Design criteria for low binder SCC, ECO-SCC

Design criteria for low binder Self-Compacting Concrete, Eco-SCC by Florian V. Mueller

Thesis submitted to the School of Science and Engineering at the Reykjavik University in partial fulfilment of the requirements for the degree of Doctor of Philosophy May 2012

Thesis Committee: Olafur H. Wallevik, Supervisor Professor in Civil Engineering, Reykjavik University & ICI Thorsteinn Ingi Sigsússon, CEO ICI Professor of Physics, University of Iceland Jon E. Wallevik Dr. Ing., Senior Research Engineer at ICI Kamal H. Khayat, Examiner Professor, Missouri University of Science and Technology

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Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC

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ICI Rheocenter


Preface This thesis is part of a larger research project called “Eco-SCC” initiated by Olafur H. Wallevik at the former Icelandic Building Research Institute. This institute was merged with the Technical Institute in 2008 to form the Innovation Center Iceland (ICI). In collaboration with Reykjavik University, the ICI Rheocenter has been created as an excellence center to conduct research related to the rheology of building materials. During my tenure in Iceland, I had the honour to join the studies at Innovation Center Iceland/Reykjavik University/ICI Rheocenter and learn in the process. This thesis reports my own contributions to the project, starting from 2006 when also considering results of my first Master’s thesis (concomitant to the Diplomarbeit for my German degree as Diplom-Bauingenieur (civil engineer)) which were partially reviewed in the light of the projects’s later findings and incorporated in chapters 3.1.1. and 3.2.1.. Additionally, the experiences I made with other projects at the ICI Rheocenter increased my knowledge in the field of rheology and Self-Compacting Concrete of various composition types. Thereby it influenced the research approach and finally the outcome of this PhD thesis. The sponsors of these projects should be gratefully acknowledged hereby. As in all theses, there remain topics not fully covered. Chapter 11 provides an overview of suggested future research. Nevertheless, one hopes that the findings in the area of challenging mix compositions for SCC can contribute to the development of a sustainable building product and broadens the applications of low-binder Self-Compacting Concrete.

Florian V. Mueller Nidau, Bern, in March 2012

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Acknowledgements I gratefully acknowledge the financial support and the provision of concrete composites for the project “Eco-SCC” at ICI Rheocenter, Reykjavik University, Innovation Center Iceland by several different companies (in alphabetical order): CEMEX Research Group AG, Switzerland CP Kelco (USA) Íbúðalánasjóður (Iceland) Landsvirkjun (Iceland) Steypustöðin (Iceland) Rannsóknasjóður Vegagerðarinnar (Iceland) Steinsteypunefnd (Iceland) Tækniþróunarsjóður (Iceland) Norstone (Norway) W.R. Grace (USA)

Additionally, I would like to especially thank my thesis supervisor and superior during my tenure at the ICI, Professor Olafur H. Wallevik. Through knowing him, I encountered rheology as a very useful and sophisticated tool to optimize binder material and special concrete mix designs. His knowledge and tasks challenged me and thereby improved my own knowledge in the field of rheology, concrete mix design and durability. Without his support, I definitely would not have taken this educational step, with all the positive consequences it has on my current profession and even private life. I personally feel honoured that Professor Kamal H. Khayat joined the steering committee of my thesis a an external examiner. His opinion about the content and quality of my work is very valuable to me. Therefore, I would like to thank him for taking the time and effort to work through this comprehensive thesis. Furthermore, I would like to thank Jon Elvar Wallevik for his feedback concerning rheological questions and the special care he always takes with his answers in this complex topic. The responsible persons of the anonymised “The Company” should be acknowledged for their support and publishing agreement concerning the concrete mixtures of chapter 3.2.2.. Additionally, the co-workers and fellow students at the institute should be acknowledged for their contributions to the project. Elisa C. Stead should be acknowledged for her valuable editorial support. Finally yet importantly, my family should be acknowledged for their support during my studies. I also wish to thank my fiancée Lena for everything.

Abstract The objective of this thesis is to establish mix design criteria for an economical and environmentally friendly Self-Compacting Concrete (SCC), here named Eco-SCC. The innovative concept results in mixtures with a cement and powder content that is considerably lower than in normal SCC, therefore achieving a low carbon footprint. In the development of Eco-SCC, a sophisticated rheological approach was used to establish the optimized mix designs, as rheology is the science of flow and deformation of matter. Furthermore, a rheograph was used as a tool to select chemical and mineral admixtures for the suspension, which is considered here as matrix volume that surrounds the aggregates beyond 0.125 mm, and how this should be composed when air is entrained or not. Thereby, two different approaches could be identified that achieve the desired matrix volume for sufficient blocking characteristics and to demonstrate rheology parameters of the fresh concrete indicating self-compactability. One approach is introduced as Air Matrix Approach (AMA) and the other is regarded as Solid Matrix Approach (SMA). The segregation resistance of Eco-SCC is restricted with the aid of a specially designed particle size distribution (PSD), verified by newly designed (static) sedimentation tests. The homogeneity of hardened concrete specimens is determined at different sample height sections, which can be affected by separation processes such as dynamic segregation and static sedimentation. The thereby determined specific gravities (SG) at the different height section are assessed and compared with theoretical sedimentation criteria computed using the mix design characteristics of the mixture. Finally, the variation of SG allows a ranking in seven Classes of Specific Gravity Distribution (CSGD) that were developed in this thesis to evaluate the sedimentation of hardened concrete, thereby assessing their stability. The other approach evaluates the appearance of coarse aggregate particles at the top part of the specimen; a concept introduced here as Aggregate Surface Appearance Index (ASAI). Since free flow without blocking and the required lubricant volume are functions of the particle packing (PP) characteristics of the granular mixture, the impact of altering PSD was evaluated for one representative aggregate type using the shear compaction in a gyratory Intensive Compaction Tester (ICT). For selected aggregate proportions, the impact of particle shape on PP was studied using spherical or cubic coarse aggregate fractions. Two parameters were identified that are sufficient for describing the loose bulk and the consolidation effort characteristics, thereby characterising the PP of one specific granular mixture. Comprehensive tests revealed superior robustness of Eco-SCC versus powder-rich SCC. The thereby experienced shear thickening materials behaviour of the powder-rich SCC was linked to rheology parameters, matrix volumes, and sedimentation stability. In addition, a model was derived to estimate the yield value of fluid concrete and SCC, based on a stable slump flow value. Correlations between results in a coaxial cylinders viscometer (ConTec Viscometer 5) and a portable impeller-based device (ConTec Rheometer-4SCC) were established.

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List of content Abbreviations

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Definitions

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1.

Introduction .............................................................................................................................................. 1 1.1. Overview on mix design approaches for SCC ....................................................................... 2 1.1.1.

Rational Mix Design approach by Okamura et al............................................................. 5

1.1.2.

Particle packing approach ..................................................................................................... 6

1.1.3.

Methods considering the role of aggregates to a bigger extend ..................................... 7

1.1.4.

Statistical approach ................................................................................................................ 8

1.1.5.

Matrix model approach ......................................................................................................... 9

1.1.6.

Rheology approach .............................................................................................................. 11

1.2.

Stability considerations ............................................................................................................ 13

1.3.

Testing SCC - An introduction .............................................................................................. 14

1.3.1.

Material models in fluid mechanics ................................................................................... 15

1.3.2.

Time dependent phenomena ............................................................................................. 17

1.3.3.

Combined phenomena ........................................................................................................ 18

1.4.

2.

Rheometry – background knowledge.................................................................................... 20

1.4.1.

Impeller-based approach .................................................................................................... 21

1.4.2.

Steady Couette flow between coaxial-cylinders .............................................................. 22

1.4.3.

Steady torsion flow in a horizontal gap ............................................................................ 27

1.4.4.

Statistical considerations ..................................................................................................... 28

1.5.

Particle packing ......................................................................................................................... 31

1.6.

Research needs .......................................................................................................................... 37

Materials and methods........................................................................................................................... 39 2.1. Materials ..................................................................................................................................... 39 2.1.1.

Binder materials.................................................................................................................... 39

2.1.2.

Binder material used in this thesis ..................................................................................... 49

2.1.3.

Admixtures ............................................................................................................................ 52

2.1.4.

Summary of admixtures used in this thesis...................................................................... 58

2.1.5.

Aggregates ............................................................................................................................. 59

2.2.

Test methods ............................................................................................................................. 64

2.2.1.

Rheometry ............................................................................................................................. 64

2.2.2.

Intensive compaction tester ............................................................................................... 66

2.2.3.

Compressive strength .......................................................................................................... 68

2.2.4.

Calorimetry............................................................................................................................ 68

2.2.5.

Time dependent volume changes ...................................................................................... 69

2.2.6.

Determination of freeze-thaw-resistance ......................................................................... 72

2.2.7.

Sedimentation and segregation .......................................................................................... 73

2.2.8.

Robustness ............................................................................................................................ 85

2.3. 2.3.1.

For Chapter 3 ....................................................................................................................... 90

2.3.2.

For Chapter 4 ....................................................................................................................... 92

2.3.3.

For robustness evaluation................................................................................................... 95

2.3.4.

Miscellaneous mixes ............................................................................................................ 96

2.4. 3.

3.1.1.

Using different cements .................................................................................................... 101

3.1.2.

Partial replacement of OPC with CaCO3 filler ............................................................. 106

3.1.3.

Partial replacement of OPC with silica fume ................................................................ 113

3.1.4.

Optimization of rheometry setup.................................................................................... 118

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Varied paste rheology due to variation of admixture ....................................................... 128

3.2.1.

Effect of different admixture combinations .................................................................. 128

3.2.2.

Varied type and dosage of stabilizer ............................................................................... 131

3.3.

5.

Mix and measurement procedures ......................................................................................... 98

Effect of paste characteristics ............................................................................................................101 3.1. Varied paste properties due to altered binder composition............................................. 101

3.2.

4.

Mix designs ................................................................................................................................ 89

Concluding remarks ............................................................................................................... 137

Effect of matrix volume ......................................................................................................................138 4.1. Matrix approaches SMA and AMA ..................................................................................... 138 4.1.1.

Effect on fresh concrete properties ................................................................................ 138

4.1.2.

Effect on hardened concrete properties ........................................................................ 142

4.2.

Effect of coarse aggregate fraction ...................................................................................... 151

4.3.

Example with 220 kg/m3 OPC and Limestone filler ....................................................... 160

4.4.


Gyratory particle packing ....................................................................................................................164 5.1. Preliminary studies ................................................................................................................. 164 5.2.

Variation of PSD .................................................................................................................... 166

5.3.

Variation of aggregate source ............................................................................................... 170

5.4.

Validation of modelling ......................................................................................................... 171

5.5.

Extension of granular skeleton with binder used in Eco-SCC ....................................... 173

5.6.

Miscellaneous .......................................................................................................................... 179

5.7.


Segregation and sedimentation ..........................................................................................................183 6.1. Segregation tests in fresh concrete ...................................................................................... 183 6.2.

Sedimentation tests in hardened concrete .......................................................................... 185

6.3.

Proposal for stability-based mixture-opimized rheograph .............................................. 192

6.4.

Concluding remarks ............................................................................................................... 196 Florian V. Mueller

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC 7.

Robustness ............................................................................................................................................198 7.1. Evaluations in fresh concrete ............................................................................................... 198 7.2.

Discussion of apparently shear-thickening binder-rich SCC........................................... 205

7.3.

Effect in hardened concrete ................................................................................................. 209

7.3.1.

Hardening monitoring with heat of hydration .............................................................. 209

7.3.2.

Compressive strength ........................................................................................................ 210

7.3.3.

Sedimentation resistance................................................................................................... 213

7.3.4.

Drying shrinkage ................................................................................................................ 215

7.4. 8.


Correlation of different rheology devices and tests ........................................................................217 8.1. Slump flow............................................................................................................................... 217 8.2.

Coaxial-cylinders viscometer versus impeller-based ......................................................... 226

8.2.1.

Yield stress and G-value ................................................................................................... 226

8.2.2.

Plastic viscosity and H-value ............................................................................................ 230

8.3.

Validation of the proposed correlation models ................................................................. 234

8.4.


9. Final remarks.........................................................................................................................................239 10. Conclusions ......................................................................................................................................244 11. Outlook into future research .........................................................................................................246 References......................................................................................................................................................228 Appendix Chapter 2......................................................................................................................................243 Appendix Chapter 3......................................................................................................................................259 Appendix Chapter 4......................................................................................................................................260 Appendix Chapter 5......................................................................................................................................265 Appendix Chapter 6......................................................................................................................................270 Appendix Chapter 7......................................................................................................................................297 Curriculum vitae............................................................................................................................................309

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Abbreviations ACI

American Concrete Institute

AEA

Air-entraining admixture

ASTM

American Society for Testing and Materials, US-American industrial standard

ASAI

Aggregate Surface Appearance Index

b

Binder material

B

Ground basalt filler, sometimes used to save space in legends, but generally GBF

b.-%

Percentage related to binder material by weight, e.g. for admixture dosages

bibm

The European Precast Concrete Organisation

c

Cement according to EN 197-1

c.-%

Percentage related to cement by weight, e.g. for admixture dosages

CVC

Conventional vibrated concrete

CSGD

Classes of specific gravity distribution

CEMBUREAU The European Cement Association

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DafStb

Deutscher Ausschuss fuer Stahlbeton (German, German Committee for Reinforced Concrete Structures)

DIN

Deutsche Industrie Norm (German industrial standard)

ECC

Easy Compacting Concrete

EFCA

The European Federation of Concrete Admixture Associations

EFNARC

The European Federation of Specialist Construction Chemicals and Concrete Systems

EN

European Norm (i.e. Standard)

ERMCO

The European Ready-Mix Concrete Organisation

eφ

Effort Coefficient to vary the solid concentration φ of the logarithmic model equation describing the densification of a bulk of solids

FA

Fly ash

f0

Rotational frequency [rps]

fib

Fédération Internationale du Béton (French) (International Federation of Structural Concrete)

FS

Fine sand used as fraction of aggregates Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC GBF

Ground basalt filler

G

Emperical yield stress value (corresponds to yield value of a viscoplastic Binghamfluid when obtained in a ConTec Rheometer-4SCC)

G1

Abbreviation used for coarse aggregate (gravel), e.g. number one

H

Emperical plastic viscosity (corresponds to plastic viscosity of a viscoplastic Bingham fluid when obtained in a ConTec Rheometer-4SCC)

HVSI

Visual stability index in hardened concrete, proposed by Lange et al. [204]

hj

Blocking step (height) of slump flow with J-ring according EN 12350-12

JRMCA

Japanese Ready-Mix Concrete Association

L

Limestone filler

OPC

Ordinary Portland Cement, i.e. CEM I according to EN 197-1

RILEM

Réunion Internationale des Laboratoires et Experts des Matériaux, systèmes de construction et ouvrages (French), International Union of Laboratories and Experts in Construction Materials

S

Slump value according ASTM C143/C143M-10a

S1

Abbreviation used for sand fraction of aggregates, e.g. number one

SCC

Self-Compacting Concrete, sometimes called: self-consolidating concrete

SCM

Supplementary Cementitious Materials

SG

Specific gravity, also known as relative density

SF

Silica fume

Sf

Slump flow value according to ASTM C1611 /C1611M-09b

Sfj

Slump flow value with J-ring according to ASTM C1621/C1621-M09b

SP

Superplasticizer, dispersing admixture, also known as High-Range-WaterReducing-Admixture (HRWRA)

ST

Stabilizing admixtures, also known as Viscosity Modifying /Enhancing Admixtures (VMA, VEA)

T

Torque

TSC

Total solid concentration of liquid admixtures after drying at 105°C

VA

Volume of air

VAGG

Volume of aggregates

VP

Volume of powder

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VW

Volume of water

VW/VP

Volume proportion of water to powder

VEA, VMA

Viscosity-enhancing admixture, viscosity-modifying admixture, both are synonyms for stabilizing admixtures

VSI

Visual stability index according to ASTM C1611/C1611M-09b

Vol.-%

Volume percentage

w

Water

w/c

Weight ratio of water to cement according EN 197-1

w/cm

Weight ratio of water to cementitious materials, without considering the strength equivalence of the materials used, also known as w/b

w/p

Weight ratio of water to powder materials, without considering the strength equivalence of the materials used

Greek letters

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𝛾

Shear velocity

𝛾̇

Shear rate

ε

Void content

ζ

Relative specific gravity

η

Apparent viscosity

μ

Plastic viscosity of a Bingham fluid

ρ

Density

τ

Shear stress

τ0

Yield stress value of a Bingham fluid

φ

Solid concentration

φi

Initial Solid Concentration

φn

Solid concentration after a number n of gyratory cycles

ω

Angular velocity

Florian V. Mueller

ICI Rheocenter


Definitions Some words and phrases are used differently in the literature, so the following section lists how the terms are used in this thesis: 1. Binder

All materials allowed in composite cements according to EN 1971, used in this thesis without considering maximum allowed contents of each material

2. Burst slump flow

The slump flow of a clearly segregated mixture in which the inner part, containing the coarse aggregates and mortar, is surrounded by a distinct segregation halo of paste, equivalent to VSI ≥ 2

3. Dynamic sedimentation

Dynamic force-driven particle migration, e.g. centrifugal forces and acceleration forces during batching and handling of fresh concrete (commonly: dynamic segregation /sedimentation)

4. Hyper-fluid paste

SP-oversaturated paste that exhibits segregation

5. Matrix

Aggregate-embedding lubricant phase, consisting of paste, air, and fine material from aggregates smaller than 125 μm. Hereby the aggregates would be considered as being larger than 125 μm

6. Paste

Mixture of binder and water

7. Powder

Particles with a maximum particle size below 125 μm, independent of its reactive nature1

8. Relative specific gravity

Specific gravity of one particular part i (SGi) of the cut sedimentation specimen related to the median SG of all members of the cut specimen

9. Segregation

Inhomogeneity and instability including bleeding for fresh concrete with insufficiently high SP-dosage for one specific water content and w/b

10. Specific gravity

The ratio of the mass of a substance to the mass of a given reference material (in this case water) occupying the same volume, also known as relative density

11. Static sedimentation

Gravity force-driven settlement of larger particles when fresh concrete is at rest

The term powder is understood in this thesis in the same broad meaning as the general description in the SCC guideline of the Japanese Society of Civil Engineers [172] implies, beyond the five materials mentioned in as example, such are: cement, granulated and ground slag, fly ash, silica fume and limestone filler. 1

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1. Introduction Eco-SCC is a mix design concept for Self-Compacting Concrete (SCC) for which the first approach was described in public by Professor Olafur H. Wallevik in 2003 [116]. It is considered an economical and ecological alternative to common Self-Compacting Concrete (SCC) and conventional vibrated concrete (CVC). The ecology aspect might be more considered for SCC, as binder content will be drastically reduced and therefore the carbon footprint per volume unit concrete. The economic aspect might be considered for both, SCC application through the reduction of cement, which is costly, and CVC, through taking advantage of using Self-Compacting Concrete. Only in 2009, a definition was made considering the content of cementitious material2 used in a concrete mixture [363]. The restrictive definition was established since some other approaches also claimed “Eco” benefits, but only considered cement clinker content and not any filler material. The objective of the Eco-SCC approach is to reduce the paste volume as such. Our approach aims for the application in ready mix concrete, which requires a certain workability retention and robustness. Furthermore, the common concrete strength class3 of C25/30 should be enabled with the mix design. This could broaden the applications for Self-Compacting Concrete and remove it from the special concrete application, the corner in which SCC has been placed in most countries around the world, and allow SCC to be utilizedas a general concrete. Our approach does not mean to replace high-strength SCC that may have to penetrate densely reinforced structures during the casting and might reveal architectural surface finishing. No concrete of such low strength class as we are aiming for is designed to perform in such a manner. The intention of this project is to develop the knowledge and tools to design low-binder SCC as a robust solution for the average daily work concrete, which is in most countries somewhat similar to a C25/30 strength class concrete. Starting from this, it is relatively easy to enhance some properties in order to meet specific durability or strength requirements simply by adapting the mix design accordingly. Usually such enhanced requirements are at the cost of higher powder content due to relatively reduced volumetric ratio of water/binder (Vw/Vb) and the use of supplementary cementitious materials in the binder. This thesis aims to increase the background knowledge of parameters that affect the production of concrete with self-compacting properties by experimenting with mix design parameters on the edge of current usage. The effect of the matrix, varied in composition and total volume, is analysed in fresh concrete with a rheology approach. Rheology parameters are determined with a stationary coaxial-cylinders viscometer as well as a portable impeller-based system that supports the more simplified information extracted from the common slump flow test. Two different approaches to obtain matrix volumes will be proposed and used as a design approach concept, namely the Air Matrix Approach (AMA) and the Solid Matrix Approach (SMA). They vary in the composition of the matrix volume that is used to retrieve sufficient blocking and filling ability with minimum binder content. Such variations require some adjustments of the mix design and will have an effect on the rheology parameters obtained in concrete. The interaction of these variations with some hardened concrete properties, such as compressive strength, homogeneity, shrinkage, and freeze-thawresistance is made. The role of aggregates, in particular their packing characteristics as a function of particle size distribution (PSD) and particle shape is analysed through the rheology of fresh concrete and in measurements with a gyratory Intensive Compaction Tester (ICT). The sedimentation of solids is analysed for different PSD’s and matrix designs and a new approach to quantify static The term “cementitious materials” refers to all materials that are allowed to be combined in order to compose a “cement” according to EN 197-1, which includes the different reactive SCM’s but also inert filler material. 3 The term “compressive strength class” always refers to the ranking system according to EN 206. 2

sedimentation is utilized. This is based on a measured and therefore objective criterion, which is obtained with relatively little effort. It utilizes the variation of specific gravity (also called as relative density) due to sedimentation. For the first time for such stability rankings, the rating is extended to cover the quality of self-compacting properties including not-yet self-compacting, i.e. easy compacting concrete (ECC). The robustness of one suitable low-binder (Eco-) SCC is tested with respect to variations of ±10% and 20% of superplasticizer and ± 5 and 10 kg/m3 of water. The response on these variations is compared with those obtained from one lean SCC and one binder-rich SCC, using the terms according to our definition of SCC mixtures in [363]. Based on the results from the different fields of measurement, a recommendation for a refined workability box (rheograph) of SCC is proposed, also considering the different matrix methods, robustness, and stability aspects.

1.1.

Overview on mix design approaches for SCC

The major step in the mix design of Self-Compacting Concrete came when the team working with Professor Okamura introduced their concept of self-compacting high-performance concrete [243, 244, and 245]. Previous applications of concrete casted without compaction energy have been reported by other researchers in the 1970´s and 80´s for some rare cases, e.g. Collepardi [71 and 72] and Sharp & Watson [300], mostly for underwater concreting. This is indicated by several statements mentioned e.g. in [300], such as “the originally intended vibration does not needed to be executed” and the described effect of excessive bleeding water “that had to be pumped away to allow surface finishing-off to proceed”, which has to be avoided in a designed SCC. Nevertheless, the publications have not triggered a systematic SCC mix design approach for many years. The Rational Mix Design Approach (RMDA) by Okamura et al. [243, 244, and 245] was established in the booming period of 1980’s Japan, when high-rise buildings and industrial and infrastructure structures demanded big volumes of high strength concrete (HSC). It was later renamed to High Performance Concrete (HPC) when the overall enhanced durability features were also discovered and emphasised (in [9]). Accordingly, this new approach was considered first as self-compacting high performance concrete. Due to the relation between strength and w/c, the most effective way to increase strength considerably was (and still is) to lower the w/c. To satisfy the lubricant volume demand for workability, such HPC requires a larger amount of binder material and the additional aid of dispersant admixtures4 to achieve a workable concrete mixture. A combination like this, with a high powder content and low w/p results in a very viscous concrete mixture that is barely workable. In order to ensure the consistent quality of the hardened HPC structure, the team came up with an approach to produce a concrete that was very fluid, did not require any additional consolidation energy and could fill even densely reinforced structures completely. Their approach is based on excessive amounts of powder and a very low water content causing enough inner resistance (i.e. plastic viscosity) against sedimentation to stabilize the coarse aggregates.

Dispersant admixtures are also called in accordance to the water reducing capacity for similar consistency (i.e. slump) and ranked as water reducer (WR), mid-range water reducer (MRWR) and high-range water reducer (HRWR), concomitant to fluidizer (FM), plastisizer (P), and superplasticizer (SP). 4

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC Subsequent recommendations for SCC mix design often rate the approaches according to the amount of binder (powder) used, for instance [171]: 1. Binder type SCC; 2. Stabilizer (or VMA5) type, having “lower” content of powder and significant contribution from any Stabilizer/VMA; and 3. Mix type, having “intermediate” content of powder and the aid of VMA. A review by Domone [91] about SCC mixtures used for larger volume applications with different mix designs, revealed the average powder content used in bigger projects for SCC (Figure 1-1) is about 500 kg/m3 and only very few mixtures contained below 400 kg/m3 at all. It is noteworthy that the use of VMA reduces the powder content only by about 50kg/m3 in average for the mixes being richer on binder (Figure 1-2). Then, the w/p is between 0.26 and 0.48 and the paste volume about 30% to 42% with the majority in the middle 30th. This usually leads to relatively high strength, even when a considerable content of inert filler is used.

Figure 1-1: Powder content for case studies [91] considering aggregate type and maximum size

Figure 1-2: Powder content for case studies [91] considering VMA usage

The ICI Rheocenter classification of SCC, first reported during a conference in 2009 [363] is given in Table 1-1. The majority of recently accomplished SCC, i.e. already somehow optimized, would be rated into group two and three. According to my experience, only the last group requires a stabilizer in its mix design, whereas it is optional in groups three and four in order to enhance the robustness of the concrete mixture [31, 32, and 33]. Of course, such low powder content would require special emphasis on the selection of aggregates, in particular particle shape and particle size distribution. At several occasions the superior homogeneity, and therefore reduced fluctuation of properties, of a stable SCC compared with fluid CVC was observed, e.g. in [355]. Thereby the acceptance of SCC might be increased when this thesis can provide an approach to produce ready-mix concrete in the lower to medium strength class, about C25/30. Some requirements for several different composites and a range of suitable contents of each of them will be proposed as one outcome of this thesis. Often one refers to Viscosity Modifying Agent (VMA), or also Viscosity Enhancing Agent (VEA). Experience shows that a stabilizer does not always result in varied plastic viscosity, which is the type of viscosity one experiences during handling and casting of concrete. Instead, often the apparent viscosity at a specific shear rate is altered, which basically affects the yield value when concrete is considered as a Bingham fluid and thereby increasing it. Therefore, the more general term Stabilizer (ST) should be favoured. 5

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They are based on several test methods applied in order to meet the different mandatory characteristics, which are described in chapter 1.2. Table 1-1: ICI Rheocenter classification of SCC [363]

Classification 1 Rich SCC 2 regular powder content SCC 3 Lean SCC 4 Green SCC 5 Eco-SCC

Cementitious powder content ≥ 550 kg/m3 500 +/- 50 kg/m3 415 +/- 35 kg/m3 350 +/- 35 kg/m3 ≤ 315 kg/m3

Range of powder or from 550 kg/m3 and upward or from 450 to 550 kg/m3 or from 380 to 450 kg/m3 or from 315 to 380 kg/m3 up to 315 kg/m3

To provide an economical alternative to SCC in its common design methods, see chapter 1.1., usually excessive amounts of reactive and/or inert filler are used to replace the cement clinker, e.g. (in alphabetical order) Bouzoubaa & Lachemi [47], Bosiljkov [48], Buchenau [57], Chodbury & Basu [67], Esping [100], Felekoglu [104], and Ferraris & Hill [110]. These materials are considered “economical” due to their usually lower prices. Additionally, the hydration induced thermal stresses can be lowered due to the differences in reactivity of the various SCM and OPC. In hardened state, their effect on microstructure and porosity of hardened concrete can potentially enhance various durability aspects, e.g. Locher [208], Mehta & Monteiro [222], Stark & Wicht [320], and Stark [324]. Sometimes it can be an economically beneficial solution for a concrete producer to use alternative materials from various waste recovery systems. Such materials are not considered in this thesis. Such materials usually have variable properties and can therefore reduce the robustness of a mix design. Due to costs arising from the use of admixtures at all, it has to be calculated for the specific local conditions whether the concept of Eco-SCC would represent an economical advantage compared to CVC that in some countries can be produced with relatively small contribution from labour costs. Several approaches to reduce the cement content were made in recent years. Very often it is based on replacing cement by other filler types, whether of reactive nature such as fly ash (FA), granulated and ground blast furnace slag (GGBS), or (quasi-)inert filler such as limestone filler6. These three mentioned are likely the most common used in SCC in bigger quantities. Other additives, such as silica fume (SF), Metakaolin (Mk), ultrafine limestone (i.e. calcium carbonate) filler (UCC) and others are used in smaller quantities as they have significant impact in small dosages. Rice-producing countries sometimes add rice husk ash, see e.g. [293], and ground stone powder, see e.g. [57 and 130], is employed occasionally. Nevertheless, one can find little information about approaches that aim to reduce the total powder content at all, for instance Su et al. [326 and 327] and Brouwers et al. [53, 54, (including Hunger [163])]. Occasionally, approaches aim to optimize the packing characteristics of the powder fraction, such as Fennis-Huijben [105] and Geisenhanslüke [131 and 296], the latter focusing more on UHPC. Of course, the paste particles are important to consider, as they are numerously present in concrete and have thereby an important impact. On the other hand, it is the whole sum of aggregate particles that require a specific lubricant volume for a specific consistency. Therefore, one approach could be to optimize the aggregate gradation. This is in order to reduce the powder content needed for both to obtain sufficient lubricant volume for the free flow of the concrete and also to provide stability. Sometimes the term “quasi-inert” is used (Stark & Wicht [319]), as CaCO3 can contribute with some dissolved ions, see e.g. Bonavetti et al. [50], Cyr eet al. [77], Matschei et al. [219], and Stark & Wicht [319], and therefore contributes actively on a chemical level to the hydration process. 6

4

Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Different mix design methodologies are briefly discussed in the following chapters.

1.1.1.

Rational Mix Design approach by Okamura et al.

The Rational Mix Design Method proposed by Okamura et al. [243, 244, and 245] is the first published approach to actually design a SCC, which differs from the common approaches for conventional concrete. They explain that a self-compacting mechanism can be achieved by reducing the content and maximum diameter of the coarse aggregate fraction and providing the mortar with certain properties/composition parameters, such as low Vw/Vp (1a) and high SP dosage (1b) and limited fine aggregate content (2). The mode of action is explained for the first (i.e. (1a) and (1b)) as to provide compatibility of deformability and viscosity and for the second (2) as to lower the pressure transfer. The mix design procedure is summarized in the following: 1. A specific air volume (VA) is chosen. Remark VFM: This might sometimes be based on freeze-thaw-resistance considerations, but probably for such mixes more importantly - to reduce plastic viscosity and thereby make the otherwise very viscous mixes workable. 2. The coarse aggregate volume is defined with about 50% of all the solids. 3. The fine aggregate volume is defined as 40% of the mortar volume. Hereby, particles less than 0.09 mm are not considered as aggregates but as powder. 4. Determining the water-powder-ratio by volume (Vw/Vp) according to the fluidity criterion of the paste with 0.9-1.0 as first approach. 5. Evaluation of the superplastizicer dosage. The evaluation of a suitable Vw/Vp according to this method is highly affected by the choice of content as well as type of binder materials and aggregate fines. It is first examined in a micro mortar including all the binder and powder material (i.e. solids less than 90 μm) of the intended mix design, which then needs to be adjusted for the final concrete solution. The parameters used for the Vw/Vp evaluation and SP dosage determination are the relative flow area Γm (Equation 1-1) and the relative funnel speed Rm (Equation 1-2). Equation 1-1 2

𝑑 Γ𝑚 = ( ) − 1 𝑑0 Here, 𝑑 is the averaged spread diameter out of two measurements that accounts for the largest dimension and the direction normal to it, and 𝑑0 the outlet diameter of the cone used. A relative flow area of about five is recommended which is equivalent to a spread of about 245 mm for the cone geometry suggested. Equation 1-2

𝑅𝑚 = 10/𝑡 [s] Here, t is the flow time out of a mortar funnel in order to distinguish it from the relative flow time of concrete Rc out of a concrete V-funnel. For the latter an optimum range between 0.50 and 1.0 has been suggested [243, 244, and 245], which is equivalent to flow times between 10 and 20 seconds and indicate rather viscous mixes.

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In order to quantify the volume of water necessary to lubricate the paste until the fluidity satisfies the SCC criterion, tests are recommended for one fixed powder content under fixed SP dosage. Then, one determines a band of different relative flow areas due to variation of water content only. The dimensionless βp is equal to a relative flow area of zero, e.g. the state before which self-initiated flow can occur. The deformation parameter Ep depends on the solid concentration and the effects from particles, such as surface charges, surface texture, and particle size distribution, and has to be determined experimentally. Equation 1-3

𝑉𝑤 = 𝛽𝑝 + Γ ∗ 𝐸𝑝 𝑉𝑝 A simplified version of this approach is recommended by the JRMCA [171] as “Standardized Mix Design Method of SCC”. In this method, but also in the approach recommended by the JSCE [172], the use of fine particles from the aggregates seems not to be considered anymore, as the definition of “powder” no longer includes it, and it is used instead to represent binder material. This change in the definition can already represent several dozen kilograms of powder material per cubic meter that is doubled by the use of binder material, in addition to the aggregate fines.

1.1.2. Particle packing approach Many researcher discussed the importance of particle packing (PP), often interpreted as particle size distribution (PSD), on the rheology of concrete, e.g. DeLarrard [85], Khayat et al. [180]. As an example, Su et al. [326 and 327] consider the importance of packing characteristics of solids on the flowability of a granular mix in the presence of a lubricant. To do this, they employ a packing factor (PF) of the aggregates as ratio of mass when tightly packed versus mass of loosely packed aggregate. They argue that with higher PF less binder would be needed due to higher content of aggregates within a specific volume. They found the fluidity, as their measure for self-compacting ability, reduces with increasing PF and thereby the compressive strength. The content of coarse aggregate (i.e. gravel) by weight (Wg) is determined according to Equation 1-4. The content of fine aggregate (i.e. sand) by weight (Ws) is determined according to Equation 1-5. In these equations, the weight proportion of the sand fraction (S) to total aggregate content (a) is considered. Equation 1-4

𝑆 𝑊𝑔 = 𝑃𝐹 × 𝑊𝑔𝐿 (1 − ) 𝑎 Equation 1-5

𝑊𝑠 = 𝑃𝐹 × 𝑊𝑔𝑆 ×

𝑆 𝑎

Remark VFM: Unfortunately, no information is given by the authors about the compaction procedure to obtain the tightly packed aggregates when determining the PF. One can see later (Chapter 5) that it has an impact on the result of the packing and thereby influences the whole mix design. They then apply the strength consideration according to experiences in CVC to determine the cement content and w/c. Here they use OPC only but do not consider composite cement as used in Europe. After this step, they utilize FA and/or GGBS in order to provide stability (i.e. by increas6

Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC ing plastic viscosity) and obtain the fluidity desired to satisfy the fluidity criterion for SCC according to the JRMCA recommendation. Since this approach aims to achieve high fluidity, self-levelling and filling ability, one will always create mix designs with relatively high powder content. Nevertheless, the powder content (i.e. binder content according to our definition) seems to be reduced compared to the original Rational Mix Design Method, see the discussion by Brouwers et al. [53 and 54]. Additionally, the approach to utilize SCM will exhibit considerable strength increase at later ages as well as a porosity refinement. This can be beneficial for durability aspects, even though it seems not to be considered for the design of concrete structures and, therefore, it would not be economical. When this approach was applied by Brouwers & Radix [53 and 54], they claimed to have “lowpowder” SCC when retrieving “only” 480kg/m3 of cement and limestone powder, while not considering the powdery particles from the aggregates as originally suggested in the Rational Mix Design Method. Professor Hwang and his co-workers developed their Densified Mix Design Algorithm (DMDA) approach originally for HPC, but also adapted it for SCC [64, 166, and 167]. It employs fly ash to densify the bulk of solids in order to reduce cement (clinker) volume required to satisfy the fluidity criterion. Since they still obey the same fluidity criterion as for highly fluid SCC and the Fuller curve (see Table 1-4 at page 32) for the aggregates, significant powder content has to be used to ensure stability since it increases plastic viscosity. More recently, Fennis-Huijben [105] studied the use of particle packing for ecological concrete by replacing cement with suitable filler material with a different origin. The studies on paste and mortars were used to extend De Larrard´s Compressible Packing Model (in [85]) to the CompactionInteraction Packing Model, including the effect of interparticle forces such as Van der Waals forces and electrostatic forces. As before, one could consider such an approach as a valuable way to reduce the use of cement clinker, thereby reducing the carbon footprint and maybe the unit price of concrete. On the other hand, one can be of the opinion that such an approach should always be combined with an approach to reduce the paste volume in general, which thereby reduces the lubricant demand of an optimized aggregate’s PSD.

1.1.3. Methods considering the role of aggregates to a bigger extend The role of aggregates was frequently studied in the past for conventional vibrated concrete: In the timeline of its appearance, important publications were produced by e.g: Fuller & Thompson [121], Abrams [1], Graf [133], Talbot & Richart [328], Furnas [122 and 123], Anderegg [14], Powers [264 and 272], Dunagan [93], Kennedy [176], De Larrard [85], Shilstone [302],. Andersen & Johansen [15], Mørtsell [224 and 226], Quiroga & Fowler [274], Rached at al. [275]. Since the particle gradation changes for several reasons in SCC, new PSD approaches had to be identified, which are briefly introduced in the following chapters. Extensive work on the theoretical background of particle lattice effect and segregation resistance was done by Bethmont et al. [26, 27, and 28], but here the work did not result in a mix design approach and is therefore not considered in this chapter.

1.1.3.1. Recommendation by ICAR at University of Texas The team around Professor Fowler [189], for SCC, in particular Koehler [192], put forward an approach to consider aggregate characteristics and their importance for an optimized mix design. They recommend applying a 0.45 power curve for the gradation, and approaching a relatively minimum paste volume with consideration of particle shape and angularity of the aggregates. The ratio

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by mass of sand/total aggregate should be in the range between 0.4 and 0.57, which should lower the paste demand to lubricate the surface area of solids. The void content is determined with dry rod consolidation of a granular skeleton. The paste volume is given in a range of 28% to 40% while the w/p (i.e. w/b according the definition used in this thesis) is in the range of 0.30 to 0.45, and could be higher when VMA is used. They do not provide information how much the w/p could be increased then, but a review of their publications [189, 190, and 192] reveals that they still have considerable powder (i.e. binder) content in order to provide sufficient lubricant volume at low w/p. They further suggest increasing the paste volume in order to increase robustness after the paste volume necessary for passing ability and filling ability has been quantified. The argument that higher paste volume would increase robustness is brought forward occasionally, e.g. Chodbury [67] and Hoeveling [155], despite the findings of others, e.g. Billberg (and coworkers) [31, 32, 33, and 34] , Lowke et al. [212], Rigueira et al. [280] and Hoeveling [155]. It is a proven fact, that the proportional increase of larger aggregate in the PSD would lower the paste volume covering it. On the opposite, the paste volume needs to exhibit sufficiently high viscosity for stabilizing the even larger volume of coarse particles, which only can be obtained with a low w/p, as what they propose when ranging it between 0.30 and 0.45. This implies the utilization of relatively large powder quantity in order to obtain the volume required as lubricant.

1.1.3.2. Dutch recommendation After significant work was done in the Netherlands in the 1990’s and early 2000’s by the team around Professor Walraven at Delft University [366], Professor Brouwers and his team from TU Twente suggest a different approach that puts more emphasis on particle packing in order to reduce paste demand [53, 54, 55, 163, and 165]. They recommend applying a particle gradation curve for the aggregates following the modified Andreasen & Andersen model (see Table 1-4 at page 32) and actually take the fines from aggregates into account, as already suggested in the Rational Mix Design Approach from Okamura and co-workers. Unfortunately, they operate with a PSD containing 44 discrete sizes for only the paste (i.e. up to 0.125 mm), and additional sizes for coarser aggregates. This implies a very theoretic approach as such a variety of sizes cannot be controlled in operations with reasonable effort. However, more importantly, they apply the fluidity criterion that was derived for powder-rich SCC originally, so significant powder content will still be required in order to satisfy the criteria concerning fluidity, blocking, and stability.

1.1.4. Statistical approach Khayat et al. [179 and 181] and Sonebi [314] propose a proportioning method with the aid of statistical design software. Analogous to the rheology response for different mix design parameters and composites, the effect would be quantified in various one-point measurements. They are to describe the different fresh concrete properties necessary to obtain a SCC, such as fluidity, filling ability, stability, and compressive strength. Therefore, slump flow and, if available, rheometric measurements are recommended to determine the fluidity and self-compactability. V-funnel flow and filling box8 are recommended to evaluate the passing ability. The surface settlement test to evaluate the static segregation resistance is further mentioned.

The ASTM and ACI considers 4.75 mm as the boundary size between sand and coarse material, differing from the European system in which 4.0 mm is used. 8 One should notice that the filling height of the various filling boxes recommended in different guidelines (Caisson test and Kajima box) differ and therefore influence the hydrostatic energy level, which might influence the results due to varied shear stresses. The results are also determined differently then. 7

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Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Since starting the alteration from one initial mix design, the method depends greatly on having experience with the specific materials used and the response for a specific range of its application. The variance between results of different mix design parameters indicate the significance of each single parameter for a range of mix designs. It allows describing the field of properties one can expect for the major mix design using small numbers of mixes, without establishing a series for each single parameter. The results mentioned in [179, 181, and 314] indicate that rather low powder content can be reached with such an approach.

1.1.5. Matrix model approach Mørtsell [224 and 226] introduced a concept, somehow similar to De Larrard [85], about estimation the workability of conventional and high performance concrete from matrix characteristics and its rheology, considering a flow time. For the latter, a modification of the mortar flow test called FLOWCYL [225] is applied. Smeplass and co-workers tried to adapt this approach for lightweight aggregate concrete [310] and Self-Compacting Concrete [253 and 311]. In this concept, concrete is considered as a two-phase system. The workability of concrete, expressed as slump value, is related to: 1. The properties of the particles; 2. The properties of the matrix; and 3. Their volume proportioning. The matrix phase is defined then as water, possible chemical additives, and all fines, including cement, pozzolans, and aggregate fines considering particles less than 125 μm. The particle phase is defined as aggregate particles larger than 125 μm. Therefore, the approach is called Particle-Matrix Model (PM-Model). The underlying model calculations require two main parameters, the flow resistance ratio of the matrix (λQ) and the air void modulus of the particles (Hm). The first parameter is determined in a modified mortar cone test, now called FlowCyl [225], in which the emptying rate of the matrix material out of a cylinder is measured with a permanent recording balance. The procedure is explained in detail in [225 and 310]. It is assumed that the result of this test provides reasonable correlation to the plastic viscosity, which they consider as the important parameter for the workability of the matrix and the subsequently produced concrete. Mørtsell uses the term “workability” as a synonym for fluidity, or more precisely – the slump value. Without commenting on it for the moment, I will just follow his definitions in this chapter during which I introduce the concept very briefly. He uses a workability function (Kp) (Equation 1-6) to describe the matrix volume and rheology properties required to obtain a specific slump value. Equation 1-6

𝐾𝑝 = (𝑛 − 𝑚) ∗ With:

n, m x Fp α

ICI Rheocenter

[tanh(𝑥) + 1] +𝑚 2

Lower and upper asymptote values; The workability function argument, see Equation 1-7; The volume fraction of the matrix; The steepness coefficient of the workability function, see Equation 1-8. Florian V. Mueller

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Equation 1-7

𝑥 = 𝛼 + (2𝐹𝑝 − 1 − 𝛽) β

“Zero-slump” argument according to Equation 1-9.

Equation 1-8

∝= 𝑘1 ∗ 𝑒 −𝑘2 𝜆𝑄 Equation 1-9

𝛽 = 2 ∗ 𝐻𝑚 − 1 + 1/𝛼 k1, k2 λQ

Constants found by regression analysis of a convenient data series; The flow resistance ratio.

Including all the equations above and rewriting Equation 1-6 gives the workability function in another form (Equation 1-10). Equation 1-10

𝐾𝑝 = (𝑛 − 𝑚) ∗

[tanh(2𝛼(𝐹𝑝 − 𝐻𝑚 ) − 1) + 1] +𝑚 2

One example is provided in [310] that show the different matrix volume demands for the same aggregate characteristics, but when using different matrix properties. Since aspects necessary to be considered for SCC are not accounted hereby, such as blocking behavior and stability, the matrix volumes presented in this figure are only valid for conventional vibrated concrete consistency classes (Figure 1-3). Nevertheless, it shows a general tendency that one also can observe in SCC.

Figure 1-3: Slump workability function for two concretes based on the same particle system, but different matrices (from Smeplass et al. [310])

Mørtsell [226] defined the air void space ratio (H) as the ratio of void contents of a loose bulk to the void content of the densified bulk of particles, compacted in a particular procedure see more in detail [226], or in English in [310]. It is determined separately for the fine aggregates, between 0.125 mm and 4 mm, and the coarse fraction and calculated simply by Equation 1-11.

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Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Equation 1-11

𝐻 = 1 − 𝜌𝑏𝑢𝑙𝑘 /𝜌𝑝 With:

ρbulk ρp

Bulk density at each densification step (6 á 5 sec); Particle density

The air void modulus (Hm) (Equation 1-12) is based on the air space ratio of the fine (0.125-4 mm) and the coarse (>4mm) aggregate, while using the fineness modulus (FM) and the fine/coarse aggregate distributions, see detailed in [226], or in English in [310]. Equation 1-12

𝐻𝑓 𝐻𝑐 𝐻𝑚 = 𝑣𝑓 ( 0.5 + 𝑇𝑓 ) + 𝑣𝑐 ((𝐹𝑀 )0.5 + 𝑇𝑐 ) 𝑐 (𝐹𝑀𝑓 ) With:

vf Vc Hf Hc Tf Tc FMf FMc

Volume fraction of fine aggregates (0.125-4 mm); Volume fraction of coarse aggregate (> 4 mm); Air void ratio of fine aggregate; Air void ratio of coarse aggregate; Calibration parameter for fine aggregate; Calibration parameter for coarse aggregate; Fineness modulus of fine aggregate; Fineness modulus of coarse aggregates.

When they were trying to adapt the PM-model to SCC [311], they found the measurement of the matrix properties in the FlowCyl test, which they related to the viscosity of the matrix, was “apparently insensitive” to the rheological differences of the matrix and the results in concrete. Therefore, they attributed the yield stress differences as important for the fluidity of SCC. This is later confirmed with observable correlations between yield value and slump (slump flow respectively), see e.g. DeLarrard [85], Flatt et al. [119], Hu et al. [162], Koehler et al. [192], Roussel [288], and Wallevik [347], and the less accentuated correlation between plastic viscosity and slump flow velocity. Additionally, the acceptance criteria, what they considered with 650 mm, could relatively easy cause misleading results, as for a very viscous mix such flow might be rather low and on the lower edge of the fluid properties necessary to obtain self-compacting properties. On the opposite range, such flow can exceed the capacity of low-viscous mixes without segregation. The flow target would need to be adapted considering the stability for different rheology parameters. In another attempt made in 2003 [253], the rheology of the matrix was measured in more advanced rheometric devices in addition to the FlowCyl test, using the parallel plate viscometer (Physica) from Anton Paar and a coaxial-cylinders viscometer from ConTec. For both devices, the plastic viscosity was found to coincide very well, whereas the yield values apparently vary considerably in the order of magnitude. The parallel plate device reveals only very small changes while the coaxial-cylinders viscometer reveals notable variations of the yield value. They noticed a certain relation of rheology parameters of the concrete to the ones found in the matrix and realized the importance of the matrix volume as such for the concrete rheology.

1.1.6. Rheology approach Professor Wallevik suggests a rheological approach to design SCC [355]. Hereby, a concrete mix design, which satisfies the design considerations for strength and durability requirements, is tested in a rheometric device so one can describe the rheology parameters that govern workability (i.e. ICI Rheocenter

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yield value and plastic viscosity) and fluidity (i.e. yield value). Since they are also used in this thesis, the devices used are explained in detail in chapter 2.4 and the underlaying fluid mechanic background in chapter 1.3. Additional tests to quantify the blocking and filling ability are required in order to satisfy the specific requirements for one application. Commonly, the tests are also accompanied by slump and slump flow measurements. If necessary, mix design parameters, such as aggregate gradation or paste volume, and its proportioning are adapted until the requirements are met. Therefore, the rheology response of different mix design composites and parameters may have to be quantified if they differ from common experience, as for instance is given in [4 and 356]. For the major property, the self-compaction, one considers the location of yield value and the plastic viscosity in one figure, called a rheograph. The position of these properties in such a rheograph also indicates the workability and consistency of the fresh concrete, see Figure 1-4 for the broad range of consistencies and Figure 1-5 for SCC and highly workable concrete.

Figure 1-4: Rheograph, indicating the rheology parameters for different concrete applications and consistencies, after Wallevik [356]

Figure 1-5: The area of rheograph for SCC new proposed by Wallevik & Wallevik [365]

Different devices are available optimized for a range of applications, so one can optimize a binder composition in paste or mortar, using a Rheomixer or a Viscometer 6 then, before one starts to analyse the full SCC mixture in a device suitable for concrete such as the Viscometer 4 or 5. Such a procedure is usually applied at the ICI Rheocenter and therefore it is the approach followed in this thesis.

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1.2.

Stability considerations

Saak et al. [290, 291, and 292] put special emphasis on a segregation-resistant mix design for SCC. Therefore, they introduced the concept of a rheological self-flow zone (SFZ) obeying a Segregation Control Theory (SCT). Following this, the parameters to provide sufficient fluidity have to be in balance with the parameters that they determined control static and dynamic segregation (Equation 1-13). For static segregation9 they consider the gravitation force Fgrav (Equation 1-14) and the buoyant force Fbuoy (Equation 1-15) acting as the major forces involved. Equation 1-13

𝐹𝑑𝑜𝑤𝑛 = 𝐹𝑔𝑟𝑎𝑣 − 𝐹𝑏𝑢𝑜𝑦 Equation 1-14

𝐹𝑔𝑟𝑎𝑣 = 𝑔𝜌𝑝 𝑉𝑝 With g being the gravitational constant, ρp the particle density and Vp the particle volume. Equation 1-15

𝐹𝑏𝑢𝑜𝑦 = 𝑔𝜌𝑚 𝑉𝑝 With g being the gravitational constant, ρm the matrix density and Vp the particle volume. Applying Equation 1-14 and Equation 1-15 in Equation 1-13 results in Equation 1-16. Equation 1-16

𝐹𝑑𝑜𝑤𝑛 = 𝑔(𝜌𝑝 − 𝜌𝑚 )𝑉𝑝 = 𝑔∆𝜌𝑉𝑝 For the drag force, as the restoring (i.e. settlement restricting) force, they apply a proportional relation to the yield shear stress τ0 of the matrix acting on the surface area of the particles. Here a cross-sectional area of the particle Ap has been employed. Finally, they suggest a yield stress limit to prevent static sedimentation, obeying the terms of Equation 1-17. Equation 1-17

4 𝜏0 ≥ 𝑔∆𝜌 𝑟 3 Remark VFM: In this equation, they consider the cement paste as the matrix for any aggregate where the static segregation needs to be restricted. Particle lattice effects due to optimized packing, that additionally can restrict the settlement action, are not considered. For the case when low shear rates applied overcome the yield shear stress and induce self-flow of a non-Newtonian fluid Saak et al. consider the viscosity and drag forces (Equation 1-18) as restricting parameters. Equation 1-18

𝐹𝑑𝑟𝑎𝑔 = 𝐶𝐷 𝜌𝑚 9

𝑣2 𝐴 2 𝑝

Here the term “sedimentation” is often used instead.

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in which CD is the drag coefficient and ν is the constant terminal falling velocity of the particle. Combining the equations Equation 1-16 and Equation 1-18 and re-arranging it one can express the terminal velocity according to Equation 1-19. Equation 1-19

8 𝑔∆𝜌𝑟𝑝 𝑣 = √( ) ( ) 3 𝐶𝐷 𝜌𝑚 Then they apply a relationship between Reynolds number (Re) and the drag coefficient (CD) for a sphere as formulated by Equation 1-20. Equation 1-20

𝑅𝑒 =

2𝑟𝑝 𝜌𝑚 𝑣 𝜂

Saak et al. refer to η as “the viscosity of the matrix” but it remains unknown whether they mean the plastic viscosity (μ), as it would be measured for a band of shear rates according to the Bingham fluid model, or apparent viscosity (η) occurring at the shearing state caused by the settlement velocity.

1.3.

Testing SCC - An introduction

For conventional vibrated concrete, the fresh concrete performances varied in a certain range, so consistency-related measurements were established that could sufficiently describe these performances, as e.g. the slump test according to ASTM C 143 and EN 12350-2, or see for other workability tests ACI 238.1R-08 [4]. When Self-Compacting Concrete was introduced, its performance was beyond the boundaries of the established tests and therefore suitable methods and procedures were needed in order to evaluate the main mandatory characteristics: 1. Self-compacting, i.e. full consolidation without external vibration energy10; 2. Filling capacity, i.e. a formwork that normally includes reinforcement bars, is completely filled without honeycombs and voids; 3. With mitigated blocking, i.e. for a defined content of reinforcement, the aggregates should not hinder the free flow due to blocking at obstacles; 4. With limited segregation, i.e. the mix design composition is so balanced that the composites remain homogenous and are not affected by bleeding, usually caused by a relative oversaturation of superplasticizer; 5. With limited settlement, i.e. the solid particles remain homogenously distributed in the fresh concrete, and therefore also during the hardening. Gravity induced settlement is mitigated. After much research work was conducted, dealing with the development and verification of suitable test methods, a few tests have been recommended by different research teams and organizations such as ACI [3], SCC European Project Group [102], JRMCA [171], and JSCE [172]. Some of the Usually one refers to the term consolidation “by its own weight“[171]. This holds true for casting processes based on free flow. For other batching methods, the energy levels contributed by gravity, velocity and height differences need to be in equilibrium in order to ensure full consolidation while still mitigate sedimentation. 10

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ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC tests are already incorporated in normative guidelines dealing with testing of fresh concrete in different countries. Most of the tests mentioned in these references provide (at least) one result which correlates with one of the intrinsic physical properties necessary to describe the fluid SelfCompacting Concrete in an adequate manner, see for instance in [108, 206, 238, 284, 285, 288, and 347]. The possibility to measure the different intrinsic fluid properties at once can only be done with rheometric methods that apply fluid mechanics.

1.3.1. Material models in fluid mechanics Fluid mechanics are used as the base to quantify the rheology of fluid fresh concrete and thereby describe their workability. It is a subclass of the general continuum mechanics dealing with fluids and how they mechanically behave. Fluids are defined as materials that have a definite volume but not a definite form. This is in order to distinguish it from solid and gaseous materials [170]. The ability of a fluid to form relatively easily into different shapes is due to the distinct anisotropic response to applied stressest. The background knowledge for fluid mechanics will not be repeated here in detail, as it is described elsewhere in numerous publications, as for instance: Bowen [49], Fox et al. [120], Irgens [170], Mase & Mase [216], Tattersall & Banfill [329], and Wallevik [345]. An important equation to consider in fluid mechanics is the Cauchy’s equation of motion (Equation 1-21), as it is one of the base equations of fluid mechanics, and is given by the following: Equation 1-21

𝜌

𝑑𝜐 = −∇𝑝 + ∇ ∗ 𝜏 + 𝜌𝑔 𝑑𝑡

The term 𝜌 is the fluid density, p is pressure, 𝜏 is the extra stress tensor, v is velocity and g is gravity. For Generalized Newtonian Fluid (or GNM), the extra stress tensor is given by Equation 1-22: Equation 1-22

𝜏 = 2𝜂𝜀̇ where the term 𝜂 is the apparent viscsoity (also known as shear viscosity) and 𝜀̇ is the rate of deformation tensor given by Equation 1-23. Equation 1-23

1 𝜀̇ = (∇𝑣 + ∇𝑣 𝑇 ) 2

1.3.1.1. Shear rate and von Mises shear stress for general flow For an incompressible material the von Mises shear stress (Equation 1-24) can be derived from the rate of deformation tensor (see for example Wallevik [348]. Equation 1-24

𝜏 = √𝜏𝑖𝑘 𝜏𝑖𝑘 ⁄2 Likewise, the shear rate for an incompressible fluid is given by Equation 1-25.

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Florian V. Mueller

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Equation 1-25

̇ 𝛾̇ = √2𝜀̇𝑖𝑘 𝜀̇𝑖𝑘 By combining the two above equations, the following can be attained: Equation 1-26

𝜏2 =

𝜏𝑖𝑘 𝜏𝑖𝑘 2𝜂𝜀̇𝑖𝑘 2𝜂𝜀̇𝑖𝑘 = = 𝜂 2 2𝜀̇𝑖𝑘 𝜀̇𝑖𝑘 = 𝜂 2 𝛾̇ 2 2 2

The above equation can be written as follows: Equation 1-27

𝜏 = 𝜂𝛾̇ Which is basically a shear stress equation.

1.3.1.2. Non-Newtonian fluids In general, all materials that do not obey the ideal behaviour of a Newtonian fluid are categorized as non-Newtonian. These variations can be categorized as shear rate dependent, time dependent, time independent with a plastic deformation part, and combinations thereof. In aggrement with the approach in Irgens [170], these fluids are described by the three main groups: 1. Time independent fluids 2. Time dependent fluids 3. Viscoelastic fluids Without going into detail, one example of special interest for cementitious materials within the first group are viscoplastic fluids, in which the Bingham fluid is the most used in the concrete industry. In terms of shear shar stress, it is given by Equation 1-28. Equation 1-28

𝜏 = 𝜂𝛾̇ = (𝜇 +

𝜏0 ) 𝛾̇ = 𝜇𝛾̇ + 𝜏0 𝛾̇

Another example is fluids with a non-linear respond to steady variation of the shear rate (so called: power law fluid behaviour), obeying e.g. the Herschel-Bulkley fluid model (Equation 1-29). Equation 1-29

𝜏 = (𝑘𝛾̇ 𝑛−1 +

𝜏0 ) 𝛾̇ = 𝑘𝛾̇ 𝑛 + 𝜏0 𝛾̇

When a disproportional increase of shear stresses is observed (n>1), the fluid is considered as shear thickening, whereas when disproportional decrements of the responding shear stresses are observed (n 𝜙𝑏 𝑛𝑐𝑑,𝑠 = 𝑚𝑎𝑥{6; 12} = 𝑓(𝐹𝑣 , 𝐹ℎ , ℎ𝑣 , 𝑓𝑜 )

𝑚 = 𝑚𝑎𝑠𝑠 𝜌𝑏𝑑 = 𝑚⁄𝑉𝑏𝑑 𝜀𝑏𝑑 = 1 − 𝜌𝑏𝑑 ⁄𝜌𝑝 < 𝜀𝑏 𝜙𝑏𝑑 = 1 − 𝜀𝑏𝑑 > 𝜙𝑏 𝑛𝑐𝑑 = 𝑚𝑎𝑥{2; 12} = 𝑓(𝐹𝑣 , ℎ𝑣 , 𝑓𝑜 )

Figure 1-17: Regular packing structures of equal spheres [in 218], i.e. one monosized fraction of spheres, including void content and coordination numbers.

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1.6.

Research needs

Based on the literature review it becomes evident that the definition of SCC and its mix design approaches were established for relatively binder-rich SCC. Apart from sufficient self-leveling and blocking criteria, both describing the filling ability of a mold for a reinforced concrete structure for one particular concrete mixture, the fluidity in terms of slump flow is widely considered as acceptance criteria for SCC. This value is an indirect measurement of the rheology properties of a mixture. Existing attempts to optimize the SCC mixtures still require addition of suitable filler material, commonly accepted as binder material in concrete, in order to meet this fluidity criterion. Applying a rheology approach intead, as is done in this thesis when measuring the intrinsic rheology parameters directly with e.g. a ConTec Viscometer 5, allows for optimizing the mix design without considering the fluidity as a criterion. Then the proportions of the rheology parameters (yield value and plastic viscosity) are employed, which indicate fluidity, workability, stability, and selfcompactability. In order to adapt the fluidity according to the requirements of one specific casting condition, one would need to alter the mix design accordingly, probably by adding more powder into the system and reducing the w/p further to maintain the stability of the fresh mixture. One result of the relatively high powder content used in conventional mix design approaches for SCC, together with the relative low w/p required for stability reason, is relatively high strength. The current working practice is to order concrete according to its designed compressive strength, usually determined at the age of 28d. Common powder content in SCC, even when using supplementary material, often results in exceeding the strength classes used for common concrete application, even if it is only at later age due to lower reactivity of supplementary cementitious materials. In such a case, even the deformation characteristics can be affected due to the relatively high paste volume and the relative slow developing modulus of elasticity, the one property of the concrete to resist the deformation in concrete, i.e. creep and shrinkage. Therefore, the objective of Eco-SCC is to investigate the mix design parameters necessary to allow reducing the paste content, while still obtaining a stable self-compacting concrete mixture. When aiming for a binder content of about 315 kg/m3, equivalent to 10 vol.-% of total volume, the additional lubricant volume to establish the excess paste for a free movement (i.e. flow) of the SCC would need to be obtained by utilizing water, entrained air, and/or additional powder commonly not considered as cementitious material, such as fines from aggregates. The rheology response of water and air would be a reduction of the plastic viscosity, the property commonly attributed to maintaining the stability of a mixture against settlement of larger particles. Therefore, the sedimentation stability needs to be assessed in order to maintain the homogeneity of the fresh concrete. Figure 1-18 summarizes the different topics this thesis emphasizes.

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Binder composition, Ch.3.1 Paste design & properties Admixture type & dosage Ch 3.2 Matrix proportion with solids (SMA) or accounting for entrained air (AMA), Ch. 4.1

Stable Self-Compacting Concrete

Matrix design & properties

Effect of course aggregate Ch 4.2

Particle packing in gyratory Intensive Compaction, Ch. 5

Example with very low content of OPC, CH 4.3

Sedimentation resistance, Ch.6 Robustness of ECO-SCC, Ch.7

Establishing suitable setups, Ch. 3.1.4

Discussion of shear thickening binder-rich SCC, Ch.7.2

Rheometry & Rheology Linking yield stress and slump flow, Ch. 8.1 Linking results from portable Rheometer.4SCC with stationary Viscometer 5, Ch. 8.2. Figure 1-18: Organogram of topics emphasised in this thesis

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2. Materials and methods The evaluations in this thesis consider different cementitious materials and admixtures. The focus is on fresh concrete properties and rheology aspects in particular, also considering sedimentation stability. In general, the compressive strength is determined for all mixes. For selected mixes, investigations concerning shrinkage and freeze-thaw resistance are obtained in addition. The measurement approach applied in this thesis can be summarized as follows: Slump flow value (Sf) Slump flow

Fresh concrete

Visual Stability Index (VSI) Slump flow with Jring (Sfj)

Viscometer

Blocking step (hj)

Rheology

Rheometer4SCC Visual Stability Index in Hardened Concrete (HVSI)

Mix design

Sedimentation

Hardened concrete

Compressive strength Drying shrinkage

Aggregate Surface Appearance Index (ASAI) Specific Gravity Distribution (SGD)

Classes of SGD Relative Specific Gravity (ς)

Freeze-thawresistance Figure 2-1: Organogram of tests applied in this thesis

The different tests and materials used will be introduced in the following chapters, also including some background knowledge for the different fields.

2.1.

Materials

2.1.1. Binder materials In general, the binder material used in Self-Compacting Concrete does not vary from that commonly used in conventional vibrated concrete. Only the dosage varies due to usually increased powder demand in order to obtain sufficient fluidity while sedimentation resistance needs to be ensured. The different amounts and proportions influence several fresh concrete parameters and thereby affect hardened concrete properties as well.

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Binder-related topics of major concern in fresh SCC are: 1. 2. 3. 4.

Fluidity and maintaining it for the desired time range; Limitation of bleeding; Segregation stability 21 and; Robustness of the mix design.

For hardening concrete and in particular SCC of special concern is: 1. Heat of hydration, as it is important to avoid cracking due to differences in heat induced stress distribution that is caused by high powder content at low w/b. 2. Strength development, i.e. sufficient early strength to allow for de-moulding and progressing with the construction process.

Remark: In general, the higher binder content and lower w/b employed in SCC develops considerable strength at early age, unless the mixture is not retarded considerably because of the admixtures used, or due to the approach to reduce clinker content by substituting it with material of relatively reduced reactivity. For hardened concrete: 1. The strength development in absolute and relative terms is important, in order to meet the strength requirements coming from the structural design and the construction process, and to establish restrain elements against longterm deformation (shrinkage and creep). 2. Maturity related development of microstructure in order to meet specific requirements concerning various durability aspects. Within this thesis, the term “binder material” is refered to as material that can be used according to EN 197-1 to produce composite cement, without considering the volume restrictions of EN 197-1 for each composite. In SCC, it is increasingly common not only to use cement (OPC and composite cement) but also filler of various origins and reactivity capacities in larger quantities. The ASTM has a similar system of composite cements with its ASTM C595. In addition, the culture differs quite a lot throughout the world as how to account for cement and filler material. One approach is commonly regarded as the composite approach whereas the other is based on (blended) cements. In some countries, e.g. Australia, USA, and UK, it is rather common to use a cement with relatively high clinker factor in concrete, but add SCM directly into the concrete at the mixing facility, whereas it is rather common e.g. in Europe to use blended cements with lower clinker factor instead and avoid adding filler materials. Clinker factor refers to the content of plain Portland Cement Clinker in composite cement. Then, the SCM has already been added in a cement plant. As a result, the use of the term “cement” can mean different things in different places, in particular concerning the content of clinker in a concrete mix. In ideal circumstances of a performance based mix design approach, the powder part would be proportioned, taking different materials, in order to meet specific properties in fresh and hardened state. The next chapter will briefly introduce each binder material used during the project. Therefore, the list does not claim to be complete for all existing binder materials.

2.1.1.1. Cement in the EN 197 A binder material called “opus cementitium” was used in the Roman Empire consisting of ground brick and reactive stone i.e. natural pozzolans such as volcanic ashes (in [319]). The cement, as we Segregation used in this context is related to admixture-binder interaction and possible oversaturation of SP, causing bleeding and segregation. 21

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC know it today, being the result of a burning process of suitable Calcium, Silicate, Aluminate and Ferrite sources (which are the main constituents of the actual (cement) clinker), was patented to Joseph Aspdin in 1824. Most likely his son William Aspdin created the first Ordinary Portland Cement (OPC) in 1843 (in [319]), that is burnt until sintering temperature. Due to different local traditions and availability of suitable raw materials around the world, but also to satisfy demands coming from different applications, the cement clinker can vary in its chemical composition whereas the main clinker constituents are usually present as shown in Table 2-1. Minor components are incorporated in the crystal structure of the different clinker phases that also affect their reactivity and the hydration products developed. The minor components with the biggest effect are the alkalis (Na+ and K+). Free lime (CaOfree), Periklas (MgO), glassy phase and other minor metastable /intermediate phases can also be present in cement clinker [208]. Some of them can have a significant effect on chemical stability of different hydration products and usually their content ought to be limited. Table 2-1: Major clinker phases and some properties (after [319])

Phase

Dicalcium-silicate

Abbreviation

Alite Tricalciumsilicate C3S

C2S

Aluminate Tricalciumaluminate C3A

Oxides

3 CaO SiO2

2 CaO SiO2

3CaO Al2O3

Content in clinker Size in clinker

40-80% (60) 20-60 μm

3-15% (7%) μ-scale

Reactivity

High

Very high, needs to be controlled with sulphate

Slow, but needs to be controlled with sulphate

Early strength 28d strength Long-term strength

+++ ++ ++

2-30% (15) 10-30 μm Medium (depending on mineralogy and modification) + ++ +++

Ferrite Calciumaluminate-ferrite C2(A,F) 4CaO Al2O3 Fe2O3 4-15% (8) μ-scale

-/+ -/+ -/+

Hydration product

CalciumSilicateHydrates (CSH); Portlandit (CH)

++ -/+ -Trisulphate (Tricalcium-sulphatealuminatesilicate-hydrate), Monosulphate (Monocalciumsulphatealuminatesilicate-hydrate)

Name

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Belite

Calcium-SilicateHydrates (CSH); Portlandit (CH)

Florian V. Mueller

Aft-and Afmphases (aluminate-ferritetrisulphate/ monosulphate)

41

Table 2-2: Strength characteristics [MPa] for cements according EN 197-1, tested on mortar according EN 196-1

Nominal strength class Strength development Minimum strength 28d Maximum strength 28d Strength at 2d Strength at 7d

35.5 N ≥32.5 ≤52.5

42.5 R ≥32.5 ≤52.5 ≥10

N ≥42.5 ≤62.5 ≥10

52.5 R ≥42.5 ≤62.5 ≥20

N ≥52.5 ≥20

R ≥52.5 ≥30

≥16

When cement clinker is considered for different strength classes according to EN 197-1, usually the ratio of the different clinker phases is adapted considering their reactivity, and the fineness obtained during the grinding process is varied. The smaller the particles, the faster the whole reaction can take place since more area is available for ions to dissolve in water and, thereby, to be available for further reactions. Therefore, EN 197-1 differentiates strength classes of any cement, whether it is an OPC or composite cement, which has been tested according to EN 196-1, shown in Table 2-2. According to EN 197-1, several different materials are allowed to formulate composite cements. They have to be processed together at the cement plant in order to meet the requirements for strength development and potential deleterious ion concentrations, such as of alkalis and chlorides. Then, the sulphate carrier also needs to be optimized for the specific composition. A sulphate source is added to control the reactivity of the aluminate phases. Minor additional constituents can be added without specific requirements as long their content does not exceed 5% in total and they do not interfere with the intentional performances. Commonly, additives are used to improve the manufacturing process of cement clinker production [208]. A negative effect on the clinker properties and performance then has to be mitigated. Most common additives are grinding aids to prevent agglomerations that can form during the grinding process. The most common grinding aid TEA (Triethanolamine) interacts with the ferrite phase and as zeta-potential enhancer [11], thereby potentially affecting the onset of hydration [201].

2.1.1.2. Reactive materials in the EN 197 and as filler in concrete Pozzolan materials can be used according to EN 197-1 when they conform to specific requirements regarding chemical composition and fineness. These affect the reactivity of the material. Named after potentially reactive material of volcanic origin from the Italian city Pozzuoli in Campania (in [319]), it represents materials that can hydrate after some activation and form C-S-H phases with dissolved calcium hydroxide (i.e. Portlandit). Considering the origin of the pozzolans, they are differentiated into natural and industrial by-product/processed pozzolans. Natural pozzolans, such as volcanic ashes, shale, clay or sedimentary rocks, would be abbreviated in the EN 197-1 with (P) or, when also naturally calcined, with (Q). They can be incorporated into composite cements in quantities up to a certain limit, described in Table 2-3. Table 2-3: Range of application for natural (P) and naturally calcined (Q) pozzolan according to EN 197-1

Portland-Pozzolan Cement CEM CEM II/A-P II/B-P (Q) (Q) 6-20% 21-35%

42

Portland-Composite Cement

Pozzolanic Cement

Composite Cement

CEM II/A-M

CEM II/B-M

CEM IV/A

CEM IV/B

CEM V/A

CEM V/B

6-20%

21-35%

11-35%

36-55%

18-30%

31-50%

Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Due to their limited local availability and increasing environmental protection regulations such natural pozzolan are likely more of local importance. On the other side, by-products from other industries can exhibit pozzolan reactivity and are therefore frequently used in cement or as filler in concrete. One advantage of the products being part of an industrial process is that one can already control their properties to a certain extent within the beneficiation process. Additionally, the use in concrete, where it can be beneficial for various aspects, provides an opportunity to dispose of otherwise troublesome materials. Fly ash “obtained by electrostatic or mechanical precipitation of dust like particles from the flue gas from furnaces fired with pulverized coal” [EN 197-1], can be used as industrially produced/processed reactive material. One differentiates siliceous fly ash (abbreviated with (V) according to EN 197-1 and Class F according to ASTM C618) and calcareous fly ash ((W) according to EN 197-1 or Class C according to ASTM C618). As explained for instance in Stark & Wicht [319], Class F fly ash can reveal pozzolan activity whereas Class C can also reveal hydraulic properties due to its higher content on reactive calcium oxide. The reactivity can be measured for instance according to ASTM C593-06. The use of fly ash as general filler material in concrete is allowed for material complying with EN 450-1. The main phases of siliceous FA consist of reactive silicon dioxide and aluminium oxide. Then the reactive calcium oxide should not exceed 10% and the free lime has to be restricted in order to avoid deleterious lime reaction (in [319]). Table 2-4: Range of applications for siliceous V and calcareous (W) fly ash in cements according EN 197-1

Portland-Fly ash Cement CEM CEM II/A-V II/B-V (W) (W) 6-20% 21-35%

Portland-Composite Cement

Pozzolanic Cement

Composite Cement

CEM II/A-M

CEM II/B-M

CEM IV/A

CEM IV/B

CEM V/A

CEM V/B

6-20%

21-35%

11-35%

36-55%

18-30%

31-50%

Fly ash has long been used as mineral filler material in concrete in order to meet special durability properties, for instance in HPC. Huettl confirmed recently [164] the work of Blaschke [43] or Uchikawa et al. [336] and others, that FA from anthracite source, i.e. class F fly ash, acts physically as filler material and as a nucleotide due to calcareous phases and the soluble glassy shell. At later stages, the pozzolan reactivity can be activated, mainly depending on the very high pH-value > 13 of the pore solution. Dissolved Ca(OH)2 only in the conditions occurring in cement paste cannot achieve such a basic environment alone, but it requires additional alkalis. When the conditions are not suitable, e.g. pH-value is low due to consumption of CH content caused by amorphous silica oxide beyond 10% and low alkali content, the pozzolan activity can be slowed down to a complete stop. The work of e.g. Biernacki et al. [35] supports such conclusions. Fly ash describes a group of materials that have in common that they are retrieved from the residue of coal burning furnaces. Therefore, it covers a wide range of not only chemistry composition (Table 2-5) but also varied physical properties.

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Table 2-5: Examples of main chemical composites in fly ashes

Chemistry SiO2 CaO Fe2O3 AL2O3 LOI ( 2% (AEA); VP < 135 l/m3, VW/VP ≈1.3, PSD- Mod. A&A 120

VA≤2% (no AEA); 135 160 l/m3, VW/VP ≈0.95-1.2, minimum μ when aggregates (round shape, lack of fines) causes low viscosity in mixtures with even higher VP

80

40

700 mm 750

800

0 0

20 40 60 80 100 120 Pl. Viscosity [Pa s]

(d) Binder-rich SCC, intermediate – high viscosity

VA≤2%; VP >> 160 l/m3, VW/VP ≈0.85-1.15, (JSCE compliance), PSDaggr ≈0.45 power curve

This approach does not differentiate powder-rich SCC further, as it is out of the application range of mixtures in this thesis and only few were done based on such conventional mix design. The recommendation describes rheology properties that are typically obtained with certain mix design characteristics. Considering minimum matrix volume for a specific matrix rheology, to ensure selfcompactability, sufficient blocking resistance and stability, these rheology areas should preferably only reached with such optimized mix designs. Only this would allow a minimum powder content, otherwise a larger powder content is needed to ensure stability of a mixture.

6.4.

Concluding remarks

Owing to the lower matrix volume of Eco-SCC, its sedimentation was found to be comparable, if not superior, to binder-rich SCC (Figure 6-31, p.193). Here a small alteration of the rheology properties can alter the sedimentation resistance considerably. The new test methods seem to deliver very accurate results, in particular the Classes of Specific Gravity Distribution (CSGD). This is based on measurements, without any judgments that might 196

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC be subjected to objective influences. It is therefore suggested to apply this method to more mixtures, in order to assess the statistical boundary conditions and to evaluate the usefulness /effectiveness in working practice. The geometry of the cylinder 150x300 mm, commonly used to assess some mechanical properties in hardened concrete, seems to deliver results as accurate as those obtained from columns of larger height up to 600 mm (Figure 6-7f, p. 186), as is recommended for sedimentation tests. This would require less volume and ease the handling and execution of the test, but without losing on the accuracy of the method. When linking the sedimentation stability with matrix properties, in terms of rheology, and its mix design proportioning, a proposal was made to refine the workability box “Rheograph” with several optimal zones, corresponding to minimum powder content for certain mix design conditions (Table 6-1, p.196). Then, not only the volume of powder VP is considered, as well as the air volume VA, or the volumetric proportion of water to powder VW/VP, but also the impact the aggregates with their PSD and shape have on concrete rheology and therefore stability.

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7. Robustness The robustness of three different general mix designs, representing low-binder (Eco-) SCC, lean SCC and binder-rich (br-) SCC, was evaluated with respect to the response of rheology parameters to the content alteration of some composites. These alterations considered the water content by ± 5 and 10 l/m3 (which is in the vicinity of about 3% and 6% for the mixes considered) and variations of superplasticizer dosages by ±10 and 20%62.The specific mix designs are provided in Table 2-44, p. 95 for the Eco-SCC, in Table 2-45, p. 96 for the Lean-SCC, and in Table 2-46, p. 96 for the binder-rich SCC. The first and the last contain about equal composites, despite the fly ash used to achieve the powder content needed for the binder-rich SCC. In the Lean-SCC, a different binder and admixture is chosen in addition to a different PSD of aggregates. The evaluations concern not only fresh concrete properties when assessing the rheology response on the intentionally altered mix design (Chapter 7.1), but also the impact such alterations have on various properties in hardened concrete (Chapter 7.3.) In addition, the shear thickening behaviour of binder-rich SCC is discussed and linked with sedimentation properties of these mixes (Chapter 7.2)

7.1.

Evaluations in fresh concrete

It appears that the slump flow is significantly affected by variation of SP for all mix designs. An increase can cause segregation, indicated by the double marks at higher SP-dosage for some mix designs, which represents the slump flow with and without segregation halo in Figure 7-1 and Figure 7-2. The reference’s slump flow is about 625 mm for the Eco-SCC, 640 mm for the lean SCC and 680 mm for the binder-rich SCC. Of course, the particular response for different binder compositions and SP types are unique, but the general trends appear as presented in Figure 7-1 and Figure 7-2. 30

SP_Eco-SCC 625 mm

120

SP_Eco-SCC

SP-lean SCC 640 mm 680 mm

80

SP_br-SCC

40 0 -40

-30

-20

-10

0

10

20

30

-80 -120 -160

Relative variation of SP [%]

Figure 7-1: Effect of varied SP-dosage on slump flow

Relative slump flow [%]

Slump flow difference [mm]

160

20

SP-lean SCC SP_br-SCC

10 0 -30

-20

-10

0

10

20

30

-10 -20 -30 Relative variation of SP [%]

Figure 7-2: Relative variation of Sf for varied SP-dosage

The reduced SP-dosage results in a stronger effect (i.e. greater reduction) on slump flow than the increased dosage enlarges the slump flow. It is worth mentioning that the segregation halo affects the appearance of the spread. The double marks at one specific SP dosage, or water content, indicate the paste (matrix) halo around the coarse aggregate spread. These composites were varied in the order of magnitude mentioned above in the mix design approach, but later only the recalculated values are considered, using the dosages of each composite that were actually added, and also the air volume measured. In particular an air volume different from the original assumption alters the content per one cubic meter. 62

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC Considering the effect of water, the variation of slump flow follows a more general response, as indicated by the different values appearing in a general scheme (Figure 7-3). Here, the surfaceactive considerations do not govern the rheology response, as they do for binder-admixture interaction, but rather the sheer volumes involved and therefore inter-particle distances are the essential factors. The variation can take place only until the oversaturation of free movable water results in a segregated slump flow. This is a characteristic value for each single reference mix design. The reference mix of the lean SCC might have been chosen on the border of its capacity to tolerate increase of SP or water, as indicated in both cases by the segregated slump flow occurring as soon as the dosage of one of the composites increases. 30

water-Eco-SCC water-lean SCC water_br-SCC

120 80 40 0 -40

-10

-5

0

5

10

-80 -120 -160

Relative slump flow [%]

Slump flow difference [mm]

160

water-Eco-SCC water-lean SCC water_br-SCC

20 10 0 -10

-5

0

5

10

-10 -20 -30

Relative variation of water [%] Figure 7-3: Effect of varied water dosage on slump flow

Relative variation of water [%] Figure 7-4: Relative variation of Sf for varied water content

The slump flow with J-ring reflects the blocking effect of obstacles on the free flow of coarse aggregates. This test was performed only for the most critical series, considering the smallest matrix volume of the Eco-SCC. It is common to determine the blocking step hj (Figure 7-5), but beyond this value one can also obtain the difference of slump flow with and without J-ring for simultaneously performed tests (Figure 7-6). 80

SP_Eco-SCC Water_Eco-SCC

20 10

SP_Eco-SCC Water_Eco-SCC

60

30

dSf-fj [mm]

Blocking step hj [mm]

40

40 20 0 -30

0 -30

-20 -10 0 10 20 30 Relative variation of composite [%]

Figure 7-5: Blocking step hj of both series in Eco-SCC robustness test mixes

-20

-10

0

10

20

30

-20 Relative variation of composite [%] Figure 7-6: Difference of slump flow with and without J-ring for Eco-SC robustness test series

The blocking step appears to be relatively high considering the limits established for binder-rich SCC, e.g. 20 mm [86] or even 10 mm [29]. One must consider the context of the differences in flow capability that can be obtained with the different types of SCC, which causes a differently shaped spread. The different slope of the spreads results in a higher blocking step measured, even though ICI Rheocenter Florian V. Mueller

199

the homogeneity due to blocking is not considerably affected, see e.g. Figure 7-7 and Figure 7-8. One can conclude that the acceptance limit of this test method should be reviewed for low-binder SCC. It can help to apply the slump difference of simultaneously performed tests with and without J-ring, see Figure 7-6. Here, three different possibilities can occur: A) The spread is approximately similar (with tendency of smaller Sfj); B) The spread of the J-ring is significantly smaller (i.e. more than 20 mm); or C) The spread of the J-ring test is slightly larger. The reason for the last is that a micro-mortar evolved through thinning out of coarse particles by a blocking effect gains more flowability, at least when it is in a hyper-fluid state (i.e. on the edge of segregation). A stable mix with sufficient matrix volume and dispersion state would reveal approximately similar spreads in both measurements. Mixes with a high content of aggregates, and where the matrix is too stiff for SCC, will reveal a larger difference in the spread diameters due to blocking of the whole concrete.

Figure 7-7: Appearance of slump flow with J-ring for minimum water content in Eco-SCC

Figure 7-8: Appearance of slump flow with J-ring for maximum water content in Eco-SCC

The rheology response to these variations was measured in the coaxial-cylinders viscometer, which reveals the underlying rheology parameters in addition to the slump flow value. The legends of the figures in Table 7-1 indicate a ranking about the sensitivity of each mix design to variations of superplasticizer or water. It appears that the SP has only a small effect on the variation of the references’ yield value in absolute terms. This indicates that in all cases the references’ SP dosage exceeds the saturation level, as e.g. Vikan & Kjellsen [341] defines it. Still, this a little bit surprising, since one would expect the different materials and volumes used would cause a larger deviation, analogous to the findings of Pedersen & Mørtsell [253], Vikan & Kjellsen [341], and Lowke & Schießl [212]. Then, the different matrix volumes and different admixture-binder interactions resulted in a different response to composite variations. The alteration of water content on yield value (Table 7-1) seems to have a greater effect on the particular lean SCC was tested, but only slightly affects the low-binder and binder-rich SCC. This is contrary to the findings in the slump flow results (Figure 7-3) and the therefore expected effects caused by a varied water content.

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table 7-1: Yield value response on varied SP (a) and varied water (b) for three different reference mix designs

100

SP_Eco-SCC

75

SP-lean SCC

50

SP_br-SCC

25 0 -25 -30

-20

-10

0

10

20

30

-50 -75 -100

Yield value difference [Pa]

Yield value difference [Pa]

100

water-lean SCC

75 50

water-Eco-SCC

25

water_br-SCC

0 -25 -10

(a) Yield value response on varied SP

0

5

10

-50 -75 -100


-5

Relative variation of water [%]

(b) Yield value response on varied water

Although plastic viscosity is not significantly affected by variations of SP, changes in the water content alter it significantly, see Table 7-2. Table 7-2: Plastic viscosity response on varied SP (a) and varied water (b) for three different reference mix designs

100 SP_br-SCC

75

SP-lean SCC

50

SP_Eco-SCC

25 0 -25 -30

-20

-10

0

10

20

-50 -75 -100


(a) Plastic viscosity response on varied SP

30

Pl. viscosity difference [Pa s]

Pl. viscosity difference [Pa s]

100

water_br-SCC

75

water-lean SCC

50

water-Eco-SCC

25 0 -25 -10

-5

0

5

10

-50 -75 -100 Relative variation of water [%] (b) Plastic viscosity response on varied water

The rheology response summarized in a rheograph can reveal whether the variations of SP-dosage (Figure 7-9) or water content (Figure 7-10) are critical or not, in order to achieve self-compacting properties while sedimentation risk is mitigated. The effect of SP variation mainly affects the yield value, as expected. Owing to the approximately similar matrix designs of the Eco-SCC and lean SCC mixes, see the corresponding figures in Table-Appendix 45 on page 315, there robustness is quit comparable, and self-compacting properties could possibly be achieved for a significant part of the mixes. Some are actually too stiff and therefore can only reveal semi-self-compacting properties, which could be readjusted relatively easily. Judging based on the ratio of the Bingham parameters, none of these mixes is suspected to reveal sedimentation risk, although the segregation halo that appeared around the slump flow for some of the mixes should be considered as a first indication of instability. For these mixes, it is advised to evaluate the dynamic stability and cohesiveness for the specific casting procedure. The case of the binder-rich mix is rather special, since both rheology parameters are affected and it remains uncertain from these data, whether the majority of these mixes would actually reveal stability against dynamic and static sedimentation. The occurrence of negative yield values will be discussed later in this chapter. ICI Rheocenter

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Yield value [Pa]

160 120 SP_Eco-SCC

80

SP-lean SCC SP_br-SCC

40

SCC-Boundaries

0 0

20

-40

40

60

80

100

120

140

Plastic viscosity [Pa s]

Figure 7-9: Rheograph with Bingham parameters considering varied SP-dosage

160

Yield value [Pa]

120 water_Eco-SCC

80

water-lean SCC water_br-SCC

40

SCC-Boundaries

0 0

-40

20

40

60

80

100

120

140


Figure 7-10: Rheograph with Bingham parameters considering varied water dosage

The differences in the effect of SP-dosage and water content are related to several different effects due to polymer concentration, their interactions with different binder material, and particle distance issues. For the first, the volumes and surface charges of the selected binder material contributes to the SP demand for one specific flowability state, which is also affected by the volume of lubricant available. For paste, the lubricant for the solids is mainly the water, and numerous attempts has been made to optimize the lubricant demand of paste, e.g. by Kennedy [176], Krell [200], Geisenhanslüke [129, 130, and 131], or Fennis-Huijben [105]. For concrete, the lubricant of the aggregates is the sum of all smaller particles and other volumes, i.e. water and air, which I consider as matrix volume in accordance with Mørtsell [226]. Then, the lubricant demand is governed by the aggregate’s PSD. To emphasize the differences when also employing the air content, the volume proportions of the matrices are differentiated in the following figure that shows the variation due to altered water content (Figure 7-11).

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Water (+air)/ powder per volume

1.8 1.6

Eco-SCC

1.4

Lean SCC

1.2 Vw/Vp by JSCE

1.0 0.8 0.6 170

180 190 200 Water content [kg/m3]

210

Figure 7-11: Volume proportion of water / powder (i.e. particles < 0.125 mm instead of 0.09 mm as in the JSCE) in solid marks, and (water plus air) / powder in unfilled marks.

Remark: The use of 0.09 mm according to the Okamura approach would lift up each single curve to a small extent, but this particle size is not conforming to the definition of a matrix in this thesis. Additionally, one has to consider the different volumes used in absolute terms. The stagnancy of the ratio VW+A/VP is caused by the fact that segregating mixes completely de-air, whereas for stiffer mixes the air content is governed by other effects, such as described in Table 2-14 on page 58. It remains uncertain whether a possibly occurring segregation in the case of EcoSCC is simply not observable due to the lack of matrix volume, or whether it is not occurring to such a distinctive extent. The slope of each single curve gives an indication for the robustness of a mix: the lower the slope, the less the property varies for one specific parameter alteration (Figure 7-12). Without considering the air volumes involved, the slope of the br-SCC is about two times higher than the one for Eco-SCC. This implies a superior robustness against volume alterations of the composite that affects sedimentation resistance the most- the water content.

Slump flow deviation [%]

20

y = 335.14x - 285.9 R² = 0.9859

y = 164.38x - 223.28 R² = 0.9893

10 Eco-SCC 0 0.6

0.8

1.0

1.2

1.4

1.6

Lean SCC br-SCC

-10

-20

1.8

Vw/ Vp ( Va+w/Vp)[-]

Figure 7-12: Effect of Vw/Vp (V(w+a)/Vp for unfilled marks) on deviation of slump flow, double marks at one variable indicate segregation

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The time dependent evolution of rheology parameters, as affected by the variation of the specific composite, is given below. In these figures, the data with a solid line represents the initial properties whereas the dotted line represents the parameters at 30 minutes. This time is about 10-15 minutes before the casting of the specimen. 800

800 Slump flow [mm]

Slump flow [mm]

10 min 600 30 min 400

SP_Eco-SCC

600

400

SP_br-SCC

Water_br-SCC

200 -30

-20 -10 0 10 20 Variation of composite [%]

Water_Eco-SCC

200 30

Figure 7-13: SP dependent evolution within time of slump flow for Eco-SCC and br-SCC

-10

-5 0 5 Variation of composite [%]

10

Figure 7-14: Water dependent evolution within time of slump flow for Eco-SCC and br-SCC

Even though the same materials have been used in these mixes, naturally in different quantities, the time dependent evolution differs significantly. For the case of Eco-SCC, the lower SP-dosage results in a workability loss, whereas the generally larger composite volumes in br-SCC causes an increase in slump flow. It seems to be independent whether a reduction of SP-dosage or water shortage causes the variation of initial fluidity-the differences in slump flow demonstrate a similar tendency. Based on the mode of action of dispersing admixtures, such finding has to be expected as their effectiveness is governed by the concentration. For the particular admixture used, the potential flow increase within time is actually employed to maintain workability in ready-mix concrete applications. This highlights the difficulties one faces in choosing a suitable admixture for lowbinder SCC mixes. The variations of the yield value within time (Figure 7-15 and Figure 7-16) are according to the varied fluidity. Nevertheless, the discrepancy between fluidity observed and yield value measured for the br-SCC series indicates that the measurement as such is somehow compromised, which will be discussed later in the chapter. The plastic viscosity remains stable within time (Figure 7-17 and Figure 7-18).

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC 160

80 40 0 -20

-10

0

10

20

40 0

30

-10

-5

Figure 7-15: SP dependent evolution within time of yield value for Eco-SCC and br-SCC

160

SP_Eco-SCC SP_br-SCC

80 40 0 -20 -10 0 10 20 Variation of composite [%]

30

Figure 7-17: SP dependent evolution within time of plastic viscosity for Eco-SCC and br-SCC

7.2.

5

10

Figure 7-16: Water dependent evolution within time of yield value for Eco-SCC and br-SCC


Plastic viscosity [Pa s] -30

0

-40 Variation of composite [%]

Variation of composite [%]

120

Water_br-SCC

80

-40

160

Water_Eco-SCC

120

SP_br-SCC

Yield value [Pa]

Yield value [Pa]

120

-30

160

SP_Eco-SCC

-10

120

Water_Eco-SCC Water_br-SCC

80 40 0 -5 0 5 Variation of composite [%]

10

Figure 7-18: Water dependent evolution within time of plastic viscosity for Eco-SCC and br-SCC

Discussion of apparently shear-thickening binder-rich SCC

The apparently negative yield value for the binder-rich SCC indicates that their determined flow curves do not obey a linear fluid model (i.e. Bingham). A linear model approach is used during the computation of intrinsic physical properties from measurements of resistances in the coaxialcylinders viscometer. One possibility could be to apply a different fluid model in order to be able to compare the different mixes, namely the Herschel-Bulkley model (Equation 7-1) and the Modified Bingham Model (Equation 7-2) as it is proposed by several authors, e.g. Feys et al. [113, 114, and 115], Heirman et al. [151, 152, and 153], Wallevik [348], and Yahia & Khayat [373]. Equation 7-1

𝜏𝑥𝑦 = 𝜏𝑜 + 𝐾𝛾 𝑛̇ Equation 7-2

𝜏(𝛾̇ ) = 𝜏 + 𝜇𝛾̇ + 𝐶𝛾 2̇

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205

Here, C is referred to as “second order coefficient” [114] and data published uses the ratio C/. Since C would govern the order of magnitude of shear thickening behaviour of a fluid, it reflects this property and should be named accordingly, in my opinion. One would propose to refer to it as “shear thickening/thinning coefficient”, depending on its algebraic sign. Proposals based on measurements obtained with a different technique, such as the parallel plate viscometer BTRHEOM by De Larrard [84], will not be considered here so as to mitigate any possible effect caused by using different investigation techniques, see e.g. Ferraris & Brower [109]. Despite the possibly inaccurate “nature”63 of the data of these mixes, the workability has actually varied in a similar magnitude as indicated in the figures of the previous section: for the case of SP variation, the computed plastic viscosity coincides with the experienced cohesiveness and remained more or less constant while only the fluidity was affected. For the case of water variation, both the fluidity (i.e. yield value) and the workability (here as indication for the plastic viscosity64) varied significantly. Then, the experienced resistance during handling the fresh concrete coincides with the plastic viscosity measured, ranging from highly viscous to low viscosity. The apparently detected shear thickening behaviour is indicated by the flow curves in Figure 7-19 and Figure 7-20. 10 0.72%

Torque [Nm]

8 0.81%

y = 14.122x - 0.0842 R² = 0.9927

0.90%

y = 12.896x - 0.1235 R² = 0.9916

0.99%

y = 12.38x - 0.2936 R² = 0.9859

1.08%

y = 13.238x - 0.2551 R² = 0.9898

6 4 2 0 0

0.1

0.2 0.3 0.4 Rotational velocity [rps]

y = 16.16x + 0.214 R² = 0.9918

0.5

Figure 7-19: Flow curves of different SP-dosages of binder-rich SCC, given as percent by binder content in the legend

It is a question whether this material can be assessed using the linear fluid model, assuming a homogen mix during the measurement, or whether the homogeneity criterion is invalid at all. 64 One might suggest to refer to the term “workability” as the ratio of Bingham parameters only. A high flow (i.e. low yield value) does not result in a “good workable mix” when the plastic viscosity is large and a low viscous mix does not have a “good workability” when the yield value is high (experienced as small slump). 63

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC 18 175 l

16 Torque [Nm]

14

y = 33.768x - 0.1639 R² = 0.9939 y = 25.864x - 0.1747 R² = 0.9946

180 l

12 10

185 l

y = 15.114x - 0.1287 R² = 0.9903 y = 10.715x - 0.0832 R² = 0.9933

8 190 l

6 4

195 l

2

y = 6.8218x - 0.0131 R² = 0.993

0 0

0.1


0.5

Figure 7-20: Flow curves of different water contents of binder-rich SCC

The data presented in these figures still consists of all torque points determined. Using the raw data, or one function in the operating software FreshWin, makes it possible to verify whether the stress state has reached its equilibrium at each rotational frequency. These raw data (see figures (b) in Table-Appendix 46 and -49 starting on page 316) have been analyzed using the equations from chapter 1.3. When the data were corrected for steady state, an accentuated shear thickening behavior can still be observed, compare figures (c) in Table-Appendix 46 and -49. The following Table 7-3 and 7-4 present interim65 rheology parameters obtained when different approaches (linear, power law, polynomial) were employed to evaluate the data, see in addition Table-Appendix 46 and -49 starting on page 316. Table 7-3: Interim rheology parameters of “br-SCC” for varied SP-dosage

linear (BINGHAM) model h g R2 SP0.72 SP0.81 SP0.90 SP0.99 SP1.08

K

B

R2

h2

h

G

R2

Slump flow Sf

Power law (HB) model

Polynom second order

0.43

14.7

0.997

12.2

0.78

0.994

1.51

14.1

0.5

0.997

605

-0.08

14.1

0.993

12.4

0.92

0.988

9.12

9.5

0.3

0.999

650

-0.03

12.3

0.994

11.0

0.93

0.994

8.09

8.7

0.3

0.999

680

-0.26

13.2

0.999

12.0

1.0

0.990

10.62

7.8

0.2

0.999

715

-0.09

11.1

0.980

10.4

0.98

0.995

8.82

7.5

0.2

0.996

743

66

Table 7-4: Interim rheology parameters of “br-SCC” for varied water dosage

linear (BINGHAM) model Power law (HB) model Polynom second order Slump flow h g R2 K B R2 h2 h G R2 Sf W175 W180 W185 W190 W195

32.9

-0.01

0.993

28.4

0.90

0.989

24.3 21.9 0.86 0.999

530

25.9

-0.17

0.995

23.3

0.95

0.992

14.6 18.5 0.49 0.999

573

14.3

+0.00

0.994

12.6

0.91

0.993

9.4

10.1 0.34 0.999

658

10.3

-0.01

0.995

9.2

0.93

0.994

6.5

7.4

0.22

695

6.8

-0.01

0.993

6.0

0.91

0.991

4.4

4.6

0.19 0.999

1.0

61

743

“Interim” refers to the computation step when the electrical resistances actually measured have been converted into torque resistances in [Nm]-units. 66 Here, segregation very clearly occurred, indicated by the halo surrounding the major spread. 65

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50

water_Br-SCC

40

ECC only

SP_br-SCC

Segregation

30 20 10

SCC

0 0

5 10 15 20 25 30 Shear thickening coefficent h2

Figure 7-21: Shear thickening coefficient h2 vs. sedimentation induced difference (range) of maximum and minimum SG

50

Diff. Max Min SG x1000

Diff. Max Min SGx1000

For the usual data treatment, these flow curves would still indicate an apparent shear thickening behavior of binder-rich SCC. When dealing with these mixes, my impression was that many of the mixes experienced a dynamic segregation. Unfortunately, I do not know any publication where the homogeneity of the mixes is investigated during the different shear rates or after them. The following figures depict the shear thickening exponent b of a power-law fluid-model, and the shear thickening coefficient h2 of a polynomial approach, with the corresponding slump flows and the difference dSG between maximum and minimum specific gravities (SG) induced by static sedimentation. The specific gravity is determined in hardened concrete and obtained from sedimentation cylinders as described in Chapter 6. Generally, it appears that the larger the inhomogeneity caused by sedimentation is, the more accentuated is the shear thickening effect represented by one of the parameters mentioned. Hereby it appears that the shear thickening coefficient provides a better indication of water (i.e. matrix volume) induced sedimentation (Figure 7-21). The shear thickening exponent appears to provide a better correlation when segregation due to SP-dosage for one specific matrix volume occurs (Figure 7-22).

40 ECC only

water_Br-SCC SP_br-SCC

30 20 10 SCC

0.7

0 0.8 0.9 1 1.1 1.2 Shear thickening exponent b

1.3

Figure 7-22: Shear thickening exponent b vs. sedimentation induced difference (range) of maximum and minimum SG

800

800

700

700

600 500

ECC only

400

water_Br-SCC

300

SP_br-SCC

200 0

5 10 15 20 25 30 Shear thickening coefficent h2

Figure 7-23: Shear thickening coefficient h2 vs. slump flow

Slump flow [mm]

Slump flow [mm]

When considering the same shear thickening parameters, but relating them to the slump flow, again the shear thickening coefficient provides a stronger correlation for the water series (Figure 7-23), whereas the shear thickening exponent has a stronger correlation for the SP-series (Figure 7-24).

600 500 ECC only 400 300 0.7


200 0.8 0.9 1 1.1 1.2 Shear thickening exponent b

1.3

Figure 7-24: Shear thickening exponent b vs. slump flow

Even though there is a direct correlation between the sedimentation-induced variety of specific gravities and slump flow (Figure 7-25), it is not the fluidity as such that is responsible for sedimentation and shear thickening behavior, but the methods (i.e. mix design parameters) by which this 208

Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC fluidity is achieved. When using a large paste volume that is dispersed to different extents, the yield value governs the sedimentation resistance, as the viscosity remains approximately constant. Then the ratio of the Bingham parameters τ0/μ reduces, which leads to sedimentation-, or more precisely dynamic segregation-, induced behavior that is experienced as shear thickening. When for the same high paste volume the water volume is altered, the plastic viscosity governs the area of rheology. Then, the very high viscosity of some mixes affect the flow curve, particularly at the higher shear rates, in which the homogeneity of the shear zone might be altered. At the other extreme, mixes having actually too low viscosity for stability considerations due to excessive water content, the thereby increased τ0/μ assists in the creation of a lack of dynamic stability. Caused by the measurement technique applied, the dynamic segregation would occur at the beginning of a measurement. Later on in the measurement, the settling particles cause higher solid concentration and thereby would increase the apparent viscosity detected. Hence, the resulting flow curve appears to obey a power-law fluid-model of shear thickening behaviour. This would question the oftenreported shear thickening behavior of binder-rich SCC, as it would be only a measurement artifact of either a dynamically segregated concrete mix or a slippage layer formed by a large paste volume of high viscosity. One could propose to verify this hypothesis with further investigations. One approach could be to determine the homogeneity at different velocity steps of one measurement by a suitable method. Modeling might be another approach.

Difference of Max. Min SG x 1000 [-]

50 40

Segregation / high sedimention risk

30

ECC only water_Br-SCC

20

SCClow viscosity

SP_br-SCC

10 SCC

0 200

300

400 500 600 Slump flow [mm]

700

800

Figure 7-25: Slump flow vs. sedimentation induced difference (range) of maximum and minimum specific gravity (SG) in hardened concrete

7.3.

Effect in hardened concrete

7.3.1. Hardening monitoring with heat of hydration The hydration was monitored up to 48 hours for five of the six series with a calorimeter under isothermal conditions. It is interesting to see whether the variations of composites affect the end of the dormant period of the hydration, and how eventually the acceleration period is altered. The energy is a measure for the exothermic ion dissociation during the chemical reaction of the clinker phases. It is therefore an indication of the maturity at a specific time, which corresponds to the strength obtained at this time. The generalized effects observed are shown in Figure 7-26: the waICI Rheocenter Florian V. Mueller

209

ter content affects the intensity of the acceleration period, as it provides the medium that governs the dissociation and transport properties of the reacting ions. The superplastizers used can vary the dissolution processes occurring in the dormant period, and thereby influence the onset of the acceleration period [208]. The order of magnitude of this effect depends on several different factors, such as: the amount and reactivity of binder material, the volume of water, the type and dosage of polymer used, and the particular interaction of the polymer with the binder material. The detailed curves of power consumptions in an isothermal calorimeter are given in TableAppendix 48 to -52 starting on page 321. Still one comment about the curvature should be made here: Despite the often-reported heat peak caused by the initial reaction of the aluminate phase, see for example Figure 6.9 on page 221 in Mehta & Monteiro [222], the appearance at the beginning of each single curve is in very good agreement with the different fresh concrete temperatures. These temperatures are additionally provided in the figures. It can influence the integral of the power consumption and thereby alter the cumulative energy at a specific time. Additionally, the material is placed in the calorimeter about 15 minutes after water addition, at which point a significant part of this early exothermal reaction has already finished. Then, the material needs to be brought (usually by cooling) to the reference temperature of 20.0°C. As expected, the admixture used in the lean SCC resulted in an almost insignificant set-retardation of about less than one hour between the maximum and the minimum dosage. This admixture is optimized for Pre-Cast Industry applications and favours early strength. The admixture used for both, the Eco-SCC and for the binder-rich SCC, is optimized for Ready Mix Concrete applications. The requirement for extended workability often causes a certain retardation effect. The difference between maximum and minimum dosages is about 2-3 hours for Eco-SCC, and about 5 hours for the binder-rich SCC. Since the same cement is used in both cases, only with about 1.4 times more in the br-SCC, these delays can clearly be attributed to the absolute content of SP-polymer in these mixes, which is also about 1.4 times higher. 25 Heat of hydration

20

+/- water

15

Ref Water

10

SP

+/- SP

5 0 0

4

8 Hydration 12 time 16

20

24

Figure 7-26: Generalized scheme of how water and SP variations alter the hydration heat of one reference mix

7.3.2. Compressive strength One important point to consider from these effects during hardening is whether the strength is affected. Unfortunately, the early strength of the lean SCC mixes was not determined, which excludes them from the following figures.

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SP_br-SCC

water_br-SCC

Relative strength 2d

Relative strength 2d [%]

110

SP_Eco-SCC

105 100 95 90

110

water-Eco-SCC

100 90 80

-30 -20 -10 0 10 20 Relative variation of SP [%]

30

Figure 7-27: Effect of SP-dosage variation on compressive strength at 2d

-10

-5 0 5 Relative variation of water [%]

10

Figure 7-28: Effect of water content variation on compressive strength at 2d

To interpret the trends appearing in both of the different parameter series, several factors have to be considered. First of all, the state of rheology at which the specific mix was when the specimens were cast, see Figure 7-15 to Figure 7-18 on page 205ff, which leads to different degrees of consolidation, particularly when comparing a true self-compacting mix with a mix that only demonstrates easy-compacting properties. This can occur due to loss of workability and would require some compaction energy. At the testing time of 48h (2d), any set retardation effect only contributes to a variation of strength, but is not the major reason anymore. It would only dominate the maturity (and therefore strength development) in the first hours up to about 24h, compare the figures (b) and (c) in Table-Appendix 48 to -55 starting on page 321. Additionally, density and the consolidation degree affect the compressive strength; see Table 7-5 and Table 7-6. The degree of consolidation is indicated by the ratio of specific gravity and unit weight (i.e. density), of course within a certain inaccuracy of the method in particular when determining the density. Still, insufficiently compacted specimens reveal an open porosity that leads to a larger difference than the accuracy implied by the method. Since it has not been determined, the impact of the consolidation on the formation of the Inter Transition Zone (ITZ) between paste and aggregate particles is not considered, and thereby on the micro-crack propagation, which affects strength in general. Table 7-5: Effects of cumulative heat of hydration (a), density (b) and consolidation degree (c) on compressive strength at 2d for SP-series


SP_Eco-SCC

105 100 95 90 200

220 240 260 Energy J/g CM at 2d

(a) Compressive strength affected by cumulative heat of hydration after 48h (2d)

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100

SP_br-SCC SP_Eco-SCC

80 60 40 20 2350

2400 Density [kg/m3]

2450

(b) Compressive strength affected by density

Florian V. Mueller

Compressive strength 2d [MPa]

100

SP_br-SCC

Compressive strength 2d [MPa]

110

SP_br-SCC SP_Eco-SCC

80

Ref 60 40 20 1.00

1.01 1.02 SG / Density

(c) Compressive strength affected by consolidation degree

211

For the water series, the known correlation of w/c with compressive strength might be the governing influence on the effect observed when increasing the water content. When the water content is reduced, the rate of strength increase can stagnate or even reduce, an effect probably be caused by the consolidation effect already discussed, see Table 7-6.

100

water_br-SCC water-Eco-SCC

110 100 90 80

200 220 240 260 Energy J /g CM at 2d (a) Compressive strength affected by cumulative heat of hydration after 48h (2d)

100

Br-SCC Eco-SCC

80 60 40

20 2350 2400 2450 Density [kg/m3]

(b) Compressive strength affected by density

Compressive strength 2d

120

Compressive strength 2d


Table 7-6: Effects of cumulative heat of hydration (a), density (b) and consolidation degree (c) on compressive strength at 2d for water-series, related to results of the reference mix

Br-SCC Eco-SCC

80 60 40

Ref

20 1.00

1.01 SG/Density

1.02

(c) Compressive strength affected by consolidation degree

The following figures depict the compressive strengths at a normative age in absolute terms and relative to each reference mix. 100

80

brSCC

y = -0.0303x + 89.606

60

lean SCC

y = 0.1136x + 64.626

EcoSCC

40 y = -0.0062x + 38.711

Strength 28d [MPa]

Strength 28d [MPa]

100

80 y = -0.5809x + 88.78 y = -0.3788x + 65.912

Eco-SCC

40 20

-30 -20 -10

0

10

20

30

-10 -5 0 5 10 Relative variation of water [%]


115 110 105 br-SCC

100

lean SCC

95

Eco-SCC

90 85 -30

-15 0 15 30 Relative variation of SP [%]

Figure 7-30: Compressive strength of water series

Rel. strength @ 28d [%]

Figure 7-29: Compressive strength of SP series

Rel. strength @ 28d [%]

SCC

60

y = -0.4566x + 39.057

20

115 110 105 br-SCC

100

lean SCC

95

Eco-SCC

90 85 -10 -5 0 5 10 Relative variation of water [%]

Figure 7-31: Alteration of compressive strength for SPseries

212

br-SCC

Figure 7-32: Alteration of compressive strength for water-series

Florian V. Mueller

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC For this maturity level and binder composition, the effects coming from w/c and degree of consolidation govern the alteration of strength, as can be seen when comparing the figures about the effect of density and SG/density in Table-Appendix 53 on page 326.

7.3.3. Sedimentation resistance One concern during the evaluation of the robustness of one mix design is how the altered rheology affects the stability against sedimentation resistance. Since the sedimentation approach followed in this thesis is based on evaluating the specific gravity distribution in hardened concrete, it should be discussed in this chapter. A brief discussion has already been made in the previous chapter. Only for the most critical mixes concerning the paste volume (matrix volume respectively) and biggest alteration of rheology properties, i.e. the both series for binder-rich SCC, the sedimentation was determined. They also contribute to the results and discussion of chapter 6, so they will be only briefly presented here (Figure 3-35). The effects for mixes with lower matrix volume, as for the lean SCC and the Eco-SCC, can be concluded from the other results in chapters 4 and 6, when a similar matrix volume is applied and a similar rheology obtained.

Difference Max. Min SG x1000

50 40

ECC

30 water_Br-SCC 20 10

SCC -30

Segregation / High risk of sedimentation

-20

SP_br-SCC SCClow viscosity

0 -10 0 10 Variation of composite [%]

20

30

Figure 7-33: Effect of alteration of composite on the difference of maximum and minimum specific gravity (SG)

With increasing the SP-dosage, the range between maximum and minimum specific gravity (dS) determined in the sedimentation cylinders increases. This is probably caused by the reducing ratio of Bingham parameters that also determines the stability. When altering the matrix volume with the alteration of water content, and thereby the Bingham parameters significantly (see Table 7-1 and Table 7-2 on page 201), the curve obeys a parabola. The maximum water content causes instability, whereas the minimum quantity causes an insufficient self-compacting property. Using the ranking criteria Visual Stability Index in hardened concrete (HVSI) after Shen [304] gives Figure 7-34. Based on this ranking, only the mix with the highest water content would be rated as unstable class (3), whereas the higher SP-dosage would only be referred to as unstable class (2). The reference mixes would be stable, whereas the mixes representing the reduced composites would be rated as stable. Taking into consideration the information about the mixes we collected before in this chapter, some of them actually did not obtain self-compacting properties at all. The use of several grades for “stable” and “unstable” seems to be an inappropriate approach to cover all the differences occurring during this robustness test. ICI Rheocenter

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213

Unstable

2

Unstable

HVSI

3

-30

water_Br-SCC

-20

1

Stable

0

Stable

-10 0 10 20 Variation of composite [%]

SP_br-SCC

30

Figure 7-34: Effect of alteration of composite on the Visual Stability Index of Hardened concrete

ASAI

Applying the ranking systems proposed in this thesis gives Figure 7-35 and Figure 7-36. 4

High risk of sedimentation

3

Moderate risk

2

Small risk


-30

-20

1

Stable & self-compacting

0

Insufficient SCC

-10 0 10 20 Variation of composite [%]

30

Figure 7-35: Effect of alteration of composite on the Aggregate Surface Appearing Index

The Aggregate Surface Appearance Index (Figure 7-35) provides a qualitatively similar picture as the HVSI (Figure 7-34), but provides a more detailed distinction. One could be of the opinion that the ASAI provides a reasonably good description of the capability of these mixes. Applying the limit criteria established for the Classes of Specific Gravity Distributions gives the information about the homogeneity as function of the height in a test specimen (Figure 7-36). An indication of whether or not self-compacting properties have been obtained is also included in the approach. The three mixes in question are consistent with the ratio of Bingham parameters before casting the specimen, compare Table 7-1 and Table 7-2 starting on page 195. The approach employs a broad range for specific gravities that can possibly occur, see Chapter 6. This yields the effect that the most fluid mixes of the series considered here are only ranked as class two and four. Despite the ranking from HVSI, the major concern for the most fluid mixes of the SP-series seems 214

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CSGD

to be the static sedimentation, but not the dynamic segregation resistance. The latter is mitigated due to sufficiently high viscosity that remains unaltered with the SP variation. Only when the plastic viscosity is also reduced considerably, as for the highest water content, the homogeneity is altered to such an extent that the mix is rated CSGD 4. This would actually imply that only the risk for static sedimentation increases, but not the one for dynamic segregation. Here is one limitation to judge for the dynamic segregation resistance in this approach, as it only contributes to the appearance of homogeneity in the sedimentation cylinders, although it is actually not directly measured. Therefore, tt is recommended testing it separately when the CSGD rating exceeds the stability criterion. For the specific requirements of one application, one is free to vary the limits proposed in this thesis, when judging the SGD into classes. Of course, then the information received from these mixes would be different and vary the judgment, accordingly. On the other hand, the limits have been chosen in order to compare a broad range of mixes and their possibly occurring differences.

-30

6


5

Mod. static sed.& dyn.segr.

4

Moderate risk of static sed.

3

Small static sed. & dyn, segr.

2

Small risk of static sedimentation

1


0 -20 -10 0 10 20 Variation of composite [%]


Insufficient self-compacting 30

Figure 7-36: Effect of alteration of composite on the Classes of Specific Gravity Distributions

7.3.4. Drying shrinkage The drying shrinkage was only determined for the low-binder Eco-SCC and here only for the series, where due to altered water content the biggest effect had to expected. For the SP-series, it was assumed that the effects that alter the compressive strength a little bit, does not contribute significantly to drying shrinkage. Only a severe retardation would alter the maturity at the initial reading after 24 hours, which can slightly affect the result, see e.g. Mueller et al. [232]. The results (Figure 7-37) are brought in context with those obtained from an average Icelandic CVC that ought to be a similar performance type in hardened concrete properties to the intended target for the application of Eco-SCC. The variation of 20 l water in total already reveals a strain difference of about 0.2 mm/m at the age of 56 days within the series and another about 0.2 mm/m to the Icelandic reference. With further progressing time, the result remains constant, at least for the time span considered. Analogous to the shrinkage strain, the weight loss of the specimens, considered as evaporation, can be rated accordingly (Figure 7-38). Here, the increasing evaporation corresponds directly to the increase of water. Even though one can not foresee the results of a recent study at the ICI, a hypothesis would be that the moisture entrapped in the Icelandic aggegates used is not enclosed in the porosity, but can interact with the paste. The water content in SSD conditions of these aggregates, when tested according to EN 1097-6:2008, is in the vicinity of ICI Rheocenter

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3% for the fine fraction and up to 5-8 % for the coarse fraction. Therefore, more water than commonly considered in the w/c can contribute to the performances in hardened state. This also includes higher shrinkage strain and considerably increased evaporation at a similarly nominal w/c.

0.8

+5 Ref

0.4

-5l - 10 l

6 4 2 0

0.0 0

56

112 168 Time [d]

224

Figure 7-37: Drying shrinkage strain for specimen of Eco-SCC water-series and one Icelandic CVC as commonly used (Ref Ís01)

7.4.

8

ICECVC + 10 l

Evaporation[%]

Shrinkage [mm/m]

1.2

0

56

112 168 Time [d]

224

Figure 7-38: Evaporation (i.e. weight loss) of drying shrinkage specimen in 22°C @ 50%-RH

Concluding remarks

The robustness of three general mix designs was investigated, one representing Eco-SCC, one leanSCC, and one binder–rich SCC, concerning the response of their fresh and hardened concrete properties on alteration of water by ±5 l/m3 and 10 l/m3, or alteration of PCE by ±10% and 20%. It appears that a mixture designed for relatively high w/b, the Eco-SCC, is remarkably robust against such alterations, to the contrary of binder-rich SCC. Here the rheology area covered by the altering mix designs are difficult to assess, in particullar in the consequences the fluctuating plastic viscosity has on sedimentation and segregation stability. On the opposite, Eco-SCC covers a small rheology area only, located “on the safe side”. This means that it either exhibits clearly visible segregation, allowing immediate countermeasures e.g. by adding ST, or it reveals as the other extreme semi-Self-Compacting properties only, which are also easy to detect and still allows countermeasures by adding the appropriate missing compound. The shear thickening of binder-rich SCC was briefly discussed and brought in context with the excess paste volume (i.e. matrix volume) and sedimentation risk affected by the various alterations of the mix design.

216

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8. Correlation of different rheology devices and tests This chapter deals with various rheological and rheometric aspects, when an attempt is made to adapt an existing approach for slump-yield value relation to a slump flow-yield value relationship, but for a value then determined in the co-axial cylinders viscometer, see chapter 8.1. This would allow estimating the yield value for a stable slump flow in order to utilize the assessment in a rheograph, at least as soon the plastic viscosity can also be reliably estimated in the future. Another aspect is the correlation of the results obtained in the portable impeller-based device Rheometer4SCC with the intrinsic physical properties determined in the co-axial cylinders viscometer. Owing to the different geometry and therefore a different physical – mathematical principle, for the portable device the well-known Reiner-Riwlin equation cannot be employed. Closing this gap allows utilizing the rheograph again when the results obtained in the Rheometer-4SCC can be converted into Pascal units, see chapter 8.2. Both these approaches are also validated for a set of data in chapter 8.3.

8.1.

Slump flow

One seldom has the opportunity to obtain a large data set covering a broad range of fluidity parameters. These data are used to quantify the correlation between slump and slump flow, in order to adapt existing models dealing with slump-yield value correlations. When SCC was introduced, it became obvious that the slump value cannot serve as a parameter to quantify the consistency of SCC. Therefore, it has not been determined in this project for cases of higher fluidity. All other results were screened in order to omit results of mixes where segregation has been observed. The residual results seem to create one general curvature (Figure 8-1), independent from the mix design characteristics. Still, one has to consider the data set as cumulative series of curves representing the stable mixes only, somewhat analogous to the discussion in Wallevik [347] and Flatt [118]. Slump flow values obtained from segregated mixes would broaden the variability. The maximum fluidity (in terms of slump flow) that can be obtained for one specific matrix volume is always limited by stability considerations, reflected by the ratio of rheology parameters for one specific yield value. Reducing the yield value further with a superplasticizer would result in a lack of stability and not increase the flow considerably, see the discussion about the effect of matrix volume and its’ proportioning in Wallevik [347] and in greater detail in chapter 4. When trying to extend the flowability, the matrix volume needs to be extended, its viscosity enlarged while reducing the yield value further, and eventually the maximum aggregate diameter needs to be reduced. Different mixes from different projects, such as my second MSc-thesis [232], an early phase of this project with my “Diplomarbeit” (“1st MSc-thesis”) [230], and recent results (“Eco-Project”) are used to derive Figure 8-1. A broad variety of paste volumes and w/b are used in these mixes, apart from varied binder materials, admixtures, and even aggregates. Anyway, a pure geometric effect is quantified here, which might make it irrelevant how the differences in the slump value have been obtained as long they represent a stable mix. It should be stressed here once that the slump flow test was generally conducted with the regular positioning of the cone, but not the inverse one as it is also allowed for SCC in the ASTM C1611 and EN 12350-8. To convert the slump value of a conventional density concrete into the corresponding slump flow value, the empirically derived model equation (which is added in Figure 8-1 for selected slump values) follows Equation 8-1. Since this model is derived empirically, it may not apply with different cone geometries and volumes. Nevertheless, a potential future adaption to mortar cone geome-

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217

tries might aid in estimating the yield value for mortar tests, which are commonly used to examine the effect of binder and admixtures in preliminary tests before testing concrete. Equation 8-1

𝑆𝑓 = 𝑑𝐴 ∗ 𝑒 (0.05∗(𝑆∗10) With:

dA S Sf

3.2 )

Opening diameter of the Abrams cone, i.e. 0.20 m (200 mm); Slump value [m]; Slump flow value [m] ([mm])67

800

1st MSc

Slump flow [mm]

700

y = 61.762e0.0856x R² = 0.8423

600

2nd MSc Eco Project

500

Model

400

Expon. (2nd MSc) y = 115.89e0.0563x R² = 0.8199

300

Expon. (Eco Project)

200 0

5

10

15 Slump [cm]

20

25

30

Figure 8-1: Relationship of slump flow and slump value for conventional density concrete, using stable mixes only

Emphasizing fluid concrete, Figure 8-2 reveals that this model equation represents Sf in the vicinity of an average value for all the results at one specific slump. Since it appears to describe also the low slump concrete with reasonable accuracy, this equation could be used as an extension of existing yield value-slump value models. 800

Slump flow [mm]

700

1st MSc

600

2nd MSc

500

Eco Project

400

Model

300 20

22

24 26 Slump [cm]

28

30

Figure 8-2: Zoomed Figure 8-1 into area of fluid concrete and SCC

The scientific approach would be to use [m] in both cases, but in order to apply the numbers in the dimensions that are commonly agreed on and defined in the relevant standards, this approach is chosen intead. 67

218

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC When determining the relationship with a mathematic approach, the volume V of the frustum of a cone needs to be considered according Equation 8-2 Equation 8-2

𝑉= In which:

ℎ 𝑅1 𝑅2

𝜋ℎ 2 (𝑅 + 𝑅1 𝑅2 + 𝑅22 ) 3 1

Cone height [m]; Upper radius [m], Lower radius [m].

For the conditions in the Abrams cone, with h=0.3 m, R1 = 0.05 m, and R2 = 0.10m, the Volume VAbrams cone becomes 0.0055 m3, equivalent to about 5.498 liters. Assuming the volume remains constant allows balancing the volume of the cone to the volume of the slump at any time, although for very viscous mixes, which would require very high fluidity in order to achieve self-compactability, the volume stability is sometimes not given since a large quantity remains stuck inside the cone. Applying such an approach would alter the height as a function of the slump (Equation 8-3), and the lower radius would be equal to the slump flow radius (Equation 8-4). Equation 8-3

ℎ𝑆𝑓 = ℎ𝑐𝑜𝑛𝑒 − 𝑆 = 0.3 − 𝑆 Equation 8-4

𝑅2 = 𝑆𝑓⁄2 The rheology parameters causing a certain fluidity at stable conditions are incorporated in Equation 8-5 as what is considered here in this thesis as corrective terms α1 and α2. When using them in the proposed values, the resulting curve overlaps well with the purely empirical graph (Figure 8-3) using Equation 8-1. Remark: The validity of this equation for different cone geometries ought to be done separately, potentially altering the corrective terms then. Equation 8-5

𝑆𝑓 = (2𝑅2,𝑐𝑜𝑛𝑒 )(𝑉∗𝑔∗𝜌∗𝛼𝑎∗𝑆 With:

𝑆𝑓 𝑅2,𝑐𝑜𝑛𝑒 V g 𝜌 α1 S α2

𝛼2 )

Slump flow [m]; Lower radius of the Abrams cone, i.e. 0.10 m; Volume of the Abrams cone, i.e. 0.0055 m3; Gravity force [m/s2]; Density [kg/m3] Correction factor, about 0.61; Slump [m] Correction exponent, about 3.2.

Remark: Although the density would affect the shear stresses occurring during the flow, the considerable alteration of it due to air-entrainment is not considered with respect to its effect obtaining a stable flow. ICI Rheocenter

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Empirical

𝑆𝑓 = 𝑑𝐴 ∗ 𝑒 (0.05∗(𝑆∗10)

𝑆𝑓 = (2𝑅2,𝑐𝑜𝑛𝑒 ) ∗

0.8 0.7 0.6 0.5 0.4 0.3 0.2 0.1

Density 2300 kg/m3

3.2 𝑒 (𝑉∗𝑔∗𝜌∗0.61∗𝑆 )

0.2 Slump value [m]

Slump flow value [m]

Slump flow value [m]

Mathematical

Density 2400 kg/m3

3.2 )

0.3

Figure 8-3: Model comparison purely empirical vs. semi-empirical incorporating a mathematical approach

Density 2200 kg/m3

0.8 0.7 0.6 0.5 0.4 0.3 0.2 0

0.1 0.2 Slump value [m]

0.3

Figure 8-4: Theoretical impact of density on SlumpSlump flow relation without considering stability due to density alteration using air-entrainment

Figure 8-5 depicts how the slump flow value is dependent on the yield value. In agreement with the results of my second Master’s thesis [232] (see the trend line in Figure 8-5), then also covering a broad range of rheology parameters, Esping [100] reports a relatively similar relationship following Equation 8-6. Equation 8-6

𝑦 = −124.62𝐿𝑛(𝑥) + 1071.82 𝑅 2 = 0.89 Here, 𝑦 corresponds to the slump flow value [mm], whereas 𝑥 corresponds to the yield value [Pa]. In his studies, Esping used the same viscometer model as we used during our investigations. It would not require a device-dependent adaption of the formula. Some results obtained during the early phase of the project, i.e. my Diplomarbeit [230] (“1st MSc”), might be measured inaccurately and should not be considered in detail, but only as additional information about occurring variability. Then, a measurement setup was used in the Viscometer 5 that was possibly not always suitable to reflect the rheology parameters; see the discussion in chapter 3.1.4. 800 y = -76.29ln(x) + 882.23 R² = 0.5137

Slump flow [mm]

700

1st MSc

600

2nd MSc

500

Eco Project

y = -123.1ln(x) + 1050.4 R² = 0.8878

400

Log. (2nd MSc)

300 200 0

100

200

300 400 500 Yield value [Pa]

600

700

800

Figure 8-5: General correlation between slump flow and yield value for a variety of mixes, including only mixes without obvious segregation signs and air content below 2.5%

220

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC The mixer-efficiency effect, see e.g. [227, 365, and 367], is not of any concern here. It is a different topic to estimate the fluidity achievable with a certain mix design and specific mixing equipment and procedures. In order to model the yield value from slump flow, extensive work was done at the former LCPC (now IFSTTAR) by Hu [161], Hu et al. [162], De Larrard [85], Ferraris & De Larrard [107], and also Roussel [285, 286, and 288], basing on mortar tests then. The latest modification of the original (since there is no official name for it yet let us call it) BTRHEOM yield value-slump-model (or in short: “τ0,BTSM for BTRHEOM”), provides a semi-empirical equation (Equation 8-7) to calculate the slump value (S), considering the specific gravity (ρsg) (i.e. comparing the density of concrete with those of water with 1000 kg/m3) and the yield stress value determined in a BTRHEOM (τ0, BT). Equation 8-7

𝑆 = 300 − 0.347

(𝜏0,𝐵𝑇 − 212) 𝜌𝑆𝐺

Example calculations presented in Figure 8-6 indicate that common concrete densities ranging from 2200 to 2400 kg/m3 do not have a considerable effect on the slump value, and slump flow value respectively when also applying Equation 8-1. The model extension to cover lightweight or heavyweight aggregate concrete will not be discussed here. Therefore, this model might serve well to estimate the slump flow for our cases of conventional density concrete. The correlation between the yield values determined either in a ConTec BML Viscometer (τ0, BML) or in a BTRHEOM (τ0, BT) has been determined once empirically and is given in Ferraris et al. on page 46 in [109], noted as Equation 8-8. Equation 8-8

𝜏𝑜,𝐵𝑀𝐿 = 0.5 ∗ 𝜏0,𝐵𝑇 − 122 Applying the transformation of Equation 8-8 to τ0, BT into an adapted Equation 8-7 gives the ConTec modification for the τ0S-model, or in short “τ0,BMLSM for ConTec BML”, represented by Equation 8-9. Equation 8-9

32 + 2𝜏0,𝐵𝑀𝐿 𝑆 = 300 − 347 ( ) 𝜌 Applying Equation 8-1 into Equation 8-9 it allows extending the “τ0,BMLSM” to “τ0,BMLSfM”. Both models are depicted in Figure 8-6 for a range of densities commonly occurring in conventional density concrete.

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Slump [mm], Slump flow [mm]

800 700

Slump, roh 2400

600

Slump flow, roh 2400

500

Slump, roh 2300

400

Slump flow, roh 2300 Slump, roh 2200

300

Slump flow, roh 2200

200 100 0 0

200

400

600

800

Yield value , τ0,BML [Pa] Figure 8-6: ConTec modified BTRHEOM yield value-slump model by Hu et al. [162] (“τ0,BMLSM for ConTec”) and its extension for slump flow (“τ0,BMLSfM for ConTec”)

This equation differs from the equation mentioned by Wallevik [347] (see Equation 8-10), even though he also refers to the same source when considering the device adaption. Equation 8-10

𝑆 = 300 − 0.416 (

394 + 𝜏0,𝐵𝑀𝐿 ) 𝜌𝑠𝑔

In the same paper, Wallevik suggests a correction term for this equation, in order to take into account the matrix effect, as described by Equation 8-11. Equation 8-11

𝑆 = 300 − 0.416 (

394 + 𝜏0,𝐵𝑀𝐿 𝑟𝑒𝑓 𝑟𝑒𝑓 ) + 𝛼(𝜏0 − 𝜏0 )(𝑉𝑚 − 𝑉𝑚 ) 𝜌𝑠𝑔

Here, α is an empirical coefficient and 7.7*10-3 mm/ (Pa l) is used in his paper. Since the reference terms are established to compare the effect of matrix volume for a mix relative to a reference mix, this equation cannot be applied for the case discussed in here and only Equation 8-10 is considered further. Combining the Equation 8-1 (i.e. slump-slump flow) and Equation 8-9 (i.e. slump -yield value) (summarizing expressed as “Model 01” in the following figures) does not lead to a satisfying computation of the slump flow measured for the fluid mixes. Since any obvious segregation has been omitted for the correlation of slump flow and slump value, the Equation 8-1 can be considered more or less exact. What is contributing to the misleading computational results can only come from either the correlation of the different devices BTRHEOM or ConTec Viscometer, or when the results obtained with the BTRHEOM affected the original slump-yield value model by Hu et al., i.e. Equation 8-7. The device-correlation Equation 8-8 was derived for a variety of mixes ranging from 9 cm to 24 cm slump; compare Table 8 on page 36 in [109]. Taking into consideration the distinct non-linear decrease of the yield value for increasing slump flow, the range evaluated in [109] might be insufficiently small to derive the relationship for the whole range of consistencies. When changing only the parameters regarding the device-correlation Equation 8-8 through Equation 222

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC 8-12, the results for the fluid mixes can be estimated with a much higher accuracy, see “Model 02” in Figure 8-8. Unfortunately, the lower slump and stiff concrete consistencies cannot be expressed with this equation, as their slump and slump flow would both be considerably underestimated, see in Figure 8-7. Equation 8-12

𝜏𝑜,𝐵𝑀𝐿 = (1/3) ∗ 𝜏0,𝐵𝑇 − 320/3 A third deviation, expressed as “Model 03”, would be a combination of another alteration of Equation 8-8 through Equation 8-13 instead and an alteration of Equation 8-7 through Equation 8-14, summarized in Equation 8-15. Equation 8-13

𝜏𝑜,𝐵𝑀𝐿 = (1/3.8) ∗ 𝜏0,𝐵𝑇 − 600/3.8 Equation 8-14

𝑆 = 300 − 180 Equation 8-15

(𝜏0,𝐵𝑇 − 400) 𝜌

200 + 3.8𝜏0,𝐵𝑀𝐿 𝑆 = 300 − 180 ( ) 𝜌 Table 8-1: Overview of equations combined in different models

Relation S-Sf τ0, BML - τ0, BT S-τ0, BML

Model 01 Equation 8-1 Equation 8-8 Equation 8-9



800

Slump flow [mm]

700 1st MSc

600

2nd MSc 500

Eco Project

400

Model 01 Model 02

300

Model 03

200 0

100

200

300 400 500 Yield value [Pa]

600

700

800

Figure 8-7: Measured and modeled slump flow values from determined yield values, using several different models, see Table 8-1

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800

Slump flow [mm]

700 1st MSc 2nd MSc 600

Eco Project Model 01 Model 02

500

Model 03 400 0

40

80 Yield value [Pa]

120

160

Figure 8-8: Zoom of Figure 8-7 into area of fluid concrete and SCC

An early approach to estimate the slump based on a given yield value by Murata & Kukawa [235] (Equation 8-16) has already been discussed by Flatt et al. [119] and Wallevik [347]. When combining it with Equation 8-1, the obtained slump flow is considerably underestimated, as it can be observed in Figure 8-9. The approach proposed by Roussel et al. [286] and validated by Flatt et al. [119] seems to consider reasonable properties when they include the density of the material (ρ), the test volume (V) and a term related to “...liquid vapor interfacial energy and the wetting angle on the plate...”; see Equation 8-17. Nevertheless, when plotting the results of this equation in Figure 8-9 it reveals a considerable overestimation of the slump flow, up to about 465 mm when it crosses the curvature described by the original LCPC approach. It is possible that the model’s validity is limited to its original combination (using the Viskomat NT from Schleibinger then), and the different conditions. The equation from Wallevik [347], adapted for slump flow by using Equation 8-1, was derived for low and intermediate slump values, and only works there. Applying my own modification with Equation 8-15 (i.e. “Model 03”) seems to serve reasonably well over the full range of fluidity. Nevertheless, the equations of “Model 02” should be favored for fluid concrete mixes beyond 400 mm slump flow. Equation 8-16

𝜏0 = 729 − 483log(𝑆⁄10) With:

S

Slump value [mm]

Equation 8-17

With:

224

ρ V R λ

𝜏0 = 1.747𝜌𝑉 2 𝑅−5 − 𝜆 𝑅 2⁄𝑉

Density; Test volume, i.e. 5.498 x 10-3 m3 for the Abrams cone; Radius of spread [m]; Term linked to the “liquid vapor interfacial energy and the wetting angle on the plate”

Florian V. Mueller

ICI Rheocenter


800

700

600

500 0

40

80

120

160

Model 01 (LCPC) Model 02 Model 03 Flatt et al. Murata et al. Wallevik Figure 8-9: Estimated slump flow for adapted slump-yield value relationship according to various models

Turning the perspective now, when estimating the yield value instead based on one measured slump flow value, the abscissa has been converted into the ordinate (Figure 8-10), even though such an approach is contrary to the physical relationship. Here, there could not one single equation be found to estimate the yield value for the full range of slump flow values. The trend equation obtained for the fluid mixes could be used up to a slump flow of 790 mm beyond which the computed yield value would become negative. Mixes exceeding such fluidity reveal competing effects from either an increased stability risk or being affected by thixotropy due to their mix design parameters. Therefore, it becomes difficult to the actually measure their rheology parameters exactly and one should use appropriate additional tests to quantify whether they fulfill the acceptance criteria of the target application. The difference between the trend equations of “2nd MSc” [232] and “Eco Project” probably have to be attributed to the different codomains, although for instance Flatt [118] and Wallevik [347] show a certain correlation between yield value and solid concentration. This relationship, then derived for paste in the case of Flatt’s work, could be interpreted as the previously mentioned matrix effect when it is extended to concrete. Of course, the solid concentration not only affects the yield value but also the plastic viscosity, as we recall in the simplest approach made by Einstein more than a century ago [97], and thereby affects the maximum stable flowability one can reach with one matrix volume and composition.

ICI Rheocenter

Florian V. Mueller

225

800

Yield value [Pa]

700 600

1st MSc

y = -436.1ln(x) + 2820.5 R² = 0.8339

500 400

2nd MSc Eco Project

300

y = -159.6ln(x) + 1065.5 R² = 0.6156

200 100

Log. (2nd MSc) Log. (Eco Project)

0 200

300

400 500 600 Slump flow [mm]

700

800

Figure 8-10: Measured yield value vs. measured slump flow for non-air entrained mixes without segregation signs

8.2.

Coaxial-cylinders viscometer versus impeller-based

8.2.1. Yield stress and G-value Several different publications of Geiker et al. [127 and 128], O.H. Wallevik [353 and 354] and J.E. Wallevik [345, 347, and 348] indicate that one can determine the intrinsic rheology parameters with the coaxial-cylinders viscometer with reasonably good agreement from torque measurements. All viscometers need a calibration in order to allow calculations of the electrical resistances to torque resistances in Nm-units. A second computation step is needed to convert the torque points into Pascal–units for which the Reiner-Riwlin equation (Equation 1-47 on page 24) is used in the coaxial-cylinders viscometer. Due to the different physical principles applied in the impeller-based system, this equation cannot be applied and I am not aware of any suitable replacement equation that could be used instead. Despite some extensive computational modelling procedures that could be applied [348 and 351], a semi-empirical approach could be used to correlate the torque from the Rheometer-4SCC (R4) to the results obtained in the Viscometer 5 (Vi5). Despite the original approach to correlate Ampere-units (from R4) with Pascal-units (from Vi5), see [361], it has been observed in [232] that the electrical resistances in Ampere-units can differ for different Rheometer4SCC devices due to slightly altering engine capacities. The particular device used at the ICI laboratory follows Equation 8-18. Therefore, the calibration constant always needs to be applied to convert the results initially given at the software interface into Nm-units according to Equation 8-19 and Equation 8-21. Only these parameters should be used in the first approach to compare results of different devices. The computation into Pascal-units follows for the G-value Equation 8-20 and for the H-value Equation 8-22, for which the coefficients k1 and k2 are empirically derived. Equation 8-18

𝑇𝑁𝑚 = 0.0138 𝑇𝑚𝐴 − 0.7934 Equation 8-19

𝐺𝑁𝑚 = 𝑘𝑔 ∗ 𝑇0,𝑚𝐴 With kg = 1/503.52 = 0.001986

226

Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Equation 8-20

𝐺𝑃𝑎 = 𝑘1 ∗ 𝐺𝑁𝑚 Equation 8-21

𝐻𝑁𝑚 𝑠 = 𝑘ℎ ∗ 𝑇 𝑚𝐴 (𝑓0 ) With kh = 1/503.52 = 0.001986 Equation 8-22

𝐻𝑃𝑎 𝑠 = 𝑘2 ∗ 𝐻𝑁𝑚 𝑠 Figure 8-11 depicts all results68 overlapping from the coaxial-cylinders viscometer and the impellerbased Rheometer-4SCC. This figure is separated for stable mixes (Figure 8-12) and mixes for which obvious segregation signs were reported (Figure 8-13). 300

Yield value [Pa]

250 200

y = 116.5x R² = 0.6516

150

All Linear (All)

100 50 0 0.0

0.5

1.0

1.5 2.0 G- value [A]

2.5

3.0

Figure 8-11: Yield value (Vi 5) versus measured G-value (R-4SCC)

Here, all results used are obtained from recent parts of the Eco-project, indicated by the label “Ecoproject” in the previous chapter. 68

ICI Rheocenter

Florian V. Mueller

227

300 y = 105.83x R² = 0.8796

Yield value [Pa]

Yield value [Pa]

300

200

100 All stable

200 Segregated >100Pa

100

0 0

1G-value [A]2

0

3

Figure 8-12: Figure 8-11 screened for stable mixes only

y = 111.53x R² = 0.0119

0

1 2 G-value [A]

3

Figure 8-13: Figure 8-11 for mixes with segregation signs and yield value > 100 Pa

Using the calibration coefficient from Equation 8-19 gives Figure 8-14, which has been separated into stable mixes (Figure 8-15) and unstable mixes (Figure 8-16). In order to emphasise the different ranges the results are placed into, the axis of abscissae is proportionally enlarged. 300

Yield value [Pa]

250 200

y = 53.207x + 6.5413 R² = 0.6607

150

All Linear (All)

100 50 0 0

1

2

3 G-value [Nm]

4

5

6

Figure 8-14: Yield value – G-value relation after second computation step for all results

228

Florian V. Mueller

ICI Rheocenter


Yield value [Pa]

250 200

y = 45.718x + 6.7667 R² = 0.9102

150

All stable Linear (All stable)

100

Linear (All stable)

y = 53.285x R² = 0.8796

50 0 0

1

2

3 G-value [Nm]

4

5

6

Figure 8-15: Yield value – G-value relation after second computation step for results of stable mixes

Yield value [Pa]

300

Segregated 200 >100Pa y = 29.416x + 27.497 R² = 0.3599

100

Linear (Segregated)

y = 56.156x R² = 0.0119

Linear (Segregated)

0 0

1

2

3 G-value [Nm]

4

5

6

Figure 8-16: Yield value – G-value relation after second computation step for results of unstable mixes

Multiplying the results of the second computation step by a factor k1 of 55 according to Equation 8-20 gives Figure 8-17, which again has been separated into Figure 8-18 and Figure 8-19.

ICI Rheocenter

Florian V. Mueller

229

300

Yield value [Pa]

250 200

y = 1.0666x R² = 0.6516

150

All Linear (All)

100 50 0 0

50

100 150 200 G-value [Pa]

250

300

Figure 8-17: Yield value (Vi 5) vs. G-value (R-4SCC), using k1 of 55

300

y = 0.9688x R² = 0.8796

Yield value [Pa]

Yield value [Pa]

300

200 y = 0.8312x + 6.7667 R² = 0.9102

100 All stable

0 0

100 200 G-value [Pa ]

300

Figure 8-18: Yield value vs. G-value for stable mixes only (VSI =0)

200 Segregated >100Pa

100

y = 1.021x R² = 0.0119

0 0

100 200 G-value [Pa]

300

Figure 8-19: Yield value vs. G-value for mixes with higher VSI >0 and yield value >100 Pa

Remarkably, the impeller based Rheometer-4SCC obtains similar results as the coaxial-cylinders viscometer for a variety of mixes where segregation signs were reported, see Figure 8-19. Nevertheless, a separation of the results into two branches can be observed. They have to be attributed to mixes of larger segregation risk. Then, the segregation already affects the samples taken and the one used for the coaxial-cylinders viscometer is of different composition. Thereby, the “real” both (G-) yield values either are under- or over-determined. Additionally, even for similar samples in both devices, it would alter the results when segregation occurs during a test; see the discussion in Jon E. Wallevik’s thesis [345]. The somewhat misplaced computational G-values obtained for some mixes having a higher yield value are caused by the viscosity of these mixes, which might be considered relatively high. It seems that such a combination of rheology parameters are beyond the suitable application range for the impeller used, which is specially designed and optimized to detect the very small resistances occurring in SCC.

8.2.2.

Plastic viscosity and H-value

Figure 8-20 depicts the correlation of plastic viscosity determined in a Viscometer 5 to the Hviscosity determined with the portable Rheometer-4SCC. This figure has been separated into Fig-

230

Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC ure 8-21 and Figure 8-22, applying the corresponding data to those results used in the previous chapter.


120 y = 11.542x R² = 0.7876

80

All Linear (All)

40

0 0

4 8 H-value [A s]

12

Figure 8-20: Plastic viscosity (Vi 5) versus H-value (R-4SCC) as determined from electrical resistances measured

120

y = 11.945x R² = 0.7162

100



120

80 y = 10.096x + 4.3788 R² = 0.7451

60 40 20

Stable

0 0

2

4 6 8 H-value [A s]

10

100 80 60

Segregated

40

>100 Pa

y = 11.078x R² = 0.7094

20 0

12

0

Figure 8-21: Figure 8-20 screened for stable mixes only

2

4 6 8 H-value [A s]

10

12

Figure 8-22: Figure 8-20 screened for mixes with obvious segregation signs or yield value >100 Pa

Using the calibration coefficients from Equation 8-21 on page 227 gives the results of the second computation step, shown in Figure 8-23 and separated in Figure 8-24 and Figure 8-25.


120 100 80

y = 5.2262x + 4.0033 R² = 0.8009

60

All

40

Linear (All)

20 0 0

4

8

12 16 H-value [Nm s]

20

24

Figure 8-23: Plastic viscosity (Vi5) vs. H-value (R4-SCC) for all results at second computation step (torque) ICI Rheocenter

Florian V. Mueller

231


120 100 y = 6.0147x R² = 0.7162

80

Stable

60

Linear (Stable)

y = 5.0837x + 4.3788 R² = 0.7451

40 20 0 0

4

8

12 16 H-value [Nm s]

20

24

Figure 8-24: Figure 8-23 screened for stable mixes


120 100 80 y = 4.4707x + 7.6147 R² = 0.7694

60

Segregated

40

>100 Pa

20 0 0

4

8

12 H-value [Nm s]

16

20

24

Figure 8-25: Figure 8-23 screened for unstable mixes

Multiplying the derived results of the H-value by a factor k2 of 6.16 according to Equation 8-22 gives Figure 8-26, which has been separated for stable mixes (Figure 8-27) and segregated mixes (Figure 8-28).


160 y = 0.9434x R² = 0.7876 120

80

All Linear (All)

40

0 0

40

80 120 H-value [Pa s]

160

Figure 8-26: Plastic viscosity versus computed H-value [Pa s], using k2 of 6.16

232

Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC 160 y = 0.9764x R² = 0.7162

120



160

y = 0.8253x + 4.3788 R² = 0.7451

80 40

Stable

0 0

40


160

Figure 8-27: Figure 8-26 screened for stable mixes

120 80

Segregated >100 Pa

40

y = 0.9055x R² = 0.7094

0 0


120

160

Figure 8-28: Figure 8-26 screened for mixes with obvious segregation signs or yield value >100 Pa

Similar to the yield value, the evaluation with stable mixes allows the determination of the H-value with a distinct narrow variance; see Figure 8-27. For the mixes considered as unstable (Figure 8-28), the variability does not increase to the same extent as it does for the yield value. As long the overall rheology is suitable for the “SCC-impeller”, which is optimized for low resistances, the Hvalue determined in the portable Rheometer 4-SCC seems to be only little affected by segregation effects and is in good agreement with the results obtained in the coaxial-cylinders viscometer. The coefficients k1 and k2 (Table 8-2), used to convert the data from the calibrated Nm-units into the physical Pascal-unit commonly considered for rheology aspects are established using the data of only all stable mixes. Then, the average value of all the yield values determined in the coaxialcylinders viscometer (𝜏̅0,𝑉𝑖5 ) is set equal to the average value of the computed G-values obtained from the Rheometer-4SCC (𝐺̅𝑇0,𝑅4 ). The accuracy one aims for is that both terms are equal up to the second decimal. The coefficient k2 for the plastic viscosity and H-plastic viscosity is determined accordingly. Table 8-2: Statistical parameters concerning the conversion from Nm-units into Pascal-units

k1 k2 55.0 6.16 Number of data points Maximum Minimum

𝝉̅𝟎,𝑽𝒊𝟓 [Pa]

40.08 56 94.0 22.0

̅ 𝑻𝟎,𝑹𝟒 𝑮

̅ 𝑽𝒊𝟓 𝝁

̅ 𝑹𝟒 𝑯

[Pa]

[Pa s]

[Pa s]

40.07 56 105.8 22.5

25.04 56 55.4 13.7

25.03 56 51.7 10.6

In [232], already a first approach was made to establish these correlation factors, using an analysis of the trend curves at the time. When applying the same approach as introduced before, the following coefficients are derived (Table 8-3), again only considering stable mixes. Table 8-3: Statistical parameters concerning the conversion from Nm-units into Pa-units for data from Mueller [232] using a thixotropy setup

k1 k2 78.73 7.25 Number of data points Maximum Minimum ICI Rheocenter

𝝉̅𝟎,𝑽𝒊𝟓 [Pa] 97.17 357 497.0 8.0

̅ 𝑻𝟎,𝑹𝟒 𝑮 [Pa] 97.17 357 490.2 7.8

Florian V. Mueller

̅ 𝑽𝒊𝟓 𝝁 [Pa s] 41.99 357 139.4 17.1

̅ 𝑹𝟒 𝑯 [Pa s] 42.00 357 233.6 10.1 233

Remarkably, the coefficients k1 and k2 differ between Table 8-2 and Table 8-3, considering the large range of mix design parameters that were covered in both projects and which also overlapped in w/b and solid concentrations to some extent. The only obvious difference is that during the evaluations to derive the parameters in Table 8-3, a thixotropy setup was used in the Viscometer 5 for all measurements. For the evaluations of the parameters in Table 8-2 usually a shorter setup was used due to stability considerations, see the discussion in chapter 3.1.4. The thereby slightly altered rheology parameters (compare Table 3-11 on page 126) are suspected to cause the differences of the converting coefficients mentioned above. In summary, we established two sets of correlation factors - one set when using the thixotropy setup, and another when using the shorter setup.

8.3.

Validation of the proposed correlation models

For all the mixes from chapter 4, even for which segregation signs were observed, the slump flow was computed with all three approaches mentioned above in Table 8-1 on page 223. The difference between the measured slump flows and the corresponding computed ones are presented in Figure 8-29. In order to mitigate the effect of potentially occurring workability loss, the average yield value was not used, but rather the measurements that were the closest in time to the slump flow test. For each mix, the actually measured density has been employed. As indicated before, the original approach from LCPC69, incorporated in “Model 01,” overestimates the slump flow considerably for the majority of the mixes. “Model 02” serves particularly well in the region of SelfCompacting Concrete with a variance below 10% (see Figure 8-30), whereas for “Model 03” the variability increases slightly but improves the prediction of the mixes with higher yield value.

Difference calc. - meas. Sf [mm]

300 200 100

Model 01 Model 02

0 0

20

40

60

80

Model 03

-100 -200

Yield value [Pa]

Figure 8-29: Difference of measured to computed Sf in respect to its yield value

When normalizing the differences occurring to the slump flow actually measured for each, the increasing variability (see Figure 8-30) for the lower ratio of Bingham parameters (compare also Figure 8-31) might have to be attributed to the instability caused by it, i.e. increasing sedimentation risk with smaller parameters (Figure 8-32). Due to the relatively very-low plastic viscosity for our mixes in general, the mixes with reducing yield value can exhibit an unfavorably low ratio of Bingham Refering in a generalizing attempt to the affiliation of Hu [161 and 162] and DeLarrard [85], at least the affiliation at the time when the work was published. 69

234

Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC parameters and therefore potentially reveal a lack of both static and dynamic stability. Such instability affects both measurements - the determination of slump flow and the Bingham parameters determined in any rheometric device.

Normalized diff. Sf [%]

50 40 30 20

Model 01

10

Model 02

0 -10

0

1

2

3

4

5

6

7

8

Model 03

-20 -30 Yield value / pl. Viscosity

Figure 8-30: Normalized difference between calculated and measured Sf, with respect to yield value /plastic viscosity as indication of stability and self-compacting properties

Pl. viscosity

60

30

50 40

20

30 20

10

10 0

0 0

1

2 3 4 5 6 7 Yield value / pl. Viscosity

8

Figure 8-31: Overview on rheology parameters measured for the mixes in Figure 8-30

ECC

70 Yield value [Pa]

Yield value [Pa]

70

80

40

Yield


80

All SCCArea

60 50 40 30

SCC

20

Sedimentation

10 0 0

10 20 30 40 50 Plastic viscosity [Pa s]

60

Figure 8-32: Rheograph of all results discussed in the figure above

When estimating the yield value from measured slump flow values, the results (Figure 8-33) are located in the vicinity of the actually measured yield values; compare Figure 8-10 on page 226. When comparing the measured with the estimated yield values in greater detail (Figure 8-34), it reveals that the differences for those mixes considered as stable are within an accuracy of ±10 Pa. Some mixes that were considered stable might have revealed a dynamic sedimentation risk due to the very low viscosity. This is attributed to air-entrainment into rather small matrix volumes of 331 l/m3 (“AEA_B0_180”). Therefore, the variability of their differences increases to the vicinity of ±20 Pa (Figure 8-34). Only for mixes for which obvious segregation signs affected the results, the computation reveals a larger scattering beyond ±20 Pa. Such mixes were for instance the SPoversaturated reference mixes of the SMA series, the mixes with too high dosage of AEA but with an appropriate water content (e.g. in series “AEA_B0_180”), or any dosage of AEA with inappropriately high water content (i.e. “AEA_B0_185”).

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Florian V. Mueller

235

Yield value, estimated [Pa]

800 700 600 GBF

500

L

400

FA

300

AEA_B150_190

200

AEA_B0_185

100

AEA_B0_180

0 200

300

400 500 600 700 Slump flow, measured [mm]

800

Figure 8-33: Estimated yield value for measured slump flow

Differences Yield value [Pa]

60 40 GBF L

20

FA AEA_B150_190

0 550

600

650

-20 -40

700

AEA_B0_185 AEA_B0_180

Slump flow, measured [mm]

Figure 8-34: Difference of estimated and measured yield value in respect to its measured slump flow

The correlation between Bingham parameters obtained in the portable (impeller-based) Rheometer4SCC and the stationary (coaxial-cylinders-based) Viscometer 5 has been analyzed for the same mixes. When considering all results in Figure 8-35 and Figure 8-36, they show a reasonably good trend and relatively narrow variance, considering the fact that segregating mixes are also included. When excluding these mixes, the average yield value of all 61 results considered misses the estimated G-yield value by only about 6 Pa (i.e. 16%), whereas the estimated H-plastic viscosity can be expressed in even higher accuracy of only 1.8 Pa s ( i.e. 8%) , see Table 8-4. Consequently, the rheograph for both devices agrees well for all matrix volume mixes, see Table 8-5. The small computational deviations of Bingham parameters and their location in the different rheographs seem to exaggerate the increased sedimentation and segregation tendency for some of the mixes, compare Figure 8-31. Using the portable Rheometer-4SCC could thereby assist the identification of mixes with increased risk of sedimentation.

236

Florian V. Mueller

ICI Rheocenter


Yield value [Pa]

80 60

GBF L

40

FA AEA_B150_190 AEA_B0_185

20

AEA_B0_180 0 0

20

40 G- value [Pa]

60

80

Figure 8-35: Measured yield value versus computed G-value, using k1 of 55.0


60 50 GBF

40

L

30

FA AEA_B150_190

20

AEA_B0_185

10

AEA_B0_180

0 0

10

20 30 40 H-value [Pa s]

50

60

Figure 8-36: Measured plastic viscosity versus computed H-value, using k2 of 6.16 Table 8-4: Statistical parameters considering stable mixes only

K1 K2 55.0 6.16 Number of data points Maximum Minimum

𝜏̅0,𝑉𝑖5 [Pa]

35.1 61 77 15

𝐺̅𝑇0,𝑅4

𝐺̅𝑇0,𝑅4 [Pa s]

[Pa s]

29.1 61 64 11

22.3 61 29.1 10.6

20.5 61 31.5 5.1

[Pa]

̅𝑅4 𝐻

Of course, the results in Table 8-5 are not identical, but given the many influences on a measurement, particularly when using different testing methods, the results might be regarded as accurate enough.

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Florian V. Mueller

237

Table 8-5: Rheograph for stable mixes of matrix series (chapter 4) obtained either in a coaxial-cylinders Viscometer (a) or in an impeller-based Rheometer-4SCC (b) GBF

80

80

60

FA

40

AEA_B150_19 0 AEA_B0_185

G-Yield value [Pa]

Yield value [Pa]

L

AEA_B0_180

20

SCC-area

0 0

10 20 30 40 50 Plastic viscosity [Pa s]

40 20 0

60

0

(a) Rheograph, measured results in Viscometer 5

8.4.

60

10 20 30 40 50 60 H-Plastic viscosity [Pa s]

(b) Rheograph computed results

Concluding remarks

It was successfully shown how to establish the equations that are required to estimate the yield value based on stable slump flow and vice versa. While the original approach alone (“Model 1”) does not yield satisfactory results for the fluid mixes, a modification was made to assess only the fluid mixes (“Model 2”). Unfortunately, this model was particularly weak for stiffer consistencies, yielding a second modification (“Model 3) that covers the full range in a satisfactory manner. Each model contains the following: slump flow, or an equation to derive the slump flow from slump value (Equation 8-5, p. 219), a term to consider the relation between yield value and slump flow (Equation 8-6, p. 220), a term to transform the results obtained in a BTRHEOM into ConTec Viscometer 5 results (Equation 8-12, p. 223 for “Model 2”, Equation 8-13, p. 223 for “Model 3”), in order to adapt ((Equation 8-15, p. 223), and the existing approach from Hu and DeLarrard (Equation 8-9, p. 221), empirically derived from the BTRHEOM. Thereby, the “τ0,BMLSfModel for ConTec” could be derived. Verification with a set of mixes revealed an accuracy of about ±10 Pa for stable mixes. For the correlation of the results obtained in the portable impeller-based Rheometer-4SCC with the intrinsic physical results obtained in the co-axial cylinders viscometer, the following remark can be made: It was found to be necessary to distinguish whether a thixotropy setup was used to derive the Bingham parameters, having a longer total shearing time then, or whether a flow curve of torque points was recorded when only a down-curve was executed. The G-value can be transformed using the coefficients k1 of 55.0 (78.73 for thixotropy setup) in Equation 8-20, p. 227, and the H-value can be transformed into Pascal-units using the coefficient k2 of 6.13 (7.25 for thixotropy setup) in Equation 8-22, p. 227. Verifications reveal not identical, though satisfyingly similar results given the multiple effects affecting the measurement, in particular when different methods are employed.

238

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9. Final remarks The original approach of Eco-SCC according to the idea formulated by O.H. Wallevik represents an economical and ecological alternative to commonly designed SCC and CVC. The advantage arises through limiting the binder volume that is considerably reduced compared to common SCC and in the vicinity of common CVC. According to the definition made in [363], the binder material content is set to about 10 vol.-percent, equivalent to about 315 kg/m3. In order to provide sufficient lubricant volume, the w/b is relatively high when compared with common SCC, and is approximately 0.57 to 0.60, that is equivalent to 180 to 190 kg/m3 of water. Thereby, the concrete strength class can be lower than for common SCC and the often-used day-to-day strength class of about C25/30 could be approached as ready-mix application, but with the advantages when using Self-Compacting Concrete. The main aspect when designing a low-binder SCC is to determine the lowest matrix volume suitable for a given aggregate PSD, considering its shape, texture, and maybe other practical limitations. The reduction of cement clinker content alone, when replacing it with SCM or inert filler as it is done in many approaches published, can potentially cause difficulties in durability and volume deformation aspects due to the relatively reduced strength and E-modulus, when the high paste volume as such remains constant. It has been shown in this thesis that it is possible to design SCC with 315 kg/m3 of binder or even less, but it has been also shown that it is actually the matrix volume that needs to be considered as the major parameter, see Chapter 4. This is in order to fulfil a minimal blocking criteria while revealing self-compacting rheology properties with a reasonably good static and dynamic sedimentation stability. Therefore, the definition of Eco-SCC would need to be reviewed. The minimum matrix content for an optimized PSD has been determined to be around 360 ± 10 l/m3. The difference between paste volume and matrix volume is the consideration of air, e.g. by applying the Air Matrix Approach (AMA), and of fines from the aggregates that are smaller than 125 micrometres. Such consideration of aggregate fines was actually already suggested in the original Rational Mix Design Approach (RMDA) by Okamura et al. (although a different size of 0.09 mm was suggested then), but it has appeared lately to not be considered very important when designing SCC in a “traditional” way (compare the definition of powder in the JSCE recommendation finally published based on the RMDA). To follow the matrix approach instead of the paste volume approach results in additional volume of 20-40 litres (for non-AEA) mixes that are considered for the mix design. Not accounting for it will always require relatively larger powder content in order to increase the stabilizing plastic viscosity of the mix. When the matrix volume for a smaller paste volume is established by using filler material instead, independent of the reactive nature of the material used, I suggest interpreting the design approach as Solid Matrix Approach (SMA). It is differentiated in such a way as with the AMA approach the paste volume can be considerably reduced and thereby the binder content. Additionally, the freeze-thaw resistance can be obtained for such low-paste SCC and air-entrainment into binder-rich SCC, as the JSCE recommendation suggests, can be avoided. For such mixes, it can cause a higher risk of static sedimentation due to the sheer volumes involved. The compressive strength was found to vary a lot, depending on the binder material used and the w/b actually applied. Some mixes would fulfil the strength criteria according EN 206 as low as C16/25, some of the AMA approach would be rated as C40/50, whereas others of the SMA approach would be rated up to C50/60. It implies that by adapting the binder material according to the needs a broad range of compressive strength classes could be reached for a similar matrix volume. The only challenge will be to maintain the stability and workability retention. Due to the diversified properties of binder materials and their influence on the admixture interaction, one cannot make a proposal for the most suitable binder and admixtures, even though favourites from among ICI Rheocenter Florian V. Mueller

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those used in the project were identified. One would prefer materials where the main constituents are of high purity, such as a combination of OPC (clinker), silica fume (SiO2) and limestone filler (CaCO3). The effect of the fluctuating parameters in fly ash would demand a higher control effort, but unground and classified FA has been used successfully. Due to the improved workability retention of PCE against most of the conventional superplasticizers, one would recommend using a PCE-based SP with low retardation effect. For the stabilizer, very good experiences were made with densified silica fume in a dosage that does not increase the 28d strength, but has a stabilizing effect on the fresh concrete. The effectiveness of a suitable stabilizer should not vary considerably during the period considered for ready-mix concrete, as was observed for some products. The gyratory Intensive Compaction Tester (ICT) was used in Chapter 5 to verify the optimum PSD that would require only a low matrix volume to obtain a stable SCC. Some of the different PSD’s were also verified in fresh concrete tests. It was concluded that the PSD for reasonable shaped solids should follow the formula according to the (by Funk and Dinger) modified Andreasen & Andersen model, applying the packing exponent of 0.20. The bulk density obtained during the PSD optimization follows a parabola curve with an optimum, i.e. highest solid concentration, when manipulating different fractions and determining the resulting solid concentration. When adding the powder, water, and air content as lubricant volume to the aggregate particles, the flowability is obtained for a solid skeleton that is near its optimum density, and not already oversaturated with powder as the representative of a common SCC was. Such oversaturation would yield a granular bulking and a reduction of the densest packing without revealing a superior blocking characteristic, but only a larger flow. For applications where such high flowability is not required, the use of lowbinder SCC represents a suitable alternative, making it possible to maintain the advantages of selfcompacting properties, while reducing the powder content. Through activating the particle lattice effect for stabilization when applying such optimized PSD allows for a low-viscous matrix in which the water and air content can be relatively increased compared to the matrix proportion in a conventionally designed SCC. Such an approach could also reduce the carbon footprint of the concrete volume and would therefore contribute to the sustainability of the cement and concrete industry. Going from the proposed Modified A&A-model to a continuous grading according to the Fuller & Thompson equation, it demands a larger powder content to establish the stabilizing plastic viscosity. This is needed in order to prevent the sedimentation of the relatively increased content of larger aggregates when the supporting lattice effect of a well-graded PSD according to the modified Andreasen & Anderson model cannot be utilized. Table 9-1 provides a recommendation for matrix volume proportions of stable and robust lowbinder SCC. The possible further reduction of matrix volume when using a low-viscosity approach, as indicated by Mørtsells Particle-Matrix Model for HPC, can still be observed when comparing the total volumes used in the AMA and SMA, although it is not a large volume difference when applying the optimized PSD. Such an approach was used for the mix presented in a talk at the conference “RheoIceland” (Figure 9-1), which was held in Reykjavik in 2009. Since the colour pigments affected the AEA, causing some adaptions of the mix design, the transformation of the mix design from laboratory to ready-mix industry is not presented in this thesis.

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table 9-1: Approach for recommended volume proportions of minimized matrix composites for stable and robust SCC

Solid Matrix Approach Vpowder Vwater Vair Vtotal

Air Matrix Approach

[l/m3]

[% per matrix]

[l/m3]

[% per matrix]

160 190 20 370

43.4 51.3 5.4 100

135 180 50 365

37.0 49.3 13.7 100

Figure 9-1: The author, supervising the filling of a mold representing the institute’s acronym ICI with redcoloured Eco-SCC during the conference “RheoIceland 2009”

One important result of this thesis is the proposal of a new sedimentation test and new ranking methods in chapter 6. Since the approach is applied in hardened concrete, and therefore is more suitable during the development phase of a mix, it provides very reasonable results to determine the static sedimentation resistance and might additionally have a potential to indicate the dynamic segregation resistance. One approach to quantify the dynamic segregation resistance is suggested, but needs to be verified with more tests, and comparison to the results of other tests in fresh concrete on dynamic segregation. Within this thesis, the static sedimentation was studied based on the effect it has on the variation of the specific gravity in a sedimentation specimen. It is ranked in Classes of Specific Gravity Distributions (CSGD) of cylinders with a diameter of 150 and a height of 300 mm. Comparison studies revealed that this dimension is sufficient to describe the static sedimentation effect and larger sized columns (100/600 and 150/600) according to the common tests are actually not needed. This ranking includes not only the possibility to consider the effect of dynamic sedimentation, but also whether the mix is fully consolidated at all when introducing one group for semi-SCC, sometimes considered as Easy Compacting Concrete (ECC). Additionally, an Aggregate ICI Rheocenter

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Surface Appearance Index (ASAI) has been developed that shows, due to the subjective nature of ranking by an operator, a similar weakness as the Hardened Concrete Visual Stability Index (HVSI) proposed by Shen [304] or the Visual Stability Index (VSI) according to ASTM 1611. Still, for mixes containing very low matrix volumes it can bring valuable qualitative information about the selfcompactability. The approach used for the determination of the robustness of one reference mix design included the variation of superplasticizer dosage in steps of ± 10 and 20%, and the variation of the water content in steps of ± 5 and 10 l/m3. The comparison of three mixes, representing one Eco-SCC, one lean SCC, and one binder-rich SCC, revealed a superior robustness of mixes that were originally designed with a smaller matrix volume and higher relative water content, in agreement with the reports of Lowke et al. [212] and Billberg [34]. Therefore, the robustness of Eco-SCC is considerably improved compared to the binder-rich SCC. In particular the non-reversible effects, i.e. segregation and increased sedimentation risk, when increasing the water content or SP dosage in binderrich SCC, are of bigger disadvantage than the reversible effects commonly occurring in low-binder SCC. Here, a lack of lubricant would result in insufficient self-compacting properties, but it could be easily readjusted by adding the appropriate composite according to conclusions drawn from the rheology status. For the reference mix tested as Eco-SCC and Lean SCC, an over-dosage of SP or water did not yield a severe segregation or increased sedimentation risk. When combining the rheology response of the robustness test of binder-rich SCC series and their sedimentation properties it raises the question whether the often reported shear thickening behaviour of binder-rich SCC is an intrinsic material property accurately described by one of the existing non-linear fluid models, e.g. Herschel-Bulkley [84] or Modified Bingham model [114, 115, 151, 152, and 153]. In particular the sedimentation test results are indicators that it could be an artefact measurement when a lack of dynamic segregation resistance occurs in mixes with relatively high paste volume with probably too low plastic viscosity. More research could clarify the situation so more tests are recommended. In conclusion, a rheograph with refined areas of Bingham parameters is proposed (Figure 6-34 on page 195), which considers not only the matrix design and its effect on rheology, blocking characteristics and the implications it has on sedimentation stability, but also some effects of aggregate properties. It indicates that the very low-viscous SCC should only be obtained by air-entrainment using a sufficiently small matrix volume for stability reasons by applying the Air Matrix Approach. Mixes in which the aggregates already cause such very low viscosity (between 10 and 20 Pa s), e.g. due to their even surface texture, particle shape and lack of fine particles, should actually contain a higher powder content than defined for Eco-SCC in order to provide a robust solution while mitigating the risk of static sedimentation and/or dynamic segregation. It was found that low paste volume alone only causes a lack of stability when the PSD is not adapted sufficiently. I consider the approach as Solid Matrix Approach (SMA), as the common gradation according to the Fuller and Thompson equation to the recommended version of the modified Andreasen and Andersen equation is changed by adding solid particles. The reactive nature of the added solids has an impact on various performance parameters in fresh and in hardened state. Therefore, it can be used to manipulate these properties in order to meet additional specific durability requirements. The rheology tests included the common slump flow test accompanying the determination of the intrinsic Bingham model parameters in a coaxial-cylinders viscometer. Most measurements were performed simultaneously with the portable impeller-based Rheometer 4-SCC. Due to a lack of mathematical equations to derive the rheology parameters from any impeller-based rheometric equipment, an empirical relation of the measured result to those obtained with the ConTec Vis242

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC cometer 5 could be established and successfully verified. Additionally, an existing approach from (the former) LCPC to estimate the slump value from measured yield values for the BTRHEOM (“τ0,BTSM for BTRHEOM”) was adapted for the ConTec Viscometer and extended for largerspread slump flow values (“τ0,BMLSfM for ConTec”), see Equation 8-15 on page 223. The validation for selected results indicated a very good agreement between measured and estimated results, of course only for stable mixes. In the contrary approach, the estimation of yield value from a measured stable slump flow value could only be established for more fluid concrete and SCC. It was validated for several cases that when applying the model one is able to predict the yield value with an accuracy of ± 10 Pa for stable mixes. As expected, the accuracy weakens when using unstable mixes. When in future work an approach can be found to estimate the plastic viscosity from simple tests with reasonable accuracy, it could serve to estimate these rheology parameters in order to allow their ranking in a rheograph. Thereby, a brief overview can be made if these parameters match the requirements for one specific application and it could reduce the control effort needed for other acceptance tests. For the ConTec Viscometer 5 itself, the need to adapt setup parameters for such low-viscous SCC has been identified, various setup parameters tested, one suitable set proposed and consequently used thereafter. Under the circumstances, the very low viscosity can cause a lack of dynamic segregation resistance, which affects the rheology measurement through particle migration. In such cases, the results determined would not be reliable and can easily yield a misleading interpretation of the capabilities of the mix.

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10. Conclusions From the work presented here, the following conclusions can be drawn: 1. Self-Compacting Concrete requires a specific critical matrix volume in order to satisfy the blocking criteria at the maximum fluidity of the fresh concrete. This matrix volume is independent from the paste volume and governed by the packing characteristics of the aggregates. 2. The paste volume is only part of the matrix volume and is not the important factor to be considered to assess the quality of SCC. 3. For stability reasons, the maximum fluidity should only be obtained with a stable matrix, i.e. not based on oversaturation with a dispersant admixture. Therefore, the rheology properties of the matrix volume need to ensure stability of the completely concrete system and its components need to be arranged accordingly. The rule of thumb is the lower the plastic viscosity, the lower the matrix volume should be chosen for stability reasons. To the contrary, when designing for low viscosity due to high w/b and /or air-entrainment, or achieving it based on aggregate characteristics, a lower matrix volume should be chosen. 4. Since SCC properties are governed by the matrix –aggregate interaction, the variation of its components allows for covering a broad range of strength- and durability-performance classes. Investigations on the matrix alone might not result in an optimized mix design of the concrete due to insufficient consideration of the coarse aggregates’ effect. 5. The critical matrix volume can be achieved using three different approaches: a. Considering the paste volume only, which results in relatively large binder content as powder material in order to provide sufficient viscosity for stability (such approach is basically used in common SCC mix design methods); b. With relatively small content of paste and additional air entrainment of about 5±1 vol.-% that could be used additionally to meet freeze-thaw requirements, so called Air Matrix Approach (AMA); c. With relatively small content of paste and additional filler, so called Solid Matrix Approach (SMA). Hereby the reactive nature of the different filler materials available can be utilized to fulfill additional durability requirements. Remark: While in approach A a stabilizer (ST) can be added to improve the robustness of a mixture, it is mandatory to add ST in approaches B and C. 6. The conventional fluidity approach for SCC in terms of slump flow values used in standards and official guidelines is valid for binder-rich and lean SCC only, for which it has been developed. Applying advanced rheology approaches considering the completely concrete system allows optimizing SCC mix designs in respect to their rheology parameters, in order to obtain self-compaction and sedimentation stability. 7. An optimized PSD is proposed for which rheology, blocking, and sedimentation tests were considered together with gyratory intensive compaction tests. This PSD would activate the particle lattice effect to stabilize the coarse aggregate, while requesting the smallest matrix volume. With the stabilizing aid of the particle lattice affect, a lower viscosity can be applied which can lead to lower strength classes due to possibly higher w/b. Mixes with PSD varying from the proposed one would require additional matrix volume due to granular bulking that occurs when oversaturating a rather monosized fraction (i.e. powder), which is on the other side needed to ensure stability with the viscosity-increasing effect of powder addition. Stabilizing admixtures seem not to enhance plastic viscosity considerably for the 244

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8.

9.

10.

11.

cases evaluated here, which have relatively low powder content and relatively high interparticle distances due to relatively high w/p. A new sedimentation test in hardened state is proposed, including two new evaluation- and ranking-methods, when rating the sedimentation-induced variation of specific gravity into classes according to their distribution in vertical direction of the sedimentation specimen (CSGD). The main advantages are that the subjective character of common visual tests can be mitigated, the effect of dynamic sedimentation is included, the rating system is extended to identify insufficiently self-compacting mixes, and the test can be easily executed with minimal requirements for complex equipment used. It was shown that common cylinder of diameter 150 and height 300 mm could be successfully used to evaluate the static sedimentation of SCC with maximum aggregate size up to 16 mm. The recommended larger size of d100 (150) / h 450 (500) mm according to guidelines is actually not needed. The new method might also be extendable to evaluate the influence of dynamic sedimentation. An existing slump – yield value model has been further developed when adapting it to slump flow and the yield value one obtains when using a ConTec Viscometer 5 (“τ0,BMLSfM for ConTec”). It allows for estimating the yield value based on a stable slump flow with an accuracy of ± 10 Pa, which can be used to assess the mix in a rheograph when also the plastic viscosity can be estimated. For the latter, still further research is needed. The intrinsic rheology parameters obtained by the coaxial-cylinders viscometer could successfully be correlated with the corresponding parameters obtained when using the portable impeller-based Rheometer -4SCC. It allows for calculating the intrinsic rheology parameters from torque measurements of the impeller-based system of the portable device, for which the alternative mathematical approach would be a complex and time consuming iteration process since a simple equation such as the Reiner-Riwlin equation does not exists.

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11. Outlook into future research The main aspect when designing a low-binder SCC is to determine the lowest matrix volume suitable for a given aggregate, considering its shape, texture, and PSD, and maybe other practical limitations. Hereby the specific casting condition needs also to be considered. Therefore, more research is suggested in order to quantify the packing characteristics for a broader range of particle shapes. Finally, it should enable the designer to preselect the paste volume required to achieve a specific minimum matrix volume, based on some key parameters obtained from a packing test of the aggregates. Additionally, the quantification of packing characteristics depending on PSD, particle shape, and surface texture could enable the aggregate producers to adapt and optimize their conditioning process during the production of aggregates. The water in excess of the content chemically required for full reaction of the binder materials is known to result in a capillary pore system. It could be reasonable to quantify the variation of porosity for different binder compositions at one specific matrix volume for the conditions in SCC, in order to access data needed for the lifetime prediction of the concrete structure considering different durability requirements. For mixes with relatively high w/b and in which SCM is incorporated in the binder, the superplasticizer content needed to achieve the fluidity state of an SCC is often relatively reduced compared to mixes with higher clinker content. The clinker phases usually have the highest polymer adsorption capacity due to the high Stern potential that immediately develops when adding water to the reactive binder. Nevertheless, the range of admixtures suitable to achieve the desired workability and maintaining it for requirements of ready-mix concrete seems to be somewhat limited. It could be beneficial for a widespread use of low-binder SCC to increase the research effort in order to develop more dispersing and stabilizing admixtures suitable for the specific demands of low-binder SCC. Additionally, the effect on time dependent slump loss of different setups concerning the shear history should be investigated. Some preliminary tests (not reported here) indicated significant differences for the same mix when treated differently during the investigation. Since thixotropy appears to be a function of both rheology parameters, the yield value as well as the plastic viscosity, the low viscous Eco-SCC should be assessed with respect to formwork pressure, both in absolute terms and in its decay. Then, also different casting techniques should be applied such as pumping from the bottom or filling from top. The mechanical properties of low binder SCC should be studied in detail. Hereby it is suggested to put special emphasis on different casting techniques to obtain vertical and horizontal elements, in particular concerning homogeneity and eventually occurring so-called “top-bar-effect”. In addition, the bond properties should then be studied. The compaction tests in the gyratory Intensive Compaction Tester revealed relatively higher densification than usually reported when using one of the more common methods, as for instance by De Larrard [85]. One hypothesis, see Table 1-5 on page 36, is that the shear compaction might slightly re-orientate some particles that were placed in the loose bulk with a very low contact number and thereby have a high degree of freedom, concomitant a low degree of stability. This supposed reorientation would cause the denser packing characteristics observed in gyratory ICT. This hypothesis could be studied with suitable image analysis, such as a computer tomography CT, which allows studying the particle orientation before and after the ICT test.

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC 100. 101. 102.

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC 146. 147.

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Stark, Jochen, Bernd Möser, and Frank Bellmann (2004): Hydratation von Portlandzement (German, Hydration of OPC), Compendium from F.A. Finger Institute für Baustoffkunde an der Bauhaus Universität Weimar, Germany (26.02.2011: http://www.uniweimar.de/Bauing/fib/lehre/skripte.htm#baustoff) Stark, Jochen, U. Frohburg and P. Nobst (2006): Uebertragbarkeit von Frost-Laborprüfungen auf Praxisverhältnisse, (German, Adaption of laboratory freeze-thaw-tests to real conditions) Abschlussbericht (Final report) DBV-Forschungsvorhaben Nr. 234, Fraunhofer IRB Verlag, ISBN 3-8167-7004-5 Stark, J. B. Möser, F. Bellmann, C. Röβler (2006): Quantitative Charakterisierung der Zemenhydratation (German, Quantitative characterization of cement hydration),in: Proceedings of XVI. IBAUSIL, Weimar, Germany, pp. 1-0047 to 1-0066 Stark, Jochen (2008): Alkali-Kieselsäure-Reaktion (German, ASR), Schriftenreihe des F.A. FingerInstitutes für Baustoffkunde an der Bauhaus-University Weimar No. 3, (in German) Stokes, G.G. (1851): On the Effect of Internal friction of fluids on the motion of pendulums, Transactions of Cambridge Philosophical Society, Cambridge UK, Vol. 9, Part II, pp. 8-106 Su, Nan, Kung-Chung Hsu, and His-Wen Chai (2001): A simple mix design method for Self-Compacting Concrete, Cement and Concrete Research, Vol. 31, pp. 1799-1807 Su, N. and B. Miao (2003): A new method for mix design of medium strength concrete with low cement content, Cement and Concrete Composites, Vol. 25, pp. 215-222 Talbot, Arthur N. and Frank E. Richart (1923): The strength of concrete - its relation to the cement aggregates and water, Bulletin No. 137, Engineering Experiment Station of University of Illinois, UrbanaChampagne, 122 page Tattersall, G.H. and P.F.G. Banfill (1983): The rheology of fresh concrete, Pitman Advanced Publishing Program Taylor, H.F.W. (1997): Cement Chemistry, 2nd Edition, Thomas Telford Services LTD, ISBN 07277-2592-0 Teichmann, T. and M. Schmidt (2004): Influence of the packing density of fine particles on structure, strength and durability, Proceedings of: UHPC Symposium, University Kassel, Germany, pp. 313322 Teichmann, T. (2007): Einfluss der Granulometrie und des Wassergehaltes auf die Festigkeit und die Gefügedichtigkeit von Zementstein,(German, Effect of granulometry and water content on strength and microporosity of hardened cement paste) Dissertation University Kassel, Structural Materials and Engineering Series No. 12, Kassel University Press Termkhajornkit, P. and T. Nawa (2004): The fluidity of fly ash-cement paste containing naphthalene sulfonate superplasticizer, Cement and Concrete Research, Vol.34, No.6, pp. 1017-1024 Toutou, Z. and N. Roussel (2006): Multi scale experimental study of concrete rheology. From water scale to gravel scale, Materials and Structures, Vol. 73, pp. 167-176 Turcry, Philippe and Ahmed Loukili (2003): A study of plastic shrinkage of Self-Compacting Concrete, in: [357], pp. 576-858 Uchikawa, H. and S. Uchida (1980): Influence of pozzolana on hydration of C 3A, in: Proceedings of the 7th International Conference on the Chemistry of Cement, Vol. III, pp. IV.24-29 Ulm, F.-J., Z.P. Bažant and F. H. Wittmann (Eds.) (2001): Creep, shrinkage and durability mechanics of concrete and other quasi-brittle materials, Proceedings of the Sixth International Conference CONCREEP-6@MIT, Cambridge (MA), Elsevier Various (2005): Measurement of properties of fresh Self-Compacting Concrete – Final report, European Union Growth Contract no G6RD-CT-2001-00580, 62 pages Verwey, E.J.W. and J.Th.G. Overbeek (1948): Theory of the stability of lyophobic colloids, Elsevier Publishing New York, ISBN 0-486-40929-5 Vieira, Manuel and Antonio Bettencourt (2003): Deformability of hardened SCC, in: [357], pp. 637-644 Vikan, Hedda and Knut O. Kjellsen (2008): Influence of cement composition on concrete rheology, Proceedings (CD) of The Third North American Conference on the Design and Use of Self-consolidating Concrete “SCC2008” Chicago, Ill. , paper No. 1086, session A5 Vogt, Carsten (2010): Ultrafine particles in concrete, Doctoral Thesis, Royal Institute of Technology, Stockholm, ISSN 110-3-4270

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343.

344.

345. 346. 347. 348. 349. 350. 351. 352.

353.

354.

355. 356. 357.

358. 359. 360. 361. 362.

363.

364.

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Wang, K., Z. Ge, J. Grove, J.M. Ruiz, and R. Rasmussen (2006): Developing a simple and rapid test for monitoring the heat evolution of concrete mixtures for both laboratories and field applications, Final Report Phase I FHWA DTF61-01-00042 Wang, K., Z. Ge, J. Grove, J.M. Ruiz, R. Rasmussen, and T. Ferragut (2007): Developing a simple and rapid test for monitoring the heat evolution of concrete mixtures for both laboratories and field applications, Final Report Phase II FHWA DTF61-01-00042 Wallevik, Jón Elvar (2003): Rheology of Particle Suspensions; Fresh Concrete, Mortar and Cement Paste with Various types of Lignosulphonate, Dissertation NTNU Trondheim, Norway Wallevik, Jon E. (2004): Thixotropy behaviour of cement pastes, Annual Transactions of The Nordic Rheology Society, Vol. 12, pp. 21-28 Wallevik, Jon Elvar (2006): Relationship between the Bingham parameters and slump, Cement and Concrete Research Vol. 36, pp. 1214-1221 Wallevik, Jón Elvar (2008): Minimizing end effects in the coaxial-cylinders viscometer: Viscoplastic flow inside the ConTec BML Viscometer 3, Journal of Non-Newtonian Fluids, Vol. 155, No.3, pp. 116-123 Wallevik, J.E. (2008): Methods of retrieving the Herschel-Bulkley parameter when using the (wide-gap) coaxialcylinders viscometer, in. [8], paper No. 1153 Wallevik, Jon Elvar (2009): Rheological properties of cement paste: Thixotropic behavior and structural breakdown, Cement and Concrete Research Vol. 39; pp. 14-29 Wallevik, Jon Elvar (2009): Development of parallel plate-based measuring system for the ConTec Viscometer, in: [364], pp. 18-24 Wallevik, Olafur H. (1983): Beskrivelse av fersk betongs egenskaper ved bruk av “ to punkts konsistensmåler” (Norwegian, Describing the fresh concrete properties using a two-point viscosmeter) MS thesis, NTNU Trondheim, Norway Wallevik, Olafur H. (1990): Den ferske betongens reology og anvendalse på betong med og uten tilsetning av Silicastøv, (Norwegian, Fresh concrete rheology and its application to concrete with and without silica fume) Dr.-Ing. Thesis NTH (NTNU) Trondheim, Norwegen Wallevik, O.H. and O.E. Gjørv (1990): Development of a Coaxial Cylinder Viscometer for Fresh Concrete, Properties of Fresh Concrete, Proceedings of the RILEM Colloquium, Chapman & Hall Hanover, pp. 213-224 Wallevik, O.H. and I. Nielsson (1998): Self-Compacting Concrete – A Rheological Approach, Proceedings of the International Workshop on Self-Compacting Concrete in Kochi, Japan, pp. 136-159 Wallevik, Olafur H. (2003): Rheology – a scientific approach to develop Self-Compacting Concrete, in [357], pp. 23-31 Wallevik, O. and I. Nielson (Eds.)(2003): Self-Compacting Concrete, Proceedings of the 3rd International RILEM Symposium on Self-Compacting Concrete, Reykjavik 2003, RILEM Proceedings Pro 33, ISBN: 2-9122143-42-X Wallevik, Olafur H., Índriði Nielson and Jón Elvar Wallevik (2004): TESTING SCC - Report on work package 5 – Rheology, Growth Contract No. GRD2-2000-30024 Wallevik, Olafur H. (2006): Rheology of Cementitious Material, Compendium of Rheology Course T845-Rheo at Reykjavik University Wallevik, O.H., S. Kubens, and F. Mueller (2007): Influence of cement-admixture interaction on stability of production of SCC, in: [87], pp. 211-216 Wallevik, Olafur H. and Björn Hjartarson (2008): A novel field instrument to measure rheological properties of Self-Compacting Concrete, in: [8], paper No. 1155 Wallevik, Olafur H. and Wolfgang Kunther (2008): Comparison of the Effects of Densified Silica Fume and Micro Silica on mortar rheology, project report Icelandic Building Research Institute IBRI, 35 pages, restricted Wallevik, Olafur H., Florian V. Mueller, Björn Hjartarson, and Stefan Kubens (2009): The green alternative of Self-Compacting Concrete – Eco-SCC, Proceedings XVII. Ibausil, Weimar, Germany, Vol. 1, pp.1105 -1116, ISBN 978-3-00-027265-3 Wallevik, Olafur H., Stefan Kubens and Sonja Oesterheld (Eds.) (2009): Rheology of Cement Suspensions such as Fresh Concrete, Proceedings of the Third International RILEM Symposium, RILEM Pro 68, ISBN: 978-2-35158-091-2

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC 365. 366. 367. 368. 369. 370. 371. 372.

373. 374.

375. 376. 377.

Wallevik, Olafur, H. and Jon Elvar Wallevik (2011): Rheology as a tool in concrete science: The use of rheographs and workability boxes, Cement and Concrete Research, Vol. 41, No. 12, pp. 1279-1288 Walraven, Joost C. (1998): The development of Self-Compacting Concrete in the Netherlands, Proceedings of the International Workshop on Self-Compacting Concrete in Kochi, Japan, pp. 87-96 Williams, David A., Aaron W. Saak, and Hamlin M. Jennings (1999): The influence of mixing on the rheology of fresh cement paste, Cement and Concrete Research, Vol. 29, pp. 1491-1496 Wittmann, F. (1968): Physical and chemical causes of creep and shrinkage of concrete, Rheologica Acta, Vol. 7, No. 4, pp. 397-400 Wittmann, F.H. (2001): Mechanisms and mechanics of shrinkage, in [337], pp. 3-12 White, H.E. and S.F. Walton (1937): Particle packing and particle shape, Journal of the American Ceramic Society, Vol. 20, Issue 1-12, pp. 155-166 Wright, P.J.F. (1953): Entrained air in concrete, Proceedings of Institution of Civil Engineers, Part I, 2, No. 3, London, pp. 337-358 Wuestholz, Timo (2005): Experimentelle und theoretische Untersuchungen der Frischbetoneigenschaften von selbstverdichtendem Beton, (German, Experimental and theoretical evaluations of fresh concrete properties in SCC), Dissertation, University Stuttgart (in German) Yahia, A. and K.H. Khayat (2001): Analytical models for estimating yield stress of high-performance pseudoplastic grout, Cement and Concrete Research, Vol. 31, pp. 731-738 Yahia, A. and K.H. Khayat (2003): Applicability of rheological models to high-performance grouts containing supplementary cementitious materials and viscosity enhancing admixtures, Materials and Structures, Vol. 36, pp. 402-412 Yajun, Ji and Jong Herman Cahyadi (2003): Effect of densified silica fume on microstructure and compressive strength of blended cement pastes, Cement and Concrete Research, Vol. 33, No.10, pp. 1543-1548 Yu, A.B. and N. Standish (1991): Estimation of the porosity of particle mixtures by a linear-mixture packing model, Ind.Eng.Chem, Vol. 30, No. 6, pp. 1272-1385, Yu, A.B. and N. Standish (1993): Limitation of proposed mathematical models for the porosity estimation of nonspherical particle mixtures, Ind.Eng.Chem.Res., Vol. 32, pp. 2179-2182

Norms and standards AASHTO T312:2009

Standard method of test for preparing and determining the density of hot mix asphalt (HMA) specimens by means of the Superpave Gyratory Compactor, American Association of State Highway and Transport Officials, January 2009

ASTM C29/C29M-09

Standard test method for bulk density (“unit weight” and voids in aggregate

ASTM C143/C143M-10a

Standard test method for slump of hydraulic cement concrete

ASTM C157

Standard test method for length change of hardened hydraulic cement mortar and concrete

ASTM C457 / C457M-10a

Standard test method for microscopically determination of air void system in hardened concrete

ASTM C597-09

Standard test method for pulse velocity through concrete

ASTM C666 / C666-M -03

Standard test method for resistance of concrete to rapid freezing and thawing

ASTM C1610 /C1610-06a:

Standard test method for static segregation of self-consolidating concrete using column technique

ASTM C1611 /C1611M-09b:

Standard test method for slump flow of self-consolidating concrete

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ASTM C16121 /C1621M-09b:

Standard test method for passing ability of self-consolidating concrete by J-ring

ASTM D698/AASHTO T990:

Standard test method for laboratory characteristics of soil using standard effort (12400ft-lbf/ft3)(600 kN m/m3)

ASTM D1557/AASHTO T180:

Standard test method for laboratory characteristics of soil using modified effort (56000 ft-lbf/ft3)(2700 kN m/m3)

ASTM C1712-09

Standard Test Method for Rapid Assessment of Static Segregation Resistance of Self-Consolidating Concrete Using Penetration Test

DAfStb (2003): SVB Richtlinie

Additional recommendation to DIN 1045 of The German Committee for Reinforced Concrete Structures, 44 pages (in German)

DIN 53015:2001

Viscometry – Measurement of viscosity by means of rolling ball viscometer by Höppler

(DIN) EN 196-1: 2005

Testing cement- Part 1: Determination of compressive strength

EN 197-1: 2000

Cement –Part 1, Composition, specification and conformity criteria for common cements

EN 206-1:2001

Concrete-Part 1; Specifications, performance, production and conformity

prEN 206-09:2007

Concrete – Part 9: Additional rules for Self-Compacting Concrete (SCC)

EN 480-11: 2008

Admixtures for concrete, mortar and grouts – test methods – Part 11: Determination of air void characteristics in hardened concrete

(DIN) EN 450-1:2008-05 (E)

Fly ash for concrete – Part 1: Definitions, specifications and conformity criteria (includes Amendment A1:2007)

EN 934-2:2009

Admixtures for concrete, mortar and grouts – Part 2: Concrete admixtures

EN 1097-4:2008

Tests for mechanical and physical properties of aggregates. Determination of the voids of dry compacted filler.

EN 1097-6:2008

Tests for mechanical and physical properties of aggregates. Part 6: Determination of density and water of absorption

(DIN) CEN/TS 12390-9:2006

Pruefung von Festbeton –Teil: Frost-und Frost.Tausalz-WiderstandAbwitterung (Testing of hardened concrete – part: Freeze-thaw resistance)

EN 12350-2:2009

Testing fresh concrete - Part 2: Slump test

EN 12350-3:2009

Testing fresh concrete - Part 3: VEBE test

EN 12350-6:2009

Testing fresh concrete - Part 6: Density

EN 12350-8: 2010

Testing fresh concrete - Part 8: Self-compacting concrete– slump flow test

EN 12350-11: 2010

Testing fresh concrete – Part 11: Self-Compacting Concrete- Sieve segregation test

EN 12350-12: 2010

Testing fresh concrete – Part 12: Self-compacting concrete – J-ring test

EN 12390-1:2009

Testing hardened concrete: Shape, dimensions and other requirements for specimens and moulds

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC EN 12397-31: 2004

Bituminous mixtures. Test methods for the hot mix asphalt specimen preparation gyratory compactor

(DIN) EN 13263-1:2007-07

Silicastaub für Beton – Teil 1: Definitionen, Anforderungen und Konformitätskriterien (German, Silica fume for concrete, Part 1: Definitions, requirements, and conformity criteria), German Version EN 123631:2005+A1:2009

NORDTEST NT BUILD 427:1994-11: Concrete, fresh: Compactibility with IC-Tester SS 137244:

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Betonprovning, Hårdnad betong, Avlagning vid frysning, (Swedish, Concrete testing, Hardened Concrete, Freeze-Thaw-Resistance) Faststaelld 95-03-08

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Appendix Chapter 2 Materials and methods 100

PSD [%]

80 60 Del 01 Del 02

40

Del 03 Del 04

20 0

Sieve size [mm]

Figure-Appendix 1: PSD of various deliveries of sand 0-8 over a period of 2 years Table-Appendix 1: Sieve curves of aggregates

Sieve size

FS1 [%]

[%]

S1

G1

G2

G3

G4

G5

32 22.4 16 11.2 8 4 2 1 0.5 0.25 0.125 0.063

100 100 100 100 99.7 95.1 86.1 73.8 59.5 40.7 23.2 11.8

100 100 100 100 98.9 77.2 61.1 44.3 27.1 13.8 6.1 2.8

100 100 95.5 54.2 10.5 0.4 0.3 0.3 0.2 0.2 0.2 0.1

100 100 100 91.8 19.2 2.3 2.1 2.0 1.9 1.9 1.8 1.7

100 100 94.8 56.6 23.4 7.3 5.7 4.8 4.1 3.4 2.5 1.8

100 100 98.7 69.0 35.7 7.6 4.0 2.5 1.7 1.1 0.8 0.6

100 100 100 71.7 22.4 2.8 2.1 2.0 1.9 1.7 0.5 0.2

[mm]

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[%]

[%]

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[%]

[%]

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Table-Appendix 2: Images of aggregates

Abbreviation

Image

B

(a) Ground basalt filler (B) from Semenstverksmidjan Ehf., Akranes, Iceland

FS

Particle appearance: see S1

S1

(b) Granite /Gneiss sand 0-8 mm (S1) from Årdal, Norway ICI Rheocenter

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G1

(c) Granite/Gneiss coarse gravel (G1) from Årdal, Norway

G2

(d) Quartz-diorite coarse gravel 8-11.2 mm (G2) from Tau, Norway

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G3

(e) Icelandic Basalt “Hardikampur Perla” , 8-16 mm (G3) from Iceland

G4

(f) Icelandic (porous) basalt "Björgun Perla"

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G5

(g) Icelandic basalt "Hardikampur Perla"

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Fines

100% 80%

AIR

60%

W

40%

SF

20%

CEM

0% 178 181 186 193 200 Water content

Matrix volume [l/m3]

Matrix proportion vol-%

400 350

Fines

300

AIR

250

W

200

SF

150

CEM

100 50 0 170 180 190 200 210 Water content [kg/m3]

[kg/m3]

Figure-App. 1: Matrix proportions (vol.-%) of Eco-SCC robustness test, water series

Figure-App. 2: Evolution of matrix volume for Eco-SCC robustness tests, water series

450 Fines

350 Fines Air water L FA CEM

100% 80% 60% 40% 20% 0%


Matrix proportion vol.-%

400

Air

300

water

250

L

200

FA

150

CEM

100 50 0

177 183 189 195 201

170 180 190 200 210

Water content [kg/m3]


Figure-App. 3: Matrix proportions (vol.-%) of lean SCC robustness test, water series

Figure-App. 4: Evolution of matrix volume for lean SCC robustness tests, water series

450 400 Fines

Fines

100% 80%

Air

60%

water

40%

FA

20%

CEM

0%



350 250

water

200

FA

150

CEM

100

175 177 186 191 196

50 0 170 180 190 200 210



Figure-App. 5: Matrix proportions (vol.-%) of br-SCC robustness test, water series ICI Rheocenter

Air

300

Figure-App. 6: Evolution of matrix volume for br-SCC robustness tests, water series

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Operation Procedure for Cylinder Sedimentation Test A. Casting of specimen Equipment: Suitable container for batching, e.g. 5l plastic bucket Cylinder mould, with a 150 mm diameter and 300 mm height Mould cover70 Procedure: Remark: The moulds should be placed in a position that makes it possible to let the specimen rest undisturbed. Since vibrations could occur that would affect the settlement the sample should not be moved after casting. 1. A representative sample (considering homogeneity and fluidity state) of fresh concrete should be taken with a suitable container of a size that requires the fewest possible filling steps. 2. Place the container directly above the cylinder, in order to keep the filling height of the concrete determined. Preferably, keep the mould positioned vertically during the filling and tilt only for very viscous mixes to obtain a complete de-aired sample. 3. Fill the mold while adapting the filling rate to the rheology demands of the material so as to obtain a representative sample. A high-viscous mix therefore needs a low filling rate in order to allow de-airing of the sample, whereas a low-viscous mix with less fluidity requires a higher filling rate in order to overcome the yield stress value. 4. To mitigate evaporation during the hydration process, cover the sample. Be cautious not to apply any energy to the mold or specimen.

B. Specimen preparation Equipment: Water-cooled concrete saw Ruler Optional: moveable table, water resistant marker, a photo camera mounted on a tripod Time demand: About 20-30 minutes to quarter and slice one cylinder (D150/H300)

It is important to avoid any impact to the material when placing the cover, so a plastic bag put over the mould or a lid loosely placed on it will suffice. 70

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Procedure: 1. Record, e.g. by taking photos, the surface appearance of the specimen, also including the top view. 2. Cut the specimen into two halves along its central axis. 3. Make records, e.g. by taking photos, of the aggregate distribution. 4. Optional: cut the two halves further, along their central axes. 5. Mark the halves (quarters), in order to identify their location relative to each other, while considering the filling direction. In my experience, drawing an off-center line (e.g. always on the right side of the residual outer cylinder) allows further specimen numbering later in the cutting process. 6. Cut the halves (quarters) into slides of about 40 mm thickness, starting with the bottom portion. Try to avoid making top slides with a thickness less than half (i.e. 20 mm) of the other specimen. If necessary, use the last cut to halve the remaining specimen. Nevertheless, the top area of the whole specimen, i.e. for all two (four) halves (quarters), should be obtained in a similar thickness, if possible. 7. Mark each single specimen with a water resistant marker. It has proven to be a worthwhile effort to number the specimen slides according to which half (or quarter) it is, as well as a number to indicate the specimen’s position. 8. Clean the remains that adhered during the cutting process from all specimens.

C. Specific gravity evaluation Equipment: Balance, allowing underwater measurement, with accuracy ± 0.1 g Towel Ruler, allowing an accuracy of ± 0.1 mm Suitable equipment to record the readings, e.g. laptop, pen & paper. Time demand: About 30-40 minutes for one cylinder (D150/H300) that was quartered and cut into 7 slices per quarter Procedure: Remark: In order to connect the specific gravity to its particular height (i.e.“horizon”), measure the thickness of each slice with the ruler. Doing so for one half (or quarter) might suffice as long as their thicknesses are evenly distributed. In this case, one half (quarter) can be considered as representative. 1. Measure the weight of each single specimen separately while under water. Record the mass under water (muw) once the reading has been stable for 30 seconds to ± 0.1 g. 2. After one specimen is measured under water, dry it gently with a towel until the shiny surface, caused by the water film, has disappeared and no more water drains out from eventually occurring voids. This state is defined as saturated and surface dry condition (SSD), in accordance with the treatment of coarse aggregates in EN 1097-6. ICI Rheocenter

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3. Immediately measure the particular specimen in SSD and record the mass (mSSD) to ± 0.1 g. D. Data evaluation Equipment: Suitable calculation tools (reasonable effort is obtained when using a template in a spreadsheet program, such as MS Excel.) Procedure: Preliminary – limit evaluation: 1. Determine the volume proportion of each single concrete composite. The polymer content of admixture might be negligible. If fibers are present, they should be considered as well and can easily be added if needed. Equation 11-1

𝑉𝑡𝑜𝑡 = 𝑉𝑎𝑖𝑟 + With:

𝑚𝐶𝑀 𝑚𝑤 𝑚𝑎𝑔𝑔 + + 𝜌𝐶𝑀 𝜌𝑤 𝜌𝑎𝑔𝑔

mcm (averaged71) mass of cementitious materials, also incorporating all filler; ρcm (averaged) density of cementitious materials; mw mass of water; ρw density of water; (averaged) mass of aggregates; (averaged) density of aggregates (if the densities allow it).

magg ρagg

2. Employ the particle size distribution (PSD) and density of the aggregates to determine the volume proportion of each single size fraction. Equation 11-2

𝑉𝑎𝑔𝑔𝑟 =

𝑚𝑎𝑔𝑔𝑟 𝑚0.063 𝑚0.125 𝑚0.25 𝑚0.5 𝑚1 𝑚2 𝑚4 𝑚8 𝑚16 = + + + + + + + + +⋯ 𝜌𝑎𝑔𝑔𝑟 𝜌 0.063 𝜌0.125 𝜌0.25 𝜌0.5 𝜌1 𝜌2 𝜌4 𝜌8 𝜌16

3. Determine the averaged density of the composite up to a specific aggregate size fraction i, e.g. paste, matrix, micro-mortar (i < 2 mm), and so on. Equation 11-3

𝜌𝑖 =

𝑚𝐶𝑀 + 𝑚𝑤 + ∑𝑖0.063 𝑚𝑎𝑔𝑔𝑟,𝑖 𝑉𝑎𝑖𝑟 + 𝑉𝐶𝑀 + 𝑉𝑤 + ∑𝑖0.063 𝑉𝑎𝑔𝑔𝑟,𝑖

4. Determine the density of the modified composite according to the sedimentation limits defined. For instance in this thesis I chose to apply a variation of ± 10 and 20 % by mass of all aggregates larger than 8 mm (4 mm, as a more rigid criterion). When replacing these

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Hereby, the volume proportions used in 1 m3 of fresh concrete are considered.

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC volumes with material that has the density of the matrix, it provides the most severe case and would allow a judgment of the sedimentation risk “on the safe side”. 5. The density of the modified composite is the basis for comparison with the specific gravity distribution one determines in the sedimentation specimen.

Calculation of specific gravity: 1. Calculate the specific gravity of each single specimen according to Equation 11-4. Equation 11-4

𝑆𝐺 =

𝑚𝑆𝑆𝐷 𝑚𝑆𝑆𝐷 − 𝑚𝑈𝑊

2. Calculate the location of the different horizons hi, also considering the loss of material due to the blade’s thickness, commonly about 3 mm. 3. Calculate the algebraic average value for the specific gravity at each horizon i considering all halves (quarters). 4. Calculate the median value of the averaged specific gravities. This is done in order to reduce the effect of single extreme values, as it would occur when applying the algebraic average value. 5. Calculate the standard deviation of all averaged specific gravities to the median value. 6. Identify the extreme values out of the averaged specific gravities per horizon, and determine their difference from the median value (i.e. “range”) and ratio of these ranges to the median value. Determination of sedimentation parameters: 1. Apply the appearance of the largest aggregate particle at the top to the ranking of the Aggregate Surface Appearance Index (ASAI) (Table-Appendix 3 & Table-Appendix 5) 2. Rank the distribution of the specific gravities along the specimen height (i.e. SGD) into classes (CSGD) (Table-Appendix 4), also considering the extreme values determined in respect to the defined limit. Table-Appendix 3: Aggregate Surface Appearance Index (ASAI) for SCC

Rank

Rating

0

Insufficently self-compacting

1


2

Small risk of sedimentation

3

Moderate risk of sedimentation

4


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Criteria Numerous coarse aggregates visible at the surface that are not embedded in mortar, or volume incompletely filled. Small numbers of coarse aggregates are visible at the surface, but embedded into mortar to a great extent. A few numbers of coarse aggregates are visible near the surface and covered by a mortar layer. Only morar and small aggregates covered by mortar are visible at the surface. In the cut area larger particles remain within the distance of dmax. Only mortar and paste are visible at the surface and in the cut area for several millimeters (>> dmax).

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Table-Appendix 4: Classes of specific gravity distribtions (CSGD) of SCC

Rank

Rating

Criteria

0

Insuffiently self-compacting

1


2

Small risk of static sedimentation

3

Small risk of static sedimentation affected by risk of dynamic segregation

4

Moderate risk of static sedimentation

5 6

Moderate risk of static sedimentation affected by risk of dynamic segregation High risk of sedimentation

Increasing specific gravity within height due to insufficient consolidation of concrete No slope of SGD, while SG and theoretical density of the de-aired /fully consolidated fresh concrete coincide well Straight slope of SGD. Variation of SG is below “10% criterion”72, i.e. about 1% Waved curvature of SGD. Variation of SG is below “10% criterion” Straight curvature of SGD. Variation of SG is below “20% criterion”, i.e. bout 2-2.5% 73 Waved curvature of SGD. Variation of SG is below “20% criterion”. Variation of SG exceeds “20% criterion”.

The criteria 10% and 20% refers to segregation of 10 and 20% of particles bigger than 8 mm. The exact limit depends on the mix design, in particular matrix volume and air proportion and needs to be evaluated for each mix design. 72 73

276

Florian V. Mueller

ICI Rheocenter


Table-Appendix 5: Examples for Aggregate Surface Appearance Index (ASAI) in relation to specific gravity distribution (SGD)

Surface

Cut area

ASAI 0

Spec.gravity*1000[-]

(Insufficiently self-compacting

ASAI 1


(Stable & selfcompacting)

ASAI 2


(Small risk of sedimentation)

ASAI 3


(Moderate risk of sedimentation)

ASAI 4

(High risk of sedimentation)

ICI Rheocenter

SGD (example) Spec. gravity *1000[-]

Rating

Florian V. Mueller

2500 2450 2400 2350 2300 0

10 20 30 Column height [cm]

0


0


0


0


2500 2450 2400 2350 2300

2500 2450 2400 2350 2300

2500 2450 2400 2350 2300

2500 2450 2400 2350 2300

277

Appendix Chapter 3 Paste composition Table-Appendix 6: Viscometer 5 raw data for mixes delivering Bingham parameters presented in chapter 3.1.

Resistance [V]

1.5

1 2s3sx9 1s2sx8 +4min 2s3sx9 +8min 0.5

1s2sx8 +12min

0 0

50

100

150

200

250 300 Data point

350

400

450

500

(a) Raw data for Table 3-9 on page 125

Resistance [V]

1.5

1 2s3sx9 1s2sx8 +4min 2s3sx9 +8min

0.5

0 0

50

100

150

200

250 300 Data point

350

400

450

500

(b) Raw data for Table 3-10 on page 125

278

Florian V. Mueller

ICI Rheocenter


Appendix Chapter 4 Effect of matrix volume Table-Appendix 7: Isothermal calorimeter data for effects obtained by Solid Matrix Approach

Power mW/g

6

21.0°C

4

2

B000

20.9°C

B050

22.0°C

B100

21.7°C

B150

0 0

4

8

12 16 Hydration time [h]

20

24

(A) Effect of ground basalt filler addition, Icelandic origin

Power mW/g

6

4

2

21.0°C

Ref

21.4°C

L050

23.0°C

L100

22.5°C

L150

0 0

4

8


20

24

(B) Effect of limestone filler addition, Swedish Limestone

Power mW/g

6

4

21.0°C

Ref

22.1°C

FA050

22.3°C

FA100

23.4°C

FA150

2

0 0

4

8 12 16 Hydration time [min]

20

24

(C): Effect of fly ash addition, Danish origin

ICI Rheocenter

Florian V. Mueller

279

Table-Appendix 8: Isothermal calorimeter data for effects obtained by Air Matrix Approach

Power mW/g

6

22.6°C

AEA0.0_180

21.6°C

AEA0.4_180

20.7°C

AEA0.7_180

4

21.8°C

2

AEA1.0_180

0 0

4


20

24

(A): Effect of air entrainment on mixes without additional fines, w/b about 0.57

6

23.3°C

AEA_B150_0.3

Power mW/g

21.6°C 4

AEA_B150_0.4

20.7°C

AEA_B150_0.7

22.8°C

AEA_B150_0.9

21.6°C

2

AEA_B150_1.0 23.2°C

AEA_B150_1.2

0 0

4

8


20

24

(B) Effect of air entrainment with additional ground basalt filler, w/c about 0.59

280

Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table-Appendix 9: Drying shrinkage data of chapter 4.2.

1.4 Shrinkage [mm/m]

1.2 1.0

Ref-Ís1

0.8

G1_B0

0.6

G1_B50

0.4

G1_B100

0.2

G1_B150

0.0 0

70

140

210

280 350 420 Hydration time [d]

490

560

630

700

(a) Shrinkage strain of series with GBF as filler material, inclusive Icelandic reference

Weight loss [%]

5 4

Ref-Ís1

3

G1_B0

2

G1_B50 G1_B100

1

G1_B150

0 0

70

140

210


490

560

630

700

(b) Weight loss (evaporation) of GBF-series in 23°C @ 50 %-RH

Shrinkage [mm/m]

1.4 1.2 1.0

Ref-Ís1

0.8

G3_B150

0.6

G3_B100

0.4

G3_B50

0.2

G3_B0

0.0 0

70

140

210


490

560

630

700

(c) Shrinkage strain of coarse aggregate G3 series, including Icelandic reference

Weight loss [%]

5 4 Ref-Ís1

3

G3_B0 G3_B50

2

G3_B100

1

G3_B150

0 0

70

140

210


490

560

630

700

(d) Weight loss (evaporation) of shrinkage specimen G3 series in 23°C @ 50 %-RH ICI Rheocenter

Florian V. Mueller

281

Shrinkage [mm/m]

1.4 1.2 1.0

Ref-Ís1

0.8

G4_B150

0.6

G4_B100

0.4

G4_B50

0.2

G4_B0

0.0 0

70

140

210


490

560

630

700

(e) Shrinkage strain of coarse aggregate G4 series, inclusive Icelandic reference

Weight loss [%]

5 4

Ref-Ís1

3

G4_B0

2

G4_B50 G4_B100

1

G4_B150

0 0

70

140

210


490

560

630

700

(f) Weight loss (evaporation) of G4 series in 23°C @ 50 %-RH

282

Florian V. Mueller

ICI Rheocenter


Appendix Chapter 5 Gyratory particle packing Table-Appendix 10: PSD of untested and tested material in chapter 5.1: “200_G1_B0”, n=1536

Sieve: Tested I Untested I PSD Theory

22.4

16

8

4

2

1

0.5

0.25

100 100 100

99.3 95.5 98.9

77.5 80.2 77.7

60.3 63.8 58.8

48.0 52.1 46.5

35.1 38.7 33.7

21.8 24.6 20.7

11.2 13.1 10.5

0.125

0.063

mcoll

min

74

4.9 6.1 4.7

2.2 2.8 2.2

1642.0

1642.8

708.3

709.2

Table-Appendix 11: PSD of untested and tested material in chapter 5.1: “420_G1_B0”= “76/24” = “G1_B0”, test III after n=1536

Sieve: Tested I Tested II Untested I Untested II PSD Theory

22.4 100 100 100 100 100

16 97.5 98.2 96.9 97.6 98.9

8 77.1 77.7 75.4 76.7 77.7

4 62.3 60.1 58.6 59.8 58.8

2 50.4 47.5 46.9 47.9 46.5

1 37.3 34.2 33.9 34.9 33.7

0.5 23.4 20.7 20.7 21.5 20.7

0.25 12.2 10.4 10.5 11.0 10.5

0.125 5.6 4.7 4.7 4.9 4.7

0.063 2.5 2.1 2.1 2.2 2.2

mcoll

min

1610.4

1610.8

1639.7

1641.1

1117.2

1118.2

1059.3

1060.3

Table-Appendix 12: PSD of untested and tested material in chapter 5.1: “600_G1_B0”, n=1536


22.4 1000 100 100

16 97.5 95.5 98.9

8 76.9 80.2 77.7

4 59.1 63.8 58.8

2 46.8 52.1 46.5

1 34.5 38.7 33.7

0.5 21.9 24.6 20.7

0.25 11.8 13.1 10.5

0.125 5.6 6.1 4.7

0.063 2.6 2.8 2.2

mcoll

min

1633.1

1633.5

708.3

709.2

Table-Appendix 13: PSD of untested and tested material in chapter 5.1: “800_G1_B0”


22.4 100 100 100

16 97.9 95.5 98.9

8 76.5 80.2 77.7

4 57.8 63.8 58.8

2 45.1 52.1 46.5

1 32.5 38.7 33.7

0.5 20.0 24.6 20.8

0.25 10.4 13.1 10.5

0.125 4.8 6.1 4.7

0.063 2.2 2.8 2.2

mcoll

min

1616.7

1617.6

709.3

709.2



22.4 100 100 100

16 97.8 95.5 98.9

8 78.4 80.2 77.7

4 64.1 63.8 58.8

2 52.7 52.1 46.5

1 39.7 38.7 33.7

0.5 25.4 24.6 20.7

0.25 13.7 13.1 10.5

0.125 6.5 6.1 4.7

0.063 3.0 2.8 2.2

mcoll

min

1628.2

1629.3

708.3

709.2



22.4 100 100 100

16 97.8 97.7 99.0

8 81.1 73.1 79.4

4 63.6 57.0 62.1

2 51.2 48.2 50.8

1 38.7 37.8 39.1

0.5 26.6 26.8 27.2

0.25 17.7 18.2 17.9

0.125 12.5 13.0 12.4

0.063 9.2 9.6 9.0

mcoll

min

1651.1

1652.9

740.5

740.4

The term mcoll reflects mass of collected material [g], whereas min stands for mass of material input in sieving process. 74

ICI Rheocenter

Florian V. Mueller

283

Table-Appendix 16: PSD of untested and tested material in chapter 5.1: “420_G1_B150”=”G1_B150”


22.4 100 100 100 100 100

16 98.9 97.8 98.5 97.6 99.0

8 83.2 75.5 78.8 78.6 79.4

4 67.6 59.3 63.7 64.2 62.1

2 56.7 49.0 53.5 53.9 50.8

1 44.2 38.1 41.5 41.9 39.1

0.5 30.5 26.5 28.9 29.1 27.2

0.25 19.6 17.2 18.9 19.0 17.9

0.125 13.1 11.7 12.8 12.9 12.4

0.063 9.2 8.3 9.1 9.1 9.0

mcoll

min

1619.7

1621.0

1161.5

1161.8

1159.1

1160.0

1121.6

1123.9



22.4 100 100 100

16 98.6 97.7 99.0

8 76.8 73.1 79.4

4 59.2 57.0 62.1

2 47.4 48.2 50.8

1 35.8 37.8 39.1

0.5 24.7 26.8 27.2

0.25 16.3 18.2 17.9

0.125 11.4 13.0 12.4

0.063 8.2 9.6 9.0

mcoll

min

1598.6

1600.6

740.5

740.4



22.4 100 100 100

16 98.2 97.7 99.0

8 79.3 73.1 79.4

4 61.4 57.0 62.1

2 49.7 48.2 50.8

1 38.0 37.8 39.1

0.5 26.4 26.8 27.2

0.25 17.4 18.2 17.9

0.125 12.1 13.0 12.4

0.063 8.7 9.6 9.0

mcoll

min

1645.2

1648.0

740.5

740.4



22.4 100 100 100

16 97.6 97.7 99.0

8 78.5 73.1 79.4

4 62.5 57.0 62.1

2 50.8 48.2 50.8

1 39.1 37.8 39.1

0.5 27.3 26.8 27.2

0.25 18.1 18.2 17.9

0.125 12.5 13.0 12.4

0.063 8.9 9.6 9.0

φi

eφ

1656.6

1656.3

740.5

740.4

Table-Appendix 20: PSD of untested and tested material in chapter 5.2: “84/16”

Sieve: Tested I Tested II Tested III Untested I Untested II PSD Theory

22.4 100 100 100 100 100 100

16 99.6 99.2 97.3 98.7 99.3 99.3

8 84.9 83.1 84.9 83.3 85.5 84.8

4 64.5 61.9 66.5 66.1 66.8 64.9

2 51.0 48.9 54.1 53.8 54.3 51.4

1 37.6 36.2 40.6 39.9 40.5 37.3

0.5 23.7 22.9 25.7 24.8 25.5 22.8

0.25 12.3 11.9 13.3 12.7 13.3 11.6

0.125 5.5 5.4 6.0 5.7 6.0 5.2

0.063 2.4 2.4 2.7 2.5 2.6 2.4

mcoll

min

1677.6

1677.5

1689.7

1691.3

1680.7

1681.7

1098.1

1099.2

1088.3

1090.4

Table-Appendix 21: PSD of untested and tested material in chapter 5.2: “68/32”, test III after n=1536


284

22.4 100 100 100 100 100 100

16 95.1 97.3 99.2 97.4 99.4 98.9

8 68.9 65.2 75.9 70.8 71.2 77.7

4 50.3 46.4 58.9 51.2 52.4 58.8

2 39.6 35.7 46.8 39.4 41.2 46.5

1 29.0 25.6 34.2 28.2 29.8 33.7

0.5 18.1 15.9 21.2 17.5 18.2 20.7

Florian V. Mueller

0.25 9.6 8.3 11.1 9.5 9.3 10.5

0.125 4.5 3.9 5.1 4.8 4.2 4.7

0.063 2.1 1.8 2.3 2.6 1.9 2.2

mcoll

min

1676.6

1675.7

1688.2

1689.7

1645.0

1646.6

1139.9

1129.8

1092.7

1093.4

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table-Appendix 22: PSD of untested and tested material in chapter 5.2: “G1_B50”


22.4 100 100 100 100 100

16 97.7 98.3 96.4 99.3 98.9

8 81.6 77.0 74.0 74.7 78.2

4 66.2 60.8 57.0 57.9 59.8

2 54.0 48.6 46.1 45.9 47.9

1 40.5 35.9 34.2 33.8 35.5

0.5 26.2 23.0 22.0 21.6 22.8

0.25 14.7 13.0 12.5 12.3 13.0

0.125 8.0 7.3 7.1 6.9 7.2

0.063 4.6 4.4 4.3 4.1 4.4

mcoll

min

1598.4

1599.3

1618.3

1619.1

1114.6

1114.6

1133.4

1133.9

Table-Appendix 23: PSD of untested and tested material in chapter 5.2: “G1_B100”


22.4 100 100 100 100 100

16 98.6 97.2 98.6 98.5 99.0

8 80.2 76.6 80.7 77.5 78.9

4 64.2 60.5 66.1 62.9 61.0

2 53.1 49.7 54.7 52.3 49.5

1 40.6 38.1 42.0 40.3 37.4

0.5 27.4 25.9 28.4 27.5 25.0

0.25 16.8 16.0 17.6 17.1 15.5

0.125 10.3 9.9 11.0 10.7 9.8

0.063 6.7 6.5 7.2 7.1 6.8

mcoll

min

1657.4

1660.0

1599.4

1600.6

1040.4

1041.1

1178.4

1178.2

Table-Appendix 24: PSD of untested and tested material in chapter 5.2: “G1_B50_50”, test III after n=1536

Sieve: Tested I Tested II Tested III Untested I PSD Theory

22.4 100 100 100 n.a. 100

16 97.8 97.0 98.0

8 77.7 78.6 77.8

4 62.1 62.6 62.0

2 50.5 51.3 50.8

1 38.1 38.8 38.4

0.5 25.4 26.0 25.6

0.25 15.5 15.8 15.6

0.125 9.6 9.9 9.8

0.063 4.3 4.5 4.3

99.0

78.6

60.7

49.2

37.2

24.9

15.4

9.9

4.4

mcoll

min

1654.8

1655.7

1595.5

1597.3

1627.9

1628.7

Table-Appendix 25: PSD of untested and tested material in chapter 5.2: “G1_B50_100”


22.4 100 100 100 100 100

16 99.0 99.6 96.7 98.1 99.0

8 79.9 80.0 74.7 79.5 79.6

4 63.3 63.4 60.8 63.7 62.2

2 51.4 51.4 49.9 52.4 50.9

1 39.3 39.4 38.1 40.6 39.1

0.5 27.2 27.2 26.2 28.3 27.1

0.25 17.8 17.9 17.3 18.7 17.8

0.125 12.2 12.2 12.0 12.7 12.3

0.063 4.5 4.6 4.4 4.4 4.2

mcoll

min

1634.3

1634.1

1650.9

1651.2

1065.5

1066.6

1101.2

1101.7

Table-Appendix 26: PSD of untested and tested material in chapter 5.2: “G1_B50_150”


22.4 100 100 100 100 100

16 99.6 98.6 98.3 99.2 99.0

8 84.2 79.0 77.9 82.1 80.1

4 64.9 60.8 61.5 65.5 63.2

2 54.4 50.4 51.0 54.7 52.3

1 43.1 39.5 40.0 43.1 40.8

0.5 30.7 28.1 28.6 30.9 29.1

0.25 20.8 19.0 19.8 21.2 20.1

0.125 14.7 13.7 14.5 15.5 14.8

0.063 4.4 4.4 4.3 4.5 4.1

mcoll

min

1598.6

1600.6

1609.4

1611.3

1104.5

1108.8

1130.6

1132.1

Table-Appendix 27: PSD of untested and tested material in chapter 5.2: “G1_B50_50_100”


ICI Rheocenter

22.4 100 100 100 100 100

16 97.4 99.2 99.2 98.8 99

8 78.8 79.3 78.4 80.0 80.0

4 63.6 62.4 61.8 63.1 63.1

2 52.7 51.4 51.3 53.0 52.1

1 41.3 39.9 40.3 41.5 40.7

0.5 29.3 28.3 28.8 29.7 29.0

0.25 19.8 19.3 19.8 20.4 19.9

Florian V. Mueller

0.125 9.0 8.9 9.3 9.5 9.1

0.063 3.9 4.0 4.3 4.3 4.1

mcoll

min

1629.7

1631.1

1638.2

1638.0

1053.4

1053.6

1130.6

1131.4

285



22.4 100 100 100 100 100

16 98.3 99.6 98.9 95.6 99.1

8 79.6 82.0 78.7 77.3 80.8

4 63.9 67.7 62.5 60.6 64.5

2 53.6 56.4 51.8 49.3 53.9

1 42.9 44.8 41.0 38.6 42.9

0.5 31.8 32.8 30.0 28.0 31.6

0.25 23.1 23.6 21.7 19.9 22.8

0.125 9.4 9.9 9.3 8.5 9.5

0.063 4.1 4.4 4.3 3.9 4.2

mcoll

min

1691.2

1692.3

1686.8

1687.6

1032.6

1035.5

1037.0

1036.9

Table-Appendix 29: PSD of untested and tested material in chapter 5.3: “G2_B000”, test III after n=1536


22.4 100 100 100 100 100 100

16 100 100 100 100 100 100

8 77.2 81.9 81.2 82.5 82.7 79.1

4 51.0 60.8 56.8 61.5 59.3 58.6

2 39.8 49.4 43.9 49.3 47.0 46.5

1 28.7 37.0 31.9 36.2 34.3 33.8

0.5 17.5 23.4 19.5 22.5 21.1 20.8

0.25 8.9 12.2 10.0 11.4 10.9 10.8

0.125 4.1 5.4 4.6 5.1 4.8 5.0

0.063 1.9 2.4 2.1 2.2 2.1 2.5

mcoll

min

1631.6

1630.2

1675.8

1677.3

1671.0

1671.3

1163.3

1165.8

1177.9

1178.1

Table-Appendix 30: PSD of untested and tested material in chapter 5.3: “G5_B000”, test III after n=1536


22.4 100 100 100 100

16 100 100 100 100

8 75.0 81.1 82.5 78.0

4 54.9 59.3 62.3 55.8

2 44.0 47.2 50.2 44.3

1 32.9 34.8 37.6 32.2

0.5 20.9 21.9 23.9 19.8

0.25 11.0 11.4 12.6 10.1

0.125 5.0 5.1 5.6 4.5

0.063 2.2 2.3 2.6 1.9

mcoll

min

1662.9

1663.8

1669.2

1667.6

1649.8

1649.5

1129.7

1130.6



22.4 100 100 100 100 100 100

16 98.6 99.1 98.6 98.1 99.5 99.1

8 78.7 81.3 75.9 83.3 82.6 81.4

4 65.2 66.1 59.9 68.1 66.8 65.6

2 55.6 55.5 50.3 56.8 55.6 55.3

1 45.3 44.9 40.4 45.0 44.1 44.7

0.5 34.5 33.9 30.6 33.5 32.8 33.7

0.25 25.9 25.4 22.9 24.8 24.4 25.3

0.125 17.9 17.3 15.8 17.0 16.8 17.4

0.063 9.3 8.7 8.2 8.5 8.6 8.8

mcoll

min

1617.2

1615.4

1662.2

1663.3

1689.1

1689.1

2035.6

2036.2

2125.6

2127.6



22.4 100 100 100 n.a. 100

16 100 100 100

8 84.7 81.4 84.1

4 66.5 60.8 64.1

2 54.8 51.2 53.6

1 44.8 41.6 43.5

0.5 34.2 31.8 33.1

0.25 25.6 24.0 25.0

0.125 17.2 16.7 17.3

0.063 8.5 8.6 9.0

100

82.5

65.4

55.2

44.6

33.7

25.4

17.6

9.0

mcoll

min

1641.0

1643.3

1672.8

1671.2

1659.8

1661.0

Table-Appendix 33: PSD of untested and tested material in chapter 5.3: “G5_B150”, test IV after n=1536

Sieve: Tested I Tested II Tested III Tested IV Untested I Untested II PSD Theory

286

22.4 100 100 100 100 100 100 100

16 100 100 100 100 100 100 100

8 80.8 80.1 85.5 79.8 83.6 82.4 81.5

4 68.9 65.4 71.3 63.1 69.3 67.4 68.8

2 59.5 55.8 60.8 52.0 57.5 56.3 59.8

1 47.6 44.2 48.2 40.3 44.5 43.8 47.6

0.5 33.2 30.7 33.4 28.0 30.4 30.0 32.8

Florian V. Mueller

0.25 20.6 19.1 20.8 17.5 18.9 18.5 20.4

0.125 13.1 12.4 13.3 11.5 12.2 12.0 13.2

0.063 8.6 8.4 8.6 7.8 8.3 8.1 9.0

mcoll

min

1687.4

1691.0

1672.1

1670.0

1663.5

1666.5

1606.4

1606.9

1488.1

1488.8

1482.2

1483.9

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table-Appendix 34: PSD of untested and tested material in chapter 5.4: “G1_B0_Binder”, test III after n=1536

Sieve: Tested I Tested II Tested III Untested I Untested II

22.4 100 100 100 100 100

16 97.2 99.5 96.4 98.3 99.2

8 82.3 79.9 74.0 81.2 81.9

4 69.8 68.0 60.8 68.2 68.5

2 59.6 58.6 51.9 58.4 58.5

1 48.3 47.6 42.1 47.3 47.3

0.5 36.0 35.2 31.4 35.3 35.3

0.25 25.9 25.1 22.4 25.6 25.5

0.125 19.6 18.8 16.8 19.2 19.2

0.063 16.0 15.4 13.8 15.8 15.6

mcoll

min

1666.4

1668.1

1676.8

1677.3

1684.8

1686.5

1749.8

1749.4

1719.0

1720.4

Table-Appendix 35: PSD of untested and tested material in chapter 5.4: “G1_B150_Binder”, test III after n=1536

Sieve: Tested I Tested II Tested III Untested I Untested II

ICI Rheocenter

22.4 100 100 100 100 100

16 98.4 99.6 97.6 97.0 100

8 81.3 81.9 81.3 79.6 83.4

4 65.7 65.9 64.1 64.5 67.7

2 55.2 55.6 53.8 54.6 57.4

1 45.2 45.7 44.1 44.9 47.3

0.5 35.4 36.0 34.6 35.6 37.5

0.25 28.1 28.7 27.5 28.6 30.1

Florian V. Mueller

0.125 23.5 24.2 23.2 24.1 25.3

0.063 20.1 20.6 19.8 20.7 21.8

m coll

min

1668.8

1670.3

1689.8

1690.2

1687.9

1689.8

1852.6

1853.3

1707.6

1709.1

287

Appendix Chapter 6 Sedimentation If not explicitly stated otherwise, the rheology parameters given in the following table are the last before casting the specimen. Table-Appendix 36: Sedimentation and specific gravity distribution of specimens from Chapter 3.2.2

Blank_10 min Sf = 695 mm VSI 1.5 τ0 = 21 Pa μ = 12.4 Pa s Air = 2.0%

(bottom)

Surface area

(top)

Specific gravity distribution Spec.gravity x 1000 [-]

Mix

2500 2400 SGD

2300

Median

2200 0

10

20

30

40

Column height [cm]

Blank_30 min Sf = 575 mm VSI 0 τ0 = 46 Pa μ = 13.4 Pa s Air = 2.3%

Spec. Gravity x 1000 [-]

2500 2400 SGD

2300

Median

2200 0

10

20

30

40

High-ST6 _10 min Sf = 700 mm VSI 1.5 τ0 = 23 Pa μ = 11.9 Pa s Air = 2.0%

Spec.gravity x 1000 [-]

Column height [cm] 2500 2400 SGD

2300

Median

2200 0

10 20 30 40 Column height [cm]

High-ST6 _30 min Sf = 500 mm VSI 0 τ0 = 76 Pa μ = 16.9 Pa s Air = 2.7%


2500 2400 SGD

2300

Median

2200 0

10

20

30

40

Column height [cm]

288

Florian V. Mueller

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Low-ST6 _10min Sf = 695 mm VSI 1.5 τ0 = 18 Pa μ = 12.1 Pa s Air = 2.0%


2500 2400 SGD

2300

Median

2200 0


Low-ST6 _30min Sf = 585 mm VSI 0 τ0 = 45 Pa μ = 12.6 Pa s Air = 2.7%


2500 2400 SGD

2300

Median

2200 0

10

20

30

40

AEA-mix Sf = 600 mm VSI 0 Air = 5.35



2300

Median

2200 0


AEA-mix + SP & ST & Def Sf = 625 mm VSI 0 Air =2.0%


2500 2400 SGD

2300

Median

2200 0

10

20

30

40

AEA-mix + SP & ST & Def + water Sf = 680 mm VSI 0.5 Air =2.0%



2300

Median

2200 0

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289


High-ST7 _10min Sf = 690 mm VSI 1.5 τ0 = 20 Pa μ = 12.8 Pa s Air = 2.0%

2500 2400 SGD

2300

Median

2200 0




High-ST7 _30min Sf = 500 mm VSI 0 τ0 = 81 Pa μ = 18.0 Pa s Air = 3.4%


2500 2400 SGD

2300

Median

2200 0


2500 2400 SGD

2300

Median

2200 0




2500 2400 SGD

2300

Median

2200 0

10

20

30

40


290


Column height [cm]

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2500 2400 SGD

2300

Median

2200 0


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2500 2400 SGD

2300

Median

2200 0


2500 2400 SGD

2300

Median

2200 0




2500 2400 SGD

2300

Median

2200 0

10

20

30

40

High-ST10 _10min Sf = 675 mm VSI 1.5 τ0 = 26 Pa μ = 13.1 Pa s Air = 2.0%



2300

Median

2200 0




2500 2400 SGD

2300

Median

2200 0

10

20

30

40

Column height [cm]

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291



2500

2400 SGD

2300

Median

2200 0


BR-SCC Sf = 735 mm VSI 0.5 τ0 = 15 Pa μ = 20.4 Pa s Air = 2.0%




2500

2400 SGD

2300

Median

2200 0


2500 2400 SGD

2300

Median

2200 0


BR-SCC; PCE overdosed Sf = 860 mm VSI 3

292


2500

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2400 SGD

2300

Median

2200 0


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Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table-Appendix 37: Sedimentation results of SMA series (specific gravity x 1000)

G1_B0, τ0= 45 Pa, µ= 18.6 Pa s, Sf, initial =625 mm, VSI 2

(bottom)

Surface area

(top)

SG distribution

2500 Specific gravity

Mix

2400 2300 G1_B0

2200 0

20 30 Height [cm]

40

50

60



10

2400 2300 G1_B50

2200 0

20

30 Height [cm]

40

50

60



10

2400 2300 G1_B100

2200 0

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20 30 Height [cm]

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2400 2300 G1_B150_0,0

2200 0

10

20 Height [cm]

30

40

50

60

G1_L50, τ0= 27 Pa, µ= 23.2Pa s, Sf, initial =640 mm, VSI 2.5

Specific gravity

2500 2400 2300 100/600 150/300

2200 0

294

10

20 30 Height [cm]

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G1_L100, τ0= 22 Pa, µ= 21.9Pa s, Sf, initial =690 mm, VSI 0.5

G1_L150, τ0= 17 Pa, µ= 22.8 Pa s, Sf, initial =703 mm, VSI 1.5

ICI Rheocenter

Specific gravity

2500 2400 2300 100/600 150/300

2200 0

10

20 30 Height [cm]

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60

295

G1_F50, τ0= 32 Pa, µ= 16.3 Pa s, Sf, initial =673 mm, VSI 1.5

Specific gravity

2500

2400

2300 100/600 150/300

2200 0

10

20 30 Height [cm]

40

50

60

G1_F100, τ0= 22 Pa, µ= 15.9 Pa s, Sf, initial =715 mm, VSI 2

Specific gravity

2500

2400

2300 100/600 150/300

2200 0

296

10

20 30 Height [cm]

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60

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G1_FA150, τ0= 23 Pa, µ= 23.4 Pa s, Sf, initial =680 mm, VSI 0.5

Specific gravity

2500

2400

2300 100/600 150/300

2200 0

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20 30 Height [cm]

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40

50

60

297

Table-Appendix 38: Sedimentation results of AMA series, low powder mixes (specific gravity x 1000)

Mix

(bottom) Surface area

(top)

SG distribution

G1_B0_180, τ0= 50 Pa, µ= 21.7 Pa s, Sf, initial = 613 mm, VSI 2.5

Specific gravity

2500 2400 2300 100/550 150/300

2200 0

10

20 30 Height [cm]

40

50

60

G1_180_0.4, τ0= 44 Pa, µ= 19.9 Pa s, Sf, initial =608 mm, VSI 0.5

298

Specific gravity

2500 2400 2300 100/550 150/300

2200 0

10

20 30 Height [cm]

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40

50

60

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2500

G1_180_0.7, τ0= 54 Pa, µ= 15.8 Pa s, Sf, initial =623 mm, VSI 1.5

Specific gravity

100/550 150/300

2400

2300

2200 0

10

20 30 Height [cm]

40

50

60

2500

G1_180_1.0, τ0= 71 Pa, µ= 11.1 Pa s, Sf, initial = 645 mm, VSI 1

Specific gravity

100/550 150/300

2400

2300

2200 0

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20 30 Height [cm]

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50

60

299

2500 100/550

G1_0.7_185, τ0= 67 Pa, µ= 11.3 Pa s, Sf, initial =660 mm, VSI 1.5

Specific gravity

150/300

2400

2300

2200 0

10

20 30 Height [cm]

40

50

60

2500 100/550

G1_1.0_185, τ0= 59 Pa, µ= 11.6 Pa s, Sf, initial =598 mm, VSI 1.5

Specific gravity

150/300

2400

2300

2200 0

300

10

20 30 Height [cm]

Florian V. Mueller

40

50

60

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table-Appendix 39: Sedimentation results of AMA series, powder-rich mixes (specific gravity x 1000)

G1_B150_0.3, τ0= 37 Pa, µ= 25.7 Pa s, Sf, initial =635 mm, VSI 0

(Bottom)

Surface area

(top)

SG distribution

2500 G1_B150_0.3

Specific gravity

Mix

2400

2300

2200 0

10

20

30 Height [cm]

40

50

60

2500

G1_B150_0.4, τ0= 43 Pa, µ= 27.6 Pa s, Sf, initial =583 mm, VSI 0.5

Specific gravity

100/550 150/300

2400

2300

2200 0

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20 30 Height [cm]

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40

50

60

301

2500


Specific gravity

100/550 150/300

2400

2300

2200 0

10

20 30 Height [cm]

40

50

60

2500 G1_B150_0.9

Specific gravity


2400

2300

2200 0

302

10

20 30 Height [cm]

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40

50

60

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2500 100/550


Specific viscosity

150/300

2400

2300

2200 0

10

20 30 Height [cm]

40

50

60

2500 G1_B150_1.2

Specific gravity


2400

2300

2200 0

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10

20 Height [cm]

Florian V. Mueller

30

40

50

60

303

Table-Appendix 40: Sedimentation specimen of “G1_B80_0.4”; τ0 =38 Pa, μ =18.8 Pa s, Sf, initial = 625 mm, VSI 0.5 (specific gravity x 1000)

150 mm diameter column (h=600 mm)


d150/h600-I 2400 2300 2200 0

10

20

30 Height [cm]

40

50

60


d150/h600-II 2400 2300 2200 0

304

10

20

30 Height [cm]

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60

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d150/h600-III 2400

2300

2200 0

10

20

30 Height [cm]

40

50

60

2500

Specific gravity

d150/h600-Mean3

2400

2300

2200 0

10

20 30 Height [cm]

40

50

60

150 mm diameter cylinder (h=300 mm)

2500

Specific gravity

d150/h300-I 2400

2300

2200 0

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20

30 Height [cm]

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40

50

60

305


d150/h300-II 2400

2300

2200 0

10

20

30 Height [cm]

40

50

60


d150/h300-III 2400

2300

2200 0

306

10

20

30 Height [cm]

Florian V. Mueller

40

50

60

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Average

2500

Specific gravity

d150/h300-Mean3

2400

2300

2200 0

10

20

30

40

50

60

Height [cm]

100 mm diameter column (h=600 mm) 2500 Specific gravity

d100/h600-I 2400 2300 2200 0

10

20

30 Height [cm]

40

50

60


d100/h600-II 2400 2300 2200 0

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20

30 Height [cm]

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307


d100/h600-III 2400 2300 2200 0

Average

10

20

30 Height [cm]

40

50

60

Specific gravity

2500

2400

2300 d100/h600-Mean3

2200 0

308

10

20 30 Height [cm]

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40

50

60

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table-Appendix 41: Sedimentation results of gap-graded mixtures; see Oesterheld et.al. [242]

Mix

(bottom)

Surface area

(top)

SG distribution

OPC525_0.66 _0.87 τ0= 21 Pa, µ= 41.5 Pa s, Sf, initial =755 mm, VSI 3

Specific gravity

2500 2400 300 er

2300

CStr-cyl

2200

0

10 Height [cm]

20

30

OPC600_0.66 _0.87, τ0= 5 Pa, µ= 56.9 Pa s, Sf, initial =725 mm, VSI 3

Specific gravity

2500 2400 2300

300er Cyl CStr

2200 0

10 20 Height [cm]

30

OPC675_0.0_ 0.6, τ0= 6 Pa, µ= 50.0 Pa s, Sf, initial =730 mm, VSI 0.5

Specific gravity

2500 2400 2300

300er Cyl CStr

2200 0

10 20 Height [cm]

30

OPC675_0.0_ 0.80, τ0= 15 Pa, µ= 40.6 Pa s, Sf, initial =765 mm, VSI 3

Specific gravity

2500 2400 300er

2300

Cyl CStr

2200 0

10 20 Height [cm]

30

OPC675_0.54 _0.80, τ0= 15 Pa, µ= 50.5 Pa s, Sf, initial =730 mm, VSI 1

Specific gravity

2500 2400 300er

2300

Cyl CStr

2200 0

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Florian V. Mueller

10 20 Height [cm]

30

309

Specific gravity

2500 OPC675_0.54 _0.87, τ0= 12 Pa, µ= 66.7 Pa s, Sf, initial =730 mm, VSI 1

2400 2300

300er Cyl CStr

2200 0

10 20 Height [cm]

30

Specific gravity

2500 SRC675_0.54_ 0.83, τ0= 40 Pa, µ= 61.3 Pa s, Sf, initial =835 mm, VSI 0

2400 2300

Mean of 3 Cyl CStr 6

2200 0

10 20 Height [cm]

30

Table-Appendix 42: Sedimentation data from robustness evaluation of binder-rich SCC

Mix

(bottom)

Surface area

(top)

SG distribution

-20%SP τ0= 12 Pa, µ= 65.5 Pa s, Sf, initial =605 mm, VSI 0

Specific gravity

2500 2400 2300 -20%SP 2200 0

10 20 Height [cm]

30

-10% SP, τ0= -575 Pa, µ= 56.6 Pa s, Sf, initial =650 mm, VSI 0

Specific gravity

2500 2400 2300 -10%SP 2200 0

10 20 Height [cm]

30

For convenience, the data are used when the Bingham model was applied by default, which might yield relatively inaccurate values for this mix design. 75

310

Florian V. Mueller

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Ref I , τ0= -16 Pa, µ= 57.0 Pa s, Sf, initial = 680 mm, VSI 0

Specific gravity

2500 2400 2300 Ref I 2200 0

10 20 Height [cm]

30

+10% SP, τ0= -22 Pa, µ= 54.0 Pa s, Sf, initial =715 mm, VSI 1

Specific gravity

2500 2400 2300 +10%SP 2200 0

10 20 Height [cm]

30

+20% SP, τ0= -20 Pa, µ= 48.6 Pa s, Sf, initial =743 mm, VSI 1.5

Specific gravity

2500 2400 2300 -20%SP 2200 0

10 20 Height [cm]

30

-10l w, τ0= -13 Pa, µ= 122.8 Pa s, Sf, initial =530 mm, VSI 0

Specific gravity

2500 2400 2300 -10l W

2200 0

10 20 Height [cm]

30

-5l w, τ0= -8 Pa, µ= 109.1 Pa s, Sf, initial =573 mm, VSI 0

Specific gravity

2500 2400 2300 -5l W 2200 0

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10 20 Height [cm]

30

311

Ref II, τ0= -13 Pa, µ= 62.2 Pa s, Sf, initial =658 mm, VSI 0

Specific gravity

2500 2400 2300 Ref II 2200 0

10 20 Height [cm]

30

+5l w, τ0= -6 Pa, µ= 45.3 Pa s, Sf, initial =695mm, VSI 1

Specific gravity

2500 2400 2300 +5l W 2200 0

10 20 Height [cm]

30

+10l w, τ0= -5 Pa, µ= 33.7 Pa s, Sf, initial =743 mm, VSI 2

Specific gravity

2500 2400 2300 +10l W

2200 0

10 20 Height [cm]

30

Note: Four other mix designs will not be given in detail, since they were not used in one of the chapters apart from their sedimentation results in chapter 6. For questions about their details, please contact the author if necessary.

312

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table-Appendix 43: Relation of VSI, blocking step hJ and slump flow difference dS to the Separation point of the rheology measurement.

-30

3

2 1 0 -20 -10 0 Separation point [%]

10

(a): Generalized VSI vs. Separation point from Viscometer measurement

-30

40 30 20 10 0 -20 -10 0 Separation point [%]

10

(c): Generalized hJ versus Separation point

2 High air content 1 0 -20 -10 0 10 Separation point [%]

-30

50 40

10 0 -20 -10 0 10 Separation point [%]

(d): Detailed hJ versus Separation point

-20

0

10

-60 Separation point [%] (e): Generalized dS(f-fj) versus Separation point

ICI Rheocenter

Difference dS (f-fj) [mm]

Difference dS (f-fj) [mm]

20 -10

High air content Insufficent PSD / matrix

20

100

60

-20

All

30

100

-30

Insufficent PSD / matrix

(b): Detailed VSI vs. Separation point

Height step J-ring [mm]

50

Height step J-ring [mm] -30

All

VSI ASTM 1611 [-]

VSI ASTM 1611 [-]

3

-30

All

60 20 -20

-10

-20

0

10

High air content Insufficent PSD / matrix

-60 Separation point [%] (f): Detailed dS(f-fj) versus Separation point

Florian V. Mueller

313

Table-Appendix 44: Correlation of sedimentation values to blocking values

4 4 All

All

3 2

SCC gap Eco-AIR

1

HVSI [-]

HVSI [-]

3

SCC gap

2 EcoAIR

1 0 -60 -40 -20 0 20 40 60 80 100 dS (f-fj) [mm]

0 0

Figure A 1: Hardened Visual Stability Index vs. difference of Sf with and without J-ring (dS (f-fj)

5

All

4 3

SCC gap

2

Eco-AIR

1

CSGD [-]

CSGD [-]

50

6

5

All

4

SCC gap

3

EcoAIR

2 1

0 -60 -40 -20 0 20 40 60 80 100 dS (f-fj) [mm]

0 0

Figure A 3: Classes of specific gravity distribution vs. dS (f-fj)

10

20 30 40 Hj [mm]

50

Figure A 4: CSGD vs. Blocking step J-ring (hj)

4

4 All SCC gap

2

Eco-AIR

1 0 -60 -40 -20 0 20 40 60 80 100 dS (f-fj) [mm]

Figure A 5: Aggregate surface appearance index vs. dS (f-fj)

Florian V. Mueller

All

3 ASAI [-]

3 ASAI [-]

20 30 40 Hj [mm]

Figure A 2: HVSI vs. Blocking step J-ring (hj)

6

314

10

SCC gap

2

EcoAIR

1 0 0

10

20 30 40 Hj [mm]

50

Figure A 6: ASAI vs. Blocking step J-ring (hj)

ICI Rheocenter


Appendix Chapter 7 Robustness

100%

500

80%

Fines

60%

AIR W

40%

SF

20%

CEM

0% 178

181

186

Water content

193

200


Matrix proportion vol-%

Table-Appendix 45: Matrix proportions (left column) and cumulative volumes (right columns) for water-seriesmixes during the robustness evaluation for Eco-SCC (line 1), lean SCC (line 2), and binder-rich SCC (third line)

AIR W

200

SF

100

CEM

0 Water content [kg/m3] 500

Fines

80%

Air

60%

water L

40%

FA CEM

20% 0%



300

170 180 190 200 210

177 183 189 195 201 Water content [kg/m3]

Fines Air water L FA CEM

400 300 200 100 0 170 180 190 200 210 Water content [kg/m3]

100%

500 Fines

80%

Air

60%

water

40%

FA

20%

CEM

0% 175 177 186 191 196 Water content [kg/m3]



Fines

[kg/m3]

100%

ICI Rheocenter

400

400 Fines

300

Air water

200

FA

100

CEM

0 170 180 190 200 210 Water content [kg/m3]

Florian V. Mueller

315

Table-Appendix 46: Evolution of fluid model parameters of binder-rich SCC of the robustness test, water-series.

18 16

175 l

y = 33.768x - 0.1639 R² = 0.9939

180 l

y = 25.864x - 0.1747 R² = 0.9946

185 l

y = 15.114x - 0.1287 R² = 0.9903 y = 10.715x - 0.0832 R² = 0.9933

Torque [Nm]

14 12 10 8 6

190 l

4 2

195 l

0 0

0.1


y = 6.8218x - 0.0131 R² = 0.993

0.5

(a): Flow curves with linear model, not omitted

5

Resistance [mV]

4 175 L 180 L

3

185 L 2

190 L 195 L

1 0 0

50

100

150 200 250 300 350 Data point @ varied velocities

400

450

(b): Torque response at different rotational velocities, accumulated as data points

316

Florian V. Mueller

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Design criteria for low binder Self-Compacting Concrete, Eco-SCC 18 16

175 l

y = 32.851x - 0.0123 R² = 0.9932

180 l

y = 25.864x - 0.1747 R² = 0.9946

185 l

y = 14.315x + 0.0033 R² = 0.9944

190 l

y = 10.287x - 0.0125 R² = 0.995

195 l

y = 6.8218x - 0.0131 R² = 0.993

175 l

y = 28.375x0.9004 R² = 0.9892

180 l

y = 23.28x0.945 R² = 0.9923

185 l

y = 12.627x0.9124 R² = 0.993

Torque [Nm]

14 12 10 8 6 4 2 0 0

0.1


0.5

(c): Flow curves with linear model, omitted

18 16

Torque [Nm]

14 12 10 8 6

190 l

y = 9.1749x0.9251 R² = 0.994

195 l

y = 5.9747x0.9072 R² = 0.9909

4 2 0 0

0.1


0.5

(d): Flow curves with power law model, omitted

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Florian V. Mueller

317

18 16

175 l

y = 24.282x2 + 21.915x + 0.8578 R² = 0.9993

180 l

y = 14.615x2 + 18.452x + 0.4863 R² = 0.9993

Torque [Nm]

14 12 10

2 185 l y = 9.4394x + 10.064x + 0.3415 R² = 0.9992

8 6

2 190 l y = 6.5197x + 7.351x + 0.2211 R² = 0.9995

4 2

2 195 l y = 4.4472x + 4.5662x + 0.188 R² = 0.9994

0 0

0.1


0.5

(e): Flow curves with polynomial model second order, omitted

Table-Appendix 47: Evolution of fluid model parameters of binder-rich SCC during robustness test, SP-series.

10 0.72%

Torque [Nm]

8 0.81%

y = 14.122x - 0.0842 R² = 0.9927

0.90%

y = 12.896x - 0.1235 R² = 0.9916

0.99%

y = 12.38x - 0.2936 R² = 0.9859

1.08%

y = 13.238x - 0.2551 R² = 0.9898

6 4 2 0 0

0.1


y = 16.16x + 0.214 R² = 0.9918

0.5

(a) Flow curves with linear model, not omitted

318

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Resistance (V)

3

0.72%

2

0.81% 0.90% 0.99%

1

1.08%

0 0

50

100

150 200 250 300 350 Data point @ varied velocities

400

450

(b) ConTec Viscometer 5 raw data of robustness tests binder-rich SCC, SP-series

10 0.72%

y = 14.714x + 0.4349 R² = 0.9969

0.81%

y = 14.122x - 0.0842 R² = 0.9927

0.90%

y = 12.303x - 0.0255 R² = 0.9942

0.99%

y = 13.238x - 0.2551 R² = 0.9898

1.08%

y = 11.057x - 0.0882 R² = 0.99

0.72%

y = 12.2x0.7833 R² = 0.9938

0.81%

y = 12.41x0.9245 R² = 0.9884

0.90%

y = 10.993x0.9295 R² = 0.9937

0.99%

y = 12.038x0.9995 R² = 0.9903

1.08%

y = 10.36x0.9886 R² = 0.9946

Torque [Nm]

8 6 4 2 0 0

0.1


0.5

(c) Flow curves with linear model, omitted

10

Torque [Nm]

8 6 4 2 0 0

ICI Rheocenter

0.1


0.5

Florian V. Mueller

319

(d) Flow curves with power law model, omitted

10 0.72%

Torque [Nm]

8

y = 1.5118x2 + 14.11x + 0.478 R² = 0.997

2 0.81% y = 9.1152x + 9.4985x + 0.328 R² = 0.9989

6

2 0.90% y = 8.0864x + 8.661x + 0.2642 R² = 0.9989

4

2 0.99% y = 10.624x + 7.8496x + 0.2254 R² = 0.9993

2

1.08% y = 8.8151x2 + 7.5314x + 0.1631 R² = 0.9955

0 0

0.1 0.2 0.3 0.4 Rotational velocity [rps]

0.5

(e) Flow curves with polynomial model second order, omitted

320

Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table-Appendix 48: Results of isothermal calorimetry for Eco-SCC, water-series

Power mW/g

6

4

21.8°C

w178

22.4°C

w181

22.5°C

w186

21.9°C

w193

2

w200

0 0

4


20

24

21.8°C 22.4°C

2.5°C

(a) Power-equivalent heat of hydration of Eco-SCC, water series

120

8h 12h

100

24h

80

48h

60 -4

-3 0 3 7 Variation of water [%]

(b) Varied heat of hydration at various ages for variation of water during a robustness test

ICI Rheocenter

Energy [%] per 48h

100

21.8°C 22.4°C 22.5°C 21.9°C 21.7°C

Energy J/g per reference

140

21.9°

21.7°C

80 8h

60

12h

40

24h

20

48h

0 -4


(c) Heat of hydration related to consumption at 48h

Florian V. Mueller

321

Table-Appendix 49: Results of isothermal calorimetry for Eco-SCC, SP-series

Power mW/g

6

21.4°C

4

SP0.88%

21.0°C

SP0.99%

22.7°C

SP1.10%

20.9°C

SP1.21%

2

SP1.32%

0 0

160 140


21.4°C 21.0°C 22.7°C 20.9°

100

21.8°C

120

8h

100

12h

80

24h

60

48h

24

21.4°C 21.0°C

22.7°C 20.9°

21.8°C

80 8h

60

12h

40

24h

20

48h

0

40 -21

-10 0 9 Variation of SP [%]

-21

20


322

20

Power-equivalent heat of hydration of Eco-SCC, SP-series

Energy [%] per 48h


(a)

4


20


Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table-Appendix 50: Results of isothermal calorimetry for Lean SCC, SP-series

Power mW/g

6 19.4°C

SP0.55%

19.8°C

SP0.61%

19.6°C

SP0.68%

20.2°C

SP0.75%

20.8°C

SP0.85%

4

2

0 0

160


20

24

Power-equivalent heat of hydration of lean SCC, SP-series 19.4°C 19.8°C 19.6°C

100

20.2°C 20.8°C

140 120

8h

100

12h

80

24h 48h

60 40

Energy [%] per 48h


(a)

4

19.4°C

19.8°C 19.6°C 20.2°C 20.8°C

80 8h

60

12h

40

24h

20

48h

0 -19


26

-19

(b) Varied heat of hydration at various ages for variation of SP dosage during a robustness test

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323

Table-Appendix 51: Results of isothermal calorimetry for br-SCC, water-series

6 20.1°C

w175

Power mW/g

20.0°C 4

w180

21.2°C

w185

20.3°C

w190

19.9°C

2

w195

0 0

4


20

24

(a) Power-equivalent heat of hydration of br-SCC, water series 20.3°C

100

19.9°C

120

8h 12h

100

24h

80

48h

60 -5

20.1°C 20.0°C 21.2°C

20.3° C 19.9°C

80 60

8h 12h

40

24h

20

48h

0


-5


324

Energy [%] of 48 h

20.1°C 20.0°C 21.2°C


140



Florian V. Mueller

ICI Rheocenter

Design criteria for low binder Self-Compacting Concrete, Eco-SCC Table-Appendix 52: Results of isothermal calorimetry for br-SCC, SP-series

Power mW/g

6

4

20.4°C

SP0.71%

21.2°C

SP0.81%

21.6°C

SP0.90%

20.5°C

SP0.99%

20.7°C

2

SP1.08%

0 0

4

8

12 16 Hydration time [min]

20

24

140

20.4°C 21.2°C

100

21.6°C 20.5°C 20.7°C

120

8h

100

12h 24h

80

48h

60 -21

20.4°C 21.2°C 21.6°C

20.5°C 20.7°C

80 60

8h 12h

40

24h

20

48h

0

-10 0 9 20 Variation of SP [%]

-21


ICI Rheocenter

Energy [%] of 48h


(a) Power-equivalent heat of hydration of br-SCC, SP-series


20


Florian V. Mueller

325

Table-Appendix 53: Compressive strength at 28d of robustness test series, as being affected by: alteration of composite quantities (a1) and (a2), density (b1) and (b2), and the degree of consolidation (c1) and (c2)

100

80 br-SCC

60

lean SCC Eco-SCC

40 20

Compr. strength 28d [MPa]


100

80

SCC

20

(a2) Effect of altered water content


100 80 60

br-SCC lean SCC

40 20 2350

Eco-SCC

2375 2400 2425 Density [Kg/m3]

-10 -5 0 5 10 Relative variation of water [%]


(a1) Effect of altered SP dosage

2450

br-SCC lean SCC

20 1.000

Eco-SCC

1.005 1.010 1.015 SG / Density

1.020

(c1) Effect of consolidation degree as ratio of specific gravity to density, SP-series


80

40

80 br-SCC

60

lean SCC

40 20 2350

Eco-SCC

2375

2400

2425

2450

(b2) Effect of density, water-series

100

60

100

Density [kg/m3]

(b1) Effect of density, SP-series


Eco-SCC

40

-30 -20 -10 0 10 20 30 Relative variation of SP [%]

326

br-SCC

60

100 80 br-SCC

60

lean SCC Eco-SCC

40 20 1.000

1.005

1.010 1.015 SG/Density

1.020

(c2) Effect of consolidation degree as ratio of specific gravity to density, water-series

Florian V. Mueller

ICI Rheocenter


Curriculum Vitae Name: Birth:

Vitus Florian Mueller June 8, 1976 in Jena, Germany

Education: 10/1996-05/2007

Bauhaus-University Weimar, Germany Graduated as Diplombauingenieur (concomitant to Master of Science in Civil Engineering)

04/2007-01/2009:

Reykjavik University, Iceland Graduated as Master of Science in Concrete Technology

Work experience: 04/2007-12/2009 01/2010-09/2011 01/2012 – current

Research engineer at Innovation Center Iceland Concrete Industrialization Specialist at CEMEX Research Group AG, CH Consultant at Holcim Group Support Ltd., CH

Publications: S. Kubens, F. Müller, O.H. Wallevik: Influence of cement-admixture interaction on the stability of production properties of SCC, in: G. de Schutter and V. Boel (Eds.): Proceedings of the fifth international RILEM Symposium “SCC 2007” in Gent, RILEM Proceedings PRO 54 Vol. 1, pp. 211-216, 2007 Florian V. Mueller, Olafur H. Wallevik: Benefits of filler material on rheology in Eco-SCC, in: Proceedings of the Third North American Conference on the design and use of SCC “SCC 2008” in Chicago, IL., ACBM, 2008 Sonja Oesterheld, Florian V. Mueller, Olafur H. Wallevik: The influence of workability loss and thixotropy on formwork pressure of self-compacting concrete containing stabilizers, in: Proceedings of the Third North American Conference on the design and use of SCC “SCC 2008” in Chicago, IL., ACBM, 2008 Olafur H. Wallevik, Florian V. Mueller, and Björn Hjartarson: Eco-SCC, an environmental and economical alternative, in: Proceedings of the ICCX Oceania, Sydney, 2009 Florian V. Mueller and Olafur H. Wallevik: Effect of maximum aggregate size in air-entrained Eco-SCC; in: C. Shi, Z. Yu, K.H. Khayat and P. Yan (Eds.): Rilem PRO 65 Proceedings of the 2nd International Symposium on Design, Performance and Use of Self-Consolidating Concrete SCC’2009, Beijing, pp. 664-673, 2009 O.H. Wallevik, F.V. Mueller, B. Hjartarson, and S. Kubens: The green alternative of self-compacting concrete, Eco-SCC; Ibausil 2009, Weimar, Germany, Vol. 1, pp. 1-1105 to 1-1116, 2009 Florian V. Mueller and Olafur H. Wallevik: Robustness of very low binder; self-compacting concrete: Eco-SCC, in: O.H. Wallevik, St. Kubens and S. Oesterheld (Eds.): Rheology of Cement Suspensions such as Fresh Concrete, Rilem Pro 68 Proceedings of the 3rd International RILEM Symposium, Reykjavik, Iceland, 2009 Olafur H. Wallevik, Florian V. Mueller, Björn Hjartarson: Rheology of Self Compacting Concrete and Eco-SCC; Proceedings of the 2nd International Conference Advances In Concrete Technology In the Middle East, Abu Dhabi UAE, 2009 Florian V. Mueller and Olafur H. Wallevik: Effect of limestone filler addition in Eco-SCC, in: K. Khayat and D. Feys (Eds.): Proceedings of 6th International RILEM Symposium on SCC & 4th North American Conference on the Design and Use of SCC, September 2010, Montreal, Quebec, Supplementary papers (CD-ROM), pp.107-116 S. Oesterheld, F.V. Mueller, S. Kubens, and O.H. Wallevik: Rheological Evaluation of Ingredients for SCC, in: K. Khayat and D. Feys (Eds.): Proceedings of 6th International RILEM Symposium on SCC & 4th North American Conference on the Design and Use of SCC, September 2010, Montreal, Quebec, Supplementary papers (CD-ROM), pp. 445-452 Florian V. Mueller and Olafur H. Wallevik: Particle packing by Gyratory Intensive Compaction as tool to optimize the aggregate gradation of low binder SCC, Eco-SCC, Invited special presentation at “OW11” Our World in Concrete & Structures, 14-16. August, Singapore

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327