time curve. Used in creep equation of Gilbert

(2002).

k3………………..maturity coefficient. Used in creep equation of AS3600-2001.

k3………………..function of the strength ratio depends on the age at first loading.

Used in creep equation of AS3600-2001.

k3………………..maturity coefficient. Used in creep equation of AS3600-2009.

k4………………..factor accounts for different environmental conditions. Used in

shrinkage equation of AS3600-2009.

k4………………..factor accounting for the environment. Used in creep equation of

AS3600-2009.

k5………………..factor accounting for the reduction of the influence of humidity and

specimen size on creep when there is an increase of concrete

strength. Used in creep equation of AS3600-2009.

K………………...additional constants. Used in creep equation of McHenry (1943).

K…………………factor accounting for cement types specified as, 1 for type I, 0.7

for type II and 1.15 for type III. Used in shrinkage equation of

GL2000 model.

m…………………constant. Used in creep equation of Lorman (1940).

m ………………..additional constants. Used in creep equation of McHenry (1943).

x

m ……………….material specific constants. Used in shrinkage equation of Bazant

and Panula (1978).

n…………………constant. Used in creep equation of Lorman (1940).

n………………….number of samples.

ne(s)………………time dependent parameter of modular ratio n at time s. Used in

AEMM method.

ne(t)………………time dependent parameters of modular ratio n.

‾ne(t)……………...time dependent parameter.

n‾e(t)……………..time dependent parameters of modular ratio n.

p………………….percentage reinforcement. Used in AEMM method.

p …………………additional constants. Used in creep equation of McHenry (1943).

P………………….applied load. Used in AEMM method.

q2…………………ageing viscoelastic compliance. Used in creep equation of B3

model.

q3…………………non-ageing viscoelastic compliance. Used in creep equation of B3

model.

q4…………………flow compliance. Used in creep equation of B3 model.

Q(t,t’)…………….binomial integral component. Used in creep equation of B3 model.

r…………………...material specific constants. Used in creep equation of McHenry

(1943).

R2………………...coefficient of determination.

RH………………..percent relative humidity. Used in shrinkage equation of Sakata

(1993).

RH………………..relative humidity (%). Used in creep equation of Sakata (1993).

R(t,s)……………..the relaxation function. Used in AEMM method.

s…………………..standard deviation of the compressive strength result.

S(t)………………..time curve of the shrinkage. Used in shrinkage equation of B3

model.

S(t)……………….shrinkage time curve. Used in creep equation of B3 model.

t…………………..time after pouring of concrete, in days. Used in shrinkage equation

of AS3600-2009.

t…………………...age of concrete, in days. Used in creep equation of Sakata (1993).

t …………………..age of concrete at loading, in days. Used in creep equation of

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Sakata (1993).

tc…………………..age of concrete when drying started, in days. Used in shrinkage

equation of GL2000 model.

ts…………………..age of concrete at the commencement of shrinkage. Used in

shrinkage equation of CEB-FIP.

t0….........................age of concrete at time of loading, in days. Used in creep equation

of GL2000 model.

t0………………….age of concrete when drying commenced in days. Used in

shrinkage equation of Sakata (1993).

t0………………….age of concrete when drying commenced, in days. Used in creep

equation of Sakata (1993).

v/s…………………volume-surface ratio in inches.

V/S………………..volume-to-surface ratio. Used in shrinkage equation of Sakata

(1993).

V/S………………..volume-to-surface ratio, in cm. Used in creep equation of Sakata

(1993).

V/S………………..volume-to-surface ratio, in mm. Used in shrinkage equation of

GL2000 model.

w………………….water content of the concrete in kg/m3. Used in shrinkage

equation of Sakata (1993).

w………………….concrete density, in kg/m3.

w………………….water content of concrete, in lb/ft3. Used in shrinkage equation of

B3 model.

w/c………………...ratio of water to cementitious material. Used in B3 model.

W………………….water content, in kg/m3. Used in creep equation of Sakata (1993).

x …………………material specific constants.Used in creep equation of Thomas

(1930).

x…………………..drying time, in days.

y…………………..shrinkage, in .

(t,s)……………...creep function

(t,t )……………..creep function at a concrete age t for a load applied at time t .

Used in Superposition method.

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(tc)………………accounting for the Pickett effects or the drying before loading,

which reduced both basic and drying creep, thus it is taken as

1 if t0 = tc. Used in creep equation of GL2000 model.

α1………………….coefficient depending on type of cement. Used in shrinkage

equation of B3 model.

α2………………….coefficient depending on curing condition. Used in shrinkage

equation of B3 model.

2…………………material specific constants. Used in creep equation of McHenry

(1943).

β ………………….additional constants. Used in creep equation of McHenry (1943).

βsc…………………coefficient depending on type of cement. Used in shrinkage


βRH………………coefficient for relative humidity. Used in shrinkage equation of

CEB-FIP.

c(t-t0)…………….development of creep with time. Used in creep equation of CEB-

FIP.

βs(t-ts)…………….development of shrinkage with time. Used in shrinkage equation

of CEB-FIP.

(fcm)……………..coefficient for concrete strength; Used in creep equation of CEB-

FIP.

β(h)………………..coefficient accounting for the effect of humidity on shrinkage.

Used in shrinkage equation of GL2000 model.

β(t)………………...coefficient accounting for the effect of time on shrinkage. Used

in shrinkage equation of GL2000 model.

(t0)……………….coefficient for the age at loading. Used in creep equation of

CEB-FIP.

χ(t,s)………………aging coefficient. Used in AEMM method.

εcs……....................design shrinkage strain. Used in shrinkage equation of AS3600-

2001.

εcs…………………total shrinkage. Used in shrinkage equation of AS3600-2009.

εcsd………………..drying shrinkage. Used in shrinkage equation of AS3600-2009.

εcse………………...endogenous shrinkage. Used in shrinkage equation of Gilbert

xiii

(2002).

εcso………………...notional or ultimate shrinkage strain, in . Used in shrinkage


cs.b………………..ultimate shrinkage strain. Used in shrinkage equation of AS3600-

2001.

csd.b……………….basic drying shrinkage strain, in . Used in shrinkage equation

of AS3600-2009.

εsp………………….specific elastic strain occurring at the initial application of a load

εsh …………………shrinkage strain, in .

sh ………………...basic shrinkage strain, in .

εe(t)…………………instantaneous strain at time t.

εel(s)………………...elastic strain occurring at the initial application of a load.

cse…………………final endogenous shrinkage, in . Used in shrinkage equation

of AS3600-2009.

csd.b ……………….ultimate shrinkage strain, in . Used in shrinkage equation of

AS3600-2009.

εshu…………………ultimate shrinkage strain, in . Used in shrinkage equation of

GL2000 model.

sh …………………ultimate shrinkage strain. Used in shrinkage equation of Sakata

(1993).

sh …………………ultimate shrinkage strain (mm/mm). Used in shrinkage equation

of B3 model.

sh …………………ultimate shrinkage strain, in .

sh …………………ultimate shrinkage strain, in . Used in creep equation of B3

model.

ε2……………………deformation at test load divided by the gauge length, in .

Used in AS1012-17.

ε2……………………strain.

ε(t,s)…………………total strain at time t of a reinforced concrete member loaded

first time at time s. Used in AEMM method.

εs(fcm)……………….function accounts for the effects of concrete compressive

strength and cement type on shrinkage. Used in shrinkage

xiv


εsh(t,t0)………………shrinkage strain ( ) at time t (days) of a specimen that

commences drying at time t0 (days). Used in shrinkage equation

of B3 model.

creep(t,t0)……………creep strain.

sh(t,tsh)……………..shrinkage at age t for drying occurred at age tsh.

εsh(t,tsh,0)……………shrinkage strain (mm/mm) at time t (days) of a specimen that

commences drying at time tsh,0 (days);

sh (tsh)………………ultimate shrinkage strain. Used in shrinkage equation of Lyse

(1960).

(t,t0)…………………creep coefficient at age t for concrete loaded at age t0.

(t,t0)…………………creep coefficient at time t of a concrete member loaded at time

t0. Used in EM method.

(t,s)………………….creep coefficient at time t of a concrete loaded first time at

time s. Used in AEMM method.

0……………………..ultimate creep coefficient, in . Used in creep equation of

CEB-FIP.

28……………………creep coefficient. Used in creep equation of GL2000 model.

RH…………………...coefficient for relative humidity. Used in creep equation of

CEB-FIP.

c……………………..product of the applicable correction factors. Used in creep

equation of ACI209R-92.

2……………………adjustment coefficient accounts for the differences of relative

humidity. Used in shrinkage equation of ACI209R-92.

3……………………adjustment coefficient accounts for the differences of average

thickness and volume/surface ratio. Used in shrinkage equation

of ACI209R-92.

4……………………adjustment coefficient accounts for the differences of slump.

Used in shrinkage equation of ACI209R-92.

5……………………adjustment coefficient accounts for the differences of fine

aggregate/total aggregate ratio. Used in shrinkage equation of

ACI209R-92.

xv

6……………………adjustment coefficient accounts for the differences of air

content. Used in shrinkage equation of ACI209R-92.

7……………………adjustment coefficient accounts for the differences of cement

content. Used in shrinkage equation of ACI209R-92.

8……………………adjustment coefficient accounts for the differences of duration

of initial curing. Used in shrinkage equation of ACI209R-92.

……………………..modification factor. Used in elastic modulus equation

suggested by Mendis et al (1997).

φcc……………………creep coefficient. Used in creep equation of AS3600-2001.

cc……………………creep coefficient. Used in creep equation of AS3600-2009.

φcc.b…………………..basic creep factor. Used in creep equation of AS3600-2001.

cc.b…………………..basic creep factor. Used in creep equation of AS3600-2009.

t……………………..creep coefficient. Used in creep equation of ACI209R-92.

u……………………..ultimate creep coefficient. Used in creep equation of

ACI209R-92.

σ0……………………..initial stress. Used in AEMM method.

σ0……………………..initial stress applied to concrete member. Used in AEMM

method.

σ(t)……………………concrete stress at time t. Used in AEMM method.

σ(t)…………………....applied stress at time t.

σip……………………..initial concrete stress due to applied load (component 1).

AEMM

∆σp(t)…………………changes in concrete stress due to applied load (component 2).

AEMM

σiε……………………..initial concrete stress due to shrinkage strain (component 3).

AEMM

∆σε(t)…………………changes in concrete stress due to shrinkage strain

(component 4). AEMM

τsh……………………..size dependence. Used in shrinkage equation of B3 model.

………………………correction factor taken equal to 0.5(t-s). Used in AEMM

method.

∆0……………………..correction factor, taken as 0.85. Used in AEMM method.

xvi

xvii

ABBREVIATIONS

ACI……………………………………………..American Concrete Institute

AEMM………………………………………….Age-Adjusted Effective Modulus

Method

AS………………………………………………Australian Standard

ASCE …………………………………………..American Society of Civil Engineers

ASTM…………………………………………..American Society for Testing and

Materials

CBD…………………………………………….Central Business District

CEB…………………………………………….Comite Euro-Intemational Du Beton

CIA …………………………………………….Concrete Institute of Australia

DEMEC………………………………………...Demountable Mechanical Strain

Gauge

EMM……………………………………………Effective Modulus method

FIP………………………………………………Federation International de la

Precontrainte

GP………………………………………………General purposes

GPa……………………………………………..GigaPascal

HES …………………………………………….High Early Strength Concrete

HSC……………………………………………..High Strength Concrete

HPC……………………………………………..High Performance Concrete

IDM …………………………………………….Improved Dischinger Method

MOS…………………………………………….Method of Superposition

MPa……………………………………………..Megapascal

NATA…………………………………………..National Association of Testing

Authorities, Australia

xviii

xix

Table of Contents

Abstract .......................................................................................................................... i

Statement of Originality ............................................................................................. iii

Acknowledgements ..................................................................................................... iv

List of Publications ..................................................................................................... vi

Notations…………………………………………………………………………….vii

Abbrevations………………………………………………………………………..xvi

Table of Contents ...................................................................................................... xix

List of Tables ............................................................................................................ xxii

List of Figures ......................................................................................................... xxiv

1 Introduction ........................................................................................................ 1 1.1 Background and Overview .............................................................................. 1

1.2 Current Knowledge and Research Rational ..................................................... 6

1.3 Scope and Objectives ....................................................................................... 9

1.4 Thesis Outline ................................................................................................ 10

1.5 References ...................................................................................................... 12

2 Literature reviews ..................................................................... 16 2.1 Introduction .................................................................................................... 16

2.2 High strength/ high performance concrete ..................................................... 17

2.2.1 Definitions of high strength/high performance concrete ...................... 17

2.2.2 Production of high strength/high performance concrete ...................... 21

2.2.3 Implications and application of High strength/high performance

concrete in tall buildings ..................................................................................... 21

2.2.4 Summary ............................................................................................... 23

2.3 Deformations in concrete structure. ............................................................... 23

2.3.1 Instantaneous shortening ...................................................................... 25

2.3.2 Creep and Shrinkage ............................................................................. 25

2.3.3 Reinforcement restraint ........................................................................ 35

2.3.4 Summary ............................................................................................... 35

2.4 Prediction of axial shortening components. ................................................... 36

2.4.1 Prediction models for Young’s Modulus of Elasticity ......................... 36

2.4.2 Prediction models for creep and shrinkage........................................... 42

2.4.3 Summary ............................................................................................... 66

2.5 Analysis of axial shortening in structural members under constant and

variable stress .......................................................................................................... 67

2.5.1 Effective modulus method (EM) .......................................................... 67

2.5.2 Rate of creep method (RC) ................................................................... 68

2.5.3 Superposition method ........................................................................... 68 2.5.4 Age-adjusted effective modulus method (AEMM) .............................. 69

2.5.5 Summary ............................................................................................... 71

2.6 Experimental and analytical studies on axial shortening ............................... 72

2.6.1 Axial Shortening measurements in Australia ....................................... 73

2.6.2 Axial Shortening measurements in United States of America ............. 75

2.6.3 Axial Shortening measurements in the United Kingdom ..................... 90

2.6.4 Axial Shortening measurements in other countries .............................. 91

2.6.5 Summary ............................................................................................... 94

xx

2.7 Summary to the chapter ................................................................................. 95

2.8 References ...................................................................................................... 96

3 Methodology .......................................................................... 110 3.1 Introduction .................................................................................................. 110

3.2 In situ instrumentation and data collection at the Q1 Tower ....................... 110

3.2.1 The Q1 Tower ..................................................................................... 110

3.2.2 Data collection at the Q1 Tower ......................................................... 119

3.3 Laboratory testing ........................................................................................ 132

3.3.1 Concrete .............................................................................................. 132

3.3.2 Specimens Sampling........................................................................... 134

3.3.3 Specimen preparation ......................................................................... 136

3.3.4 Compressive Strength Test ................................................................. 138

3.3.5 Modulus of Elasticity Testing ............................................................. 139

3.3.6 Shrinkage Testing ............................................................................... 142

3.3.7 Creep Testing ...................................................................................... 143

3.4 Summary to the chapter ............................................................................... 146

3.5 References .................................................................................................... 147

4 Concrete Testing Results .......................................................... 151 4.1 Introduction .................................................................................................. 151

4.2 Compressive strength ................................................................................... 152

4.2.1 Analysis of Results ............................................................................. 153

4.2.2 Test results as compared with predicted values .................................. 156

4.2.3 Discussion ........................................................................................... 158

4.3 Modulus of Elasticity ................................................................................... 164

4.4 Drying shrinkage .......................................................................................... 168

4.4.1 Analysis of Results. ............................................................................ 168

4.4.2 Shrinkage: Grade 65 MPa Concrete. .................................................. 169

4.4.3 Shrinkage: Grade 50 MPa Concrete. .................................................. 171

4.4.4 Test results compared with predicted values ...................................... 173

4.5 Creep ............................................................................................................ 175

4.5.1 Analysis of Results ............................................................................. 175

4.5.2 Creep results: Compressive Strength at 28 days ................................ 176

4.5.3 Creep results: Elastic Modulus ........................................................... 177

4.5.4 Creep Results: Drying Shrinkage of creep specimens ........................ 181

4.5.5 Creep Results: Creep. ......................................................................... 182

4.5.6 Test results as compared with predicted values .................................. 186


4.7 References .................................................................................................... 191

5 Measured Axial Shortening of Q1 Columns and Wall ...................... 193 5.1 Introduction .................................................................................................. 193

5.2 The data processing methodology ............................................................... 193

5.2.1 Field documentation ........................................................................... 194

5.2.2 Database documentation ..................................................................... 194

5.2.3 Computation of Results ...................................................................... 194

5.3 Computation of Stress in Columns and Core Walls .................................... 196

5.3.1 Defining the typical construction sequences ...................................... 196

5.3.2 Actual construction history ................................................................. 199

5.3.3 Stress computation .............................................................................. 201

xxi

5.4 Axial Shortening Results ............................................................................. 202

5.4.1 Results and Discussion ....................................................................... 202

5.4.2 Measured axial shortening of columns and walls at level B2 ............ 205

5.4.3 Measured axial shortening of columns and walls at level B1 ............ 220

5.4.4 Measured axial shortening of columns and walls at level L2, L31, L49

and L63 231

5.4.5 Implication of the long-term axial shortening results ......................... 238

5.5 Comparisons of Measured Axial Shortening and Predictions ..................... 243

5.5.1 Computation of axial shortening ....................................................... 243

5.5.2 Effects of the Ambient Environment .................................................. 246

5.5.3 Prediction of the material properties .................................................. 248

5.5.4 Axial shortening calculation ............................................................... 250

5.5.5 Axial shortening spreadsheet .............................................................. 252

5.5.6 Axial shortening results versus predictions at level B2 ...................... 257


5.7 References .................................................................................................... 269

6 Parametric evaluation of the AEMM adopted model ...................... 272 6.1 Introduction .................................................................................................. 272

6.1.1 Sensitivity analysis – concrete grade .................................................. 274

6.1.2 Sensitivity analysis – concrete density ............................................... 275

6.1.3 Sensitivity analysis – construction cycle ............................................ 275

6.1.4 Sensitivity analysis – environment condition ..................................... 277

6.1.5 Sensitivity analysis – creep coefficient .............................................. 280

6.1.6 Sensitivity analysis – percent of reinforcement .................................. 282


7 Conclusion and recommendations ............................................... 285 7.1 Conclusion ................................................................................................... 285

7.2 Recommendations ........................................................................................ 286

Appendix A ........................................................................... 287 A.1 Details of instrumented Columns and Core Walls ....................................... 287

A.2 Axial shortening results by locations ........................................................... 294

A.2.1 AXS results – Level B2 ...................................................................... 294

A.2.2 AXS results – Level B1 ...................................................................... 305

A.2.3 AXS results - Level 31 ....................................................................... 315

A.2.4 AXS results – Level 49 ....................................................................... 318

A.2.5 AXS results – Level 63 ....................................................................... 323

A.3 Axial shortening results versus predictions ................................................. 329

A.3.1 Level B2 columns and core walls – ACI 209R-1992 model .............. 329

A.3.2 Level B2 columns and core walls – AS3600-2001 model ................. 337

A.3.3 Level B2 columns and core walls – AS3600-2009 model ................. 341

A.3.4 Level B2 columns and core walls – Best fitted model ....................... 349

A.4 Axial shortening field data sheet.................................................................. 357

A.5 Views of Database for storage of data reading ............................................ 358

A.6 Views of spreadsheet for calculating and graph plotting of measured axial

shortening .............................................................................................................. 361

xxii

List of Tables

Table 2.1. Types of HPC according to SHRP (Russell, 1999). .................................. 18

Table 3.1. Material Specifications for the Q1 Tower applied to this research were as

follows........................................................................................................................ 113

Table 3.2. Microstrains detected by one division on the dial gauge (Mayes, 2001) 121

Table 3.3. Specifications of the Kestral 3000 (Nielsen-Kellerman, 2004) ............... 123

Table 3.4. Identification of instrumented elements at each selected level and their

corresponding duration of strain measurement .......................................................... 125

Table 3.5. Number of DEMEC stations using 200mm DEMEC gauge ................... 130

Table 3.6. Number of DEMEC stations using 600mm DEMEC gauge. .................. 130

Table 3.7. Concrete mix compositions (Courtesy of Readymix Pty, 2004). ............ 133

Table 3.8. List of material sources/brand names (Courtesy of Readymix Pty, 2004).

.................................................................................................................................... 133

Table 3.9. Concrete grades of vertical members at various levels and corresponding

number of sampling ................................................................................................... 134

Table 3.10. Details of specimens collected at each sampling ................................... 135

Table 3.11. Number and test ages of compressive strength specimens .................... 138

Table 3.12. Number of elastic modulus specimens for each sampling. .................... 142

Table 3.13. Shrinkage measurement reading schedule. ............................................ 143

Table 3.14. Creep measurement schedule................................................................. 145

Table 4.1. Number of Tested Compressive Strength Samples at Various Ages. ...... 154

Table 4.2. Specified and Actual Compressive Strength at 28 days at Q1 (this

research) and World Tower (Bursle et al, 2003). ...................................................... 158

Table 4.3. Deviations in Compressive Strength of Models and Test Results. ......... 163

Table 4.4. Deviations of the Various Elastic Modulus Predictions. ......................... 167

Table 4.5. Number of Shrinkage Samples and Test Schedule. ................................. 169

Table 4.6. Summary of Tests for Creep Specimens.................................................. 176

Table 4.7. Compressive strength at 28 days and corresponding test loads. .............. 177

Table 4.8. Elastic Modulus of Creep Specimens and Standard Specimens. ............. 181

Table 4.9. Pearson correlation of various model predictions as compared to

corresponding measured values. ................................................................................ 190

Table 5.1. Actual construction history of Q1 Tower ................................................ 200

Table 5.2. Summary of Set B2/C1 results................................................................. 209


Table 5.4. Summary of Set B2/W1 and B2/W2 results. ........................................... 215

Table 5.5. Summary of differential shortening between various columns/core walls at

level B2. ..................................................................................................................... 219



Table 5.8. Summary of Set B1/W2 results. .............................................................. 227


level B1. ..................................................................................................................... 230

Table 5.10. Summary of Short-term results at Level L2, L31, L49, L63. ................ 237

Table 5.11. Average monthly weather data, Surfers Paradise CBD (Gold Coast

Seaways Station). Source: Australian Bureau of Meteorology. ................................ 247

xxiii

Table 6.1. The variation of investigated parameters. ................................................ 273

Table A.1. Details of Column TC01, TC02. ............................................................. 287



Table A.4. Details of Column TC07, TC08 .............................................................. 288




Table A.8. Details of Core Wall TW01 .................................................................... 290

Table A.9. Details of Core wall TW06 ..................................................................... 291

Table A.10. Details of Core Wall TW07, TW08 ...................................................... 291

Table A.11. Details of Core Wall TW09 .................................................................. 292






xxiv

List of Figures

Figure 1.1. Axial shortening effects on cladding and pipe systems (Neville, 2002) ..... 1

Figure 1.2. Frame distortion due to differential shortening (Rangan and Warner, 1996)

........................................................................................................................................ 2

Figure 1.3. Details of special moveable connections (Rangan and Warner, 1996) ....... 4 Figure 1.4. Additional moment due to slab tilting (Wong, 2003).................................. 5 Figure 1.5. Computational model for determining slab camber (Beasley, 1997) .......... 6

Figure 2.1. Strain development in plain concrete members (ACI 209.1R-05) ............ 25

Figure 2.2. Creep concrete with different cements, loaded at 28 days (Neville, 1970)

...................................................................................................................................... 28

Figure 2.3 Creep and shrinkage strain variables (Pauw and Chai, 1967). ................... 30

Figure 2.4. Measured column strains, level 23, World Tower Building (Bursle et al,

2003). ........................................................................................................................... 74

Figure 2.5. Lake Point Tower

(http://www.aviewoncities.com/chicago/lakepointtower.htm) .................................... 77

Figure 2.6. Lake Point Tower- levels of instrumentation and measurement locations.

(Pfeifer et al, 1971) ...................................................................................................... 78

Figure 2.7. Lake Point Tower- measured axial shortening of columns. (Pfeifer et al,

1971) ............................................................................................................................ 79

Figure 2.8. Lake Point Tower- measured axial shortening of core walls in millionths

unit (aka microstrain), (Pfeifer et al, 1971). ................................................................ 80

Figure 2.9. Member size correction charts for creep and shrinkage of concrete.

(Pfeifer et al, 1971) ...................................................................................................... 82

Figure 2.10. Lake Point Tower- measured and predicted axial shortening of column at

1st and 30

th storeys. (Pfeifer et al, 1971) ...................................................................... 82

Figure 2.11. Lake Point Tower- measured and predicted axial shortening of core wall

at 10th

and 30th

storeys. (Pfeifer et al, 1971) ................................................................ 83

Figure 2.12. Lake Point Tower- differential shoretning between instrumented columns

and core walls. (Pfeifer et al, 1971) ............................................................................. 84

Figure 2.13. Water Tower Place (http://www.aviewoncities.com) ............................. 85

Figure 2.14. Water Tower Place- measurement locations. (Russell and Larson, 1989)

...................................................................................................................................... 86

Figure 2.15. Water Tower Place- measured axial shortening of columns and core wall

during the first 5 years. (Russell and Larson, 1989) .................................................... 87

Figure 2.16. Water Tower Place- caisson shortening and measured axial shortening at

the fourth basement. (Russell and Larson, 1989) ........................................................ 88

Figure 2.17. Water Tower Place- measured and predicted axial shortening of column

D2. (Russell and Larson, 1989) ................................................................................... 89

Figure 2.18 Measured column strains of the 19-storeyed building by SOFO system. 93

Figure 3.1. The Architect’s perspective of the Q1 Tower (courtesy of Sunland Group,

2002) .......................................................................................................................... 111

Figure 3.2. Illustration of facilities at podium floor (courtesy of Sunland Group, 2002)

.................................................................................................................................... 114

Figure 3.3. Plan of podium floor (Sunland Group, 2003) .......................................... 114

Figure 3.4. Bored piles and pile caps layout of the Q1 Tower (Whaley, 2002) ........ 115

Figure 3.5. An isometric view of the Q1 Tower’s structural elements (Whaley, 2002)

.................................................................................................................................... 117

xxv

Figure 3.6. Basement floor plan of the Q1 Tower (based on Whaley, 2001) ............ 117 Figure 3.7. A typical floor plan of the Q1 Tower (based on Whaley, 2001). ............ 118

Figure 3.8. DEMEC gauges (Mayes, 2001) ............................................................... 120

Figure 3.9. The Kestrel 3000 Pocket Weather Meter

(http://www.nkhome.com/ww/3000/3000.html) ....................................................... 122

Figure 3.10. Instrumented levels and concrete grades (Whaley, 2002) ..................... 124

Figure 3.11. Typical positioning of DEMEC points on columns/walls. .................... 128

Figure 3.12. Instrumented column and DEMEC gauge in use .................................. 129

Figure 3.13. Compressometer setup for Elastic Modulus tests. ................................. 140

Figure 3.14. Creep test ............................................................................................... 145

Figure 4.1. Histogram of 28 days Compressive Strength Results, Grade 65 MPa. ... 155

Figure 4.2. Compressive Strength Results of Grade 65 MPa concrete. ..................... 156

Figure 4.3. 28 days’ Compressive Strength Results of Grade 50 MPa concrete. ...... 156

Figure 4.4. Compressive Strength Results, Grade 50 MPa concrete. ........................ 156

Figure 4.5. Measured Compressive Strength against Model Results from Codes,

65MPa. ....................................................................................................................... 157


50MPa. ....................................................................................................................... 158

Figure 4.7. Regression of the Actual/Specified f’c values of Concrete Grade 50-80

MPa ............................................................................................................................ 160

Figure 4.8. Measured compressive strength over times and predictions by codes in

conjuction with η, series 1. ........................................................................................ 161

Figure 4.9. Measured compressive strength over times and predictions by codes in

conjuction with η, series 2. ........................................................................................ 161

Figure 4.10. Strength development of actual concrete and those due to codes

recommendations. ...................................................................................................... 162

Figure 4.11. Strength development of actual concrete data against those generated by

Equation 4.7. .............................................................................................................. 163

Figure 4.12. Elastic Modulus Results of Q1 Tower Concrete Samples versus AS

3600-2001 predictions. .............................................................................................. 165

Figure 4.13. Elastic Modulus Results of Q1 Tower Concrete Samples compared with

AS 3600-2001, CEB-FIP and AS3600-2009 Models. ............................................... 166

Figure 4.14. Shrinkage Results for Concrete Grade 65 MPa. .................................... 170

Figure 4.15. Shrinkage Results for Concrete Grade 50 MPa. .................................... 171

Figure 4.16. Combined Shrinkage Results of Concrete Grade 65 MPa and 50MPa. 172

Figure 4.17. Drying Shrinkage Results and AS 3600-2001 predictions. ................... 174

Figure 4.18. Drying Shrinkage Results and Predictions based on Various Models. . 174

Figure 4.19. Average Elastic Strain of Creep Specimens. ......................................... 179

Figure 4.20. Average Elastic Modulus of Creep Specimens and Prediction based on

Equation 4.8 ............................................................................................................... 180

Figure 4.21. Drying Shrinkage of Creep Specimens. ................................................ 182

Figure 4.22. Average Measured Strain of Creep Specimens. .................................... 184

Figure 4.23. Creep Coefficient for Q1 Tower Concrete. ........................................... 185

Figure 4.24. Specific Creep of Q1 Tower Concrete. ................................................. 186

Figure 4.25. Measured and predicted creep coefficient of concrete grade 65 MPa. .. 187



Figure 4.28. Creep coefficient comparison with ACI 209. ........................................ 188

Figure 4.29. Creep coefficient comparisons with AS 3600-2001. ............................. 189

xxvi

Figure 4.30. Creep coefficient comparison with AS3600-2009. ............................... 189 Figure 5.1. Simplified loading sequences for columns and core walls (adopted from

Beasley, 1987)............................................................................................................ 199

Figure 5.2. Stress computation procedures (stage 3-4). ............................................. 201

Figure 5.3. A typical plot of measured strains at a column of Q1 Tower (TC06-B2

data)............................................................................................................................ 202

Figure 5.4. Average AXS of column TC06-B2 and linear fitted line at various

intervals ...................................................................................................................... 204

Figure 5.5. Mean Axial Shortening Results, Level B2 Columns – Set B2/C1, Q1

Tower. ........................................................................................................................ 206

Figure 5.6. Computed Stress, Level B2 Columns – Set B2/C1, Q1 Tower. ............. 207

Figure 5.7. Mean Axial Shortening Results, Additional Level B2 Columns – Set

B2/C2 (TC01, TC03, TC04), Q1 Tower. ................................................................... 210


B2/C2 (TC07, TC08, TC09), Q1 Tower. ................................................................... 211


B2/C2 (TC11, TC13, TC14), Q1 Tower. ................................................................... 212

Figure 5.10. Computed Stress, Additional Level B2 Columns – Set B2/C2, Q1

Tower. ........................................................................................................................ 212

Figure 5.11. Mean Axial Shortening Results, Level B2 Core Walls – Set B2/W1 + Set

B2/W2, Q1 Tower. ..................................................................................................... 214

Figure 5.12. Computed Stress, Level B2 Core Walls – Set B2/W1+ Set B2/W2, Q1

Tower. ........................................................................................................................ 215

Figure 5.13. Axial shortening of TW09 plotted against TC06, TC10 and TC12. ..... 216

Figure 5.14. Axial shortening of set B2/W2 core walls plotted against set B2/C2

columns. ..................................................................................................................... 217


Tower. ........................................................................................................................ 221

Figure 5.16. Computed Stress, Level B1 Columns – Set B1/C1, Q1 Tower. ........... 221


B1/C2 (TC01, TC03, TC04), Q1 Tower. ................................................................... 223


B1/C2 (TC07, TC08, TC09), Q1 Tower. ................................................................... 224


B1/C2 (TC11, TC13, TC14), Q1 Tower. ................................................................... 224


Tower. ........................................................................................................................ 225

Figure 5.21. Mean Axial Shortening Results, Level B1 Core Walls – Set B1/W1

(TW06, TW07, TW08), Q1 Tower. ........................................................................... 226


(TW10, TW16, TW17, TW18), Q1 Tower ................................................................ 227

Figure 5.23. Computed Stress, Level B1 Core Walls – Set B1/W1, Q1 Tower. ...... 227


columns. ..................................................................................................................... 229

Figure 5.25. Mean Axial Shortening Results, Level L2 Columns, Q1 Tower. ........ 232

Figure 5.26. Mean Axial Shortening Results, Level L31 Columns (TC01, TC03,

TC05), Q1 Tower. ...................................................................................................... 232


TC08, TC14), Q1 Tower. ........................................................................................... 233

xxvii

Figure 5.28. Mean Axial Shortening Results, Level L31 Columns and core wall

(TC07, TC13, TW18), Q1 Tower. ............................................................................. 233 Figure 5.29. Mean Axial Shortening Results, Level L49 Columns (TC03, TC05,

TC07, TC08, TC09), Q1 Tower................................................................................. 234


TC13, TC14), Q1 Tower. ........................................................................................... 235

Figure 5.31. Mean Axial Shortening Results, Level L63 Columns, Q1 Tower. ...... 235 Figure 5.32. Mean Axial Shortening Results, Level L63 Core Walls, Q1 Tower. ... 236

Figure 5.33. Plot of Axial Shortening from TC05 and TC06 at level B2 .................. 239



Figure 5.36. Plot of Axial Shortening from TW09, TW16 and TW18 at level B2 ... 241

Figure 5.37. Calculation of Axial Shortening flow chart........................................... 246

Figure 5.38. Graphs of average monthly relative humidity, Surfers Paradise CBD

(Gold Coast Seaways Station). Source: Australian Bureau of Meteorology. ............ 247

Figure 5.39. Compressive strength prediction procedures. ........................................ 248

Figure 5.40. Axial shortening calculation procedures ............................................... 257

Figure 5.41. Measured vs predicted axial shortening (using ACI 209R-1992 material

models), set B2/C1 columns and their opposite columns, and TW09-B2. ................ 259

Figure 5.42. Measured vs predicted axial shortening (using AS3600-2001 material

models), set B2/C1 columns and TW09-B2. ............................................................. 260


models), set B2/C1 columns and TW09-B2. ............................................................. 261

Figure 5.44. Measured vs predicted axial shortening (using best material models), set

B2/C1 columns and TW09-B2................................................................................... 262


models), set B2/C2 columns. ..................................................................................... 264






B2/C2 columns........................................................................................................... 267

Figure 6.1. Sensitivity of AEMM adopted model due to compressive strength. ....... 274

Figure 6.2. Sensitivity of AEMM adopted model due to concrete density................ 275

Figure 6.3. Sensitivity of AEMM adopted model due to construction cycle (concrete

grade = 65 MPa). ....................................................................................................... 276


grade = 45 MPa). ....................................................................................................... 277

Figure 6.5. Sensitivity of AEMM adopted model due to environment condition

(concrete grade = 65 MPa). ........................................................................................ 278



Figure 6.7. Sensitivity of AEMM adopted model due to final shrinkage value




Figure 6.9. Sensitivity of AEMM adopted model due to creep coefficient (concrete

grade = 65 MPa). ....................................................................................................... 281

xxviii


grade = 45 MPa). ....................................................................................................... 282

Figure 6.11. Sensitivity of AEMM adopted model due to percentage of reinforcement



(concrete grade = 45 MPa). ........................................................................................ 283 Figure A.1. AXS results, column TC01, Level B2, Q1 Tower.................................. 294

Figure A.2. AXS results, column TC02, Level B2, Q1 Tower.................................. 295








Figure A.10. AXS results, column TC10, Level B2, Q1 Tower................................ 299





Figure A.15. AXS results, core wall TW06, Level B2, Q1 Tower. ........................... 301

















Figure A.32. AXS results, column TC12 (front face), Level B1, Q1 Tower. ........... 310

Figure A.33. AXS results, column TC12 (side face), Level B1, Q1 Tower. ............. 310









Figure A.42. AXS results, column TC01, Level 31, Q1 Tower. ............................... 315


xxix






Figure A.49. AXS results, core wall TW18, Level 31, Q1 Tower. ........................... 318















Figure A.64. AXS results, column TC10, Level 63, Q1 Tower. ............................... 326 Figure A.65. AXS results, core wall TW01, Level 63, Q1 Tower. ........................... 326




Figure A.69. Measured vs predicted axial shortening (TC01-B2, TC02-B2: ACI

209R-1992 material models) ...................................................................................... 329













Figure A.76. Measured vs predicted axial shortening (TW09-B2: ACI 209R-1992

material models)......................................................................................................... 336

Figure A.77. Measured vs predicted axial shortening (TC05-B2, TC06-B2 : AS3600-

2001 material models)................................................................................................ 337


2001 material models)................................................................................................ 338


2001 material models)................................................................................................ 339

Figure A.80. Measured vs predicted axial shortening (TW09-B2 : AS3600-2001

material models)......................................................................................................... 340

xxx

Figure A.81. Measured vs predicted axial shortening (TC01-B2, TC02-B2: AS3600-

2009 material models)............................................................................................... 341


2009 material models)............................................................................................... 342


2009 material models)................................................................................................ 343


2009 material models)............................................................................................... 344


2009 material models)................................................................................................ 345


2009 material models)................................................................................................ 346


2009 material models)............................................................................................... 347


material models)......................................................................................................... 348

Figure A.89. Measured vs predicted axial shortening (TC01-B2, TC02-B2: best

selected material models) ........................................................................................... 350



Figure A.91. Measured vs predicted axial shortening (TC05-B2, TC06-B2 : Best










Figure A.96. Measured vs predicted axial shortening (TW09-B2: Best selected

material models)......................................................................................................... 356

Figure A.97. A field worksheet used for recording of in-situ data reading. .............. 357

Figure A.98. A typical format of data reading input for the database. ...................... 359

Figure A.99. List of queries for data reading at locations ......................................... 360

Figure A.100. A typical view of data sorted by a query. ........................................... 360

Figure A.101. A typical view of column details input in “Cover” worksheet. .......... 361 Figure A.102. A typical view of “Data” worksheet. .................................................. 362

Figure A.103. A typical view of “Cumulative” worksheet ........................................ 363

Figure A.104. A typical view of “Cumulative2” worksheet. ..................................... 363

Figure A.105. A typical view of “Graph A” worksheet. ........................................... 364

Figure A.106. A typical view of “Graph B” worksheet. ............................................ 365

xxxi

1

1 Introduction

1.1 Background and Overview

In tall buildings, vertical structural members such as columns and core walls are

continuously subjected to a certain amount of vertical movement or axial shortening.

For steel structures, axial shortening is rather fundamental as steel columns only

undergo elastic shortening at the instance of load application. However, in the case of

reinforced concrete columns and core walls, the focus of this thesis, the phenomenon

becomes more complicated as the columns also undergo long-term shortening due to

creep and shrinkage (Elnimeiri and Joglekar, 1989).

According to Fintel and Ghosh (1984), axial shortening in concrete buildings is

defined as the combined shortening due to elastic, shrinkage and creep effects of a

column and can be as high as 1 inch for every 80 feet of building height

(approximately 1040x10-6

mm/mm or 1040 microstrains). This axial shortening of

any single column can be problematic to non-deformable (non-structural) elements

that are attached to it as seen in Figure 1.1. In Figure 1.1 buckling of cladding and

pipes that resulted from axial shortening of a column is presented.

Pipes buckling

Claddings buckling

Figure 1.1. Axial shortening effects on cladding and pipe systems (Neville, 2002)

2

Although axial shortening occurs on individual columns or core walls, for design

engineers, it is the differential shortening of two adjacent columns influenced by the

differences in stress levels, loading histories, reinforcement ratios, volume-to-surface

ratios and exposed environment conditions, which becomes a crucial matter of

concern (Neville et al, 1983). It is acknowledged that differential shortening can

theoretically create numerous problems in any given structure. This could include

excessive unexpected movements in the slab-column frame system (as shown in

Figure 1.2) where excessive movement in the exterior column results in frame

distortion and failure of non-structural elements such as cladding, partitions, finishes,

piping, etc. Also excess forces and moments can result in surrounding concrete

elements which, if unaccounted for, can cause overstressing.

Figure 1.2. Frame distortion due to differential shortening (Rangan and Warner, 1996)

Studies on axial shortening or the effects of creep and shrinkage on concrete had

commenced long before the era of high-rise buildings. Back then, the height of

concrete buildings was limited and only a few buildings surpassed a height of 20

storeys. However, such buildings were sturdily built with heavy claddings, partitions

and low-stressed structural members of substantial sizes (oversized columns) giving

the overall system a large capacity in case of overstressing. Conclusively, concerns of

differential shortening effects in early concrete buildings were therefore considered

3

secondary and negligible due to the fact that their effects in those structures were

minimal (Fintel and Ghosh, 1986).

In the early 1960s, a new trend in building construction led to greater awareness on

axial shortening induced problems. These trends included the height of concrete

buildings rising from 20 to 60 storeys which consequently led to a step up from

working stress design to ultimate stress design and a sharp increase in the strength of

concrete (Fintel and Ghosh, 1984).

The increased building height magnified the differential shortening issue as the axial

shortening of a few millimeters per floor which, when multiplied by a large number of

floors, would be accumulated up to a few inches (50 mm) over the total height. For

example, an average differential shortening of 200 microstrains per floor can lead to a

cumulative differential of 48 mm at the top floor of an 80-storey building with floor

height of 3.00 m (Rangan and Warner, 1996).

The other trends i.e. change in design method and increased concrete strengths

resulted in smaller sections of vertical structures where stiffness and rigidity are

reduced due to relatively high axial stresses. These factors altogether therefore made

modern high-rise buildings more sensitive to differential shortening effects and thus

this problem was no longer a secondary issue but at the forefront of structural designs.

It is therefore considered a necessity to accommodate the effects of differential

shortening appropriately as negligence would lead to serviceability deficiencies or, in

extreme cases, structural distress.

The implications of differential shortening in a high-rise building can be structural as

well as non structural depending on its magnitude. Where minor differential

shortening occurs, its effects are non structural and generally we are only concerned

with the impaired serviceability of the building that would result from the tilted slabs

(Russell, 1985). These non structural adversities can be eradicated effectively by

using specially designed connections which provide a certain degree of tolerance

between joist connections such as the ones shown in Figure 1.3.

4

Figure 1.3. Details of special moveable connections (Rangan and Warner, 1996)

5

Figure 1.4. Additional moment due to slab tilting (Wong, 2003) However, when major differential shortening occurs, structural effects are induced by

large extra moments, ±M, that are generated from the distorted slabs or beams as

shown in Figure 1.4. Additional moment due to slab tilting (Wong, 2003), this

induced moment together with the load redistribution due to the differential

shortening could overstress the more rigid columns as load will transfer to the

columns that shorten less and away from the column that shorten more, particularly

where columns are located close to each other (Fintel and Ghosh, 1984; Smith and

Coull, 1991).

Reports of damages due to differential shortening in buildings constructed after the

1960s can be found in the literature (Plewes, 1976 and 1977; Neville, 1998 and 2002).

It is therefore essential that during the course of designing high-rise buildings, in

order to prevent such distortions, the designer has to determine the magnitude of

differential shortening beforehand. By this means, the anticipated differential length

can be counteracted by providing adequate formwork cambering during construction.

An example of a computational model for determining slab camber is shown in

Figure 1.5.

Traditionally, differential shortening predictions are made through mathematical

expressions derived from short-term test data, recommended formulae of material

models and analytical procedures of concrete members. These methods offer a

practical level of accuracy and have been widely used to account for differential

shortening. However, the majority of these methods are considered to be outdated as

there is a deficiency of verification data for its application in concrete applications

today, which tend to be more sophisticated in terms of their composition and

performance. Therefore, these methods and verifications need to be addressed

accordingly.

6

Figure 1.5. Computational model for determining slab camber (Beasley, 1997)

1.2 Current Knowledge and Research Rational

At present, the demand for residential units and office suites is growing rapidly,

resulting in the construction of numerous high-rise buildings throughout Australia and

around the world. In the quest to have their buildings recognized as major city

landmarks, investors and owners are in favor of superior and innovative structures. A

majority of these structures are of reinforced concrete which, by far is the most widely

accepted construction material.

7

With the new age concrete structures, especially those exceeding 60 storeys, there has

been a trend to specify concrete of very high strengths (28 days strength greater than

65MPa) in order to minimize the sizes of vertical members (columns and core walls),

and improve constructability due to the consequent minimal steel content when High

Strength Concrete (HSC) is employed (Beasley, 1997). As a result, these members

are subject to large imposed loadings leading to high-sustained stress and causing

time dependent shortening.

It is desirable to make accurate axial shortening predictions in order to prevent

excessive deformation in members and to avoid subsequent serviceability problems.

In addition to this, with the increasing height of buildings, data on differential

shortening is limited.

In order to reasonably predict axial shortening, in the absence of test data, predictions

of creep, shrinkage and elastic modulus of plain concrete can be used to arrive at a

rough estimation of long-term deformation of concrete. Within the past 30 years,

various models replicating concrete properties have been widely used and studied.

They include models recommended by ACI 209 (American Concrete Institute), CEB-

FIP (Comite Euro-International Du Beton), AS3600 (Australian Standard 2001 and

2009) and those developed by independent researchers such as the BP model (Bazant,

1972; Bazant and Panula, 1978; Bazant et al, 1991, 1992; Bazant and Kim, 1991,

1992a, 1992b, 1992c), B3 model (Bazant and Baweja, 1995a, 1995b, 1995c), and the

B3 model-short form (Bazant and Baweja, 2000)

More precise predictions can be made through reinforced concrete column models

and methods which generally involve complex analytical procedures and the effects

of concrete composition, stress history, age at loading, intensity of stress, shape and

size of members, longitudinal reinforcement, etc into the analysis. According to

Neville et al, 1983, some of the best known methods are: the Effective Modulus

method (EMM), Rate of Flow method (RFM), Rate of Creep method (RCM), the

Improved Dischinger Method (IDM), the Method of Superposition (MOS) and the

Age-Adjusted Effective Modulus Method (AEMM).

8

For specific applications of high-rise concrete buildings, where high strength concrete

(HSC) is a requirement, the use of such analytical models may lead to inaccurate

shortening predictions due to the perceived errors in creep and shrinkage models for

higher strength concretes used in such analysis. In the recent years, researchers have

attempted to verify these models and consequently traced prediction errors. It was

acknowledged that these material models need to be modified and verified for use in

predicting shortening in tall concrete buildings which use HSC. However, models

recently derived from high strength concrete data such as Gilbert (2002, 2004, and

2005), Huo et al (2001) and updated code such as AS3600-2009 may help provide

better material predictions.

Although, a rather reasonable prediction of shortening in tall buildings could be

achieved by any of the above-mentioned methods together with updated material

models for HSC, the accuracy is still in doubt and needs to be justified since all data

available is based on investigations of small specimens which, are not the true

representative of concrete used in the proposed structure.

Given the limitations of existing knowledge and in order to provide certain and more

accurate axial shortening predictions, an experimental study on concrete axial

shortening in tall buildings, by recording of strain data experienced by concrete in real

structures, is therefore necessary.

Research on the recently constructed 80-storeys Q1 Tower (located in the Gold Coast

region, Australia) was conducted to investigate and obtain the required data. By

monitoring strains on selected members, a vast amount of data representing axial

shortening in the Q1 Tower was obtained during the three (3) years of its

construction. The project is considered unique as only four buildings were fully

instrumented: Water Tower Place (Russel and Corley, 1978), Lake Point Tower

(Pfeifer et al, 1971), UTS Building, (Bakoss et al, 1984) and the University of

Technology Sydney Building (Brady E.A, 1980) have been monitored for axial

shortening during the last 30 years. The Q1 Tower measurements and data collection

are unique as it is by far the tallest residential building in the world constructed using

9

the commercially available high strength concrete mixes. Although the contributions

of previous works are of great significance, they were based on concrete strengths and

heights much lower than those of today’s taller structures. Therefore, the Q1 Tower

measurements encompass the prevailing conditions in terms of material properties and

loadings.

The validation and consequent modification of material models and the development

of an analytical shortening procedure using results from the Q1 Tower will be

significant for future implementations of reliable shortening methods. Furthermore,

instrumentation at Q1 Tower is thus far, the most extensive and comprehensive as it

comprises of a large number of monitored elements, along with a vast number of

smaller specimens that were lab-tested for concrete properties.

1.3 Scope and Objectives

In order to make a substantial contribution in alleviating the above mentioned

predicaments, the ultimate aim of this research is to contribute significantly to the

experimental and analytical studies of axial and differential shortening. Reliable

creep, shrinkage and elastic modulus models would be verified and/or modified and

utilized to calculate the axial shortening of columns and core walls. This will be

compared to the actual field measurements from the Q1 Tower.

In order to achieve this main aim, the following sub-objectives are herewith outlined:

To obtain data on time dependent axial shortening of reinforced concrete

columns and core walls of a high-rise structure (i.e. using the 80 storey Q1

Tower, in Surfers Paradise, Gold Coast, Australia). Additional standard

concrete tests such as compressive strength, elastic modulus, shrinkage and

creep would also be simultaneously conducted in the lab to determine actual

shortening-related mechanical properties of concrete used in the monitored

members (concrete specimens were cast from the site delivered concrete).

10

To compare the experimental creep and shrinkage data obtained from lab

specimens with the prediction results obtained from current material models in

order to justify the most suitable expression to be used in enhancing the

development of a shortening analytical model.

To develop a reliable axial shortening prediction method by incorporating the

material models and idealized construction histories into an existing analytical

method for the prediction of differential shortening of columns and core walls

in tall concrete buildings and subsequently verify the developed model with

data obtained from the Q1 Tower.

To assess the limits and capability of the developed axial shortening analytical

models as well as its sensitivity to changes of various factors.

1.4 Thesis Outline

This research study is presented in seven (7) chapters wherein this chapter has

outlined the background of differential shortening problems through statements of

current knowledge, research significance, scope and aims of the study.

Chapter Two presents a comprehensive review of literature related to the research.

The reviewed topics include high strength/high performance concrete,

deformation/axial shortening in concrete structures and its components i.e. elastic,

creep, shrinkage strain as well as effects of reinforcement and climate changes,

prediction models/methods for axial shortening components (material models) from

mathematical expressions derived from a range of test data to widely adopted generic

models recommended by international design codes and alternative models proposed

by researchers.

The Chapter Two also gives a review on analytical methods commonly used in

computing axial shortening in structural members as well as the principle of

superposition and its relevance to the calculation of axial shortening. A review of

11

published axial shortening data obtained from existing structures and those findings is

also presented.

Chapter Three gives a detailed description of the experiments, including illustrations

and detailed information on Q1 Tower, details of the implemented methodologies for

the DEMEC gauge instrumentation and data collection used, as well as details of lab

experiments undertaken to determining concrete properties.

Chapter Four deals with standard test results obtained from the laboratory. The four

tested properties of concrete samples i.e. compressive strength, elastic modulus,

shrinkage and creep are reported therein. Comparisons between the test results and

prediction by widely accepted international standards are made. For each studied

property, a model which gives the best prediction is improved by using modified

parameters and justified using the test data.

Chapter Five presents axial shortening results obtained from the Q1 Tower. An axial

shortening analytical method incorporating modified material models (international

standards/well established models verified/modified to fit the Q1 Tower lab results)

and a typical loading sequence (a modified loading sequence based on work presented

in Beasley (1987) incorporating actual construction history of the Q1 Tower) is

adopted. In addition, comparisons between the measured axial shortening of the Q1

Tower and the predicted values are also made.

Chapter Six emphasizes the effects of various factors on the axial shortening

prediction of the method previously outlined in Chapter Five. The impact of changing

each factor to the axial shortening results is carefully examined by means of

sensitivity analysis.

Chapter Seven concludes the thesis by providing a summary of the findings and gives

recommendations for further studies.

12

1.5 References

ACI Committee 209, “Prediction of Creep, Shrinkage, and Temperature Effects in

Concrete Structures (ACI 209R-92),” American Concrete Institute, Farmington Hills,

Mich., 1992, 47 pp.

AUSTRALIAN STANDARD AS 3600-2001, “Concrete structures,” Standards

Association of Australia, Sydney, Australia, 2001, 165 pp.



Bakoss, S.L., Burfitt, A., Cridland, L., Heiman, J.L., “Measured and Predicted Long-

Term Deformations in a Tall Concrete Building,” Deflection of Concrete Structure,

SP-86, American Concrete Institute, Detroit, 1984, pp. 63-94.

Bazant, Z.P., “Prediction of Concrete Creep Effects Using Age-Adjusted Effective

Modulus Method,” ACI Journal, V.69, April. 1972, pp. 212-217.

Bazant, Z.P., and Beweja, S., “Creep and Shrinkage Prediction Model for Analysis

and Design of Concrete Structures-Model B3,” Materials and Structures, V.28,

No.180, 1995a, pp. 357-365, 415-430, and 488-495.

Bazant, Z.P., and Beweja, S., “Justification and Refinements of Model B3 for

Concrete Creep and Shrinkage – 1 Statistics and Sensitivity,” Materials and

Structures, V.28, 1995b, pp. 415-430.

Bazant, Z.P., and Beweja, S., “Justification and Refinements of Model B3 for

Concrete Creep and Shrinkage – 1 Updating and Theoretical Basis,” Materials and

Structures, V.28, 1995c, pp. 488-495.

Bazant, Z.P., and Beweja, S., “Creep and Shrinkage Prediction Model for Analysis

and Design of Concrete Structures (Model B3),” In: AlManesseer, A., (Ed.), Creep

and Shrinkage of Concrete, ACI Special Publication, 2000, in press.

13

Bazant, Z.P., Kim, J.K., and Panula, L., “Improved Prediction Model for Time-

Dependent Deformations of Concrete: Part 1: Shrinkage,” Materials and Structures,

V.24, 1991, pp. 327-345.

Bazant, Z.P., and Kim, J.K., “Improved Prediction Model for Time-Dependent

Deformations of Concrete: Part 2: Basic Creep: Part 3: Creep at drying: Part 4:

Temperature Effects: Part 5: Cyclic Load and Cyclic Humidity,” Materials and

Structures, Part 2, V.24, 1991, pp. 409-421; Part 3, V.25, 1992a; Part 4, V.25, 1992b,

pp. 84-94; Part 5, V.25, 1992c, pp. 163-169.

Bazant, Z.P., Panula, L., and Kim, J.K., “Improved Prediction Model for Time-

Dependent Deformations of Concrete : Part 6: Simplified Code-Type Formulation,”

Materials and Structures, V.25, 1992, pp. 219-223.

Bazant, Z.P., and Panula, L., “Practical Prediction of Time-Dependent Deformations

of Concrete”, Materials and Structures, Parts I and II, Vol. 11, No. 65, pp. 307-328;

Parts III and IV, Vol. 11, No. 66, pp. 415-434; Parts V and VI, Vol. 11, No. 67, pp.

169-183, 1978.

Beasley, A.J., “A Numerical Solution for the Prediction of Elastic, Creep, Shrinkage

and Thermal Deformations in the Columns and Cores of Tall Concrete Buildings,”

Research Report CM – 87/2, Civil and Mechanical Engineering Department, Faculty

of Engineering, University of Tasmania, 1987.

Beasley, A.J., “High Performance Concrete-Implications for Axial Shortening in Tall

Buildings,” Department of Civil and Mechanical Engineering, University of

Tasmania-Hobart, Tasmania, 1997.

CEB-FIP 1993., “CEB-FIP Model Code 1990,” Bulletin d’Information No 213/214,

Comite Euro-International Du Beton, 1993, 436 pp.

Elnimeiri, M.M., and Joglekar, M.R., “Influence of Column Shortening in Reinforced

Concrete and Composite High-Rise Structures,” Long-Term Serviceability of

Concrete Structure, SP-117, American Concrete Institute, Detroit, 1989, pp.55-86.

14

Fintel, M., and Ghosh, S.K., "High Rise Design: Accounting for Column length

Changes,” The journal of the American Society of Civil Engineering ASCE, Vol 54,

No.4, April 1984, pp. 55-59.

Fintel, M., and Ghosh, S.K., "Column Length Changes in Ultra-High-Rise Buildings,”

Advances in Tall Buildings, Council on Tall Buildings and Urban Habitat, Van

Nostrand Reinhold, New York, 1986.

Gilbert, R.I., “Creep and Shrinkage Models for High Strength Concrete – Proposal for

inclusion in AS3600,” Australian Journal of Structural Engineering, IE.Aust, Vol. 4,

No. 2, 2002, pp. 95-106.

Gilbert, R.I., “Design properties of materials – changes proposed for AS3600 – creep

and shrinkage clauses,” Concrete in Australia, Concrete Inistitute of Australia, Vol.

30, No. 3, September-November 2004, pp. 24-28.

Gilbert, R.I., “AS3600 Creep and Shrinkage Models for Normal and High Strength

Concrete”, Shrinkage and Creep of Concrete, SP-227, American Concrete Institute,

Farmington Hills, Mich, 2005, pp. 21-40.

Huo, X.S., Al-Omaishi, N., and Tadros, M.K., “Creep, Shrinkage, and Modulus of

Elasticity of High Performance Concrete,” ACI Materials Journal, Volume 98, No.6,

November-December. 2001, pp.440-448.

Neville, A.M.; Dilger, W.H,; and Brooks, J.J., “Creep of Plain and Structural

Concrete,” Construction Press, London, 1983.

Neville. A.M., “Concrete Technology – An Essential Element of Structural Design,”

Concrete International, Vol 20, No.7, July 1998, pp. 39-41.

Neville. A.M., “Creep of Concrete and Behavior of Structures, Part I: Problems,”

Concrete International, Vol 24, No.5, May 2002, pp. 59-66.

Pfeifer, D.W., Magura, D.D., Russell, and H.G., Corley. W.G., “Time Dependent

Deformations in 70 Storey Structure,” Designing for effects of Creep and Shrinkage

15

Temperature in Concrete Structure, SP-27, American Concrete Institute, Detroit,

1971, pp.159-185.

Plewes, W.G., “Vertical Movement of Building Frames and Cladding, in Cracks,

Movements and Joints in Buildings,” Nat. Ras. Council of Canada, Div. Bldg. Res.,

NRCC 15477, September, 1976.

Plewes, W.G., “CBD-185: Failure of Brick Facing on High-Rise Buildings,”

Canadian Building Digest, 1977, [on-line] Available from, http://irc.nrc-

cnrc.gc.ca/cbd/cbd185e.html; Accessed 20/10/2005.

Rangan, B.V., Warner, R.F., “Large Concrete Building.” Longman, Essex, 1996,

269pp.

Russell, H.G., “High-Rise Concrete Buildings: Shrinkage, Creep, and Temperature

Effects.” Analysis and Design of High-Rise Concrete Buildings, SP-97, American

Concrete Institute, Detroit, 1985, pp.125-137.

Russel, H., and Corley, W., “Time-Dependent Behavior of Columns in Water Tower

Place,” Douglas Mchenry International Symposium on Concrete and Concrete

Structures, SP-55, American Concrete Institute, Detroit, 1978, pp.347-373.

Smith, B.S., and Coull, A., “Tall Building Structures: Analysis and Design.” John

Wiley and Sons Inc, 1991, Canada.

Wong, C.Y., “Time Series Modelling of Column Shortening in Tall Buildings due to

Creep and Shrinkage of High Performance Concrete.” Master of Engineering Thesis,

RMIT University, Melbourne, 2003.

http://irc.nrc-cnrc.gc.ca/cbd/cbd185e.html

http://irc.nrc-cnrc.gc.ca/cbd/cbd185e.html

16

2 Literature Reviews

2.1 Introduction

Axial shortening of concrete is one topic of interest that covers a large scope of

studies including mechanisms of creep and shrinkage.

The scope of literature reviews covered the following areas only:

- The identification and the implication of high strength/ high performance

concrete

- Theories and mechanisms of axial shortening

- Prediction models available for estimation of related concrete properties

- Development of axial shortening prediction models

- Field experiments to collect axial shortening data from buildings

An anthology of definitions and required performance critiria of high strength/ high

performance concrete as set by researchers are reviewed in Section 2.2. This is

followed by Section 2.3 that contains a concise summary on axial shortening

machanisms in buildings and the knowledge on each of its components. The

prediction methods/code models available for estimation of time dependent properties

of concrete that are required for further axial shortening analysis, i.e. compressive

strength, elastic modulus, shrinkage and creep when verified test data of concrete used

is not available are outlined in Section 2.4. Creep and shrinkage models taking

various mathematical forms are also examined. In Section 2.5, well-accepted generic

models for axial shortening predictions are presented. The pros and cons of each

models are also discussed. Finally, a survey on experimental studies of axial

shortening in actual buildings is presented in Section 2.6. Data taken from various

buildings with different types of instrumentation are investigated with a most suitable

instrumentation technique justified to be used in this study.

17

2.2 High strength/ high performance concrete

The popularity of high-rise buildings made from concrete gives rise to the demand for

concrete with greater compressive strength and workability in order to maintain

sufficient stiffness of the vertical members while maintaining their suitable size

according to the Architectural requirements. These concrete are usually referred to as

high strength/high performance concrete. However, as the height of the new towers

increased, the definition and requirements of high strength/high performance concrete

have been changed from time to time.

In order to establish a mutual understanding on the definitions, common properties

and the trend and justification of use for high strength/high performance concrete, its

definitions; productions and usage history in tall buildings as reported in the literature

are outlined in the sub-sections to follow.

2.2.1 Definitions of high strength/high performance concrete

In terms of general understanding, High Strength Concrete (HSC) usually refers to

structural concrete produced with relatively higher 28-day strength than the usual

normal strength mix in order to meet requirements due to some special circumstances

(Aitcin and Neville, 1993). This requirement is used to satisfy the need for a taller

structure with maximised rentable space given minimum column size and core wall

thickness (Moreno, 1990). The benchmark 28-day strength of HSC has risen steadily

over the last century alongside the steady increase of 28-day strength of the

commonly used concrete. For example, in the 1960s, according to Aitcin and Neville

(1993), the term HSC was applied to concrete with strengths in excess of 40 MPa

whilst some thirty years later, HSC refers to 65 MPa concrete and above.

The definitions of HSC also vary depending on geographical usage due to the

variances of the local building codes. HSC shall be defined henceforth according to

the definition given in AS3600-2001, which is the current Australian Standard for

18

designing of concrete structures, as the concrete with compressive strength greater

than 65 MPa.

Before defining high performance concrete (HPC), it is important to give explanation

on the apparent difference between the above mentioned HSC and HPC. This is

because many people misunderstand that HPC is the new name of HSC but it is

actually much more (Papworth and Ratcliffe, 1994). The difference between the two

terms being that while the properties of HSC emphasis is only in terms of 28-day

strength, HPC also covers other aspects of a concrete’s physical properties such as

durability (Russell, 1999). HPC also refers to concrete designed to have enhanced

physical properties in order to perform specific functions (Monreno, 1990). According

to the Strategic Highway Research Program or SHRP (Zia et al, 1991 and 1993) HPC

refers to concrete that meets the following requirements:

1. Ratio of water-to-cementitious materials not exceeding 0.35

2. Minimum durability factor, as determined by ASTM C666 Method A, greater

than 80 percent

3. Minimum strength; 21 MPa within 4 hours, 34 MPa within 24 hours, 69 MPa

within 28 days

Based on SHRP requirements, HPC has been categorized into four types as follows;

Very early strength (VES), high early strength (HES), very high strength (VHS) and

Fiber Reinforced. Properties of the four types of HPC are presented in Table 2.1.

Table 2.1. Types of HPC according to SHRP (Russell, 1999).

Type of HPC Strength criteria Water-to-

cementitious ratio

Minimum

durability factor

Very Early Strength (VES) 14 MPa in 6 hours ≤0.4 80%

High Early Strength (HES) 34 MPa in 24 hours ≤0.35 80%

Very High Strength (VHS) 69 MPa in 28 days ≤0.35 80%

There have been many other definitions of HPC available in the literature. These

definitions are proposed individually and differently by researchers/organizations

19

from many countries such as those of, to name but a few; from USA, Forster (1994),

and Mather (1996); from England, Neville (1995), and Swamy (1996); from

Australia, Papworth and Ratcliffe (1994), Ryan and Potter (1996), and Rangan

(1996); from Japan, Tomosawa (1996). Definitions of HPC as given by these

researchers are listed as follows:

“Concrete made with appropriate materials combined according to a selected

mix design and properly mixed, transported, placed, consolidated, and cured

so that the resulting concrete will give excellent performance in the structure

in which it will be placed, in the environment to which it will be exposed, and

with the loads to which it will be subjected for its design life” (Forster, 1994)

– USA.

“Any concrete that has properties needed for some objectives that are not

possessed by ordinary concrete” (Mather, 1996) – USA.

“Concrete in which its ingredients and proportions are specifically chosen so

as to have particular appropriate properties for the expected use of the

structure; these properties are usually a high strength or low permeability”

(Neville, 1995) – England.

“Concrete which is designed to give optimized performance characteristics for

a given set of loads, usage, and exposure conditions consistent with the

requirements of cost, service life, and durability” (Swamy, 1996) – England.

“Concrete that meets the requirements or goes beyond the limits of the normal

performance range” (Papworth and Ratcliffe, 1994) – Australia.

“Concrete that meets multiple performance criteria that are significantly more

stringent than those required for normal structural concrete” (Ryan and Potter,

1996) – Australia.

20

“There is no unique definition of HPC; it can be defined only with reference to

the performance requirements of the intended use of the concrete” (Rangan,

1996) – Australia.

“Concrete with high workability and very high fluidity with a density that can

be controlled as required, to meet the diversified requirements for design and

construction environments of a wide variety of structures” (Tomosawa, 1996)

– Japan.

In recognition of many definitions available throughout the world, Russell (1999)

states that HPC users seem to think that they know what HPC is; however, it is more

difficult for them to put a globalised definition for HPC on paper. However, according

to Russell (1999)’s review, definitions of HPC as proposed by researchers mentioned

above “all have validity but all mean something different”.

To develop a working definition for HPC, a subcommittee on high-performance

concrete (THPC) was established in 1991 by ACI’s Technical Activities Committee

(TAC). The official ACI definition of HPC as represented in Russell (1999) is

reproduced as follow: High-Performance concrete is “concrete meeting special

combinations of performance and uniformity requirements that cannot always be

achieved routinely using conventional constituents and normal mixing, placing, and

curing practice”. To describe HPC in more simple terms, TAC comments that HPC is

“a concrete in which certain characteristics are developed for a particular application

and environment”. Examples of certain characteristics mentioned in the TAC

commentary (Russell, 1999) and other report (Papworth and Ratcliffe, 1994) are

reproduced as follows:

Ease of placement (workability)

Compaction without segregation

Early age strength

Strength

Setting time

Long-term mechanical properties

21

Permeability

Density

Durability

Ductility

Heat of hydration

Toughness

Volume stability

Long life in severe environment

2.2.2 Production of high strength/high performance concrete

In general, the production of HSC/HPC is relatively more stringent as compared to

conventional concrete. To meet special requirements in terms of both strength

development and workability, the raw materials must be of high quality and control of

the manufacturing process of HSC/HPC must be strictly enforced.

Mix design of HSC/HPC must be user friendly, precise, consistent and

economy to reproduce.

Certain admixtures have been added to the HSC/HPC mix to give additional

workability/strength.

Set controllers and superplasticizers have been largely used to give HSC/HPC

good characteristics i.e. high workability, long slump retention and controlled

setting. Polymer based admixtures may be added too.

2.2.3 Implications and application of High strength/high performance concrete in

tall buildings

In general, HSC/HPC has been used globally in constantly increasing volume during

the past twenty years. It has been popular because it is economical by means of cost

per resisting stress. Also because of its greater strength, the size of HSC/HPC

columns and core walls are relatively smaller. The benefit of HSC/HPC is not limited

only in terms of strength as major report (ACI Committee 363R-1992) indicated that

22

long-term serviceability is also enhanced. Also a reason for choosing HSC/HPC

includes specific needs such as early formwork removal, better water retention, and

greater stiffness (Webb, 1993). Although most HSC/HPC applications to date have

been found in building structures, it should be noted that, this state-of –the-art

concrete has also been used sparingly in other types of structures such as bridges,

dams, piles for marine foundations, grandstand roof, modular bank vaults, and

industrial manufacturing panels and floors (ACI Committee 363, 1987).

The economic advantage of HSC/HPC has been reported by many researchers (ACI

Committee 363, 1987; Burnett, 1989; Moreno, 1990; Webb, 1993; Papworth and

Ratcliffe, 1994) particularly when it is used in primary structural members i.e.

columns and core walls.

In Chicago, HSC/HPC has been used since the 1960s, started with concrete strength

of 41 MPa. Since then the strength of HSC/HPC used in Chicago gradually increased

extensively to which 117 MPa concrete has been used recently. The use of HSC/HPC

in some well known Chicago buildings during the 1960s – 1990s has been reported in

literature (Moreno, 1990). These buildings were Outer Drive East (circa 1962, 41

MPa), Lake Point Tower (circa 1965, 52 MPa), Mid-Continental Plaza (circa 1972, 62

MPa), River Plaza (circa 1976, 76 MPa), Chicago Mercantile Exchange (circa 1982,

96 MPa), Construction Technology Laboratories (circa 1988, 117 MPa), and 225 W.

Wacker Drive (circa 1990, 96 MPa).

In Australia, HSC/HPC with compressive strengths around 50 to 65 MPa have been

used in major building projects in major cities since the 1970s. In Melbourne, tall

buildings built with HSC/HPC include some notable structure i.e the Collins Place (55

MPa), the Rialto (60 MPa), Bourke Place (55-level, 60 MPa), and 530 Collins Street

(39-level, 65 MPa) (Burnett, 1989). Concrete with greater strengths, 80 MPa concrete,

was first used on Melbourne Central, a 55-level tower completed in 1991, by the

initiation of Connell Wagner, Lewis Construction and CSR Readymix. The concrete

mix and its special ingredient i.e. densified silica fume were then considered new to

Australia. After the completion of Melbourne Central, the use of HSC/HPC in

23

Australia continues and now 100 MPa concrete can be ordered for delivery in most

capital cities (Papworth and Ratcliffe, 1994).

Examples of buildings located elsewhere that used HSC are Scozia Plaza (circa 1987,

Toronto), La Laurentienne Building (circa 1984, Montreal), KLCC (80 MPa,

Malaysia), and Biyok Tower II (80 MPa, Bangkok).

2.2.4 Summary

Various definitions of HSC and HPC as given in the literature were reviewed. For

HSC, compressive strength is the only defining criterion. But for HPC, the criteria

depend on special characteristics of concrete that are specified for a particular job.

In general, the minimum strength for concrete to be considered as HSC increases over

years as the quality of concrete continues to improve and the use of concrete once

considered as HSC becomes normal practice. For this study, it is justified that

concrete used in the Q1 Tower building having compressive strength greater than 50

MPa be recognised as HSC/HPC. In this research, the term HPC is disregarded to

avoid confusion between the definitions of HSC and HPC since the performance of

HSC usually meets with the requirement of HPC.

Records of HSC used in well-known tall buildings located in various parts of the

world were reported with strengths up to 100 MPa being used. The nominal concrete

strengths used in the building in this research was either 40, 50, and 65 MPa. As

described later greater actual strengths were achived.

2.3 Deformations in concrete structure.

The formation process of axial deformation that occurs in a plain concrete member

under a constant loading condition can be represented diagrammatically as shown in

Figure 2.1. After the casting time, the day zero, concrete is cast into a mould and

remains there until the mould is removed at day t0. During this period the concrete

24

swells due to the absorption of the excess water thus negative strain (expansion) can

be detected if measured.

After time t0, the mould is removed and the loss of water commenced. The shrinkage

strains of both autogenous and drying types therefore start to eradicate the negative

strain caused by the swelling. At some points when shrinkage strain becomes greater

than the swelling, the concrete therefore begins to deform accordingly and the strain

becomes positive (contraction) (ACI 209.1R-05).

At a particular time, the stress is applied to the concrete member causing a sharp

strain increase immediately. This strain occurs due to the elastic properties of the

concrete and is defined as the initial strain, elastic strain or instantaneous strain. After

the emergence of the initial strain, the development of creep strain starts. Creep

consists of two sub-types i.e. basic creep and drying creep which both continue to

increase indefinitely (ACI 209.1R-05).

The total axial shortening at selected time can be determined by adding up the known

values of the components previously mentioned. However, in case of concrete with

reinforcement, the negative strain caused by the reinforcement restraint (not shown in

Figure 2.1) must be included (O’Moore, 1996).

25

Figure 2.1. Strain development in plain concrete members (ACI 209.1R-05)

The various stages of axial deformation in a section are described in detail in the sub-

sections to follow.

2.3.1 Instantaneous shortening

The instantaneous or elastic strain is defined as the strain that occurs immediately

when a stress is applied on a concrete specimen. It consists of an elastic component

which is recoverable and an inelastic component which is irrecoverable. However, the

inelastic component is relatively small if the subjected stress is low. This means that

the instantaneous strain in concrete can be assumed to be purely elastic since applied

stress in most structures rarely exceeds half of the characteristic compressive strength

(0.5f’c). The instantaneous strain is mostly influenced by the magnitude of stress and

elastic modulus of concrete, and the relationship of the three variables is expressed in

the following form:

tE

tt

c

e Equation 2.1

where: εe(t) is the instantaneous strain at time t; σ(t) is the applied stress at time t; and

Ec(t) is the elastic modulus of concrete at time t.

2.3.2 Creep and Shrinkage

2.3.2.1 Shrinkage

Shrinkage is a time-dependent deformation that occurs in concrete whether subjected

to load or not. It is caused by physical-chemical changes that occur without stress and

the loss of liquid in the cement gel during the drying process which can last for a

number of years after the making of that concrete (Seed, 1948).

26

Shrinkage is affected by a large number of variables that affect the drying of concrete

such as environment ambient humidity, age of concrete, size of member, water and

cement content, the mix characteristics etc. (Branson, 1977; Neville, Dilger and

Brook, 1983). In general, the rate of shrinkage is usually high at early ages in the

hours and days after drying commences and gradually decreases when age increases

until reaching a final value after years (see Figure 2.1). However, initial shrinkage

depends on the degree of initial moist curing before the drying commenced and the

age of concrete at its first drying for it is known that concrete with poor curing has

more potential for large shrinkage. Shrinkage is usually expressed as a dimensionless

strain εsh (mm/mm) (Neville, 1995).

The basic shrinkage strain sh refers to the assumed ultimate or final value of

shrinkage. In most practical cases, its value should fall in a range of 500x10-6

to

1000x10-6

mm/mm. The basic shrinkage strain and the rate of shrinkage of a concrete

specimen are dependent mostly upon properties of concrete, size of specimen and the

environment conditions. This value can be obtained from the testing or, in absence of

test data, major code recommendations.

There are four major types of shrinkage in concrete as follows;

Drying Shrinkage – is the reduction in concrete volume with time due to

moisture loss or, in similar meaning, and may be defined as the volumetric

change due to the drying of concrete. Drying shrinkage is mainly caused by

the movement of water from the cement gel to the atmosphere. This type of

shrinkage is considered significantly larger than the others;

Autogenous Shrinkage – is the reduction of concrete volume due to hydration

of the cement as well as various chemical reactions within the cement paste

(sometime also called Chemical Shrinkage). This kind of shrinkage is mainly

associated with the loss of water from the capillary pores due to the hydration

of cement so called self-desiccation that take place even if the specimen is

sealed or has not been exposed to the drying condition. Such shrinkage is

independent of the size of the specimens and usually smaller than the drying

shrinkage;

27

Plastic Shrinkage – is the shrinkage of concrete due to early age moisture loss

before, or shortly after, the setting of concrete. This can be minimised through

good construction practices;

Carbonation Shrinkage – is the kind of shrinkage that is caused by the

chemical reaction of various cement hydration products with carbon dioxide

in the air.

2.3.2.2 Creep

Creep is defined as the time dependent increase of strain in hardened concrete

subjected to sustained stress. The rate of creep and the final values of creep depend on

similar factors that simultaneously affect shrinkage, as well as applied stresses and the

age at loading of the specimen (Branson, 1977; Neville, Dilger and Brook, 1983). In

general, the quality of concrete has a direct effect on creep as higher quality concrete

tends to demonstrate less creep when given similar environmental exposure and stress

conditions. Also, the dependence of creep on major influencing factors such as

environmental humidity and volume-to-surface ratio of the exposed member are

reduced. The mix design and its constituent material also affect creep. For example,

creep is less when water-to-cement ratios are low or the maximum size of aggregate

increases (Gilbert, 2005).

The magnitude of the final creep and the rate of creep are influenced significantly by

loading history, particularly the magnitude and time of the first loading as young

concrete loaded aggressively creeps more as compared to similar concrete with mild

loading history (Gilbert, 2005). The rate of creep is usually high at early ages

following the loading, and decreases as time progresses until approaching near

constant values where creep reaches a final value (see Figure 2.2). Creep and axial

stress are linearly related for stresses up to 40 percent of the ultimate strength, which

covers most practical cases. For concrete subjected to stress exceeding that limit, the

creep and axial stress relationship becomes non-linear (Neville, 1995).

28

Figure 2.2. Creep concrete with different cements, loaded at 28 days (Neville, 1970)

There are two major components of creep: basic creep, cb, and drying creep, cd. Basic

creep is defined as the creep that occurs under hygral equilibrium condition where

there is no moisture exchange in the environment such as in the case of sealed

specimens or concrete stored in water. Drying creep is the additional creep that takes

place when drying starts. It results from exchange of moisture between the stressed

members and the ambient medium. Both creep components naturally exist in most

concrete structures. Total creep, ct, is thus equal to:

dbt ccc Equation 2.2

Creep is usually expressed as a dimensionless quantity (mm/mm). Sometimes, it is

more convenient to express creep in terms of creep per unit stress or specific creep,

C(t,s). Where c(t,s) represents creep at time t of a concrete member loaded first time

at time s and σ is applied stress (MPa). Specific creep can be written as:

),(),(

stcstC Equation 2.3

The summation of the specific creep and the specific elastic strain, εsp, occurring at

the initial application of a load is specified as the creep function, , which is given

by:

29

),()(, stCsst sp Equation 2.4

with

sE

s

c

elsp

1 Equation 2.5

where: εel(s) is the elastic strain occurring at the initial application of a load; Ec(s) is

the elastic modulus at the initial application of a load.

The ratio of the creep strain to elastic strain occurring at the initial application of a

load is defined as the creep coefficient, Ø(t,s), which is given as:

s

stcst

el

,, Equation 2.6

The reduction of stress in concrete kept under constant strain due to the creep effects

is defined as the relaxation. Relaxation and creep are closely linked physically.

2.3.2.3 Shrinkage and creep parameters

The parameters affecting creep and shrinkage have been

extensively reviewed by many researchers based on numerous

experimental data (Branson, 1977). These parameters and their

relating aspects are outlined in a diagram published in Pauw and

Chai (1967) as reproduced below in

Figure 2.3.

30

Figure 2.3 Creep and shrinkage strain variables (Pauw and Chai, 1967).

From

Figure 2.3, it can be seen that the parameters affecting creep and shrinkage can be

categorized in to three groups depending on their ground, i.e. factors related to

mechanical properties of concrete, i.e. material properties and curing conditions;

factors related to various environment conditions; and factors related to loadings

condition, i.e. loading history and stress condition of the interested members.

The first group includes: size of the interested member; water-cement ratio; cement

type; type of aggregate; and quantity of cement and aggregate which are related to

material properties, and duration; temperature and humidity of curing which are

related to the curing condition when the member was produced.

The second group includes factors related to the environmental condition surrounding

the member, i.e. relative humidity and temperature of the environment.

31

The third group being: age of concrete at first loading and when drying commenced;

number of load cycle; and duration of loading and unloading which are related to the

loading history, and distribution; magnitude and rate of stress which are depicting the

stress condition of the members.

The general trends of resulting creep and shrinkage behaviors by means of each of the

factors are described as follows;

Size and shape of member – For analysis purposes the effect of the shape and

size of member is usually referred in terms of volume/surface ratio following

the work of Hanson and Mattock (1966). Size of a member has significant

effects on both creep and shrinkage of concrete. Generally the larger specimen

displays less value of shrinkage and creep as compared to smaller specimens.

For creep, the sensitivity to a variation of size and shape is less than shrinkage

and the size effects exist only for specimens with thickness up to about 0.9 m.

The effects of size on creep occur only during the initial period during first

several weeks after the loading. After this the creep rate becomes equal in

specimens of all sizes. As for the effect of shape on creep and shrinkage,

logically, the shape of the specimen affects the moisture distribution within it.

In general, it was found that a cruciform section would have much higher rate

of initial creep than a circular section of a similar crossed sectional-area

(Chivamit, 1965). However, the shape factor is very much lesser important

than the size factor and can be neglected in most cases (Neville et al, 1983);

Water-cement ratio – It has been known that varying the water-cement ratio

results in a variation of the cement paste. Concrete with greater water-cement

ratio exhibits higher creep and shrinkage values as compared to those of

concrete with lower water-cement ratio (Neville et al, 1966). Note that creep is

approximately the square of the water-cement ratio (Lorman, 1940);

Cement type – the concerns on the effect of cement types are due to different

ingredients which have influence on the quality of the hydrated cement paste.

In general, creep and shrinkage are found to be inversely proportional to the

rate of hydration of the concrete. This means that at a given applied stress, the

32

concrete with a faster hardening capacity would yield a lower amount of creep

(Neville et al, 1983);

Type of aggregate – The presence of aggregate has influence on creep and

shrinkage of concrete due to its mineralogical character although the aggregate

itself does not undergo creep. The influence of aggregate type on creep and

shrinkage is due to the elastic modulus of aggregate. In general, the use of

aggregate with higher elastic modulus results in less creep and shrinkage


Quantity of cement and aggregate – This factor can be expressed as the

aggregate-cement ratio because their volumes as a part of concrete are

complementary. Thus increasing either of them would result in the decrease of

another and vice versa. This reason makes the effect of varying aggregate-

cement ratio to creep and shrinkage somewhat non-conclusive and unclear


Duration and type of curing – The extended curing duration reduces creep and

shrinkage of concrete. The type of curing has major influence on creep and

shrinkage particularly when steam curing is used. Steam curing reduces creep

and shrinkage by up to 50% due to the accelerated hydration of cement;

Curing humidity – The amount of water present when load is applied has some

influence on creep. A smaller amount of water present at the loading time

means a lesser creep potential and therefore lower curing humidity means

lesser creep;

Curing temperature – Elevated temperature occurs for the period of curing

causing a reduction in viscosity of the deformable phase in concrete (Ross et

al, 1958). This phenomenon results in higher amount of creep;

Environment humidity – The influence of relative humidity on creep and

shrinkage is large, however it decreases in larger specimens. The linear

relationship between creep and relative humidity was observed by the

investigators (Troxell et al, 1958; L’Hermite and Mamillan, 1968). In general,

concrete subjected to lesser relative humidity has more creep and shrinkage

potential than concrete under more humid conditions. However, this trend is

valid only when hygral equilibrium of specimen is not attained;

33

Environmental temperature – After the initial period of hydration, the

temperature of concrete would be depending upon the ambient temperature. It

has been observed that a greater rate of creep occurs in hot weather even if it is

quite humid. In general, creep increases with an increase in temperature up to

about 70˚C;

Age of concrete at first loading and when drying started – It has been found

that the creep of concrete loaded at relatively younger age is much greater

during the first few weeks after loading than those of concretes loaded at older

ages (Davis et al, 1934) but after about 28 days, creep is found independent of

the age of loading (Glanville, 1933). Also, for an age of loading exceeding 28

days, the influence of age of loading on creep can be neglected;

Number of loading cycles – Generally, creep of concrete specimen subjected

to cyclic loading is greater than creep of concrete under a sustained load,

equivalent to the mean stress of the cyclic load;

Stress distribution – Creep in tension is initially higher than in compression

under the same stress (Davis et al, 1937). Long term tension creep is equal to

long term compression creep of the same stress (Glanville and Thomas, 1939);

Magnitude of stress and rate of stress – The magnitude of strength can be

conveniently expressed in terms of a ratio of stress to strength of the concrete

member. The relationship of stress/strength ratio and creep is stated in Neville

(1959) that “for constant mix proportions and the same type of aggregate,

creep is proportional to the applied stress and inversely proportional to the

strength at the time of application of the load”. The rate of stress applied on

the concrete also affects creep. The creep of concrete loaded with higher rate

(loaded faster) is to be higher than those of concrete loaded relatively slower.

Major factors that have influence on creep of a concrete specimen include (after

Bazant and Chern, 1982);

Age at loading

Ambient humidity

Size of specimen

Compressive strength of concrete

34

Cement type

Water-cement ratio

Many other factors that are also reported to have some influence on creep are (after


Shape of specimen

Age at the start of drying

Drying duration (for creep)

Delay of loading after drying

Temperature effect on creep rate

Simultaneous effects of temperature and drying

Temperature effect on aging rate

Age at the start of heating

Amplitude of stress cycling (if any)

Mean stress in case of cycling

Nonlinearity at high stress

Cement content

Unit weight of concrete

Sand-aggregate ratio

Aggregate-cement ratio

Major factors that have influence on shrinkage of a concrete specimen include (after


Relative Humidity – Shrinkage decreases as the relative humidity increases;

Temperature – Shrinkage tends to increase at higher temperature;

Age at first exposure – Specimens with an older age when first exposed to

drying condition will exhibit less shrinkage than younger specimens;

Size of Specimen – Larger specimens display less shrinkage as compared to

smaller specimens;

Type and Curing Duration

35

Type of Cement

Cement Content – Concrete with high cement content tends to have a higher

shrinkage value;

Coarse Aggregate Content

Properties of Chemical Admixture – The influence of chemical admixture on

shrinkage varies depending mostly on the type and composition of the

admixture.

2.3.3 Reinforcement restraint

The effects of the presence of vertical reinforcement in concrete members include: the

reduction of creep and shrinkage strains as compared to the plain concrete of a similar

quality (Neville and Dilger, 1970), and the redistribution of stress from the concrete to

the steel causes gradual changes in the load carrying proportion of the two materials

over a long period of time, even though the overall load-bearing capacity of the

member remains constant. According to testing (Troxell et al, 1958), in a heavily

reinforced column, concrete stress will be fully transferred to the vertical

reinforcement; on the other hand, in a lightly reinforced column, stress in all steel will

be increasing until yielding.

2.3.4 Summary

The development process of axial shortening in columns and core walls was

diagrammatically presented and the definition for each of its components was

described. Theoritical background on instantanous shortening, creep, shrinkage and

reinforcement restraint were outlined. Also factors affecting axial shortening were

discussed.

From the review given in this section, it is noted that axial shortening can be broken

down into four major components each of which is affected by various factors. The

resultant axial shortening is taken as the sum of shortening due to each component.

36

It is evident that axial shortening is not a simple phenomena. Therefore axial

shortening could not be estimated precisely without taking into account all these

factors in a proper manner.

2.4 Prediction of axial shortening components.

The various prediction models for determining material properties of concrete which

are related to axial shortening, i.e. modulus of elasticity, creep and shrinkage, are

described in detail in the sub-sections to follow.

2.4.1 Prediction models for Young’s Modulus of Elasticity

Young’s Modulus of Elasticity can be defined as linearized instantaneous (1 to 5

minutes) stress-strain relationship determined as the slope of the secant drawn from

the origin to a point of 0.45 'cf on the stress-strain curve (ACI 209-92).

It is known that the modulus of elasticity increases in line with an increase in the

compressive strength of concrete. The inclusion of a density variable in elastic

modulus prediction equations decreases the variance of the modulus (Billinger and

Symons, 1997).

Although, many factors such as; volumetric proportion and modulus of elasticity of

the aggregate (Neville, 1995), loading conditions and loading rate, moisture state of

the specimen, aggregate stiffness, elastic modulus of cement-paste matrix, porosity

and composition of the transition zone are said to have influence on modulus of

elasticity. However, it is difficult to represent the relationship between elastic

modulus and time by a single equation over the whole range of measured value.

The following international codes, standards and researches have been examined:

37

2.4.1.1 Australian Standard

AS3600 (1994 and 2001) proposes Equation 2.7 for predicting the modulus of

elasticity for concrete grade below 50MPa.

cE = '043.05.1

cfw Equation 2.7

In Australia, knowing that the AS3600-2001 elastic modulus equation often

overestimates modulus of elasticity for high strength concrete; as per Mendis et al

(1997) suggested using Equation 2.7 with a modification factor as follow:

cE = '043.05.1

cfw Equation 2.8

where w is concrete density in kg/m3 and 'cf in MPa, 0.1'002.01.1 cf

Gilbert (2002, 2004, 2005) proposed the use of AS3600 equation when MPafc 40'

and the following Equation 2.9 when MPaf c 40' . This recommendation has been

adopted in the current Australian Standard for Concrete Structure, AS3600-2009.

5.112.0'024.0 wfE cc Equation 2.9

where w is concrete density in kg/m3 and 'cf in MPa

2.4.1.2 American Standard

ACI Committee 209 (ACI 209R-92, released in 1992 and re-approved in 2008) and

the ACI 318 building code’s recommendation (ACI 318-08) propose the following

equation, which was originally developed by Pauw (1960). This is based mainly on

test data of lightweight to normal weight concrete for estimation of concrete elastic

modulus.

38

cE = '043.05.1

cfw Equation 2.10

Where w is concrete density in kg/m3, 'cf in MPa.

A revised equation for determining modulus of elasticity was presented in the ACI

committee-363 report (ACI 363R-92). Equation 2.11 is recommended for concrete

with strength up to 83MPa.

9.6'32.3 cc fE Equation 2.11

Where cE is expressed in GPa and 'cf in MPa

2.4.1.3 European Standard

Current CEB- FIP Model Code (CEB-FIP, 1990) gives Equation 2.12 for estimating

elastic modulus for normal weight concrete ( w = 2400 kg/m3) valid for concrete with

compressive strengths between 12 to 80 MPa.

3/14 10/8'1015.2 cc fxE Equation 2.12

For 'cf in MPa and cE in MPa

In BS 8110: Part2, modulus of elasticity is given by Equation 2.14 as follows:

Given that:

cuc ff 85.0' Equation 2.13

Then:

39

cuc fE 2.020 Equation 2.14

where cE is in GPa, cuf = characteristic concrete cube strength in MPa and

2.4.1.4 International Research (listed by chronological order)

Branson and Christianson (1971) studied modulus of elasticity of concrete with

various strengths from 1500 psi to 6180 psi (10.3 MPa to 42.6 MPa) and unit weights

from 89 to 147 pcf (1425 to 2355 kg/m3). Equation 2.15 which gives elastic modulus

results slightly lower than the ACI equation was proposed.

'301635005.1

cc fwE Equation 2.15

where cE and 'cf are expressed in psi, w is concrete density in pcf

Carrasquillo et al (1981) published results from an investigation of the properties of

high strength concrete. The compressive strength of test concrete ranged from 21 to

76 MPa. By comparing prediction values with test results reported by the Portland

Cement Association, the verification of ACI 318-77 modulus of elasticity equation

was made. As a result, Equation 2.16 has been proposed to give better correlation

between the modulus of elasticity and the compressive strength of concrete compared

to the ACI 318-77 equation which has been found to overestimate modulus of

elasticity of concrete with compressive strengths over 41 MPa.

6900'3320 cc fE Equation 2.16

where cE is expressed in MPa and 21MPa < 'cf < 83MPa

40

The expression developed by Martinez et al (1984), Equation 2.17 was reported:

5.1)2320/)(6900'3320( wfE cc Equation 2.17

For 'cf , cE in MPa and w is concrete density in kg/m3

Ahmad and Shah (1985) reported test results on the structural properties of concrete

with strengths greater than 42 MPa. It has been found that an equation proposed

previously by ACI (ACI committee 363, 1984) overestimated modulus of elasticity of

concretes with compressive strengths up to 83 MPa by 20 percent.

An improved empirical formula has been proposed as follow:

325.05.26 )'()(108.33 cc fwxE Equation 2.18

where w is concrete density in kg/m3 and 'cf in MPa

Shih et al (1989) conducted an experimental study on the static Young’s modulus of

normal-strength (38 MPa) and high-strength concrete (62 MPa). A total number of 73

high-strength concrete cylinders, 18 normal-strength concrete cylinders (4 inch

diameter and 8 inch length) and 38 normal-strength concrete cylinders (6 inch

diameter and 12 inch length) were tested.

Based on data from other researchers and test results obtained, the following

empirical equation has been proposed for determining static modulus of elasticity.

1370'4660 cc fE Equation 2.19

Where cE is expressed in MPa and 21MPa < 'cf < 83MPa

41

Shih et al (1989) commented that the prediction of elastic modulus of light weight

concrete should be made with a formula postulated exclusively for light weight

concrete. As proposed formulae are recommended for normal weight concrete thus

the prediction using this formula may be inaccurate. Shih et al (1989) also questions

the suitability of the original equation proposed by Pauw (1960) due to advances of

modern concrete.

Tomosawa (1990) proposed the following relationship for elastic modulus prediction:

437.0)'(5627 cc fE Equation 2.20

For 'cf in MPa

Kakizaki (1992) proposed the approximate expression for modulus of elasticity of

concrete with strength between 80 to 140 MPa.

Equation 2.21

where cE is expressed in GPa and 'cf in MPa

Gardner and Zhao (1993) proposed Equation 2.22 for predicting modulus of elasticity

for design purposes

cmc fE 43003500 Equation 2.22

where cmf is mean compressive strength in MPa

The mean compressive strength can be related to characteristic compressive strength

at 28 days by the following relationship recommended by CEB-FIP (1990).

'65.3 cc fE

42

828

'

28 ccm ff Equation 2.23

where

28cmf = 28-days mean compressive strength, MPa

28

'

cf = 28-days characteristic, or specified, compressive strength, MPa

Billinger and Symons (1997) proposed Equation 2.24 which was rewritten as

Equation 2.25, a power relationship with 'cf of high strength concrete in the range of

49.2 to 108.5 MPa.

5.11617.0'0199.0 wfE cc Equation 2.24

2545.05.1'1113.0 cc fwE Equation 2.25

where w is in kg/m3 and 'cf in MPa

2.4.2 Prediction models for creep and shrinkage

2.4.2.1 Mathematical models.

This approach uses mathematical expressions, which are derived from short-term

experimental data to undertake long-term prediction of creep and shrinkage. In

general, the creep-time functions can be categorized into two categories: functions

with or without a finite value. The former are expressions in the form of exponential

and hyperbolic function, these forms are somewhat more commonly adopted. The

latter are the Logarithmic and Power expressions. Based on Neville et al (1983), a

succinct review of the four kinds of the creep-time expression is represented hereafter;

43

Creep prediction : Hyperbolic function

The early popular creep-time expression in the form of a hyperbolic function was

proposed by Ross (1937) and Lorman (1940). Ross’s equation being:

0

0

0''

,ttBA

tttt Equation 2.26

where:

0, tt = creep coefficient at age t for concrete loaded at age t0

'A , 'B = material constants

0, tt = duration of loading

Lorman’s equation takes form as shown:

0

0

0,ttn

ttmtt Equation 2.27

where: m, n are constants; σ is applied stress in psi

Ross and Lorman’s hyperbolic expressions were commonly used because they

provided a simple means for determining experimental constants and for computing

creep deformation. Also they give good prediction for long-term creep despite the

underestimation of creep at early ages. The modifications of Ross’s equation found in

literature include a creep expression proposed by Ali and Kesler (1964), and creep

models found in the CEB-FIP code model (1978, 1990).

The major variation of Ross’s equation was the hyperbolic-power expression for

creep adopted by Branson and Christianson (1971) and Bazant et al (1976), i.e.

uc

c

Cttd

tttt

0

00, Equation 2.28

44

where:

Cu = ultimate creep coefficient

c, d = constants

The hyperbolic-power expression of Branson and Christiason (1971) was later used as

a basis of the creep equation of the ACI 209 code models (1971 and 1992), as well as

the work presented in Bazant et al (1976) which was later incorporated into the BP

model (Bazant and Panula, 1978).

Creep prediction : Exponential function

The relationship of creep strain, applied stress and some material constants in the

form of exponential function was first proposed by Thomas (1930). This form of

equation assumes that creep is proportional to the amount of the remaining potential

creep. Thomas’s equation takes the form:

xxaatA

creep eCtt 21, 20 Equation 2.29

where:

2C , 2A , a , x = material specific constants; e = the napierian base

By means of visco-elastic behaviour, McHenry (1943) proposed the following

expression for creep:

rt

creep ett 1, 20 Equation 2.30

where:

= applied stress

2 , r = material specific constants

45

McHenry (1943), having difficulty verifying his equation with the experimental data,

declared that an extra exponential term be added to achieve a better accuracy.

McHenry’s creep expression then became:

tmKprt

creep eeett ''

20 1'1, Equation 2.31

where:

β’, p’, K and m’ represent additional constants.

According to Neville et al (1983), McHenry (1943)’s equation gives more accurate

prediction compared to equation proposed by Thomas (1930). However, McHenry

(1943) noted that the equation is not recommended for concrete loaded at age less

than five days and some concrete subjected to the load more than one year as it

appears to give inaccurate prediction.

As noted by Neville et al (1983), there have been many creep-time expressions

proposed by researchers after McHenry (1943). However, none of the current code

models were found using exponential expression as the basis for their equation

(O’Moore, 1996).

Creep prediction : Logarithmic function

By assuming that the rate of specific creep is inversely proportional to time, a creep-

time expression in the form of Logarithmic function was proposed by the US Bureau

of Reclamation (1956) for estimating creep when the ratio of stress to strength is not

exceeding 0.35 (Neville et al, 1983), i.e.

1log, 000 tttAttC Equation 2.32

where:

46

0, ttC = specific creep

0tA = specific coefficient of concrete and age at loading

The logarithmic expression generally underestimates creep at early age in a similar

way to the hyperbolic expression. The logarithmic equation does not have an ultimate

value. None of the codes were found based on this function (O’Moore, 1996).

Creep prediction : Power function

The power function for creep prediction was originally presented by Straub (1930).

This form of creep-time expression tends to overestimate the long-term creep strains

(O’moore, 1996). The equation takes the form:

1

010,B

creep ttAtt Equation 2.33

or in rectified form being:

)log(),(log 0110 ttBAttcreep Equation 2.34

where:

0, ttcreep = creep strain

1A , 1B = material specific constants

The major adoption of Straub’s equation was conducted by Bazant and Osman (1976)

to give the Double Power Law for basic creep of concrete which was later used to

form the integral part of the BP model (Bazant and Panula, 1978). Subsequent

research by Bazant and Chern (1984, 1985a, 1985b) were conducted to improve the

accuracy of the Double Power Law resulting in multiple revisions of the original basic

creep model. Extra complexity and sophistication were added to the original

expressions in order to boost extra fit to the data. The outcomes were the Double

Power Logarithmic Law (Bazant and Chern, 1984), the Log Double Power Law

(Bazant and Chern, 1985a) and the Log Double Power Law (Bazant and Chern,

47

1985a). However, these equations are not applicable to loads of very short duration

and drying creep is not included.

Shrinkage prediction : Hyperbolic-Power function

Unlike creep, it has been agreed that shrinkage of a concrete specimen will be liable

to a final values. The shrinkage-time functions were therefore found expressed only in

the form of the Hyperbolic-Power function and exponential function (Neville et al,

1983).

ACI209 (1982) and Bazant et al (1976) both based their creep expressions on the

Hyperbolic-Power function. Work proposed by Bazant et al (1976) later found its way

into the famous BP model (Bazant and Panula, 1978). The general form of

hyperbolic-power shrinkage time function takes the form:

''

23

,

m

sh

sh

shshttBA

tttt Equation 2.35

where:

shsh tt, = shrinkage at age t for drying occurred at age tsh

3A , 2B , ''m = material specific constants

Shrinkage prediction : Exponential function

According to a review by O’Moore (1996), early researchers that derived shrinkage

expressions based on exponential function were Aroutiounian (1959) and Lyse

(1960). Their expressions were based on the following function:

shttc

shshshsh ettt 31, Equation 2.36

where:

48

)( shsh t = ultimate shrinkage strain

C3 = material specific constants

Based on Equation 2.36, Wallo and Kesler (1968) proposed another shrinkage

expression with the experimental constants as follows:

svtt

shshshshhettt

//1.065.0

1, Equation 2.37

where: v/s is the volume-surface ratio in inches.

Unlike the hyperbolic-power expression, exponential expression has never been

adopted by major codes for shrinkage prediction.

2.4.2.2 Generic/Code models.

ACI 209R-92 model (American Concrete Institute code model)

In the 1970s, American Concrete Institute released a report consisting of methods for

predicting creep, shrinkage and temperature effects in a concrete structure based

largely on the work presented by Branson and Christiason (1971). This work has been

updated several times and the most recent revision was proposed in 1992 (ACI209R-

92). The American code model aimed at providing a unified method for designers as a

comprehensive approach to the problem of volume changes in concrete. Its vast scope

included the prediction of strength and elastic properties of concrete, theory for

predicting creep and shrinkage of concrete, assumptions and theories for the analysis

of time dependent effect in concrete structure and some numerical examples.

In this method, the effect of moisture changes, sustained loading and temperature

changes are taken in to account while factors which are difficult to quantify such as

non-homogeneous nature of concrete caused by multi stages of construction, the

history of water content are not included in the prediction. Acknowledging that the

nature of the problem is highly complex due to a relatively large number of factors

involved, the procedures presented in the ACI 209R-92 report are simplified and good

49

for design purposes. In 2008, the long existing ACI 209R-92 still remains the most

up-to-date American code for creep and shrinkage as it was recently reapproved in

2008 by ACI 209 committee.

ACI 209R-92 shrinkage equations

The ACI 209 code model defines the shrinkage strain εsh(t,tsh,0) in the form of

hyperbolic function. For specimens kept under the standard condition, shrinkage at

any time, t, for moist curing concrete is expressed as:

sh

sh

sh

shshtt

tttt

)(35

)(),(

0,

0,

0, Equation 2.38

For steam curing, shrinkage at any time, t, the expression becomes:

sh

sh

sh

shshtt

tttt

)(55

)(),(

0,

0,

0, Equation 2.39

where: εsh(t,tsh,0) = shrinkage strain (mm/mm) at time t (days) of a specimen that

commences drying at time tsh,0 (days); sh = ultimate shrinkage strain (mm/mm).

For non-standard conditions, values of ultimate shrinkage strain sh can be

determined by using correction factors as directed in Equation 2.40.

'

8

'

7

'

6

'

5

'

4

'

3

'

2780sh Equation 2.40

The existing six adjustment coefficients intend to account for the differences of

relative humidity )( '

2 , average thickness and volume/surface ratio )( '

3 , slump )( '

4 ,

fine aggregate/total aggregate ratio )( '

5 , air content )( '

6 , cement content )( '

7 and

50

duration of initial curing )( '

8 . They are used to cover for the effect of different

concrete strength, composition and physical condition. However, the effects of coarse

aggregate to shrinkage have not been addressed in this model. Although all correction

factors are applied to an ultimate value, it is noted that they may be used in

conjunction with data from tests.

ACI 209R-92 creep equations

Creep coefficient t of normal strength concrete under standard conditions at time t

(days) after loading can be predicted using the following equations:

utt

t60.0

60.0

10 Equation 2.41

cu 35.2 Equation 2.42

where: γc = product of the applicable correction factors.

Correction factors for creep take into account the effects of loading age, initial moist

curing, volume-to-surface ratio, ambient relative humidity and temperature other than

21 degrees Celsius.

It can be seen that the ACI 209R-1992 model is a sophisticated model that takes into

account the effects of many significant factors. However, the use of ACI 209 model

with high strength concrete may induce significant error as the ACI model was

derived from material data of normal strength concrete and the effects of concrete

strength have not been taken into account. In general, the ACI 209R-1992 model

underestimates early shrinkage as well as creep but it overestimates long-term

shrinkage (Garder and Zhao, 1993).

CEB-FIP 1990 (Euro-International Concrete Committee code model)

51

In 1990, CEB-FIP proposed a major revision of the original 1978 model code, which

had been used extensively in European countries as a basis of their independently

established national design codes. Despite its eighteen years of age in 2008, CEB-FIP

model code 1990 remains the most up-to-date version of this series since there is no

modification of creep and shrinkage prediction since then.

Comparing the CEB-FIP model code 1990 with ACI 209-1992, it has been found that

both methods express their creep and shrinkage prediction formulae in a hyperbolic

power form (O’Moore and Dux, 1995). However, input data from experiments are not

necessary for the prediction, as is the case in ACI 209-1992. Parameters taken into

account for creep and shrinkage prediction are; characteristic compressive strength,

dimensions of the member, exposed relative humidity, age at loading, duration of

loading and type of cement.

Creep and shrinkage prediction methods as presented in CEB-FIP model code 1990

are represented herein. However, it is noted that the practice of this method is limited

by the following conditions: The model is valid for structural concrete with

compressive strength between 12 and 80 MPa, noted that using the code for concrete

with compressive strength greater than 50 MPa should be with caution; Applied stress

must be less than 40 percent of compressive strength at the day of loading; Exposed

relative humidity between 40 and 100 percent; Exposed temperatures should be in the

range of 5°C to 30°C. Also the CEB-FIP model code 1990 does not cover concrete

made with chemical admixtures.

CEB-FIP 1990 shrinkage equations

The expression for the shrinkage strain of concrete εcs(t,ts) at time t is given as follow:

)(),( sscsoscs tttt Equation 2.43

52

where: εcso is the notional or ultimate shrinkage strain (mm/mm); βs(t-ts) represents the

development of shrinkage with time; ts is the age of concrete at the commencement of

shrinkage.

The notional shrinkage can be obtained from:

))(( RHcmscso f Equation 2.44

with:

610))1450/9(10160()( cmsccms ff Equation 2.45

ARHRH 55.1 Equation 2.46

3)100/(1 RHARH Equation 2.47

where: εs(fcm) = function accounting for the effects of concrete compressive strength

and cement type on shrinkage; βRH = coefficient for relative humidity; fcm = mean 28-

day compressive strength (MPa); βsc = coefficient depending on type of cement.

The development of shrinkage with time is given by:

5.0

2100/350

)(s

sss

tth

tttt Equation 2.48

where: h = effective thickness of the concrete member.

CEB-FIP 1990 creep equations

The creep coefficient ),( 0tt of a concrete specimen loaded at time t0 at any time t

may be calculated from:

53

)(),( 000 tttt c Equation 2.49

where: 0 is the ultimate creep coefficient (mm/mm); )( 0ttc represents the

development of creep with time.

The ultimate creep coefficient is given as:

)()( 00 tf cmRH Equation 2.50

with:

3/1)100/(46.0

100/11

h

RHRH Equation 2.51

5.0)10/(

3.5)(

cm

cmf

f Equation 2.52

2.0

0

0)(1.0

1)(

tt Equation 2.53

where: RH = coefficient for relative humidity; RH = relative humidity (%); h =

notational size of member (mm); )( cmf = coefficient for concrete strength; )( 0t =

coefficient for the age at loading.

The CEB-FIP90 code model states that Equation 2.49 and 2.50 would be the best

representative of creep at a notional temperature of 20ºC up to 70 years of age

according to the reviewed data. Adjustments for concrete at other temperatures may

be made according to the given process, which in explicit forms is not reproduced

here but can be found in the code.

54

According to Gardner and Zhao (1993), the CEB-FIP90 model underestimates

shrinkage, however creep prediction was found accurate.

AS3600-2001 (Australian Standard code model for concrete structures)

The Australian standard for concrete structure AS3600-2001 represents the third

major revision of the Australian code in the last twenty years. The contents are still

mainly based on the superseded draft code SAA Committee BD/2 (1986) and

AS3600-1994.

The scope of the standards set upon two criterions: characteristic compressive

strength and saturated surface-dry density of the concrete. The applicable range of

compressive strength of concrete was originally from 20 to 50 MPa and density from

1800 kg/m3 to 2600 kg/m

3 in the draft code SAA Committee BD/2 (1986). The

range of applicable concrete strength remained so in the AS3600-1994 whilst the

upper limit for concrete density was extended to 2800 kg/m3. Thereafter the

amendment was made into the current Australian code (AS3600-2001) for the

extension of the upper limit of the applicable concrete strength from 50 MPa to 65

MPa.

The Australian codes have always been using a graphical method, mostly similar to

the method used previously in CEB-FIP 1970 code model, to express coefficients that

applied on the ultimate value of shrinkage strain and creep coefficient to

accommodate for the influence of local Australian concretes and environments. This

made these codes relatively simple to use.

AS 3600-2001 shrinkage equations

According to AS3600-2001, the ultimate shrinkage strain bcs. which is called

differently therein as the basic shrinkage strain is modified by the application of the

55

environmental factor k1. The design shrinkage strain εcs at any time may be calculated

from:

bcscs k .1 Equation 2.54

Factor k1 takes into account the effects of both the type of exposed environment and

the size of specimen to shrinkage. It can be determined directly by interpolation from

the charts given in Section 6.1.7.2 of the code. In the absence of test data, AS3600-

2001 recommends the ultimate value of shrinkage strain to be taken as 850x10-6

.

AS3600-2001 states that the accuracy of its shrinkage prediction may be within a

range of ± 40% of the predicted values. However, Gilbert (2002) argues that actual

error experienced when predicting shrinkage with AS3600’s equation would be much

greater than that as, AS3600 model does not take into account the effects of the

quality and compositions of the concrete. Gilbert (2002) also states that, in reality,

factor k1 tends to overestimate the effects of member size and underestimate the

amount of early-age shrinkage.

According to the view point of Gilbert (2002), shrinkage prediction using model

suggested by ACI209R-92 for steam curing condition (Equation 2.39) may be

justified for concrete with a low water-cement ratio (<0.4) and with good quality plus

well graded aggregates or otherwise Equation 2.54 may underestimate shrinkage by at

least 50%.

AS 3600-2001 creep equations

The definition of the φcc.b or the basic creep factor as called by the AS3600-2001 code

model is given therein as the ratio of final creep strain to elastic strain for a specimen

loaded at 28 days under a constant stress of '4.0 cf . To estimate the value of creep

coefficient at any time, the standard specifies modifying the value of the ultimate

creep coefficient using two factors: the creep-time curve k2 and the maturity

coefficient k3. The two factors are used to accommodate the influence of the local

56

conditions of the concrete to creep in the similar way of modifying the ultimate

shrinkage strain using factor k1.

The creep coefficient at any time φcc, known in AS3600-2001 as the design creep

factor, is given as:

bcccc kk .32 Equation 2.55

Factor k2 takes into account the influences of the hypothetical thickness, the

environment and the time after loading to creep. Factor k3, expressed as a function of

the strength ratio (fcm/f’c), depends on the age at first loading. Values of these two

factors can be obtained from the design charts given in section 6.1.8.2 of the code.

AS3600-2001 states that the accuracy of its creep prediction should be within a range

of ± 30% of the predicted values providing that the cement used has sulphate content

within the range of 2% to 3.5%; the concrete member is not exposed to a prolonged

temperature over 25ºC; and sustained stress is not over '5.0 cf . Gilbert (2002) suggests

that a major revision of the code be made to narrow down the range of prediction

error.

The AS3600-2001 creep model is known to underestimate the early development of

creep in structural members. This is due to the inaccurate shape of factor k2 and the

factor k3 that tends to underestimate creep of specimens loaded at very early age. The

model also found underestimating the creep of concrete with compressive strength in

the range of 40 MPa to 50 MPa (Gilbert, 2002).

AS 3600-2009 (current Australian Standard code model for concrete

structures)

The empirical formulas for creep and shrinkage in the new code were originally

proposed by Gilbert (2002) and subjected to minor refinement as presented in Gilbert

57

(2004) and Gilbert (2005). The advancements of the AS3600-2009 model include the

extension of the limit of applicable range of concrete grade from 65 MPa to 100 MPa

and revisions of the procedures for estimating strength and deformation characteristics

of concrete.

AS 3600-2009 shrinkage equations

The shrinkage model proposed in Gilbert (2002 and 2005) which is adopted by AS

3600-2009 includes two major amendments as compared to AS3600-2001. The first

amendment is that Gilbert’s shrinkage model redefines total shrinkage εcs as the

summation of two shrinkage components: the endogenous shrinkage εcse which is the

summation of thermal and chemical shrinkage; and the drying shrinkage εcsd. Total

shrinkage is therefore given as follow:

csdcsecs Equation 2.56

The Endogenous shrinkage depends on compressive strength '

cf and may be taken as:

)0.1( 1.0* t

csecse e Equation 2.57

6'* 1050)0.106.0( ccse f Equation 2.58

where: *

cse = final endogenous shrinkage (mm/mm); '

cf = compressive strength

(MPa); t = time after pouring of concrete (days).

The second amendment is that the effect of concrete strength to drying shrinkage is

now accounted for in Gilbert’s shrinkage model. The drying shrinkage is now given

by:

58

bcsdcsd kk .41 Equation 2.59

with

*.

'

. )008.00.1( bcsdcbcsd f Equation 2.60

in which: k1 is the size factor; k4 is the factor for different environmental conditions;

bcsd . is the basic drying shrinkage strain (mm/mm); *.bcsd is the ultimate shrinkage

strain (mm/mm) depending on the quality of local aggregates.

Values of k1 can be either interpolated from design charts or computed using formulae

given in section 5.3 in Gilbert (2002) which is also given in Figure 3.1.7.2 on page 40

of the AS 3600-2009 code. The values of k4 were specified as 0.7, 0.65, 0.60 and 0.50

for an arid environment, an interior environment, a temperate environment and a

tropical/coastal respectively. The values of k4 were specified as 800x10-6

, 900x10-6

and 1000x10-6

for concrete located in Sydney and Brisbane, Melbourne and elsewhere

respectively.

AS 3600-2009 creep equations.

Also adopted entirely from Gilbert (2002), the AS 3600-2009 creep model is based on

the model recommended in AS 3600-2001. However, the existing factors k2 and k3

have been revised and two additional factors k4 and k5 are now included in the new

model.

The creep coefficient is given by

bcccc kkkk .5432 Equation 2.61

where: k2 = factor of the creep-time curve; k3 = maturity coefficient; k4 = factor

accounting for the environment; k5 = factor accounting for the reduction of the

59

influence of humidity and specimen size on creep when there is an increase of

concrete strength.

Values of k2 and k3 can be estimated using charts or formulae given in Gilbert (2002)

which also can be found in Figure 3.1.8.3(A) on page 42 in AS 3600-2009 code.

Values of k4 are specified as for shrinkage given in previous section. Factor k5 can be

taken as 1.0 for concrete with '

cf up to 50 MPa. When '

cf is in between 50 MPa to

100 MPa, k5 is expressed as:

'

335 )0.1(02.0)0.2( cfk Equation 2.62

This creep model and the AS3600-2001 model are in very close agreement for low

strength concrete (Gilbert, 2002). However, Gilbert (2002) states that his proposed

creep model provides relatively more accurate prediction for higher strength concrete

despite its suggestive range of ±30%.

Sakata Model (Sakata, 1993)

Sakata (1993) proposed new prediction equations for creep and shrinkage of concrete

which were derived using a statistical method on many experimental data from both

the literature and his own experiment. Sakata’s test data comprises the test results of

160 different beam specimens for shrinkage and another 104 beam specimens for

creep respectively. Note that compressive strength of the test specimens is in a range

of 10 MPa to 58MPa.

Sakata Model shrinkage equations

According to Sakata (1993), shrinkage can be predicted from the product of an

ultimate shrinkage and a shrinkage-time exponential expression of the shrinkage-time

relationship, i.e.

60

56.0

00 )(108.0exp1),( tttt shsh Equation 2.63

5210))10/)/(ln(5))(ln(38)100/exp(1(7850 SVwRHsh Equation 2.64

where: sh = ultimate shrinkage strain (x10-5

mm/mm); RH = relative humidity (%);

w = water content of the concrete (kg/m3); V/S = volume-to-surface ratio; t0 = age of

concrete when drying commenced (days).

Sakata Model creep equations

Similar to shrinkage, the creep-time curve of the Sakata model is also expressed by an

exponential equation. Sakata’s creep equation takes the following form:

6.0

0 '09.0exp1,', ttDCBCtttC Equation 2.65

where: C(t,t’,t0) = specific creep (x10-5

cm2/kgf); BC = basic creep (x10

-5 cm

2/kgf);

DC = drying creep (x10-5

cm2/kgf); t = age of concrete(days); t’ = age of concrete at

loading (days) and t0 = age of concrete when drying commenced (days).

The expressions of the basic creep and the drying creep are given, viz.

67.04.225 'ln/105.1 tCWWCBC Equation 2.66

3.0

0

36.02.24.12.4100/1/ln/0045.0 tRHSVWCCWDC Equation 2.67

where: C= cement content (kg/m3); W = water content (kg/m3); RH = relative

humidity (%) and V/S = volume-to-surface ratio (cm)

From Sakata’s investigation, creep and shrinkage predictions using the proposed

model agree well with both CEB-FIP experimental data and data collected by Sakata

61

(1993). Better fit when using Sakata model as compared to other models (ACI209,

CEB-FIP 70, CEB-FIP 78, CEB-FIP 90 and BP model (Bazant and Panula, 1978) was

observed.

B3 Model (Bazant and Baweja, 1995a, 1995b, 1995c)

The B3 model is one model amongst the many that has received publicity in the

literature in recent years. It represents the third and the most current updated of creep

and shrinkage models previously developed at Northwestern University, i.e. BP-KX

model (Bazant et al, 1991, 1992; Bazant and Kim, 1991, 1992a, 1992b, 1992d) and

BP model (Bazant and Panula, 1978). The application of the model is limited for

concrete with the following parameter ranges: compressive strength, 17 MPa ≤ f’c ≤

69 MPa; ratio of water to cementitious material, 0.3 ≤ w/c ≤ 0.85; cement content, 160

kg/m3

≤ c ≤ 721 kg/m3; ratio of aggregate to cement, 2.5 ≤ a/c ≤ 13.5.

B3 Model shrinkage equation

Mean shrinkage strain in the cross-section at time t is given by:

)(),( 0 tSktt hshsh Equation 2.68

628.0'1.2

21 10)270)()(26( csh fw Equation 2.69

31 hkh Equation 2.70

sh

tttS 0tanh)( Equation 2.71

62

where: εsh(t,t0) = shrinkage strain (mm/mm) at time t (days) of a specimen that

commences drying at time t0 (days); sh = ultimate shrinkage strain (mm/mm); α1 =

coefficient depending on type of cement; α2 = coefficient depending on curing

condition; w = water content of concrete (lb/ft3); kh = humidity dependence; h =

relative humidity (%); S(t) = time curve of the shrinkage; τsh = size dependence

B3 Model creep equations

Based on material behavior phenomenological theory, the B3 model expresses creep

in term of the compliance function and the total creep strain was divided into two

components, basic creep )',(0 ttC and drying creep ),',( 0tttCd .

The compliance function of the basic creep is given by:

'ln'1ln)',()',( 4

1.0

320t

tqttqttQqttC Equation 2.72

where: q2 = ageing viscoelastic compliance; q3 = non-ageing viscoelastic compliance;

q4 = flow compliance; Q(t,t’) = binomial integral component.

The drying creep is expressed by:

2/1

50 '8exp8exp),',( tHtHqtttCd Equation 2.73

6.01'5

5 1057.7 shcfq Equation 2.74

)()1(1)( tShtH Equation 2.75

where: h = relative humidity (0≤h≤1); sh = ultimate shrinkage strain (mm/mm); S(t)

= shrinkage time curve.

63

The use of a compliance function in B3 model tends to reduce the risk of error due to

inaccurate modulus values thus improve the accuracy of the prediction. However, the

model is not much favored by practicing engineers due to its relative complexity and

the lack of some required input data that may not be available during the design

period (Huo et al, 2001).

GL2000 model (Gardner and Lockman, 2001)

GL2000 model has been proposed by Gardner and Lockman (2001) to be used as a

simple design-office procedure for creep and shrinkage prediction. Therefore, the

given formulae are simple and explicit (even though according to Huo et al (2001),

the equations of correction factors in the GL2000 model are more complicated as

compared with the ACI 209R-1992), whilst most of the required inputs are normally

available at the design stage such as: specified concrete strength; volume-to-surface

ratio; age of loading; ambient relative humidity. For convenience, specified values are

given for other inputs that are also required but not known to the design engineer such

as: the ultimate shrinkage and creep coefficients.

The model proposed is the development of GZ model (Gardner and Zhao, 1993) and

its revision by Gardner (2000). The old GZ model takes into account the effects of

ambient relative humidity, strength of concrete at 28 days, drying age after moist

curing and size of members. It had better accuracy when compared to the ACI 209R-

1992 model and was good for concrete with mean compressive strength up to 75 MPa

whilst the updated version the GL2000 model is good for concrete with mean

compressive strength not exceeding 81 MPa providing that self-desiccation is not

experienced.

GL2000 Model shrinkage equations

Like the CEB-FIP 1990 model code, the 28-days mean compressive strength fcm28 is

used instead of the 28-days characteristic strength fck28. In absence of test data, the 28-

days mean compressive strength can be determined using the following relationship:

64

51.1 2828 ckcm ff Equation 2.76

The shrinkage model as given in Gardner and Lockman (2001) is reproduced here as

follow:

)()( thshush Equation 2.77

6

2/1

'

28

104350

1000cm

shuf

K Equation 2.78

418.11)( hh Equation 2.79

6

2/110

)/(97

)()(

SVtt

ttt

c

c Equation 2.80

where: εshu = ultimate shrinkage strain (mm/mm); β(h) = coefficient accounting for the

effect of humidity on shrinkage; β(t) = coefficient accounting for the effect of time on

shrinkage; h = humidity expressed as a decimal (h≤1); tc = age of concrete when

drying started (days); V/S = volume-to-surface ratio (mm); K = factor accounting for

cement types specified as, 1 for type I, 0.7 for type II and 1.15 for type III.

GL2000 Model creep equations

Like B3 model (Bazant and Beweja, 1995a, 1995b, 1995c), the GL2000 model also

expresses creep in terms of a compliance function, i.e.

00 ,1

. ttCE

ttJcmto

Equation 2.81

with

65

28

28

0,cmE

ttC Equation 2.82

where: Ecmto = elastic modulus at time of loading (MPa); C(t,t0) = specific creep; 28 =

creep coefficient; Ecm28 = elastic modulus at 28 days (MPa); t0 = age of concrete at

time of loading (days).

Assuming that drying creep would be zero when relative humidity reaches 96%

(hygral equilibrium), creep coefficient 28 is given by:

5.0

2

0

02

5.0

0

0

5.0

0

3.0

0

3.0

0

28

/15.0086.115.2

7

7

142

SVtt

tth

tt

tt

ttt

tt

tc Equation 2.83

The term ct accounting for the Pickett effects or the drying before loading, which

reduced both basic and drying creep, thus it is taken as 1 if t0 = tc. When ctt0 ,

ct is given by:

5.05.0

2

0

0

/15.01

SVtt

ttt

c

c

c Equation 2.84

According to comparisons conducted by Gardner and Lockman (2001), the prediction

using Equation 2.83 and 2.84 fits well with extensive data of creep and shrinkage

from RILEM databank. Error of the most predictions fall in the range of ±30% as

compared to the test data. Comparisons of the predictions of other major models such

as: ACI209, CEB-FIP1990 and B3 with the same test data showed that GL2000 is

more superior than the others.

66

2.4.3 Summary

From the various mathematical functions given above, it was evident that there are

many models and codes proposed for prediction of time dependent properties of

concrete by researchers from many regions. In this respect, certain codes of practice

have been well established (for example, ACI standards for America, CEB-FIP for

European and AS for Australia) and continue to be updated from time to time when

new proposals are justified (such as Gilbert model adopted in AS 3600-2009 and

GL2000 model expected to be included in the next edition of ACI 209, respectively).

For practices conducted in other regions, national codes may be adopted if available.

Theses codes are usually well-known regional codes as mentioned above which are

modified to suit the properties of local concrete. In the absence of national codes,

creep and shrinkage may be estimated by using any proposed equation in conjunction

with short-term test data.

However, the use of any codes/models to estimate time dependent properties of

concrete are limited due to the following constrains:

The equations are limited in terms of application since they were derived from

small batches of specific concrete.

Changing of mix proportions, material sources, construction environment

affects the performance and accuracy of these equations.

These equations are not recommended for direct prediction of total axial

shortening because uncertainties occurred during construction such as loading

history is not taken into account.

67

2.5 Analysis of axial shortening in structural members under constant and

variable stress

There have been studies on effects of axial shortening on concrete structure where

actual construction conditions are applied and the effects of instantanous shortening

and creep and are considered as a combined variable. The well-established axial

shortening models proposed in the literature i.e. Effective modulus method (Faber,

1927), Rate of creep method (Dischinger, 1937), Superposition method (McHenry,

1943) and Age-adjusted effective modulus method (Bazant, 1972) are reviewed in the

sub-sections to follow.

These models are different from material models reviewed earlier in Section 2.4 as

they involved more rigorous analysis and the output is the resultant axial shortening

not just any of its component. Together with concrete properties predicted using a

selected material model, these axial shortening models are employed when axial

shortening analysis is required.

2.5.1 Effective modulus method (EM)

Developed by Faber (1927), the concept of EM method was centered on the use of the

then newly proposed effective modulus Ee in an elastic analysis of time-dependent

effects of concrete in structural members. The effective modulus can be obtained after

the reduction of elastic modulus by a factor [1+Ø(t,t0)] as follows:

00 ,1/ tttEEe Equation 2.85

where: E(t0) is elastic modulus at the application of the loading; 0, tt is the creep

coefficient at time t of a concrete member loaded at time t0.

The EM method is the simplest but least accurate as the effects of varying stress

history have not been taken into account (Branson, 1977). As such, only a single

strain-time curve under constant stress applied at age t = t0 has been used. Typically,

68

enormous error can be expected when using the EM method except where stress is

consistent or when the aging of the concrete is negligible as in the case of old concrete

(Neville et al, 1983).

2.5.2 Rate of creep method (RC)

This method was developed by Dischinger (1937) based on a finding of Glanville

(1930) that rate of creep is independent of the age at loading. Dischinger (1937)

developed a first-order differential equation to estimate structural creep by assuming

that the rate of creep is constant with time and shrinkage is developed at the same rate

as creep. The former assumption leads to an implication that since creep curves for all

ages at loading are parallel thus only one curve depicting either specific creep or creep

coefficient at initial loading (t0) would be required for the estimation of creep for

other subsequent loading (t>t0). The differential equation for solving axial shortening

problems, according to Dischinger (1937), takes the following form:

d

d

d

d

tEtE

t

d

d sh1

0

Equation 2.86

In contrary to the deficiency of the EM method discussed earlier, the RC method

generally gives good prediction for concrete loaded when young. The major

shortcoming of the RC method was the overestimation of creep under a decreasing

stress and vice versa. These are caused by the assumption of the parallel lines which

was found not agrees with observations and the fact that no allowance was made in

the method for creep recovery.

As discussed by Neville et al (1983), the RC method is now considered obsolete due

to the following deficiencies: under estimation of creep of old concrete, lack of

accounting for creep recovery, over estimation of creep under a decreasing stress.

2.5.3 Superposition method

69

McHenry (1943) applied the principal of superposition in order to postulate a

shortening analysis method for concrete under varying stress. McHenry (1943)’s

assumption is “Additional strains at time t caused by an incremental stress at age 't

are independent of any effects from stress applied earlier or later” which means each

applied load increment would initiate a creep strain as if it was the only loading

applied to that column. The expression of total creep strain at time t, due to the

varying stress is given by:

''

',,1

0

'

0

0

0 dtt

ttttt

tEt

t

t Equation 2.87

where: ', tt = creep function at a concrete age t for a load applied at time 't

The concept of this method is to take every stress increment to the stress history by

using a known creep function in given differential equation which can be solved by

the help of the numerical solution. As stated by Neville et al (1983), the accuracy of

this method is dictated on representation of creep function for a particular concrete

together with number of steps applied in the numerical solution.

The flaws of the superposition method include: the underestimation of creep under

decreasing stress, this is due to the assumption that creep recovery is equivalent to the

corresponding creep itself; and the laborious processes involved in the calculation,

requiring computer assistance. However, the superposition method usually yields

better prediction than the EM and RC methods. After all, most of the current

computation methods deal with axial shortening of concrete structure assumes the

validity of the superposition of linear viscoelasticity (Staquet and Espion, 2005).

2.5.4 Age-adjusted effective modulus method (AEMM)

The AEMM method was developed by Bazant (1972), based on work conducted by

Ross (1968) who proposed major refinements to the well known EM method, i.e. the

allowance for the variation of elastic modulus and an unbounded final value of creep,

70

and the introduction of the concept of an aging coefficient, χ(t,s). As an extension of

the EM method, the AEMM method is made capable for computing strain of concrete

members under a varying stress and, vice versa, stress under a varying strain.

In the AEMM method, the aging coefficient is used to account for the effects of aging

on the ultimate value of creep for varying stress history after the application of the

original load. The aging coefficient is dependent on the value of the creep coefficient

and on the age of loading (Trost, 1967). According to Bazant (1972), The aging

coefficient is a product of the relaxation function of the concrete, R(t,s), which is

related with creep. The aging coefficient is defined as;

),(

1

),()(

)(),(

ststRsE

sEst

c

c Equation 2.88

where: Ec(s) = elastic modulus of concrete at first loading; R(t,s) = the relaxation

function; Ø(t,s) = creep coefficient at time t of a concrete loaded first time at time s.

The value of the aging coefficient may be determined from curves given in the

literatures, viz. Bazant (1972); CEB-FIP; 1978; CEB-FIP (1993); Neville et al (1983).

Under some conditions where the published curve is not valid or available, the value

of the aging coefficient can be determined using the method proposed by Bazant and

Kim (1978) in which the relaxation function was rewritten as a function of the creep

function st, , and takes the following form:

1,

,

1,

115.0

,

1, 0

tt

ss

ttststR Equation 2.89

where: 0 is a correction factor, taken as 0.85; is equal to 0.5(t-s).

The formulation of the AEMM method involves the following conditions: shrinkage

strains are treated independently and additively from creep; elastic and creep

components are combined into a single function of initial stress, σ0, that applied to the

71

member and the creep coefficient; restraining effects due to the presence of the

reinforcement and effects of stress increments applied over time are accounted for in a

function containing the aging coefficient, χ (t,s). Total strain can be written as:

),(),(1)(

)(),(),(1

)(, 0

0,

0 ststsE

tttst

sEst

c

shsh

c

Equation 2.90

where: ε(t,s) is total strain at time t of a reinforced concrete member loaded first time

at time s; χ(t,s) is the aging coefficient; σ0 = initial stress applied to concrete member;

σ(t) = concrete stress at time t.

The AEMM method is more theoretically exact than the EM method. Despite its

mathematical simplicity, the accuracy of the AEMM method has been found more

superior to both the EM and the RC methods (Bazant, 1972). This reason makes the

AEMM method widely used in the literature.

2.5.5 Summary

From the reviews of the four methods, it can be seen that there have been attemps to

improve the accuracy of axial shortening prediction by deriving prediction models

with more realistic assumptions. The first model available for calculation of axial

shortening was EM method in which its only breakthrough innovation was an

introduction of a term namely effective modulus.

The EM method was crude as the applied stress in vertical members was assumed to

be constant, the most important known flaw of this method that leads to a low

precision and an opportunity for other researchers to rectify this problem. RC and

Superposition method provided improvement over the EM method as they take into

acount the effects of additional incremental loading by means of a differential

equation, obviously a more realistic approach to reproduce stress in vertical members

as they actually occur in most structures.

72

Finally, AEMM method was proposed as a definitive improvement over the long time

obsoleted EM method. The major improvements include the use of an aging

coefficient which can be determined by means of a newly proposed relaxation

function in order to account for the effects of aging, making the AEMM method the

most sophisticated and theoritical exact axial shortening model available. The AEMM

method is therefore widely used and frequently quoted in literature. It is justified that

the AEMM method be employed in this study for axial shortening analysis.

2.6 Experimental and analytical studies on axial shortening

Although the creep and shrinkage effects on concrete structures have been researched

since the early 1900s, it was not until about sixty years later when field measurements

on reinforced concrete buildings dedicated to the investigation of axial shortening in

columns and core walls were first undertaken. Since then, structural monitoring is

more and more accepted as a suitable approach to study the behaviour of structures

particularly when the concerned topic may substantially influence the serviceability

and safety of the structure.

In general, early works were mainly conducted in the USA during 1960s and 1980s.

However, it was found that more publications from elsewhere including Australia

became available after 1980s. The two most recent column shortening monitoring

projects were conducted by Australian researchers. They were the experiments at

World Tower in Sydney and Q1 Tower in Gold Coast as presented in this thesis.

Generally, the total or resultant axial shortening which is the combination of

instantaneous or elastic shortening, creep, shrinkage, temperature and humidity effects

as well as methods used for separating the different of each effect are reported.

Details of researches completed in the past vary in many aspects such as duration of

the data collection, instrumented details and implemented prediction methods as well

as construction schedule and rate of loading. Some of these factors have not been in

the control of researchers and all they could do was to interpret their influence after

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the field data has been obtained. The employed monitoring approaches were also

found varying vastly. They depended on factors such as the type of the structure, its

dimensions and type of construction materials, duration of data collection, and the

amount of research budget.

The following sections of this chapter give a review of literature regarding field

measurements of axial shortening on columns and core walls in buildings from

different countries.

2.6.1 Axial Shortening measurements in Australia

The UTS City Campus Building - Sydney

Bakoss et al (1984) conducted a comprehensive study on the long-term deformations

in a 120 m tall building. Five years results of columns and core walls deformation

were reported. The building which is now accommodating the University of

Technology Sydney was monitored for long-term strains and deformations in its

columns and cores. The instrumentation comprised a complete monitoring system

which includes many gauges of vary types including embedded and surface strain and

temperature transducers, survey reference points, wind pressure tapping and

anemometer above the roof of the building to measure wind characteristics.

In Bakoss et al (1984), results from one column were presented together with

predicted values computed with the methods presented in Fintel and Khan (1969) and

Warner (1975). Reasonably good agreement was observed when using both methods.

However, it was noted that accuracy of the two methods was indifferent and mainly

influenced by the reliability of creep and shrinkage data that used in the axial

shortening predictions. It was stated that heavy reinforcement and the delay of the

first significant load application resulted in the relatively low elastic strains. Effects of

seasonal ambient temperature changes was investigated and found insignificant. The

differential movements between columns and core walls were found inconsequential.

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World Tower (Sydney, Australia)

The most recent Australian research on collecting axial shortening data from a real

structure found amongst the published literatures was conducted at the World Tower

Building in Sydney, Australia by researchers of the University of New South Wales

(Bursle et al, 2003). The World Tower Building is a 78-storeyed building located in

the CBD area of Sydney. The columns and core walls of this building were made

from concrete with design strengths in a range of 32 to 90 MPa. The axial shortening

of three selected columns has been measured using DEMEC gauge techniques. In

addition, internal strain gauges were installed to measure the stress in the vertical

reinforcement.

A typical plot of the measured column strains taken on level 23 of the building has

been reported (see Figure 2.4) as well as examination of some material properties

adopted in the shortening prediction, i.e. characteristic strength of concrete; the use of

density in the prediction of elastic modulus; as well as creep and shrinkage. It appears

that column deformation increases with the increasing applied load.

Figure 2.4. Measured column strains, level 23, World Tower Building (Bursle et al,

2003).

75

Bursle et al (2003)’s finding includes: compressive strength at 28 days of delivered

concretes is usually higher than their designated strength by up to 20%. Therefore

measured strength should be used in the axial shortening analysis to avoid errors; the

elastic modulus equations are not sensitive to the variation of concrete density

therefore a standard value of density should be used in absence of actual concrete

density data. Bursle et al (2003) concludes that work on the influence of material

properties is required in order to determine the axial shortening more accurately, also

existing code models needed to be critically reviewed.

2.6.2 Axial Shortening measurements in United States of America

A 38-storeyed Building – Chicago

Fintel and Khan (1971) conducted field measurements of strains in columns and walls

of a 38 storeys building. The employed measurement apparatus was a Demountable

Strain Gauge (DEMEC) with 10 inch (254 mm) gauge length. Note that details on

DEMEC will be given later in Chapter 3. The DEMEC measuring points were half

inch in diameter brass knob with hole drilled in their centre. At the second floor of

the building which serves as a parking garage, nine DEMEC measuring points were

installed on two isolated columns and two shear walls several days after the formwork

removal. The concrete grade of the instrumented elements was 5000 psi (34.5 MPa).

The measurements were then conducted over 79 months; nine months and three

weeks were during the construction. Movements due to ambient temperature and

relative humidity were neglected due to the lack of information thus no corrections

were made. An established procedure to analyze differential shortening of multi

storey buildings when an equivalent one bay frame is facilitated was also presented

together herein the paper. The method was verified with the measured results as the

measured creep and shrinkage strain of columns and core walls compare acceptably

with the computed values.

76

According to the results published by Fintel and Khan (1971), it has been concluded

that; differential shortening exists between the columns and core walls, this should be

considered structurally and accommodated in the design where the presented

simplified analysis may be used satisfactorily. It was noted that resulting frame

moment may be relaxed about 50% over a prolonged period of time due to the creep

of the slabs. It was observed that size and shape in terms of volume-to–surface ratio

has lesser effects on differential shortening as compared to the stress and amount of

reinforcement. Therefore a column and its adjacent walls should be designed to have

the same amount of stress to avoid problem due to differential shortening.

Lake Point Tower – Chicago

In the late 1960s, the Portland Cement Association initiated an extensive experimental

program at Lake Point Tower, a 70 storeyed reinforced concrete structure in Chicago,

Illinois, where it would be monitored for time dependent shortening of concrete

columns and core walls. Four concrete grades with compressive strength of 3500 psi

(24 MPa), 5000 psi (35 MPa), 6000 (41 Psi) and 7500 psi (51 MPa), depicted as type

A, B, C and D respectively, were used in the making of vertical structural members

which diameters and thicknesses vary over the total height. An image of the Lake

Point Tower is given in Figure 2.5.

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Figure 2.5. Lake Point Tower

(http://www.aviewoncities.com/chicago/lakepointtower.htm)

To obtain substantial amount of data, columns and core walls in twenty-seven

different storeys were monitored for axial strains during and after construction. After

casting, Selected columns and core walls were attached with reference discs using

epoxies, allowing length change to be determined using a mechanical strain gauge

with a gauge length of 20 inch. The instrumented levels and measurement locations

are presented in Figure 2.6.

http://www.aviewoncities.com/chicago/lakepointtower.htm

78

Figure 2.6. Lake Point Tower- levels of instrumentation and measurement locations.

(Pfeifer et al, 1971)

At selected storeys where depicted with asterisks as seen on the left side of the front

view of the tower, three columns and two locations on core walls were instrumented

as shown in a floor plan reproduced in Figure 2.6, with three gauge lines spaced at

about 90 degree intervals of column perimeter on each location. For the rest of the

instrumented storeys, only column 1C and core wall 1W were instrumented with a

single gauge line for each location. In addition, a humidity probe was installed into

column 1C at selected storeys to measure relative humidity within the column.

Interim results that contained measured axial shortening of columns at two levels

were reported by Magura et al (1968). Measured axial shortening of the instrumented

columns and core walls for the selected 27 storeys were fully reported later in Pfeifer

et al (1971). These measured results are presented herein Figure 2.7 and Figure 2.8.

Note that all strains have been adjusted to a reference temperature of 73°F using the

measured thermal coefficient.

79

Figure 2.7. Lake Point Tower- measured axial shortening of columns. (Pfeifer et al,

1971)

80

Figure 2.8. Lake Point Tower- measured axial shortening of core walls in millionths

unit (aka microstrain), (Pfeifer et al, 1971).

It can be seen from Figure 2.7 and Figure 2.8 that the premature termination of the

strain measurement can be found in many storeys. This was due to the loss of points

or access to the locations. In general, total deformations of core walls was less than

that of those investigated in columns and axial shortening in upper floors are to some

extent greater than those of the lower storeys. The first was due to the relatively light

load on the core walls, and the latter due to influence of different concrete strengths

used. The deformations are almost independent of story.

Properties of the field concrete were determined by the laboratory testing conducted

simultaneously. The tested properties include: compressive strength, elastic modulus,

creep, shrinkage and the coefficient of thermal expansion. Results obtained were to be

used in the axial shortening analysis.

A data analysis conducted to estimate the magnitude of axial shortening that would

occur theoretically, was also outlined in Pfeifer et al (1971). When neglecting

temperature effects, the total axial shortening was assumed to be consisting of three

components: instantaneous, creep and shrinkage shortening. The instantaneous

81

shortening was calculated using elastic modulus of concrete corresponding to age of

concrete as in Hanson and Mattock (1966) and Pfeifer (1970). A method suggested by

Dischinger (1937 and 1939) was favored for predictions of creep and shrinkage

shortening. For better accuracy, it has been modified for the inclusion of

reinforcement effects. The modified equation proposed in Pfeifer et al (1971) takes

the following form:

caCE

gs

epE

C

Sf

10

Equation 2.91

In Pfeifer et al (1971), the tested creep and shrinkage data were used in Equation 2.91

to yield predictions of axial shortening. However, to account for the effect of volume

to surface area effects, Pfeifer et al (1971) recommended using correction factors,

based on the work of Hanson and Mattock (1966), to modified values of creep and

shrinkage from the laboratory test in order to make them suitable with much larger

actual member size. A chart developed by Pfeifer et al (1971) for this purpose is

shown in Figure 2.9.

82

Figure 2.9. Member size correction charts for creep and shrinkage of concrete.

(Pfeifer et al, 1971)

Corrected values from Figure 2.9 were then applied into the Equation 2.91. Assuming

a series of loading increments was applied at various times, with the employment of

the superposition theory (McHenry, 1943), prediction of creep shortening was made.

In a much simpler approach, shrinkage prediction was made for different age of

concrete but in only one increment. The three components were then added together

to give total axial shortening.

Comparisons of measured and predicted column shortening at 1st and 30

th storeys of

Lake Point Tower are shown in Figure 2.10. Similar comparisons for core walls at

10th

and 30th

storeys are presented in Figure 2.11.

Figure 2.10. Lake Point Tower- measured and predicted axial shortening of column at

1st and 30

th storeys. (Pfeifer et al, 1971)

83

Figure 2.11. Lake Point Tower- measured and predicted axial shortening of core walls

at 10th

and 30th

storeys. (Pfeifer et al, 1971)

It can be seen from Figure 2.10 and Figure 2.11 that the prediction gives moderately

good agreement with the measured values during the construction period. However,

major error was originated over long term prediction. This finding indicated that

analysis assumptions and tested concrete properties used may not be entirely valid.

For columns both measured and predicted shortenings were fairly similar for both 1st

and 30th

storeys. However, for lower storeys elastic shortening was the main

contributor to the total shortening and vice versa for upper storeys. This was due to

the larger column sizes, more reinforcement and stronger concrete in column at lower

storeys that resulted in a much lower creep and shrinkage coefficient. Also lower

columns carried their maximum load at a much later age as compared to those of

upper columns.

For core walls, a similar analysis was made. However, predicted and measured axial

shortening as shown in Figure 2.11 did not match very well like in columns. It was for

84

the reason that the load was assumed to be distributed uniformly although the strains

were only measured at one location.

Plots of differential shortening of each storey are shown in Figure 2.12. The analysis

was made for the age of 571 days (end of construction) and 2000 days after the first

floor was cast. The analysis for differential shortening at 2000 days was achieved by

extrapolating the calculated strain-time curves for selected storeys.

Figure 2.12. Lake Point Tower- differential shoretning between instrumented columns

and core walls. (Pfeifer et al, 1971)

Differential shortening was found increasing linearly up to about the 37th

storey where

it remains relatively constant. This means that the result of each storey affects total

differential shortening equally. It was also found that differential shortening becomes

steady once the age of 2000 day is reached.

Water Tower Place – Chicago

Water Tower Place is a 76-storeyed reinforced concrete building with a height of 262

metres. Started during the construction, long-term deformations in selected columns,

85

core walls and caissons were monitored for over 13 years. Results and findings of this

research were published in Russell and Corley (1978) for the initial three-year result

and Russell and Larson (1989) for the thirteen-years end result. A photograph of

Water Tower Place is shown in Figure 2.13.

Figure 2.13. Water Tower Place (http://www.aviewoncities.com)

The structure consists of a 4-storeyed basement, a 13-storeyed lower part measured

65.2x161.8 m in plan, and an extending 63-storeyed tower that is located at one

corner of the base structure and measured 28.7x67.4 m in plan. Structural design for

lower part was conventional with reinforced concrete frame acting in combination

with two massive lift shafts in resisting lateral forces, whilst a concept combining

tubular design with perimeter columns and the transverse walls forming two equal

tubes that act together in resisting horizontal forces, was used for the tower structure.

(Russell and Corley, 1978)

Axial shortening measurement was conducted at one selected level in the basement

and five selected levels in the tower where member size, concrete strength and

percentage of reinforcement were changed over the height. The instrumented levels

and their specified strengths were level B4 (62 MPa), L14 (62 MPa), L32 (52 MPa),

http://www.aviewoncities.com/

86

L44 (41 MPa), L57 (34.5 MPa) and L73 (28 MPa) respectively. The framings and

location of instrumentations are shown in the following Figure 2.14.

Figure 2.14. Water Tower Place- measurement locations. (Russell and Larson, 1989)

Following the method used at the Lake Point Tower, a similar mechanical strain

gauge with 20 inch gauge length was once again used at the Water Tower Place. A

minimum of three sets of gauge were installed on circular columns and four on

rectangular columns. Air temperatures were also measured. All gauge readings were

adjusted for changes in temperature of concrete with the reference temperature of

73°F, the same as Lake Point Tower.

Measured axial shortening of columns and walls during the first 5 years as reported in

Russell and Larson (1989) is given in Figure 2.15. Note that results shown in one

curve are average values of all gauge readings at one location.

87

Figure 2.15. Water Tower Place- measured axial shortening of columns and core

walls during the first 5 years. (Russell and Larson, 1989)

As shown in Figure 2.15, findings were in good agreement with axial shortening data

from the Lake Point Tower (Pfeifer et al, 1971). Firstly, it can be seen that

measurements at level 15, 44 and 57 were terminated hastily due to the loss of points,

similar problem as experienced in the research at the Lake Point Tower. Secondly,

regardless of level 73, magnitudes of observed axial shortening were roughly in the

same range at all levels. Finally, axial shortening in core walls was less as compared

to those measured at columns.

88

In addition to the initial 5-year measurements, measurements were continued at the

fourth basement and in selected caissons. Axial shortening in caissons was measured

using strain meters cast into the structures. Long-term results from columns at the

fourth basement and instrumented caissons are shown in Figure 2.16. It can be seen

from Figure 2.16 that axial shortening still continues over the long period.

Figure 2.16. Water Tower Place- caisson shortening and measured axial shortening at

the fourth basement. (Russell and Larson, 1989)

In this present study, initial data analysis was made using a microcomputer program

in which the method that had been used previously in the research at Lake Point

Tower (Pfeifer et al, 1971) was computerized to facilitate analysis at each storey.

In brief, instantaneous and creep shortening was calculated in several increments

depending on the age of concrete and the storey under consideration. The total

instantaneous strain was the summation of increments up to the target age. These

instantaneous increments were calculated using the elastic modulus from the

laboratory test.

89

Creep increments were computed based on the creep test data. The curve

recommended by Pfeifer et al, (1971) that has been presented previously in Figure 2.9

was used to accommodate for effects of size. The effect of the presence of

reinforcement was taken care of using Dischinger’s equation (Dischinger 1937 and

1939; Pfeifer et al, 1971). Creep increments were superimposed for the total creep

shortening. Calculation was made the same way for shrinkage deformation, however

only one increment was used.

Comparisons between measured and predicted axial shortening for selected column

D2 is given in Figure 2.17.

Figure 2.17. Water Tower Place- measured and predicted axial shortening of column

D2. (Russell and Larson, 1989)

90

The observations were again in good agreement with the previous study (Pfeifer et al,

1971). Accurate predictions were observed during the construction period, however,

under predictions observed thereafter.

The ratio of instantaneous, creep and shrinkage shortening to the total shortening was

depended on levels of the building and progress of the construction. At basement

levels, instantaneous strains were the largest proportions of shortening whilst at upper

levels creep and shrinkage were the largest proportion due to the relatively light load.

It was found that the effect of instantaneous shortening was minimal at level 73.

During construction, particularly at early ages, instantaneous strain was almost the

only factors that affected axial shortening as shrinkage and creep still had not quite

developed yet. After construction, instantaneous shortening did not vary whilst creep

and shrinkage increased constantly.

2.6.3 Axial Shortening measurements in the United Kingdom

A Six-Storeyed and An Eighty-Storeyed Building

Arumugasaamy and Swamy (1989) studied the long–term behavior of concrete

columns in two structures during construction, occupancy and service. Strain results

obtained from the buildings during a period of ten years were reported. In the first

structure, a six-storeyed reinforced concrete building depicted as structure A, strain

measurements were conducted at seven columns in the basement. The concrete

strength was 31 MPa. In the second structure, an eighty-storeyed building depicted as

structure B, deformations in an internal column (575x575 mm) and its external

adjacent column (435x355 mm) were measured. The concrete strength of the two

columns was 42 MPa.

91

The instrumentation as reported in Arumugasaamy and Swamy (1989) was extensive.

A set of different gauges was used to collect data including surface strain, internal

strain, stress and temperature. The Briquetted Acoustic Gauges were used to measure

internal concrete strain of the columns. The strains in longitudinal reinforcement of

the columns of structure A were measured using a Demec gauge, while in structure B

surface mountable acoustic gauges were used. The temperatures in structure A and

structure B were measured by an acoustic gauge and copper-constantin thermocouples

respectively. Lastly moisture content in concrete was measured by using humidity

probes and stress of concrete in the columns of structure B was measured using a

Carlson stress meter.

The findings, according to Arumugasaamy and Swamy (1989) were; all the columns

showed performance in response to actual variations occurred in term of loading, and

the load transfer from concrete to steel reinforcement was noticed particularly in

columns with a higher percentage of steel.

2.6.4 Axial Shortening measurements in other countries

The Library Building of Concordia University – Canada

The long-term deformations of reinforced concrete columns in the Library Building of

Concordia University in Montreal, Canada have been reported by Miao et al (1993

and 1994) and Miao and Aitcin (1995).

In this study, two structural columns and two control columns, which were never

loaded, were instrumented and monitored for strain development. Concrete grade of

columns were 80 and 100 MPa. A fully automated data acquisition system has been

employed. Eight IRAD vibrating wire extensometers were installed in each column

for the strain reading of different internal location, i.e. four extensometers were placed

vertically at different distances from the center of the mid-height section; one

extensometer was placed horizontally in the centre of the column; and three

extensometers were welded to three vertical reinforcement bars. In addition, a

92

thermocouple was added to the acquisitor to measure the ambient environmental

temperature.

The temperature development and corresponding column strains due to the hydration

process during the first days after casting have been published in Miao et al (1993). In

general, the concrete temperature rises up and reaches its peak at approximately 20-

hours after the casting. The maximum temperature has been measured at centre of the

column and the resulting expansion was about 450x10-6

mm/mm. Afterwards,

temperature and expansion strains decreased with time. The expansion strain

stabilized at about 150x10-6

mm/mm after 3.5 days. The variation of the coefficient

of thermal expansion of concrete was also noticed to be relatively high during the first

seven-hours and decreased progressively to stabilize at around 10x10-6

per one degree

Celsius after setting.

The one-year results have been subsequently reported by Miao et al (1994). The one-

year investigation reveals that the temperatures in different locations of the columns

are indifferent under environmental conditions; therefore, the mechanical effects of

thermal gradient can be discarded. Column strains and stresses in the reinforcing steel

have been found to be influenced by the temperature fluctuation. It has been observed

that columns containing concrete with relatively higher compressive strength

(100MPa) have greater strains than 80 MPa columns due to higher elastic modulus

and lower creep potential. Stresses in the reinforcement of the two loaded columns are

approximately similar with a value of about 110MPa at one year.

Drying shrinkage of the control columns has been reported in Miao et al (1994). It has

been found that: drying shrinkage of the 100 MPa concrete which contains silica fume

is greater than that of 80 MPa concrete; concrete with longer curing time has less

drying shrinkage (this phenomenon is more pronounced for 80 MPa concrete); and the

strains measured from columns were different from those of smaller lab samples due

to the differences in terms of volume to surface ratio.

Finally, five-year results for the monitoring of HSC columns on the new library of

Concordia University (Montreal, Canada) have been reported by Miao and Aitcin

93

(1995). The results on strain development and stress in concrete and reinforcement

bars are also presented. They stated that the temperature variation has greatest

influence on the strain development as compared to other influencing factors. Miao

commented that lower strength concrete had higher load induced strain because of its

lower modulus of elasticity and its higher creep coefficient. The specified structure

design loads were compared to measure values and found to be overestimating the

latter.

A 19-storeyed Building - Singapore

Glistic et al (2003) performed long-term structural monitoring of a 19-storeyed

residential building in Singapore. The monitoring of 10 ground level columns was to

be performed during the whole life-span of the building including the construction

and in-use periods. In this paper, the monitoring methodology of the embedded fiber

optic strain gauge system (SOFO) employed was outlined. Results from

measurements during the construction of the 19-storeys were reported. The plot of

total strain in columns as reported in Glistic et al (2003) is reproduced in Figure 2.18.

It can be seen that all columns developed similar level of deformation except Column

9 which had relatively lighter load.

Figure 2.18 Measured column strains of the 19-storeyed building by SOFO system.

94

The irregularities of the measured axial shortening has been noted by Glistic et al

(2003) that sometimes column strain is constant even if there is an increase of load

due to the additional floors, also high jump is sometimes noticed. These irregularities

may be due to some unknown applications of temporary loads on the floors or strain

initiated by temperature variations.

For analysis purposes, the CEB-FIP model code 1990 has been used to calculate

shrinkage and creep component of the monitored columns; thermal strain has been

neglected; and elastic strain has been taken as the residual strain after deducting creep

and shrinkage strains from the measured column strain. It has been found that

measured and predicted elastic strains are much different particularly at initial

periods, afterward the difference decreases as the construction progresses.

Glistic et al (2003) concluded that the performance of the SOFO monitoring system is

excellent even though only one gauge has been installed in each column. The

accuracy of the CEB-FIP model code 1990 has been found to be approximately within

a range of ±25% and this may be improved if data on temperature variation is

available and incorporated into the analysis.

2.6.5 Summary

Works on axial shortening measuring in buildings were comprehensively reviewed. It

was found that axial shortening data were obtained from buildings by means of strain

gauges devices either being mechanical or electrical powered. By far, the embedded

fiber optic strain gauge system (SOFO) reported in Glistic et al (2003) appears to be

the most up-to-date and technically advanced system that was employed in this

particular kind of experiment whereas the DEMEC gauge is the most simple and

oldest. Despite its simplicity, comparison between data obtained using electronic

strain gauge and DEMEC gauge shows that no difference in term of accuracy was

evidenced. However, it appeared that DEMEC gauge was far more frequently used to

collect axial shortening data from buildings perhaps due to its ease of use and much

lower cost.

95

In Australia, a recent work (Bursle et al, 2003) was conducted on the World Tower

building in Sydney where DEMEC gauge technique was employed. Data reported in

Bursle et al (2003) displayed a similar trend of axial shortening as compared to data

of Glistic et al (2003). Similar irregularities of axial shortening data were also

reported in both reports despite their different choice of instrument used.

Consequently, it is justified that DEMEC gauge technique as adopted by Bursle et al

(2003) be replicated for this research in order to obtain axial shortening of the Q1

Tower columns and core walls.

2.7 Summary of the chapter

In this chapter, based on the comprehensive reviews of the related literatures, a

complete picture about the development and what had been done in the area of axial

shortening studied both analytically and experimentally are consequently clearly

established. Firstly, aspects of HSC/HPC are outlined. It was found that the

increasing compressive strength of concrete used in modern high-rise buildings has

consequently reflected the need to address axial shortening issues as columns and core

walls in high-rise buildings are designed to take high service stress. There were

numerous reports on the use of HSC/HPC in buildings where discrepancies of

HSC/HPC requirements were found. For this study, concrete grade 50 MPa and

higher are classified as HSC.

Axial shortening is considered as a result of instantaneous and creep strain occuring

due to an applied stress plus shrinkage strain occuring due to drying of cement paste

minus restraint due to presence of rebars. In order to assess axial shortening, time

dependent properties of concrete, i.e. stregth, elastic modulus, creep and shrinkage

need to be known. Test data of these properties are plentiful and are not directly

important for this study. Therefore, only a selection of major material models were

discussed. In Australia, AS3600 code model is well established and has been regularly

updated. The most recent update version of the AS3600 code, AS3600-2009 which

adopted creep and shrinkage model from Gilbert (2002 and 2005) and the superceded

code AS3600-2001 are both reviewed and investigated.

96

Contrary to the abundant concrete test data, researches on axial shortening model are

extremely scarce as only 4 methods were identified, all from overseas. The most up-

to-date axial shortening model was AEMM method (Bazant, 1972), since then a more

up-to-date model which is also well-accepted could not be found. The AEMM method

is also widely acknowledged to be the most realistic model available. Therefore it is

adopted for further axial shortening analysis in this study.

Another area focused on this chapter is researches emphasized on experimental study

of axial shortening on real buildings. There were a small number of published

researches in this area over the last 30 years. However, it is observed that attempts

have been made repeatedly in order to obtain in-situ data and to understand more

about axial shortening. Several instrumentation methods were employed and none

was clearly proved to be the excellent in terms of accuracy but DEMEC gauge was

found to be the most versatile thus it is employed for this study. In Australia, a few

works in this area were identified and the latest Australian research was the

experiment conducted on the World Tower in Sydney (Bursle et al, 2003).

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100


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110

3 Methodology

3.1 Introduction

An experimental program includes axial shortening data collection on the construction

site of the Q1 Tower and the laboratory concrete testing was implemented in order to

enable detailed study of real-time axial shortening that occurred under real-life

conditions.

For in-situ data collections, DEMEC gauge techniques as adopted earlier for the

World Tower research (Bursle et al, 2003) were implemented. This method was found

overall suitable for building site condition. The lab testing program was conducted to

verify major concrete properties that required for further axial shortening analysis.

Tested properties were compressive strength, elastic modulus, shrinkage and creep.

A detailed description of the Q1 Tower as well as required instrumentations and

standard laboratory testing involved in this research is given hereafter.

3.2 In situ instrumentation and data collection at the Q1 Tower

3.2.1 The Q1 Tower

3.2.1.1 General information

Located on a site encircled by Gold Coast Highway, Clifford Street, Hamilton Avenue

and Northcliffs Terrace at Surfers Paradise, Q1 Tower is the world’s tallest residential

tower on Surfers Paradise, Australia (Whaley, 2002). The site occupies almost an

entire CBD block with an area of 11,500 m2. The tower which faces the world’s

famous Surfers Paradise beach has been built at the northwest corner of the site. A

perspective of the Gold Coast’s skyline with the Q1 Tower is shown in Figure 3.1.

111

The structure includes an eighty-storey residential tower of approximately five

hundred apartments, two basement levels of parking with a capacity of eight hundred

cars; a podium level consisting of restaurants and retail outlets as well as extensive

landscaping and sports facilities.

Figure 3.1. The Architect’s perspective of the Q1 Tower (courtesy of Sunland Group,

2002)

A typical floor of the tower consists of approximately 1500m2 and decreases at

regular intervals down to approximately 1200m2 at level 74. The tower consists of

nine apartments per floor at the lower level and reduces to four apartments per floor at

the higher levels. There are six plant rooms, of which four plant rooms are located at

levels 38 up to 41 and occupy half the area of each floor. The other two plant rooms

take up the entire floors at levels 75 and 76. Other uninhabitable areas include a roof

garden at the rear of the building between levels 60 and 69 and two observation decks

at levels 77 and 78 (Whaley, 2001).

The building is served by nine elevators: two express lifts up to the observation decks;

four lifts to the low-rise zone up to level 41; and three lifts to the high-rise zone from

levels 42 to 74. The building is encapsulated in a glass curtain wall; although balcony

112

windows are operable to ensure year-round comfort. The height of the tower’s

structure from ground level to level 78 is approximately 275 metres. The single mast

extended from the crown at the top level further boosts the total height to 325 metres

(Whaley, 2001).

3.2.1.2 Dynamic response

According to Whaley (2002), the dynamic response under service conditions is

critical to the design of the Q1 Tower which is a tall and slender structure as it may

affect the comfort of residents of the upper floor. It was decided to limit the dynamic

response of the tower to less than 12 milli-g for a one year return period, i.e. the

lateral movement was limited to a height/600 under serviceability wind loading,

equivalent to a movement of 400mm at level 78. For a return period of ten years, the

allowable acceleration is considered equal to 15.8 milli-g.

3.2.1.3 Structural system

The Q1 was designed based on code recommendations of the Standard Association of

Australia. The codes referred to in this project are listed as follow:

AS3600-2001: Concrete Structures

AS1170.1-1989: SAA Loading Code - Permanent, imposed and other actions.

AS1170.2-1989: SAA Loading Code – Wind actions.

AS1170.4-1993: SAA Loading Code – Earthquake loads.

AS1720.1-1997: SAA Timber Structures Code – Design methods.

AS1538-1988: SAA Cold-formed Steel Structure Code.

AS3700-2001: Masonry Structures.

AS4100-1998: Steel Structures.

113

Table 3.1. Material Specifications for the Q1 Tower applied to this research were as

follows.

Grade f'c (MPa) 25 32 40 50 65

Young Modulus E (MPa) 27x103

31x103

34x103

38x103

43x103

Coefficient of thermal expansion

Shrinkage strain

Basic Creep Factor 4.2 3.4 2.5 2 2

Poisson's Ratio

Density

Plain Bars

Deformed Bars

Steel Wire, Plain and Deformed

Welded Wire Fabric

Modulus of Elasticity

fsy = 450 MPa

fsy = 450 MPa

200x103MPa

24kN/m3

Reinforcement

fsy = 250 MPa

fsy = 500 MPa

Concrete

10x10-6

700x10-6

at 8 weeks

0.2

Podium and basement structure

According to Whaley (2002), the podium structure comprises of reinforced concrete

frames including a slab-on-ground subjected to hydrostatic uplifts due to variations in

ground water level, and two levels of reinforced concrete suspended floors. The floor

system of the podium has been designed to carry a weight of up to 750 mm of

landscaping and 1500 mm of swimming pool water. The podium consists of a 225

mm thick conventional reinforced concrete flat slab with 400mm drop panels. A

perspective of various facilities at the podium level is shown in Figure 3.2.

114

Figure 3.2. Illustration of facilities at podium floor (courtesy of Sunland Group, 2002)

The column grids of the podium and the basement were aligned to suit their function

as a car park (see

Figure 3.3). Retaining walls at the basement levels are cast in-situ in order to ensure

adequate strength and enhance water retention whereas retaining walls at the upper

levels are block-work. The podium and the main tower are separated both vertically

and horizontally by transverse joints and placed-pour joints. Lateral stability of the

podium structure is mainly dependent on the perimeter basement walls.

Figure 3.3. Plan of podium floor (Sunland Group, 2003)

115

Tower structure

The foundation of the tower consists of large diameter piles founded in weathered

rock. This ensures minimal differential settlement due to variations of soil properties

at the site. The bored piles and pile caps layout is shown in

Figure 3.4.

Figure 3.4. Bored piles and pile caps layout of the Q1 Tower (Whaley, 2002)

The tower structure consists of cast in-situ floors supported by cast in-situ columns

and walls (Whaley, 2002). The typical horizontal structure contains cast in-situ flat

slabs with varying thickness of 200 to 225mm and perimeter band beams. Floor spans

are generally 9 to 10 meters and the slabs are partially pre-stressed where slab

116

deflection exceeds the stipulated values. Grade 40 MPa concrete was used in all cast

in-situ floor slabs over the height of the entire tower. The designs of floor slabs were

optimized to provide a less complex reinforcement layout within acceptable deflection

limits. Minimum required reinforcement is N16@300mm and compression

reinforcements were incorporated in some cantilevers to reduce deflections.

The vertical structure consists of outrigger walls and columns designed and positioned

to suit lateral stability requirements of the tower by connecting the central lift cores to

the massive perimeter columns via concrete walls (a shear core-outrigger-perimeter

frame system). The core walls are connected across the floor by link beams of lintel

along corridors. There are approximately fifteen link beams per each typical floor.

Note that majority of the outrigger walls terminate at higher levels to allow for

relatively larger apartments while the lateral stability of the tower relies mainly on its

central core. A cross-sectional view of the structure elements is given in Figure 3.5.

The columns and core walls of the main tower have been specifically arranged to suit

lateral stability, architect’s layout, and typical floor slab soffit requirements. The

structural floor plan of the tower’s basement and typical floors are shown in Figure

3.6 and Figure 3.7, respectively.

117

Figure 3.5. An isometric view of the Q1 Tower’s structural elements (Whaley, 2002)

Figure 3.6. Basement floor plan of the Q1 Tower (based on Whaley, 2001)

118

Figure 3.7. A typical floor plan of the Q1 Tower (based on Whaley, 2001).

Columns and core walls in the tower have been optimized under vertical loads to

achieve “balance compressive stress” (Whaley, 2002). This is to induce an equal

magnitude of stress into all vertical members within any given floor by varying its

cross-sectional areas. Therefore, an equal magnitude of axial shortening is assumed in

all vertical members.

It should be noted that the perimeter columns have been designed to take most of the

additional loads arising from wind effects. Thus, they are designed to resist either

compression under normal circumstances or tension under wind loading. On the other

hand, the shear core and outrigger walls bear only compression. Also worthy of note

is that differential shortening between adjacent columns under serviceability wind

loads was limited to L/500 (Whaley, 2001).

Most importantly for this research are the design details of columns and core walls

such as dimensions, grade of concrete, and amount of reinforcement varying at certain

119

levels throughout the height of the tower. The dimensions reduce steadily over its

height to a minimum thickness of approximately 200 mm and 300mm for walls and

columns, respectively. Similarly, the grade of concrete decreases from 65 MPa at the

basement level to 40 MPa at the top. The percentage of reinforcement in the columns

ranges from 1% to 4%. Full instrumented details of columns and core walls in the

tower are enclosed in Appendix A.

3.2.2 Data collection at the Q1 Tower

The main objective of the data collection is to extensively collect actual

measurements of time dependent axial shortening developed within the columns and

core walls of the Q1 Tower.

3.2.2.1 Apparatus and equipment

The data collection utilised a simple mechanical device that required a team of people

to manually undertake measurements. This is a similar process used by Bursle et al

(2003).

The Demountable Mechanical Strain Gauge (DEMEC)

The demountable mechanical strain gauge or DEMEC gauge is a precision instrument

originally developed for strain measurement and crack monitoring on any concrete

structure.

Mayes Instrument Limited (2001), (http://www.mayes.co.uk/demec.htm), the

manufacturer of DEMEC gauges used in this study, describes a typical DEMEC

gauge as “A device consisting of an invar main beam with two conical locating points

one fixed and the other pivoting on a special knife edge. The points are located in pre-

120

drilled stainless steel discs, which are attached to the structure with adhesive.

Movement of the pivoting point is measured by the strain gauge which is attached to a

base plate on the invar beam”.

Figure 3.8. DEMEC gauges (Mayes, 2001)

The DEMEC gauge comes with an Invar reference bar to control measurements and a

setting out bar to enable accurate positioning of the locating discs as shown in Figure

3.8 (Mayes, 2001). The name “Invar” refers to the component of the main beam and

reference bar of DEMEC which is a special alloy of Iron (64%) and Nickel (36%)

(http://en.wikipedia.org/wiki/Invar). Invar has a very small coefficient of thermal

expansion i.e. 1.22 microstrain per 1 degree celcius. (The use of invar facilitates and

provides a low coefficient of thermal expansion to the DEMEC gauge at room

temperatures). This allows users to achieve maximum accuracy of measurements at

all time with minimum interference from ambient environmental temperature.

The DEMEC gauge is fully portable, thus it is ideal for strain measurements at several

locations within a structure. DEMEC gauges are equiped with a precision mechanical

dial gauge of 1/500 mm gradation. The change in length measured in microstrains

(mm/mm) for any DEMEC gauge can be calculated by dividing 1/500mm by the

gauge length. Therefore, when a similar mechanical dial gauge is affixed on any

DEMEC gauge, the smallest strains it can measure depend solely on the gauge length.

As such, a DEMEC gauge with a greater gauge length would be capable of detecting

relatively finer strains when compared to a shorter one. The most common gauge

lengths of DEMEC gauges available from the manufacturer are listed with a

corresponding measurable unit (microstrains) in Table 3.2.

121

Table 3.2. Microstrains detected by one division on the dial gauge (Mayes, 2001)

For the Q1 Tower project, two sets of DEMEC gauges were utilised for the strain

monitoring process; one with a 200mm gauge length; and the other with a 600 mm

gauge length. According to the menufacturer (Mayes, 2001), one gradation on the dial

gauge of the two above-mentioned DEMEC gauges is equivalent to 8.0 and 2.7

microstrains, respectively.

The Araldite epoxy glue

“Araldite 5 Minute” is a widely used two-part epoxy adhesive famous for its

relatively quick setting time and excellent adhesion. Manufactured by Selleys Pty.

Ltd, it can be purchased at any local hardware shop in Australia. Due to the following

properties stated by the manufacturer, this epoxy glue was used to provide the

required bonding between the DEMEC points place on columns/core walls surfaces

(Selleys Pty.Ltd, 2002);

Rapid setting time (4-7 minutes)

High bond strength

Good chemical resistance

For use on surface of most materials except thermoplastics

122

The manufacturer also noted some limitations in using Araldite 5 Minute as follows:

Bond strength weakening on exposure to temperature above 60°C

Not suitable for long-term immersion in water

Both of these were not seen as problem for the Q1 Tower.

Kestrel 3000 Pocket Weather Meter

A pocket weather meter, the Kestral 3000, was used in measuring humidity and

temperature at monitoring locations on the site simultaneously with the strain

readings. The Kestral 3000 has been designed to measure environmental conditions

immediately and accurately, as well as cope with abrupt changes. Furthermore, it is

portable and easy to use. An image of the Kestral 3000 is given in Figure 3.9.

Figure 3.9. The Kestrel 3000 Pocket Weather Meter

(http://www.nkhome.com/ww/3000/3000.html)

The Kestral 3000 is equipped with a quick response external thermistor for detecting

temperatures, a thermally adjusted humidity sensor for sensing the environmental

humidity and a user-replaceable impeller for measuring of wind speed. Environmental

factors measured by the Kestral 3000 include: wind speed; maximum wind gust;

average wind speed; air, water and snow temperatures; wind chill; relative humidity;

heat stress index; and dew point. Precision of the Kestral 3000 is given in Table 3.3.

http://www.nkhome.com/ww/3000/3000.html

123

Table 3.3. Specifications of the Kestral 3000 (Nielsen-Kellerman, 2004)

Measurement Unit of MeasurementAccuracy Range

Current, Max and Avg

Wind Speed

m/s, km/hr, mph,

fpm, kts, Beaufort

Force

±3% of reading

or ±0.1 m/s

0.3 to 40 m/s

(0.7 to 89 mph)

Temperature, Wind Chill,

Heat Stress, Dewpoint ºC, ºF ±1ºC (±2ºF)

-29 to 70ºC (-

20 to 158ºF)

Relative Humidity % ±3% or scale 5 to 95%

3.2.2.2 Instrumentation details

During the course of data collection, a large number of columns and core walls in

various locations of the Q1 Tower were selected to monitor and measure axial

shortening using the DEMEC gauge. Both types of vertical elements are considered

independently due to their properties and functions, i.e. size, shape, concrete grade,

applied stress level and amount of steel reinforcement. Thus, it was anticipated that

each element would display different magnitudes of axial shortening.

Unlike most laboratory experiments, field measurements greatly depend on

accessibility. The degree of accessibility to the columns and walls being monitored

depends on the construction activities in progress of the given time of measurement.

In order to avoid common problems of inaccessibility, most foreseen obstructions

during the course of data collection were analysed. Available information such as the

construction programme of the contractor (Sunland Group Pty) was studied, and the

following data collection scheme was proposed.

Instrumented levels and locations at the Q1 Tower

Prior to commencement of construction, the levels at which columns and walls are to

be monitored were determined. The following levels were considered: basement

124

levels B2 and B1, levels L1 to L8, level L31, level L49 and level L63. Approximate

location, elevation (metres) and nominal concrete grade (MPa) used at each

instrumented level is given in Figure 3.10.

Figure 3.10. Instrumented levels and concrete grades (Whaley, 2002)

The number and locations of instrumentated elements on each desired level were

determined based on accessibility, estimated duration of data collection and the

125

element’s proneness to axial shortening. The locations at selected levels are

summarized in Table 3.4.

Table 3.4. Identification of instrumented elements at each selected level and their

corresponding duration of strain measurement

126

Columns Core walls

B2 14 7 TC01, TC02, TC03, TC04, TC05,

TC06, TC07, TC08, TC09, TC10,

TC11, TC12, TC13, TC14, TW06,

TW07, TW09, TW10, TW16,

TW17, TW18

432-906

B1 13 8 TC01, TC02, TC03, TC04, TC06,

TC07, TC08, TC09, TC10, TC11,

TC12, TC13, TC14, TW06,

TW07,TW08, TW09, TW10,

TW16, TW17, TW18

259-884

L1 5 2 TC02, TC05, TC06, TC10, TC12,

TW08, TW09

49-209

L2 2 2 TC10, TC12, TW08, TW09 56-159

L3 5 3 TC02, TC05, TC06, TC10, TC12,

TW08, TW09, TW10

22-141

L4 5 2 TC02, TC05, TC06, TC10, TC12,

TW08, TW09

13-105

L5 2 3 TC06, TC10, TW08, TW09, TW10 48-83

L6 2 3 TC06, TC10, TW08, TW09, TW10 35-77

L7 1 3 TC06, TW08, TW09, TW10 49-56

L8 1 3 TC06, TW08, TW09, TW10 42-49

L31 10 2 TC01, TC03, TC04, TC05, TC06,

TC07, TC08, TC11, TC13, TC14,

TW08, TW18

89-139

L49 9 0 TC03, TC05, TC07, TC08, TC09,

TC10, TC11, TC13, TC14

49-104

L63 6 5 TC02, TC04, TC06, TC08, TC09,

TC10, TW01, TW09, TW10,

TW16, TW18

101-121

Number of

instrumented

elementsLevel Identification Code

Total duration

of strain

measurement

(days)

Number of measurement stations per location

127

At each desired location, the required numbers of DEMEC points are marked as soon

as access is gained. The age of concrete when the DEMEC points are marked varies

on the construction phase at each location. For circular columns, one to three rows of

DEMEC stations determined from the three pairs of DEMEC points, are set up around

its circumference to obtain reliable axial shortening measurements. However, for

rectangular columns and core walls, one to three rows of DEMEC stations were

mounted on one side only. This is based on the assumption that columns and core

walls of Q1 Tower are predominantly axially loaded and as such, the properties of

concrete are uniform throughout the elements and axial shortening is equally

developed on all sides.

According to past research (Swamy and Potter, 1976; Arumugasaamy and Swamy,

1989), variations of (measured) concrete surface strains of columns are noted to be

unsystematic. This may be attributed to haphazard distribution of concrete properties

owing to segregation, bleeding and other factors that arise during the course of

production and placing of in-situ concrete. Thus, the variation tends to be uniformly

distributed on any single side of the column or wall and therefore, an average reading

of many points should ascertain reliable results. Therefore, measurements taken from

many stations located at one side of the columns and walls are considered adequate.

Some typical layouts of the DEMEC points marked on the face of columns and walls

are shown in Figure 3.11.

128

Figure 3.11. Typical positioning of DEMEC points on columns/walls.

Instrumented rectangular columns and the recording of axial shortening

measurements using a 200mm DEMEC gauge is shown in Figure 3.12.

129

Figure 3.12. Instrumented column and DEMEC gauge in use

The number of stations at each individual location was determined based on the

precision of the DEMEC gauge used and the accessible surface area of the element. In

general, when comparing DEMEC gauges of the same type, those with longer gauge

lengths usually result in higher-accuracy measurements. Due to this reason, the

number of stations at a single element varies from 3 to 9 pairs, using a 200 mm

DEMEC gauge. The number was reduced to 1 to 2 pairs, when using a larger 600 mm

DEMEC gauge. The strain was calculated as the average of all DEMEC gauge station

readings for each element.

130

Table 3.5. Number of DEMEC stations using 200mm DEMEC gauge

Level Instrumented elements

Number of

DEMEC

stations

B2 TC10-4, TC12-4, TW06, TW07, TW10, TW16, TW17 3

TC01, TC07, TC08, TC09, TC11, TC12, TC13, TC14, TW18 6

TC02, TC03, TC04, TC05, TC06, TC10, TW09 9

B1 TC10-4, TC12-4, TW06, TW07,TW08, TW10, TW16, TW17 3

TC04, TC07, TC08, TC09, TC11, TC12, TC13, TC14, TW18 6

TC01, TC02, TC03, TC06, TC10 9

L1 TC10-4, TC12-4 3

TC02, TC02-4, TC05, TC05-4, TC06, TC06-4, TW08, TW09 6

TC10, TC12 9

L2 TC10-4, TC12-4 3

TC12, TW08, TW09 6

TC10 9

L3 TC02, TC02-3, TC05-2, TC06-2, TW09, TW10 3

TC05, TC06, TC10, TC12, TW08 6

L4 TC02, TW09 3

TC05, TC06, TC10, TC12, TW08, 6

L5 TC06-2, TW09, TW10 3

TC06, TC10, TW08 6

L6 TC06 3

TC10, TW08, TW09, TW10 6

L7 TW09 3

TC06, TW08, TW10 6

L8 TC06, TW08, TW09, TW10 6

L63 TC02, TW01, TW09, TW10, TW16, TW18 3

TC04, TC06, TC08, TC09, TC10 6

Table 3.6. Number of DEMEC stations using 600mm DEMEC gauge.

131

Level Instrumented elements

Number of

DEMEC

stations

L31 TC01, TC03, TC04, TC05, TC06, TC07, TC11, TC13,TW08,

TW18 1

TC08, TC14 2

L49 TC05, TC11, TC14 2

TC03, TC07, TC08, TC09, TC10, TC13 3

General procedures of DEMEC station installation and measurements

The basic steps in the installation of DEMEC stations and data reading are as itemized

below:

- DEMEC station installation commences as soon as the desired location is

accessible, and the formwork stripped.

- The highest pair of DEMEC points is initially attached to the column surface

at marked locations using Araldite. The adhered points are left undisturbed for

at least 1 hour to ensure rigid bond strength between the concrete surface and

DEMEC points.

- To finalise a complete set of DEMEC point instrumentation, the lower pair of

DEMEC points is adhered to the column surface using the invar setting-out

bar to ensure precise and original gauge lengths. The levels of each set of

DEMEC points are checked using a plummet. The instrumented column is left

undisturbed for at least 1 day prior to the initial gauge value reading.

- The value of the first gauge reading is taken of the initial length of the gauge

point where the strain occurred and will be subsequently monitored. The

gauge readings were also taken on the Invar reference bar as a control.

- At specific intervals, subsequent gauge readings as well as the reading on the

Invar reference bar are taken. The readings continue until access to the

location is terminated as construction work progresses or until the structure is

completed.

132

3.3 Laboratory testing

In addition to the in-situ data collection described earlier in Section 3.2, a large

number of control specimens were prepared to determine the material properties of

the concrete used on site. The tested properties include compressive strength, elastic

modulus, creep, and shrinkage.

3.3.1 Concrete

Concrete used in the numerous columns and core walls of Q1 Tower are of three

nominal concrete grades in line with their target compressive strengths, i.e. 65 MPa,

50 MPa and 40 MPa. All concrete specifications meet the requirements of the

Australian Standard AS1379-1997; Specification and Supply of Concrete. All

concretes were produced in accordance with the typical mix design of commercial

concrete mixes which are locally available and widely used in the Gold Coast and its

vicinity. Note the slump for all vertical elements was 150 mm.

The concrete was batched at the Readymix concrete plant located on Hinde Street,

Southport, Gold Coast, approximately five kilometers from the job site thus ensuring

a travel time of not more than 30 minutes. Upon delivery, a slump value of 150 mm

was verified according to AS1012.3.1–1998 prior to placement. Where the slump test

failed, a rejection was ordered.

The concrete components met the requirement of the Australian Standards, i.e.

AS3972-1997, Portland and blended cements; AS2758.1-1998, Aggregates and rock

for engineering purposes – Part 1: Concrete aggregates; AS3582.1-1998,

Supplementary cementitious materials for use with Portland and blended cement –

Part 1: Fly ash; AS3582.3-1994, Supplementary cementitious materials for use with

Portland cement – Part 3: Silica fume; and AS1478.1-2000, Chemical admixtures for

concrete, mortar and grout – Part 1: Admixtures for concrete.

Ordinary Portland cement, Type GP for general purposes was used in all concrete

mixes. The coarse aggregate consists of finely grained crushed stone with a maximum

133

size of 20mm. Other aggregates used were coarse river sand and fine sand.

Admixtures used consisted of the commercially available materials. The concrete mix

designs of vertical elements are given in Table 3.7 and material sources/brand names

and specifications are summarized in Table 3.8.

Table 3.7. Concrete mix compositions (Courtesy of Readymix Pty, 2004).

Concrete grade 40 50 65

Mix constituents

GP cement 280 345 460

Fly Ash 95 115 150

20mm aggregate 660 680 720

10mm aggregate 340 340 340

Coarse sand 590 550 460

Fine sand 290 250 190

Admixtures

Air entraining agent 0 0 0

Water reducer 1500 1840 1920

Superplasticizer 375 460 480

retarder 375 460 480

Weight (kg/m3)

Volume (cm3/m

3)

Table 3.8. List of material sources/brand names (Courtesy of Readymix Pty, 2004).

Materials Descriptions Source/supplier Standard Spec

Cement QCL Bulwer type GP cement QCL AS 1379

Coarse aggregate 10mm Beenleigh quarry AS 2758.1

20mm Beenleigh quarry AS 2758.1

Fine aggregate Fine sand Jacob well AS 2758.1

Coarse river sand Ipswich quarry AS 2758.1

Fly Ash Fine grade Tarong AS 3582.1

Daratard 77 Water-reducing initial set retarder Grace Co.Ltd. AS 1478.1

Darex AEA Air entraining admixture Grace Co.Ltd. AS 1478.1

Daramene R Ste retarding admixture Grace Co.Ltd. AS 1478.1

Daracem 19A Superplasticizer Grace Co.Ltd. AS 1478.1

134

3.3.2 Specimens Sampling

The sampling scheme was carefully planned to ensure that the properties of the

sampled concrete well represent actual elements of the structure, given that a

minimum number of samples were taken for each test. On the other hand, the scheme

was planned such that a broader population was covered with a smaller number of

specimens as possible.

Specimens for each of the three concrete grades used in the columns and core walls,

i.e. 65 MPa, 50 MPa and 40 MPa were taken from the batches of concrete at the

different levels based on the in-situ concrete usage plan stipulated by the designer. It

is assumed that 5 samplings, for each concrete grade would well represent the

concrete placed on site to a certain acceptable degree of accuracy. The in-situ concrete

usage plan for the vertical structural members at the Q1 Tower is presented in Table

3.9 below.

Table 3.9. Concrete grades of vertical members at various levels and corresponding

number of sampling

Concrete Grade (MPa) Level Number of Floors Number of Samplings

65 B2-L09 11 5

50 L10-L39 29 5

40 L40-L78 38 5

Specifically, specimens of the three concrete grades were randomly taken from the

following levels.

Specimens for Grade 65 MPa concrete were taken from levels L2, L3, L4, L5,

and L6.

Specimens for Grade 50 MPa concrete were taken from levels L15, L20, L25,

L30 and L35.

135

Specimens for Grade 40 MPa concrete were taken from levels L45, L50, L55,

L60 and L65.

In summary, the required samples commenced at level L2 with samplings of Grade 65

MPa concrete and ended at level L65 with sampling of Grade 40 MPa. A total number

of 15 samplings were taken during the course of construction which lasted

approximately 3 years.

The number of test samples taken for each sampling varied depending on the

specified test type and age. In general, thirteen to fourteen 200mmx100mm diameter

cylinders and three 75x75x75mm prisms were obtained for each sampling. These

samples were for testing of compressive strengths; elastic modulus and drying

shrinkage. In addition, for each of the three concrete grade samples, an extra set of 15

samples of 300mmx150mm diameter cylinders were taken for creep test from one of

the five grade samplings arranged.

The actual sampling locations at each specified level, where in-situ concrete was set

aside for casting the specimens, are randomly picked on site. This would not affect the

accuracy of tests given that all vertical members on any given floor were cast with the

same batch of concrete. The number of specimens for each sampling is detailed in

Table 3.10.

Table 3.10. Details of specimens collected at each sampling

136

Sampling

No.Level

Concrete

Grade

(MPa)

Number of

Ø100x200mm

cylinders

Number of

75x75x280mm

prisms

Number of

Ø150x300mm

cylinders

1a L2 65 13 3 0

1b L3 65 13 3 0

1c L4 65 14 3 0

1d L5 65 13 3 0

1e L6 65 13 3 15

2a L15 50 13 3 0

2b L20 50 13 3 0

2c L25 50 14 3 0

2d L30 50 13 3 0

2e L35 50 13 3 15

3a L45 40 13 3 0

3b L50 40 13 3 0

3c L55 40 14 3 0

3d L60 40 13 3 0

3e L65 40 13 3 15

3.3.3 Specimen preparation

All specimens were prepared on site by Readymix Pty according to AS1012.1-1993 at

the specified times during the course of construction. The code describes the method

for obtaining samples of freshly-mixed concrete directly from mixers and from

concrete deposited in readiness for casting. During concrete placement of the selected

concrete mix, a portion of fresh concrete was collected, using a 150 litre capacity.

This concrete was then set-out for the casting of test specimens and the required field

consistency tests, i.e. slump test, compacting factor test, vebe test and compactibility

index test.

The moulds conform to AS1012.8.1-2000, Section 7.2 for cylinder specimens and

AS1012.13-1992, Section 7.2.1 for prism specimens (shrinkage) specimens. Concrete

in both moulds were compacted using the rodding technique as described in

AS1012.8.1-2000, Section 7.3 and tamping technique as described in AS 1012.13-

1992, Section 7.2.2. Both the moulding and compacting processes were strictly

monitored and completed within 20 minutes of concrete dumping.

137

Soon after casting, all freshly cast specimens in the moulds were moved to the site

storage area where they were kept overnight prior to demoulding. The cylindrical

specimens were covered with individual metal caps whereas plastic sheets were used

for the prismic specimens, in order to prevent premature loss of moisture and potential

strength gradients, due to drying.

The moulds were stripped off at site approximately 24±2 hours from the time of

moulding and then transported to the test facility at NATA (National Association of

Testing Authorities, Australia), a certified laboratory of Readymix Technical Services

in Beenleigh, Queensland, to test for compressive strength, elastic modulus and

shrinkage specimens, with exception of the 300mmx150mm diameter samples that

were sent to the Civil Engineering Laboratory at the University of Queensland in

Brisbane for creep tests.

In order to meet the standards of the laboratory, it is mandatory that the time between

moulding and entry into the standard moist-curing condition must never exceed 36

hours. In addition, the concrete must be well saturated and in a good physical

condition.

At the Readymix laboratory, the compressive strength and elastic modulus specimens

were cured in accordance with AS1012.8.1-2000, Section 9 for standard tropical zone,

i.e. they were stored in lime-saturated water at a controlled temperature of 27±2ºC

providing the required standard moist curing conditions until the date of testing.

Shrinkage specimens were also placed in the standard moist curing conditions in a

similar way but they were kept in lime-saturated water for only 7 days and then

moved to the drying room which provides a controlled environment that meets the

standard drying conditions (temperature 23±1ºC, humidity 50±5%).

At the University of Queensland’s laboratory, the creep specimens were cured under

standard moist curing conditions for 7 days and thereafter, under the standard drying

conditions until the commencement of tests at age 28 days. The standard moist curing

138

and drying conditions were in accordance with AS 1012.8.1-2000 and AS 1012.13-

1992 respectively.

3.3.4 Compressive Strength Test

Compressive strength specimens (200mmx100mm diameter concrete cylinders) were

prepared for each concrete mix as described in Section 3.3.3, which is in accordance

with AS 1012.8.1-2000: Method for Making and Curing Concrete-Compression and

Indirect Tensile Test Specimens. The specimens were then tested in accordance with

AS 1012.9-1999: Method for Determination of the Compressive Strength of Concrete

Specimens.

The number of specimens allocated for the test per sampling depends on the required

number per test age. Generally, two specimens were required for each test age thus,

the number of specimens range between ten to twelve specimens per sampling in

which five to six test ages are specified (see Table 3.11)

Table 3.11. Number and test ages of compressive strength specimens

139

Sampling number Number of Specimens Test Age (days)

1a 10 7, 28, 56, 90, 180

1b 10 7, 28, 90, 180, 365

1c 14 7, 28, 90, 180, 365, 550, 730

1d 10 7, 28, 56, 90, 180

1e 10 7, 28, 90, 180, 365

2a 10 7, 28, 56, 90, 180

2b 10 7, 28, 90, 180, 365

2c 14 7, 28, 90, 180, 365, 550, 730

2d 10 7, 28, 56, 90, 180

2e 10 7, 28, 90, 180, 365

3a 10 7, 28, 56, 90, 180

3b 10 7, 28, 90, 180, 365

3c 14 7, 28, 90, 180, 365, 550, 730

3d 10 7, 28, 56, 90, 180

3e 10 7, 28, 90, 180, 365

At each designated test age, two samples were removed from the curing tank and

when visibly dried, the dimensions and weight of each sample were recorded. The

samples were thereafter capped with a filling of a sulphur mixture in accordance with

the procedure described in AS 1012.9-1999, Section 6. Note that a rubber cap was not

allowed to be used in this study.

Following final inspection of the samples, they were placed in the compression testing

machine, one at a time. The sample was then stressed smoothly and continuously at a

rate equivalent to 20±2 MPa per minute (this is equal to a loading rate of 157±16

kN/min for 200mmx100mm diameter cylinder samples) until failure. The maximum

force applied to the specimen was manually recorded.

The compressive strength (MPa) of test specimens was then calculated by dividing the

maximum force (106N) by the cross-sectional area (mm). The average of the two

specimens was then taken as the final value, representative of the compressive

strength at the test age.

3.3.5 Modulus of Elasticity Testing

140

The elastic modulus of field concrete specimens was measured in accordance with

procedures outlines in AS 1012.17-1997, Method for the Determination of the Static

Chord Modulus of Elasticity and Poisson’s Ratio of Concrete Specimens. The test

specimens were prepared in the same manner as done for the compressive strength

specimens. Similarly, the specimens were Ø100x200 mm concrete cylinders.

Once, the modulus specimens were removed from the curing tank, they were left to

air-dry and then, capped with sulphur mixtures at the date of test. Simultaneously, the

dimensions and weight of each specimen were recorded. After final inspection, the

specimens were placed along with the compressometer, which is a deformation-

measuring apparatus designed particularly for the modulus test.

The compressometer consists of two yokes, a mechanical averaging device, and a

gauge, which fully complies with AS 1012.17-1997, Section 2.3.2, Clause (c). An

image of the compressometer arrangement is shown in Figure 3.13 below.

Figure 3.13. Compressometer setup for Elastic Modulus tests (AS 1012.17-1997).

The whole compressometer arrangement was then placed in the compression testing

machine. AS 1012.17-1997, Method 1 specifies the test load as 40 percent of the

mean compressive strength of more than 2 specimens, tested in accordance with AS

141

1012.12 immediately prior to the modulus test. However, in this experiment, control

specimens for determining compressive strength was not necessary since the

compressive strength test of the same concrete was carried out on the same day prior

to the modulus test. Therefore, the strength test results obtained from the test

described in Section 3.3.4 was used.

The loading procedure strictly followed AS 1012.17-1997; Section 2.5.2, i.e. the

specimen has to be loaded three times. The rate of loading was taken as 15±2 MPa per

minute, which is equal to a loading rate of 118±16 kN/min for 200mmx100mm

diameter cylinders. The first loading served as a dummy run, for the purpose of

seating the gauge. Therefore, only results of the two subsequent readings were

recorded and the average was taken as the final value and representative of the elastic

modulus at the test age.

The results read from the gauge are the applied load at longitudinal strains of 50

microstrain and respective deformations at the test load. The elastic modulus (MPa)

can therefore be calculated from the following relationship:

00005.02

12 GGE Equation 3.1

where:

G1 = applied load at a strain of 50 microstrain divided by the cross-sectional area of

the unloaded specimen (MPa);

G2 = test load divided by the cross-sectional area of the unloaded specimen (MPa);

ε2 = deformation at test load divided by the gauge length (microstrain).

Three specimens per 1 sampling were used for determination of the modulus of

elasticity. However, it is noted that there were four out of the five samplings for each

concrete grade used for the modulus test (see Table 3.12). The specified ages at which

the modulus test was conducted were 7, 28, 56, 90, 180, 550 and 730 days.

142

Table 3.12. Number of elastic modulus specimens for each sampling.

Sampling number Number of Specimens

1a 3

1b 3

1c 0

1d 3

1e 3

2a 3

2b 3

2c 0

2d 3

2e 3

3a 3

3b 3

3c 0

3d 3

3e 3

3.3.6 Shrinkage Testing

Unrestrained shrinkage testing of samples taken from the concrete supplied to Q1

Tower was carried out in accordance with AS1012.13-1992; Method for

Determination of the Drying Shrinkage of Concrete for Samples Prepared in the Field

or in the Laboratory. This test method involves measuring the length changes of the

75x75x280 mm concrete prisms.

Three specimens per sample were taken for shrinkage test, as earlier mentioned in

Table 3.10. A total number of fifteen specimens were prepared for shrinkage test for

each concrete grade (forty-five specimens in total). Shrinkage test measurements were

taken from these specimens for duration of two years. The designated ages of

specimens at the times of shrinkage readings are listed in Table 3.13.

143

Table 3.13. Shrinkage measurement reading schedule.

Reading Age of Specimens

(days)Remarks

1 7 Initial measurement

2-13 14-90 Weekly measurement

14-34 120-720 Monthly measurement

As seen from Table 3.13, the initial reading was taken 7 days after casting the

specimens. Subsequent measurements were carried out on a weekly basis until the

specimens were ninety days old. Thereafter, monthly readings were taken until the

specimens are two years old.

As described earlier in Sub-Section 3.3.3, shrinkage specimens were cured in a lime-

saturated water tank at 27±2ºC for 7 days. The specimens were then removed from the

tank for the initial measurement of the horizontal length and thereafter stored in the

drying room at 23±1ºC and 50±5% relative humidity throughout the remaining test

duration. The measurements were taken using a horizontal length comparator in the

drying room at particular intervals.

The drying shrinkage of a sample over any period of drying time can be calculated by

subtracting its length (mm) at the given time, from the mean initial length (mm) and

divided by the difference in the original effective gauge length (taken as 250mm for

the apparatus used in this study). The resulting drying shrinkage is expressed in

microstrain.

3.3.7 Creep Testing

A series of creep tests were carried out on concrete specimens undergoing creep under

constant stress. The test was carried out in accordance with AS1012.16-1996, Method

for Determination of the Creep of Concrete Cylinders in Compression. The test

specimens were also prepared in accordance with AS1012.8.1-2000. The standard

cylinders of Ø150x300mm were used for the creep test. One set of specimens for one

144

creep test consists of 15 concrete cylinders. This comprises of 3 cylinders to be tested

in the loading frame, 2 cylinders for the preparation of dummy blocks, 3 cylinders to

be tested for compressive strength at 28-days, 3 cylinders to be used as controls for

deformations from causes other than loads, and 4 cylinders as spares.

The specimens for creep tests were cured as described in Sub-Section 3.3.3. At 28

days, compressive strength of three cylinders was determined; dummy blocks were

cut out from two cylinders; and three test cylinders were stacked in the loading frame.

Three sets of loading frame were used with each frame providing a means for creep

reading for each of the three concrete grades. It was required that the creep rigs be

stored in standard drying conditions at all time during the test. A schemetic diagram

of a loading frame given in AS 1012.16-1996 is represented in Figure 3.14.

145

Figure 3.14. Creep test (AS 1012.16-1996)

The DEMEC gauge was used to measure deformation in both the creep specimens as

well as the control specimens. To facilitate the use of the DEMEC gauge, gauge

points were fixed prior to the test, on the three creep specimens stacked in the frame

as well as the three control specimens stored in the same room. On each specimen,

three gauge points at approximately 120 degree of the cylindrical circumference were

marked out. The average of the three is taken as the final value.

According to AS1012.16-1996, the test specimens were then loaded using a portable

jack with a load equivalent to 40 percent of its compressive strength determined

earlier on the same day. It is worth to note that this does not actually happen on site.

Strain values of the test and control specimens were taken immediately after loading.

The load was then maintained by a system of springs.

Creep strain can be calculated by subtracting the average strain of control specimens

and average immediate strain at loading from the total strain measured from the

loaded specimen. The creep test was conducted over a period of two years and the

specified timings for creep readings are given in

Table 3.14.

Table 3.14. Creep measurement schedule.

Reading Age of Specimens

(days)Remarks

1 28 Immediately after loading

2 28 2 hours after loading

3 28 6 hours after loading

4-10 29-35 Daily readings for 7 days

11-14 42-63 Weekly readings for 4 weeks

14-36 90-720 Monthly readings until age 2 years

As seen from

Table 3.14, the initial creep readings were taken twenty-eight days after casting of the

specimens, immediately after application of the test load. Subsequent readings were

146

taken within the same day at two and six hours, after loading. Thereafter, readings

were taken on a daily basis for seven consecutive days followed by weekly readings

for four weeks. Finally, monthly readings were taken until the specimens were two

years old.


The experiments conducted on the Q1 Tower and test conducted at Readymix

laboratory and elsewhere were fully described. It is evident that the Q1 Tower is

definitely one of the most important buildings recently built in terms of its

architectural and structural techniques employed during its design and construction.

The layout of basement floors with many selections of column size and shape enable

axial shortening study on columns and core walls with variable factors.

The most comprehensive in-situ data collection and laboratory testing were conducted

for this study. The work covers a large scope, i.e. instrumentation and monitoring of

axial shortening of approximately more than 70 columns and 30 core walls located on

different levels of the building, sampling of fresh concrete at site, and long-term

testing of more than 200 concrete specimens for compresive strength, elastic modulus,

shrinkage and creep.

The commonly used DEMEC gauge technique was employed for in-situ data

collection. This method was widely adopted in other overseas researches (see review

in Chapter 2, Section 2.6) and recently adopted by Australian researchers (Bursle et

al, 2003) for similar kinds of data collections. The experiment was more emphasized

in basement level locations where highest stress was applied but also provided a vast

amount data from other levels in which axial stresses were relatively lighter. The

required lab testing programme covers major concrete properties necessary for axial

shortening analysis. At certified laboratories, the tests were carried out strictly

according to current relevent Australian Standards. The long-term test duration for the

laboratory testing ensures the reliability of concrete data during the course of the in-

situ data collection.

147

3.5 References

Arumugasaamy, P., and Swamy, R.N., “In Situ Performance of Reinforced Concrete

Columns,” Long-Term Serviceability of Concrete Structures, ACI SP-117, 1989, pp.

87-111.

AUSTRALIAN STANDARD AS 1012.1., “Method of testing concrete – Sampling of

fresh concrete,” Standards Association of Australia, Sydney, Australia, 1993, 6 pp.

AUSTRALIAN STANDARD AS 1012.3.1., “Method of testing concrete –

Determination of properties related to the consistency of concrete – Slump test,”

Standards Association of Australia, Sydney, Australia, 1998, 8 pp.

AUSTRALIAN STANDARD AS 1012.8.1., “Method for Making and Curing

Concrete-Compression and Indirect Tensile Test Specimens,” Standards Association

of Australia, Sydney, Australia, 2000, 13 pp.

AUSTRALIAN STANDARD AS 1012.9., “Method for Determination of the

Compressive Strength of Concrete Specimens,” Standards Association of Australia,

Sydney, Australia, 1999, 13 pp.

AUSTRALIAN STANDARD AS 1012.12.1., “Method for Determination of the mass

per unit volume of hardened concrete – Rapid measuring method,” Standards


AUSTRALIAN STANDARD AS 1012.13., “Method for Determination of the Drying

Shrinkage of Concrete for Samples Prepared in the Field or in the Laboratory,”


AUSTRALIAN STANDARD AS 1012.16., “Method for the Determination of Creep

of Concrete Cylinders in Compression,” Standards Association of Australia, Sydney,

Australia, 1996, 9 pp.

148

AUSTRALIAN STANDARD AS 1012.17., “Method for the Determination of the

Static Chord Modulus of Elasticity and Poisson’s Ratio of Concrete Specimens,”


AUSTRALIAN STANDARD AS 1170.1-1989., “Minimum design loads on

structures (known as the SAA Loading Code) - Dead and live loads and load

combinations,” Standards Association of Australia, Sydney, Australia, 1989, 24 pp.


structures (known as the SAA Loading Code) - Wind loads,” Standards Association



structures (known as the SAA Loading Code) - Earthquake loads,” Standards


AUSTRALIAN STANDARD AS 1379., “Specification and supply of concrete,”


AUSTRALIAN STANDARD AS 1478.1-2000., “Chemical admixtures for concrete,

mortar and grout - Admixtures for concrete,” Standards Association of Australia,


AUSTRALIAN STANDARD AS 1538-1988., “Cold-formed Steel Structures Code,”


AUSTRALIAN STANDARD AS 1720.1-1997., “Timber structures - Design

methods,” Standards Association of Australia, Sydney, Australia, 1997, 177 pp.

AUSTRALIAN STANDARD AS 2758.1-1998., “Aggregates and rock for

engineering purposes - Concrete aggregates,” Standards Association of Australia,


149

AUSTRALIAN STANDARD AS 3582.1-1998., “Supplementary cementitious

materials for use with portland and blended cement - Fly ash,” Standards Association


AUSTRALIAN STANDARD AS 3582.3-1994., “Supplementary cementitious

materials for use with portland cement - Silica fume,” Standards Association of

Australia, Sydney, Australia, 1994, 9 pp.

AUSTRALIAN STANDARD AS 3600-2001., “Concrete structures,” Standards


AUSTRALIAN STANDARD AS 3700-2001., “Masonry structures,” Standards


AUSTRALIAN STANDARD AS 3972-1997., “Portland and blended cements,”


AUSTRALIAN STANDARD AS 4100-1998., “Steel structures,” Standards





Mayes Instrument Limited., “The DEMEC mechanical strain gauge,” [on-line]

Available from, http://www.mayes.co.uk/demec.htm; Accessed 20/11/2002.

Nielsen-Kellerman. Co.Ltd., “Kestral 3000 Pocket Weather Meter,” [on-line]

Available from, http://www.nkhome.com/ww/3000/3000.html; Accessed 15/2/2004.

Wikipedia Foundation, Inc., “Invar,” [on-line] Available from,

http://en.wikipedia.org/wiki/Invar; Accessed 21/08/2003.

http://www.mayes.co.uk/demec.htm

http://www.nkhome.com/ww/3000/3000.html

http://en.wikipedia.org/wiki/Invar

150

Readymix Pty. Ltd., Assorted Information on Concrete Supplied to The Q1 Tower,

2004.

Selleys Pty. Ltd., “Adhesives – Selleys Araldite Super Strong” [on-line] Available

from, http://www.sellys.com.au/content_306.htm; Accessed 20/11/2002.

Sunland Group Pty. Ltd., “Design Development Report of The Q1 Tower,” Ove Arup

and Partners, Hong Kong, China, Volumn 1, 2002.

Sunland Group Pty. Ltd., Structural Drawings and Documents associated with the Q1

Tower Construction, 2003.

Swamy, R.N., and Potter, M.M.A., “Long-term movements in an eight-storeyed

reinforced concrete structure,” Proceedings. Conference on Performance of Building

Structures. Glasgow, Scotland, 1976, pp. 413-429.

Whaley, B., “Design Development Report of The Q1 Tower,” Sunland Group Ltd,

Gold Coast, 2001.

Whaley, B., “Q1 Tower-The proposed world’s tallest residential building ar Sufers

Paradise, Australia,” Presented at the seventeenth Australasian Conference on the

Mechanics of Structures and Materials 12th

-14th

June 2002 in Gold Coast, 17ACMSM

Proceedings, Griffith University Gold Coast Campus, Gold Coast, Volume 1, Pages

23-29.

http://www.sellys.com.au/content_306.htm

151

4 Concrete Testing Results

4.1 Introduction

The concrete laboratory testing results acquired through standard tests, which partially

accounts for the complete experimental work for this research (i.e. testing of concrete

properties in the laboratory and measurement of axial shortening on columns and core

walls of Q1 Tower), is herewith presented.

Compressive strength, modulus of elasticity, drying shrinkage and creep are the major

components of axial shortening and are required to be known when axial shortening is

estimated using any methods that was reviewed in section 2.5. However, before an

accurate analysis of axial shortening can be made, it is essential that the computed

values of concrete properties using models are justified and accurate. If major

discrepancies occur, adjustment to the adopted model by using available test data of

the actual concrete used may be necessary in order to improve the accuracy of the

prediction.

In order to assess the time dependent properties of concrete used in the Q1 Tower and

to obtain data for justification of the methods used to predict the material properties,

standard testing for the four concrete properties fully described and mentioned in

Chapter 3 was conducted.

Results of the four properties tested are examined and compared to predictions

obtained using selected methods/codes which were previously reviewed in section

2.4. For each of the four properties, after the material model comparisons are made, a

preferred model was obtained and adjusted to give the best prediction for the test

results. These modified material models are adopted later in the process of axial

shortening modeling to ensure minimum discrepancy of the material property

predictions, when the axial shortening model itself is examined and adjusted.

152

4.2 Compressive strength

In practice, compressive strength is considered to be the most significant concrete

property. It is directly linked to the quality of concrete and therefore the value of

characteristic compressive cylinder strength of concrete at 28 days, f’c, is usually

specified by the design engineer for compliance purposes and generally used for

analysis purposes in structural design as well as in the axial shortening analysis.

According to Australian Standard AS 3600-2001 and AS 3600-2009, definition of f’c

is the material strength that taken as equal to the specified concrete grade or assessed

by standard test carried out in accordance with AS 1012.9. The CEB-FIP 1990 code

model states that f’c maybe estimated using the following Equation 4.1

8'

cmc ff Equation 4.1

where fcm = the mean compressive strength at the relevent age

Note that AS1379-1997 also gives the following expression;

skff cccm

' Equation 4.2

where kc = the assessment factor that vary between 1.25-3.2, s = standard deviation of

the compressive strength result.

When the compressive strength at a particular age other than 28 days is required, it

can be determined using any of the available concrete design code recommendations.

These recommendations include time series models with ratios of strength gains at

any ages (strength development curve) as compared to f’c.

Theoretically, the predictions achieved using these code recommendations should

give satisfactory results when two critical criteria are met: f’c of the supplied concrete

is similar to the specified values; the strength development rate is not significantly

153

different from the prototype concrete used. According to Neville (1995), factors

affecting compressive strength include Water-to-cement ratio, porosity of the cement

paste, type and volume of coarse aggregate.

However, the properties of concrete supplied to the job site, particularly in the case of

HSC, are usually governed by such performance measures as slump, pumpability and

durability rather than by the specified f’c. AS3600 and AS1379 adopt f’c as a

minimum criterion for strength to ensure compliance. Therefore, in order to meet both

criteria, the strength of the actual concrete delivered at site may well exceed the

designated value.

The excess f’c in the actual concrete has a direct effect on the accuracy of concrete

property predictions, and consequently may result in reduction of the accuracy of

axial shortening prediction because the characteristic strength, f’c, is frequently used

in the estimation of the stress/strength ratio and elastic modulus required in the axial

shortening prediction.

4.2.1 Analysis of Results

Given that compressive strength testing is a mandatory requirement to ensure good

quality control of the pre-mix concrete supplied to the site, compressive strength of a

total number of 836 samples which were tested are reported.

The number of compressive strength samples of concrete grades 65 MPa and 50 MPa

at various ages is given in Table 4.1. For grades 65 MPa and 50 MPa concrete, results

from 462 and 374 samples are reported respectively, for which the compressive

strength results are reported for 0 days up to 180 days. The number of test samples at

28 days is relatively high (147 and 151 samples for 65 and 50 MPa respectively) as

compared to the number of test sample at other ages. This indicates that the value of

f’c (5% upper bound results of all test samples) will be in closer agreement to the

value of fcm (average compressive strength).

Note that f’c can be calculated using the following expression:

154

n

sff cmc 96.1' Equation 4.3

where f’c = characteristic compressive strength, fcm = mean compressive strength, s =

standard deviation of the compressive strength result, and n = number of samples.

From the histogram of strength results at 28 days for grade 65 MPa concrete, as seen

in Figure 4.1, the test data is approaching a normal distribution. In addition, Figure

4.2, which plots compressive strength against the age of concrete, it can be seen that

strength development depends on the elapsed time. In all cases, early strength

develops at relatively faster rate particularly under 28 days and thereafter, this rate

decreases as concrete’s age increases. The strength development then slows

significantly as the concrete reaches an age of 56 days.

Table 4.1. Number of Tested Compressive Strength Samples at Various Ages.

Grade 65 MPa

0 1 2 3 4 5 6 7 days

2 1 4 15 12 16 6 100 count

8 9 10 14 21 28 31 32 days

4 2 1 1 2 147 1 10 count

56 57 90 91 179 180 days

99 1 23 3 2 10 count

Total 462

Grade 50 MPa

0 1 2 3 4 5 6 7 days

6 1 3 2 3 4 3 100 count

8 9 28 29 31 32 56 57 days

1 3 151 2 2 6 66 2 count

90 92 180 days

7 2 10 count

Total 374

155

0

5

10

15

20

25

30

35

40

56 63 70 77 84 91

Mor

e

f'c at 28 days (MPa)

Fre

qu

ency

Figure 4.1. Histogram of 28 days Compressive Strength Results, Grade 65 MPa.

Similarly, the histogram of the strength results at 28 days for grade 50 MPa concrete

plotted in Figure 4.3, show a more evident normal distribution of the results. Plots of

strength against time for grade 50 MPa concrete is shown in Figure 4.4, wherein a

similar trend of strength development as previously observed for the grade 65 MPa

concrete, is noted.

0

20

40

60

80

100

120

0 50 100 150 200

Age of concrete (days)

f' c (

MP

a)

156

Figure 4.2. Compressive Strength Results of Grade 65 MPa concrete.

0

5

10

15

20

25

30

53 58 64 69 74 80

Mor

e

f'c at 28 days (MPa)

Fre

qu

ency

Figure 4.3. 28 days’ Compressive Strength Results of Grade 50 MPa concrete.

0

20

40

60

80

100

0 50 100 150 200

Age of concrete (days)

f' c (

MP

a)

Figure 4.4. Compressive Strength Results, Grade 50 MPa concrete.

4.2.2 Test results as compared to predicted values

Compressive strength test results obtained from the 836 specimens of the grade

65MPa and grade 50 MPa concrete are evaluated and plotted in Figure 4.5 and Figure

4.6, respectively. Note that these specimens were collected and tested by Readymix

Pty, Australia, for quality control purposes. The plotted data are the 5% lower bound

157

values of the tested strength results in which at least 95% of the individual test results

exceeded these values. Also included are the predictions of strength against the age of

concrete using the current compressive strength models recommended in ACI 209R-

1992, CEB-FIP 1990 and AS3600-2009.

In both figures, it is seen that the actual f’c of the concrete supplied for construction

were notably higher than those predicted values obtained by any of the three codes at

any given age. However, it is clearly seen that all expected values from all codes

follow the same trend and there are no permanent variations.

0

10

20

30

40

50

60

70

80

90

100

0 10 20 30 40 50 60 70 80 90

Age (Days)

f' c

(MP

a)

f'c for Series1

ACI 209R-1992

CEB-FIP 1990

AS 3600-2009


65MPa.

158

0

10

20

30

40

50

60

70

80

90

100

0 10 20 30 40 50 60 70 80 90

Age (Days)

f' c

(MP

a)

f'c for Series2

ACI 209R-1992

CEB-FIP 1990

AS 3600-2009


50MPa.

4.2.3 Discussion

In both the 65 and 50 MPa cases, the actual f’c of concrete delivered to Q1 Tower

were higher than the specified f’c. Similar deviations were observed by Bursle et al

(2003) for the concrete delivered to the World Tower Building in Sydney, Australia

where axial shortening of columns was also investigated. For completeness, the actual

28 day compressive strengths of concrete used at the Q1 Tower and World Tower

Buildings is compared with the specified f’c and summarised in Table 4.2 below:

Table 4.2. Specified and Actual Compressive Strength at 28 days at Q1 (this

research) and World Tower (Bursle et al, 2003).

159

Source Specified f'c Actual f'c

f'c

deviation

(%)

Q1 (Series 2) 50 65 30

World Tower 60 72 20

Q1 (Series 1) 65 75 15

World Tower 80 90 12.5

As highlighted, the concrete strength for grade 50 MPa concrete deviated as high as

30%. However, this deviation decreases with higher grades. To calibrate the value of

the specified f’c, a correction factor, η , is introduced to account for the “additional

strength”:

'/ cc ff Equation 4.4

where fc = actual compressive strength at 28days (MPa)

For concrete mixes with available test data, η can be determined directly using

Equation 4.4. However, in the absence of test data, an equation for η is necessary.

Figure 4.7 shows the power regression analysis obtained through actual strength data

of all the buildings investigated. The following equation was derived for this

investigation as follows:

3103.0'3071.4 cf Equation 4.5

where 50 MPa≤ f’c ≤ 80 MPa:

160

n= 4.3071*f'c-0.3103

R2 = 0.898

1.05

1.1

1.15

1.2

1.25

1.3

1.35

0 20 40 60 80 100

Specified f'c

Actual/Specified f'c

Power (Actual/Specified f'c)

Figure 4.7. Regression of the Actual/Specified f’c values of Concrete Grade 50-80

MPa

Consequently, the adjusted 28-day compressive strength of concretes, f’ca, where 50

MPa≤ f’c ≤ 80 MPa is obtained through:

''

cca ff Equation 4.6

Prediction values of the adjusted compressive strength, f’ca, using ACI 209R-1992,

CEB-FIP 1990 and AS 3600-2009 code models incorporating the correction factor, η

as defined in Equation 4.4 and 4.5 is plotted in Figure 4.8 and Figure 4.9 for grades

65MPa and 50 MPa concrete respectively.

On comparing, f’c, plotted in Figure 4.5 and Figure 4.6 and the modified values, f’ca,

plotted in Figure 4.8 and Figure 4.9, it is noted that accuracy of the compressive

strength prediction is significantly improved in the latter. It is therefore recommended

that f’ca is used in compressive strength predictions.

161

0

10

20

30

40

50

60

70

80

90

100

0 10 20 30 40 50 60 70 80 90

Age (Days)

f' c (

MP

a)

f'c for Series1

ACI 209R-1992

CEB-FIP 1990

AS 3600-2009

Figure 4.8. Measured compressive strength over time and predictions by codes in

conjuction with η, series 1.

0

10

20

30

40

50

60

70

80

90

100

0 10 20 30 40 50 60 70 80 90

Age (Days)

f' c

(MP

a)

f'c for Series2

ACI 209R-1992

CEB-FIP 1990

AS 3600-2009

Figure 4.9. Measured compressive strength over time and predictions by codes in

conjuction with η, series 2.

Although more accurate predictions are obtained by incorporating the proposed

additional-strength factor, η, differences in the strength development rate of concrete

may be affected by the level of prediction accuracy. This prediction may be improved

162

by replacing the original time-ratio components recommended in the major codes,

with closely similar time-ratio relationships which is more difficult to accomplish.

The percentage strength development of the actual concrete and those calculated

using η, the excess-strength factor, in conjunction with the compressive strength

equations recommended in the current ACI, CEB-FIP and AS 3600 codes are shown

in Figure 4.10.

0

10

20

30

40

50

60

70

80

90

100

110

120

130

0 10 20 30 40 50 60 70 80 90

Age (Days)

Str

eng

th (

% o

f 2

8 d

ay)

Series 1

Series 2

ACI 209R-1992

CEB-FIP 1990

AS 3600-2009

Figure 4.10. Strength development of actual concrete and those due to codes

recommendations.

Apart from the rapid strength development of series 1 concrete, the percentages of

strength development based on ACI 209R-1992, CEB-FIP 1990 and AS 3600-2009

are similar as they satisfactorily agree with the test data. Therefore, any of these

models may be used in conjunction with Equation 4.6. To further improve the

prediction accuracy, data from Q1 Tower and World Tower Building were analysed

using the S-curve model with a curve estimation regression. This led to:

100

'/2.271.4' c

t

c

fetf Equation 4.7

163

where f’c (t) = compressive strength at time t (in days).

The improved level of agreement between Equation 4.7 and the measured data is

significant in Figure 4.11 especially for the 14 and 28-day strengths.

0

10

20

30

40

50

60

70

80

90

100

110

120

130

0 10 20 30 40 50 60 70 80 90

Age (Days)

Str

eng

th (

% o

f 2

8 d

ay)

Series 1

Series 2

Equation 4

Figure 4.11. Strength development of actual concrete data against those generated by

Equation 4.7.

Further comparisons were made with the 3 major codes and based on the actual test

results, the percentage errors were computed for all the four prediction formulas at

various ages, as shown in Table 4.3. It is evidenced that by using Equation 4.7

percent deviation of compressive strength predictions reduced significantly.

Table 4.3. Deviations in Compressive Strength of Models and Test Results.

Equation 4.7

164

Tested f'c(MPa) f'c (MPa) Dev (%) f'c (MPa) Dev (%) f'c (MPa) Dev (%) f'c (MPa) Dev (%)

3/series 1 44.9 29.8 33.7 38.9 13.4 39.0 13.2 39.9 11.2

7/series 1 56.5 45.7 19.0 50.6 10.4 48.0 15.0 60.6 -7.3

28/series 1 75.0 65.5 12.7 65.0 13.3 65.0 13.3 76.7 -2.4

56/series 1 83.6 70.5 15.6 69.9 16.3 70.2 16.0 79.8 4.5

90/series 1 89.6 72.7 18.9 72.6 18.9 76.7 14.4 81.0 9.6

3/series 2 35.2 27.0 23.2 33.1 5.9 30.0 14.8 34.7 1.5

7/series 2 47.5 35.2 26.0 38.9 18.0 37.0 22.1 52.7 -10.9

28/series 2 65.0 50.4 22.5 50.0 23.1 50.0 23.1 66.7 -2.6

56/series 2 68.0 54.3 20.2 53.8 20.9 54.0 20.6 69.4 -2.1

90/series 2 69.6 55.9 19.6 55.8 19.7 59.0 15.2 70.4 -1.3

Proposed EquationAge

(days)/series

ACI 209 CEB-FIP AS3600

4.3 Modulus of Elasticity

Measured elastic modulus of the Q1 concrete samples and their corresponding

compressive strengths are plotted in Figure 4.12. This plot is a direct representation of

the non-linear relationship of elastic modulus and measured compressive strength

based on the codes used in modeling.

With reference to Figure 4.12, the elastic modulus increases with compressive

strength. The average elastic modulus as concrete strength increases from 40 MPa to

90 MPa varies from 30000 MPa to 45000 MPa. However, data used in this figure are

results from concrete grades 65 and 50 MPa at various ages.

A prediction line of elastic modulus based on AS3600-2001 for concrete strength

varying from 20 MPa to 90 MPa is also indicated in Figure 4.12 and this may be used

for the Q1 Tower concrete to a certain level of accuracy (approximately ±20%). For

concrete strengths less than 70 MPa, the prediction based on AS3600 is conservative

as over-estimation is obtained. Regardless, the AS3600-2001 gives very good

prediction for concrete strengths greater than 70 MPa.

To improved the accuracy of prediction, a curve fitted to the test result (see Figure

4.12) is derived using a non-linear regression analysis, based on an logarithmic elastic

modulus equation:

6567924866 cmc fLnE Equation 4.8

165

y = 24866Ln(x) - 65679

R2 = 0.7853

0

10,000

20,000

30,000

40,000

50,000

60,000

0 20 40 60 80 100

fcm (MPa)

Ec

(MP

a)

Q1 Result

AS3600-2001

Log. (Q1 Result)

Figure 4.12. Elastic Modulus Results of Q1 Tower Concrete Samples versus AS

3600-2001 predictions.

Based on the measured compresive strength at various ages (refer to section 4.2) of

concrete grades 65 MPa and 50 MPa with density of 2350 kg/m3, a plot of the

measured elastic modulus results versus compressive strengths and predictions based

on the selected international standards/methods i.e. AS3600-2001, ACI 209-1992,

CEB-FIP (1990), AS3600-2009 and the proposed method (Equation 4.8) is presented

in Figure 4.13. It is noted that the Elastic Modulus formula recommended by ACI 209

is identical to that of AS3600-2001 thus resulting in identical modulus predictions.

In general, all models under investigation give reasonably good predictions as all their

prediction curves fall within the range of measured elastic modulus results. When

compared to the results based on AS3600-2001, the predictions based on CEB-FIP

tends to be higher for concrete grades below 70 MPa and lower for concrete grades

above 70 MPa.

When comparing the accuracy of the elastic modulus model adopted by AS3600-2009

and the earlier AS3600-2001 code, it is seen that the new code produces no better

166

results than the AS3600-2001 code for the concrete supplied. It is also noted that the

prediction line generalted by AS3600-2009 has the least slope as compared to other

models. However, despite the consistency observed for concrete strengths less than 60

MPa, the AS3600-2009 prediction line tends to divert from the data group as soon as

strength reaches 60 MPa and beyond.

Regardless of the trend of other adopted models, the proposed elastic modulus

equation gives the best fit of the test data.

0

10,000

20,000

30,000

40,000

50,000

60,000

0 20 40 60 80 100

fcm (MPa)

Ec

(MP

a)

Q1

AS3600-2001/ACI 209

CEB-FIP (1990)

AS3600-2009

Proposed Equation

Figure 4.13. Elastic Modulus Results of Q1 Tower Concrete Samples compared with

AS 3600-2001, CEB-FIP and AS3600-2009 Models.

The discrepancies between the various elastic modulus predictions based on the 3

models and the proposed equation for concrete strengths ranging from 40 to 90 MPa

is presented on Table 4.4. A deviation of up to 20% is noted for the mid-range

concrete strength (50 to 55 MPa) using the AS3600-2001/ACI 209 model. However,

prediction for other concrete strength using the same model yields reasonably

accurate results with maximum deviations less than 8%.

167

Table 4.4. Deviations of the Various Elastic Modulus Predictions.

Tested

fcm

Tested

Ec

(MPa) (MPa) Ec (MPa) Dev (%) Ec (MPa) Dev (%) Ec (MPa) Dev (%) Ec (MPa) Dev (%)

41.3 30670 31483 2.7 34334 11.9 31241 1.9 26844 -12.5

46.3 31644 33335 5.3 35668 12.7 32274 2.0 29686 -6.2

50.7 29095 34883 19.9 36763 26.4 33138 13.9 31943 9.8

54.6 30458 36200 18.9 37683 23.7 33873 11.2 33786 10.9

59.7 36724 37853 3.1 38821 5.7 34796 -5.3 36006 -2.0

64.9 36811 39467 7.2 39917 8.4 35696 -3.0 38083 3.5

70.7 39534 41192 4.2 41073 3.9 36660 -7.3 40212 1.7

74.9 43265 42391 -2.0 41866 -3.2 37329 -13.7 41638 -3.8

80.5 42698 43955 2.9 42889 0.4 38201 -10.5 43439 1.7

85.9 46368 45392 -2.1 43819 -5.5 39003 -15.9 45039 -2.9

89.9 45428 46437 2.2 44489 -2.1 39587 -12.9 46172 1.6

AS 3600-

2001/ACI 209CEB-FIP AS 3600-2009 Proposed Equation

Elastic Modulus prediction based on the CEB-FIP model appears to be less accurate

than those of the ACI 209 predictions. Deviations of both the mid-range concrete

strength and those less than 50 MPa are higher, at 26% (against 20%) and 12%

(against 5%) respectively. Note that for the AS3600-2009 model, deviations for

concrete strength less than 55 MPa is significantly less than those of AS3600-

2001/ACI 209 and CEB-FIP. However, prediction using the AS3600-2009 model for

concrete strength greater than 75 MPa yields larger deviations of 16%. As compared

to the other three models, the proposed equation reduces the magnitude of maximum

deviations over the total range of data. It is observed that only two points indicate

deviations greater than 10% (12.5% for 41.3 MPa and 10.9% for 54.6MPa).

Therefore the proposed equation should be used to represent elastic modulus

behaviors of the Q1 Tower concrete for purposes of axial shortening analyses of this

building.

168

4.4 Drying shrinkage

Drying Shrinkage is another important property of concrete, which is considered a

part of quality control along with other properties such as compressive strength. In

practice, the actual and final drying shrinkage of in-situ concrete should be as close to

the design specification as possible. This is to avoid adverse shrinkage effects such as

loss of prestress in slab, excess deflections in beams and furthermore, axial shortening

in columns and walls may be not effectively accounted for and, consequently this may

affect overall performance and durability of the structure.

Shrinkage is considered a key factor contributing to axial shortening. In columns and

walls, early age axial shortening occurs as a result of shrinkage since elastic

shortening and creep are by far in lesser amount because the columns/walls have not

been subjected to heavy load until later stages of the construction. Once the concrete

sets and develops its full strength its shrinkage also stabilizes or in other words

shrinkage development stops. Despite the various reports on long-term shrinkage

development, it is widely agreed that the amount of such shrinkage is relatively small

when compared to shrinkage that develops earlier. In addition, creep also becomes a

major contributor to long-term axial shortening after structure is subjected to its full

service load.

4.4.1 Analysis of Results.

In order to obtain actual shrinkage data for the in-situ Q1 Tower concrete to use for

further axial shortening analysis, a total number of 33 shrinkage samples were tested

for shrinkage properties based on AS 1012.13 as previously outlined in Section 3.3.6.

The tests were conducted for a period of 422 and 331 days for grades 65 and 50 MPa

concrete respectively. The number of samples, frequencies of shrinkage reading and

the duration of tests for both concrete grades are given in Table 4.5. Given that the

rate of shrinkage and its sensitivity to internal/external changes vary with time starting

at the commencement of drying therefore the interval of shrinkage readings for the

169

initial period of the test up to 90 days is set as weekly. After 90 days, the interval is

increased to once a month.

Table 4.5. Number of Shrinkage Samples and Test Schedule.

Concrete Grade

(MPa) No. of Samples Age at Testing

65 15

7 to 91 days (once a week),

91 to 422 days (once a month)

50 18

7 to 91 days (once a week),

91 to 331 days (once a month)

4.4.2 Shrinkage: Grade 65 MPa Concrete.

Plots of shrinkage results for the concrete samples of grade 65 MPa are shown in

Figure 4.14. Note that the plotted data points represent an average value of three

readings taken from a shrinkage sample at a given test age. The total number of data

for this set of tests is 323.

In addition, Figure 4.14 also outlines a fitted logarithmic curve for this data set. From

the regression analysis, the best fitted equation with R2

of 0.6818 results in a

logarithmic equation as follows:

24.322ln489.90 xy Equation 4.9

where y = shrinkage (microstrain) and x = drying time (days)

From Figure 4.14, it is seen that the fit line is traced all the way through the middle of

the band of plotted data. This therefore implies that the measured shrinkage data of

170

these samples is a band of data that extend approximately ± 25% from the median of

shrinkage values at any given times. This approximate percentage of scattering shows

that the quality of tested concrete in terms of shrinkage is well controlled since

AS3600 suggests that the shrinkage value of concrete may vary up to ± 40% from the

mean value.

y = 80.489Ln(x) + 322.24

R2 = 0.6818

0

100

200

300

400

500

600

700

800

900

1000

0 100 200 300 400 500

days

mic

rost

rain

65MPa

Log. (65MPa)

Figure 4.14. Shrinkage Results for Concrete Grade 65 MPa.

It is clearly observed that the magnitude of shrinkage of the 65 MPa concrete depends

on the elapsed drying time. As expected, early shrinkage progresses aggressively soon

after drying of concrete commences and the rate of shrinkage development thereafter

varies against time. A greater part of the shrinkage strain (80% of the final shrinkage

measured at 422 days) was generated within the first 56 days. At 100 days, the

average shrinkage strain was closer to its final value at approximately 700

microstrains. Beyond 100 days, shrinkage increased only slightly. And beyond 400

days, measured shrinkage remained at approximately the same value.

171

4.4.3 Shrinkage: Grade 50 MPa Concrete.

In addition to the shrinkage results of grade 65 MPa concrete previously discussed in

section 4.4.2, shrinkage results of grade 50 MPa concrete are presented in Figure 4.15.

Again, the graph represents an average value of three readings taken from a shrinkage

sample at a given test age. The total number of data for this set of test is 275.

A fitted logarithmic curve for this data set is also included in

Figure 4.15. From the regression analysis, it was found that the best fitted equation

with R2

of 0.566 also taking form of logarithmic equation and has an expression as

follows;

77.339ln85.79 xy Equation 4.10

where y = shrinkage (microstrain), x = drying time (days)

y = 79.85Ln(x) + 339.77

R2 = 0.566

0

100

200

300

400

500

600

700

800

900

1000

0 100 200 300 400 500

days

mic

rost

rain

50MPa

Log. (50MPa)

Figure 4.15. Shrinkage Results for Concrete Grade 50 MPa.

172

In general, the test results for grade 50 MPa concrete is observed to be very similar to

those of grade 65 MPa concrete. Moreover, the fitted curves of both data sets also

appear to be following a similar path. When comparing Equation 4.9 against Equation

4.10, it was noted that both fitted lines differ slightly in terms of their mathematical

expressions.

In order to investigate the similarities of both data sets further, shrinkage results of

both concrete grades and their fitted lines are combined together, as represented in

Figure 4.16.

For 65 MPa

y = 80.489Ln(x) + 322.24

R2 = 0.6818

For 50 MPa

y = 79.85Ln(x) + 339.77

R2 = 0.566

0

100

200

300

400

500

600

700

800

900

1000

0 100 200 300 400 500

days

mic

rost

rain

65MPa

50MPa

Log. (65MPa)

Log. (50MPa)

Figure 4.16. Combined Shrinkage Results of Concrete Grade 65 MPa and 50MPa.

It is evident from Figure 4.16 that shrinkage results of both groups are similar. This

leads to the conclusion that shrinkage values for 65 MPa and 50 MPa concrete used

are not dependent on the concrete grades. The only uncontrolled factor that could

induce differences in shrinkage of the two groups is the different compositions of the

65 MPa and 50 MPa concrete mixes. It is obvious that the effect of a difference in mix

composition on drying shrinkage is minimal in this particular study.

173

It is worthy to note that although the fitted lines may be used to represent shrinkage

values at any given times, its use beyond an age of 100 days may cause over-

estimation. Therefore, the validity of these fitted equations should be limited up to

100 days only. Beyond 100 days, a definite final shrinkage value should be defined

for the analysis purposes. For these particular cases, a final shrinkage value of 700

microstrains is considered suitable.

4.4.4 Test results compared with predicted values

Shrinkage results of grade 65 MPa, and 50 MPa concrete, and their corresponding

fitted lines are now plotted along with the shrinkage curve generated from the AS

3600-2001 code recommendation with the assumptipon that the final shrinkage equals

700 microstrains (see Figure 4.17).

Figure 4.17 indicates that AS 3600-2001 gives a relatively more accurate prediction in

accordance with the trend lines of both data sets. Also similar to the log equations,

prediction by AS 3600-2001 yields an over-estimation of shrinkage beyond 100 days.

Thus, the maximum predicted shrinkage values should be limited to 700 microstrain

once the age of concrete reaches 100 days.

For further comparison with other international codes, plots of the shrinkage results,

corresponding fitted lines, and the AS 3600-2001 shrinkage curve and other

predictions based on ACI 209, and AS3600-2009 is presented in

Figure 4.18.

174

For 65 MPa

y = 80.489Ln(x) + 322.24

R2 = 0.6818

For 50 MPa

y = 79.85Ln(x) + 339.77

R2 = 0.566

0

100

200

300

400

500

600

700

800

900

1000

0 100 200 300 400 500

days

mic

rost

rain

65MPa

50MPa

"AS 3600 (700)"

Log. (65MPa)

Log. (50MPa)

Figure 4.17. Drying Shrinkage Results and AS 3600-2001 predictions.

For 65 MPa

y = 80.489Ln(x) + 322.24

R2 = 0.6818

For 50 MPa

y = 79.85Ln(x) + 339.77

R2 = 0.566

0

100

200

300

400

500

600

700

800

900

1000

0 100 200 300 400 500

days

mic

rost

rain

65MPa 50MPa AS 3600-2001 (700)

ACI 209 (780) AS 3600-2009 Log. (65MPa)

Log. (50MPa)

Figure 4.18. Drying Shrinkage Results and Predictions based on Various Models.

175

4.5 Creep

In addition to the elastic modulus and shrinkage, creep is another important concrete

property that plays a vital role in the overall quality of the concrete as well as its long-

term serviceability as creep is a major factor contributing to axial shortening in

structures. Creep is taken as a strain in excess of the elastic strain at the time of

determination. This gradual increase in strain with time under load conditions is due

to creep. The increase in strain over time can be several times larger than the initial

strain.

4.5.1 Analysis of Results

During the course of design, creep and shrinkage values are incorporated based on the

estimated values recommended by the relevant codes. However unlike shrinkage

which can be verified by quality control tests, verification of actual creep for in-situ

concrete is rarely available due to the complicated nature of the test process and the

required time period of the tests.

The actual creep data of Q1 tower concrete is crucial to the estimation and modeling

of axial shortening. Creep and shrinkage occur simultaneously in built structures and

the amount of each individual property cannot be determined without actual test data.

Therefore, in order to accomplish this, three sets of creep rigs containing concrete

samples of grade 65, 50, and 40 MPa respectively were set up as outlined in the

testing procedures of AS 1012.16 (refer to section 3.3.7) for a test duration of up to

1100 days. Note that the creep test for this research was conducted at the University

of Queensland.

In addition to the creep test conducted on the specimens set in creep rigs, companion

specimens (6 nos. 300mmx150mm diameter cylinders for each concrete grade) were

tested for their compressive strength and shrinkage. For each concrete grade, 3

specimens were tested at 28 days for their compressive strength and the remaining 3

specimens were monitored simultaneously with the creep test, for their drying

shrinkage.

176

The type of test conducted, frequencies of creep and shrinkage reading and the

duration of test for the three concrete grades are summarised in Table 4.6. It is widely

known that the rate of creep development and its sensitivity to changes vary with time

starting at the commencement of sample loading. Therefore, the interval of creep

readings at the initial period of the test (up to approximately 90 days) is set for once a

week, whereas from 90 and 400 days, the interval is set at approximately once a

month. After 400 days, creep reading is taken at approximately every 100 days.

Table 4.6. Summary of Tests for Creep Specimens.

Concrete Grade

(MPa)Tested Property Testing Period

Frequency of

monitoring

65, 50, 40 Compressive strength 28 days

once only at

28th day

65

Drying Shrinkage and

Creep

28 to 99 days 99

to 378 days 378

to 1149 days

once a week

once a months

every 100 days

50


Creep

28 to 76 days

76 to 398 days

398 to 1073 days

once a week

once a months

every 100 days

40


Creep

28 to 59 days

59 to 227 days

227 to 662 days

once a week

once a months

every 100 days

4.5.2 Creep results: Compressive Strength at 28 days

With reference to AS1012.16, the Method for Determination of the Creep of Concrete

Cylinders in Compression requires that concrete cylinders set in a conformed test rig

177

being tested for creep shall be stressed once the age of concrete reaches 28 days at 40

percent of the compressive strength value and this applied stress shall be maintained

throughout the entire duration of the test.

In order to determine the test stress as specified by AS 1012.16, the 28-days

compressive strength of the three 300mmx150mm diameter cylinder specimens for

each concrete grade tested are determined. The average result of these three tested

specimens and the resulting test stresses for creep (40% of the tested 28-days

compressive strength) is presented in Table 4.7.

Table 4.7. Compressive strength at 28 days and corresponding test stresses.

Concrete

Grade (MPa)

28-days Strength

(MPa)

Test Stress

(MPa)

65 69.3 27.2

50 55 22.2

40 47 18.8

From this experiment, the compressive strength at 28 days of the 300mmx150mm

diameter cylinders appears to be less than results of the 200mmx100mm diameter

cylinders which was earlier reported in section 4.2. The effect of sample size on the

compressive strength are noted which is in good agreement with observation reported

by other researcher (Almudaiheen and Hanson, 1987). However, due to the small

number of 300mmx150mm diameter cylinders tested, the observed trend cannot be

taken as an absolute ruling as testing of more samples is necessary if further

verification of this phenomenon is required.

Regardless, the corresponding strength tests reported in this section is aimed at

obtaining the relevant test stresses which would be applied to samples in the creep

rigs. The test stresses for concrete grade 65, 50 and 40 MPa were 27.2, 22.2 and 18.8

MPa respectively.

4.5.3 Creep results: Elastic Modulus

178

Once the age of concrete reaches 28 days, the test load as calculated in section 4.5.2

was applied to the creep samples installed in the creep testing rigs, by means of a

hydraulic machine. The applied stress was then maintained by means of a spring

system in the creep rig.

Prior to the load application, the initial strain on the creep sample was recorded. The

magnitude of strain immediately upon application of the test loads for each of the

three concrete grade samples was calculated with reference to the initial recorded

strain. It is assumed that this instantaneous strain is purely elastic and directly related

to the elastic modulus of the concrete as creep and shrinkage has not yet developed.

Another assumption made is that the instantaneous strain remains constant beyond 28

days therefore any strain increases are due to the combined effects of creep and

shrinkage. Plots of the average instantaneous strains of the creep specimens obtained

from the three readings for each concrete grade is presented in Figure 4.19. The

instantaneous strain at 28 days and beyond are 848.9 microstrain, 854.7 microstrain

and 640.4 microstrain for grades 65 MPa, 50 MPa and 40 MPa concrete respectively.

179

0

200

400

600

800

1000

0 50 100 150 200

Number of Days Loaded

mic

rost

rain

Elastic Strain 65 MPa Elastic Strain 50 MPa Elastic Strain 40 MPa

Figure 4.19. Average Elastic Strain of Creep Specimens.

It is evident from Figure 4.19 that elastic strain of grade 40 MPa concrete is relatively

less than those of 65 MPa and 50 MPa, given that the applied test load is also lower.

However, the elastic strain of the grades 65 and 50 MPa concrete is observed to have

similar magnitudes.

To determine the elastic modulus values of the same, the elastic strain can be

incorporated in the formula as recommended in AS 1012.17-1997, reproduced

hereafter as Equation 4.11.

2

2GE Equation 4.11

180

where; E = elastic modulus (MPa), G2 = applied stress (MPa), ε2 = strain (10-6

mm/mm)

Elastic modulus of the creep samples and predicted values using the previously

proposed elastic modulus (Equation 4.8, referred to section 4.3) derived from the test

data of the same concrete but with relatively smaller specimen size (200mmx100mm

diameter cylinders) are plotted altogether with the test result in Figure 4.20.

0

10,000

20,000

30,000

40,000

50,000

60,000

0 20 40 60 80 100

fcm (MPa)

Ec (

MP

a)

Test Data

Prediction

Figure 4.20. Average Elastic Modulus of Creep Specimens and Prediction based on

Equation 4.8

It is noted from Figure 4.20 that the measured elastic modulus of the creep samples

are less than the elastic modulus given by Equation 4.8 for all tested concrete grades.

Furthermore, the elastic modulus of concrete grade 50 MPa is found to be

unexpectedly lower than those of grade 40 MPa concrete. This is quite unusual given

that higher strength concrete should have a greater modulus of elasticity which is

directly related to the concrete strength (Neville, 1995).

Numerical comparison of elastic modulus of the creep specimens and the

200mmx100mm diameter specimens is presented in Table 4.8. Deviations of the

181

elastic modulus for grade 40, 50 and 65 MPa concrete are evaluated at 2.34, 23.53 and

19.32 percent respectively. This implies that elastic modulus of concrete with

relatively higher compressive strength tend to be more sensitive to the variations.

Therefore, results obtained from a small number of samples may not be accurate. In

other words, elastic modulus values obtained from these creep samples (3 cylinders

for each concrete grade) may greatly deviate from their true values. In the case of

grade 50 MPa concrete, the measured elastic modulus of the creep samples is likely to

be about 5000-6000 MPa lower than the expected value.

Table 4.8. Elastic Modulus of Creep Specimens and Standard Specimens.

Concrete

Grade

(MPa)

Actual

f'c

(MPa)

Ec

Ø150x300mm

(MPa)

Ec

Ø100x200mm

(MPa)

Deviation (%)

40 47 29356 30058 2.34

50 55 25974 33967 23.53

65 69.3 32041 39714 19.32

Despite the discrepancies and uncertainties observed in the elastic modulus obtained

from the creep samples, the instantaneous strains reported in this section will be used

to determine the creep strain. However, the more reliable elastic modulus data

obtained from the rather rigorous experiment (refer to section 4.3) will be used in the

course of axial shortening modeling which is further discussed in detail on chapter

5.5.

4.5.4 Creep Results: Drying Shrinkage of creep specimens

In order to determine the drying shrinkage of the tested creep specimens, three

unloaded 300mmx150mm diameter cylinders for each concrete grade tested were

measured for drying shrinkage under similar controlled environment during the

duration of creep tests. The average drying shrinkage results of creep specimens

grade 40, 50 and 65 MPa are given in Figure 4.21.

182

0

100

200

300

400

500

600

0 200 400 600 800 1000 1200


mic

rost

rain

Shrinkage Strain 65 MPa Shrinkage Strain 50 MPa Shrinkage Strain 40 MPa

Figure 4.21. Drying Shrinkage of Creep Specimens.

Results from Figure 4.21 indicates that drying shrinkage after one year period ranges

from aproximately 285 microstrain to 420 microstrain for for concrete grade 65 MPa

to 40 MPa respectively. Given the elastic strain (see Figure 4.19) and the shrinkage

strain (see Figure 4.21), it is therefore possible to determine creep strain of the

Ø150x300mm creep specimens when measured strain on loaded specimens is known.

4.5.5 Creep Results: Creep.

The average strain readings taken from the creep samples of concrete grade 65 MPa,

50 MPa and 40 MPa along with the tested shrinkage results (as earlier presented in

183

Figure 4.21) and the resulting creep are presented in Figure 4.22. Due to the

constraints of limited duration of the project and the different frequencies of

collecting the test specimens, creep results were taken for approximately 650 days

after loading are available for all concrete grades. Thereafter, test data of a longer

period were available; approximately 1100 days for grades 65 and 50 MPa concrete.

It is noted that the magnitude value of creep for each concrete grade at a particular

time as plotted in Figure 4.22 was determined by subtracting the measured values of

shrinkage and elastic strain taken from the strain readings of the concrete samples in

each of the creep rigs, whereas the other curves are derived from the average values of

actual test data of the 150mmx300mm diameter cylinders.

As presented in Figure 4.22, the total strain, as well as creep and shrinkage increase

rapidly within the first 100 days after loading. During this period, increases in

shrinkage as well as creep all contribute to the increase in total measured strain.

Thereafter, shrinkage strains become stabilize and therefore only increases in creep

contribute to strain development. Although creep strains continue to increase with

time and at varying rates, the final value of creep was not reached at the termination

of the tests.

184

0

500

1000

1500

2000

2500

3000

3500

0 200 400 600 800 1000 1200


mic

rost

rain

Total Strain 65 MPa Creep Strain 65 MPa Shrinkage Strain 65 MPa



Figure 4.22. Average Measured Strain of Creep Specimens.

Of the three tested concrete grades, it is observed that average creep of the grade 50

MPa concrete is the highest, followed by the grades 40 MPa and 65 MPa concrete at

any given time. The measured creep at 600 days is 1420, 1750 and 1890 microstrains

for grades 65, 40 and 50 MPa concrete respectively. Subsequently, creep increases at

a relatively lower rate as noted by the test data of grades 65 and 50 MPa concrete. The

measured creep strain at 1050 days is 1530 and 1990 microstrain for grades 65 and 50

MPa concrete with an increase of 7.7 and 5.3 percent only in 450 days.

The two most common representations of creep results are presented in Figure 4.23

and Figure 4.24. These are the creep coefficient and specific creep of the three

concrete grades plotted against time. Coefficient or characteristic creep is the ratio of

creep to the initial elastic deformation. This expression of creep takes into account the

elasticity of aggregate which affects creep as well as the elastic modulus of concrete.

185

Specific creep or unit creep, on the other hand, is the value of creep per unit stress in

units of microstrain per MPa. This expression of creep takes into account the applied

load which affects creep.

0.0

0.4

0.8

1.2

1.6

2.0

2.4

2.8

0 200 400 600 800 1000 1200


65 MPa

50 MPa

40 MPa

Figure 4.23. Creep Coefficient for Q1 Tower Concrete.

With reference to Figure 4.23, it is noted that creep coefficient increases with time.

The creep coefficient of grade 40 MPa concrete is the highest at any given time, and

then followed by grades 50 MPa and 65 MPa concrete respectively. At 600 days, the

creep coefficient is 2.7, 2.2 and 1.7 for grades 40, 50 and 65 MPa concrete

respectively. Beyond 600 days, the creep coefficient increases at a relatively smaller

rate, as noted by the average of the grades 65 and 50 MPa concrete. The measured

creep coefficient at 1050 days is 2.3 and 1.8 for grades 50 and 65 MPa concrete

respectively, and increases by 4.5 and 5.9 percent only in 450 days.

It is noted that creep coefficient significantly depends on the age and grade of

concrete. Also lower strength concrete has relatively higher creep coefficient, and

thus, it tends to develop more creep strain than the initial elastic strain.

186

0

10

20

30

40

50

60

70

80

90

100

0 200 400 600 800 1000 1200


mic

rost

rain

/MP

a

65 MPa

50 MPa

40 MPa

Figure 4.24. Specific Creep of Q1 Tower Concrete.

Similarly, with reference to Figure 4.24, it is noted that specific creep increase with

time, in a similar manner as the creep coefficient. Specific creep of the grade 40 MPa

concrete is the highest at any time, followed by the 50 MPa and 65 MPa concrete

respectively. At 600 days, the measured specific creep is 91.9, 85.0 and 52.2

microstrain per MPa for grades 40, 50 and 65 MPa concrete respectively. The final

measured creep coefficient at 1050 days is 89.6 and 56.4 for 50 and 65 MPa concrete,

and increases by 5.4 and 8.0 percent only in 450 days. From these findings, it is noted

that lower strength concrete has higher sensitivity to creep when subjected to a similar

stress.

4.5.6 Test results as compared with predicted values

The measured creep results are compared with the predicted values obtained using

creep models of ACI 209R-92, AS 3600-2001, AS 3600-2009. Plots of predicted and

measured creep against time for concrete grades 65 MPa, 50 MPa and 40 MPa are

presented in Figure 4.25, Figure 4.26, and Figure 4.27 respectively.

187

0.0

0.4

0.8

1.2

1.6

2.0

2.4

2.8

0 200 400 600 800 1000 1200


Cre

ep C

oef

fici

ent

Measured ACI 209 AS 3600-2001 AS 3600-2009

Figure 4.25. Measured and predicted creep coefficient of concrete grade 65 MPa.

0.0

0.4

0.8

1.2

1.6

2.0

2.4

2.8

0 200 400 600 800 1000 1200


Cre

ep C

oef

fici

ent



188

0.0

0.4

0.8

1.2

1.6

2.0

2.4

2.8

0 200 400 600 800 1000 1200


Cre

ep C

oef

fici

ent



0.0

0.4

0.8

1.2

1.6

2.0

2.4

2.8

0.0 0.4 0.8 1.2 1.6 2.0 2.4 2.8

Creep Coefficient (Measured)

Cre

ep C

oef

fici

ent

(AC

I 209

)

65 MPa 50 MPa 40 MPa

Figure 4.28. Creep coefficient comparison with ACI 209.

189

0.0

0.4

0.8

1.2

1.6

2.0

2.4

2.8

0.0 0.4 0.8 1.2 1.6 2.0 2.4 2.8


Cre

ep C

oef

fici

ent

(AS

3600

-2001

)


Figure 4.29. Creep coefficient comparisons with AS 3600-2001.

0.0

0.4

0.8

1.2

1.6

2.0

2.4

2.8

0.0 0.4 0.8 1.2 1.6 2.0 2.4 2.8


Cre

ep

Co

eff

icie

nt

(AS

36

00

-20

09

)


Figure 4.30. Creep coefficient comparison with AS 3600-2009.

190

In order to arrive at the most accurate creep model for the test data, Pearson’s

correlation of the three creep prediction models is carried out and results are shown in

Table 4.9. In conclusion, ranking of these models based on Pearson’s correlation is as

follows; AS3600-2009 model gives the best creep prediction (R2=0.997), AS3600-

2001 model gives the second best (R2=0.995) and ACI 209 model is the least accurate

model for Q1 concrete (R2=0.981).

Table 4.9. Pearson correlation of various model predictions as compared to

corresponding measured values.

Concrete Grade/Code

ACI 209

(1992)

AS 3600

(2001)

AS

3600

(2009)

65 0.999 0.999 0.999

50 0.96 0.997 0.998

40 0.984 0.988 0.993

Average Value 0.981 0.995 0.997

Pearson Correlation


Properties of concrete specimens obtained from the field and tested in the laboratory

and prediction values computed by using various code models were presented. By

this means, the performance of major code model on prediction of compressive

strength, elastic modulus, creep and shrinkage is evaluated.

For compressive strength, it was found that concrete delivered to the Q1 Tower

always has higher 28-days strength than the grade that was specified for both grade 65

MPa and 50 MPa. This results in all model significantly underestimating compressive

strength, especially ACI209R-92 model. To correct this error, equation 4.7 has been

proposed for compressive strength prediction.

From the prediction value of elastic modulus compared to measured values, it is

observed that all models give estimation of elastic modulus with limited reliability

depending on the compressive strength. It is clear that none of the model under

191

investigation shows good prediction results through the range of compressive strength

from 40 to 90 MPa. A non-linear equation fit to the test result is therefore proposed

to eradicate this problem.

Shrinkage results of grade 65 MPa and 50 MPa concrete were reported. It was found

that shrinkage rerults of both concrete grades are not different. Shrinkage prediction

results show that AS3600-2001 gives accurate shrinkage prediction for the Q1 Tower

concrete while other models particularly AS3600-2009 model underestimated

shrinkage.

Test results of creep specimen are reported. This also includes results of control

specimens provided for the creep test, some of which were tested at 28-days for

compressive strength and elastic modulus, and the rest were kept in a control

environment for shrinkage readings that were conducted when creep data were taken.

The creep coefficient of grade 65 MPa, 50 MPa, and 40 MPa concrete were reported.

It was found that AS3600-2009 model gives accurate creep prediction.

In conclusion, it was found that there is no code model that gives best prediction for

each of all concrete properties being studied. To achieve the most precise estimation

possible, compatibility of the material models used must be verified with actual test

data. The model can be also easily derived by means of regression if a sufficient

amount of test data is available.

4.7 References



Mich., 1992, 47 pp.

Almudaiheem, Jamal A., and Hansen, W., “Effect of Specimen Size and Shape on

Drying Shrinkage of Concrete,” ACI Materials Journal, v 84, n 2, 1987, pp 130-135.

192

AUSTRALIAN STANDARD AS 1012.9., “Method for Determination of the

Compressive Strength of Concrete Specimens,” Standards Association of Australia,


AUSTRALIAN STANDARD AS 1012.16., “Method for the Determination of Creep

of Concrete Cylinders in Compression,” Standards Association of Australia, Sydney,

Australia, 1996, 9 pp.

AUSTRALIAN STANDARD AS 1012.17., “Method for the Determination of the

Static Chord Modulus of Elasticity and Poisson’s Ratio of Concrete Specimens,”


AUSTRALIAN STANDARD AS 1379., “Specification and supply of concrete,”











Neville, A.M., “Properties of Concrete- Fourth Edition.” Longman, Essex, 1995,

844pp.

193

5 Measured Axial Shortening of Q1 Columns and Walls

5.1 Introduction

In addition to the laboratory test results, the axial shortening results collected at the

Q1 Tower through field measurement results are reported in herein. Similarly, details

of the methodology for recording data readings on site and storing of such data during

the course of data collection is elaborated on. The method of converting DEMEC

gauge readings to cumulative strain over time is also outlined.

From initial plots of axial shortening against time, it is observed that, unlike plots of

laboratory concrete testing results, there are significant fluctuations of the axial

shortening-time curves in most cases. To verify this trend, a plot of axial shortening

against time, representing the typical axial shortening results obtained in this study

was compared with data from other buildings including the Water Tower Place

(Russell and Larson, 1989) and World Tower (Bursle, 2003 and 2006).

Finally, axial shortening results at every instrumented level are discussed. Due to the

complication of axial shortening models, comparisons of axial shortening results and

prediction values will be discussed at length in Section 5.5.

5.2 The data processing methodology

The data processing methodology outlines the process as used in calculating the axial

shortening values from the DEMEC gauge.

The data processing involves the following simplified steps;

1. Field documentation

2. Database documentation

3. Result computation

194

5.2.1 Field documentation

Data readings and other related information were recorded in field sheets when taking

measurements on site. The following were recorded therein: date of reading, ID of

data collection location, gauge values at DEMEC points on the instrumented

members, readings of the reference invar bar, ID of the DEMEC gauge operator, ID

of the DEMEC gauge and reference invar bar in operation, and in some cases, room

temperature and relative humidity. An example of a field data sheet showing recorded

DEMEC readings obtained at a column is presented in Appendix A.

5.2.2 Database documentation

At the close of each day of measurement, the field sheets were taken back to the

School of Engineering, Griffith University, Gold Coast Campus where data recorded

therein were then manually transferred into a database developed using Microsoft

Access™. Data readings from each instrumented member were stored, using an SQL

which was automatically updated when each subsequent data reading became

available. This therefore yields a reliable database containing real time updates of data

readings from all instrumented members. A Printscreen view of the database with

records of DEMEC data readings obtained from a column during the course of data

collection is presented in Appendix A.

5.2.3 Computation of Results

During the course of data collection, data readings stored in the database query were

imported into a Microsoft Excel™ spreadsheet customised to carry out the required

and necessary computations of the raw data. Output of the spreadsheet consists of

numerical figures and graphs representing axial shortening of the required member

during the monitoring period.

The following steps are involved in the spreadsheet analysis:

- Parameters: The columns/walls properties such as geometry, size and number

of reinforcement bars, grade of concrete, maximum design stress and the

195

number of DEMEC points/Gauge length are entered in the spreadsheet first.

Simultaneously, the volume to surface ratio of each column and the percentage

of reinforcement per sectional area are automatically computed with the

customized equations, therein.

- Imported Data: DEMEC gauge readings data are imported to the Microsoft

Access™ database (as described in Section 5.2.2). The imported data consist

of Invar bar reading, date of data reading and DEMEC reading data for each

DEMEC point.

- Data format: the imported data are organized into a more concise format, for

easy access and presentation.

- Data calibration: DEMEC data are calibrated using the Invar bar reading

data. The “true reading”, which is the calibrated DEMEC data is obtained by

subtracting the value of the Invar bar reading from the gauge reading. The

“true reading” is thus termed as it is independent of surrounding ambient

conditions.

- Strain calculation: For succeeding set of DEMEC readings, the increase in

strain is computed and is defined herein as “incremental strain”. The

cumulative sum of the incremental strains is therefore computed and is defined

herein as “cumulative strain”. Note that all computed strains are measured in

units of microstrains.

- Loading input: the details of loading and stress on members at each date of

DEMEC reading during the course of construction are manually input. Such

details include the number of floors built above the member and the estimated

applied load, in kilo-Newtons. The applied load is estimated through the

method outlined in Section 5.3. The corresponding sectional stress in the

member and its compressive strength is thereafter automatically computed.

Stress to strength ratio of each member at any given date during the course of

data collection is then determined. The real time compressive strength of the

member is computed.

- Axial shortening calculation: The cumulative axial shortening (in

millimeters) of each member is computed.

- Graph plots: The following five graphs are plotted using the above-

mentioned data; average strain of each DEMEC point (in microstrains) against

196

time (in days), average strain of the member (in microstrains) against time (in

days), average axial shortening of the member (in millimeters) against time (in

days), member stress (in Mpa) against age of member (in days) and stress to

strength ratio against age of member (in days).

An excerpt of the spreadsheet with values, and graphs of axial shortening of a column

computed from DEMEC reading data is provided in Appendix A.

5.3 Computation of Stress in Columns and Core Walls

5.3.1 Defining the typical construction sequences

The stress history of the tower was simulated to permit precise analysis of the field

data. The simulation process involved collation and interpretation of field records

prepared by the contractor such as concrete registry and weekly progress reports to

reconstruct the most realistic construction history. Finally, a typical scenario of the

incremental loading sequence used in replicating the actual occurrences during

construction was established.

Column loads were computed using the structure’s self-weight, construction load and

in-service load at the tributary area measured from as-built drawings.

The self weight includes gravity load of the column and its supporting slab whereas

the construction loads contain additional dead load and live load due to construction

activities, including heavy equipment and machineries. The in-service load is defined

as the additional load increment leading to the designed full occupancy and

serviceability load after the completion of the building.

The following assumptions have been made: Self weight and construction load during

construction equals 80% of the design load. Application of the in-service load, which

is equal to 20 % of the design load, would be effected 3 months after completion of

the structure, thus permitting an adequate transition period for load changes. The

additional dead load arising from finish furnishings plus service live loads applied

197

after the commissioning of the building gradually replaces the construction loads with

time. The simplified construction cycle used in simulating sequences of loading

application for the instrumented members is illustrated in Figure 5.1. The process is as

follows.

For the concerned column C1 at level 1, stage 1a represent the completion of casting

of C1. Formwork is assumed to be removed 3 days after casting and column self

weight, SWC1, is applied. The age of column and total column load are denoted as TC1

and WC1 respectively.

At stage 1b, the first set of props is put in place on S1 to provide full support for slab

S2. The column and slab at the floor above (C2 and S2) are then constructed. These

processes require TC2 days, and thus, the self weight of C2 is applied to C1 at

TC1=3+TC2 days.

Stage 1c is similar to stage 1b and begins TC3 days later. Assuming 3 sets of props

are used, the last set is installed on S3 at stage 1d, and SWC4 is applied at time

TC1=3+TC2+TC3+TC4 days.

Stage 2a begins when the props on S1 supporting S2 have been replaced with back-

props and 50% of construction loads applied. Props from S1 are then re-used, on S4,

as S5 and C5 are constructed. The total column loads now consist of self weight of C1

to C5 and 50% of construction loads from S2 to S5. Total elapsed time

TC1=3+TC2+TC3+TC4+TC5 days.

Stage 2b and 2c are similar to stage 2a. The incremental loading includes self weight

of the additional columns and 50% construction load of the next level slab. At the

end of stage 2c, back-props are placed on slabs S1, S2 and S3 whilst props are on

slabs S4, S5 and S6.

At stage 3, back-props on S1 are removed as C1 bears its self weight and full

construction load of columns and slabs above. With the construction of each new slab

198

and column, construction loads and column self weight are applied simultaneously.

This stage is repeated until the completion of the building.

C1

Stage 1a

TC1= 3

WC1=SWC1

C1

Stage 1b

TC1= 3+TC2

WC1=SWC1+SWC2

C2

C1

Stage 1c

C2

S1 S1

S2

S1

S2

C3

S3

C1

Stage 2a

C2

S1

S2

C5S5

C4

C1

Stage 2b

C2

S1

S2

C6S6

C5

C1

Stage 2c

C2

S1

S2

C7S7

C6

3

2

1 3i

CiC TT

5

2

1 3i

CiC TT

5

2

)(

5

1

12

1

i

sic

i

CiC WSWW

3

11

iCiC SWW

6

2

1 3i

CiC TT

6

2

)(

6

1

12

1

i

sic

i

CiC WSWW

7

2

1 3i

CiC TT

7

2

)(

7

1

12

1

i

sic

i

CiC WSWW

C1

Stage 3

C2

S1

S2

C8S8

C7

8

2

1 3i

CiC TT

8

2

)(

8

1

1

i

sic

i

CiC WSWW

C1

Stage 4

C2

S1

S2

S(j+1)

Cj

j

2

1 3i

CiC TT

J+1

2

)(

j

1

1

i

sic

i

CiC WSWW

C1

Stage 5

C2

S1

S2

S(j+1)

Cj

Ws

9032

1

j

i

CiC TT

1

2

)(

1

2

)(

1

1

j

i

sis

j

i

sic

j

i

CiC WWSWW

C1

Stage 1d

C2

S1

S2

C4

S4

4

2

1 3i

CiC TT

4

1

1

i

CiC SWW

C3

199

Figure 5.1. Simplified loading sequences for columns and core walls (adopted from

Beasley, 1987).

Stage 4 represents C1 loading at completion of the building where the top floor

column and the roof are denoted as Cj and S(j+1). Total column loads for C1

comprises of column self weight of C1 to Cj and construction load of slabs S2 to

S(j+1).

At stage 5, it is assumed that the in-service load, defined as Ws, is applied at once on

each slabs over the building at 90 days after stage 4. In other words, in-service loads

of S2 to S(j+1) are applied to C1 instantaneously. Thus, the Total column load for C1

comprises of the column self weights of C1 to Cj, and construction loads, plus in-

service load of slabs S2 to S(j+1). The active structure is assumed to carry its full

design loading at the completion of this stage.

5.3.2 Actual construction history

The construction history of Q1 Tower as recorded from the observations at the site is

given in Table 5.1. The construction of the super structure of Q1 Tower was

commenced on 25 January 2003 and completed on 4 October 2005. The construction

period per one floor varied from 4 to 44 days. The structural work was considered to

be completed (full dead load applied) on 4th

May 2005 upon the completion of level

78. However, it is assumed that the full live loads were not taken into the structure

until the grand opening in October 2005. This information was taken into account

when column stresses were computed.

From Table 5.1, it is clearly shown that the construction cycle per 1 floor of Q1

Tower varied during the course of the construction. Built cycle at early stage of the

construction is relatively slow however the target built cycle of 7 days per floor was

achieved after the 10th

level was built. Although, the fastest cycle achived was 4 days

per floor, construction was delayed at some level. After the 78th

level was built, the

200

interior and architectural works were completed and the opening ceremony

commenced 153 days later.

Table 5.1. Actual construction history of Q1 Tower

LevelPour Date

of Slabs

Built

Cycle

(days)

LevelPour Date

of Slabs

Built

Cycle

(days)

LevelPour Date

of Slabs

Built

Cycle

(days)

B2 25/1/2003 25 27 9/2/2004 8 55 24/8/2004 8

B1 19/2/2003 27 28 17/2/2004 8 56 1/9/2004 6

1 18/3/2003 27 29 25/2/2004 5 57 7/9/2004 7

2 14/4/2003 36 30 1/3/2004 9 58 14/9/2004 6

3 20/5/2003 34 31 10/3/2004 7 59 20/9/2004 4

4 23/6/2003 21 32 17/3/2004 7 60 24/9/2004 8

5 14/7/2003 11 33 24/3/2004 6 61 2/10/2004 5

6 25/7/2003 13 34 30/3/2004 4 62 7/10/2004 9

7 7/8/2003 9 35 3/4/2004 5 63 16/10/2004 7

8 16/8/2003 44 36 8/4/2004 15 64 23/10/2004 7

9 29/9/2003 12 37 23/4/2004 6 65 30/10/2004 9

10 11/10/2003 11 38 29/4/2004 15 66 8/11/2004 4

11 22/10/2003 6 39 14/5/2004 12 67 12/11/2004 5

12 28/10/2003 6 40 26/5/2004 5 68 17/11/2004 7

13 3/11/2003 5 41 31/5/2004 4 69 24/11/2004 6

14 8/11/2003 5 42 4/6/2004 13 70 30/11/2004 11

15 13/11/2003 7 43 17/6/2004 5 71 11/12/2004 5

16 20/11/2003 5 44 22/6/2004 6 72 16/12/2004 29

17 25/11/2003 6 45 28/6/2004 7 73 14/1/2005 7

18 1/12/2003 8 46 5/7/2004 4 74 21/1/2005 14

19 9/12/2003 6 47 9/7/2004 7 75 4/2/2005 28

20 15/12/2003 4 48 16/7/2004 5 76 4/3/2005 31

21 19/12/2003 22 49 21/7/2004 5 77 4/4/2005 30

22 10/1/2004 5 50 26/7/2004 4 78 4/5/2005 153

23 15/1/2004 6 51 30/7/2004 7 End 4/10/2005

24 21/1/2004 7 52 6/8/2004 6

25 28/1/2004 7 53 12/8/2004 7

26 4/2/2004 5 54 19/8/2004 5

201

5.3.3 Stress computation

Given the construction history of the building (see Table 5.1) the increases of column

stress during the time of back prop removals and the completion of the building (stage

3 to 4 as defined in 5.3.1) was computed in accordance with the flow diagram

algorithm presented and illustrated in Figure 5.2.

START

STOP

i, j

5. Loading input for column segment

at i storey in j storey building.

5a. Level no. “i”.

5b. Self weight (N).

5c. Construction load (N).

5d. Occupancy load (N).

Is

i = j

YES

i = i+1 NO

6. Construction history for column

segment at I storey in j storey

building.

6a. Construction cycle (days).

Figure 5.2. Stress computation procedures (stage 3-4).

202

5.4 Axial Shortening Results

5.4.1 Results and Discussion

The full set of axial shortening time curves generated from all the available data is

presented in Appendix A. Due to the vast number of data and graphs, only a selected

plot of axial shortening strain test results, representing the typical characteristics of

results, is presented in Figure 5.3 for further discussion. Figure 5.3 represents results

obtained from column TC06-B2 on second basement level, in which nine pairs of

DEMEC points were instrumented and thus considered as one of the most

comprehensive and consistent axial shortening data sets.

Figure 5.3. A typical plot of measured strains at a column of Q1 Tower (TC06-B2

data).

203

In general, test results indicate that axial shortening increases with an increase in time

as expected, for all gauge positions. However, negative values of data were noted at

early ages and this may be attributed to the effect of thermal expansion, caused by

hydration heat.

The variations of axial strain at different gauge positions within a single column are

clearly observed. For instance, axial shortening of Series 6 is approximately double

that of Series 8 and Series 9. This behavior is considered typical using this kind of

measurement system because of variations in concrete properties of the column

surface after concrete placing. Furthermore the use of DEMEC gauges allow for strain

readings to be taken at the column surface only, where the sensitivity of readings may

be affected by unknown internal voids caused by segregation during concreting.

Further observation identifies that axial shortening measured from a lower level of the

same column is more likely to be of a higher magnitude. For example, the axial

shortening (at 400 days) of series 3, 2, 1 is approximately 400, 300, 170 microstrain

respectively. It was also observed that the values of measured axial shortening would

sometimes fluctuate over a short period. This could be attributed to the exposed

condition of the concrete members as they are subjected to daily variations of

seasonal temperatures and humidity.

In order to capture such variations, more than one set of DEMEC points were used on

a single column where applicable. A representative of each column was therefore

utilized by obtaining the mean of results from all measurement stations for further

analysis. The average axial shortening-time from values of Figure 5.3 is plotted, and

four straight lines fitting the data between the limits of 0-100, 100-400, 400-600, 600-

900 days respectively (see Figure 5.4) are obtained.

It can be seen that early AXS developments (0-100 days) can be expressed as a linear

time dependent function. As referred to construction history of the Q1 Tower (Table

5.1), shrinkage results (Section 4.4) and creep results (Section 4.5), it is found that

shrinkage has significant effect to AXS at this stage since the column has not been

204

subjected to high level of axial stress (note that only 3 out of 80 floors were built at

100 days). Other kinds of strain i.e. elastic strain and creep strain were therefore still

small.

Figure 5.4. Average AXS of column TC06-B2 and linear fitted line at various

intervals

There was non-linearity in AXS-time relationship during 100-400 days. During this

period, shrinkage was stabilized as the shrinkage curve reaching its zenith thus its

effect to AXS is very little. It can be seen that AXS developments also slow down

despite the higher speed of construction (note that only 28 more floors were built

during this period, an average cycle of 11 days per floor). This is due to the relatively

smaller effect of elastic shortening as compared to effect of early shrinkage coupled

with sluggish development of early creep.

205

During the third stage (400-600 days), AXS-time relationship can be expressed in

linear terms again. The AXS development speeded up again but still not as fast as it

was in the first stage. The additional AXS occurred is due to the effect of faster

construction (now construction cycle reduced to around 7 days per floor) coupled with

greater creep development.

Finally after 600 days, the construction was reaching its end. The average

construction cycle was 10 days per floor and the last floor was built at 800 days

(referred to Table 5.1). The AXS development was relatively lower and the AXS-

time relationship took a non-linear form again. This is because creep development at

this stage slow down before reaching its final value while the elastic shortening also

increased at a slightly slower rate (according to the increasing construction cycle

time).

Note that the discussion given in this subsection is only for AXS results obtained from

column TC06/B2 which intended to represent general characteristic of AXS results

taken from other locations/floor levels. Full results of AXS taken at Q1 Tower are

presented in Appendix A. To downsize the number of AXS data, the average AXS

results of each column/wall is used for further study.

5.4.2 Measured axial shortening of columns and walls at level B2

- Long-term results, columns set B2/C1

Axial shortening readings of 4 columns located at the 2nd

basement level, B2, are

plotted in Figure 5.5. These columns are TC05, TC06, TC10, TC12 and their

locations are shown on the floor plan in Figure 5.5. Columns are chosen as: note that

columns TC05 and TC06 are identical but only located symmetrically on either side

of the structure. This is done to verify method as it is expected that shortening will be

symmetrical. The plotted data points represent the average of 9 DEMEC readings

taken from a single column at the given age. The corresponding stress of the 4

206

columns during the period of data collection is computed based on Section 5.3, and is

presented in Figure 5.6.

These columns are amongst the few instrumented locations which were accessed

since their early ages (axial shortening measurement commenced at the age of 9, 11,

12, and 46 days at TC06, TC12, TC10, and TC05 respectively) and data readings

were taken until the structure was completed. Therefore, axial shortening data

obtained from these columns are the most comprehensive. The axial shortening data

of these columns is denoted herein as Set B2/C1 data. Data for these columns

including the size and volume to surface ratio, duration of strain reading, age of

column when data reading commenced and terminated, as well as a summary of the

measured axial shortening and its corresponding stress at particulars ages are given in

Table 5.2.

Columns at Level B2

TC05

TC06

TC10

TC12

-100

0

100

200

300

400

500

600

700

800

0 200 400 600 800 1000

Age of Column (days)

Ax

ial

Sh

orte

nin

g (

mic

ro

stra

in)


Tower.

TC06

TC05

TC10

TC12

207

Columns at Level B2

0

2

4

6

8

10

12

14

16

0 200 400 600 800 1000


Str

ess

(M

Pa

)

TC05,06

TC10

TC12

Figure 5.6. Computed Stress, Level B2 Columns – Set B2/C1, Q1 Tower.

From Figure 5.5 and Figure 5.6, it is observed that axial shortening and stress increase

with time in all columns. Stress at TC05, TC06 and TC12 are general comparable

until the age of column reaches 700 days, when the stress at TC12 becomes constant.

Prior to 500 days, the stress at TC10 increases at a relatively higher rate than those of

TC05, TC06, and TC12 although it is relatively lower after 500 days. The

observations during the course of data collection of these columns are reported based

on three periods as follows:

Early age observation - In general, axial shortening increases rapidly during

the first 150 days of data reading despite the corresponding lower rate of stress

development. Inevitably, during this period, the gradient of the axial

shortening curves is higher than those of their corresponding stress curves. It

can therefore be implied that the early shortening is essentially due to drying

shrinkage given that the subjected stress is minimal.

Mid period observation - From 150 days onward, the gradients of the axial

shortening curves become less steep and tend to approximately line up with

the stress curve. This therefore implies that the effect of drying shrinkage

208

reduces (refer to shrinkage test results reported in Sections 4.4 and 4.5.4)

whereas the columns are subjected to gradually increased stresses arising from

the additional completed floors above. Given that instantaneous shortening

occurs immediately, an additional stress is applied on the column, axial

shortening of columns increases simultaneously with increase in stress.

Although creep due to the sustained stress on the columns has developed, its

effect is not as significant as the instantaneous shortening during this period of

observation. This is because the structure is yet to be fully loaded and the

creep fully developed (refer to creep results reported earlier in Section 4.5.5).

Final period observation - It is observed that the rate of axial shortening

changes again from 700 days onward, as soon as the structure is completed

and the stress remains constant. The rate of axial shortening varies due to the

creep of concrete at the structure’s full dead load. It is observed from Figure

5.5 and Figure 5.6 that axial shortening continues to increase, regardless of the

constant state of stress

Considering the age at which readings commenced, as shown in Figure 5.5 only axial

shortening results at columns TC06, TC10 and TC12 may be considered to deduce

more accurate conclusions. This is because the data reading at column TC05

commenced at a later age (46 days). If column TC05 is to be included in the

comparison of axial shortening at column set B2/C1, then its values for 9 to 46 days

must be determined and included in the measured results.

From Figure 5.5 and Figure 5.6 and the summary in Table 5.2, it is evident that, at

893 days, TC06 has the highest magnitudes of axial strain and stress equal 733

microstrains and 13.4 MPa respectively, TC12 ranks second with axial strain and

stress values equal to 642 microstrains and 11.8 MPa respectively. TC10 ranks lowest

with axial strain and stress values equal to 606 microstrains and 10.7 MPa

respectively.

However, at 100 days, the stress in all 3 columns is similar with TC12 exhibiting the

highest axial strain of 348 microstrains whereas TC06 and TC10 have similar

magnitudes of 236 and 231 microstrains respectively. Note that TC12 has the lowest

209

volume to surface ratio of 320. This observation verifies the effect of shrinkage,

which is more significant during the earlier age of the column.

Table 5.2. Summary of Set B2/C1 results.

when

reading

started

when

reading

ended

at 100

days

at 600

days

at 893

days

TC05 Ø2000/500 2.4 924 46 970 44/0.33 396/10.4 485/13.4

TC06 Ø2000/500 2.4 909 9 918 236/0.33 644/10.4 733/13.4

TC10 4000x900/367 2.7 906 12 918 231/0.31 534/9.0 606/10.7

TC12 3200x800/320 2.3 882 11 893 348/0.32 600/10.7 642/11.8

Age of column (days) AXS/Stress (microstrain/MPa)

Note: Columns/Walls height = 2250mm

ColumnSize (mm)/

V/S ratio

Rebar

Area

(%)

DEMEC

reading

duration

(day)

- Short-term results, columns set B2/C2

Axial shortening of the remaining columns located in the 2nd

basement level, B2, is

plotted in Figure 5.7 to Figure 5.9. Axial shortening of TC01, TC03, and TC04 is

presented in Figure 5.7. Axial shortening of TC07, TC08, and TC09 is presented in

Figure 5.8. Axial shortening of TC11, TC13, and TC14 is presented in Figure 5.9. The

axial shortening data of these columns is denoted herein as Set B2/C2 data.

Note that columns TC03 and TC04; TC07 and TC08; TC13 and TC14 are identical

columns which they are located opposite each other. Once again, the plotted data

points represent an average value of 9 DEMEC readings taken from a column at the

given age. The corresponding stresses of these columns, during the period of data

collection, is also computed (based on Section 5.3), and is presented in Figure 5.10.

It is worth noting that axial shortening at B2/C2 columns were measured at later ages

of 362 to 371 days and continued until completion of the structure. Although axial

shortening data of B2/C2 columns is not available at ages less than 362 days, the data

is considered comprehensive due to the long-term period of measurement (from 504

to 608 days), which covers the mid-period and final-periods. Details including the

size, and volume to surface ratio, duration of strain reading, age of column when data

reading commenced and terminated, as well as a summary of the measured axial

shortening and its corresponding stress at particular ages are given in Table 5.3.

210

Following the same trend as B2/C1 set, it is observed from Figure 5.7 to Figure 5.10,

and the summary in Table 5.3 that axial shortening and stress increases with time in

all columns. Stresses at all columns are generally similar except at TC13 and TC14, in

which the stresses are obviously greater than other columns. Column TC01, which is

the smallest circular column, has the maximum axial shortening. Thereafter, columns

TC13 and TC14 rank second in higher values of axial shortening, as they are subject

to the highest stress. All other columns lie within the same range of axial shortening

values. Axial shortening of TC09 is the lowest.


B2/C2 (TC01, TC03, TC04), Q1 Tower.

211



212



Columns at Level B2 TC13,14

0

2

4

6

8

10

12

14

16

18

0 200 400 600 800 1000


Str

ess

(M

Pa

)

TC01,02

TC03,04

TC07,08

TC09

TC11

TC13,14


Tower.

At any age, TC01 has the highest axial shortening values of 117, 382, 529

Microstrains at 404, 602, 893 days respectively, regardless of the stress which is

relatively lower than those of TC13 and TC14 columns. Note that TC01 has

approximately the same volume to surface ratio as TC13 and TC14 (338 and 333,

respectively), although the shape is circular. From this observation, it can be implied

that circular columns are more sensitive to axial shortening than rectangular columns

with the same stress and volume to surface ratio.

213


when

reading

started

when

reading

ended

at 404

days

at 602

days

at 893

days

TC01 Ø1350/338 3.0 555 362 917 117/5.21 382/9.2 529/11.6

TC03 Ø2200/550 2.4 608 362 970 20/6.4 271/9.0 394/11.0

TC04 Ø2200/550 2.4 555 363 918 54/6.4 329/9.0 418/11.0

TC07 3400x1000/386 3.1 600 370 970 5/6.82 260/8.9 340/11.4

TC08 3400x1000/386 3.1 523 371 894 2/6.82 234/8.9 392/11.4

TC09 4000x900/367 2.7 504 370 874 5/6.56 221/9.5 298/10.7

TC11 3200x800/320 2.3 504 370 874 10/5.9 250/10.7 342/11.8

TC13 4000x800/333 5.3 600 370 970 77/9.4 355/14.1 403/15.3

TC14 4000x800/333 5.3 547 370 917 38/9.4 313/14.1 369/15.3

Note: Columns/Walls height = 2250mm

ColumnSize (mm)/

V/S ratio

Rebar

Area

(%)

DEMEC

reading

duration

(day)


- Long-term and short-term results, core walls set B2/W1 and C2/W2

The mean axial shortening data of the instrumented core walls located at the 2nd

basement level, B2, is plotted in Figure 5.11. The plotted data points represent the

average values of up to 9 DEMEC readings taken from the core walls at the given

age. The designations and locations of these walls i.e. TW06, TW07, TW09, TW16,

TW17, TW18 are shown on the floor plan inset also in Figure 5.11. It should be noted

that early axial shortening readings were taken at TW09 (measurements commenced

at the age of 15 days) only. Axial shortening data of the remaining core walls were

recorded approximately after 365 days. Therefore, axial shortening data of TW09 is

defined as set B2/W1 and those of the remaining walls are defined as set B2/W2.

It can be seen that long-term measurement only taken at TW09 when the age of core

walls was 15 to 917 days, comparable to those of set B2/C1 columns particularly

TC06, TC10 and TC12. The measurement at the remaining core walls (set B2/W2)

was taken during the age of 369 to 853 days, approximately on par with the

measurement time of set B2/C2 columns.

214

The corresponding stresses at these columns during the period of data collection is

computed based on the method outlined in Section 5.3 and is presented in Figure 5.12.

Core Walls at Level B2

-100

0

100

200

300

400

500

600

700

800

0 200 400 600 800 1000

Age (days)

Axia

l S

ho

rten

ing

(d

ays)

TW06 TW07 TW17

TW16 TW18 TW09

Figure 5.11. Mean Axial Shortening Results, Level B2 Core Walls – Set B2/W1 + Set

B2/W2, Q1 Tower.

TW09

TW06

TW16

TW17 TW18 TW07

215

Figure 5.12. Computed Stress, Level B2 Core Walls – Set B2/W1+ Set B2/W2, Q1

Tower.

Table 5.4. Summary of Set B2/W1 and B2/W2 results.

when

reading

started

when

reading

ended

at 175

days

at 404

days

at 628

days

at 805

days

TW09* 600/187.5 902 15 917 298/4.32 320/9.4 520/12.0 610/12.91

TW06** 400/142.9 476 377 853 N/A 33/4.8 180/7.4 217/9.24

TW07** 400/142.9 436 369 805 N/A 3.3/4.8 196/7.4 196/9.24

TW16** 300/115.4 454 369 823 N/A 13.3/6.4 213/9.9 223/12.31

TW17** 300/115.4 436 370 806 N/A -3.3/6.4 250/9.9 266/12.31

TW18** 350/129.6 436 370 806 N/A 0/6.34 235/11.5 246/13.9

Remark * Set B2/W1, ** Set B2/W2

AXS/Stress (microstrain/MPa)

Core

wall

Thickness

(mm)/ V/S

ratio

DEMEC

reading

duration

(day)

Age of core wall (days)

- Differential shortening between columns and core walls

In order to determined the differential shortening occurred between columns and core

walls, axial shortening data of TW09 is plotted against those of TC06, TC10 and

TC12 in Figure 5.13 and axial shortening data of core walls set B2/C2 are plotted

against those of columns set B2/C2 in Figure 5.14 respectively.

216

TW09

-100

0

100

200

300

400

500

600

700

800

0 200 400 600 800 1000


Ax

ial

Sh

orte

nin

g (

mic

ro

stra

in)

TC06

TC10

TC12

TW09

Figure 5.13. Axial shortening of TW09 plotted against TC06, TC10 and TC12.

TW09

TC06 TC12

TC10

217

-100

0

100

200

300

400

500

600

300 400 500 600 700 800 900 1000


Ax

ial

Sh

orte

nin

g (

mic

ro

stra

in)

TC01 TC03 TC04 TC07 TC08

TC09 TC11 TC13 TC14 TW06

TW07 TW17 TW16 TW18


columns.

From visual inspection of data in both Figure 5.13 and Figure 5.14, it was found that

core walls tend to have less axial shortening as compared to columns despite a similar

magnitude of the applied stress. This difference is due to the relatively more solid

geometry of core walls which virtually takes the shape as a hollow box with a massive

perimeter as compared to perimeter of columns.

In general, differential shortening between one column and another column as well as

between one column and one core wall was found not visibly found at very early ages

(less than 50 days), however, the rate of differential increased during 50-150 days and

if not slightly increased, it remained approximately constant beyond 150 days until

the end of the data collection. The worst case differential shortening between TW09

and the three columns was TW09 against TC06 (approximately 150 microstrain at

950 days).

218

However, as seen in Figure 5.13, the axial shortening of TW09 gives particularly

good agreement at all given times with the axial shortening data of column TC10

which is its direct opposite perimeter column. This implied that the differential

shortening between TW09 and TC10 occurred at B2 level is obviously at minimum

and it should not affect the integrity of the structure.

The comparison of axial shortening data of core walls set B2/W2 against columns set

B2/C2, however, reveal greater chance of differential shortening as the worst case

being the differential shortening between axial shortening at TC01, the upper limit of

the set B2/C2, against axial shortening at TW07, the lower limit of the set B2/W2,

was approximately equal to 300 microstrain different at 800 days.

The axial shortening of columns and core walls and the absolute value of differential

shortening between columns and adjacent core walls or columns is summarized in

Table 5.5. The case comparisons no.1 to no.4 was conducted using data set B2/C1

and B2/W1 and they are listed as follows;

- Case no.1 represents the worst case differential shortening between adjacent

columns in set B2/C1. This occurred between TC06 and TC10 and the

maximum magnitude was up to 110 microstrain at 603 days.

- Case no.2 represents worst case differential shortening between columns in set

B2/C1 and core wall TW09. This occurred between TC06 and TW09 and the

maximum magnitude was up to 124 microstrain at 603 days.

- Case no.3 represents a typical case of differential shortening between a

column of set B2/C1 and its adjacent column. The selected case was between

TC06 and TC12 which the maximum magnitude of differential shortening was

up to 117 microstrain at 106 days.


column of set B2/C1 and its adjacent core wall. The selected case was between

TC10 and TW09 which the magnitude of differential shortening at any given

times was virtually negligible.

219

The case comparisons no.5 to no.8 was conducted using data set B2/C2 and B2/W2

and they are listed as follows;

- Case no.5 represents worst case differential shortening between adjacent

columns set B2/C2. This occurred between TC01 and TC09 and the maximum

magnitude was up to 255 microstrain at 840 days.

- Case no.6 represents worst case differential shortening between core walls set

B2/W2. This occurred between TW07 and TW17 and the maximum


- Case no.7 represents a worst case differential shortening between a column of

set B2/C2 and core walls set B2/W2. The selected case was between TC01 and

TW07 which the maximum magnitude of differential shortening was up to 302

microstrain at 804 days.


column of set B2/C2 and its adjacent column. The selected case was between

TC09 and TC11 which the magnitude of differential shortening at any given

times was relatively low.


level B2.

at106

days

at 405

days

at 603

days

at 840

days

1 worst case B2/C1-B2/C1 TC06-TC10 6 96 110 106

2 worst case B2/C1-B2/W1 TC06-TW09 5 86 124 99

3 Adjacent members B2/C1-B2/C1 TC10-TC12 117 93 77 73

4 Adjacent members B2/C1-B2/W1 TC10-TW09 1 10 14 7

5 worst case B2/C2-B2/C2 TC01-TC09 N/A 112 160 255

6 worst case B2/W2-B2/W2 TW07-TW17 N/A 6 76 46

7 worst case B2/C2-B2/W2 TC01-TW07 N/A 114 185 302

8 Adjacent members B2/C2-B2/C2 TC09-TC11 N/A 5 13 36

Case

no.Remark/comparing set

Comparing

Members

Differential shortening (microstrain)

220

It was found that axial shortening of core walls at level B2 tends to be equal across

the floor plan and is relatively less than those of columns. Given example differential

shortening between core walls is very small whilst differential shortening between

columns or between columns and core walls can be equally greater.

For a floor height of 2250mm at level B2, the maximum differential occurred was

approximately 300 microstrain which is equivalent to 0.675mm or h/3333, far lower

than the allowable limit of h/500 set by the structural engineer of Q1 Tower (Sunland

Group Ltd, 2002). Therefore, it is evident that differential shortening in this building

is well designed for; its existence is minimal, and should not have any effects on

performance of the structure.

5.4.3 Measured axial shortening of columns and walls at level B1

- Long-term results, columns set B1/C1

Similarly to the axial shortening data previously presented in 5.4.2, the average long-

term axial shortening data of 2 columns were taken at the 1st basement level, B1,

namely TC10 and TC12. The axial shortening readings were taken at early age until

approximately 850 days is presented in Figure 5.15. The axial shortening data of

TC10 and TC12 at level B1 is denoted herein as Set B1/C1 data. The corresponding

stress of columns TC10 and TC12 during the period of data collection is computed

based on the method given in section 5.3 and is presented hereafter in Figure 5.16.

221

Columns at Level B1

TC10

TC12

0

100

200

300

400

500

600

0 200 400 600 800 1000


Ax

ial

Sh

orte

nin

g (

mic

ro

stra

in)


Tower.

Columns at Level B1

0

2

4

6

8

10

12

14

16

18

0 200 400 600 800 1000


Str

ess

(M

Pa

)

TC09,10

TC11,12

Figure 5.16. Computed Stress, Level B1 Columns – Set B1/C1, Q1 Tower.

TC10 TC12

222

Details of column TC10 and TC12 including the size and volume to surface ratio,

duration of strain reading, age of column when data reading commenced and

terminated, as well as a summary of measured axial shortening and its corresponding

stress at particular ages of columns, is given in Table 5.6.


when

reading

started

when

reading

ended

at 115

days

at 430

days

at 848

days

TC10 4000x900/367.4 2.7 841 7 848 176/0.3 273/6.9 449/10.6

TC12 3200x800/320.0 2.3 884 7 891 163/0.7 287/6.9 498/11.6


Note: Columns height = 3850 mm

ColumnSize (mm)/ V/S

ratio

Rebar

Area

(%)

DEMEC

reading

duration

(day)

- Short-term results, columns set B1/C2

Average axial shortening of the remaining instrumented columns located on the 1st

basement level, B1, denoted herein as Set B2/C2 data, is presented in Figure 5.17 to

Figure 5.19. Similarly to the instrumented location of the column set B2/C2, columns

of the set B1/C2 also includes TC01, TC03, TC04, TC07, TC08, TC09, TC11, TC13,

TC14.

It is noted that B1/C2 columns were also instrumented at later ages in addition to

B1/C1 columns (axial shortening measurement commenced at the age of 330 to 351

days, slightly earlier than the same columns at level B2) and the data readings were

also taken until the structure was completed. Although, axial shortening data of B1/C2

columns is not available for age of columns less than 330 days but the data is

considered to be comprehensive due to the long-term data reading duration (from 555

to 614 days, in exception of TC01 where the data reading ceased after a duration of

280 days) which cover most timing of mid-period and final-period observations.

Based on the method given in section 5.3, the corresponding stress of B1/C2 columns

during the period of data collection were computed and presented in Figure 5.20.

223



224





Columns at Level B1TC13,14

0

2

4

6

8

10

12

14

16

18

0 200 400 600 800 1000


Str

ess

(M

Pa

)

TC01,02

TC03,04

TC07,08

TC09

TC11

TC13,14

225


Tower.

Column details of B1/C2 including the size and volume to surface ratio, duration of

strain reading, age of column when data reading commenced and terminated, as well

as a summary of measured axial shortening and its corresponding stress at particular

ages of columns, is given in Table 5.7.


when

reading

started

when

reading

ended

at 399

days

at 610

days

at 891

days

TC01 Ø1350/337.5 3.0 280 330 610 47/5.3 281/9.6 N/A

TC03 Ø2200/550.0 2.4 614 330 944 60/6.5 159/9.3 314/10.8

TC04 Ø2200/550.0 2.4 561 330 891 18/6.5 163/9.3 308/10.8

TC07 3400x1000/386.4 3.1 608 336 944 78/6.9 260/9.2 382/11.2

TC08 3400x1000/386.4 3.1 561 330 891 42/6.9 178/9.2 271/11.2

TC09 4000x900/367.4 2.7 608 336 944 55/6.6 215/9.6 316/10.6

TC11 3200x800/320.0 2.3 608 336 944 77/6.1 288/10.9 343/11.6

TC13 4000x800/333.3 4.2 593 351 944 55/8.8 248/13.4 301/14.4

TC14 4000x800/333.3 4.2 555 336 891 68/8.8 200/13.4 282/14.4


Note: Columns/Walls height = 3850 mm

ColumnSize (mm)/ V/S

ratio

Rebar

Area

(%)

DEMEC

reading

duration

(day)

- Short-term results, core walls set B1/W2

At level B1, no data reading was taken at core walls area during the early age.

However, axial shortening data reading at seven core walls locations were taken

during the approximate ages of 336 days to 891 days, which was approximately the

same duration as core walls set B2/W2. The instrumented core walls at basement level

B1 were TW06, TW07, TW08, TW16, TW17 and TW18. Average axial shortening

data of these core walls, designated as data set B1/W2 (account for the relatively short

term data), are presented in Figure 5.21 and Figure 5.22.

The corresponding stress of these columns during the period of data collection is also

computed based on the method given in section 5.3 and is presented in Figure 5.23.

Details of core walls set B1/W2 including the size and volume to surface ratio,

duration of strain reading, age of column when data reading commenced and

226

terminated, as well as a summary of measured axial shortening and its corresponding

stress at particulars ages of columns, is given in Table 5.8


(TW06, TW07, TW08), Q1 Tower.

227


(TW10, TW16, TW17, TW18), Q1 Tower

Figure 5.23. Computed Stress, Level B1 Core Walls – Set B1/W1, Q1 Tower.

Table 5.8. Summary of Set B1/W2 results.

228

when

reading

started

when

reading

ended

at 399

days

at 610

days

at 848

days

at 891

days

TW06 400/285.7 1.4 512 336 848 80/4.5 220/7.1 354/8.8 N/A

TW07 400/285.7 1.5 259 351 610 198/4.5 267/7.1 N/A N/A

TW08 400/285.7 1.5 670 344 867 140/4.5 243/7.1 386/8.8 N/A

TW10 300/230.8 1.4 555 336 891 17/4.9 207/6.9 347/8.4 332/8.4

TW16 300/230.8 1.1 547 344 891 30/6.1 210/9.6 355/11.8 361/11.8

TW17 300/230.8 2.7 547 344 891 20/6.1 210/9.6 341/11.8 343/11.8

TW18 350/259.3 0.9 555 336 891 20/5.9 190/11.1 259/13.2 253/13.2

Age of core wall (days) AXS/Stress (microstrain/MPa)

Note: Core walls height = 3850 mm

Core

wall

Thickness

(mm)/ V/S

ratio

Rebar

Area

(%)

DEMEC

reading

duration

(day)

- Differential shortening between columns and core walls

The axial shortening results of TC10 and TC12 at level B2 and B1 are previously

presented in Figure 5.5 and Figure 5.15 respectively where minimal differential

shortening was observed. It was seen that differential shortening between TC10 and

TC12 at level B1 was approximately between 50 to 80 microstrain only which was

relatively smaller than differential shortening between TC10 and TC12 at level B2.

In order to determine the differential shortening occurring between columns and core

walls during age 336 and 891 days, axial shortening data of columns set B1/C2 and

core walls set B1/W2 are plotted together in Figure 5.24.

229

-100

0

100

200

300

400

500

600

300 400 500 600 700 800 900 1000


Ax

ial

Sh

orte

nin

g (

mic

ro

stra

in)

TC01 TC03 TC04 TC07 TC08 TC09

TC11 TC13 TC14 TW06 TW07 TW08

TW10 TW16 TW17 TW18


columns.

From visual inspection of data presented in Figure 5.24, it was found that axial

shortening of both B1/C2 and B1/W2 fall within a similar range. The measured axial

shortening at 336 days and 891 days was 0±50 microstrains and 300±50 microstrains

respectively. At level B1, it should be noted that there is no evidence that core walls

tend to have less axial shortening as compared to columns despite the similar

magnitude of the applied stress as observed earlier in level B2.

It can be seen that differential shortening between one column and another column as

well as between one column and one core wall was minimal at all ages. The rate of

differential shortening slightly randomly fluctuated over the course of the readings,

common pattern observed in this study. The worst case differential shortening

amongst these columns and core walls during the data readings was approximately

150 microstrain, 50% less than those amongst B2/C2 and B2/W2 at level B2.

230

The axial shortening of columns and core walls and the absolute value of differential

shortening between columns and adjacent core walls or columns at level B1 is

summarized in Table 5.9. The five case comparisons are listed as follows;

- Case no.1 represents differential shortening between adjacent columns set

B1/C1. This was between TC10 and TC12 and the maximum magnitude was

approximately 73 microstrain at 603 days.

- Case no.2 represents the worst case differential shortening between columns

of set B1/C2. This occurred between TC07 and TC08 and the maximum


- Case no.3 represents the worst case differential shortening between core walls

of set B1/W2. This was between TW16 and TW18 which the maximum

magnitude of differential shortening was up to 96 microstrain at 804 days.

- Case no.4 represents the worst case of differential shortening between a

column of set B2/C1 and a core wall of set B1/W2. This was between TC07

and TW18 which the maximum magnitude of differential shortening was up to

128 microstrain at 804 days

- Case no.5 represents a typical case of differential shortening between a core

wall of set B1/W2 and its adjacent columns. The case represented was

between TC14 and TW10 and the maximum magnitude was up to 97

microstrain at 840 days.


level B1.

at106

days

at 405

days

at 603

days

at 840

days

1 Adjacent members B1/C1-B1/C1 TC10-TC12 13 46 73 49

2 worst case B1/C2-B1/C2 TC07-TC08 N/A 58 64 111

3 worst case B1/W2-B1/W2 TW16-TW18 N/A 10 37 96

4 worst case B1/C2-B1/W2 TC07-TW18 N/A 58 70 128

5 Adjacent members B1/C2-B1/W2 TC14-TW10 N/A 41 17 97

Case

no.Remark/comparing set

Comparing

Members

Differential shortening (microstrain)

Similarly to the behavior observed at level B2, axial shortening of core walls at level

B1 tends to be equal across the floor plan. However, the axial shortening of core walls

231

is equal to those of columns. There is no evidence suggesting that differential

shortening between columns and core walls could be equally greater than those

occurred between columns which was observed at level B2.

For a floor height of 3850mm at level B1, the maximum differential occurred was

approximately 150 microstrain which is equivalent to 0.493mm or h/7809, far lower

than the allowable limit of h/500. Therefore, it is evident that differential shortening at

level B1 is in the allowable limit.

5.4.4 Measured axial shortening of columns and walls at level L2, L31, L49 and

L63

In addition to the instrumented locations at basement levels B2 and B1, a vast number

of columns and core walls located at other levels i.e. L2 (transfer floors), L31; L49;

L63 (typical floors) were also instrumented, although only short-term data was

available due to limited access.

Mean axial shortening data of these columns and core walls are presented in Figure

5.25 to Figure 5.31 respectively. A 600mm DEMEC gauge was used for axial

shortening measurements at levels L31 and L49 whereas a 200mm DEMEC gauge

was used at other levels.

232

Figure 5.25. Mean Axial Shortening Results, Level L2 Columns, Q1 Tower.


TC05), Q1 Tower.

233


TC08, TC14), Q1 Tower.

Figure 5.28. Mean Axial Shortening Results, Level L31 Columns and a core wall

(TC07, TC13, TW18), Q1 Tower.

234


TC07, TC08, TC09), Q1 Tower.

235


TC13, TC14), Q1 Tower.

Figure 5.31. Mean Axial Shortening Results, Level L63 Columns, Q1 Tower.

236

Figure 5.32. Mean Axial Shortening Results, Level L63 Core Walls, Q1 Tower.

Details of the instrumented columns and walls at levels L2, L31, L49 and L63

including the size, volume to surface ratio, duration of strain reading, age of column

when data reading commenced and terminated, as well as a summary of the measured

axial shortening and its corresponding stresses at particular ages are listed in Table

5.10.

The strain measurements were taken at these upper levels earliest at the age of 20

days (TC08/L31) and the latest, at the age of 127 days (TW08/L2). The measurements

at most of these columns and walls started at later ages than at those of Level B2

Columns – Set B2/C1. This is due to the actual progress of construction during which

the formwork consisting of slip-form and jump-form which facilitates construction of

columns and walls at level L2 and above, 3 levels in advance of the floor slab. Access

to these columns and walls was therefore not available at their early ages whereas the

basement levels did not have this limitation as floor slabs were cast simultaneously

with columns and walls.

237

Therefore, the data collection at these levels ended prematurely as compared to strain

readings taken at basement levels. The duration of strain readings was from 49 to 159

days only. The loss of access was due to the interior works which required removal of

DEMEC points from the stations. After the completion of the interior work, the

completed apartment rooms were not to be entered, thus the experiment had to be

ceased.

From the short-term results measured at these levels, it was found that axial

shortening at most stations has similar magnitude at approximately 200±80

microstrains at the age of 180 days. The precision of the measurements at level L49

compared against all other levels did not show any superiority although the 600-mm

DEMEC gauges were used. It can be concluded that it is more suitable to use 200 mm

DEMEC gauge for measuring axial shortening of columns and walls on the site which

was larger than 600 mm DEMEC gauge. This is because 200 mm DEMEC gauge is

lighter, easier to be used and actually yields results with the same level of accuracy as

the 600 mm DEMEC regardless of the factory specification claiming that the 600 mm

DEMEC is accurate up to a division of 2.7 microstrains whereas the 200 mm DEMEC

is good up to a division of 8 microstrains (referred to Table 3.2 in Sub-Section

3.2.2.1).

Table 5.10. Summary of Short-term results at Level L2, L31, L49, L63.

238

when

reading

started

when

reading

ended

at 42 days at 75 daysat 156

days

TW08/L2 thk 300/285.7 56 127 183 - - -38

TW09/L2 thk 600/375 119 24 143 75 55 45

TC10/L2 vary 159 36 195 - 102 118

TC12/L2 vary 159 38 197 - 90 105

TC01/L31 1200x1200/300 96 72 168 - -30/1.3 140/3.1

TC03/L31 2800x800/311.1 139 29 168 - 120/1.8 223/2.6

TC04/L31 2800x800/311.1 139 29 168 - 37/1.8 266/2.6

TC06/L31 1600x1400/373.3 105 20 125 - -6/1.7 216/3.9

TC07/L31 3400x800/323.8 132 27 159 - 53/1.4 160/2.5

TC08/L31 3400x800/323.8 139 20 159 - 108/1.6 180/2.5

TC13/L31 3500x800/325.6 139 35 174 73/0.7 130/1.6 243/4.3

TC14/L31 3500x800/325.6 139 35 174 - 80/1.6 198/4.3

TW18/L31 thk 300/115.4 96 85 181 - - 217

TC03/L49 2200x800/293.3 104 54 158 - 68/2 138/4.8

TC05/L49 2200x800/293.3 70 54 124 - 105/4.3 -

TC07/L49 3400x600/255 49 54 103 - 53/1.6 -

TC08/L49 3400x600/255 83 54 137 - -17/1.6 -

TC09/L49 4000x400/181.8 83 54 137 - 74/3.4 -

TC10/L49 4000x400/181.8 104 54 158 - 10/3.4 198/6.7

TC11/L49 3200x600/252.6 104 54 158 - 82/3.6 138/5.2

TC13/L49 3000x800/315.8 104 54 158 - 27/4.0 21/5.1

TC14/L49 3000x800/315.8 104 54 158 - 25/4.0 42/5.1

TC02/L63 1000x1000/250 121 41 162 0/1.8 89/4.7 63/7.2

TC04/L63 2200x400/169.2 121 35 156 30/1.5 61/3.9 166/6.7

TC06/L63 3000x400/176.5 121 35 156 18/1.6 59/4.3 135/7.1

TC08/L63 1800x600/225 121 35 156 36/1.6 67/4.9 127/6.8

TC10/L63 1800x400/163.6 121 35 156 26/2.0 57/4.9 123/5.6

TW01/L63 thk 300/115.4 121 54 5 - 66 164

TW09/L63 thk 400/142.9 121 54 175 - 28 84

TW16/L63 thk 300/115.4 101 67 168 - 17 24

TW19/L63 thk 300/115.4 108 67 17 - 22 134

AXS/Stress (microstrain/MPa)

Column

/levelSize (mm)/ V/S ratio

DEMEC

reading

duration

(day)

Age of column (days)

5.4.5 Implication of the long-term axial shortening results

Long-term axial shortening of columns and core walls measured from the site at level

B2 are presented in Figure 5.33 to Figure 5.35 respectively.

In Figure 5.33, measured strains from day zero due to shortening of column TC06, are

presented. Also the results from the similar opposite columns TC05 for which the

measurement started at later age are given. As the building is symmetrical, the

missing early results of these columns were assumed equal to those of their opposite

columns. Measured results of TC05 were then projected.

239

Similar procedure also applied to axial shortening data of TC09 and TC10; TC11 and

TC12. The long-term axial shortening of TC10 is ploted with projected data of TC09

and TC12 is ploted with projected data of TC11 as presented in Figure 5.34 and

Figure 5.35 respectively.

Figure 5.33. Plot of Axial Shortening from TC05 and TC06 at level B2

TC06

TC05

0

100

200

300

400

500

600

700

800

0 200 400 600 800 1000

Age (days)

Ax

ial

Sh

ort

enin

g (

mic

rost

rain

)

TC06

TC05

TC06

TC05

240

TC09

TC10

0

100

200

300

400

500

600

700

800

0 200 400 600 800 1000

Age (days)

Ax

ial

Sh

ort

enin

g (

mic

rost

rain

)

TC09

TC10


TC11

TC12

0

100

200

300

400

500

600

700

800

0 200 400 600 800 1000Age (days)

Ax

ial

Sh

ort

enin

g (

mic

rost

rain

)

TC11

TC12


TC122

TC119

TC10

TC09

241

From Figure 5.33 to Figure 5.35, It is observed that axial shortening results of TC06,

TC10 and TC12 correlate well with those of their companion columns. Axial

shortening of columns range from 480 microstrain to 750 microstrain at 750 days.

This clearly indicates symmetrical behavior. It was found that size and shape of

columns affects the amount of shrinkage occurring. This result in greater early strain

measured at TC11 and TC12 as compared to TC05 and TC06 despite their similarity

in stress history.

Measured axial shortening of core walls TW16 and TW18 are treated in the same way

as the results of TC05, TC09 and TC11. They are projected and plotted together with

results of TW09 in Figure 9.

TW09

TW16

TW18

0

100

200

300

400

500

600

700

800

0 200 400 600 800 1000

Age (days)

Ax

ial

Sh

ort

enin

g (

mic

rost

rain

)

TW09

TW16

TW18

Figure 5.36. Plot of Axial Shortening from TW09, TW16 and TW18 at level B2

All core walls shorten at the same amount after 400 days as clearly seen in Figure

5.36. Results of TW09 correspond particularly well with the stress history. The effects

242

of rapidly developed stress over the first 200 days are evident. Generally, axial

shortening of core walls are slightly lower than those of columns regardless of the

higher subjected stresses. They are around 480 to 500 microstrain at 750 days.

For all cases, it can be seen that axial shortening progresses over time in a certain

pattern. In the first 200 days, it arises rapidly with or without significant load

application. The axial shortening curves of this time period more or less resemble

typical shrinkage and creep curves. Then during the age of 200 days to 400 days, the

curve is more in line with the stress history and after 400 days, the trends of stress and

axial shortening development are linearly matched.

These observed findings identify that the contribution of axial shortening components

such as creep, shrinkage and loading history varies over time. According to the

measured results and the stress history, most of the strain presence within the first 200

days is due mostly to shrinkage than afterwards where creep and elastic components

take over the part as a result of high cumulative stress and the aging of concrete.

When comparing axial shortening results of the Q1 Tower to data from other building

with the same level of height (see Section 2.6.2) i.e. Water Tower Place (Russell and

Larson, 1989) the column strains measured at the Q1 Tower were generally lower

than those measured at the Water Tower Place. This is possibly due to the higher

concrete strength, slower loading rates and lower stress level in columns that were

present at the Q1 Tower building.

Referred to differential shortening results within level B2 of the Q1 Tower reported

earlier (see Figure 5.14 and Table 5.5), no major differential shortening between

columns and their adjacent core walls can be observed in this study unlike the

observations at Water Tower Place where lesser amount of axial shortening in core

walls were evidenced. This may be a result of equal design stress in all vertical

members of the Q1 Tower building whilst the other two had significantly lower

stresses in core walls.

243

5.5 Comparisons of Measured Axial Shortening and Predictions

5.5.1 Computation of axial shortening

The computation of axial shortening prediction applies the principles of superposition

and linear theory of viscoelasticity. This is based on the assumption that concrete

creep strains vary linearly in accordance with the magnitude of applied stress and

creep strains due to each applied stress, which can be superimposed directly.

Axial shortening is assumed to consist of three contributing components namely,

instantaneous, creep and shrinkage (these components effect a positive strain over

time) and one resisting component: reinforcement restraint, which gives a negative

strain. Strain values of these components are summed up together at selected time

intervals to obtain the magnitude of axial shortening. If the prediction is accurate then

measures for accommodating axial shortening can be implemented effectively.

Key steps for the computation of axial shortening are presented in the flow chart

shown in Figure 5.37. Each step is briefly described as follows:

1. The analysis starts with the input in two categories: “General input” and “Loading

input”.

1.1. General input consists of information on geometry and properties of material

of the member being analyzed, i.e. size, shape, height, concrete grade, size

and number of rebars. This information for columns and cores are presented

in Appendix A.1. General input also includes information on the ambient

environment to which the member is exposed. The types of environment, as

defined in AS3600-2001 are adapted in the analysis. In compliance, the

average values of daily relative humidity recorded in the Bureau of

Meteorology’s annual local weather report was used. The nearest weather

station is located about 5km from the Q1 Tower at Southport, Gold Coast,

Australia (http://www.bom.gov.au/). Further details, including effects of the

ambient environment on axial shortening are elaborated in Section 5.5.2.

244

1.2. Loading input consists of information required for the estimation of axial

loading in the member being analyzed, i.e. self weight, construction load and

occupancy load whereas Self weight is calculated from the member size given

in Appendix A, and calculation of construction and occupancy loads is

discussed in Section 5.3.

2. Applied loading at specific time intervals is calculated. The most realistic loading

sequence is pre-defined in order to give an accurate simulation of the magnitude

and rate of loading on the member. The loading sequence is described in Section

5.3.

3. Axial shortening components, i.e. creep, shrinkage, instantaneous and

reinforcement restraint are calculated using material models recommended in

AS3600-2001 (previously reviewed in Section 2.4 and the prediction accuracy is

compared with other models using actual laboratory data outlined in Section

4.4.4) and axial shortening models proposed by Bazant (1972), Age-Adjusted

Effective Modulus Method (previously reviewed in Section 2.5.4).

3.1. The Shrinkage is considered independent of the effects of other components

as it is not dependent on the loading and affected only by environmental and

size factors of the drying surface. In practice, shrinkage of a reinforced

concrete member can be directly calculated by using formulae for shrinkage

of plain concrete, recommended in literatures. This is based on the

assumption that the reduction due to reinforcement restraint in the member is

taken into account when calculating instantaneous and creep components.

The selected model and calculation procedures for the shrinkage strain are

given in Section 5.5.3.3.

3.2. Time dependent properties of concrete in the member at time intervals of

each individual load increment, i.e. compressive strength, elastic modulus and

creep, are calculated in accordance with AS3600-2001. The selected model

and calculation procedures are discussed in Section 5.5.3.

245

3.3. Axial shortening components except shrinkage due to each load increments,

i.e. instantaneous, creep and reduced effects of reinforcement restraints are

calculated using Age-Adjusted Effective Modulus Method using values of

concrete properties obtained in the earlier step. Details of the axial shortening

calculation are given in Section 5.5.4.

4. Finally, axial shortening of the member at the required time intervals is calculated.

Given that the superposition theory (McHenry, 1943) is valid, instantaneous and

creep strain induced by each load increment, reduction due to reinforcement

restraints and shrinkage, is summed up to obtain the value of axial shortening at

the selected time interval. Further details are also given in Section 5.5.4.

General input

- column size, shape

- reinforcement size, quantity

- type of environment

Loading input

- self weight

- resulting construction load from

each floors

- resulting occupancy load from

each floors

Calculation of applied loading

- using assumed construction

sequence +actual history

- stage construction applied

(3,7,14,etc days of ages)

Calculations of time dependent

properties of the concrete (AS3600-

2006, draft code)

-strength

-elastic modulus

-creep

For each applied loading at defined

stages

Calculation of shrinkage

(AS3600-2006, draft code)

Calculation of axial shortening

components (Superposition + Age

Adjusted Modulus Method)

- instantaneous (elastic) shortening

- creep shortening+relaxation

- reinforcement restraint

Total axial shortening computed

246

Figure 5.37. Calculation of Axial Shortening flow chart.

5.5.2 Effects of the Ambient Environment

Changes in ambient environment of the concrete member may affect the rate of axial

shortening, especially if the member is fully exposed. In addition, the rate of creep

and shrinkage may be affected by changes in temperature and relative humidity as

they directly influence the amount of water loss from the member, which is the key

factor inducing creep and shrinkage.

While relative humidity influences the rate of creep and shrinkage, temperature

changes influence expansion and/or contraction in the exposed concrete member. The

effect of temperature changes depend on the thermal coefficient of the concrete which

is normally considered equal to 10 microstrains per each degree Celsius of

temperature change.

Wind conditions also have some influence on axial shortening of the concrete

member. Strong winds may induce extra evaporation of surface moisture thus

increasing the rate of shrinkage and creep. Internal forces due to wind loads may also

induce extra axial strain in some members and restraining strain in others. However,

effects of wind action is usually neglected in axial shortening analyses since the effect

is temporary and wind conditions such as magnitude and direction of wind change

constantly. Wind conditions recorded by the Australian Bureau of Meteorology are

shown in Table 5.11. The annual average wind speed taken from the Gold Coast

Seaway is 22.6 km/h.

The average monthly relative humidity of Surfers Paradise, Gold Coast recorded by

the Australian Bureau of Meteorology is shown in Figure 5.38. It was seen that the

daily differences between 9 a.m. and 3 p.m. is not more than 10% as the monthly

relative humidity does not rapidly change. The average monthly relative humidity

ranges between 56% and 72%.

Similarly, monthly temperature data obtained from the Australian Bureau of

Meteorology is shown in Table 5.11. The table shows that average differences in

247

temperature between 9 a.m. and 3 p.m. is 1.5°C, as the average monthly temperature

ranges from 16.5°C to 26.7°C. This indicates that the daily effect of temperature

changes can be neglected as the induced strain is very small (15 microstrains).

Average monthly relative humidity - Surfers Paradise, Gold Coast

0

10

20

30

40

50

60

70

80

90

100

Jan Feb Mar Apr May Jun Jul Aug Sep Oct Nov Dec

Month

Mea

n P

erce

nta

ge

Rel

ati

ve

Hu

mid

ity

9:00 AM

3:00 PM

Figure 5.38. Graphs of average monthly relative humidity, Surfers Paradise CBD

(Gold Coast Seaways Station). Source: Australian Bureau of Meteorology.

Table 5.11. Average monthly weather data, Surfers Paradise CBD (Gold Coast

Seaways Station). Source: Australian Bureau of Meteorology.

Jan Feb Mar Apr May Jun Jul Aug Sep Oct Nov Dec Annual

25.8 25.7 24.6 22.5 19.5 17.1 16.5 17.6 20 21.9 23.6 24.8 21.6

71 72 72 69 69 69 66 62 63 64 65 67 67

17.9 17 18.1 16.6 14.2 13.5 12.8 15.2 16.5 17.4 18.4 17.4 16.3

26.4 26.7 25.6 24 21.8 20 19.7 20.1 21.5 22.5 23.9 25.2 23.1

71 71 70 66 63 59 56 57 62 67 69 69 65

24.8 23.3 24.5 22 19.3 18.6 19.1 22.2 23.8 24.7 24.4 24.5 22.6

Mean 3pm relative humidity (%)

Mean 3pm wind speed (km/h)

Mean 9am temperature (Degrees C)

Mean 9am relative humidity (%)

Mean 9am wind speed (km/h)

Mean 3pm temperature (Degrees C)

248

5.5.3 Prediction of the material properties

5.5.3.1 Prediction of Compressive Strength

Compressive strength of columns at time, t, in days was determined using AS3600-

2009 code recommendations if concrete grade is lower than 50 MPa. For concrete

grade 50 MPa and above, Equation 4.7 which was proposed in Section 4.2 is used.

The algorithm of concrete strength calculation and the equations adopted are

presented in a flow chart shown in Figure 5.39 below

START

STOP

t, f’c

'

cc ftf

t

t

85.00.4

Is

f’c < 50 MPaYES

3103.0'3071.4 cf

100

'

2.271.4

c

t

c

fetf

NO

Figure 5.39. Compressive strength prediction procedures.

249

5.5.3.2 Prediction of Elastic Modulus

The elastic modulus at time, t, in days was calculated by using methods by ACI 209-

1992, AS 3600-2001, AS 3600-2009 (referred to Equation 2.10, 2.7 and 2.9 in Section

2.4.1) and the proposed Equation 4.8 which gives best fit to the tested elastic modulus

results (see Section 4.3).

Note that the concrete density is taken as equal to 2400kg/m3 when using prediction

methods.

5.5.3.3 Prediction of Shrinkage

The drying shrinkage of columns at time, t, in days was calculated by using methods

by ACI 209-1992, AS 3600-2001, AS 3600-2009 (referred to Equation 2.38, 2.54 and

2.56 in Section 2.4.2)

The moist curing condition is assumed and the final shrinkage value is taken equal to

780, 850 and 800 microstrain for shrinkage prediction by ACI 209-1992, AS 3600-

2001 and AS 3600-2009 respectively.

5.5.3.4 Prediction of Creep

The creep of columns at time, t, in days was calculated by using methods by ACI 209-

1992, AS 3600-2001, AS 3600-2009 (referred to Equation 2.41, 2.55, 2.61 in Section

2.4.2)

For ACI 209-1992 and AS3600-2001, relative humidity is taken equal to 67% and

type of exposed environment is type 4 (tropical or a near coastal climate)

250

For AS 3600-2009, Average annual humidity is taken equal to 50% and k4 equal to

0.50 (tropical or a near coastal climate)

5.5.4 Axial shortening calculation

Based on the AEMM method (Bazant, 1972) as adopted in the COLECS computer

program by Beasley (1987) incorporated with updated material models (referred to

Chapter 4) and modified construction sequences using a stress history derived from

actual construction history (previously derived in Section 5.3), the axial shortening of

columns and core walls studied at Q1 tower was predicted and comparison were made

to the data readings.

The constitutive law adopted in COLECS (Beasley, 1987) which represents total axial

shortening at any given time, t, in days is represented as follows

stCsE

tst

sE

t

tstCsE

tst

sE

tt

c

i

c

i

s

c

pip

c

pip

,,

,,

1

1

Equation 5.1

where:

σip = initial concrete stress due to applied load (component 1)

∆σp(t) = changes in concrete stress due to applied load (component 2)

σiε = initial concrete stress due to shrinkage strain (component 3)

∆σε(t) = changes in concrete stress due to shrinkage strain (component 4)

χ(t,s) = aging coefficient

It should be noted that the first term in Equation 5.1 deals with elastic strain of the

columns/walls under variable loading conditions. The second term represents creep

strains and the third terms represent shrinkage strain. The fourth and the fifth term

take into account the effect of reinforcement to creep-shrinkage-reinforcement

251

interaction, these terms therefore equal to zero in the case of plain concrete

applications.

The expression for component 1, σip, was given as follows

ce

ipA

P

psn

1

1 Equation 5.2

where:

P = applied load

p = percentage reinforcement

ne(s) = time dependent parameter of modular ratio n at time s

The time dependent parameter of modular ratio n at time s, ne(s), can be obtained

from the equation below:

Equation 5.3

The component 2, ∆σp(t), takes the following form:

ip

e

ep C

ptn

psnt 1

1 Equation 5.4

The time dependent parameters, ‾ne(t), is given as:

stCstsE

Etn

c

se ,,1 1 Equation 5.5

The term aging coefficient, χ(t,s), for 2 ≤ s ≤ 100 days, can be determined using the

following relationship:

ssCsE

Esn

c

s

e ,1 1

252

st

stst

20

11. Equation 5.6

Where

sk

sk

2

1

and

sCek

,33.1

114.078.0

sCek

,33.1

218.016.0

The component 3, σiε, is as follows

pEspsn

ss

e

i 1

1 Equation 5.7

Finally the component 4, ∆σε(t), can be determined as follows

tEpsn

Cptn

psnt sci

ee

e 11

11 Equation 5.8

The algorithm for axial shortening calculation based on Equation 5.1 using a

spreadsheet program is presented in Figure 5.40 (see Section 5.5).

5.5.5 Axial shortening spreadsheet

253

A Microsoft Excel spreadsheet was developed in order to compute axial shortening of

selected columns and core walls by using a modified AEMM method as previously

outlined in 5.5.4 incorporate with actual stress history (referred to Section 5.3) and

various material models including best-fitted model based on the tested data (see

Section 5.5.3).

The spreadsheet consists of the twelve worksheets as follow;

1. Input: Details of column (i.e. ID, shape, diameter, length, width, height),

concrete properties (i.e. concrete grade, slump, cement content, air content,

aggregate content, fine aggregate content), reinforcement properties (i.e.

quantity and size of rebars, elastic modulus of steel), condition of exposed

environment (i.e. type of exposed environment or percent relative humidity)

and age of concrete at first loading are entered into this worksheet. Then

theoretical thickness, rebar area, percentage of fine aggregates which are data

required for further axial shortening computation is determined.

2. Loading: Loading data includes column self weight, construction load and

occupancy load subjected to each floors are entered into the space allocated in

this worksheet. A total applied load at various designated days is then

computed according to the pre-entered construction history and the loading

data. Each load increment applied to the column is automatically transferred

into an independent column in this worksheet. This allows further calculation

of axial shortening due to each separate loading increment according to the

superposition theory.

3. Strength: Concrete strength at age 1 day to 3650 days and at the age of initial

loading are computed based on any of the following methods: ACI 209-1992,

AS3600-2001, AS3600-2009 and Equation 4.7.

4. Elastic Modulus: Elastic modulus at age 1 day to 3650 days and at the age at

application of each load increment are computed based on any of the

254

following methods: ACI 209-1992, AS3600-2001, AS3600-2009 and

Equation 4.8.

5. Creep Coefficient: Creep coefficient at age 1 day to 3650 days due to each

load increment are computed based on any of the following methods: ACI

209-1992, AS3600-2001, AS3600-2009.

6. Shrinkage Coefficient: Shrinkage coefficient at age 1 day to 3650 days due to

each load increment are computed based on any of the following methods:

ACI 209-1992, AS3600-2001, AS3600-2009.

7. Component 1: initial concrete strain due to applied load, σip, (component 1) at

application of each load increment and time dependent parameters of modular

ratio n, ne(t), at age 1 day to 3650 days due to each load increment are

determined.

8. Component 2: changes in concrete stress due to applied load, σp(t),

(component 2); aging coefficient, χ(t,s); and time dependent parameters of

modular ratio n, n‾e(t); at age 1 day to 3650 days due to each load increment

are determined.

9. Component 3: initial concrete stress due to shrinkage strain, σiε, (component

3); and changes in concrete stress due to shrinkage strain, ∆σε(t), (component

4) at age 1 day to 3650 days due to each load increment are determined.

10. AXS: elastic strain, creep strain, and reinforcement interaction at age 1 day to

3650 days due to each load increment are determined. The summation of these

strain components due to all load increments and shrinkage strain as

determined earlier in worksheet “shrinkage” finally gives axial shortening

strain of column at time 1 day to 3650 days. The algorithm for calculation of

axial shortening is given in Figure 5.40.

255

11. Measured Data: Axial shortening data obtained from data reading at Q1 tower

are entered into this worksheet for comparison purposes.

12. Graph measured vs. predicted: axial shortening data obtained from Q1 tower

(age of column up to 1000 days) and the followings graphs of strain against

age of column in days are presented in this worksheet:

Elastic shortening

Creep shortening

Shrinkage shortening

Reinforcement restraining

Total axial shortening

13. Graph 10 yrs: similar type of graphs as presented in worksheet “Graph

measured vs. predicted” are represented in this worksheet for age of columns

up to 3650 days (10 years)

256

START

STOP

Input

stCsE

tst

sE

t

tstCsE

tst

sE

tt

c

i

c

i

sh

c

pip

c

pip

,,

,,

1

1

ip

tp

sEc

st,

stC ,1

tsh

i

t

257

Figure 5.40. Axial shortening calculation procedures

5.5.6 Axial shortening results versus predictions at level B2

In order to evaluate the accuracy of modified AEMM method when different material

models are incorporated, measured axial shortening of columns and core walls in

microstrains are presented together with 10 year predicted values in Figure 5.41 to

Figure 5.48. Similar comparison curves for prediction period of 1100 days (the data

collection period) are also presented in Appendix A in larger sizes for clarity. Note

that estimated values of axial shortening components which are elastic shortening,

shrinkage, creep, and reinforcement restraint are also given within the same plots.

In general, it can be seen that axial shortening at any time is a summation of elastic,

creep and shrinkage strain subtracting by strain in opposite direction due to

reinforcement restraint. It is observed that elastic and creep strain are the two largest

axial shortening components respectively (the combined two make approximately 70-

80% of positive strain). Shrinkage is the third largest component which is relatively

smaller (approximately 20-30% of positive strain). Reinforcement restraint helps

reducing axial shortening by a small quantity which is equal to approximately 5-10%.

The four graphs consisting of axial shortening results of set B2/C1 columns i.e. TC05,

TC06, TC10, TC12 and core wall TW09 which have the most comprehensive data

readings are presented in Figure 5.41 along with their predicted axial shortening using

ACI 209R-1992 models for estimation of concrete properties. Given that the floor

plan of Q1 tower is symmetrical therefore opposite columns are identical and having

similar stress history, data reading of TC09 and TC11 (data reading commenced at

age 370 days) are then interpolated with data reading of TC10 and TC12 at age 370

days and presented therein.

Measured and predicted axial shortening of similar columns and core walls i.e. set

B2/C1, TC09, TC11 and TW09 compared against predicted axial shortening when

other material models are incorporated into AEMM method are presented in Figure

5.42 to Figure 5.44. The material models used are AS 3600-2001, AS 3600-2009 and

258

combination of models providing best fit to the test data of concrete (referred to

Chapter 4, Section 4.2, 4.3, 4.4, and 4.5)

From Figure 5.41 to Figure 5.44, it can be seen that axial shortening of columns can

be predicted with similar accuracy when AS3600-2001 and AS3600-2009 material

models are adopted. ACI 209R-92 is the only material model that leads to under

estimation of columns axial shortening at all ages when used thus least suitable for Q1

concrete. Most accurate prediction of columns shortening is achieved when

combination of material models which proven to be most compatible to test data of

Q1 concrete are used (so called best material models).

Similar accuracy also observed after 1000 days for columns as ACI 209R-1992

clearly under-estimates axial shortening. For the other models, exact comparisons

could not be made due to the lack of long-term data reading. However, from the

observed trend of axial shortening developed before 1000 days when the building

takes constuction load and selfweight compared with the trend after 1000 days when

the building takes full occupancy load, it is more likely that the adoption of AS 3600-

2001 and AS 3600-2009 would lead to over-estimation again as approximately 20%

over-estimation already occurred at 1000 days (see Figure 5.42 and Figure 5.43). The

most precise axial shortening estimation after 1000 days is therefore likely to be

achieved in cases that the most suitable material models are adopted into AEMM

method. It can be seen that significant prediction discrepancy at 1000 days does not

exist when best material models are adopted.

Prediction of axial shortening at core wall TW09 is not as accurate as compared to

those of columns as over-estimation is observed in all cases. However, using best

material models also clearly improve prediction accuracy for TW09.

259


models), set B2/C1 columns and their opposite columns, and TW09-B2.

260


models), set B2/C1 columns and TW09-B2.

261


models), set B2/C1 columns and TW09-B2.

262


B2/C1 columns and TW09-B2.

263

Measured and predicted axial shortening of the remaining columns of set B2/C1 i.e.

TC01, TC02, TC03, TC04, TC08, TC08, TC13, TC11 compared against predicted

axial shortening when using various models i.e. AS 3600-2001, AS 3600-2009 and

combination of models providing best fit to the test data of concrete (referred to

Chapter 4, Section 4.2, 4.3, 4.4, and 4.5) are incorporated into the modified AEMM

method are presented in Figure 5.45 to Figure 5.48 respectively. It can be seen that

data readings of most columns with the exception of TC01 and TC02 are available

after approximate age 370 days when measured axial shortening are interpolated on

the prediction lines.

In general, similar findings as found in the previously mentioned case study of

column set B2/C1 and core wall TW09 are also noted for this set of columns.

Referring to data reading of TC01 and TC02, it was found again that ACI 209R-92

under-estimates axial shortening of all columns whereas AS3600-2001 and AS3600-

2009 gives improved prediction. Again, it was found that using suitable adjusted

material models again lead to more accurate axial shortening prediction over the

entire study period.

It is noted that the plots of result comparison of columns without early age data

(TC03, TC04, TC07, TC08, TC13, and TC14) reveals that effects of each material

model to the axial shortening prediction after 370 days is not tremendous. However,

using AS3600-2001 model appears to yield the most accurate after 370 days.

264


models), set B2/C2 columns.

265



266



267


B2/C2 columns.

268


In this chapter, handling process of axial shortening data from field sheets to a

database, conversion of gauge reading data stored in database queries into axial

shortening data through processes that preset in a spreadsheet, and prediction of axial

shortening using AEMM method are fully described.

It is without doubt that experiment conducted at the Q1 Tower has produced a huge

amount of axial shortening data. It is evident that the handling of data in this project

would not be possible without assistant of the computerised database facility. A

systemetic calculation process performs by means of a spreadsheet is also crutial to

the task of turning countless series of gauge reading which accumulated over the three

years data collection into plots of axial shortening.

To facilitate the axial shortening predictions, a spreadsheet containing full calculation

process according to AEMM method was developed based on algorithm previously

used for COLECS (Beasley, 1987). However, there has been a modification on

construction sequence to match the actual construction history of the Q1 Tower. The

simulation of stress in columns and core walls over the period of observation is

another sophisticate matter that was also presented. ACI209R-92, AS3600-2001 and

AS3600-2009 models as were employed for the calculation of compressive strength,

elastic modulus, creep, and shrinkage. Equation fitted to the actual data of

compressive strength and elastic modulus as presented in Section 4.2 and 4.3 are also

provided.

From the axial shortening results presented , it was found that axial shortening curves

is fairly sensitive to chances which can be due to uncontrolled factors that exist on

site. This sensitivity was also observed on the World Tower data as well as data from

overseas researches. Data at level B2 and B2 are the most comprehensive eventhough

early age data are availably at some locations. It was found that axial shortening is

affected mostly by shrinkage at early stage, elastic shortening couples with early

creep occurred at the application of each load increment and the remaining shrinkage

269

contributes to axial shortening after 200 days upto the end of the construction, and

creep becomes the main contributor to axial shortening beyond.

Symmetrical behaviour of the building was confirmed as columns located at the

opposite side of the tower had exactly similar axial shortening curves. However, It

was found that axial shortening in different columns occurred at different rate

depending on the loading and size. But, at the end of the data collection, it appeared

that there is no major difference between columns that located in one level. At 893

days , maximum axial shortening occurred at TC06-B2 was 733 microstrain or 1.64

mm, a mere L/1370 of the 2.25 metres floor height.

Also there is no significant differential shortening occurred amoungst core walls. The

axial shortening of core walls was even less. At 893 days, axial shortening at TW09-

B2 was 610 microstrain or 1.37 mm, equal to L/1640. The worst case differential

shortening between columns and adjacent core walls at level B2 occurred between

column TC07 and core wall TW18 and was only 128 microstrain or 0.3mm, L/7500

of the 2.25 metres floor height. The differential shortening found was within the

acceptable limit of L/500 which was specified in the Q1 Tower design report

(Sunland Group, 2002).

From the long-term data, it was evident that AEMM method adopted give axial

shortening prediction with varying accuracy depending on the material model

adopted. The improvement in terms of prediction accuracy was achieved when

material model suggested for Q1 Tower concrete were used.

5.7 References



Mich., 1992, 47 pp.

270







Beasley, A.J., “A Numerical Solution for the Prediction of Elastic, Creep, Shrinkage

and Thermal Deformations in the Columns and Cores of Tall Concrete Buildings,”

Research Report CM – 87/2, Civil and Mechanical Engineering Department, Faculty

of Engineering, University of Tasmania, 1987

S. Bursle, N. Gowripalan, D. Baweja and K. Kayvani, “Refining Differential



S.Bursle., “Influence of Material Related Parameters on the Axial Shortening of High

Strength Concrete Columns,” Masters of Engineering Thesis, School of Civil and

Environmental Engineering, University of New South Wales, Sydney, Australia,

2006.

McHenry, D., “A new aspect of creep in concrete and its application to design,”

Proc.ASTM, Vol.43, 1943, pp. 1069-1084.

Russell, H.G., and Larson, S.C., “Thirteen Years of Deformations in Water Tower

Place,” ACI Structural Journal, V.86, No.2, Mar-Apr.1989, pp.182-191.

271

272

6 Parametric evaluation of the AEMM adopted model

6.1 Introduction

The sensitivity analysis was employed in order to facilitate the evaluation of the

degree of influence due to variation of each input factors to major axial shortening

components and the resultant shortening. This detailed study was conducted using the

spreadsheet developed earlier (see Section 5.5.1) as a tool to compute value of axial

shortenings with variations of parameter investigated.

To perform the sensitivity analysis, input variables of column TC06-B2 was adopted.

Best material model and AEMM was employed for axial shortening computation. The

variations of any of these parameters are input into the spreadsheet one at a time while

the remaining parameters remained constant. The resulting axial shortening value due

to altered parameter is benchmarked against axial shortening value obtains using

default parameter.

The investigated parameters and the variation adopted are listed as follow:

Specified concrete grade (f’c): ranging from 40 to 80 MPa. Note that this value

is used since the excess value of tested 28-days strength of concrete delivered

to Q1 Tower has been taken into the consideration of material model

adjustment as outlined in Section 4.2. Results are benchmarked against the

default value taken as 65 MPa.

Concrete density: ranging from 2300 to 2500 kg/m3. Reference value taken

equal to 2400kg/m3.

Construction/stress history: regular construction cycles ranging from 7 days to

15 days are employed. Results were compared against the actual construction

cycle of Q1 Tower which having irregular cycle for each floors.

Average stress level: service stress ranging from 0.1f’c to 0.3f’c comparing

against 0.2f’c (service stress of column TC06-B2).

273

Environment condition: Axial shortening when Arid and temperate inland

conditions are specified is benchmarked against result using coastal condition

(also Q1 Tower’s condition).

Average annual humidity: ranging from 40 to 60 percent. Results are

benchmarked against average annual shrinkage value of 50% (approximation

of Gold Coast annual humidity, refer to Section 5.5.2)

Final shrinkage value: ranging from 700 to 900 microstrain. Reference value is

taken as 800 microstrain (as per AS3600-2009 recommendation for Brisbane

vicinity)

Basic creep coefficient: the creep coefficient value of 1.6 and 2.0 are

employed, compared against 1.88 (computed value for concrete grade 65

MPa, AS3600-2009 creep model)

Percentage of reinforcement: from 1 to 5 percent compared against 2.4 percent

(actual steel ratio of TC06-B2)

The variations adopted in the study are summarised in Table 6.1. The first and the

second columns represent the lower and upper bound of each parameter while default

value being benchmarked against is given in the third column respectively. When a

parameter was analysed, the value of the remaining factors remained constant and

taken as per “Reference Value” column shown in Table 6.1 below and the factor

being investigated was substituted by value within the specified lower and upper

limit.

Table 6.1. The variation of investigated parameters.

Parameters Unit

Variations

Lower

Limit

Upper

Limit

Reference

Value

Specified concrete grade (f’c) MPa 35 85 65

Concrete density Kg/m3 2300 2500 2400

Construction cycle

days per

floor 7 10 varies

Average stress level 0.1f'c 0.3f'c 0.2f'c

Environment condition Arid Temperate Coastal

Final shrinkage value microstrain 600 1100 800

Basic creep coefficient 1.4 2.2 1.88

Percentage of reinforcement 1 5 2.4

274

6.1.1 Sensitivity analysis – concrete grade

The sensitivity of the predicted axial shortening as a function of compressive strength,

concrete density, construction cycle, environment condition, final shrinkage value,

basic creep coefficient, and precentage of reinforcement are presented in Figure 6.1 to

Figure 6.11 respectively. Note that only axial shortening at 1, 5, 9, 28, 56, 140, 270.

365, 550, 720, 930, 1278, 1825, 2555, and 3650 days are determined.

Figure 6.1. Sensitivity of AEMM adopted model due to compressive strength.

It is shown in Figure 6.1 that compressive strength has an increasing effect on

predicted axial shortening with an increasing time. Higher grade concrete was found

having lower amount of axial shortening. It is noted that the predicted axial shortening

values do not appear to be different before 140 days. The model is also more sensitive

at later ages. There is a greater change between 45 and 55 MPa, then there is between

35 MPa and 45 MPa. It is evident that concrete with relatively lower grade i.e. 35

MPa and 45 MPa has significant effect on sensitivity of AEMM adopted model .

275

However, the sensitivity of the model is less as compressive strength increase,

particularly in case of concrete grade 55 MPa and higher.

6.1.2 Sensitivity analysis – concrete density

The sensitivity of the predicted axial shortening as a function of concrete density is

presented in Figure 6.2. It is clear that changing of concrete density has small effect to

value of predicted axial shortening particulary when concrete is younger than 1 year.

This is due to the fact that the change of concrete density affect very little on the value

of compressive strength.

Figure 6.2. Sensitivity of AEMM adopted model due to concrete density.

6.1.3 Sensitivity analysis – construction cycle

The sensitivity of predicted axial shortening as a function of construction cycle is

presented in Figure 6.3 and Figure 6.4. The assumed construction cycle per 1 floor of

7 days, 10 days and actual cycles (varying between 4 days to 36 days, average 11

days) are adopted while keeping the opening date of the building at 930 days. Figure

276

6.3 presents the influence of construction cycle while keeping concrete grade at 65

MPa. Figure 6.4 presents the sensitivity while concrete grade is 45 MPa.

Note that the actual construction cycle refers to as-built construction history of Q1

Tower according to Table 5.1 presented earlier in Section 5.3.2.


grade = 65 MPa).

277


grade = 45 MPa).

It is evident from Figure 6.3 and Figure 6.4 that changing construction cycles has

effects on the prediction of axial shortening only during the construction after load

application commenced at 56 days. It is also shown that faster construction leads to

more rapid axial shortening development. However, changing construction cycle has

no effect to long-term predicted axial shortening

The only difference between Figure 6.3 and Figure 6.4 is the concrete grade (65 MPa

and 45 MPa) by which the results indicate that the 45 MPa concrete is more sensitive

to the changing of construction cycle. In total, The effect of changing construction

cycles is greater than the effect of changing compressive strength for concrete grade

55 MPa and above that occurred during a similar construction period.

6.1.4 Sensitivity analysis – environment condition

The sensitivity of the predicted axial shortening as a function of environment

condition is presented in Figure 6.5 and Figure 6.6. Figure 6.5 presents the sensitivity

of the predicted axial shortening as a function of environment condition while keeping

concrete grade constant at 65 MPa. In Figure 6.5, the sensitivity while concrete grade

is kept constant at 45 MPa is presented.

278


(concrete grade = 65 MPa).



279

Figure 6.5 and Figure 6.6 show that predicted axial shortening values over times by

AEMM adopted model show small sensitivity to the changing of the environmental

condition. The model is more sensitive at later ages. Arid environment which has a

dry climate condition induced highest axial shortening. Again, the two figures also

show that the sensitivity of predicted axial shortening value increase as the concrete

grade decrease. A greater change between Coastal and Temperate condition is noted

for both 65 MPa and 45 MPa concrete. Effect of concrete grade to the model

sensitivity are also noted as the overall sensitivity increased when concrete grade

decreased.

The sensitivity of the predicted axial shortening as a function of final shrinkage value

is presented in Figure 6.7and Figure 6.8. Figure 6.7 presents the sensitivity of the

predicted axial shortening as a function of final shrinkage value while keeping

concrete grade constant at 65 MPa. In Figure 6.8, the sensitivity while concrete grade

is kept constant at 45 MPa is presented.



280



As can be observed in Figure 6.7 and Figure 6.8, it is founded that the AEMM

adopted model has little sensitivity to changing of final shrinkage. The effect of this

factor to predicted axial shortening value is minimal for concrete grade 45 MPa. The

change of predicted axial shortening value appears to be the same for each increment

of 100 microstrain increase in final shrinkage value. Similarly to the effect of

changing compresive strength and environment condition, this factor also has an

increasing effect on axial shortening with increasing time. Also the predicted axial

shortening values do not appear to be different at early ages.

6.1.5 Sensitivity analysis – creep coefficient

The sensitivity of the predicted axial shortening as a function of creep coefficient is

presented in Figure 6.9 and Figure 6.10. In Figure 6.9, the sensitivity of the predicted

axial shortening as a function of creep coefficient while keeping concrete grade

281

constant at 65 MPa is presented. In Figure 6.10 the sensitivity while concrete grade is

kept constant at 45 MPa is presented.


grade = 65 MPa).

282


grade = 45 MPa).

Figure 6.9 and Figure 6.10 show that the predicted axial shortening values are

sensitive to changes in creep coefficient over times. Greater value of creep coefficient

means greater creep, consequently induces greather axial shortening. Similar to the

changing of final shrinkage value, the change appears to be the same for each

increment of 0.2 increase in creep coefficient. However, the changes are sensitive

only at later ages. Overall sensitivity increased when concrete grade decreased. The

sensitivity of changing creep coefficient are on par with changing final shrinkage

values and changing environment condition.

6.1.6 Sensitivity analysis – percent of reinforcement

Figure 6.11 and Figure 6.12 present the sensitivity of the predicted axial shortening as

a function of reinforcement percentage. Figure 6.11 presents the sensitivity of the

predicted axial shortening as a function of reinforcement percentage while keeping

concrete grade constant at 65 MPa. In Figure 6.12, the sensitivity while concrete

grade is kept constant at 45 MPa is presented.

283





284

Figure 6.11 and Figure 6.12 show that the predicted axial shortening values are

sensitive to changes in percentage of reinforcement over times. The sensitivity are

greater at later ages. Again, overall sensitivity increased when concrete grade

decreased. The sensitivity of changing percentage of reinforcement is the highest of

the factors analysed.


Based on the sensitivity analysis, it was found that axial shortening prediction by

AEMM method is affected by many factors by which the idividual sensitivity are all

different. It was found that the AEMM method is most sensitive to the changing of

compressive strength, particulaly if concrete grade decreases from 55 MPa to 45 MPa.

Compressive strength do influence the sensitivity of AEMM method due to other

factors, particularly if concrete grade is less than 65 MPa. This is because the lower

grade concrete has large water content (higher w/c ratio) and will produce a larger

shrinkage as compared to those of higher grade concretes. It was found that the

sensitivity of AEMM method due to construction cycle, environmental condition,

creep coefficient, and percentage of reinforcement clearly increases when concrete

grade decreases from 65 MPa to 45 MPa.

Three factors was found not critical for final result of the AEMM prediction.

Construction cycle is only sensitive to changes of construction cycles. Changing of

concrete density did not change the results of AEMM prediction at all. Percentage of

reinforcement, eventhough appears to have significant effects to axial shortening

prediction for 45 MPa concrete, its exact value can be determined from the as-built

drawing thus the chance of prediction error due to discrepancy of reinforcement

percentage is neglect.

285

7 Conclusions and Recommendations

7.1 Conclusions

1. A few number of axial shortening study on tall buildings were reported.

DEMEC gauge was successfully used to take axial shortening data for these

projects. However, only two recent works on buildings that are equivalent to

Q1 Tower in terms of concrete strength used and number of floors are

published. These buildings were the Water Tower Place, Chicago and the

World Tower, Sydney.

2. For compressive strength, a correction factor, η, as per Equation 4.5 was

assigned to improve prediction accuracy of the existing code model.

3. Accuracy of all elastic modulus models are valid for Q1 Tower concrete only

in limited range of compressive strength. A non-linear eqation derived from

the test data as per Equation 4.7 are proposed for prediction of elastic

modulus.

4. The AS3600-2001 model is the best predictor of drying shrinkage for

concrete used at Q1 Tower followed by the ACI209R-92, and AS3600-2009

models . AS3600-20001 model are proposed for shrinkage prediction.

5. The AS3600-2009 model is the best predictor of creep for concrete used at Q1

Tower followed by the AS3600-2001, and ACI209R-92 models . AS3600-

2009 model are proposed for creep prediction.

6. AEMM method incorporated with a simulated construction sequence derived

from actual construction history of the Q1 Tower was used to gives axial

shortening prediction values. Material model as per ACI209R-92, AS3600-

2001, AS3600-2009 and Equation 4.7 and 4.8 were used to calculated concrete

strength, elastic modulus, shrinkage, and creep.

286

7. Based on the long-term predicted and measured axial shortening results,

accuracy of material model has significant effects to the accuracy of AEMM

method. More accurate axial shortening was achieved when recommended

combination of material models for each concrete properties were used

instead of using model entirely adopted from a single code model.

8. The axial shortening prediction model adopted is sensitive to changes in

compressive strength and the sensitivity increases for lower grade concrete.

Other factor also have an influence on axial shortening but the sensitivity of

the axial shortening model to these factors are much smaller. Therefore, the

adopted axial shortening can be improved by replacing compressive strength

model with Equation 4.7.

7.2 Recommendations

1. An average construction cycle may be used for axial shortening prediction by

using AEMM method instead of varying cycles according to actual

construction history.

2. There is a need to develop a simplified prediction model suitable for hand

calculation of axial shortening since AEMM method was found exhaustive.

287

Appendix A

A.1 Details of instrumented Columns and Core Walls

Table A.1. Details of Column TC01, TC02.

B2-L1 1350 diam. 34N40 65 17.3

L1-L3 Varies 65 16.8

L3-L10 1400x1400 34N28 65 11.2

L11-L20 1400x1400 34N28 50 11.2

L21-L30 1400x1400 34N28 50 12.1

L31-L40 1200x1200 34N28 50 12.1

L41-L50 1200x1200 32N24 40 13.5

L51-L59 1000x1000 20N24 40 13.5

L60-L70 800x800 20N20 40 15.7

L71-L77 800x800 20N20 40 12.5

Compressive

Stress

(MPa)

Level Column Size Reinforcement

Details

Concrete

Grade

(MPa)


N/bh N/bh

(DL+LL+Wu) (DL+LL)

B2-B1 2200 72N40 65 21.8 16.9

L1-L2 Varies 65 25.9 19.9

L3-L10 3800x800 60N32 65 26.3 20.2

L11-L20 3800x800 60N32 50 26.3 20.2

L21-L30 3800x800 40N32+20N24 50 15.7 12.8

L31-L40 2800x800 48N32 50 15.7 12.8

L41-L50 2200x800 40N32 40 13.5 11.3

L51-L60 2200x400 40N24 40 13.5 11.3

L60-L70 2200x400 30N24 40 10 12.4

L70-L76 2200x400 30N20 40 4.8 6.7


Details

Concrete

Grade

(MPa)

288


B2-L1 2000 diam. 60N40 65 20.5

L1-L3 Varies 65 19.9

L3-L10 3400x800 44N36 65 21.3

L11-L20 3400x800 44N36 50 21.3

L21-L28 3400x800 44N36 50 16.1

L29-L40 1600x1400 44N36 50 22

L41-L54 1600x1400 44N36 40 20.8

L55-L59 3600x400 36N36 40 18.5

L60-L71 3000x400 8N32+20N28 40 14.2

L72-L74 3000x400 8N32+20N28 40 5.8

L75-L76 900dia 14N36 40 5.8

Compressive

Stress

(MPa)


Details

Concrete

Grade

(MPa)

Table A.4. Details of Column TC07, TC08

N/bh N/bh

(DL+LL+Wu) (DL+LL)

B2-L1 3400x1000 84N40 65 26.7 17.4

L2 3400x1000 84N40 65 32.9 21.1

L3-L10 3400x800 84N40 65 29.4 18.4

L11-L20 3400x800 84N36 50 28 18

L21-L30 3400x800 44N32 50 18.2 12.3

L31-L40 3400x800 44N28 50 18.2 12.3

L41-L54 3400x600 44N28 40 14.9 10.7

L55 1800x600 22N32 40 10.8 13.8

L56-L59 1800x600 22N28 40 10.8 13.8

L71-L75 1200X600 22N24 40 5.6 7.9

Level Column

Size

Reinforcement

Details

Concrete

Grade

(MPa)

289


N/bh N/bh

(DL+LL+Wu) (DL+LL)

B2-L1 4000x900 96N36 65 26.1 16.5

L2 4000x900 96N36 65 32.1 20.1

L3-L10 4000x700 96N32 65 27.6 16.6

L11-L20 4000x700 96N28 50 26.5 17

L21-L40 4000x600 72N24 50 18.5 12.1

L41-L50 4000x400 72N24 40 16.8 14.6

L51-L60 4000x400 40N20+22N24 40 16.8 14.6

L61-L70 1800x400 22N24 40 9.9 12.5

L71-L75 1800x400 22N24 40 1.6 1.9


Details

Concrete

Grade

(MPa)


B2-L1 3200x800 46N40 65 18

L2 3200x800 46N40 65 19.4

L3-L10 3200x800 36N36 65 18.8

L11-L20 3200x700 36N32 50 15.5

L21-L30 3200x700 36N32 50 13.6

L31-L40 3200x600 34N28 50 13

L41-L50 3200x600 34N28 40 11.8

L51-L59 3200x400 32N24 40 11.8

L60-L70 1600x400 16N24 40 11.4

Compressive

Stress

(MPa)


Details

Concrete

Grade

(MPa)

290


N/bh N/bh

(DL+LL+Wu) (DL+LL)

B2-L1 4000x800 108N40 65 27.6 25.3

L2 4000x800 108N40 65 28.4 20.4

L3-L10 4000x800 84N32 65 27.1 20

L11-L20 4000x800 76N28 50

L21-L30 4000x800 36N28+36N24 50 17.4 13.77

L31-L40 3500x800 64N24 50

L41-L50 3000x800 54N24 40 12.3 14.5

L51-L60 2400x800 46N24 40

L60-L71 40 7.3 7.8

L72-L78 40 10.8 12.6


Details

Concrete

Grade

(MPa)

Table A.8. Details of Core Wall TW01

Axial Stress

(MPa)

(DL+LL)

B2-L1 300 65 18.5

L2 300 65 16.3

L3-L10 250 65 16.5

L11-L19 250 50 16.5

L20-L41 250 50 12.0

L42-L59 250 40 8.7

L60-L71 200 12@200mm 40 5.7

L72-L76 200 12@200mm 40 3.7

Level Core Wall

Thickness

(mm)

Reinforcement

Details

Concrete

Grade

(MPa)

291

Table A.9. Details of Core wall TW06

Axial Stress

(MPa)

(DL+LL)

B2-L1 350 20@200mm 65 12.1

L2 300 20@200mm 65 17.7

L3-L10 250 16@200mm 65 17.4

L11-L20 250 16@200mm 50 17.4

L21-L30 250 16@200mm 50 12.2

L31-L40 250 16@200mm 50 12.2

L41-L50 250 16@200mm 40 9.1

L51-L59 250 16@200mm 40 9.1

L60-L70 250 16@200mm 40 7.4

Level Core Wall

Thickness

(mm)

Reinforcement

Details

Concrete

Grade

(MPa)

Table A.10. Details of Core Wall TW07, TW08

Axial Stress

(MPa)

(DL+LL)

B2-L1 400 28@200mm 65 17.3

L1 400 24@200mm 65 17.3

L2 300 24@200mm 65 19.3

L3-L10 250 20@200mm 65 17.8

L11-L19 250 20@200mm 50 17.8

L20-L41 250 20@200mm 50 16.4

L42-L59 250 20@200mm 40 9.1

Level Core Wall

Thickness

(mm)

Reinforcement

Details

Concrete

Grade

(MPa)

292


Axial Stress

(MPa)

(DL+LL)

B2-L1 600 20@200mm 65 18.0

L2 600 36@200mm 65 18.0

L3-L10 300 16@200mm 65 14.0

L11-L19 300 16@200mm 50 14.0

L20-L30 300 16@200mm 50 12.0

L31-L41 300 12@200mm 50 12.0

L42-L59 250 12@200mm 40 11.8

L60-L71 250 12@200mm 40 4.0

Level Core Wall

Thickness

(mm)

Reinforcement

Details

Concrete

Grade

(MPa)


Axial Stress

(MPa)

(DL+LL)

B2-L1 1000 20@200mm 65 12.7

L2 700 20@200mm 65 14.6

L3-L10 500 16@200mm 65 16.8

L11-L19 500 16@200mm 50 16.8

L20-L41 500 16@200mm 50 13.2

L42-L59 400 16@200mm 40 7.7

L60-L71 500 16@200mm 40 3.9

Level Core Wall

Thickness

(mm)

Reinforcement

Details

Concrete

Grade

(MPa)

Table A.13.. Details of Core Wall TW16

293

Axial Stress

(MPa)

(DL+LL)

B2-L1 300 20@200mm 65 11.8

L2 300 20@200mm 65 15.3

L3-L10 250 20@200mm 65 32.5

L11-L19 250 16@200mm 50 32.5

L20-L41 250 16@200mm 50 25.1

L42-L59 250 16@200mm 40 8.5

L60-L71 200 16@200mm 40 6.0

L72-L76 200 16@200mm 40 3.5

Level Core Wall

Thickness

(mm)

Reinforcement

Details

Concrete

Grade

(MPa)


Axial Stress

(MPa)

(DL+LL)

B2-L1 300 32@200mm 65 19.0

L2 300 32@200mm 65 15.5

L3-L10 250 20@200mm 65 15.0

L11-L19 250 20@200mm 50 15.0

L20-L41 250 16@200mm 50 12.0

L42-L59 250 16@200mm 40 8.5

L60-L71 200 16@200mm 40 6.1

L72-L76 200 16@200mm 40 2.8

Level Core Wall

Thickness

(mm)

Reinforcement

Details

Concrete

Grade

(MPa)


Axial Stress

(MPa)

(DL+LL)

B2-L1 350 20@200mm 65 19.1

L2 300 20@200mm 65 18.3

L3-L10 250 20@200mm 65 19.5

L11-L19 250 20@200mm 50 19.5

L20-L41 250 16@200mm 50 14.9

L42-L59 250 16@200mm 40 8.4

L60-L71 250 16@200mm 40 6.4

Level Core Wall

Thickness

(mm)

Reinforcement

Details

Concrete

Grade

(MPa)


294

Axial Stress

(MPa)

(DL+LL)

L3-L10 300 65 14.2

L11-L19 300 50 14.2

L20-L41 300 50 12.1

L42-L59 250 40 9.3

L60-L71 250 12@200mm 40 6.2

Level Core Wall

Thickness

(mm)

Reinforcement

Details

Concrete

Grade

(MPa)

A.2 Axial shortening results by locations

A.2.1 AXS results – Level B2

Figure A.1. AXS results, column TC01, Level B2, Q1 Tower.

295



296



297



298



299



300



301


Figure A.15. AXS results, core wall TW06, Level B2, Q1 Tower.

302



303



304



305

A.2.2 AXS results – Level B1



306



307



308



309



310

Figure A.32. AXS results, column TC12 (front face), Level B1, Q1 Tower.

Figure A.33. AXS results, column TC12 (side face), Level B1, Q1 Tower.

311



312



313



314



315

A.2.3 AXS results – Level 31

Figure A.42. AXS results, column TC01, Level 31, Q1 Tower.


316



317



318


Figure A.49. AXS results, core wall TW18, Level 31, Q1 Tower.


319



320



321



322



323



324




325



326



327



328


329

A.3 Axial shortening results versus predictions

A.3.1 Level B2 columns and core walls – ACI 209R-1992 model

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000

Age of Member (days)

Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening

Shrinkage Shortening

Reinforcement Restraining

Q1 Measured Strains (TC01-B2)

Total Shortening (AEMM+Superposition)



209R-1992 material models)

330

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








331

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








332

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








333

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








334

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








335

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








336

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening



Total Shortening (AEMM)

Q1 Measured Strains (TW09-B2)

Figure A.76. Measured vs predicted axial shortening (TW09-B2: ACI 209R-1992

material models)

337

A.3.2 Level B2 columns and core walls – AS3600-2001 model

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening







2001 material models)

338

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








339

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








340

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening



Q1 Measured Values



material models)

341

A.3.3 Level B2 columns and core walls – AS 3600-2009 model

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening




AXS (AEMM+Superposition)


Figure A.81. Measured vs predicted axial shortening (TC01-B2, TC02-B2: AS 3600-


342

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








343

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening






Figure A.83. Measured vs predicted axial shortening (TC05-B2, TC06-B2 : AS 3600-


344

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








345

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








346

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








347

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








348

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening



Q1 Measured Values


Figure A.88. Measured vs predicted axial shortening (TW09-B2 : AS 3600-2009

material models)

349

A.3.4 Level B2 columns and core walls – Best fitted model

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening






350


selected material models)



351

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








352

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








353

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








354

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








355

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening








356

-200

0

200

400

600

800

1000

1200

1400

1600

0 200 400 600 800 1000


Ax

ial S

ho

rte

nin

g (

mic

ros

tra

in)

Elastic Shortening

Creep Shortening



Q1 Measured Values (TW09-B2)


Figure A.96. Measured vs predicted axial shortening (TW09-B2: Best selected

material models)

357

A.4 Axial shortening field data sheet

Figure A.97. A field worksheet used for recording of in-situ data reading.

358

A.5 Views of Database for storage of data reading

A typical input sheet as used in the database is shown in Figure A.98. In this sheet,

nine values of DEMEC gauge readings taken on 3rd

November 2004 on the south side

(station 1) of column TC04 located on basement level B2 (column/wall ID, TC04-B2)

which are stored in the 2000th

records are presented. The temperature and the relative

humidity in the environment measured when DEMEC gauge reading taken was 25.1

°C and 71.9% respectively. The DEMEC gauge no.2 was used together with INVAR

bar no.2 giving the bar reading value of 7.96. The name of DEMEC gauge operator

was Sadudee Boonlualoah. For this record, it was able to obtain data reading from

most available points except point no.6 which was temporarily inaccessible due to

construction circumstances.

After a record is completed, data readings and other information are transferred into

an individual query that contains all data readings from a similar column/wall ID that

also have similar station number. A Microsoft Access window that show list of some

queries actually used to store data reading obtained from various instrumented

location is shown in Figure A.99. It can be seen clearly that DEMEC gauge readings

were conducted in many columns and core walls of the building at different floors.

Also, in some places, data readings were taken from more than one side of the same

member.

359

Figure A.98. A typical format of data reading input for the database.

360

Figure A.99. List of queries for data reading at locations

Figure A.100. A typical view of data sorted by a query.

361

A.6 Views of spreadsheet for calculating and graph plotting of measured axial

shortening

Figure A.101. A typical view of column details input in “Cover” worksheet.

362

Figure A.102. A typical view of “Data” worksheet.

363

Figure A.103. A typical view of “Cumulative” worksheet

Figure A.104. A typical view of “Cumulative2” worksheet.

364

Figure A.105. A typical view of “Graph A” worksheet.

365

Figure A.106. A typical view of “Graph B” worksheet.