Development of a Portable Rheometer for Fresh Portland Cement Concrete

RESEARCH REPORT ICAR –105-3F

Sponsored by the Aggregates Foundation for Technology, Research and Education Technical Report Documentation Page 1. Report No. 2. Government Accession No. 3. Recipient's Catalog No. ICAR 105-3F

4. Title and Subtitle 5. Report Date August 2004

DEVELOPMENT OF A PORTABLE RHEOMETER FOR FRESH 6. Performing Organization Code PORTLAND CEMENT CONCRETE

7. Author(s) 8. Performing Organization Report No. Eric P. Koehler and David W. Fowler Research Report ICAR 105-3F

9. Performing Organization Name and Address 10. Work Unit No. (TRAIS)

International Center for Aggregates Research The University of Texas at Austin 11. Contract or Grant No. ECJ 5.200 Project No. ICAR-105 Austin, Texas 78712-1076 12. Sponsoring Agency Name and Address 13. Type of Report and Period Covered Aggregates Foundation for Technology, Research, and Education Research Report 1415 Elliot Place NW September 2002-August 2004

Washington, D. C. 20007 14. Sponsoring Agency Code

15. Supplementary Notes

16. Abstract:

Based on a literature search in which 61 existing workability test methods were identified and on feedback from government, industry, and academia, criteria for an improved workability test device were developed. It was determined that the best approach to measuring workability would be to develop a new portable rheometer.

The ICAR rheometer—a low-cost, fully portable test device for concrete—was developed and tested. A first generation prototype was built using off-the-shelf components. The ICAR rheometer is approximately the size of a drill and can be operated by hand or positioned above a standard container. It is capable of measuring a flow curve or performing a stress growth test and is appropriate for nearly the full range of concrete workability ranging from a slump of approximately 2 inches to self-consolidating concrete.

Experimental testing on a wide range of concrete mixtures indicated that the ICAR rheometer was able to detect changes in workability and rheology successfully. As a dynamic test that adds energy to concrete, it is well suited for measuring high-microfines concrete and other highly thixotropic concrete mixtures. Field testing confirmed the portability of the ICAR rheometer. The low cost and portable form factor of the ICAR rheometer can make the routine measurement of concrete rheology in the field an economically viable solution to characterizing concrete workability.

17. Key Words 18. Distribution Statement

ICAR rheometer, fiber reinforcement, ground granulated blast No restrictions. furnance slag, aggregates, slump test, concrete workability, high microfines

19. Security Classif.(of this report) 20. Security Classif.(of this page) 21. No. of Pages 22. Price Unclassified Unclassified 306 Form DOT F 1700.7 (8-72) Reproduction of completed page authorized

DEVELOPMENT OF A PORTABLE RHEOMETER FOR FRESH PORTLAND CEMENT CONCRETE

Eric P. Koehler The University of Texas at Austin Austin, TX

and

David W. Fowler The University of Texas at Austin Austin, TX

ICAR Report 105-3F ICAR 105: Measuring the Workability of High Fines Concrete

Sponsored by: International Center for Aggregates Research The University of Texas at Austin

Aggregates Foundation for Technology, Research and Education (AFTRE)

August 2004

ACKNOWLEDGEMENTS

The research described in this report was conducted at the International Center for Aggregates Research at The University of Texas at Austin. The authors gratefully acknowledge the funding provided by the Aggregates Foundation for Technology, Research, and Education. The authors also wish to thank the following companies for contributing materials used in the research (in alphabetical order): Aggregate Industries, Boral Material Technologies, Capitol Aggregates, Edward C. Levy Company, Grace Construction Products, Holcim (US), Sika Corporation, and Vulcan Materials Company. The BTRHEOM rheometer used in the research was on loan from the Federal Highway Administration. The authors acknowledge the helpful input from the participants of the concrete workability workshop in May 2003 and Degussa Admixtures for assisting in facilitating the workshop. The authors thank the National Institute of Standards and Technology for organizing the field testing where the ICAR rheometer was used. Finally, the authors gratefully acknowledge the assistance of Chiara Ferraris of the National Institute of Standards and Technology in serving as a consultant to the research project.

ii ABSTRACT

The purpose of this research was to identify an effective field test method for measuring the workability of concrete in general and of high-microfines concrete in particular. The workability of fresh concrete has traditionally been measured with the slump test, which provides an inadequate indication of workability. For certain concrete mixtures—such as those containing fiber reinforcement, ground granulated blast furnace slag, or high contents of aggregate microfines—the slump test can provide inaccurate and misleading results. The need for a better test method for workability is well established within the concrete industry. Based on a literature search in which 61 existing workability test methods were identified and on feedback from government, industry, and academia, criteria for an improved workability test device were developed. It was determined that the best approach to measuring workability would be to develop a new portable rheometer. The ICAR rheometer—a low-cost, fully portable test device for concrete—was developed and tested. A first generation prototype was built using off-the-shelf components. The ICAR rheometer is approximately the size of a drill and can be operated by hand or positioned above a standard container. It is capable of measuring a flow curve or performing a stress growth test and is appropriate for nearly the full range of concrete workability ranging from a slump of approximately 2 inches to self-consolidating concrete. Experimental testing on a wide range of concrete mixtures indicated that the ICAR rheometer was able to detect changes in workability and rheology successfully. As a dynamic test that adds energy to concrete, it is well suited for measuring high-microfines concrete and other highly thixotropic concrete mixtures. Field testing confirmed the portability of the ICAR rheometer. The low

iii cost and portable form factor of the ICAR rheometer can make the routine measurement of concrete rheology in the field an economically viable solution to characterizing concrete workability.

iv TABLE OF CONTENTS

LIST OF TABLES ...... xi

LIST OF FIGURES...... xiv

CHAPTER 1: INTRODUCTION ...... 1 1.1 Research Background...... 1 1.2 Historical Perspective...... 4 1.3 Research Objectives ...... 8 1.4 Project Scope...... 9

CHAPTER 2: FLUID RHEOLOGY...... 11 2.1 Introduction ...... 11 2.2 Properties of Fluids ...... 11 2.2.1 Definition of a Liquid...... 12 2.2.2 Constitutive Equations for Fluid Flow...... 14 2.2.3 Thixotropy and Anti-Thixotropy...... 17 2.2.4 Dilatancy ...... 18 2.2.5 Viscosity...... 18 2.2.5.1 Definition ...... 18 2.2.5.2 Origin of Viscosity in Fluid Suspensions...... 21 2.2.6 Yield Stress ...... 23 2.3 Measurement of Rheology ...... 25 2.3.1 Capillary Tube Viscometers...... 25 2.3.1.1 Description ...... 25 2.3.1.2 Derivation of Equations for Capillary Tube Viscometers...... 26 2.3.2 Rotational Rheometers ...... 29 2.3.2.1 Description ...... 29 2.3.2.2 Derivation of Equations for Coaxial Cylinders Rheometers...... 31

v 2.3.2.3 Effects of Dead Zones in Coaxial Cylinders Rheometers...... 37 2.3.2.4 Methods to Correct for Dead Zones in Coaxial Cylinders Rheometers ...... 45 2.3.2.5 End Effects...... 51 2.3.3 Special Topics in the Measurement of Rheology...... 52 2.3.3.1 Deborah Number...... 52 2.3.3.2 Estimation of Shear Rate...... 53 2.3.3.3 Relative Rheometers ...... 54 2.3.3.4 Slip at Boundaries (Wall Effect) ...... 56 2.4 Rheology of Concentrated Suspensions...... 57 2.4.1 Direct Yield Stress Measurements in Rotational Rheometers ... 57 2.4.2 Use of Vane for Direct Yield Stress Measurements ...... 60 2.4.2.1 Overview ...... 60 2.4.2.2 Vane Dimensions ...... 61 2.4.2.3 Location of Yielding...... 63 2.4.2.4 End Effects and the Calculation of Yield Stress ...... 65 2.4.2.5 Effect of Rotation Speed on Stress Growth Test Results ...... 69 2.4.2.6 Measurement System Stiffness ...... 70 2.4.3 Extension of Vane Geometry to Viscosity Measurements...... 70 2.5 Application of Fluid Rheology Concepts to Concrete ...... 72

CHAPTER 3: FACTORS INFLUENCING CONCRETE RHEOLOGY AND WORKABILITY...... 77 3.1 Introduction ...... 77 3.2 Effects of Cement...... 78 3.2.1 Cement Content...... 78 3.2.2 Cement Characteristics...... 78 3.3 Effects of Water Content...... 79 3.4 Effects of Aggregates ...... 80

vi 3.4.1 Aggregate Volume Fraction...... 80 3.4.2 Sand-to-Aggregate Ratio...... 80 3.4.3 Shape and Texture...... 81 3.4.4 Gradation...... 81 3.4.5 Microfines Content...... 82 3.5 Effects of Chemical Admixtures ...... 83 3.5.1 Water-Reducing Admixtures...... 83 3.5.2 Air Entrainment Agents ...... 86 3.5.3 Viscosity Modifying Admixtures...... 86 3.6 Effects of Supplementary Cementitious Materials...... 87 3.6.1 Fly Ash ...... 87 3.6.2 Silica Fume...... 88 3.6.3 Ground Granulated Blast Furnace Slag...... 89 3.7 Effects of Fibers ...... 90 3.8 Summary ...... 90

CHAPTER 4: EVALUATION OF POTENTIAL APPROACHES TO WORKABILITY CHARACTERIZATION ...... 93 4.1 Introduction ...... 93 4.2 Assessment of Existing Workability Test Methods ...... 93 4.3 Feedback from Industry, Government, and Academia...... 98 4.4 Criteria for New Workability Test Methods ...... 100 4.5 Selection of the Most Promising Approach for Workability Characterization ...... 102

CHAPTER 5: DEVELOPMENT OF THE ICAR RHEOMETER ...... 105 5.1 Introduction ...... 105 5.2 Overview of Development Process...... 105 5.3 Selection of Potential Impellers ...... 107 5.3.1 General Concepts ...... 108 5.3.2 Evaluation of Existing Impellers...... 111

vii 5.3.2.1 Tattersall Two-Point Device ...... 111 5.3.2.2 IBB Rheometer...... 116 5.3.2.3 Fresh Concrete Tester (FCT 101)...... 117 5.3.2.4 Commercially Available Mixing Impellers...... 118 5.3.3 Impeller Selection ...... 121 5.4 Evaluation of Conventional Hand Drill Technology ...... 121 5.5 Selection of Components ...... 123 5.6 Software ...... 129 5.6.1 Graphical User Interface ...... 129 5.6.2 Internal Software Operation ...... 136

CHAPTER 6: EXPERIMENTAL DETERMINATION OF OPERATING CHARACTERISTICS OF THE ICAR RHEOMETER...... 139 6.1 Introduction ...... 139 6.2 Materials, Mixtures, and Test Procedures...... 139 6.3 Impeller Type...... 141 6.3.1 Series I Tests – Qualitative Observations ...... 143 6.3.2 Series II Tests – Quantitative Performance...... 151 6.3.3 Series III Tests – Spiral Impellers...... 163 6.3.3.1 Background ...... 163 6.3.3.2 Qualitative Descriptions of Observed Flow Behavior ...... 166 6.3.3.3 Quantitative Measurements...... 170 6.3.3.4 Conclusions ...... 173 6.3.4 Final Selection...... 174 6.4 Optimum Speed for Stress Growth Tests...... 175 6.5 Gap Sizes...... 178

viii 6.6 End Effects...... 185 6.7 Conclusions ...... 193

CHAPTER 7: LABORATORY TESTING PROGRAM...... 195 7.1 Introduction ...... 195 7.2 Materials...... 195 7.3 Mixture Proportions...... 201 7.4 Mixing and Testing Procedures...... 202 7.5 Test Results for Conventional Concrete...... 207 7.5.1 Fly Ash ...... 207 7.5.2 Ground Granulated Blast Furnace Slag...... 209 7.5.3 Silica Fume...... 211 7.5.4 Water-to-Cement Ratio ...... 213 7.5.5 Water-Reducing Admixtures...... 215 7.5.6 Air-Entraining Agent...... 217 7.5.7 Blends of Natural and Manufactured Sand ...... 219 7.5.8 Sand-to-Aggregate Ratio...... 221 7.5.9 Aggregate Microfines...... 223 7.5.10 Slag Aggregate ...... 226 7.6 Rheometer Performance...... 228 7.6.1 Repeatability of Test Results...... 228 7.6.2 Calculation of Test Results ...... 231 7.7 Relationships between Rheology and Workability ...... 235 7.8 Self-Consolidating Concrete...... 239 7.8.1 Mixture Proportions and Test Procedures...... 239 7.8.2 Test Results ...... 242 7.8.3 Relationships between Rheology and Workability ...... 250 7.8.4 Rheometer Performance...... 257

ix 7.9 Field Testing...... 262 7.10 Workability Range...... 266 7.11 Conclusions ...... 268

CHAPTER 8: SUMMARY, CONCLUSIONS, AND RECOMMENDATIONS ...... 271 8.1 Summary ...... 271 8.2 Conclusions ...... 272 8.2.1 Concrete Workability Characterization...... 272 8.2.2 Development of the ICAR Rheometer ...... 274 8.2.3 Testing of ICAR Rheometer ...... 275 8.3 Recommendations for Future Work...... 277

REFERENCES...... 281

APPENDIX A: LABORATORY TEST DATA FOR CONVENTIONAL CONCRETE MIXTURES ...... 291

APPENDIX B: LABORATORY TEST DATA FOR SELF- CONSOLIDATING CONCRETE MIXTURES...... 303

x LIST OF TABLES

Table 2.1: Sample Calculation for Effective Annulus Method...... 49 Table 2.2: Recommended Vane Dimensions from Nguyen and Boger (1985) .... 62 Table 2.3: Requirements for Vane for Soil Undrained Shear Strength Measurements (ASTM D 2573)...... 62 Table 3.1: Summary of Factors Influencing Concrete Rheology...... 91 Table 4.1: Variability in Slump Test Measurements in Highway Applications (Baker and McMahon 1969) ...... 96 Table 5.1: Preliminary Design Requirements for ICAR Rheometer ...... 106 Table 5.2: Gap Sizes in Existing Concrete Rheometers...... 109 Table 6.1: Mixture Proportions for Chapter 6 Tests ...... 140 Table 6.2: Qualitative Observations for Mix 1 (5-Inch Slump)...... 144 Table 6.3: Qualitative Observations for Mix 2 (1/2-Inch Slump)...... 146 Table 6.4: Concrete Properties for Series II Impeller Type Tests...... 151 Table 6.5: Average Percentage Reduction in Torque for Each Impeller ...... 153 Table 6.6: Torque after 90 Seconds of Breakdown...... 155 Table 6.7: Average Relative Flow Curve Parameters for Each Impeller...... 157 Table 6.8: Flow Curve Measurements for Mix 1 ...... 158 Table 6.9: Flow Curve Measurements for Mix 2 ...... 159 Table 6.10: Flow Curve Measurements for Mix 3 ...... 160 Table 6.11: Flow Curve Measurements for Mix 4 ...... 161 Table 6.12: Manufacturer’s Recommended Uses for Spiral Impellers...... 165 Table 6.13: Concrete Properties for Series III Impeller Type Tests ...... 166

xi Table 6.14: Observations for Spiral Impellers ...... 167 Table 6.15: Average Percentage Reduction in Torque for Each Spiral Impeller 171 Table 6.16: Average Relative Parameters for Each Impeller (Forward Direction) ...... 172 Table 6.17: Average Relative Parameters for Each Impeller (Mix 4, Reverse Direction) ...... 173 Table 6.18: Concrete Properties for Tests of Optimum Stress Growth Test Speed ...... 175 Table 6.19: Results of Tests for Optimum Stress Growth Test Speed...... 177 Table 6.20: Concrete Properties for Gap Size Tests ...... 179 Table 6.21: Testing of Horizontal Side Gap ...... 180 Table 6.22: Effect of Horizontal Side Gap...... 180 Table 6.23: Effects of Top Cover and Bottom Gap ...... 181 Table 6.24: Concrete Properties for End Effects Testing...... 185 Table 6.25: Test Data for End Effects Testing...... 186 Table 6.26: Determination of End Effects for Stress Growth Test ...... 188 Table 6.27: Yield Stress Values Determined from Method 1 ...... 189 Table 6.28: Yield Stress Values Determined from Method II ...... 190 Table 6.29: Torque Attributable to Side of Effective Cylinder Based on Method III for 5-Inch Vane...... 191 Table 6.30: Determination of End Effects for Flow Curve Measurements...... 192 Table 6.31: Summary of Results from Chapter 6...... 194 Table 7.1: Chemical and Physical Properties of Cement ...... 196

xii Table 7.2: Aggregate Properties...... 197 Table 7.3: Aggregate Particle Size Distributions ...... 198 Table 7.4: Proportions of Base Mixtures ...... 202 Table 7.5: Visual Observations of Fresh Concrete ...... 206 Table 7.6: Coefficients of Determination from Flow Curve Measurements...... 228 Table 7.7: Self-Consolidating Concrete Mixture Proportions ...... 240 Table 7.8: Visual Stability Index Ratings (Daczko 2002) ...... 242 Table 7.9: Mixture Proportions for Field Testing ...... 263 Table 7.10: Field Testing Results...... 265 Table A.1: Fly Ash Test Data...... 292 Table A.2: Ground Granulated Blast Furnace Slag Test Data ...... 293 Table A.3: Silica Fume Test Data ...... 294 Table A.4: Water-to-Cement Ratio Test Data...... 295 Table A.5: Water-Reducing Admixtures Test Data ...... 296 Table A.6: Air-Entraining Agent Test Data ...... 297 Table A.7: Natural Sand/Manufactured Sand Blends Test Data...... 298 Table A.8: Sand-to-Aggregate Ratio Test Data ...... 299 Table A.9: Aggregate Microfines Test Data ...... 300 Table A.10: Slag Aggregate Test Data...... 301 Table B.1: Self-Consolidating Concrete Test Data (Section 1 of 3)...... 304 Table B.2: Self-Consolidating Concrete Test Data (Section 2 of 3)...... 305 Table B.3: Self-Consolidating Concrete Test Data (Section 3 of 3)...... 306

xiii

LIST OF FIGURES

Figure 1.1: Storage of Limestone Microfines at a Quarry ...... 3 Figure 1.2: Apparatus for Slump Flow Test (Ahlers and Walker 1924)...... 7 Figure 2.1: Two-Dimensional Representation of Viscous Flow...... 12 Figure 2.2: Basic Constitutive Relationships for Flow ...... 14 Figure 2.3: Thixotropy ...... 18 Figure 2.4: Schematic of Viscosity Decrease at Low Shear Stresses as Measured in Controlled-Stress Rheometer ...... 24 Figure 2.5: Extrusion Capillary Tube Viscometer ...... 26 Figure 2.6: Generalized Capillary Tube Viscometer...... 26 Figure 2.7: Typical Rotational Rheometer Geometries ...... 30 Figure 2.8: Top View of a Coaxial Cylinders Rheometer...... 32 Figure 2.9: Flow of a Bingham Material in a Coaxial Cylinders Rheometer – No Dead Zone ...... 38 Figure 2.10: Flow of a Bingham Material in a Coaxial Cylinders Rheometer – Dead Zone Present ...... 39 Figure 2.11: Minimum Rotation Speed to Eliminate Dead Zone ...... 41 Figure 2.12: Influence of Dead Zone on Measured Torque versus Rotation Speed Curve ...... 42 Figure 2.13: Influence of Ratio of Yield Stress to Plastic Viscosity on Errors due to Neglecting Dead Zone (Rotation Speed = 10 rpm to 60 rpm) ...... 43

xiv Figure 2.14: Influence of Radii Ratio on Errors Due to Dead Zone (Rotation Speed = 10 rpm to 60 rpm)...... 44 Figure 2.15: Sample Calculation for Effective Annulus Method...... 50 Figure 2.16: Typical Relative Rheometer Results for Concrete ...... 56 Figure 2.17: Wall Effect...... 57 Figure 2.18: Typical Results for a Stress-Controlled Stress Growth Test ...... 58 Figure 2.19: Typical Results for a Rate-Controlled Stress Growth Test ...... 59 Figure 2.20: Typical Vane Impeller ...... 60 Figure 2.21: Yield Surface Superimposed on Finite Element Mesh for a Four- Bladed Vane (Keentok, Milthorpe, and O’Donovan 1985) ...... 64 Figure 2.22: Schematic Representation of Yielding Process in a Stress Growth Test (Yan and James 1997) ...... 65 Figure 2.23: Effect of Rotation Speed on Yield Stress (Saak, Jennings, and Shah 2001)...... 69 Figure 2.24: Finite Element Solution Streamlines for Narrow Gap Vane-in- Cup Rheometer (Barnes and Carnali 1990) ...... 71 Figure 2.25: Shear Stress Distribution in Annulus Based on Finite Element Analysis (Barnes and Carnali 1990)...... 72 Figure 2.26: Workability Box for a Specific Application...... 76 Figure 5.1: Mk I Apparatus (Tattersall and Banfill 1983) ...... 112 Figure 5.2: Square Anchor for Mk I Apparatus (Tattersall and Banfill 1983).... 113 Figure 5.3: Mk II Apparatus and Interrupted Helix Impeller (Tattersall and Banfill 1983) ...... 115

xv Figure 5.4: Two-Blade Impeller (Tattersall and Banfill 1983) ...... 115 Figure 5.5: H-Shaped Impeller for the Mk III Apparatus (Tattersall and Banfill 1983)...... 116 Figure 5.6: IBB Rheometer and Impeller...... 117 Figure 5.7: FCT 101 (from Product Literature) ...... 118 Figure 5.8: “Egg Beater” Mixing Impeller for Mortar...... 119 Figure 5.9: Drum Mixer for Mortar, Plaster, and Granulates ...... 120 Figure 5.10: Mixing Impeller for Drywall Joint Compound, Drywall Texture, and Paint...... 120 Figure 5.11: First Generation Prototype of ICAR Rheometer ...... 124 Figure 5.12: ICAR Rheometer Mounted in Frame and Positioned above Container ...... 126 Figure 5.13: Positioning Rod in Channel...... 127 Figure 5.14: Rendering of Prototype and Envisioned Mass-Production Version128 Figure 5.15: Graphical User Interface for the ICAR Rheometer Software...... 130 Figure 5.16: Typical Test Summary File ...... 134 Figure 5.17: Typical Speed versus Time Plot for Flow Curve Measurement in Concrete ...... 137 Figure 6.1: Potential Impellers...... 142 Figure 6.2: Flow of Concrete around Impellers – Mix 1 (5-Inch Slump)...... 145 Figure 6.3: Flow of Concrete around Impellers – Mix 2 (1/2 Inch Slump) ...... 147 Figure 6.4: Flow around Vane in Uncovered Case ...... 148 Figure 6.5: Voids Left by Impellers – Mix 2 ...... 150

xvi Figure 6.6: Typical Reduction in Torque over Time (Rotation Speed = 40 rpm) ...... 152 Figure 6.7: Percentage Reduction in Torque for Each Mixture ...... 154 Figure 6.8: Torque after 90 Seconds of Breakdown ...... 155 Figure 6.9: Average R2 Terms for Each Impeller, 16-Inch Container ...... 162 Figure 6.10: Spiral Impellers...... 164 Figure 6.11: Intended Flow Direction for Spiral Impellers (From Manufacturer’s Packaging) ...... 165 Figure 6.12: Flow of Concrete around Spiral Impellers – Mix 1 (5-Inch Slump) ...... 168 Figure 6.13: Flow of Concrete around Spiral Impellers Operated in Reverse – Mix 4 (3.5-Inch Slump)...... 169 Figure 6.14: Typical Stress Growth Plots at Various Rotation Speeds...... 176 Figure 6.15: Results of Tests for Optimum Stress Growth Test Speed ...... 178 Figure 6.16: Effect of Top Cover at 0.025 rev/sec (Bottom Gap = 5 Inches)..... 182 Figure 6.17: Effect of Bottom Gap at 0.025 rev/sec (Top Cover = 5 Inches)..... 183 Figure 6.18: Effect of Top Cover at 0.667 rev/sec (Bottom Gap = 5 Inches)..... 184 Figure 6.19: Effect of Bottom Gap at 0.667 rev/sec (Top Cover = 5 Inches)..... 184 Figure 6.20: Determination of End Effects: 0.025 rev/sec, 16-Inch Container .. 187 Figure 6.21: Determination of End Effects: 0.025 rev/sec, 10-Inch Container .. 188 Figure 6.22: Determination of End Effects: 0.667 rev/sec, 16-Inch Container .. 192 Figure 6.23: Determination of End Effects: 0.667 rev/sec, 10-Inch Container .. 193 Figure 7.1: Aggregate Particle Size Distributions...... 198

xvii Figure 7.2: Amount of Microfines Remaining in Each Size Fraction after Standard Sieving Operation for Limestone and Granite Manufactured Sands...... 201 Figure 7.3: ICAR Rheometer Dimensions ...... 205 Figure 7.4: Influence of Fly Ash on Rheology...... 208 Figure 7.5: Influence of GGBFS on Rheology ...... 210 Figure 7.6: Influence of Silica Fume on Rheology ...... 212 Figure 7.7: Influence of Water-to-Cement Ratio on Rheology...... 214 Figure 7.8: Influence of Water Reducers on Rheology...... 216 Figure 7.9: Influence of Air Content on Rheology ...... 218 Figure 7.10: Influence of Blended Sands on Rheology ...... 220 Figure 7.11: Influence of Sand-to-Aggregate Ratio of Rheology...... 222 Figure 7.12: Influence of Microfines Content on Rheology ...... 224 Figure 7.13: Influence of Changing Sand Particle Size Distribution in High- Microfines Mixtures...... 225 Figure 7.14: Influence of Slag Aggregate ...... 227 Figure 7.15: Comparison of Rheometer Results from First and Second Tests... 230 Figure 7.16: Comparison of Yield Value and Yield Stress (First Test)...... 232 Figure 7.17: Comparison of Viscosity Value and Plastic Viscosity (First Test) 233 Figure 7.18: Comparison of Yield Torque and Yield Stress (First Test)...... 234 Figure 7.19: Yield Stress and Plastic Viscosity Measurements for All Conventional Concrete Tested ...... 235 Figure 7.20: Workability Boxes for Segregation Resistance Rating...... 236

xviii Figure 7.21: Workability Boxes for Overall Workability Rating ...... 237 Figure 7.22: Workability Boxes for Average of All Ratings ...... 237 Figure 7.23: Influence of Water-to-Cement Ratio on SCC Rheology ...... 243 Figure 7.24: Influence of Water-to-Cement Ratio on Slump Flow...... 244 Figure 7.25: Influence of Cement Content on SCC Rheology...... 245 Figure 7.26: Influence of Cement Content on Slump Flow ...... 246 Figure 7.27: Influence of Fly Ash on SCC Rheology...... 247 Figure 7.28: Influence of Fly Ash on Slump Flow ...... 248 Figure 7.29: Influence of Sand Type on SCC Rheology...... 249 Figure 7.30: Influence of Sand Type on Slump Flow...... 250 Figure 7.31: Yield Stress and Plastic Viscosity from ICAR Rheometer for All SCC Measurements ...... 251 Figure 7.32: Slump Flow Test for Excellent SCC Mixture (Mix 5, Test 4,

Slump Flow = 27 Inches, VSI = 0, T50 = 5 sec) ...... 252 Figure 7.33: No Segregation Evident in Slump Flow Test (Mix 5, Test 4, Slump Flow = 27 Inches, VSI=0) ...... 253 Figure 7.34: Slight Mortar Halo Evident in Slump Flow Test (Mix 8, Test 2, Slump Flow = 30.5 Inches, VSI=2)...... 253 Figure 7.35: Severe Segregation (Mix 13, Test 1, Slump Flow = 31 Inches, VSI = 3)...... 254 Figure 7.36: Relationships between Slump Flow and Rheological Parameters.. 255

Figure 7.37: Relationships between T50 and Rheological Parameters ...... 256 Figure 7.38: Comparison of Parameters from ICAR Rheometer...... 258

xix Figure 7.39: Evidence of Flow throughout ICAR Rheometer Container for SCC Mix...... 258 Figure 7.40: Comparison of Rheological Parameters from BTRHEOM and ICAR Rheometer...... 259 Figure 7.41: Evolution of Yield Stress and Plastic Viscosity over Time (Mix 10)...... 261 Figure 7.42: Evolution of Yield Stress and Plastic Viscosity over Time (Mix 12)...... 261 Figure 7.43: Field Testing with Rheometer ...... 264 Figure 7.44: Results of Field Testing– Series I...... 265 Figure 7.45: Results of Field Testing – Series II...... 266 Figure 7.46: Zero Slump Concrete Tested with Vane Impeller (Concrete Filled to One Inch below Top of Vane)...... 267

CHAPTER 1: INTRODUCTION

1.1 RESEARCH BACKGROUND The workability of fresh concrete—that is, the ease with which it can be transported, placed, consolidated, and finished—is a critical property that has a direct impact on the strength, durability, appearance, and cost of concrete. For more than eighty years, the workability of fresh concrete has been measured predominately throughout the world with one simplistic test method—the slump test. In the slump test, a sample of fresh concrete is placed in a 12-inch tall cone mold. The mold is removed and the vertical distance the concrete subsides, or slumps, is recorded as a measure of workability. Whereas workability is a broadly defined term, the slump test measures only one aspect of workability, namely, consistency. Other aspects of workability are commonly described in subjective, qualitative terms. Although the introduction of the slump test as an ASTM standard test method in 1922 represented an important advance in the design and control of concrete mixtures, the slump test is now viewed as incapable of providing an adequate characterization of the workability of today’s much more advanced concrete mixtures. Modern concrete production systems have not eliminated the need to monitor concrete workability in the field. To the contrary, the advent of new so-called high-performance concrete mixes, which can incorporate a wide array of different materials and can be susceptible to small changes in mixture proportions, has made the monitoring of workability even more critical. Concrete mixtures that include certain materials—such as aggregate microfines, fiber reinforcement, ground granulated blast furnace slag, and new classes of chemical

1 admixtures—cannot be characterized properly with the slump test. Such materials can have profound effects on workability, yet the slump test often produces false or misleading results for concretes made with these materials. The lack of a viable concrete workability test method has engendered reluctance in the concrete industry in specifying non-standard materials. For instance, the use of high contents of aggregates microfines are prohibited in current specifications, such as ASTM C 33, in part because of concerns about workability. ASTM C 33 was developed based on the use of natural river sand in the production of concrete. As the availability of natural river sand has decreased, the use of manufactured sand has become more important. Manufactured sand is produced extracting rock from a quarry and crushing it down to the size of sand. This crushing process creates dust-of-fracture microfines, which by definition have a maximum particle size of 75 µm. Although research has shown that manufactured sands with up to 20% dust-of-fracture microfines by mass of the fine aggregate can be used to make high-quality concrete (Ahn and Fowler 2001; Quiroga 2003), the workability of such concrete can be much different than concrete made with natural river sands. Specifically, concrete with a high microfines content may appear to be stiff and unworkable when observed in a static state (such as with the slump test); however, once flow is initiated in the concrete, the workability is much better. In order to meet the current ASTM C 33 specification, aggregates producers must wash their crushed aggregates to remove microfines. The microfines must be stored, as shown in Figure 1.1, and then sold or disposed. Although limited uses do exist for microfines, much of the material becomes a waste product. If higher percentages of microfines were permitted in concrete, the process of washing and storing aggregate microfines could be sharply reduced, or even eliminated. Therefore, the increased utilization of high- microfines concrete, enabled by an enhanced test method for concrete

2 workability, would reduce aggregate production costs and create a new market for a current waste product.

Figure 1.1: Storage of Limestone Microfines at a Quarry

The slump test is also ineffective for self-consolidating concrete (SCC), an advanced type of highly flowable concrete that can exhibit superior hardened properties and result in significantly reduced labor costs. SCC is engineered to flow readily under its own weight through the incorporation of chemical admixtures and the utilization of improved particle size distributions. In order for SCC to be used successfully, its workability must be carefully monitored and controlled. A robust field test method for concrete workability is clearly needed. Yet the challenges of developing a new test method that is simple, relevant, and accurate has proven daunting. A myriad of test procedures for determining workability has been developed for research, mixture proportioning, and field use; however, the vast majority of these test methods has never found any use beyond one or two initial studies. The concrete industry has made strides in applying the concepts of rheology—that is, the scientific study of the flow and deformation of matter—to the workability of fresh concrete. By providing a scientific description

3 of the fundamental flow properties of cement paste, mortar, and concrete, rheology represents a useful method of characterizing concrete workability. Previous attempts to use rotational rheometers—which measure the amount of stress needed to generate a range of flow rates in a fluid—for concrete have proven promising. The existing rheometers that have been built, however, are too large and expensive for routine use, especially in the field. Even with the increase in knowledge of concrete rheology, the slump test remains the most commonly used test method for measuring concrete workability. The need for improved workability testing technology is well established. According to the “Roadmap 2030” study for the US concrete industry, which was facilitated by the Strategic Development Council of the American Concrete Institute, the development of “portable testing technologies for use at the jobsite” is a “high priority research need” (SDC 2002, 23). The roadmap specifically states that “field tests are needed to measure fundamental rheological properties of fresh concrete” (SDC 2002, 23). A separate survey conducted by the National Ready Mixed Concrete Association identified the need for a better method to characterize the workability of high-performance concrete (Ferraris and Lobo 1998). A viable, accurate, and relevant concrete workability test method would foster an improved understanding of concrete workability, reduce concrete construction costs, improve hardened concrete quality, and promote the use of new and underutilized materials.

1.2 HISTORICAL PERSPECTIVE Workability has always been a key characteristic of fresh concrete. As such, considerable effort has been expended to develop rapid, relevant, and robust test methods for characterizing workability. A brief but instructive look at the history of workability measurement helps to establish an understanding of the

4 difficulty of measuring concrete workability and to place current rheology-based test methods in proper perspective. The American concrete industry at the beginning of the 1900s had no standard test methods to measure workability. Instead, subjective, qualitative descriptions of “consistency” were typically given. While recognizing that concrete consistency was of “utmost importance,” Taylor and Thompson (1922, 250) divided concrete consistency into three simplistic and vague categories: “dry” consistency, “medium or quaking” consistency, and “wet or mushy” consistency. Wig (1912) presented research that was representative of the state of the art of the time. In his study, three different concrete contractors were given identical quantities of cement and aggregates and instructed to prepare concrete mixes as they normally would. Beams cast of the resulting concrete differed in strength by more than 200% for the hand-mixed batches and 90% for the machine-mixed batches. Summarizing the results, Wig conceded that the water content “perhaps had the greatest effect upon the strength [emphasis added]” but stressed the importance of mixing and handling. This lack in understanding of workability and of the importance of water content on concrete properties severely impaired concrete quality. The concrete industry took a dramatic step forward in 1918 when Duff Abrams, a leading concrete researcher at the time, published his seminal Design of Concrete Mixtures, in which he showed unequivocally that concrete strength was directly related to the ratio of water-to-cement—what Abrams called the water ratio. Whereas other mixture proportioning procedures of that time focused mainly on achieving an optimum packing of aggregates and considered the water content to be subordinate, Abrams showed that the water-to-cement ratio was the most important parameter in developing mixture proportions and that it should be set as low as possible on the condition that proper workability could be achieved. To define workability, Abrams suggested determining a “relative consistency”

5 term calculated by measuring the distance fresh concrete slumped after being molded in a 6-inch by 12-inch cylinder—the slump test. Although it may be difficult to appreciate today, the concrete industry of the 1920s truly did not understand the importance of water content. Abrams would later write, “I believe that the big thing [the slump] test can do is to bring home the importance of water control in building construction” (Abrams 1922, 178). The slump test was adopted quickly. In a survey of 40 state highway departments conducted in 1918 and 1919, only three states had requirements for slump while all of the other states only had general provisions. By 1924, 25 of 47 state highway departments surveyed had explicit requirements for the slump of concrete used in concrete road building (Johnson 1924). In comparison, 39 of 50 state departments of transportation had specific requirements for slump for concrete pavements in 2004 (ACPA 2004). Although the slump test was quickly accepted due to its simplicity, the concrete industry immediately recognized that the slump test was inadequate for fully and properly characterizing workability. From the 1920s to the present, a profusion of test methods has been introduced to improve workability characterization. The majority of tests developed since the 1920s are empirical— that is, they attempt to simulate a field placement condition and measure a value, such as time or distance, which serves as an index of workability. One early empirical test, pictured in Figure 1.2, consisted of measuring the horizontal spread of a concrete specimen formed in a slump cone and subjected to jolting on a drop table. Attempts have also been made to measure the fundamental properties of concrete instead of empirical values. Fundamental parameters provide a standard, scientific description of concrete that is independent of the test device used. Indeed, devices introduced as early as the 1920s attempted to measure concrete viscosity, although the thinking behind such devices was inchoate (Smith and Conahey 1928; Bates and Dwyer 1928). Prior to World War II, several

6 unsuccessful attempts were made to develop rheometers to measure concrete (Powers 1968). Not until the 1970s, though, were rheometers developed that could successfully provide reasonably accurate measurements of concrete (Tattersall and Banfill 1983). Many of the concrete rheometers in use today are based on the work done in the 1970s. Through the use of these devices, concrete rheology has emerged as a viable technique for characterizing the workability of fresh cement paste, mortar, and concrete. Despite the advancements made in concrete rheology, much more work remains to be done in order to allow rheology to be used on a widespread basis.

Figure 1.2: Apparatus for Slump Flow Test (Ahlers and Walker 1924)

While the concrete industry has changed significantly over the last eight decades, the slump test has remained unchanged. The slump test was developed at a time when water content was directly related to both strength and workability. By giving an indication of the water content for a given mixture, the slump test also provided a means of ensuring adequate hardened strength. With the wide

7 array of materials available for use in modern concrete, the relationships between water content and both strength and workability are no longer as simple as they once were. While increasing the water content typically increases fluidity and reduces strength, increasing fluidity does not necessarily reduce strength. This fact demonstrates the limited usefulness of the slump test. In particular, the slump test should not be used to define workability in absolute terms or to predict strength. Instead, it is best suited as a quality control device. According to Kosmatka, Kerkhoff, and Panarese (2002), the slump test should only be used to compare concretes with similar mixture proportions in order to detect changes in the water content, characteristics of materials, mixture proportions, or mixing operations. In reality, the concrete slump test is used to measure much more, despite the clear limitations of the test. While the value of ready-mixed concrete produced annually in the United States now exceeds $100 billion (SDC 2002), the industry relies on a simple test, for which the equipment cost less than $30, to attempt to ensure adequate workability. The direct and indirect costs of poor concrete workability, however, can be significant and point to the clear need for an improved technique for characterizing workability.

1.3 RESEARCH OBJECTIVES The purpose of the research described in this report is to identify an effective field test method for measuring the workability of concrete in general and of high-microfines concrete in particular. Such a test method should be robust, portable, relevant to current concrete materials, low-cost, and accurate. The test method can be an existing test method, a modification to an existing test method, or an entirely new test method.

8 1.4 PROJECT SCOPE To accomplish the research objectives, the research project was divided into three parts. First, a literature review was conducted to assess the current state of the art and to determine needs for future workability test methods. The literature review focused on fluid rheology, concrete workability, and existing concrete workability test methods. Feedback was sought from industry, government, and academia. Second, the ICAR rheometer, a new portable test device, was developed based on the requirements determined in the first stage of the project. Finally, the ICAR rheometer was tested on a series of concrete mixtures covering nearly the full range of concrete workability in order to verify its ability to characterize workability adequately and accurately. This report is split into eight chapters, including the introduction. The results of the literature review are presented in Chapters 2 and 3, with Chapter 2 covering fluid rheology and Chapter 3 describing factors influencing rheology and workability. An associated summary of existing workability test methods was published separately as ICAR Report 105.1 (Koehler and Fowler 2003). Chapter 4 provides an evaluation of the lessons learned from the literature review and identifies the ICAR rheometer as the most promising approach for workability characterization. The development of the ICAR rheometer is described in detail in Chapter 5. The initial testing conducted to determine certain operating characteristics of the ICAR rheometer is described in Chapter 6. Chapter 7 describes the series of tests conducted to validate the ability of the ICAR rheometer to provide an adequate and accurate description of workability. Chapter 8 presents the summary and conclusions for the research project and list recommendations for future work.

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CHAPTER 2: FLUID RHEOLOGY

2.1 INTRODUCTION Fluid rheology is a well-established, widely used science that is directly applicable to the workability of fresh concrete. Although fresh concrete can be considered a fluid, the characterization of the rheology of fresh concrete is complicated by the fact that concrete is a complex material with time-dependent properties and a wide range of particle sizes. Essentially, fresh concrete is a concentrated suspension of aggregates in cement paste. The cement paste itself is a concentrated suspension of cement grains in water. Fortunately, the rheology literature is replete with experiences in measuring the rheology of concentrated suspensions. The lessons learned by others in characterizing materials similar to concrete are highly relevant to the application of rheology to fresh concrete workability. This chapter describes the basic concepts of fluid rheology germane to fresh concrete, including basic fluid properties and the techniques available for measuring these properties.

2.2 PROPERTIES OF FLUIDS Although rheology is concerned with the flow and deformation of matter—including liquids, solids, and gasses—the term rheology is mainly used to refer to the study of liquids. Topics such as elasticity are not typically considered in the formal study of rheology.

11 2.2.1 Definition of a Liquid To define a liquid, it is useful to make a distinction between an elastic solid and a viscous liquid. If a constant stress is applied to an elastic solid, the material will undergo a finite deformation; this deformation will be fully recovered upon removal of the load. For a linear-elastic solid material subjected to an applied stress, σ , and undergoing a strain, ε , the constitutive equation is given in Equation (2.1). The coefficient of proportionality relating stress and strain is the modulus of elasticity, E.

σ = Eε (2.1)

In terms of shear properties, the shear stress, τ , on an elastic solid is related to the shear strain, γ , by the shear modulus, G:

τ = Gγ (2.2)

In contrast to an elastic solid, a viscous liquid deforms continuously due to an applied shear stress for as long as the shear stress is applied; this deformation will not be recovered once the load is removed. The two-dimensional case for the flow of a liquid between two parallel plates of sufficient length such that end effects can be ignored is represented in Figure 2.1.

F v y

Figure 2.1: Two-Dimensional Representation of Viscous Flow 12

In viscous flow, the shear stress and the time rate at which shear stress is applied are related—the faster the fluid is sheared the greater the shear stress that is required. For the case of constant flow, the shear stress, τ , is related to the shear rate, γ& , by the coefficient of viscosity, η :

τ = ηγ& (2.3)

In the case shown in Figure 2.1, the shear rate is equal to the velocity gradient:

dv γ& = (2.4) dy

As will be shown later in Section 2.3.2.2, Equation (2.4) does not hold true for all cases, such as for circular flow in a coaxial cylinders rheometer. The distinction between a solid and a liquid is not as clear as the above distinction would make it seem. Certain inelastic solid materials can undergo irrecoverable deformations over a certain range of strains. Likewise, viscous fluids may behave as elastic solids at very low shear strains. A viscoelastic material exhibits both a viscous and elastic response under a constant stress. Thus, from a rheological standpoint, it is not important to distinguish between solids and liquids in a formal sense but to consider the relevant type of behavior under a particular loading of a given material. Indeed, Whorlow (1992, 9) states that a “formal distinction between solids and liquids…is of little practical value.” Further, it has been argued that given sufficient time, all materials flow (Barnes 1999).

13 2.2.2 Constitutive Equations for Fluid Flow Equation (2.3) represents one combination of shear stress and shear rate during steady flow; however, in measuring the rheology of a material it is important to know the flow properties over a range of shear stresses and shear rates. The relationship between shear stress and shear rate is represented graphically in a flow curve. Fluids may be distinguished by their flow curves. Various models—or constitutive equations—have been developed to idealize flow curves. Six of the most common constitutive relationships associated with concrete are plotted in Figure 2.2.

Casson

Herschel-Bulkley

Bingham Power Law (Shear Thickening) Newtonian Shear Stress Shear

Power-Law (Shear Thinning)

Shear Rate

Figure 2.2: Basic Constitutive Relationships for Flow

The most basic constitutive equation is for a Newtonian fluid, where the linear relationship between shear stress and shear rate—given above as Equation

14 (2.3)—applies for the entire range of shear rates. As such, viscosity is a material constant:

τ = ηγ&

Although the Newtonian model is a simple equation that represents basic fluid flow, it fails to represent adequately the flow response of most fluids, including many concrete mixtures. Specifically, the Newtonian model assumes that the relationship between shear stress and shear rate is linear and that the flow curve intercepts the shear stress axis at the origin. In reality, most fluids do not behave linearly. Many fluids possess some minimum stress—namely, a yield stress—that must be exceeded before flow occurs. The concept of a yield stress is readily seen in the concrete slump test. When the slump cone is first removed, the stress induced by gravity is sufficient to exceed the yield stress. As a result, the concrete flows briefly until the height of the concrete is low enough that the stress induced by gravity no longer exceeds the yield stress. Materials that exhibit a yield stress are considered to be viscoplastic materials. The Bingham model incorporates a yield stress term, τ 0 , but maintains a linear relationship between shear stress and shear rate:

τ = τ 0 + µγ& (2.5)

In the Bingham equation, viscosity, η , is often replaced with plastic viscosity, µ . Technically, the terms viscosity and plastic viscosity refer to the same physical relationship despite the fact that different terminology is used. One way to represent nonlinearity is with the power-law model, which still assumes a zero yield stress but represents the shape of the flow curve as an exponential relationship:

b = aγτ & (2.6)

In the above equation, both a and b are material constants. If the exponent, b, is set to less than unity, the flow curve will be concave downward. Such a relationship is known as shear-thinning, or pseudo-plastic, behavior. If b is set to greater than unity, the resulting flow curve will be concave upward, representing shear-thickening behavior. The Herschel-Bulkley model essentially combines the Bingham equation and power-law equation to represent both a yield stress and a nonlinear flow relationship:

b 0 += aγττ & (2.7)

Setting the exponent, b, equal to unity simply results in the Bingham equation. In concrete, the Hershcel-Bulkley equation may be either concave upward or downward over the range of shear rates commonly tested. Another way to represent nonlinearity of flow in fluids with a yield stress is given by the Casson model:

21 21 2121 0 += γµττ & (2.8)

Additional listings of constitutive equations for particulate systems are given by Hackley and Ferraris (2001) and by Whorlow (1992). Although any of the above constitutive equations can potentially be used for concrete, the Bingham equation is used most commonly because of its accuracy in representing most concrete mixtures and its simplicity in only requiring the determination of two parameters, yield stress and plastic viscosity. Even for fluids that exhibit nonlinear flow behavior, the linear Bingham equation

16 may be an acceptable equation over a narrow range of low shear rates. Concrete mixtures subjected to vibration behave as Newtonian fluids for low shear rates (Tattersall 1991). Self-consolidating concretes have yield stress values near zero, and thus approach the behavior of Newtonian or power-law fluids. Whorlow (1992) has identified three drawbacks of the Bingham model for measuring fluids in general. First, flow curves are rarely linear in reality. Second, flow hysteresis often occurs when measuring ascending and descending flow curves. Third, the yield stress is not a well-defined property.

2.2.3 Thixotropy and Anti-Thixotropy The measured flow curve often depends on the shear history of the sample. A thixotropic material experiences a reversible, time-dependent decrease in viscosity when subjected to constant shearing whereas an anti-thixotropic, or rheopeptic, material experiences an increase in viscosity (Hackley and Ferraris 2001). Depending on the material, this time-dependence of flow properties can influence readings for time periods ranging from seconds to days. Figure 2.3 depicts the effect of thixotropy for a test where the shear rate is gradually increased from zero to a maximum value and then decreased back to zero. For all thixotropic materials, the up-curve will be above the down-curve. Conversely, the down-curve will be above the up-curve for anti-thixotropic materials. The area between the up- and down-curves depends in part on the material and the amount of time for the reading of each shear rate point—longer readings allow for a more pronounced thixotropic effect—and should not be taken as a representation of the degree of thixotropy (Whorlow 1992).

Shear Stress

Shear Rate Figure 2.3: Thixotropy

2.2.4 Dilatancy Dilatancy is the increase in volume of a fluid during shearing. In concentrated suspensions, particles sliding past each other lead to such an expansion in volume. Dilatancy should not be confused with shear-thickening behavior (Whorlow 1992; Hackley and Ferraris 2001).

2.2.5 Viscosity

2.2.5.1 Definition Viscosity was defined previously as the coefficient of proportionally relating shear stress and shear rate under a state of steady shear. For Newtonian fluids, this viscosity is a material constant that applies for all shear rates. If the viscosity varies as a function of shear rate, it is referred to as the non-Newtonian

18 viscosity. A variety of viscosity terms are used in practice. Those related to concrete are described in this section. The viscosity term previously indicated as η and most commonly used in concrete technology is the dynamic viscosity, which is given in SI units of Pa.s or CGS units of poise (1 Pa.s = 10 poise). The kinematic viscosity, v , which is expressed in units of m2/s or stokes (1 m2/s = 10,000 stokes), is the dynamic viscosity divided by density, ρ :

η v = (2.9) ρ

The apparent viscosity is equal to the shear stress divided by the shear rate for cases where this quotient is dependent on the shear rate. For nonlinear flow curves, the apparent viscosity is simply the slope of a line drawn from the origin to a point on a flow curve.

The differential viscosity, ηdiff , is defined as the derivative of shear stress with respect to shear rate:

∂τ η diff = (2.10) ∂γ&

The plastic viscosity, µ , is defined as the limit of the differential viscosity as the shear rate approaches infinity:

∂τ pl µη == lim (2.11) γ& ∞→ ∂γ&

For Bingham materials, the plastic viscosity is equal to the differential viscosity for all shear rates. For non-ideal Bingham materials that are nonlinear at 19 low shear rates but that reach a linear portion at high shear rates, the plastic viscosity is the slope of this linear portion of the curve.

For non-Bingham materials, the infinite shear viscosity, η∞ , can be used to represent the slope of the flow curve at high shear rates. For instance, power- law fluids may be idealized as two separate linear portions over relevant shear ranges. In such a case, the slope of the second linear region may be considered as the infinite shear viscosity.

The zero shear viscosity, η0 , is the limit of the differential viscosity as the shear rate approaches zero:

∂τ η0 = lim (2.12) γ&→0 ∂γ&

For nonlinear flow curves, the zero shear viscosity can be used to represent the linearized first-Newtonian region for a narrow range of shear rates.

The relative viscosity, ηr , is given as the ratio of the viscosity of a suspension or solution to the viscosity of the suspending medium of solvent:

η η r = (2.13) η s

The specific viscosity, η sp , is then defined as the relative viscosity minus one:

η = ηrsp −1 (2.14)

20 Fluidity, φ , is the reciprocal of viscosity:

1 φ = (2.15) η

2.2.5.2 Origin of Viscosity in Fluid Suspensions Three types of forces act on particles in fluid suspensions: colloidal particle interaction forces, Brownian forces, and viscous forces (Barnes et al. 1989). Each of these forces contributes to viscosity. Particle interaction forces are the result of attraction or repulsion between particles. Attractive forces may be due to van der Waals forces or unlike electrostatic charges while repulsive forces may be caused by steric repulsion or like electrostatic charges. These forces exist simultaneously and can result in complex behavior. Attractive forces tend to cause flocculated structures, which are broken down by shearing and rebuilt once shearing is ceased. As a result, flocculated systems typically are thixotropic and exhibit a yield stress. The flocs trap the suspending fluid and increase the apparent solids volume concentration. The viscosity of a flocculated structure generally increases with the degree of flocculation. In contrast, repulsive forces tend to result in a dispersed structure. In dispersed systems, the viscosity is generally related to the solids concentration but not to the degree of dispersion. Brownian forces result in the randomization of the radial distribution and spatial orientation of particles. Since Brownian forces are strongly size- dependent, with the biggest effects on particles smaller than 1 µm, their influence on concrete is minimal. Viscous forces are proportional to the local velocity difference between a given particle and the surrounding fluid. Therefore, a change in the viscosity of

21 the suspending medium results in a change in the viscosity of the overall suspension. Multiple models have been developed to relate suspension properties to viscosity. The Einstein model, developed in 1911, relates the viscosity of a

suspension, η , to the viscosity of the suspending medium, η s , and the solids volume concentration, φ :

η = η s + φ)5.21( (2.16)

The model is based on spherical particles with uniform size. Higher order terms of φ can be added to account for particle interaction. Due to the simplicity of the Einstein model, it is best suited for suspensions with solids volume concentrations of less than 10% (Barnes 1989). Modifications to the Einstein model must be made for concentrated suspensions. The Krieger-Dougherty equation, which is suitable for concentrated suspensions, is based on the Einstein

equation and incorporates the maximum packing fraction (φm ) and the intrinsic viscosity ( η][ ), as shown in Equation (2.17):

φ − ][ φη m ηη s −= )1( (2.17) φm

The maximum packing fraction is defined as the solids volume concentration at which the particle concentration results in three-dimensional contact throughout the suspension and the viscosity approaches infinity. The maximum packing fraction is a function of the type of packing, the particle shape, and the particle size distribution. Intrinsic viscosity is a dimensionless number defined by Barnes et al. (1989) as the limiting value of the reduced viscosity as the solids volume concentration approaches zero. The intrinsic viscosity equals

22 2.5 for spheres and higher values for other shapes. The values of maximum packing fraction and intrinsic viscosity used in the model can be varied at different shear stresses for a single shear thinning material. Although the Krieger-Dougherty model has been applied to concrete (Szecsy 1997), Ferraris (1999) states that the Einstein and Krieger-Dougherty models can be used for cement paste but should not be used for concrete.

2.2.6 Yield Stress Yield stress commonly occurs in multiphase fluids such as concentrated suspensions. The solid particles interact to form a flocculated, three-dimensional network structure that resists flow at sufficiently low stresses. The yield stress is related to the force required to break down this structure and initiate flow. The questions of how yield stress should be defined and of whether a yield stress even exists have been debated widely. An extensive history of the debate is provided by Barnes (1999). The question of whether a yield stress exists is based on the concept that all materials flow if given sufficient time. The yield stress determined from a flow curve is based on an extrapolation of the measured points back to the shear stress at a zero shear rate. Flow models—such as the Bingham or Herschel-Bulkley model—commonly assume that if lower shear rates could be measured, the flow curve would continue back to and intercept the shear stress axis. Accurate data at low shear rates are needed to confirm that flow curves can be extrapolated to a zero shear rate. The accuracy of such low-shear measurements is reduced by slip, fracture, and expulsion of the sample. The advent of modern controlled-stress rheometers has allowed measurements to be made at significantly lower shear rates than previously possible. These rheometers have shown that at low shear stresses, the viscosity is very high. When the shear stress is increased to a certain point, the viscosity begins to decrease rapidly, as depicted in Figure 2.4. This transition from high viscosity to low is related to what 23 conventionally has been thought of as the transition from elastic behavior to viscous flow. Despite this new information, some researchers still suggest that the yield stress exists in certain cases and can be verified.

Viscosity (log)

Shear Stress (log)

Figure 2.4: Schematic of Viscosity Decrease at Low Shear Stresses as Measured in Controlled-Stress Rheometer

Whether it truly exists or not, yield stress is a term with practical significance. Nguyen and Boger (1983) state that yield stress is a model parameter but not a true material property. As Whorlow (1992) points out, the question of the existence of yield stress may be important in certain applications; however, for most practical applications, the existence of a true yield stress does not need to be questioned. Indeed, in many practical cases, the value of the yield stress— however it is defined—must be known. Liddell and Boger (1996, 236), in summarizing the yield stress debate, conclude that “a knowledge of the yield stress and its measurement is essential in industry.” What matters, then, is the definition of yield stress. If it is assumed that the yield stress is a relevant parameter to be measured, the actual measured value can vary significantly depending on the test method used. Static measurements of yield stress—measured when the material is initially

24 at rest—are typically higher than the measurements of yield stress obtained from dynamic measurements of flow curves (Whorlow 1992). This discrepancy further complicates the discussion of the meaning of the yield stress.

2.3 MEASUREMENT OF RHEOLOGY Rheological properties can be measured in capillary tube viscometers or rotational rheometers. In concrete, rotational rheometers are used predominately in cases where the rheological parameters are to be determined in fundamental units while variations on capillary tube viscometers are used in limited cases. According to Hackley and Ferraris (2001), rotational methods are generally better for concentrated suspensions, gels, and pastes despite the fact that capillary tube methods tend to be more precise in measuring viscosity. Rotational methods offer the advantage of being able to shear a sample indefinitely in order to achieve equilibrium and to monitor changes over time. For non-Newtonian fluids, the distribution of shear rate and shear stress is typically better defined in a rotational device than a capillary tube device. The problem of temperature rise due to shearing can be more of a problem in a rotational rheometer, although methods are available to limit temperature change. Capillary tube viscometers are typically cheaper and simpler than rotational rheometers (Whorlow 1992).

2.3.1 Capillary Tube Viscometers

2.3.1.1 Description Capillary tube viscometers can be used to measure Newtonian and non- Newtonian fluids. A variety of viscometer types are available; however, since viscometers are not widely used for concrete, only a brief description will be given here. For concrete, capillary tube viscometers or variations on the concepts

25 of capillary tube viscometers are typically used to determine the viscosity of highly fluid concrete or to assess pumpability (Tattersall and Banfill 1983). One common type of capillary tube viscometer is an extrusion capillary tube viscometer, such as the one depicted schematically in Figure 2.5. In this arrangement, a piston at one end forces material to flow through a pipe and out the other end. By measuring the force applied by the piston and the flow rate of material out of the pipe, viscosity can be determined.

Q F 2a L

Figure 2.5: Extrusion Capillary Tube Viscometer

2.3.1.2 Derivation of Equations for Capillary Tube Viscometers The derivation of equations for a capillary tube viscometer is based on a tube of finite length, L, and radius, a, shown in Figure 2.6. It is assumed that all flow occurs parallel to the longitudinal axis of the tube. Provided no slippage occurs at the walls, the fluid velocity varies as a function of radius, with the maximum velocity at the center of the tube and zero velocity at the walls.

2πrLτ a r πr2P

Figure 2.6: Generalized Capillary Tube Viscometer

The force due to a pressure, P, acting at the end of a discrete cylinder of radius r can be set equal to the opposing viscous force due to shear stress,τ , acting along the entire length of the cylinder:

2 = 2 rLPr τππ (2.18)

Rearranging the above equation in terms of shear stress results in the following relationship:

Pr τ = (2.19) 2L

The shear rate in the tube is defined as the derivative of velocity, v, with respect to radius. In this case, a negative sign is added because velocity decreases with increasing radius:

dv γ& −= (2.20) dr

Next, the derivative of velocity with respect to shear stress is determined by using the chain rule:

dv dv dr dr 2L γ −=−== γ&& (2.21) dτ dr dτ dτ P

In order that any constitutive equation can be used, it is desirable to substitute γ& = f ()τ into the above equation to represent a generic constitutive equation:

27 dv 2L = f ()τ (2.22) dτ P

Next, Equation (2.22) can be integrated to determine an expression for velocity:

Pa 0 2L = 2L fdv () dττ (2.23) v ∫∫ τ P

For a Newtonian fluid, the viscosity varies in the tube as a function of radius as given by Equation (2.24):

P rv )( ( −= ra 22 ) (2.24) 4Lη

The flow rate through the tube can be calculated by integrating the product of area and velocity over the radius of the tube:

a = 2π vdrrQ (2.25) ∫0

The result for a Newtonian fluid is the well-known Poiseuille equation:

π 4 Pa Q = (2.26) 8ηL

While it is technically possible to make one pressure measurement at one flow rate for a Newtonian fluid, it is more precise to make a range of measurements and plot pressure versus flow rate. The slope of the resulting line,

28 which passes through the origin, can be used to determine viscosity. It is also possible to keep the pressure constant and change the length of the tube. For a yield stress material, the stress in the center of the tube is equal to zero, as evident in Equation (2.19). Therefore, a region of material in the center of the tube—where the stress is less than the yield stress—passes through the tube as a solid plug. This phenomenon is known as plug flow. Because the shear stress is always zero at the center of the tube, plug flow can never be completely eliminated for materials with a yield stress. For a Bingham material, the Buckingham-Reiner equation accounts for the presence of plug flow:

πa 4 ⎛ 8 L 16(ττ L)4 ⎞ Q ⎜ P 0 +−= 0 ⎟ (2.27) ⎜ 43 ⎟ 8µL ⎝ 3a 3 aP ⎠

By plotting Q versus P, it is possible to determine both yield stress and plastic viscosity.

2.3.2 Rotational Rheometers

2.3.2.1 Description Rotational rheometers are able to apply shear stress to a single sample of material continuously. By measuring a series of combinations of shear stress and shear rate, a flow curve can be determined. It is possible to impose a range of shear rates and determine the resulting shear stresses (controlled-rate rheometer) or to impose a range of shear stresses and measure the resulting shear rates (controlled-stress rheometer). Compared to a controlled-rate rheometer, a controlled-stress rheometer typically has higher sensitivity, particularly at very low shear rates, and can better differentiate between highly non-Newtonian fluids (Schramm 1994).

29 Multiple geometrical configurations of rotational rheometers are available; three main rotational rheometer geometries are shown in Figure 2.7. Numerous variations of each geometry exist; the particular test set-up selected depends on the properties of the material to be tested. In a coaxial cylinders rheometer, the fluid is placed between two cylinders. It is possible either to rotate the outer cylinder at a series of fixed speeds while measuring the resulting torque on the fixed inner cylinder or to keep the outer cylinder fixed while torque and rotation speed are measured at the rotating inner cylinder. In a parallel plate rheometer, the bottom plate is fixed while the top plate rotates. Both torque and rotation speeds are measured at the top blade. The cone and plate rheometer is similar to the parallel plate rheometer, with the exception that a cone is used instead of a top plate. Coaxial cylinders and parallel plate rheometers have been used to measure the rheology of concrete and cement paste. The cone and plate rheometer configuration is less applicable to concrete due to the difficulty of fitting aggregates around the cone. Oscillatory rotational rheometers, which measure the flow properties by rotating in alternating directions at a fixed frequency, have been used in limited instances for concrete and for cement but are not discussed in this report.

Measured Applied/ Applied/ Inner Bob Torque Top Plate Measured Cone Measured Outer Torque Torque Cylinder Fluid Fluid Fluid

Bottom Plate Bottom Plate Applied Torque (Fixed) (Fixed)

Coaxial Cylinders Parallel Plate Cone and Plate Figure 2.7: Typical Rotational Rheometer Geometries

In any of the above rheometer geometries, an assumption is made about the distribution of fluid velocity throughout the material. Therefore, using the dimensions of the rheometer, it is possible to develop analytical equations relating the torque and rotation speed measured by the rheometer to the specific parameters of a given constitutive equation. The equations for a coaxial cylinders rheometer are provided in Section 2.3.2.2 for Newtonian and Bingham materials. Equations for other geometries and constitutive equations are available in rheology texts.

2.3.2.2 Derivation of Equations for Coaxial Cylinders Rheometers The top view of a coaxial cylinders rheometer is shown in Figure 2.8. Only the material in the space between the inner and outer cylinders—namely, the annulus—is considered. Any end effects at the top or bottom of the cylinder or concrete specimen are ignored. The derivations are based on the assumptions that laminar flow occurs, that inertial effects can be ignored, and that the velocity of the material in contact with the surface of a cylinder is equal to the velocity of that cylinder. If the width of the annulus is very narrow compared to the radius of the cylinder, it is possible to use an average radius to compute the shear stress and shear rate. This case of a “narrow gap” is not applicable to concrete and is not considered here. Instead, the case of a “wide gap” is considered.

31 Outer Cylinder R1

Inner Cylinder

h = cylinder height

Figure 2.8: Top View of a Coaxial Cylinders Rheometer

The velocity gradient at any point within the annulus is defined as the derivative of velocity with respect to the radius. From the product rule, the velocity gradient is the sum of the angular velocity, ω , and the shear rate, as indicated in Equation (2.28).

dv rd ω)( dω ω +== r (2.28) dr dr dr

The shear rate, shown in Equation (2.29), is needed for the derivation of flow parameters.

dω γ& = r (2.29) dr

The shear rate and velocity gradient vary throughout the radius. The torque generated at a particular shear rate is considered to act along an imaginary

32 cylinder of radius r, as indicated by the dashed line in Figure 2.8. This value of torque, T, is the product of the surface area of the imaginary cylinder, the shear stress acting over the surface area of the cylinder,τ ; and the radius of the cylinder:

= π τ rrhT ))()(2( (2.30)

Therefore, the shear stress at any point is simply the torque divided by the area of the imaginary cylinder:

T τ = (2.31) 2πr 2 h

The rheometer records the torque acting on either the inner or outer cylinder—these torques are equal and opposite—and the angular velocity of one cylinder relative to the other. Using these data, the flow parameters can be derived based on a known constitutive equation. The derivations for Newtonian and Bingham fluids are provided as follows. First, for a Newtonian fluid, the derivation begins with the constitutive equation, which was given previously as Equation (2.3) and is repeated below:

τ = ηγ&

The Newtonian equation is next rearranged so that shear rate is on the left side of the equation. Equations (2.29) and (2.31) are plugged in for shear rate and shear stress, respectively, as shown in Equation (2.33).

τ γ& = (2.32) η

dω T r = (2.33) dr 2 2 hr ηπ

Equation (2.33) is integrated across the annulus from ϖ = Ω at r = R1 to

ϖ = 0 at r = R2, as shown in Equation (2.34).

0 T R2 1 dϖ = 3 dr (2.34) Ω 2 ηπ ∫∫ R1 rh

The result of the above integration is an expression for a straight line that is a function of viscosity, cylinder height, and cylinder radii, as shown in Equation (2.35). The same result would have been obtained if the outer cylinder rotated while the inner cylinder was fixed, such that the limits of integration would have

been ϖ = 0 at r = R1 to ϖ = Ω at r = R2.

T ⎛ 11 ⎞ =Ω ⎜ − ⎟ (2.35) 4 hηπ ⎜ 2 2 ⎟ ⎝ 1 RR 2 ⎠

If the output from the rheometer is plotted as torque versus angular velocity, the slope of the linear fit can be used in Equation (2.36) to solve for viscosity.

4πhT η = Ω ⎛ 11 ⎞ (2.36) ⎜ − ⎟ ⎜ 2 2 ⎟ ⎝ 1 RR 2 ⎠

34 For a Bingham fluid, the derivation of the flow parameters again starts with the constitutive equation, which was given previously as Equation (2.5) and is repeated below:

τ = τ 0 + µγ&

Again, the Bingham equation is rearranged to place the shear rate on the left side of Equation (2.37). Then, Equations (2.29) and (2.31) are plugged in for shear rate and shear stress, respectively, as shown in Equation (2.38).

−ττ 0 γ& = (2.37) µ

dϖ T τ r −= 0 (2.38) dr 2 2 hr µπ µ

If it is assumed that flow occurs throughout the annulus, integration is performed between the limits of ϖ = Ω at r = R1 and ϖ = 0 at r = R2.

0 R2 ⎛ T τ ⎞ dϖ = ⎜ − 0 ⎟dr (2.39) Ω ∫∫ R ⎜ 3 ⎟ 1 ⎝ 2 hr µπ µr ⎠

The result of the above integration is the Reiner-Riwlin equation, as shown in Equation (2.40). The Reiner-Riwlin equation is an expression for a straight line where the slope is defined in terms of the plastic viscosity, cylinder height, and cylinder radii while the intercept is defined in terms of yield stress, viscosity, and cylinder radii.

35 T ⎛ 11 ⎞ τ ⎛ R ⎞ =Ω ⎜ − ⎟ − 0 ln⎜ 2 ⎟ (2.40) 4 hµπ ⎜ 2 2 ⎟ µ ⎜ R ⎟ ⎝ 1 RR 2 ⎠ ⎝ 1 ⎠

From the rheometer data of angular velocity versus torque, the slope of the straight line fit is determined and set equal to Equation (2.41), so that plastic viscosity can be determined:

4πhT µ = Ω ⎛ 11 ⎞ (2.41) ⎜ − ⎟ ⎜ 2 2 ⎟ ⎝ 1 RR 2 ⎠

With plastic viscosity determined, the only remaining unknown is yield stress, which can be calculated using Equation (2.40). To convert the intercept and slope values for the plot of torque versus rotation speed to yield stress and plastic viscosity as needed in the flow curve, the values of the intercept and slope only need to be multiplied by geometry constants:

⎛ 11 ⎞ ⎜ − ⎟ ⎜ 2 2 ⎟ ⎝ 1 RR 2 ⎠ τ 0 = intercept,( Nm) (2.42) ⎛ R2 ⎞ ⎜ ⎟4ln πh ⎝ R1 ⎠

⎛ 11 ⎞ ⎜ − ⎟ ⎜ 2 RR 2 ⎟ (2.43) µ = ⎝ 1 2 ⎠ Slope,( )Nm.s 4πh

36 2.3.2.3 Effects of Dead Zones in Coaxial Cylinders Rheometers For fluids with a yield stress, the range of shear stresses present in the annulus of a coaxial cylinders rheometer may not be sufficient to cause all material to flow. The result is a dead zone where no flow occurs. Within the concrete literature, the presence of a dead zone is often referred to as plug flow, although this term is a bit of a misnomer because the problem with plug flow is that not all of the material flows. In general rheology, the term plug flow is used to describe the movement of yield stress fluids through a pipe. The flow of a Bingham material in a coaxial cylinders rheometer is illustrated in Figure 2.9 for a case where the inner cylinder rotates and the outer cylinder remains stationary. The shear stress throughout the annulus is greater than the yield stress; therefore, a dead zone does not exist. A 10% increase in radius results in an approximately 17.4% decrease in shear stress. The angular velocity and shear rate are maximum at the inner cylinder. At the outer cylinder, the angular velocity is equal to zero while the shear rate is at its minimum value. When all material flows, the traditional Reiner-Riwlin equation is applicable. In contrast, Figure 2.10 demonstrates the case where a dead zone is present. The portion of the material that flows is subjected to shear stresses greater than the yield stress while the remainder of the material in the dead zone is subjected to shear stresses below the yield stress and does not flow. The angular velocity and shear rate are equal to zero at the transition radius between the flowing material and the dead zone.

37 Shear Stress τ τ0

Shear Rate γ &

Angular Velocity ω

Inner Radius All Material Flows

Outer Radius

Figure 2.9: Flow of a Bingham Material in a Coaxial Cylinders Rheometer – No Dead Zone

38 Shear Stress

τ τ0

Shear Rate γ &

Angular Velocity ω

Flowing Zone Inner Radius

Dead Zone

Transition Radius

Outer Radius

Figure 2.10: Flow of a Bingham Material in a Coaxial Cylinders Rheometer – Dead Zone Present

39 If the traditional Reiner-Riwlin equation is used for a case where a dead zone is present, the error in the computed rheological parameters can be significant. If a dead zone is present, the limits of integration in Equation (2.39) must be changed to cover only the portion of the annulus where flow occurs. By definition, flow ceases where the shear stress in the material equals the yield stress. Therefore, the limits of integration for Equation (2.39) should be changed 1/2 to ϖ = Ω at r = R1 and ϖ = 0 at r = ( 2πhT τ 0 ) for the case where the inner cylinder rotates and the outer cylinder is fixed. The result of the integration with these new limits is shown in Equation (2.44):

T ⎛ 1 2 h ⎞ ττπ ⎛ T ⎞ =Ω ⎜ − ⎟ − 00 ln⎜ ⎟ (2.44) 4 hµπ ⎜ 2 T ⎟ 2µ ⎜ 2 ⎟ ⎝ R1 ⎠ ⎝ 2 τπ Rh 10 ⎠

It is possible to determine the minimum angular speed at which the dead zone is eliminated by substituting the torque at which the dead zone begins to occur at the outer cylinder into the Reiner-Riwlin equation. The result is given in Equation (2.45):

2 τ ⎡1 ⎛⎛ R ⎞ ⎞ ⎛ R ⎞⎤ =Ω 0 ⎢ ⎜⎜ 2 ⎟ − ⎟ − ln1 ⎜ 2 ⎟⎥ (2.45) µ ⎢2 ⎜⎜ R ⎟ ⎟ ⎜ R ⎟⎥ ⎣ ⎝⎝ 1 ⎠ ⎠ ⎝ 1 ⎠⎦

From Equation (2.45), it is evident that the occurrence of a dead zone during a given flow curve measurement is based on three parameters: the rotation speed, the ratio of yield stress to plastic viscosity, and the ratio of the outer radius to the inner radius. The dead zone can be eliminated by increasing the rotation speed, reducing the ratio of yield stress to plastic viscosity, reducing the ratio of the outer radius to inner radius, or some combination of the three. The yield stress and plastic viscosity depend on the material being tested while the rotation speed

40 is determined from the desired shear rate; therefore, in designing a rheometer, the ratio of the outer radius to the inner radius is the only parameter that can be changed. The minimum rotation speed to eliminate the dead zone for a given ratio of yield stress to plastic viscosity is shown in Figure 2.11 as a function of the radii ratio. This graph indicates that it is advantageous to decrease the ratio of outer to inner radius.

rad/s 2.5 ,* ,* 0 µ τ 2

Speed 1.5

1 Rotation

0.5

0 Minimum 1 1.2 1.4 1.6 1.8 2 2.2 2.4 2.6 2.8 3

RR 12 Figure 2.11: Minimum Rotation Speed to Eliminate Dead Zone

It is also instructive to consider the influence of the other two parameters on the errors induced by using the traditional Reiner-Riwlin equation when a dead zone is present. The influence of the dead zone on the measured torque versus rotation speed curve is shown in Figure 2.12. If no dead zone exists, the linear dashed line will be measured. The presence of a dead zone—and concomitant reduction in the amount of material that flows—reduces the amount of torque measured at a given rotation speed. In this case, the dead zone is eliminated at a

41 rotation speed of 77.1 rpm. If all points on the flow curve are measured at speeds above 77.1 rpm, then no error results. When a portion or all of the points on the flow curve are measured at points below 77.1 rpm, the magnitude of the error due to the presence of a dead zone depends on the range of rotation speeds measured. It is clear that erroneously ignoring the dead zone will result in an underestimate of yield stress and an overestimate of plastic viscosity. While the overall curve accounting for the dead zone is nonlinear, if a narrow range of rotation speeds is measured, the portion of the curve over this range may appear linear. The fact that the curve appears linear over a certain region should not be taken to mean that the rheometer is correctly measuring the rheology of a Bingham material. For instance, over a rotation speed range of 10 to 60 rpm, the curve appears to be approximately linear (R2 = 0.997); however, if the presence of the dead zone is ignored, the yield stress will be 18% too low and the plastic viscosity will be 22% too high.

No Dead Zone 10 Dead Zone

4 Measured Torque (Nm)

R2/R1 = 2 2 τ 0 = 10 µ 0 0 102030405060708090100 Rotation Speed (rpm) Figure 2.12: Influence of Dead Zone on Measured Torque versus Rotation Speed Curve 42

Figure 2.13 shows the influence of the ratio of yield stress to plastic viscosity on the error in the measured values of yield stress and plastic viscosity due the dead zone being neglected. The ratios shown in Figure 2.13 are intended to be representative of those found in typical concrete measurements. For the case shown, the error in yield stress is greater than plastic viscosity below a ratio of 7.5 while the error in plastic viscosity is greater above a ratio of 7.5. For self- consolidating concrete with a yield stress near zero, the error due to neglecting the dead zone is near zero; in stiffer concretes with higher ratios of yield stress to viscosity, the importance of properly accounting for the possible presence of a dead zone is greater.

120 Yield Stress 100 Plastic Viscosity

R2/R1 =2 80

Percent Error Percent 20

0 0 102030405060 -20

-40 Yield Stress/ Plastic Viscosity

Figure 2.13: Influence of Ratio of Yield Stress to Plastic Viscosity on Errors due to Neglecting Dead Zone (Rotation Speed = 10 rpm to 60 rpm)

Figure 2.14 shows the effect of reducing the ratio of outer radius to inner radius at a constant ratio of yield stress to plastic viscosity. For a ratio of yield stress to plastic viscosity of 10, the diminishing return for reducing the ratio of

R2/R1 begins to be pronounced at a value of R2/R1 of approximately 1.6, where the error in yield stress and plastic viscosity are 3.7% and 3.6%, respectively.

100

80 Yield Stress Plastic Viscosity

τ 40 0 = 40 µ

τ 0 20 = 10

Percent Error µ

0 1 1.2 1.4 1.6 1.8 2

-20

-40 R2/R1

Figure 2.14: Influence of Radii Ratio on Errors Due to Dead Zone (Rotation Speed = 10 rpm to 60 rpm)

While it has been suggested (Tattersall and Banfill1983; Ferraris 1999) that the ratio of outer radius to inner radius should be no greater than 1.1 or 1.2, the plots in Figure 2.12 through Figure 2.14 indicate that it is necessary to consider also the material to be measured (ratio of yield stress to plastic viscosity) and the desired range of shear rates (minimum rotation speed). Still, a policy of

44 minimizing the ratio of outer radius to inner radius to the greatest extent practical is generally advantageous. While the calculations of stress and velocity distribution presented above are based on theoretical calculations, Raynaud et al. (2002) used magnetic resonance imaging (MRI) to measure experimentally the velocity profiles in a coaxial cylinders rheometer for two thixotropic bentonite suspensions with solids volume concentrations of 3.2% and 4.6%. The maximum velocity at the inner cylinder was found to approach the correct theoretical value. The velocity profiles indicated distinct sheared and unsheared regions in the annulus. The shape of the velocity profiles were not as expected because the shear rate did not go to zero at the critical radius between the sheared and unsheared region, but rather experienced a discontinuity. For transient flow—that is, where the rotation speed of the inner cylinder was changing over time—the location of and stress at the critical radius for a given rotation speed varied based on the flow history of the sample. When a vane was used in place of the inner cylinder, the maximum velocity occurred at a location within the vane and was below the theoretical maximum value at the tips of the vane. Distinct sheared and unsheared regions were still evident in the annulus for the vane measurements. The distribution of velocity for the vane measurements suggested significant secondary flows in the proximity of the vane. As a result of the MRI measurements, the authors concluded that the commonly assumed velocity profiles for coaxial cylinders rheometers with sheared and unsheared zones are invalid and that the true velocity profile should be determined for any test by using an MRI measurement or other suitable technique.

2.3.2.4 Methods to Correct for Dead Zones in Coaxial Cylinders Rheometers Since the errors from neglecting the dead zone can be significant and since the ratio of the inner to outer radius cannot be sufficiently reduced to avoid 45 completely the dead zone in tests of concrete, it is necessary to account properly for the dead zone. Although the actual flows in fresh concrete are complex and can not be determined precisely by analytical means, approximations can be made to improve the accuracy of rheometer results. Three such approaches to account analytically for the presence of the dead zone are described as follows.

Method 1: Point Elimination Method In the first method, the flow curve is measured over the desired range of rotation speeds and any points below the minimum speed to prevent the presence of a dead zone—determined with Equation (2.45)—are eliminated. The remaining points are used to calculate the flow curve. Such a procedure is used for the BML viscometer (Ferraris and Brower 2001). In order to use Equation (2.45), the yield stress and plastic viscosity must be known a priori. Since these parameters cannot be determined until the dead zone points are eliminated, an iterative process must be used. In such a process, the yield stress and plastic viscosity are determined based on all measured points and used to compute the speed at which a dead zone occurs. All points below this dead zone speed are eliminated and the yield stress and plastic viscosity are again computed for the remaining points. Based on these new values of yield stress and plastic viscosity, the dead zone speed is calculated and additional points removed. This procedure is continued until all points below the true dead zone speed are eliminated. Since this method eliminates points, it is possible that all or most points could be eliminated, resulting in an inadequate number of points. Further, the resulting points at higher shear rates may not reflect the properties of the material at lower shear rates.

46 Method 2: Independent Yield Stress Method In the second method, the yield stress is measured independently and used in Equation (2.44)—that is, the modified Reiner-Riwlin equation—to calculate plastic viscosity. The independent measurement of yield stress can be accomplished with a stress growth test or stress decay test, both of which are described in detail later in Section 2.4.1. This method is complicated by the need to perform two separate measurements—one for the direct yield stress and one for the flow curve. The yield stress measured by the direct method is likely to be different than the yield stress determined from the flow curve (with no errors due to a dead zone) because of different test conditions.

Method 3: Effective Annulus Method In the third method, an iterative nonlinear optimization technique is used to determine the value of yield stress and plastic viscosity based on Equation (2.44). A similar method is utilized in the CEMAGREF-IMG rheometer (Ferraris and Brower 2001). The method is based on the assumption that the transition radius between the flowing zone and the dead zone can be correctly calculated. The flowing zone, which changes for each flow point, is then used as the “effective annulus” such that only material in this region is considered in determining constitutive equation parameters. In effect, the dimensions of the outer cylinder are changing throughout the test for each rotation speed. A sample calculation of this procedure is shown in Table 2.1 for a Bingham material tested in a coaxial cylinders rheometer with an inner cylinder radius of 0.1 m, an outer cylinder radius of 0.2 m, and a height of 0.2 m. The measured values of rotation speed and torque are obtained from the rheometer. Then, in the first iteration, values of yield stress and plastic viscosity are used in Equation (2.44) to calculate the “modified” rotation speed at each point. If the yield stress and plastic viscosity are assumed to be 500 Pa and 20 Pa.s,

47 respectively, for the first iteration, the modified rotation speed for the maximum speed point shown in Table 2.1 is calculated as:

T ⎛ 1 2 h ⎞ ττπ ⎛ T ⎞ =Ω ⎜ − ⎟ − 00 ln⎜ ⎟ mod 4 hµπ ⎜ 2 T ⎟ 2µ ⎜ 2 ⎟ ⎝ R1 ⎠ ⎝ 2 τπ Rh 10 ⎠

69.13 ⎛ 1 π )500)(2.0(2 ⎞ )500( ⎛ )69.13( ⎞ =Ω ⎜ − ⎟ − ln⎜ ⎟ mod ⎜ 2 ⎟ ⎜ 2 ⎟ π )20)(2.0(4 ⎝ )1.0( )69.13( ⎠ )20(2 ⎝ π )1.0)(500)(2.0(2 ⎠

Ω mod = 000.5 rad/sec

The difference between the measured rotation speed and modified rotation speed is computed as the error for each point. The mean squared error (mse) for n flow curve points is determined as:

( Ω−Ω ) 2 mse = ∑ measured modified n

For subsequent iterations, the values of yield stress and plastic viscosity are varied using a nonlinear optimization technique to minimize the mean sum of errors. The final values from this nonlinear optimization technique are shown in Table 2.1. The same calculation could have been performed by solving Equation (2.44) for torque and minimizing the error between the measured torque and the “modified” torque. The measured flow curve points and the corrected fitted line based on the yield stress and plastic viscosity values determined from the effective annulus method are shown in Figure 2.15. Had a straight line been fitted through the measured points, the intercept would have been less and the slope greater than the theoretical line for no dead zone.

Table 2.1: Sample Calculation for Effective Annulus Method Measured Values Modified Values Effective Rotation Rotation Torque (Error)2 Outer Speed Speed Radius (rad/sec) (Nm) (rad/sec) (rad/sec)2 (m) 6.283 13.69 6.213 0.0049 0.190 5.027 12.45 5.123 0.0093 0.181 3.770 10.89 3.830 0.0036 0.170 2.513 8.92 2.326 0.0350 0.153 1.257 7.46 1.362 0.0111 0.140 mse 0.013

From Measured Values From Modified Values Yield Value 5.884 Nm Yield Stress 301.6 Pa Viscosity Value 7.994 Nm.s Plastic Viscosity 32.2 Pa.s R2 0.995 mse 0.013 Speed to Prevent 7.56 Dead Zone rad/sec

Further, the effective outer radius at a given rotation speed can be calculated based on Equation (2.30). For the maximum speed point shown in Table 2.1, the radius at which the dead zone begins is calculated as:

T R ,2 eff = 2 hτπ 0

69.13 R = ,2 eff π )6.301)(2.0(2

R ,2 eff = 190.0 m

49 20

18 Corrected for Dead Zone

16 Theoretical for No Dead Zone

8 Torque (Nm)

0 02468 Rotation Speed (rad/sec)

Figure 2.15: Sample Calculation for Effective Annulus Method

For concrete, it is likely that the actual radius at which flow ceases varies to some extent due to the presence of aggregates. It is also important to consider the size of the effective annulus relative to the maximum aggregate size. If this ratio is too small, the flow in the effective annulus will not be representative. Of these three approaches to accounting for the dead zone, the Effective Annulus Method appears to be the most appropriate. The method can be programmed into a rheometer control software program to calculate the modified values of yield stress and plastic viscosity automatically.

50 2.3.2.5 End Effects The derivations presented in Section 2.3.2.2 for coaxial cylinders rheometers were based only on the distribution of shear stress and shear rate in the annulus. If the inner cylinder is immersed into the material being tested such that material exists above or below the inner cylinder, it is clear that the additional shear stress acting on the top and bottom of the cylinder due to this additional material will contribute to the total recorded torque. Additionally, the torque per unit height of cylinder will be less near the ends of the cylinder than near the middle because the maximum velocity gradient is not directed radially outward at the ends of the cylinder (Whorlow 1992). These end effects must be taken into account when computing rheological parameters. It is possible to approximate the amount of torque attributable to end effects or to modify the inner cylinder in such a way as to eliminate end effects. To determine the amount of torque attributable to end effects, Whorlow (1992) suggests measuring the torque for a given rotation speed with the annulus filled to various heights. The plot of torque versus immersed cylinder height should be linear. The slope of the resulting line expresses the torque per unit height while the intercept is a representation of the amount of torque due to end effects. The method may not be suitable for fluids with time-dependent properties. It is also possible to use fully-immersed cylinders of different heights to plot a similar torque versus cylinder height curve. In some cases, the end effects may be independent of the rotation speed. End effects typically vary for different materials, making it necessary to develop individual calibrations for each material. As a result, it is not possible simply to use one fluid to provide a single calibration for all materials. Another approach to eliminating end effects is to modify the geometry of the rheometer to minimize or eliminate shear stresses acting on cylinder ends (Whorlow 1992). One solution for a rheometer with a fixed inner cylinder and

51 rotating outer cylinder is to place fixed guard cylinders above and below the inner cylinder. These fixed guard cylinders are adjacent to but not attached to the inner cylinder, thus allowing the inner cylinder to rotate through small angles and measure only stress acting on the side. Such an approach is used in the BML viscometer for concrete (Ferraris and Brower 2001). Another solution to eliminate torque on the bottom of the inner cylinder is to trap an air bubble under the cylinder. This solution is complicated by the difficulty in verifying the existence of the bubble and the fact that some air may escape during the test. A third solution is to use a double-gap coaxial cylinders rheometer so that the cylindrical area is much larger on the sides than the ends. A fourth solution is to assume that the volume below the cylinder acts as a parallel plate rheometer while the volume horizontally adjacent to the cylinder acts as a coaxial cylinders rheometer. Then, the torque from the parallel plate rheometer portion and the coaxial cylinders rheometer portion are added together to equal the total torque. This procedure is utilized in the FHPCM rheometer for concrete (Yan and James 1997).

2.3.3 Special Topics in the Measurement of Rheology

2.3.3.1 Deborah Number In any rheology test, it is important to consider the time scale of measurement. As indicated earlier, all materials flow if given sufficient time. A material can behave as a solid or liquid depending on the rate at which shear stress is applied. To measure a material properly, the time scale of the measurement must be appropriate for the tested material. This concept is

represented with the Deborah number, De, which is a dimensionless group that relates the characteristic time of the material,τ , to the characteristic time of the deformation process, T:

52 τ D = (2.46) e T

For an ideal linear-elastic solid material, the characteristic time is infinite while the characteristic time of an ideal Newtonian viscous liquid is zero. High Deborah numbers are related to solid-like behavior while low Deborah numbers are associated with liquid-like behavior. If the characteristic time of a material is too high relative to the characteristic time of the deformation process of a given fluid rheometer, the material cannot be measured in that particular rheometer. The concept of the Deborah number is directly applicable to concrete. The Deborah number can be used to determine the minimum workability that can be measured with a given fluid rheometer. Most available rheometers for concrete are only capable of measuring concrete mixtures with slumps greater than approximately 50 to 100 mm (2 to 4 inches). Not only is the torque required to operate a rheometer for low-slump concrete mixtures high, it is questionable whether concepts of fluid rheology can be applied to low-slump concretes. Low- slump concretes do not readily flow under their own weight without vibration and behave more like solids than liquids. As such, the Deborah numbers for low- slump concretes are too high when such materials are tested in existing concrete rotational rheometers. Therefore, it is inappropriate to attempt to measure low- slump concretes in existing concrete rotational rheometers.

2.3.3.2 Estimation of Shear Rate The range of shear rates generated in a rheometer for a given fluid should be similar to the shear rates present in actual field conditions. Due to the possibility of non-linearity in the flow curve, the selection of a proper range of shear rates will ensure that the results are relevant to the given application. According to Schramm (1994), the shear rate for a given application can be

53 estimated as the maximum speed of a fluid as it flows through a gap divided by the gap size:

maximum speed rateshear estimated rateshear = (2.47) size gap size

On a jobsite, the speed with which concrete flows through a pump, down a chute, or through the space between reinforcing bars could be determined and used in the above equation to determine the shear rate. It is not known whether an extensive study has been conducted to determine the actual shear rates in the field for different concrete construction processes. Szecsy (1997) suggests, without data, that 10 sec-1 is a maximum practical rate in the field.

2.3.3.3 Relative Rheometers Instead of measuring fluids in an absolute rheometer and determining rheological parameters in fundamental units, it is possible to use a relative rheometer to measure values related to but not necessarily equal to rheological parameters. Such relative rheometers are typically mixers instrumented to measure the amount of resistance created on mixing blades. The mixers may be those used in a particular industrial or field process or a mixer designed to simulate practical conditions. Unlike traditional rheometers where the shear rate throughout the rheometer is known, the shear rate around a mixing blade is highly complex and essentially impossible to model analytically. Therefore, converting measured torque and rotation speed values to constitutive equation parameters is not generally possible. According to Schramm (1994), the resistance of the test sample to mixing is typically proportional to viscosity. Mixers can be used to create either turbulent flow and high shear action or laminar flow in homogenous materials. The results

54 from mixer tests are specific only to a certain device geometry and manufacturer. Test results can be compared to standard samples that are known to perform well or badly in a specific application. According to Schramm (1994), it is necessary to first establish a matrix of experimental data to relate field performance to mixer sensor reading based on such factors as test temperature, rotor speeds, length of shear time, sample volume, and type of fluid. Whorlow (1992) states that while it might be possible to measure the viscosity of Newtonain fluids using the flow around a mixing blade, it is inappropriate to use such a technique to study more complex flow. Schramm (1994) states that absolute and relative rheology should be used synergistically. Relative rheological measurements can provide important practical information that absolute rheological measurements cannot provide. A variety of measurements can be made with relative rheometers. Schramm (1994) suggests measuring rheograms of torque versus time. These rheograms can be used to determine the effects of changes in fluid composition on rheological properties or how the rheological properties of a fluid with constant composition change with time. Relative rheometer measurements can be made at one speed or multiple speeds. For concrete, relative rheometers typically operate in a manner analogous to traditional absolute rheometers. The devices turn a mixing impeller at a series of fixed speeds and measure the resulting torque. A straight line is fitted to a plot of torque versus rotation speed, as shown in Figure 2.16. The intercept of this line with the torque axis, G, is related to yield stress, while the slope of the line, H, is related to plastic viscosity. Tattersall and Banfill (1983) have utilized a general method for calibrating a relative rheometer in order to convert G and H values to yield stress and plastic viscosity. Whorlow (1992) and Ferraris (1999) have asserted that the accuracies of such calibration techniques are questionable. One of the main problems with using G and H is related to the dead zone problem in coaxial cylinders rheometers. As the mixing

55 blade is rotated at faster speeds, the volume of material that flows increases. Therefore, the increase in torque is attributable both to viscosity and to the fact that a larger volume of material is flowing.

H G 1 Torque, Nm

Rotation Speed, rev/sec

Figure 2.16: Typical Relative Rheometer Results for Concrete

2.3.3.4 Slip at Boundaries (Wall Effect) An important consideration in measuring the rheology of suspensions is the slip that can occur at boundaries of rheometers or viscometers due to wall effect, which is depicted in Figure 2.17. The solid particles, such as aggregates, are normally able to pack together to achieve a certain density; however, solid particles are not able to pack as efficiently against a wall. The exclusion of solid particles in the vicinity of the wall results in a region of low solids concentration where flow is concentrated. Because the flow is not distributed across the rheometer or viscometer as assumed, the analytical equations to relate measured parameters to fundamental parameters are invalid. It is possible to reduce slippage by roughening the rheometer walls to allow improved packing.

56 Low Solids Concentration Layer

Figure 2.17: Wall Effect

2.4 RHEOLOGY OF CONCENTRATED SUSPENSIONS The rheology of concentrated suspension has been studied extensively. Though concrete is a unique material that requires special approaches to rheologocal characterization, the lessons learned from other similar concentrated suspensions are instructive.

2.4.1 Direct Yield Stress Measurements in Rotational Rheometers The yield stress of a concentrated suspension can be measured in a rotational rheometer using indirect or direct methods. Indirect methods, which were discussed previously, involve measuring a flow curve and extrapolating the curve to find the shear stress at a zero shear rate based on a known constitutive equation or a fitted curve. The yield stress determined by an indirect method is dependent on the assumed constitutive model, the accuracy and range of the experimental flow data, and the type of instrument used (Nguyen and Boger 1983). Indirect methods have been found to be unreliable in certain cases due to the presence of slip flow and the lack of data at low shear rates. In contrast, direct methods measure only yield stress, using one of several approaches. In the stress relaxation technique, the material is sheared at a constant shear rate, which is then reduced—either gradually or at once—to zero. The 57 remaining shear stress acting on the rheometer upon cessation of flow is defined as the yield stress. In a stress growth test, which may be performed as a stress- controlled or rate-controlled test, the stress in the material is gradually increased until flow beings. In the stress-controlled version of the test, a constant stress is applied to a sample and the resulting strain is measured over time. This shear stress is gradually increased—either in steps or at a constant rate—until flow occurs. A typical stress-controlled stress growth test plot is shown in Figure 2.18. In the rate-controlled version of the test, the shear rate is maintained at a low, constant level until flow initiates. The maximum shear stress that develops in the sample is recorded as the yield stress. A typical rate-controlled stress growth test plot is shown in Figure 2.19.

Yield Stress Exceeded

Shear Strain Incremental Stresses Applied

Time Figure 2.18: Typical Results for a Stress-Controlled Stress Growth Test

τ0(d)

τ0(s) Shear Stress

Time Figure 2.19: Typical Results for a Rate-Controlled Stress Growth Test

The stress growth plot in Figure 2.19 consists of three portions (Nguyen and Boger 1985; Yan and James 1997; Liddell and Boger 1996). In the initial linear region, the material behaves elastically. Next, as bonds between solid particles are broken, the plot transitions from the linear region to the curved, viscoelastic region. Finally, when a sufficient number of bonds are broken to achieve flow, the yield stress is reached, indicated as the maximum value on the curve. The portion of the curve after the peak is of little value because the amount of material flowing in the annulus is not known (Schramm 1994). The amount of time for shear stress to build up should be sufficiently long so that elastic deformation dominates. Two different yield stresses are indicated in Figure 2.19. The first yield stress, which occurs at the transition from elastic to viscoelastic behavior, is referred to as the static yield stress. The peak shear stress in referred to as the dynamic yield stress. The dynamic yield stress is generally used as the indication of yield stress because it is associated with full structural breakdown and the beginning of plastic flow (Liddell and Boger 1996; Saak, Jennings, and Shah 2001). The yielding that occurs is generally irreversible because the hydrodynamic forces at low shear rates are not strong enough to bring particles close together for the reformation of bonds (Nguyen and Boger 1983).

59 2.4.2 Use of Vane for Direct Yield Stress Measurements

2.4.2.1 Overview Although the direct yield stress methods can be performed in most conventional rheometer geometries, device artifacts can potentially distort results significantly. Instead of using a traditional coaxial cylinders rheometer with smooth or roughened walls, the inner cylinder can be replaced with a vane, such as the one shown in Figure 2.20. As it turns, the vane cuts a cylindrical volume, much like a traditional inner cylinder; however, the slippage due to wall effect is mitigated because yielding occurs within the material, not at the boundary of the material. Specifically, the yield surface occurs along a cylindrical surface defined by the tips of the blade, such that unyielded material exists on either side of the yield surface. Another advantage of using the vane is that the action of inserting the vane into the specimen creates minimal disruption to the specimen, which is particularly important for thixotropic materials where shear history influences results.

Figure 2.20: Typical Vane Impeller

60 The vane method is generally considered to be best for concentrated suspensions with yield stresses above 10 Pa (Barnes and Nguyen 2001). Nguyen and Boger (1983) compared the results from the rate-controlled stress growth test with a vane to both a stress relaxation measurement with a coaxial cylinders rheometer and an indirect measurement with a capillary tube rheometer. For tests of a red mud suspension at a range of solids concentrations greater than 64%, the three methods provided consistent results. Some discrepancies were found between the vane geometry and the other two methods at lower solids contributions. These discrepancies were attributable to a greater contribution from viscous forces. Yoshimura et al. (1987) measured a series of oil-in-water emulsions with yield stresses ranging from 50-550 dynes/cm2 and susceptible to slip. The test results indicated that the concentric cylinder, parallel disk, and vane geometries gave comparable results. For Bentonite suspensions, Alderman, Meeten, and Sherwood (1991) found that the vane geometry worked best for suspensions with yield stresses greater than 10 Pa, while the coaxial cylinders geometry worked best for suspensions with yield stresses less than 2 Pa. Between 2 Pa and 10 Pa, both geometries worked equally well. The use of a vane for rheology measurements dates to at least the 1930s. The concept originated in soil mechanics, where vanes are used to determine the shear strength of soils, as currently described in ASTM D 2573. Vanes have been commercially available since the 1950s and, since that time, have been used to measure a wide variety of concentrated suspensions, as summarized by Barnes and Nguyen (2001).

2.4.2.2 Vane Dimensions A variety of recommendations for vane dimensions have been proposed. Vanes can consist of 4 to 8 individual blades. Yoshimura et al. (1987) found no difference in measurements of oil-in-water emulsions between using a vane with 61 four blades or eight blades. Based on finite element modeling of a Herschel- Bulkley fluid, Keentok, Milthorpe, and O’Donovan (1985) determined that a continuous shearing zone did not form for a two-bladed vane, whereas the shearing zone for a three-bladed vane was irregularly shaped. The container should be sufficiently large so that the effect of its boundaries can be ignored. The recommendations of Nguyen and Boger (1985) for vane dimensions are shown in Table 2.2. Elsewhere, Nguyen and Boger (1983) state that the depth of the material should be twice the length of the vane and the diameter of the container should be twice the diameter of the vane.

Table 2.2: Recommended Vane Dimensions from Nguyen and Boger (1985) Parameter Requirements (Vane Height)/(Vane Diameter) H/D < 3.5 (Container Diameter)/(Vane Diameter) DT/D > 2.0 (Vertical Distance from Top of Blade to Z /D > 1.0 Surface)/(Vane Diameter) 1 (Vertical Distance from Bottom of Container Z /D > 0.5 to Bottom of Blade)/(Vane Diameter) 2

The required dimensions for the use of the vane in determining the shear strength of soil, based on ASTM D 2573, are shown in Table 2.3.

Table 2.3: Requirements for Vane for Soil Undrained Shear Strength Measurements (ASTM D 2573) Parameter Requirements Vane Diameter, D 35 to 100 mm (1.5 to 4 in.) Vane Shaft Diameter, d 12.5 to 16.5 mm (0.5 to 0.65 in.) Vane Height, H 1D < H < 2.5D Taper Angle, i Usually 0 (rectangular) or 45 degrees (tapered) Blade Thickness Average=2mm, max=3mm Number of Blades 4 Vane Area Ratio Area of vanes/gross area swept < 12%

62 Liddell and Boger (1983) found that yield stress was independent of vane dimension for the dimensions considered. Alderman, Meeten, and Sherwood (1991) utilized a set-up where the height below the vane was equal to the blade height, the height above the vane was half the vane height, and the diameter of the container was three times larger than the diameter of the vane. Yoshimura (1987) positioned the top of the vane at the top of the fluid in order to eliminate the torque contribution from the top of the vane. In any vane configuration, the vane size should be selected in order to achieve sufficient torque resolution.

2.4.2.3 Location of Yielding It is generally assumed that the yield surface occurs along an imaginary cylinder defined by the tips of the vane. While this assumption is generally acceptable for most materials, it has been suggested that yielding may actually occur at a diameter larger than that of the vane (Nguyen and Boger 1985). Keentok, Milthorpe, and O’Donovan (1985) showed experimentally that the ratio of the yield surface diameter to the vane diameter could range from 1.00 to 1.05 for four different automotive greases. The experimental data showed that the location of the yield surface was dependent on yield stress, plastic viscosity, elasticity, and shear history (for thixotropic materials). The vane diameter had a negligible effect on the location of the yield surface. Nguyen and Boger (1985) point out the experimental measurements of Keentok, Milthorpe, and O’Donovan (1985) were made well after yielding. An associated finite element computer simulation performed by Keentok, Milthorpe, and O’Donovan (1985) indicated that a ratio of the yield surface diameter to the vane diameter of approximately 1.025 was acceptable. Figure 2.21 shows the boundaries of the yielded material superimposed on the finite element mesh. The outer edge of the yielded material extends beyond the vane blade. The

63 approximately circular shape of the yielded surface was confirmed with experimental observations.

Boundaries of yielded material

Figure 2.21: Yield Surface Superimposed on Finite Element Mesh for a Four- Bladed Vane (Keentok, Milthorpe, and O’Donovan 1985)

The finite element mesh shown in Figure 2.21 is relatively coarse. Yan and James (1997) utilized a more sophisticated finite element model with a substantially finer mesh that incorporated more than 1,300 elements. Their work was based on a Herschel-Bulkley fluid with a yield stress ranging from 100 Pa to 400 Pa, a plastic viscosity ranging from 1 Pa.s to 50 Pa.s and power index equal to 0.5 or 1.0. Whereas Keentok, Milthorpe, and O’Donovan (1985) determined that the shear rate was at its maximum at the tip of the vane blades and lower elsewhere in the narrow annular region where yielding occurs, Yan and James (1997) determined that the shear rate and shear stress were nearly uniform in the narrow annular region. The velocity of the fluid within the cylindrical region defined by the vane dimensions was equal to the velocity of the vane while the velocity of the fluid outside the vane was zero. The result was a sharp velocity gradient along a cylindrical-shaped region defined by the tips of the blade.

64 Figure 2.22 shows schematically the yielding process as confirmed by the finite element modeling of Yan and James (1997). As the vane first begins to rotate, elastic deformation occurs and the shear stress and shear strain are concentrated only at the vane tips. When the stress in the fluid is increased further, yielding begins to occur at the blade tips and viscoelastic flow occurs due to the contributions from both elastic deformation and viscous flow. Finally, when the yielding process is complete, the shear rate and shear stress are distributed approximately uniformly around the cylinder as viscous flow occurs.

te – Elastic Flow tve – Viscoelastic Flow t2 – Viscous Flow

Figure 2.22: Schematic Representation of Yielding Process in a Stress Growth Test (Yan and James 1997)

The finite element models of Keentok, Milthorpe, and O’Donovan (1985) and of Yan and James (1997) are based on vanes of infinite length and do not consider end effects.

2.4.2.4 End Effects and the Calculation of Yield Stress In order to calculate the yield stress from torque readings, it is necessary to consider analytically the distribution and magnitude of shear stress acting on the

65 ends of the vane. From equilibrium, the total torque acting on the vane, T, is the

sum of the torques attributable to the side, Ts, and the two ends, Te, of the vane:

= + 2TTT es (2.49)

Assuming that yielding occurs at the cylindrical surface defined by the tips of the blade, the torque attributable to the side of the vane can be expressed in terms of the vane diameter, D, vane height, H, and the shear stress acting on the side of the vane, τ s . Because the distribution of stress on the ends of the vane is unknown, it can be represented with an integral in terms of an unknown function

of end shear stress, τ e (r), and a radius, r, as shown in Equation (2.49).

D 2/ ⎛ π 2 ⎞ ⎛ 2 ⎞ = ⎜ HDT ⎟ s + ⎜ τπτ e )(22 drrr ⎟ (2.49) ⎝ 2 ⎠ ⎝ ∫0 ⎠

To solve Equation (2.49), additional testing or assumptions are needed. Three methods are available for this purpose.

Method 1: Uniform Distribution of Shear Stresses In the first method, the shear stresses acting on the side and ends of the vane are assumed to be evenly distributed and equal to the yield stress when the maximum torque is reached. The total torque at yielding is thus given by:

π 3 ⎛ HD 1 ⎞ T = ⎜ + ⎟τ o (2.50) 2 ⎝ D 3⎠

The yield stress can then be calculated based on the maximum torque and the vane dimensions.

Method 2: Power-Law Distribution of End Shear Stresses In the second method, the shear stresses on the ends of the vane are assumed to vary with radius based on a power-law relationship. While the shear stress along the side of the vane is equal to the yield stress, the shear stresses at the ends of the vane vary from zero at the center of the vane to the yield stress at the tips of the vane. The end shear stresses may be expressed based on Equation (2.51):

m ⎛ 2r ⎞ D τ e ()r = ⎜ ⎟ τ s for 0 r ≤≤ (2.51) ⎝ D ⎠ 2

Using this power-law relationship, the total torque acting on the vane is given by Equation (2.52):

π 3 ⎛ HD 1 ⎞ T = ⎜ + ⎟τ o (2.52) 2 ⎝ mD + 3 ⎠

To solve for the two unknowns, τ o and m, the above equation can be rearranged to the form a straight line, where the independent variable is H/D:

2T H τ τ += o (2.53) πD 3 o mD + 3

By performing at least two measurements with vanes with different ratios of H/D, the two unknowns can be determined.

Method 3: Unknown Distribution of End Shear Stress In the third method, no assumption is made about the distribution of shear stress on the ends of the vane. Instead, Equation (2.49) can be rearranged into the form of a straight line, where the independent variable is the height of the vane:

D 2/ ⎛ π 2 ⎞ 2 = ⎜ 0 ⎟HDT + τπτ e )(4 drrr (2.54) ⎝ 2 ⎠ ∫0

The intercept of the line is equal to the total torque acting on the ends of the vane. By making measurements with at least two vanes with different heights, the value of the slope term can be determined and used to calculate yield stress. Nguyen and Boger (1985) performed experimental measurements with the above three methods and determined that the third method, which uses a torque balance to ignore end shear stress, was the most accurate method. Due to the need to make measurements with at least two different vanes, the first method was determined to be preferable for practical work. Separately, Nguyen and Boger (1983, 330) determined that as long as the ratio of H/D is greater than 2, the error resulting for using a constant end shear stress versus using an end shear stress based on a power-law distribution “should not be excessive.” Yoshimura et al. (1987) successfully used the first method with a controlled-stress rheometer. Alderman, Meeten, and Sherwood (1991) found that end effects depended on yield stress—and thus solid concentration—for bentonite suspensions. The end effects were negligible for low yield stress fluids but needed to be considered for higher yield stress fluids.

68 2.4.2.5 Effect of Rotation Speed on Stress Growth Test Results The rotation speed used in a rate-controlled stress growth test can have a significant influence on the yield stress. The effect of rotation speed on yield stress measurements for cement paste is shown in Figure 2.23.

Figure 2.23: Effect of Rotation Speed on Yield Stress (Saak, Jennings, and Shah 2001)

If the rotation speed is too high, viscous resistance of the fluid, instrument inertia and insufficient damping can lead to errors (Nguyen and Boger 1983). If the rotation speed is too low, the reformation of network bonds and reorientation of particles can increase the measured yield stress (Liddell and Boger 1996). Therefore, the optimum rotation speed should correspond to the minimum yield stress measurement. While the data presented in Figure 2.23 indicate an optimum value of rotation speed, Nguyen and Boger (1983) found that for concentrated red mud suspensions, rotation speeds below 8 rpm had no effect on yield stress. Others

69 have recommended setting the rotation speed to the minimum possible value (Barnes and Nguyen 2001). It is possible that some authors have not detected the portion of the curve below the optimum value because some rheometers are not capable of rotating at sufficiently low rotation speeds for a given material.

2.4.2.6 Measurement System Stiffness The stiffness of the measurement system in a rheometer can affect the shape of the torque versus time plot for stress growth tests. Liddell and Boger (1996) examined this effect with two different rate-controlled rheometers, one of which was 550 times stiffer than the other. The first rheometer had a stiffness of 0.002 Nm/rad while the other had a stiffness of 1.1 Nm/rad. The response of the flexible system was characterized by a mainly elastic response with minimal viscoelastic response. In contrast, the stiff system exhibited very fast elastic build-up followed by extensive viscoelastic behavior prior to yielding. Therefore, the stiffer rheometer had to be operated at a slower speed in order to extend the time of the elastic behavior.

2.4.3 Extension of Vane Geometry to Viscosity Measurements Given the success of the using the vane for determining yield stress, the use of the vane has been extended to measurements of viscosity (Barnes and Nguyen 2001). If the material within the vane moves with the vane, then a vane- in-cup rheometer should be equivalent to a coaxial cylinders rheometer, with the exception that slip is prevented. Barnes and Carnali (1990) developed and experimentally verified a finite element model of a vane-in-cup rheometer to analyze the flow of material. Figure 2.24 shows the streamlines determined for a Newtonian fluid and power-law fluid (exponent k=0.2). In the Newtonian fluid case, the streamlines are not circular,

70 indicating some fluid within the swept-out region exchanges with fluid from the annulus. In the power-law fluid case, the streamlines form concentric circles, indicating that the assumption of an equivalent cylindrical bob is reasonable. The transition between circular and non-circular streamlines was found to occur at a power-law exponent of k = 0.5.

Newtonian Fluid Power-Law Fluid (k=0.2) Figure 2.24: Finite Element Solution Streamlines for Narrow Gap Vane-in- Cup Rheometer (Barnes and Carnali 1990)

Figure 2.25 shows a comparison of the shear stress distribution in the annulus for the vane and coaxial cylinders geometries. In the vane rheometer, the shear stress distribution is essentially flat except at the vane tip. In order to prevent the occurrence of a dead zone where no flow occurs, the shear stress in this flat area must be greater than the yield stress.

71 Vane Blade Bob Tip

Stress Stress

Radius Radius Angle Cup Angle Cup

Vane-in-Cup Coaxial Cylinders Figure 2.25: Shear Stress Distribution in Annulus Based on Finite Element Analysis (Barnes and Carnali 1990)

2.5 APPLICATION OF FLUID RHEOLOGY CONCEPTS TO CONCRETE In selecting or designing a rheometer for cement paste, mortar, or concrete, it is important to take into consideration the properties that are to be measured and the expected ranges of values for these variables. The measurement of cement paste or concrete rheology is a highly specialized topic. The measurement of cement paste rheology is simpler and better established than concrete rheology. Unlike concrete, which must be measured with specially made rheometers, cement paste can be measured using conventional, commercially available rheometers with few modifications. Coaxial cylinders and parallel plate rheometers are commonly used (Ferraris, Obla, and Hill 2000; Rahman and Moncef 2003). Furthermore, measuring the rheology of cement paste instead of concrete allows the use of a substantially smaller sample size and the elimination of aggregate effects. Measuring cement paste rheology is an effective way to evaluate the effects of admixtures and supplementary cementitious materials on flow properties. The rheology of cement paste can be related to concrete rheology, although the relationship is complex (Ferraris and

72 Gaidis 1992). The measurement of cement paste is beyond the scope of this report. Concrete rheology presents several unique challenges due to the nature and composition of concrete. The main problem with properly characterizing the rheology of concrete is the large size of coarse aggregates. The general rule for rheometers is that the gap size, or distance between parts, should be at least ten times the maximum particle size (Van Wazer et al. 1963; Ferraris 1999). This rule has been supported by experiments showing that particle packing density begins to approach a steady value once the ratio of container diameter to particle diameter reaches ten (McGreary 1961). For common concrete mixtures with a maximum aggregate size of 1 inch, a 10-inch gap would be required. A second potential problem results from the dead zone that can occur in yield stress fluids. As described earlier, the ratio of outer radius to inner radius should be set to an acceptably low value for the material and range of shear rates to be tested. For concrete, the maximum value of this radii ratio has been suggested as 1.2 (Tattersall and Banfill1983) or 1.1 (Ferraris 1999). If a rheometer were constructed based on the requirements that the ratio of the gap size to the maximum aggregate size be ten and the ratio of the outer radius to inner radius be 1.2, the required sample volume for such a rheometer would be 2.6 m3 (3.4 yd3) (Tattersall and Banfill,1983). A further problem for concrete rheometers is the slippage that occurs at the walls of rheometers due to wall effect. This problem must be mitigated by selecting components with proper texture. Due to the challenges described above in building a conventional, absolute concrete rheometer—namely, one that directly measures fundamental parameters—attempts were made beginning at least as early as the 1970s to build relative rheometers which measure mixing action in order to determine parameters that are related to but not equal to fundamental parameters. Although

73 relative rheometers can be effective tools for industrial or field settings, it is desirable to measure rheological parameters in fundamental units. Further, the calculations presented in Section 2.3.3.3 indicate that the presence of a dead zone can result in significant errors in using relative parameters in place of determined fundamental parameters analytically. Given the practical impossibility of achieving an ideal geometry, it is necessary to make compromises in the determination of rheological parameters. Although absolute rheometers have been built, measurement artifacts can significantly distort results. Further, the results between different absolute rheometers for the same concrete can vary widely. Mork (1996, 373) states that it is “impossible” to determine shear stress and velocity profiles in concrete, and, therefore, recommends using relative rheometer parameters of G and H. Despite years of availability, relative and absolute concrete rheometers are used only on a limited basis. While concrete rheometers do provide useful new information about workability, several significant factors have stymied the adoption of rheometers for use on a more widespread basis. First, rheology is still largely an unfamiliar topic for the concrete industry. Rheology is a multidisciplinary subject incorporating mathematics, physics, and chemistry and can appear quite complicated. As Barnes et al. (1989) indicated, “rheology is a difficult subject.” An array of factors must be considered in selecting a rheometer, determining test settings, and interpreting test results. Presently, no ASTM International standard test method is available for concrete rheology. The availability of multiple concrete rheometers with varying features adds to the complexity of implementing rheology measurements. Second, concrete rheometers can only be used for a limited range of workability. Most rheometers are not capable of measuring concretes with slumps below 50 to 100 mm (2 to 4 inches). For these dry-consistency mixtures, the torque needed to turn the rheometer is much greater than for more fluid mixes.

74 Further, dryer mixes tend to have higher ratios of yield stress to plastic viscosity; therefore, the presence of a dead zone is more likely. While it is technically possible to build a rheometer with a sufficiently large motor and to analytically account for a potential dead zone, it remains an open question whether it is appropriate to apply the same concepts of rheology used for moderate- and high- slump concretes to low-slump concretes, which do not readily flow without vibration. Third, the cost of concrete rheometers is prohibitive for many applications. Prices of commercially available rheometers can range from approximately $20,000 to upwards of $50,000. The availability of concrete rheometers is limited. The major worldwide makers of general purpose fluid rheometers do not sell models appropriate for concrete. Instead, concrete rheometers must be custom built or purchased from a small number of specialty sources from around the world. Fourth, concrete rheometers are still mainly in the research domain and need additional development work before being used on a widespread basis. Each existing rheometer has important drawbacks—such as geometry restrictions or a tendency to cause segregation—that must be overcome prior to widespread use. Anyone using a rheometer must be familiar with these drawbacks in order to interpret test results properly. The absolute magnitudes of rheological parameters can vary significantly between different rheometers even when testing identical cement paste or concrete mixtures (Rahman and Moncef 2003; Ferraris and Brower 2001). Ferraris and Martys (2003) have shown that it is possible to relate the results from different rheometers; however, additional work is needed to develop and confirm their theory. At the time of this writing, the wide variation in absolute rheometer results means that absolute rheometers must be used as relative rheometers, that is, the results from a given absolute rheometer can only be compared to other test results obtained on that same absolute rheometer.

75 Finally, rheological parameters need to be related to practical field applications. In a sense, learning to use a rheometer is like learning to convert from English units to metric units: most workers in concrete have a good sense of how concrete with a certain slump reading should appear; however, they do not have a practical understanding of what a concrete with a certain yield stress and plastic viscosity should look like. Before concrete rheology can be used extensively, rheological parameters need to be related to actual field applications. Appropriate ranges of yield stress and plastic viscosity need to be defined for specific applications, such as for placing bridge decks or casting structural columns. The concept of a workability box, shown in Figure 2.26, that would define a zone of acceptable rheology, has been suggested for this purpose (Tattersall 1991).

Zone of Acceptable Rheology Plastic Viscosity

Yield Stress

Figure 2.26: Workability Box for a Specific Application

Despite these drawbacks, concrete rheometers do provide important information about concrete flow properties. Additional development work can address these problems and further facilitate the application of fluid rheology to the fresh concrete.

CHAPTER 3: FACTORS INFLUENCING CONCRETE RHEOLOGY AND WORKABILITY

3.1 INTRODUCTION The rheology and workability of concrete are influenced by nearly every aspect of the mixture proportions, material characteristics, and construction conditions. The effects of many of these factors on workability and slump are well known and widely reported. Less data exist for concrete rheology. This chapter presents an overview of the influence of key factors on both workability and rheology. The focus, however, is on rheology because trends in rheology are more readily quantified and can be used for direct comparison with the results from the ICAR rheometer. The generalization of trends in rheology, even for a single factor ceteris paribus, is fraught with complications. First, a trend in rheology for one variable is also a function of other characteristics of the concrete mixture. For instance, the use of an admixture may have a certain effect in one particular concrete mixture but have a reverse effect when used in a separate concrete mixture of a different composition. Second, the interactions between different admixtures can be significant. Third, materials from different sources—or even the same source— can vary widely in their composition and physical characteristics. A trend drawn from data for a single material source—such as one fly ash, one ground granulated blast furnace slag, or one aggregate—should not be extended to all fly ashes, all ground granulated blast furnace slags, or all aggregates of a particular mineralogy. Fourth, rheological measurements can be a function of measurement technique. As shown by Ferraris and Brower (2001), the rheological parameters

77 measured by different concrete rheometers, even on the same concretes, can vary between different rheometers. Finally, the historical lack of suitable techniques for quantifying concrete rheology has resulted in a paucity of data in the literature on the effects of various factors on concrete rheology. While some factors, such as high-range water reducer (HRWR) dosage, have been reported widely, others have been reported scarcely, if at all. A broad range of data from various sources is desirable for drawing general conclusions. The effects of chemical admixtures and supplementary cementitious materials are often described in the literature in terms of cement paste rheology. By measuring just cement paste, the influence of aggregates can be eliminated and smaller mixtures can be tested. The role of aggregates is important, however, in relating measurements from cement paste to concrete. In some cases, a trend in rheology for a particular mixture change in cement paste may be reverse in concrete (Tattersall and Banfill, 1983).

3.2 EFFECTS OF CEMENT

3.2.1 Cement Content An increase in the cement content, at a constant water-to-cement ratio, provides more paste to coat aggregates and to fill the spaces between aggregates, resulting in improved workability. Smeplass (1994) found that an increase in cementitious materials content (cement with 5% silica fume) relative to aggregate volume resulted in a decrease in both yield stress and plastic viscosity.

3.2.2 Cement Characteristics The chemical composition and physical characteristics of cement can significantly influence workability. Even for a single type of cement, as defined

78 by ASTM C 150 or ASTM C 1157, the changes in cement characteristics can be consequential. Hope and Rose (1990) examined the effects of cement composition on the water demand required for a constant slump. Although the correlations between composition and water demand varied between different aggregates and mixture proportions, the authors were able to draw several conclusions. The water demand increased for cements with high Al2O3 or C2S contents and decreased for cements with high loss on ignition, high carbonate addition, or high C3S content. The particle size distribution of the cement was found to be significant for concrete made with angular aggregate and less pronounced for concrete made with rounded aggregate. For the concrete with angular aggregate, the cements with a higher portion of material smaller than 10µm exhibited higher water demand. The specific surface, however, had minimal influence on water demand. Vom Berg (1979) determined that increasing cement fineness resulted in exponential increases in both yield stress and plastic viscosity for cement pastes. Mork and Gjoerv (1997) found that the ratio of gypsum-to-hemihydrate in cement could influence concrete rheology. For a cement with high contents of

C3A and alkalis, a reduction in the gypsum-to-hemihydrate ratio resulted in a decrease in yield stress but little change in plastic viscosity. When a melamine- based HRWR was used, the trend was reversed, with a lower gypsum-to- hemihydrate ratio resulting in an increase in yield stress. For a cement with lower contents of C3A and alkalis, the effects of the gypsum-to-hemihydrate ratio were less pronounced. Further, a reduction in the sulfate content from 3 to 1 percent resulted in a decrease in both the yield stress and plastic viscosity.

3.3 EFFECTS OF WATER CONTENT An increase in the water-to-cementitious materials ratio in either concrete or cement paste results in reductions in both yield stress and plastic viscosity (Tattersall and Banfill 1983; Tattersall 1991; Mork 1996; Szecsy 1997). The 79 addition of water reduces the solids concentration, resulting in less resistance to flow. Workability is improved with increasing water-to-cementitious materials ratios up to a certain point, after which segregation can become a problem.

3.4 EFFECTS OF AGGREGATES

3.4.1 Aggregate Volume Fraction An increase in the total volume fraction of aggregate in concrete results in increases in yield stress and plastic viscosity (Szecsy 1997; Geiker et al. 2002). Higher volume fractions of aggregates result in reduced spacing between aggregates and, thus, greater resistance to flow. The relationship between solids volume concentration and viscosity is well established for concentrated suspensions (Barnes et al. 1989).

3.4.2 Sand-to-Aggregate Ratio Workability can be improved by optimizing the sand-to-aggregate ratio (S/A). Optimum values of S/A exist for minimizing yield stress and plastic viscosity (Tattersall 1991; Szecsy 1997). An optimum S/A for yield stress may not be optimum for plastic viscosity. Increasing or decreasing the S/A from its optimum value results in increases in yield stress or plastic viscosity. At high values of S/A, a reduction in sand content results in a reduction in the surface area of aggregates that must be coated with cement paste and, thus, a reduction in the resistance to flow. When the sand content is reduced below the optimum value, the result is a lack of fine aggregates to fill the voids between coarse aggregates and, thus, increased resistance to flow. For tests reported by Tattersall (1991), the minimum value of yield stress occurred at an S/A of about 0.33, while the minimum value of plastic viscosity

80 was reached at an S/A of approximately 0.40. The exact value was a function of water-to-cement ratio. Szecsy (1997), when testing crushed limestone and river gravel coarse aggregates, found that the minimum yield stress was achieved at an S/A of approximately 0.40 while plastic viscosity was minimized at an S/A of approximately 0.30. In comparison, S/A values of approximately 0.50 are typical for self-consolidating concrete.

3.4.3 Shape and Texture Aggregate shape and texture strongly influence concrete workability and rheology. In concentrated suspensions, any deviation from a spherical shape results in an increased viscosity (Barnes et al. 1989). Quiroga (2003) found that aggregates with spherical, cubical, or rounded shapes and smooth textures required less cement and water to achieve the same slump as aggregates with flat, elongated, or angular shapes and rough textures. Spherical shapes are preferable because they more readily flow past each other and have reduced specific surface area (Tattersall 1991). Quiroga (2003) found that when gradation was held constant, aggregates with greater packing density, which is related to shape and texture, produced higher slumps. Tattersall (1991) suggests that particle shape has a greater influence on plastic viscosity than on yield stress and that texture has no significant effect on rheology.

3.4.4 Gradation The gradation, or particle size distribution, of aggregate plays a critical role in the workability and rheology of concrete. Ideally, the gradation should take into account all materials, including the cementitious materials and aggregates. In concentrated suspensions, increasing the polydispersity, or spread of sizes, decreases viscosity (Barnes et al. 1989).

81 Concretes produced with gap-graded aggregates, which intentionally omit certain size fractions, can be harsh and more susceptible to segregation. Quiroga (2003) found that uniform aggregate particle size distributions required less water for a given slump than other gradations. In designing a concrete mixture, the gradation can be optimized for a variety of objectives, such as slump, packing density, uniformity, or plastic viscosity. Quiroga (2003) found that mixtures optimized for maximum packing density or slump produced harsh mixtures with poor workability and high susceptibility to segregation. Concrete mixtures above the line on the 0.45 power chart resulted in stiff mixtures, while mixes below the line resulted in harsh, segregating mixtures. Quiroga (2003), therefore, recommends selecting a gradation that strikes a balance between high packing density and uniform grading.

3.4.5 Microfines Content The addition of microfines can improve or reduce workability depending on the quantity and characteristics of the microfines, as well as the composition of the rest of the concrete mixture. Like larger aggregates, the quantity, shape, texture, and particle size distribution of the microfines are important in achieving improvements in workability. The addition of microfines increases the surface area that must be wetted; however, the provision of fines can improve the particle size distribution and result in an overall improvement in flow characteristics. Ho et al. (2002) evaluated the addition of either limestone or granite powder in a cement paste intended for use in self-consolidating concrete. The limestone powder and granite powder had approximately 80% and 75% passing the #200 sieve, respectively, and were obtained as dust from the aggregate crushing process. In general, the replacement of cement with the inert powders at rates up to 55% reduced cement paste yield stress and plastic viscosity. All cement paste samples incorporated one of two different high-range water-

82 reducing admixtures and maintained a constant water-to-powder ratio (cement and filler). The reduction in Bingham parameters was less pronounced for the granite powder, which tended to have flakey and elongated shapes. Ghezal and Khayat (2002) examined the use of a limestone filler material with a Blaine fineness of 565 m2/kg and 97.2% of particles smaller than 45 µm. When used in self-consolidating concrete mixtures at rates up to 100 kg/m3 with a constant water-to-powder ratio, the limestone filler resulted in decreases in yield stress and plastic viscosity. The change was most pronounced at low cement levels. The use of limestone filler also enhanced the stability of the concrete mixtures. Quiroga (2003) found that the addition of microfines resulted in increased dosages of water-reducing admixture required to achieve a constant slump; however, the effect of microfines varied widely, with limestone microfines requiring less HRWR than granite or traprock microfines. The rate of increase in demand for HRWR became significantly higher when the percentage of microfines exceeded 15% of the total fine aggregate mass.

3.5 EFFECTS OF CHEMICAL ADMIXTURES

3.5.1 Water-Reducing Admixtures Water-reducing admixtures enhance workability by reducing the water-to- cementitious materials ratio needed to achieve a given slump. Alternatively, they can be used to increase slump for a given water-to-cementitious materials ratio, reduce cement content while keeping the water-to-cementitious materials ratio constant, or some combination of the above applications. The exact effects of water-reducing admixtures depend on the chemical composition of the admixture and the mixture proportions of the concretes to which they are added. In general,

83 however, water-reducing admixtures result in significant decreases in yield stress, while plastic viscosity typically increases or decreases modestly. Mork (1996) suggests that, in general, low-range water reducers decrease yield stress and plastic viscosity, while high-range water reducers decrease yield stress and increase plastic viscosity. For both types of admixtures, the changes in plastic viscosity are most pronounced at high admixture dosages. Similarly, Smeplass (1994) found that the use of high-range water reducer in concrete mainly reduced yield stress but had little impact on plastic viscosity. For cement paste, Ho et al. (2002) found that two high-range water reducers decreased yield stress, but resulted in minimal decreases in plastic viscosity. Tattersall (1991) reported that the use of a lignosulphonate-based low- range water-reducing admixture in concrete resulted in a reduction in both yield stress and plastic viscosity, although the effect on yield stress was more pronounced. The decrease in these values was most pronounced at low dosages and leveled off at higher dosages. In contrast, the use of melamine sulphonate-, naphthalene sulphonate-, and lignosulphonate-based high-range water-reducing admixtures in concrete all resulted in dramatic reductions in yield stress but little change in plastic viscosity. Again, the effects of using of these admixtures was most pronounced at low dosages and decreased with increasing dosage. Tattersall (1991) also presented data showing that the addition of a high- range water-reducing admixture resulted in an increase in viscosity when used in a concrete with a low sand content (S/A = 0.35), but a decrease in viscosity when used in a concrete with a high sand content (S/A = 0.45). The change in yield stress was approximately the same regardless of the sand content. Tattersall and Banfill (1983) suggest that at low sand contents, the flocculated cement paste separates coarse particles; therefore, when the cement is deflocculated, the coarse particles come closer together and generate greater resistance to flow. The result is an increase in plastic viscosity of the concrete in spite of the decrease in

84 viscosity of the cement. In mixes with a high sand content, the sand fills more of the space between coarse particles. As a result, a reduction in viscosity of the paste results in a reduction in the viscosity of the concrete because the coarse particles do not move sufficiently closer together. Billberg, Petersson, and Norberg (1996) used melamine- and naphthalene- based high-range water-reducing admixtures and found a reduction in both yield stress and plastic viscosity. The concrete tested had an S/A ratio of 0.57 and a maximum aggregate size of 16 mm. The reduction in yield stress was greater in percentage terms—whereas the yield stress was reduced from 600 Pa to approximately 100 to 200 Pa, the plastic viscosity was reduced from 30 Pa.s to approximately 15 to 20 Pa.s. According to Tattersall (1991) the effects of naphthalene- and melamine- based high-range water reducers depend on cement characteristics. Further, increasing the cement content increases the potency of high-range water-reducing admixtures. Faroug, Szwabowski, and Wild (1999) found that the effects of naphthalene- and melamine-based high-range water reducers were most pronounced at low water-to-cement ratios. The use of both types of high-range water reducers in concrete resulted in decreases in yield stress and plastic viscosity. The admixtures had essentially no effect on plastic viscosity above a water-to-cement ratio of 0.40 or on yield stress above a water-to-cement ratio of 0.50. The decline in potency with increasing water-to-cement ratios was attributed to the increase in the ratio of total water to adsorbed capillary and floc water. Although the plastic viscosity did not change when the water-to-cement ratio was increased to 0.50, the additional water released through the action of the high- range water reducers was sufficient to cause segregation.

85 3.5.2 Air Entrainment Agents Air-entraining agents improve workability, particularly for lean or harsh mixtures or mixtures with angular or poorly-graded aggregates. The presence of entrained air results in a concrete that is more cohesive; however, excessive entrained air contents can make concrete sticky and difficult to finish. Air entrainment also reduces segregation and bleeding (Kosmatka, Kerkhoff, and Panarese 2002). Tattersall (1991) showed that the use of air-entraining agent in concrete reduced plastic viscosity to a much greater extent than yield stress. The change in plastic viscosity was essentially zero above an air content of 5%, although the yield stress continued to decrease at higher air contents. Likewise, Mork (1996) suggests that, in general, low dosages of air-entraining agent mainly reduce plastic viscosity, while higher dosages mainly result in reductions in yield stress. In cement paste, air entrainment can increase yield stress (Tattersall and Banfill 1983). This increase is thought to be due to the apparent negative charge imparted on the air bubbles by the air entrainment agent. This negative charge can attract hydrating cement grains, resulting in the formation of bridges between the cement grains. In concrete, the reduction in plastic viscosity is likely due to the “ball bearing” effect of the spherical air bubbles. The yield stress of the concrete is not decreased as significantly as the viscosity due to the increase in yield stress of the cement paste.

3.5.3 Viscosity Modifying Admixtures Viscosity modifying admixtures (VMAs), also know as anti-washout admixtures, are typically used in self-consolidating concrete or for placing concrete underwater. For self-consolidating concrete, VMAs are used to improve stability by reducing segregation, surface settlement, and bleeding. In underwater concrete, VMAs reduce the washout mass loss. VMAs increase both the yield

86 stress and plastic viscosity. A thorough overview of VMAs and their effects on concrete is provided by Khayat (1998). A range of VMAs with various chemical compositions are commercially available. VMAs used for concrete typically consist of water soluble polymers, such as welam gum or cellulose derivatives. Typically, these VMAs increase the viscosity of the mixing water through a variety of mechanisms, with the precise mode of action depending on the type of polymer. The use of a VMA results in shear-thinning, or pseudoplastic, behavior in cement pastes or mortars. This behavior is advantageous for concrete because the relatively high viscosity at low shear rates prevents segregation of aggregates while the relatively low viscosity at higher shear rates ensures excellent deformability during mixing, pumping, and placing operations. VMAs also increase thixotropy.

3.6 EFFECTS OF SUPPLEMENTARY CEMENTITIOUS MATERIALS

3.6.1 Fly Ash The use of fly ash is generally recognized to improve the workability of concrete by reducing the water content needed to achieve a certain slump. In terms of rheology, fly ash reduces yield stress but has variable effects on plastic viscosity. The influence of fly ash depends on whether the cement is replaced with fly ash on a mass or volume basis. Two conflicting mechanisms influence the workability of concretes with fly ash. Since fly ash particles are smaller than the cement particles they replace, the surface area that must be wetted increases, resulting in a reduction in workability. However, the spherical shape of fly ash particles produces a “ball bearing” effect, which allows coarser particles to flow more readily and improves workability.

87 Tattersall (1991) showed that the use of a mass replacement of fly ash in concrete mixtures resulted in a reduction of yield stress, while the plastic viscosity decreased only slightly. The magnitude of reduction in yield stress depended on the initial cement content, with fly ash having the greatest improvement at lower initial cement contents. When fly ash was replaced on a volume basis instead of a mass basis, the changes in yield stress and plastic viscosity were doubled, suggesting that the increased surface area played a larger role in the incremental difference in volume between the mass and volume replacements. Szecsy (1997) found that a 10% fly ash mass replacement level in concrete mixtures resulted in an increase in yield stress. From 10 to 20%, the use of fly ash reduced the yield stress. The use of 5% fly ash resulted in a reduction of plastic viscosity; however, further replacement of cement with fly ash at rates up to 20% resulted in little additional change in plastic viscosity.

3.6.2 Silica Fume The use of silica fume can improve workability when used at low replacement rates but can reduce workability when added at higher replacement rates. The addition of 2 to 3% silica fume by mass of cement can be used as a pumping aid for concrete (Tattersall 1991). Like fly ash, the spherical shape of silica fume particles is advantageous for workability; however, the small diameter of silica fume particles can significantly increase the surface area that must be wetted. According to Tattersall (1991) and Mork (1996), a threshold value of the silica fume replacement level exists for concrete mixtures, such that below the threshold value, the use of silica fume reduces plastic viscosity but produces little change in yield stress. Above the threshold value, both yield stress and plastic viscosity increase with increasing levels of silica fume replacement.

88 Faroug, Szwabowski, and Wild (1999) measured the rheology of concrete with the silica fume used as either a replacement or addition to cement. When used as a replacement, the yield stress increased with increasing replacement levels up to 20%, above which further silica fume replacement resulted in a reduction in yield stress. The plastic viscosity decreased at up to a 10% replacement rate, but then began increasing at higher replacement rates so that the plastic viscosity was approximately unchanged from the control at a 15% replacement rate and higher than the control at further replacement rates up to 30%. When used as an admixture at levels up to 10%, silica fume resulted in increased yield stress across the tested range. Plastic viscosity increased at addition levels up to 7.5%, above which it began to decrease. Shi, Matsui, and Feng (2002) tested mortar mixtures and found that the addition of silica fume resulted in a reduction in both yield stress and plastic viscosity at replacement rates up to 6% and 9%, respectively. At higher rates, yield stress and plastic viscosity increased, such that at a 12% replacement rate, both yield stress and plastic viscosity were higher than the control mixture.

3.6.3 Ground Granulated Blast Furnace Slag The use of ground granulated blast furnace slag (GGBFS) generally improves workability, although its effects can be variable depending on the characteristics of the concrete mixture in which it is used. According to Tattersall (1991), the effect of GGBFS on workability is much less than that of fly ash for cases when a constant slump is maintained. Tattersall (1991) reported results showing that the effect of GGBFS on rheology was strongly dependent on the cement content and GGBFS type. For a low cement content (200 kg/m3), the addition of slag reduced yield stress and increased plastic viscosity for the two slags, which were used at replacement rates of 40% and 70%. At a higher cementitious materials content (400 kg/m3) the use 89 of the first slag resulted in minimal change in rheology while the use of the other slag resulted in increases in yield stress and plastic viscosity. The water content was held constant when the cementitious materials content was changed; therefore, the water-to-cementitious materials ratio decreased as the cementitious materials content increased.

3.7 EFFECTS OF FIBERS The use of steel or synthetic fibers can dramatically decrease concrete workability and increase thixotropy. Tattersall (1991) showed that increasing the content of steel and synthetic fibers resulted in increases in both yield stress and plastic viscosity. For the steel fibers, increasing the fiber length resulted mainly in an increase in yield stress but little change in plastic viscosity.

3.8 SUMMARY The factors influencing concrete rheology and workability are summarized in Table 3.1. Rheology depends on the concentration, shape, and particle size distribution of the various solid constituents as well as the use of chemical admixtures. Due to the wide variation in materials available for concrete production and the infinite number of possible combinations of these materials, the information contained herein applies only to general cases. For specific combinations of materials, trial batches can be tested to confirm trends.

90 Table 3.1: Summary of Factors Influencing Concrete Rheology Yield Stress Plastic Viscosity Cement Content Decrease Decrease Water Content Decrease Decrease Aggregates Aggregate Volume Fraction Increase Increase Sand-to-Aggregate Ratio Optimum value Optimum value Round or cubical preferred to flat, Shape elongated, or angular Texture Smooth preferred to rough Uniform gradation, high packing density Gradation preferred Microfines Content Mixed Mixed Admixtures Water-Reducing Admixtures Decrease Mixed Air Entrainment Agent Mixed Decrease Viscosity Modifying Admixture Increase Increase Supplementary Cementitious Materials Fly Ash Decrease Mixed Silica Fume (low dosage) Decrease Decrease Silica Fume (high dosage) Increase Increase GGBFS Mixed Increase Fiber Reinforcement Increase Increase

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CHAPTER 4: EVALUATION OF POTENTIAL APPROACHES TO WORKABILITY CHARACTERIZATION

4.1 INTRODUCTION Despite nearly a century of efforts to develop an effective test method to measure workability quickly and accurately, much work remains to be done. The 61 existing workability test methods described in ICAR Report 105.1 (Koehler and Fowler 2003) provide a basis for identifying the principles and concepts that have and have not worked in the past and for developing requirements for a new test method. This chapter describes the lessons learned from existing test methods; the feedback obtained from industry, government, and academia; general requirements for a new test method; and the identification of the most promising approach for workability characterization.

4.2 ASSESSMENT OF EXISTING WORKABILITY TEST METHODS The 61 test methods described in ICAR Report 105.1 (Koehler and Fowler 2003) provide more information on what has not worked than on what has worked. Still, an examination of these methods helps identify needs addressed by earlier researchers and prevents dead ends in future research. Throughout the literature, numerous references quickly rule out the possibility of using a field workability test method that does not directly measure yield stress and plastic viscosity. However, no one has proposed a simple, rugged, inexpensive field test to measure the fundamental rheological parameters of concrete. The slump test has endured for nearly 90 years because of its simplicity and accuracy. In fact, some might say that the concrete industry has become

93 complacent in measuring workability due to the simplicity of the slump test. On large construction projects where substantial sums of money are spent on producing and placing concrete, the amount of money spent to measure a property as important as workability is insignificant. Still, in order for the concrete industry to spend more money to monitor workability, a compelling alternative to the slump test in terms of cost, accuracy, and usefulness must be developed. The development of new high-performance concretes will make the need for more accurate field characterization of concrete rheology more important. This need for an improved test method along with new understanding of concrete rheology should propel the adoption of a test method that is more complex than the slump test as long as it is economically viable and provides accurate and relevant results. It is perhaps most instructive to study the slump test because of its extensive use over more than 80 years. The slump test has even been used as a simple indication of yield stress for concentrated suspensions other than concrete (Pashias et al. 1996). The lessons drawn from the slump test must not be overlooked in the future development of workability test methods. The slump test has endured longer than any other test method because of its simplicity and the relevance of its results. Various references cite different dates for the first introduction of the slump test (Chapman 1913; Abrams 1918). The slump test was first standardized into ASTM in 1922. Even at the time when the slump test was developed, the test’s simplicity was one of its key selling points. No other test is as simple, fast, and reliable as the slump test. One of the main disadvantages of the slump test, however, is that it is a static, not dynamic, test. Dynamic tests provide an indication of the flow properties of concrete after the yield stress has been exceeded and are better suited for thixotropic materials such as high-microfines concrete. Any new test method should be comparable to the slump test in terms of simplicity, speed, and

94 accuracy, while also giving a better indication of the dynamic properties of concrete under actual placing conditions. Despite criticism that the results of the slump test can be too easily “adjusted” to meet a specification, the results of the slump test have been shown to be sensitive to changes in material properties. To evaluate the ability of the slump test to detect changes in material properties, Baker and McMahon (1969) compiled the testing, sampling, and material variances for the slump test on eleven construction projects from two states. Testing variance is a measure of testing errors and uncertainties in the test; sampling variance is a measure of errors and uncertainty in the sampling process; and material variance is a measure of changes in material properties. In an ideal test, all of the variance will be attributable to material variance. The testing variances as a percentage of total variance, as reported in Table 4.1, compare favorably to other commonly used test procedures. As a percentage of total variance, the pressure method for determining air content (ASTM C233) has a 21.5% testing variance, the 28-day compressive strength test (ASTM C39) exhibits a 16.8% testing variance, and the standard unit weight determination (ASTM C138) has a 51.0% testing variance (Popovics 1994). Aside from the slump test, the remaining workability test methods can be split into two groups: tests that have been used in the concrete industry but not to the extent of the slump test and tests that have never gained acceptance beyond any initial studies. The most notable test methods that appear to have found some use by the concrete industry include the compaction factor test (Powers 1968; Neville 1981; Bartos 1992; Bartos, Sonebi, and Tamimi 2002; BS 1881-103 1993), Vebe consistometer (Bartos 1992; Scanlon 1994; Bartos, Sonebi, and Tamimi 2002; ASTM C 1170; EN12350-3), Kelly ball test (Powers 1968; Bartos 1992; Scanlon 1994; Ferraris 1999; Bartos, Sonebi, and Tamimi 2002), inverted slump cone test (Tattersall and Banfill 1983; McWhannell 1994; Johnston 1994;

95 ASTM C 995-01; Bartos, Sonebi, and Tamimi 2002), flow table test (Tattersall 1991; Bartos 1992; Bartos, Sonebi, and Tamimi 2002; EN12350-5), K-slump tester (Bartos 1992; Scanlon 1994; Ferraris 1999; US patent 3,863,494; Bartos, Sonebi, and Tamimi 2002), and trowel test (Bartos 1992; Dobrowlski 1998; Bartos, Sonebi, and Tamimi 2002). Existing rotational rheometers have gained limited acceptance. While some other devices do show promise and may be used more extensively in the future, at the time of this writing they appear to have gained little acceptance by the concrete industry.

Table 4.1: Variability in Slump Test Measurements in Highway Applications (Baker and McMahon 1969) Testing Overall Variance Testing Sampling Material Project Observations Standard Mean as % of Variance Variance Variance Deviation Total Variance 1 184 0.16 0.04 0.26 0.68 2.44 35% 2 200 0.13 0.02 0.45 0.80 1.50 22%

State 1 3 300 0.25 0.09 0.46 0.89 2.76 31% 1 216 0.074 0.00 0.15 0.47 2.04 33% 2 200 0.06 0.06 0.37 0.70 1.86 12% 3 200 0.08 0.025 0.42 0.73 2.34 15% 4 200 0.027 0.012 0.206 0.495 1.77 11% 5 204 0.066 0.03 0.305 0.633 2.37 16% State 2 6 200 0.033 0.034 0.14 0.456 2.12 16% 7 200 0.084 0.086 0.20 0.609 2.41 23% 8 200 0.158 0.047 0.50 0.844 2.26 22%

Of these seven test methods other than the slump test that have been used in the concrete industry beyond a few initial studies, all are single-point tests with at least one major disadvantage to the slump test. The compaction factor test has been used more extensively in Europe than in the United States, although overall its use has declined. While the compaction factor test is a dynamic test and does provide more information than the slump test, it is bulky for common site use and

96 the value of its results are not well understood. The Vebe consistometer is used for very low-slump concretes; however, it is too large for site use and is not typically used in cases where the slump test can be used. The Kelly ball test has been shown to be at least as accurate as the slump test, yet it has never gained widespread acceptance and has been discontinued as an ASTM standard test method. The inverted slump cone test shows promise in properly dealing with the thixotropy of fiber-reinforced concrete; however, the device needs to be better developed in order to measure a wider range of concretes and to allow comparison of test results between different individual test apparatuses. The K- slump tester is commercially available and can be slightly faster than the slump test; however, it is not as accurate as the slump test. The trowel test is useful; however, due to its subjectivity, it will likely never be specified or used on a widespread basis. Rotational rheometers are still used primarily in the research domain; none has been extensively marketed to contractors as a field test device. The concepts used by rotational rheometers do show promise for field use. However, with the exception of the BTRHEOM rheometer (Hu et al. 1996; de Larrard et al. 1997; de Larrard 1999; Ferraris and Brower 2001; Bartos, Sonebi, and Tamimi 2002), all are too bulky for site use. Even the BTRHEOM rheometer, which requires a computer and includes seals that must be replaced frequently, needs to be made more rugged for field use. All of the devices are far too expensive, even for most laboratory work. If any of the rotational rheometers could be developed in a smaller, more rugged form and on a mass production basis with inexpensive parts, the cost and practicality of such a rotational rheometer could be competitive with the slump test, given the additional information such a test device could provide.

97 4.3 FEEDBACK FROM INDUSTRY, GOVERNMENT, AND ACADEMIA Early in the project, a workshop was held to solicit feedback on existing workability test methods and needs for improved workability characterization. Workshop participants included representatives from government, industry, and academia. The workshop considered the full range of workability, with a particular focus on low-slump concrete. After the workshop, a statement of conclusions was prepared and sent to all workshop participants for review. The following conclusions were reached: • For concrete in general, it is not necessary to measure directly fundamental flow properties as long as the properties measured are relevant to the construction operation. • The measurement of yield stress and plastic viscosity is not relevant to dry-consistency concrete mixtures (with slumps less than approximately 2 to 4 inches). Instead, three factors are of greater relevance: o The energy applied to initiate movement of the concrete. o The rate of flow of concrete, expressed as the energy required to move a certain mass of concrete. o The energy required to achieve full consolidation. • Alternatively, a soil mechanics approach may be appropriate for low- slump concrete mixtures. • Although it is important to measure flow properties under vibration because such measurements are representative of actual field conditions, the meaning of viscosity under vibration is still not well understood. Indeed, the BTRHEOM rheometer, which can only be used for mixtures with slumps greater than 4 inches, is the only currently available rheometer capable of measuring viscosity under vibration without special modifications.

98 • The use of performance-based specifications will not eliminate the need to monitor workability. Ensuring proper workability is important for achieving good strength and durability. • Finishability is an issue that should be considered separately from workability. It is related to bleeding, aggregate characteristics, and the rheology of the paste at the concrete surface.

Additionally, the following assessments were made of the prospect for a portable concrete rheometer: • The prospects for building a portable rheometer are promising. The IBB rheometer, which features an impeller geometry, is capable of successfully measuring a wide range of concrete from dry sand to self- compacting concrete. Master Builders has used the IBB successfully for several years. Developing a portable field device, however, presents different challenges. Unlike the IBB, the portable drill device would not be set in a fixed position above a standard volume of concrete. • It may be possible to utilize a properly selected axial impeller to measure low-slump concrete mixtures. The existing portable drill device (FCT 101), in its current form, features an axial rotation, whereas the IBB features a planetary rotation. • It is likely with low-slump concrete mixtures that the scatter of the data would be large. • If yield stress and plastic viscosity are not relevant for low-slump concrete mixtures, using an impeller device may not be appropriate for such mixtures.

99 4.4 CRITERIA FOR NEW WORKABILITY TEST METHODS Based on the advantages and disadvantages of existing workability test methods and the feedback from the workshop, criteria for the creation and evaluation of future test methods were developed. Any new test method should provide a more complete description of workability than the slump test is capable of providing and be competitive with the slump test in terms of speed and economic viability. The criteria for any new workability test method are described as follows: • Parameters Measured: Any new test method should directly or indirectly measure yield stress and plastic viscosity. The test method should measure the dynamic properties of moderate- and low-slump concretes and should appropriately measure concretes that exhibit high thixotropy. To accomplish this, the test should add energy to the concrete, such as with vibration. • Ruggedness: Any new test device must be sufficiently rugged to be used regularly on a jobsite. Depending on the accuracy of the device, it may also be used in the lab for research and mix proportioning. • Workability Range: Any new test method should be able to measure the widest possible range of workability. The wider the range of workability, the more versatile the device will be and the greater the chances that the device will be adopted widely. In reality, no device can measure all concretes, from zero-slump concrete to self- consolidating concrete. • Aggregate Size Restrictions: The device must feature proper geometry to allow testing of concretes with a wide range of aggregate sizes. Based on existing tests, such as the slump test, the device should measure concretes with a maximum aggregate size of up to 1 to 1.5 inches.

100 • Sample Size: The sample size should be kept to a minimum while still being sufficiently large to determine test parameters accurately. • Cost: The cost of any device, when mass-produced, must be competitive with simple, currently available devices, most notably the slump test. • Test Duration: The time required to perform the test must be minimized. The slump test can be performed in several minutes. Other tests allow workability to be monitored continuously with little interruption of construction operations. • Complexity/Training: Any new test device must be sufficiently simple to be performed and interpreted by field workers. Although the test may report results in terms of yield stress and plastic viscosity, field personnel not familiar with concrete rheology must be able to interpret these values and make quick decisions. The use of nomographs or an embedded electronic device can facilitate the interpretation of results in the field. • Number of People Required to Perform Test: One person should be able to perform the test method quickly. This person should also be able to perform other duties on the job site, instead of only monitoring workability. • Data Processing: The results of the test should preferably be obtained directly without any calculations or processing. When data processing is required, an embedded electronic device should perform all calculations and display simple results that can be used directly. • Size and Weight: The device must be small and lightweight so that one person can easily move it around on a jobsite. • Electricity: Although any new test device should preferably be able to operate without electricity, devices requiring power should not be

101 eliminated. Many construction sites have power readily available. Alternatively, batteries can be used.

Most importantly with any device, it must be accepted by a wide range of parties within the concrete industry. As such, the device must satisfy the seemingly conflicting objectives of being simple and rheologically accurate. Concrete contractors will not decide to use a new test unless it clearly adds value to their construction operations. Researchers have been skeptical of simple devices that give a relevant indication of workability but do not directly measure the fundamental rheological properties of concrete. For instance, in discussing the inverted slump cone test for fiber-reinforced concrete, Tattersall and Banfill (1983, 238) write, “it is extremely unfortunate that in a new area of concrete technology it is proposed to establish yet another empirical and quite arbitrary test for workability; the long-term result can only be to add to the confusion which already exists.” A device that meets the majority of the criteria enumerated above stands the greatest chance of being adopted by all diverse parties in the concrete industry.

4.5 SELECTION OF THE MOST PROMISING APPROACH FOR WORKABILITY CHARACTERIZATION Based on the examination of the state of the art in workability measurement and the current and future needs of the concrete industry vis-à-vis workability, it was decided that a low-cost, portable rheometer would be the most promising approach for measuring concrete workability in the field. Since no such rheometers existed, it was decided to develop and test the ICAR rheometer, a completely new portable rheometer designed specifically to meet the changing needs of the concrete industry. As originally envisioned, the ICAR rheometer would be approximately the size of a hand drill and would feature a properly

102 designed impeller that could be immersed into a sample of concrete on a jobsite and rotated at various speeds in order to determine the rheological properties. The concept of the ICAR rheometer fulfills the requirements for new test methods described in Section 4.4. The ICAR rheometer can measure the scientific fundamental flow properties of fresh concrete. The ICAR rheometer can be a dynamic test by adding energy to the concrete and measuring the flow properties after the yield stress has been exceeded. As a portable device, the ICAR rheometer can be constructed to be rugged, lightweight, and compact. It can be designed to measure a wide range of workability with typical maximum aggregate sizes. The sample size can be kept to a reasonable value, given the size of concrete constituents. If properly designed, the cost of building the ICAR rheometer on a mass-production basis can be significantly less than the cost of existing laboratory concrete rheometers and can be economically viable relative to the slump test given the value of the information provided by the ICAR rheometer. The operation of the test can be designed so that it requires little training and can be completed by one person in less time than the slump test. A software program can be used to automate the operation of the test and the presentation of test results. The use of a rheometer can simplify the process of workability characterization by eliminating the need for inconsistent, subjective measurements of workability. Although operation of a rheometer requires electrical power, the ICAR rheometer can be operated from a battery, just as a portable hand drill would be operated. From this initial concept, the development process of a first generation prototype was commenced, as described in detail in the next chapter. The criteria described in Section 4.4 were considered throughout the development and testing process. The thorough consideration of both the lessons learned from past efforts in characterizing workability and the current and future needs of the concrete industry helped to enable the creation of a workability test method that balanced

103 the needs of the concrete industry with the challenges and limitations of measuring concrete workability.

104

CHAPTER 5: DEVELOPMENT OF THE ICAR RHEOMETER

5.1 INTRODUCTION With the decision made to build a portable rheometer and after a thorough study of the literature regarding fluid rheology and concrete workability characterization, the development process for the ICAR rheometer began. While the basic requirements were known, it was necessary to find a combination of components that could fulfill the objectives for the ICAR rheometer in terms of cost, capabilities, accuracy, and portability. This chapter describes the development process in detail, from the evaluation of available components to the selection of components and the development of the software program.

5.2 OVERVIEW OF DEVELOPMENT PROCESS The objective of the development process was to demonstrate the feasibility of producing and using a portable rheometer for characterizing the workability of concrete. To this end, a first generation prototype was developed using all off-the-shelf components. With each decision in the development process, attention was paid not just to selecting acceptable off-the-shelf components, but to assessing the feasibility of developing a mass-production version of the rheometer. The use of off-the-shelf components resulted in a prototype that was larger and more expensive than a future mass-production version; however, it dramatically shortened the development time and cost and permitted the experimental testing of concrete to commence sooner.

105 The development of the ICAR rheometer consisted of three major aspects: development of an impeller, selection of a motor and gearbox, and development of a control system. The impeller selection was perhaps the most important aspect. An impeller was needed that would create favorable flow conditions in the concrete in order to enable the determination of rheological parameters. The motor and gearbox needed to be selected to produce sufficient torque and to be compact and lightweight. Finally, a simple and reliable software program and associated electronics were needed to automate and control the testing process and presentation of results. As a first step, preliminary minimum design requirements were developed for the ICAR rheometer. These requirements, which are listed in Table 5.1, were based on the expected operating range of the rheometer, the operating characteristics of existing laboratory-grade rheometers, a preliminary survey of available technology, and the specific workability test device requirements enumerated in Chapter 4.

Table 5.1: Preliminary Design Requirements for ICAR Rheometer Factor Design Requirements Maximum Torque 50 Nm (442 in-lb) Rotation Speed 0 – 1.2 rev/sec Battery Voltage 18 – 24 VDC Size As compact as possible Entire operation automated to the greatest extent Control practical As small as possible while still generating Impeller representative flow in concrete and minimizing segregation

With these minimum requirements set, the search for proper components began. The first approach was to attempt to modify a commercially available hand drill to operate as a rheometer. This approach is described in Section 5.4. When

106 the components in existing hand drills were found to be unsuitable for a concrete rheometer, an entire rheometer prototype was built from scratch. Individual off- the-shelf components were procured and assembled. The initial configuration of the prototype was tested in concrete to determine its practicality and functionality. Based on this testing, which is described later in Chapter 6, the design was modified further to achieve the final design of the first generation prototype. The final design of the first generation prototype was tested in a series of concrete mixtures, as described in Chapter 7. While the prototype was found to be suitable, it will need to be modified further in the future to create a mass- production version. As with any product in any industry, there is always opportunity for further development and improvement of the rheometer. Although the rheometer was developed for measuring concrete, it could likely also be used for measuring the rheology of other similar coarse-grained suspensions.

5.3 SELECTION OF POTENTIAL IMPELLERS The key factors to consider in selecting an impeller are the movement of the concrete generated by the impeller and the dimensions of the impeller relative to the aggregate size. The geometry must be selected to avoid both segregation and trapping of aggregate. The impeller must be of sufficient size so that gap sizes are a proper multiple of the maximum aggregate size and so that the increment of torque generated at each higher speed increment can be measured accurately. Still, the torque generated should not be too large in order to avoid requiring an unreasonably large rheometer motor. It has been shown that a rheometer’s geometry, gap size, and surface friction can have a significant influence on the measurement of the rheological properties of cement paste (Rahman and Moncef 2003). Adding fine and coarse aggregate further complicates the determination of rheological parameters.

107 Therefore, careful attention must be paid to each aspect of the design of an impeller.

5.3.1 General Concepts The tendency of axial impellers when used in dry-consistency concrete mixes is to push aggregate away from the blade, creating a channel in which the impeller spins. Planetary impellers can still cause segregation or aggregate trapping; however, such action does not influence results to the same extent because the impeller is intended to come into contact constantly with concrete that has little or no shear history. The shape and speed of the impeller strongly influence the nature of the generated mixing action and, therefore, the degree of segregation or aggregate trapping that occurs during the test. Various studies have been conducted to determine the minimum ratio of gap size to aggregate size. Ratios ranging from 3 to 10 have been suggested. Although no one size has been officially accepted, most concrete rheometers feature gaps less than 4 times the maximum aggregate size. A comparison of gap sizes is presented in Table 5.2. Results from testing with these rheometers are generally accepted by concrete rheology researchers as being reasonable (Ferraris and Brower 2001). The size requirements for rheometers using a higher gap to aggregate size ratio would be impractical. Only one rheometer, the 500-liter CEMAGREF-IMG coaxial cylinders rheometer (Ferraris and Brower 2001), is capable of satisfying the gap size to maximum aggregate size ratio of 10 for concrete. The ratio of its outer radius to inner radius in the CEMAGREF-IMG, however, is too large to ensure that a dead zone will not occur in the concrete.

108 Table 5.2: Gap Sizes in Existing Concrete Rheometers Reported Max Gap Gap Size Rheometer Geometry Aggregate Size Location (Inches) Size Ratio 1.5 in. 0.5 in* Coaxial Between FHPCM (38.1 (12.7 3.0 Cylinders Cylinders mm) mm) BML Coaxial Between 1.77 in. 0.98 in** 1.8 Viscometer Cylinders Cylinders (45 mm) (25 mm) Coaxial Bertta Between 2.95 in. Cylinders Apparatus Cylinders (75 mm) (Oscillatory) 8.66 in. CEMAGREF- Coaxial Between (220 IMG Cylinders Cylinders mm) Blades to 1.85 in.

Tattersall Two- Impeller Sides (47 mm) point – Helix (Axial) 2.17 in. Blade Width (55 mm) Horizontal 2.13 in. Tattersall Two- Impeller Inside H (54 mm) point – H- (Planetary) Impeller to 2.0 in. Shaped Sides (50 mm) Horizontal 2.13 in. 2.1 Impeller Inside H (54 mm) 1.0 in.*** IBB Rheometer (Planetary) Impeller to 1.97 in. (25 mm) 2.0 Sides (50 mm) Vertical 4.0 in. BTRHEOM Height or 1.0 in.*** Parallel Plate (100 4.0 Rheometer Horizontal (25 mm) mm) Space *Yen et al. (1999) **Bartos, Sonebi, and Tamimi (2002, 3). It is also reported that larger aggregates may be used, “dependant on the mix.” ***Ferraris and Brower (2001)

Much of the early research into appropriate gap size was based on using coaxial cylinders rheometers for ceramic particle suspensions. Van Wazer et al. (1963) state that for a coaxial cylinders viscometer used to measure particle suspensions, the gap between the outer and inner cylinders should be ten times the

109 maximum particle size. Wall effect, or the fact that packing density decreases as the ratio of container dimension to particle size decreases, is a key consideration in determining this minimum gap size. McGreary (1961) conducted experiments to determine the packing density of mechanically packed mono-sized particles in different sized containers. This research indicated that the packing density was essentially independent of the ratio of container size to particle size when this ratio was at least 10. The presence of blades in coaxial cylinders rheometers limits the risk of slippage due to wall effect. The influence of slip has not been fully explored with impeller and parallel plate rheometer geometries. For coaxial cylinders concrete rheometers, it has been found that a gap dimension to particle size ratio of at least 5 is acceptable (Ferraris 1999). The work based on coaxial cylinders rheometers has also been extended to other rheometer geometries. A study done with the BTRHEOM rheometer to examine the effect of gap sizing found that when the ratio of vertical gap to maximum aggregate size was set to less than 4, the effect of gap size on the Bingham parameters became significant (Ferraris and Brower 2001). As a counterpoint to the above requirements, the practicality of any device must be considered. In addition to accuracy requirements, concrete rheometers must be small, rugged, and economically viable. It is also useful to consider the dimensions of common concrete elements, where concrete will be required to flow and where rheology measurements will be directly relevant. ACI 211.1 requires that the maximum aggregate size be no greater than one-fifth the narrowest dimension between forms, one-third the depth of slabs, and three- fourths the clear spacing between reinforcing bars. Concrete in chutes and pumping pipes will rarely have dimensions greater than 10 times the maximum aggregate size. Therefore, concrete can rarely be considered a homogenous fluid subjected to simple gradients under actual flow conditions in the field.

110 Although it is still desirable to know the fundamental properties of a homogenous fluid, clearly a compromise must be made when measuring concrete in order to have a practical test device. Given the need for practicality and the fact that measurements made with existing rheometers typically fit the Bingham model and are consistent with expectations, rheology experts have come to accept gap dimension to particle size ratios of 3 to 5.

5.3.2 Evaluation of Existing Impellers Two existing laboratory concrete rheometers, the Tattersall two-point device and the IBB rheometer, feature impeller geometries. A study of the development of the impellers for these rheometers is instructive in developing a new impeller to measure the widest possible range of workability. Other potential impellers are currently available. The following subsections describe each of the currently available impellers. Since it is now generally recognized that rheological measurements of the same concrete will vary between rheometers due in part to differences in geometry and surface friction, Ferraris and Martys (2003) have recommended a method to relate these different measurements by determining a relative viscosity. Thus, it is realized that no rheometer is likely to make true “fundamental” measurements and that additional effort will be required to compare the result from one rheometer to another.

5.3.2.1 Tattersall Two-Point Device The Tattersall two-point device was reportedly the first rheometer to measure concrete as a Bingham fluid. The device was developed after other researchers had made unsuccessful attempts to measure concrete workability using coaxial cylinders rheometers. The development of the two impellers for the

111 Tattersall device is described by Tattersall and Banfill (1983) and summarized in the following paragraphs. The original version of the two-point device, the Mk I apparatus, was first presented in 1971 and consisted of a Hobart food mixer with a hook-shaped impeller moving in a planetary motion. This original test device is shown in Figure 5.1. The intent of the design was to use an impeller that would move in planetary motion and continuously come into contact with fresh concrete. Speed was altered by changing the gear settings while torque measurements were made indirectly with a wattmeter. The three speed settings were 95, 170, and 310 revolutions per minute. The measured flow curves, when used on mixes with water-to-cement ratios from 0.40 to 0.70, were generally linear.

Figure 5.1: Mk I Apparatus (Tattersall and Banfill 1983)

112 The planetary motion and hook impeller are standard features on commercially available kitchen mixers. The hook blade is typically intended for dough. The geometry of the hook varies widely between models. Later work carried out with a larger Hunt mixer capable of operating at six different speeds showed that the hook impeller was suitable for low- to moderate- slump concretes. The device, however, could not be used for high workability mixes because the increment of power required when the bowl was loaded versus when it was empty was too insignificant to make precise rheological measurements. To overcome this problem, a larger impeller that would produce a higher torque was developed. The “square anchor” impeller was selected for this purpose after an examination of several other impeller shapes and sizes. This new impeller, shown in Figure 5.2, worked well for high-slump mixes. It also worked on lower-slump mixtures, although it produced significantly higher power readings than the original impeller.

Dimensions: mm

Figure 5.2: Square Anchor for Mk I Apparatus (Tattersall and Banfill 1983)

The results of the tests using the square anchor impeller matched the trends for the original hook. Anomalous results were obtained for highly fluid 113 mixes. These results were attributed by Tattersall and Banfill (1983) to the nonlinearity of the torque/power relationship and to the occurrence of turbulence in the concrete. While the Mk I apparatus was effective in establishing concrete as a Bingham fluid, additional work was needed to produce accurate and reliable results. Based on additional research, the Mk I apparatus was later replaced by the Mk II apparatus for high workability mixes and the Mk III for low to medium workability mixes. The goal of designing the impeller for the Mk II apparatus was to utilize axial motion with a greater range of speeds than available with the Mk I apparatus. The impeller selected for the Mk II, shown in Figure 5.3, is an interrupted helix turning in axial motion. This design was selected over a helical screw impeller. The presence of gaps between each flat blade is important for stiffer concretes because it allows concrete displaced by the blades to be replaced. If a helical screw impeller were used, concrete would move upward to the top of the impeller and not be fully replaced. The interrupted helix impeller was used for rotation speeds from 0.25 revs/sec up to 1.25 revs/sec. The impeller was found to work for mixes with slumps greater than 4 inches. When used for concretes with slumps generally less than 4 inches, the interrupted helix impeller of the Mk II pushed concrete to the outside of container and then rotated in the empty pocket. Therefore, an impeller operating in a planetary motion was selected for use with low-slump mixes. The interrupted helix impeller was used in planetary motion; however, dry-consistency concrete still moved to the outside of the container. The next attempt used a two-blade impeller, depicted in Figure 5.4, but also produced unsatisfactory results. The final attempt incorporated an H-shaped impeller moving in planetary motion, shown in Figure 5.5, and is considered the Mk III apparatus. This impeller, along

114 with the planetary gearbox, fits directly into the existing Mk II apparatus. A larger bowl is required when using the H-shaped impeller.

Dimensions: mm

Figure 5.3: Mk II Apparatus and Interrupted Helix Impeller (Tattersall and Banfill 1983)

Dimensions: mm

Figure 5.4: Two-Blade Impeller (Tattersall and Banfill 1983)

115

Dimensions: mm

Figure 5.5: H-Shaped Impeller for the Mk III Apparatus (Tattersall and Banfill 1983)

The gearboxes used for the Tattersall two-point device and the IBB rheometer to create planetary motion are simple designs that are easy to install. In designing such a gearbox, the gear ratio should be a non-integer so that when the central shaft makes a full revolution, the impeller returns in a different alignment than in the previous revolution. The offset from the central shaft should be large enough to avoid aggregate trapping but not so large that the required concrete container is excessively large. The planetary motion gearbox used in the Tattersall device and the IBB rheometer have gear ratios greater than one, resulting in an increase in speed and decrease in torque from the motor to the impeller.

5.3.2.2 IBB Rheometer The IBB rheometer (Beaupre and Mindness 1994; Bartos, Sonobi, and Tamimi 2002), shown in Figure 5.6, is a modification of the Tattersall two-point 116 test. The major components of the IBB rheometer are very similar to those of the Tattersall device. The IBB rheometer includes a planetary H-shaped impeller that is identical to the Tattersall impeller. The IBB rheometer does not use the interrupeted helix impeller. It is reported that the IBB rheometer is capable of measuring the full range of concrete, from approximately a slump of 1 inch up to self-consolidating concrete (Bartos, Sonebi, and Tamimi 2002).

Figure 5.6: IBB Rheometer and Impeller

5.3.2.3 Fresh Concrete Tester (FCT 101) The fresh concrete tester (FCT 101) was developed in Germany (Steiner 1996) and is sold commercially by at least two companies in Britain. The device, shown in Figure 5.7, is not a true rheometer because it only operates at one speed. Product literature indicates that the results of the test are related to slump.

117

Figure 5.7: FCT 101 (from Product Literature)

The impeller consists of two hemispheres on opposite sides of a shaft that moves in axial rotation. According to product literature, the device is capable of measuring concrete with slumps ranging from 7 to 250 mm and maximum aggregate sizes ranging from 7 to 32 mm. An assessment conducted by Wong et al. (2000) using a larger version of the FCT 101 impeller showed that the device was capable of predicting slump. When used to measure low-slump concrete, the impellers created a channel through the concrete; therefore, only data from the first revolution were meaningful.

5.3.2.4 Commercially Available Mixing Impellers An assortment of mixing blades designed to fit into drills is commercially available for mixing fluid materials ranging from latex paint to mortar and granular materials. Though most are not intended for concrete, the mixers are

118 intended to be simple, while providing efficient mixing. The mixers could be used to mix material in containers ranging in size from a standard 5-gallon bucket to a 55-gallon drum. The impellers are designed to mix the material efficiently and thoroughly and are not intended to cause segregation when used appropriately. It is likely that the user would move the impeller around within the material to ensure complete mixing. An impeller for mixing mortars and drywall spackling is shown in Figure 5.8. The geometry is simple—the round bars do not continuously direct material flow in one direction. When used for concrete, both segregation and aggregate trapping are possible.

Figure 5.8: “Egg Beater” Mixing Impeller for Mortar

A drum mixer intended for mortar, plaster, and granulates is shown in Figure 5.9. According to product packaging, the impeller can be operated at a maximum speed of 500 rpm. Smaller versions are available for lower viscosity materials, such as paints and inks. The blades are designed so that when the impeller rotates in the clockwise direction, material in the center of the impeller moves upwards while the material surrounding the impeller moves downward.

119

Figure 5.9: Drum Mixer for Mortar, Plaster, and Granulates

A mixer intended for drywall joint compound, drywall texture, and paint is shown in Figure 5.10. The manufacturer suggests a speed range of 450 to 800 rpm. The blades are designed to lift, fold, and mix the material.

Figure 5.10: Mixing Impeller for Drywall Joint Compound, Drywall Texture, and Paint

120 5.3.3 Impeller Selection Although several different rheometer impellers have been developed, it is clear that the statement of Tattersall and Banfill (1983) that “an improved form may yet be developed” remains true. If possible, the impeller should move in an axial motion, thereby avoiding the cost and complication of a planetary gearbox. The impeller should be of minimal size, while meeting a gap size to aggregate size ratio of 3 to 5 for 1-inch aggregate. The movement of the concrete generated by the impeller should not result in either segregation or trapping of aggregate. In order to select an impeller, it was necessary to test experimentally several impeller geometries in concrete. The evaluation of nine potential impellers, developed based on the information in this section, is described in Chapter 6. The final impeller selected was a four-bladed vane intended to act as the inner cylinder of a coaxial cylinders rheometer.

5.4 EVALUATION OF CONVENTIONAL HAND DRILL TECHNOLOGY Before selecting the components of a rheometer, the prospect of modifying an existing hand drill to operate as a rheometer was evaluated. Conventional hand drills are compact, powerful, battery operated, and readily available. In order to be used as a rheometer, a hand drill would need to be modified by adding sensors to measure torque and rotation speed, appropriate electronics to regulate rotation speed, and a data acquisition system to record sensor measurements. An impeller would be attached with the chuck. If a kit could be developed to modify an existing drill, the cost and development time of such a system would be less than the creation of an entirely new device and the kit could be a viable commercial product. A wide range of battery-operated hand drills is commercially available. Battery-operated hand drills typically include a brushed DC motor and a transmission (gearbox). The transmission gear ratio is usually manually adjustable

121 in order to vary the speed and torque for the given application, such as drilling a hole or driving a screw. Additionally, many drills allow speed to be varied further at a constant gear ratio based on how far the user pulls back on the trigger. Battery voltages commonly range from 9 VDC to 18 VDC, with several models operating from 24 VDC batteries. The models with higher voltage typically produce higher peak torques. For a typical 18 VDC drill, the peak torque can range from 350 to 500 inch-pounds at the maximum gear ratio. The maximum rotation speed is typically 400 to 500 rpm at the maximum gear ratio and 1,500 to 2,000 rpm at the minimum gear ratio. The main problem with modifying a hand drill to serve as a rheometer is the motor. The motors in many of the commercially available drills are able to provide sufficient torque to turn an impeller in concrete; however, the motors are designed to deliver this torque at high speeds. The typical low speed setting on a drill produces a speed that is approximately 5 to 7 times the maximum speed needed for a flow curve measurement in a rheometer. It would not be feasible to drastically reduce the voltage supplied to the motor in order to reduce the torque. The speed could be reduced by replacing the original motor with different motor designed to turn at slower speeds or by replacing the original transmission with a different transmission with a larger gear ratio. Both options would likely increase the size of the drill and would make the simple modification of a drill impractical. Therefore, it was decided to design and build a rheometer from the beginning with individual components selected and configured specifically for the rheometer. Despite the fact that conventional hand drill technology could not be readily modified, the evaluation of existing hand drill technology demonstrated that high torque can be delivered in a compact package and from battery power.

122 5.5 SELECTION OF COMPONENTS The main components of the first generation prototype of the ICAR rheometer are shown in Figure 5.11. The motor and gearbox provide the necessary ranges of torque and rotation speed. The motor used is a brushed DC servomotor that can be operated from 18 VDC or 24 VDC battery power. The inline planetary gearbox, which features a 50-to-1 gear ratio, reduces the rotation speed and increases the torque supplied by the motor. A ½-inch keyless chuck allows the impeller to be quickly attached to the rheometer. Rotation speed is measured with an incremental optical encoder mounted to the end of the motor. A non-contact inline torque transducer is connected with couplings between the gearbox and the keyless chuck. The torque transducer is based on magnetoelastic technology, which is more reliable than the slip ring torque transducers that Cabrera and Hopkins (1994) found to be problematic when used in the Tattersall MK II apparatus. Compared to other non-contact technologies, the magnetoelastic technology is generally smaller and less expensive. The aluminum frame and the bearing protect the torque transducer and gearbox from large lateral forces being transmitted from the impeller shaft. A handle is provided so that the rheometer can be operated by hand. The components in the first generation device exceed the requirements of Table 5.1 for rotation speed and torque capacity. The wires shown in Figure 5.11 return to a separate electronics box. In a later generation prototype, it is envisioned that all components, included the electronics, will be housed together. The electronics box consists of a DC speed controller and a compact USB-based data acquisition system. The DC speed controller is an open-loop system that regulates the DC voltage supplied to the motor based on a separate analog signal voltage. The associated computer software, described in Section 5.6, completes the closed-loop system, which is needed to adjust for changes in speed due to changes in load. The USB-based data acquisition system provides excitation voltage to and accepts signals from the

123 encoder and torque transducer and provides the analog signal voltage to the speed controller. The USB data acquisition system is connected to a laptop computer, which is used to automate the test operation and presentation of results.

Encoder Handle

Torque Transducer DC Motor Bearing Keyless Gearbox Chuck

Couplings

Aluminum Frame Figure 5.11: First Generation Prototype of ICAR Rheometer

The power for the rheometer can be supplied from a DC battery or an AC/DC converter. While the battery enables field use, the AC/DC converter is convenient for lab testing where the rheometer will be used on a frequent basis over a long duration. The prototype shown in Figure 5.11 can be operated by hand or positioned above a standard container. If the rheometer is to be operated by hand, the operator uses both hands to grasp the rheometer and resist the reaction torque generated as the impeller is turned. The concrete can be contained in any volume such that there is sufficient distance from the impeller to all boundary surfaces. Possible examples of appropriate containers include a standard 5-gallon bucket found on many jobsites, a wheelbarrow, a portable mixer, or a hopper. A depth

124 indicator rod can be provided to ensure the impeller is immersed to the correct distance. This set-up most nearly matches the original concept of a drill device. Alternatively, the components shown in Figure 5.11 can be mounted to a frame, which is subsequently positioned above a container of standard size, as demonstrated in Figure 5.12. The use of this alternative set-up makes the operation of the test easier for the user and ensures that the test is performed in a consistent manner. This alternative set-up is still sufficiently rugged and portable for use on a jobsite. Laboratory testing with the first generation prototype indicated that improved accuracy is achieved when the rheometer is mounted in a fixed frame and placed above a container of standard size. The rheometer simply slides into the frame and is attached securely. The frame then slides downward into slots in the top of the container. The slots resist rotation of the rheometer/frame assemblage. The container may include multiple slots to allow the rheometer impeller to be immersed to different depths. Provided the container is placed on approximately level surface, the mass of the portable rheometer is sufficient to maintain the position of the rheometer/frame assemblage above the container.

125

Figure 5.12: ICAR Rheometer Mounted in Frame and Positioned above Container

In order to ensure the rheometer is set above the center of the container, optional vertical positioning rods attached to each end of the frame slide through channels on the sides of the container, as shown in Figure 5.13. Concrete may be placed in the container before or after the rheometer/frame assemblage is placed above the container. An external or internal vibrator may be used to consolidate the concrete in the container or to measure the rheology of concrete under vibration. For the first generation prototype, a schedule 80 PVC pipe was used for the container. This choice of material enabled containers of different sizes to be built quickly. The mass-production version would likely be constructed of a sturdy, lightweight metal.

126

Figure 5.13: Positioning Rod in Channel

The components selected for the first generation prototype may not be the best or only available solutions. If the rheometer were produced in larger quantities, it is likely that most or all of the off-the-shelf components used in the first generation prototype would be replaced with custom parts made specifically for the ICAR rheometer. Other aspects of the design could be completely changed. For instance, torque could be measured by monitoring the power drawn by the motor instead of with an in-line transducer. The optical encoder could be replaced with a tachometer. For the mass-production version of the ICAR rheometer, the components shown in Figure 5.11 will likely be reduced in size and weight and enclosed in a case. Figure 5.14 shows a rendering of how such mass-production version may look. The components from the first generation prototype, or similar

127 replacements, would all fit into a plastic case, which would also include a handle and power source. A handheld computer (PDA) or similar embedded electronics would be used instead of a laptop to automate the operation of the device and presentation of test results. The speed control and data acquisition electronics could be placed within the case. The entire mass-production version shown in Figure 5.14 could still be mounted into a frame and secured above a standard container, just as in the first generation prototype.

Encoder Power Handle Source Motor

Gearbox

Torque Case Transducer

Keyless Chuck

First Generation Mass-Production Version

Prototype Figure 5.14: Rendering of Prototype and Envisioned Mass-Production Version

128 5.6 SOFTWARE The software to operate the rheometer and compute test results was written with LabVIEWTM 6.1 from National Instruments, Austin, TX. LabVIEWTM 6.1 is an object-oriented programming language for test and measurement applications. The rheometer software can be operated on any Microsoft Windows operating system using the freely available run-time version of LabVIEWTM 6.1. As a consequence, the marginal cost of installing the software on additional computers is zero. The versatile and inclusive software enables testing operations to be fully automated. The software can be used—with only minimal modifications—for later generation prototypes consisting of different components.

5.6.1 Graphical User Interface The graphical user interface, shown in Figure 5.15, is intuitive and user- friendly. The test inputs and results are shown on a single screen to ensure that all information is visible to the user. When information is placed in multiple menus, it may become easy for the user to overlook an important input. Although the screen may at first appear cluttered with numerous inputs, the information is laid out in distinct sections and in simple steps. The top portion of the software consists of the test settings. In the first step, the user selects the type or types of tests to be performed. If a direct yield stress test is selected, the rheometer will perform a stress growth test. For the absolute flow curve, the Bingham parameters will be calculated based on the Effective Annulus Method for the vane geometry (cf. Section 2.3.2.4). In addition, relative units of yield value and viscosity value will also be displayed. The yield value is defined as the intercept of the straight line fit to the measured torque versus rotation speed data, while the viscosity value is the slope of this line. In computing the relative parameters, no corrections are made for the

129 presence of a dead zone; therefore, the precautions of Section 2.3.2.3 should be considered. If the relative flow curve option is selected, only the yield value and viscosity value will be calculated. The relative flow curve can be measured for any impeller geometry. It is possible to perform both a yield stress test and then measure either a relative or absolute flow curve.

Figure 5.15: Graphical User Interface for the ICAR Rheometer Software

The second step allows the user to specify a pause time period from the time when the green “GO!” button is pressed until the test actually begins. This pause gives the operator time to either position the rheometer or to move to a

130 location to view the test better. The pause is particularly intended for cases where the rheometer is operated in the hand-held drill configuration. The user can press the “GO!” button and then immerse the impeller into the concrete and prepare to resist the torque generated by the rotation of the impeller. In the third step, the stress growth test options are specified. If a stress growth test is not to be performed, this section may be skipped. The test allows a breakdown period during which the rheometer operates at a constant speed for a specific duration. The impeller can then be stopped for a rest period. If the breakdown time is not needed, the breakdown time and rest time can be set to zero. The test speed can be set to any value that the rheometer components are capable of achieving. For the rheometer components of the first generation prototype, the optimum yield stress speed was determined to be 0.025 rev/sec (cf. Section 6.4), which is set as the default value. In the fourth step, flow curve test options are input. First, the speed and time of the structural breakdown regime are set. Then, the direction of speed points—either ascending or descending—is selected. Finally, the number of points on the flow curve, the time for each point, and the minimum and maximum speeds are set. The software equally distributes the specified number of points between the maximum and minimum specified speeds. Each input has minimum and maximum possible values so that if the user specifies a value outside the operating range of the rheometer components, the input value will be coerced to the appropriate maximum or minimum. The number of speed points must be an integer. The form is pre-populated with typical values as defaults. Next, in step five, the geometry of the impeller is set. These options are required for stress growth tests or absolute flow curve measurements where the results are to be computed in fundamental units. The type of impeller (vane, egg beater, etc.) is selected. For the vane geometry, the height and radius of the vane

131 and the outer radius of the container are specified. Again, the default values are set for the typical vane used in the first generation prototype (cf. Section 6.3.4). The next section of the software consists of the file settings. The user can specify a prefix for file names. The software then appends two values to this prefix. First, the software adds an underbar ( _ ) and a number. The first number added is 1. On subsequent tests for the same file prefix saved to the same folder, the number on the file name is increased by 1 for each test. The use of this number serves two purposes. First, it allows the order of subsequent tests to be identified by their file names. Second, it fully prevents existing files from being overwritten. The software automatically searches the directory where the file is to be saved and locates any files with the specified prefix. If any such files are found, the file with the highest number is identified, and the file to be written is given the next number. If such a system were not in place, the file software would instead need to prompt the user to confirm the overwrite action or to specify a new file name. The approach employed in the software saves time and reduces error. After the prefix and the number, the software then appends either “_raw” or _summary” depending on the type of file written. Therefore a typical sequence of files may appear as: mix1_1_raw mix1_1_summary mix1_2_raw mix1_2_summary mix2_1_raw mix2_1_summary etc…

The user then selects whether to write a summary file, raw data file, or both. The raw data file is a tab-delimited data file with all internal continuous data

132 generated by the software. The columns, from left to right, are: elapsed time (seconds), test type (“yield,” “pause,” or the flow curve speed point number), target speed (rev/sec), actual speed (rev/sec), torque (Nm), and counter value. The counter value is taken from the encoder and is used internally by the software to compute the rotation speed. The raw data file is not typically needed unless the internal operation of the rheometer is to be monitored. A typical summary file is shown in Figure 5.16. The file is saved with a “.txt” extension, so it may be opened quickly with an application such a Notepad in Microsoft Windows. The file consists of the test input options and test results. Only values relevant to the particular test are printed. The first two lines indicate the software version and the file name. The date and time of the test are taken from the internal computer clock. Next, the types of tests performed are shown followed by the test inputs. The yield stress inputs are preceded by “YS” while the flow curve inputs are denoted with “FC”. The geometry is indicated in terms of the type of impeller and the dimensions of the impeller and container. Numerical test results are printed for the yield stress test, relative flow curve, and absolute flow curve, as applicable. Finally, the flow curve points are listed in two columns, which are tab delimited to facilitate copying into a spreadsheet program.

133

Figure 5.16: Typical Test Summary File

With all test inputs set, the test may be started. The “Test Operation” section consists of three main buttons. The “Zero Torque” button resets the zero torque voltage offset used in the software. Depending on the electronics used, this value may vary slightly from test to test. The electronics in the mass production version should be selected to minimize any fluctuations. In the electronics used in the first generation prototype, fluctuations were rare and typically were no more

134 than 0.4 Nm. Still, it is helpful to click the zero offset button between each test. The zero offset button should only be clicked when the rheometer impeller is in air. If the impeller has been inserted into the material in the container, a net torque may be imposed on the impeller by the at-rest material. To begin the test, the user clicks to “GO!” button. If a direct yield stress test has been selected in Step 1 of the “Test Inputs” section, the stress growth test will be performed first. Otherwise, the flow curve test is started. For the stress growth test, a real-time torque versus time plot is displayed on the screen. The user must monitor this plot. Once it is clear that the peak torque has been reached, the user should press the “Done” button to stop the stress growth test. The flow curve test, if requested in Step 1, then automatically starts. The flow curve test is fully automated and requires no user intervention. During testing, the green “Test Status” light is illuminated. The test can be stopped immediately at any time by clicking the “STOP!” button. After the “STOP!” button has been selected, it turns black and must be clicked a second time to be deselected. Throughout the test, the target speed setting for either the yield stress or flow curve measurement is displayed. For the flow curve, the elapsed time of the test is displayed. When the speed must be changed quickly, such as at the beginning and end of a test or between a yield stress test and a flow curve test, the speed setting is linearly ramped over a period of 2.5 seconds. None of the data recorded during these periods is used to compute test results. This precaution protects the rheometer components from spikes in torque. If the speed is immediately increased or decreased by a significant amount, the dynamic effects can significantly increase the torque in the rheometer, possibly damaging rheometer components. When the testing is finished, all results are shown in the “Test Results” section. Three types of numerical test results may be shown. First, yield stress

135 results are shown in terms of yield torque, which is the maximum torque achieved during the stress growth test, and yield stress, which is the yield torque converted into stress based on the vane geometry. The equation to convert yield torque to yield stress is based on the assumption that the shear stress on the side and ends of the vane is uniformly distributed and equal to the yield stress (cf. Section 6.6). If a relative flow curve is measured, the yield value (Y), viscosity value (V), and R2 value will be shown under the “Relative Parameters” heading. If an absolute flow curve is measured, the yield stress, plastic viscosity, and mean squared error will be shown under the “Bingham Parameters” heading and the yield value, viscosity value, and R2 value will be shown under the “Relative Parameters” heading. Any informative or specific error messages are shown in the appropriate message box. General errors are indicated in the error box at the bottom of the window. The measured flow curve points are plotted on the graph of torque versus rotation speed. The straight line for the relative parameters (uncorrected) is plotted in red while the line for the corrected absolute parameters is displayed in green. The graph is automatically scaled to fit the given data. At the start of any tests, all results are reset to zero; the final results are not shown until all testing is completed.

5.6.2 Internal Software Operation A full description of the internal operations of the software is well beyond the scope of this report. The speed is controlled with a PID (proportional, integral, derivative) closed-loop system. This effective and widely used approach to closed-loop systems is easily implemented into LabVIEWTM. The PID loop must be “tuned” for the selected components and for concrete. Once tuned, the PID loop is effective across essentially the full range of concrete rheology. A typical speed versus time plot for a flow curve is shown in Figure 5.17. Due to the presence of 136 aggregate, some scatter is expected (as shown as deviation from the black solid line); however, the average speed over a second or longer consistently equals the desired speed to within less than 1%. When the speed is changed from one flow curve point to the next, the actual speed slightly overshoots and then stabilizes. To account for this stabilization time, the first 0.8 seconds of torque and rotation speed data are discarded each time the rotation speed is changed.

1.2

0.8

0.6

0.4 Rotation Speed (rev/sec) Rotation Speed

0.2

0 0 5 10 15 20 25 30 35 40 45 50 55 Time (Seconds)

Figure 5.17: Typical Speed versus Time Plot for Flow Curve Measurement in Concrete

The software obtains readings from the data acquisition system at 20 Hz. In order to protect the motor, gearbox, and torque transducer, the software displays an error message and automatically stops the test if a certain maximum torque is exceeded.

137

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CHAPTER 6: EXPERIMENTAL DETERMINATION OF OPERATING CHARACTERISTICS OF THE ICAR RHEOMETER

6.1 INTRODUCTION With the basic design of the rheometer complete, it was next necessary to determine experimentally certain operating characteristics of the rheometer in order to complete the first generation prototype. First, an impeller that would minimize segregation, engage surrounding material, and enable an accurate determination of rheological parameters was selected. Then, using this impeller, the optimum speed for determining yield stress in a stress growth test was determined. The horizontal and vertical gap sizes between the impeller and container boundaries were assessed. Finally, the magnitude and distribution of shear stress acting on the ends of the impeller were evaluated. This chapter describes each test that was performed, the results of these tests, and the final design selections.

6.2 MATERIALS, MIXTURES, AND TEST PROCEDURES Four different concrete mixtures, covering a range of fresh concrete characteristics, were used for the tests described in this chapter. The mixture proportions are shown in Table 6.1. The materials used in these mixtures are described in greater detail in Chapter 7. The coarse aggregate for the first two mixtures was a river gravel (RG), while mixes 3 and 4 utilized a crushed limestone coarse aggregate (LS). Both mixtures incorporated a natural sand and ASTM C 150 Type I portland cement. The water content was varied slightly for each mixture to suit the needs of the testing. 139

Table 6.1: Mixture Proportions for Chapter 6 Tests Target Batch Mass, kg/m3 (lb/yd3) Coarse Fine Mix Slump Coarse Fine Type Type Cement Water (Inches) Aggregate Aggregate 1104.3 681.5 361.7 180.9 Mix 1 7 RG NS (1861.3) (1148.7) (609.7) (304.9) 1128.8 696.6 369.8 162.7 Mix 2 3 RG NS (1902.7) (1174.2) (623.3) (274.2) 1015.5 690.8 385.2 181.1 Mix 3 7 LS NS (1711.6) (1164.5) (649.3) (305.2) 1043.6 710.0 395.9 158.4 Mix 4 3 LS NS (1759.1) (1196.8) (667.3) (266.9)

The concrete mixtures were prepared in general accordance with ASTM C 192. All materials were batched at least 12 hours prior to mixing using the batching procedures described in Chapter 7. The concrete was mixed in a rotating drum mixer. For mixing, the coarse aggregate and a portion of the mixing water were added to the mixer, which was then started to blend the aggregates. Next, the portland cement followed by the rest of the mixing water was added to the mixer. The concrete was mixed for three minutes, allowed to rest for three minutes, mixed for another two minutes, and discharged into a wheelbarrow. The slump was measured during the rest period and the water content was adjusted to ensure that the slump would be within the target range. Immediately after concrete was discharged from the mixer, the slump test was performed. Concrete was loaded by shovel into the rheometer container and testing was started as soon as practical. Multiple rheometer tests were preformed for each concrete mixture. The time limit for testing each concrete mixture was 60 minutes from the initial contact of cement and water. Slump was again measured at the end of all testing for each concrete mixture.

140 6.3 IMPELLER TYPE Six potential impellers, shown in Figure 6.1, were tested to identify the impeller or impellers that minimize segregation and create a favorable flow pattern in the surrounding concrete. The vane impeller, which measures 5 inches in diameter and 5 inches in height—is mainly intended to function in a manner analogous to the inner cylinder in a coaxial cylinders rheometer. As such, the use of a vane would enable the ICAR rheometer to serve as an absolute rheometer so that the Bingham parameters could be calculated in fundamental units. The egg- beater impeller is a commercially available impeller for mixing mortars. The egg- beater impeller was also cut in half—to create the half egg-beater impeller—so that it would create less torque and would not trap a plug of concrete within the four outer bars. The offset egg-beater impeller is a modified version of the half egg-beater impeller and is designed so that if a channel is created behind one bar, the second bar will arrive at a different location and displace concrete back into the channel created by the first bar. Each side of the offset egg-beater has the same surface area; therefore, the net lateral force in the impeller should be zero. The joint compound paddle is a large impeller designed mainly to mix materials without aggregates. Likewise, the spiral impeller is mainly intended to mix paints. In total, four different spiral impellers were tested. The fact that the joint compound paddle and the spiral impeller are designed to create certain flow paths for mixing may make them suitable for testing concrete. The potential impellers were tested in three series of tests which are described in the following three subsections. The six impellers shown in Figure 6.1 were tested in the first two series; four different spiral impellers were tested in the third series.

141

Vane Egg-Beater

Half Egg-Beater Offset-Egg Beater

Joint Compound Paddle Spiral Figure 6.1: Potential Impellers

142 6.3.1 Series I Tests – Qualitative Observations In the first series of tests with the impellers, an attempt was made to visualize experimentally the flow generated by each impeller. The flow in the concrete was visually observed with each impeller partially covered and fully covered by concrete. Each impeller was rotated at two speeds, 1.5 rpm and 40 rpm. For the first test for each speed, each impeller was allowed to protrude 1 inch above the concrete surface (“UNCOVERED”). In the second test, each impeller was covered with 4 inches of concrete (“COVERED”). For all tests, the bottom cover was 10 inches while the container diameter was 16 inches. The two river gravel concrete mixtures with different water content were tested: the high-slump mix had a slump of 5 inches while the low-slump mix had a slump of ½ inch. Tables 6.2 and 6.3 contain qualitative observations of the flow behavior for each impeller. In the descriptions, the cut volume is defined as the volume of concrete inside the outer boundaries of the impeller. Although the observations are qualitative, they provide key insights into the flow behavior around each impeller and provide part of the basis for selecting the final impeller. Photographs of the concrete around each impeller for the uncovered cases are shown in Figure 6.2 and 6.3 for mixes 1 and 2, respectively.

143 Table 6.2: Qualitative Observations for Mix 1 (5-Inch Slump) Vane UNCOVERED: Rotation of the vane creates an open void behind the vane. The void begins near the tip of vane blade and increases in size with time. The final void volume is not cylindrical in shape, but approximately defined by a linear boundary between the vane tips at any given instant. Although the vane does engage some concrete outside the cut volume, minimal flow occurs in this outer zone. A portion of concrete being pushed by the advancing side of the vane creates a force, which is approximately linear, pointing radially outward from the vane and acting on the concrete outside the cut volume. Finally, there appears to be some wall effect inside the cut cylinder adjacent to the blade vanes, resulting in a mortar-rich zone. This wall effect is exacerbated by shearing—at the end of shearing, mainly mortar remains within the cut volume. Therefore, the reduction in torque over time is due structural breakdown, the creation of a channel by the cut volume, and the creation of four mortar pockets within the vane. COVERED: During shearing, a conical indication is created above the vane as concrete from above the vane falls into the void that results when concrete is pushed outside of the cut cylinder. All material at the surface appears to be in motion locally; however, there does not appear to be any meaningful amount of flow over time. Egg-Beater UNCOVERED: Rotation creates a distinct channel behind the bars. The creation of this channel eventually results in an isolated plug of concrete that rotates with the impeller (as opposed to remaining stationary). Some aggregates from inside the plug do try to move outward. There is significantly less engagement of the concrete surrounding the egg-beater than in the case for the vane. The wall effect problem with the vane impeller is not a problem for the egg-beater impeller. COVERED: At the low speed, some material above the blade settles. This settlement is not evident at the high-speed setting. There does appear to be some engagement of the aggregate. Half Egg-Beater UNCOVERED: Rotation creates a circular channel behind the bars, eventually resulting in a plug that reduces in volume as shearing continues. The approximately linear radial force acting from the bars is fairly significant compared to the other impellers. COVERED: There is some engagement of the aggregates. The location of impeller bars at any instant is evident from viewing the top of the concrete. Offset Egg-Beater UNCOVERED: Rotation creates loose material within the cut volume; however, a distinct plug is not created. The larger radius bar creates some flow in the area immediately outside the cut volume; the shorter radius bar does not. COVERED: There does appear to be some flow evident from surface. After some shearing, spiral flow lines are visible on the surface. Joint Compound Paddle UNCOVERED: Rotation creates a very large channel as the angle on the blade attempts to reduce the volume of concrete passing through its openings. Some material from outside the cut volume flows inward but is pushed back out when the blade returns. COVERED: Maximum torque exceeded, no observations made. Spiral UNCOVERED: Rotation does create a void in the cut volume. Although some material in the cut channel is replaced, the spiral impeller does not engage flow. COVERED: Very little flow effects are visible.

144 Vane Egg-Beater

Half Egg-Beater Offset-Egg Beater

Joint Compound Paddle Spiral Figure 6.2: Flow of Concrete around Impellers – Mix 1 (5-Inch Slump) 145

Table 6.3: Qualitative Observations for Mix 2 (1/2-Inch Slump) Vane UNCOVERED: At the low rotation speed, it takes several rotations to create a channel. Some material initially sticks to the blade; however, after enough shearing all material within the cut volume is eventually removed. Egg-Beater UNCOVERED: Like the vane impeller, the bars of the egg-beater create an approximately linear force acting radially outward from the vane. A separate plug of concrete within the egg- beater forms quickly. Eventually, there is a reduction in the volume of this plug. Many spikes in the torque occur during shearing. Half Egg-Beater UNCOVERED: As a void is created within the cut volume, some materials from outside the cut volume falls into this void. There is some loosening of the material and an associated dilation. At the high speed setting, lateral and axial forces are generated in the impeller shaft, causing the rheometer to move around within the channel. Offset Egg-Beater UNCOVERED: Creates a loosening and expansion in volume. Rheometer blade moves around substantially. Due to unequal lengths, there is a preference for the rheometer to “walk” within the concrete towards the looser material. Joint Compound Paddle UNCOVERED: The main effect of rotation is a loosening of concrete. This loosening creates a loose zone and a stiffer zone. The continual turning of the blade toward the stiffer zone results in large lateral forces in the impeller. As a result, the test was quickly stopped. Spiral UNCOVERED: An overall reduction in the volume of concrete with this spiral occurs, although some material does fall in behind bars. At the higher speed, concrete moves out of the cut volume much faster and little torque is generated.

146 Vane Egg-Beater

Half Egg-Beater Offset-Egg Beater

Joint Compound Paddle Spiral Figure 6.3: Flow of Concrete around Impellers – Mix 2 (1/2 Inch Slump)

147 For the vane impeller, the visual observations confirmed that the vane engaged material outside the vane and created a cut volume, which is analogous to the inner cylinder of a coaxial cylinders rheometer. As expected for a coarse- grained suspension, wall effect did reduce the coarse aggregate volume concentration within the cut volume. Provided the ratio of the size of the cut volume relative to the maximum aggregate size is sufficiently large, the wall effect should have a negligible effect on test results because the spatial distribution of aggregates will be approximately uniform at the boundary of the cut cylinder. The fact that voids were created within the cut volume during shearing—as represented in Figure 6.4—limits the minimum range of workability that can be tested at a given rotation speed. The fact that the cut volume was not circular, but followed a linear shape, was due to the fact that some material from outside the cut volume flowed back inside during shearing. The conical indentation that occurred above the vane when it was fully covered indicated that vertical movement of material occurred and likely resulted in the replacement of material in the voids behind the vane.

Volume cut by vane more closely follows a linear shape than a cylindrical shape between blade tips.

Void Concrete adjacent to end of blade pushes tangentially outward on concrete outside Concrete of void.

Figure 6.4: Flow around Vane in Uncovered Case

148 The behavior of the egg-beater impeller was similar to the vane; however, the egg-beater impeller cut out a distinct cylindrical plug of concrete for both high and low-slump concrete mixtures. While the plug moved at approximately the same speed as the vane, the channel created behind the blades of the egg-beater impeller meant that the inner plug would only act in a manner analogous to the inner cylinder of a coaxial cylinders rheometer for fluid mixes. The observations of the covered egg-beater impeller indicated that the impeller engaged much of the concrete outside the cut volume. The rotation of the half egg-beater impeller did not result in substantial mixing of the material or the creation a distinct inner plug. Rather, the impeller loosened material as it turned, especially at low slumps. Like the half egg-beater impeller, the offset egg-beater impeller created a loosening of concrete and an associated expansion in material volume. The outer bar appeared to dominate the deformation process. It loosened a region of material as it turned, including the region where the inner bar passes. As a consequence, greater torque was contributed by the outer bar than the inner bar. The creation of lateral forces in the shaft was evidence that the forces on the two bars were not balanced. The joint compound paddle created too much torque due to its size and geometry. Because of the angle of the blades, the rotation of the impeller attempted to reduce the volume of the concrete as it passed through the impeller. As a result, torque was substantially increased. The spiral impeller appeared to be too small to engage flow in the surrounding concrete. In the low-slump concrete mixture, concrete was removed from inside the spiral, leaving behind an empty void. The uncovered cases did not fully represent the flow in the rheometer because they neglected any vertical flow of material that may occur in actual test conditions. While all impellers displaced concrete and left voids of various sizes,

149 any concrete present vertically above an impeller could presumably flow to replace any displaced concrete. Upon removal from the low-slump concrete (1/2-inch), four of the impellers left a clean void, as shown below in Figure 6.5. The concrete within the void was either expelled during shearing or formed a plug that adhered to the impeller or became trapped within the impeller. Whereas the vane and half egg- beater impellers left relatively rough void side walls, the egg-beater and spiral impellers left smooth, paste-rich surfaces.

Vane Egg-Beater

Half Egg-Beater Spiral Figure 6.5: Voids Left by Impellers – Mix 2

150 Based on these qualitative observations of flow behavior, the vane and egg-beater impeller appeared to create the most favorable flow patterns. The spiral, half egg-beater, and offset egg-beater impellers appeared to be unsuitable. With the exception of the joint compound paddle, larger versions of any of the impellers relative to the maximum aggregate size may provide more favorable flow patterns in the concrete. For the joint compound paddle, a smaller version may be acceptable.

6.3.2 Series II Tests – Quantitative Performance In the second series of tests, the performance of each of the six impellers was tested quantitatively. The properties of the four concrete mixtures used in this testing are shown in Table 6.4. Flow curves were measured with each impeller in order to assess the following factors: degree of structural breakdown, effect of flow curve measurement sequence (ascending or descending speeds), and the quality of the flow curve measurement.

Table 6.4: Concrete Properties for Series II Impeller Type Tests Slump before Slump after Final Elapsed Mix Testing Testing Time (Inches) (Inches) (Hr:Min) Mix 1 6 6 1:00 Mix 2 3.5 3.5 1:00 Mix 3 7.5 7.5 0:59 Mix 4 3.5 3.5 0:51

The first factor considered was the structural breakdown of the concrete that occurs when shearing is started in the rheometer. This structural breakdown occurs in all rheometers and is not unique to the ICAR rheometer. A typical plot of torque versus time for an impeller rotating at a constant speed is shown in Figure 6.6. The torque initially increases linearly up to a maximum value and then 151 drops—sharply at first and gradually later—to an average steady-state value. The magnitude of torque reduction and the time for it to take place are important considerations in selecting an impeller. While the reduction in torque is typically due mainly to structural breakdown, it can also be due to segregation. Therefore, assessing the reduction in torque provides an indication of the presence of segregation. An impeller that exhibits a substantially greater reduction in torque than the other impellers is likely to be creating more segregation.

8 Torque (Nm)

0 0 102030405060708090 Time (seconds)

Figure 6.6: Typical Reduction in Torque over Time (Rotation Speed = 40 rpm)

In order to quantify the degree of structural breakdown and judge the relative amount of segregation, each impeller was rotated at a constant speed of 1.0 rev/sec for 90 seconds. The vane was tested in both the 10-inch and 16-inch containers. The other impellers were tested only in the 16-inch container. The percentage change in torque from the peak torque achieved in the first few

152 seconds to the steady torque achieved after 90 seconds of shearing was determined. The percentage reduction in torque, as indicated in Table 6.5 and plotted in Figure 6.7, ranged from 45 to 94%. The reductions in torque in the half egg-beater and offset egg-beater impellers were consistently above average for each mix. On average for the four mixtures the vane and egg-beater impellers provided slightly below average reductions in torque. The spiral impeller produced significantly lower than average reductions in torque in two of the three mixtures in which it was tested and slightly below average reduction in torque in the other mixture.

Table 6.5: Average Percentage Reduction in Torque for Each Impeller Impeller Mix 1 Mix 2 Mix 3 Mix 4 Vane (10" Container) -- 71.6% 83.1% -- Vane (16" Container) 72.1% 68.9% 80.2% 80.8% Egg-Beater 75.6% 73.1% 79.3% 75.5% Half Egg-Beater 80.3% 82.4% 94.2% 88.5% Offset Egg-Beater -- 79.5% 89.5% 84.8% Joint Cmpd. Paddle -- -- 73.5% -- Spiral -- 44.7% 81.4% 62.3% Average 76.0% 70.0% 83.0% 78.4%

153 100%

90%

80%

70% Vane (10") 60% Vane (16") EB 50% HEB OEB 40% JCP Spiral 30%

Percentage Reduction in Torque in Reduction Percentage 20%

10%

0% Mix 1Mix 2Mix 3Mix 4

Figure 6.7: Percentage Reduction in Torque for Each Mixture

It is also useful to consider the total torque generated after breakdown, which is shown in Table 6.6 and Figure 6.8. If the torque measured by the impeller is too low, the resolution in torque measurements may be insufficient for accurate flow curve measurements. Further, the low torque may be an indication that only a small portion of material is flowing. If the torque is too high, the range of concrete workability that can be measured will be limited. The joint compound paddle clearly created too much torque—it was unable to measure three of the four mixtures. The spiral, half egg-beater, and offset egg-beater impellers produced torques significantly below the other impellers. Interestingly, the offset egg-impeller produced significantly higher torques than the half-egg beater, which suggests that the offset egg-beater impeller engages more material than the half- egg beater impeller.

154

Table 6.6: Torque after 90 Seconds of Breakdown Torque, Nm Impeller Mix 1 Mix 2 Mix 3 Mix 4 Vane (10-Inch Container) -- 8.23 3.45 -- Vane (16-Inch Container) 3.74 5.56 2.20 4.76 Egg-Beater 3.34 4.95 1.82 3.99 Half Egg-Beater 2.57 2.11 0.56 1.99 Offset Egg-Beater -- 3.30 1.16 3.56 Joint Cmpd. Paddle -- -- 6.61 -- Spiral -- 3.23 0.84 2.44 Average 3.22 4.56 2.38 3.35

6 Vane 10" Vane 16" 5 EB HEB 4 OEB JCP 3 Spiral

Torque After Breakdown, Nm AfterBreakdown, Torque 2

0 1234

Figure 6.8: Torque after 90 Seconds of Breakdown

The next consideration was the quality of the flow curves measured by each impeller. Relative flow curves were measured with each impeller in the four

155 different concrete mixtures. Five points were measured between 0.2 rev/sec and 1.0 rev/sec for each flow curve. One flow curve was measured with descending speeds after 5 seconds of breakdown, while two curves—the first with descending speeds and the second with ascending speeds—were measured after 90 seconds of breakdown. The breakdown speed was 1.0 rev/sec. The average relative parameters for these measurements—namely, yield value and viscosity value— are summarized in Table 6.7. It should be cautioned that some impellers were not used in all tests; therefore, the total magnitude for a given impeller could be distorted. The actual measurements for each mix are shown in Tables 6.8 through 6.11. To judge the quality of each of the flow curve measurements, the degree of linear fit of the torque versus rotation speed data was assessed using the coefficient of determination (R2). Even considering the fact that dead zones occur within the rheometer, the R2 terms should still be close to unity.

156

Table 6.7: Average Relative Flow Curve Parameters for Each Impeller Yield Viscosity Test R2 Value Value (Nm) (Nm.s) Vane (10-Inch Container) 5 sec, descending 3.92 4.24 0.75 90 sec, descending 3.39 2.64 0.93 90 sec, ascending 3.50 1.89 0.79 All 3.47 2.98 0.84 Vane (16-Inch Container) 5 sec, descending 3.22 3.54 0.96 90 sec, descending 2.77 1.50 0.97 90 sec, ascending 2.78 1.29 0.96 All 2.92 2.11 0.96 Egg-Beater 5 sec, descending 3.05 2.61 0.93 90 sec, descending 2.64 1.08 0.84 90 sec, ascending 2.58 0.87 0.71 All 2.76 1.52 0.83 Half Egg-Beater 5 sec, descending 1.59 1.49 0.70 90 sec, descending 1.27 0.78 0.68 90 sec, ascending 1.06 0.83 0.83 All 1.21 0.95 0.68 Offset Egg-Beater 5 sec, descending 3.07 3.80 0.92 90 sec, descending 1.98 1.20 0.84 90 sec, ascending 1.81 1.05 0.85 All 2.29 2.02 0.87 Joint Cmpd Paddle 5 sec, descending 5.38 8.30 0.98 90 sec, descending 4.34 2.76 0.97 90 sec, ascending 3.17 3.67 0.94 All 4.30 4.90 0.96 Spiral 5 sec, descending 1.37 1.03 0.95 90 sec, descending 1.63 0.52 0.85 90 sec, ascending 1.76 0.25 0.43 All 1.66 0.71 0.75

157

Table 6.8: Flow Curve Measurements for Mix 1 Flow Curve Container Breakdown Impeller Direction Yield Viscosity Diameter Time R2 Value Value (Inches) (Seconds) (Nm) (Nm.s) 5 Desc. 4.13 4.20 0.66 10 90 Desc. 2.88 1.93 0.88 90 Asc. 3.16 1.56 0.81 Vane 5 Desc. 2.93 2.94 0.95 16 90 Desc. 2.47 1.10 0.98 90 Asc. 2.38 0.91 0.98 5 Desc. 3.27 2.17 0.91 10 90 Desc. 2.66 1.00 0.84 Egg- 90 Asc. 2.57 1.11 0.99 beater 5 Desc. 2.92 2.24 0.88 16 90 Desc. 2.65 0.55 0.66 90 Asc. 2.15 1.34 0.98 Half 5 Desc. 2.70 2.74 0.92 Egg- 16 90 Desc. 1.69 1.22 0.47 Beater 90 Asc. 0.94 1.56 0.74 Offset 5 Desc. ------Egg- 16 90 Desc. ------Beater 90 Asc. ------Joint 5 Desc. ------Cmpd 16 90 Desc. ------Paddle 90 Asc. ------5 Desc. ------Spiral 16 90 Desc. ------90 Asc. ------

158

Table 6.9: Flow Curve Measurements for Mix 2 Flow Curve Container Breakdown Impeller Direction Yield Viscosity Diameter Time R2 Value Value (Inches) (Seconds) (Nm) (Nm.s) 5 Desc. 6.18 6.37 0.91 10 90 Desc. 5.21 4.50 1.00 90 Asc. 5.31 2.74 0.80 Vane 5 Desc. 4.35 4.53 0.93 16 90 Desc. 4.11 1.86 0.98 90 Asc. 4.32 1.26 0.94 5 Desc. 4.62 3.12 0.98 10 90 Desc. 3.96 1.46 0.99 Egg- 90 Asc. 3.53 1.08 0.69 beater 5 Desc. 1.72 1.42 0.39 16 90 Desc. 1.69 0.75 0.43 90 Asc. 1.65 0.55 0.92 Half 5 Desc. 3.16 4.99 0.97 Egg- 16 90 Desc. 2.77 1.19 0.73 Beater 90 Asc. 2.26 0.90 0.95 Offset 5 Desc. ------Egg- 16 90 Desc. ------Beater 90 Asc. ------Joint 5 Desc. 2.16 1.50 0.95 Cmpd 16 90 Desc. 2.62 0.75 0.73 Paddle 90 Asc. 3.01 0.08 0.03 5 Desc. 6.18 6.37 0.91 Spiral 16 90 Desc. 5.21 4.50 1.00 90 Asc. 5.31 2.74 0.80

159

Table 6.10: Flow Curve Measurements for Mix 3 Flow Curve Container Breakdown Impeller Direction Yield Viscosity Diameter Time R2 Value Value (Inches) (Seconds) (Nm) (Nm.s) 5 Desc. 2.23 3.45 0.96 10 90 Desc. 2.08 1.48 0.93 90 Asc. 2.04 1.36 0.78 Vane 5 Desc. 1.35 2.67 0.98 16 90 Desc. 1.42 1.17 0.98 90 Asc. 1.59 0.77 0.99 5 Desc. 1.58 2.08 0.91 10 90 Desc. 1.01 1.01 0.94 Egg- 90 Asc. 1.12 0.69 0.97 beater 5 Desc. 0.43 0.78 0.67 16 90 Desc. 0.25 0.51 0.88 90 Asc. 0.37 0.27 0.89 Half 5 Desc. 1.63 2.06 0.81 Egg- 16 90 Desc. 0.98 0.53 0.85 Beater 90 Asc. 0.68 0.83 0.87 Offset 5 Desc. 5.38 8.30 0.98 Egg- 16 90 Desc. 4.34 2.76 0.97 Beater 90 Asc. 3.17 3.67 0.94 Joint 5 Desc. 0.59 0.57 0.95 Cmpd 16 90 Desc. 0.64 0.29 0.97 Paddle 90 Asc. 0.52 0.42 0.83 5 Desc. 2.23 3.45 0.96 Spiral 16 90 Desc. 2.08 1.48 0.93 90 Asc. 2.04 1.36 0.78

160 Table 6.11: Flow Curve Measurements for Mix 4 Flow Curve Container Breakdown Impeller Direction Yield Viscosity Diameter Time R2 Value Value (Inches) (Seconds) (Nm) (Nm.s) 5 Desc. 1.45 2.15 0.68 10 90 Desc. ------90 Asc. ------Vane 5 Desc. 4.26 4.03 0.96 16 90 Desc. 3.10 1.88 0.95 90 Asc. 2.83 2.24 0.93 5 Desc. 3.07 3.01 0.96 10 90 Desc. 2.96 1.30 0.79 Egg- 90 Asc. 3.51 0.37 0.19 beater 5 Desc. 1.52 1.02 0.85 16 90 Desc. 1.46 0.65 0.94 90 Asc. 1.28 0.94 0.80 Half 5 Desc. 4.42 4.37 0.98 Egg- 16 90 Desc. 2.20 1.88 0.95 Beater 90 Asc. 2.48 1.44 0.74 Offset 5 Desc. ------Egg- 16 90 Desc. ------Beater 90 Asc. ------Joint 5 Desc. 1.48 1.60 0.90 Cmpd 16 90 Desc. 2.11 0.48 0.50 Paddle 90 Asc. 1.78 0.77 0.86 5 Desc. 1.45 2.15 0.68 Spiral 16 90 Desc. ------90 Asc. ------

The average R2 terms for each impeller are graphed in Figure 6.9 for tests performed in the 16-inch container. The R2 terms were generally highest for the 5- second breakdown time, although both the yield value and viscosity value were distorted from their true values due to the insufficient breakdown time. Specifically, the torque decreased over the duration of the flow curve test not only because the speed was reduced but also because further structural breakdown occurred. Only the vane and the joint compound paddle produced R2 terms greater

161 than 0.90 for the 90-second breakdown tests. The joint compound paddle was used in only one mix; therefore, additional testing is needed with this geometry. The spiral and half-egg beater produced R2 terms below 0.80 for descending curves measured after 90 seconds of breakdown and are, therefore, unsuitable. The egg-beater and offset egg-beater produced R2 values between 0.80 and 0.90.

1 0.9 0.8 5-Second 0.7 Breakdown 0.6 90-Second 0.5 Breakdown, R^2 0.4 Descending 0.3 90-Second Breakdown, 0.2 Ascending 0.1 0 Vane Egg- Half Egg- Offset JC Paddle Spiral Beater Beater Egg- Beater

Figure 6.9: Average R2 Terms for Each Impeller, 16-Inch Container

It should not be implied that a 90-second breakdown time is required for all tests. Such a time would not only be impractical, it would also be unnecessary. A breakdown time of 25 to 30 seconds is sufficient for the vast majority of concrete mixtures. Based on the quantitative data, the vane, when used in the 16-inch container, provided the best results. The vane produced high R2 terms for flow curve measurements, it produced sufficient torque for adequate torque resolution, and its degree of structural breakdown was comparable to the other impellers. The

162 joint compound geometry was promising; however, its size was too large for use in concrete. A smaller version may be suitable. While the spiral impeller produced unacceptably low values of R2 for flow curve measurements, it did produce the smallest reduction in torque, possibly suggesting that it was too small to engage sufficient flow in the surrounding concrete. A larger version of the spiral may produce higher R2 terms while also producing a favorable flow pattern in concrete.

6.3.3 Series III Tests – Spiral Impellers

6.3.3.1 Background After the initial spiral impeller appeared to be too small, three additional spirals were obtained and tested for comparison with the original spiral. The four spirals tested are shown in Figure 6.10. Like before, the purpose of the testing was to identify the impeller that created the least segregation while also creating a favorable flow pattern. Spiral 1, which has a bottom diameter of 4 inches, is the same impeller used in the Series I and II tests. Spiral 2 is identical to Spiral 1; however, the bottom ring was removed. Spiral 3, which has a bottom diameter of 4 5/8 inches, is a larger version of Spiral 1. Unlike the first three spirals, which create a downward direction of flow, Spiral 4 is larger and creates an upward direction of flow. The vertical gap between blades in Spiral 4 is 3 inches, while the outside diameter is 4 5/8 inches. Like Spiral 2, the original ring at the bottom of Spiral 4 was removed.

163

Spiral 1: Spiral 2: Original Spiral Original Spiral, Bottom Removed

Spiral 3: Spiral 4: Larger Version of Original Spiral Alternate Geometry Figure 6.10: Spiral Impellers

The manufacturer’s recommended uses for Spirals 1, 3, and 4 are indicated in Table 6.12. Figure 6.11 shows the manufacturer’s representation of the intended flow direction for each spiral.

164 Table 6.12: Manufacturer’s Recommended Uses for Spiral Impellers Flow Maximum Spiral Recommended Uses Direction Speed Fast mixing paints, inks, Spiral 1 Downward 800 rpm liquids Interior/exterior paints, Spiral 3 Downward 600 rpm latex paints, wallpaper adhesives, concrete sludge Spiral 4 Ready-mix mortar, plaster, (including Upward 500 rpm gypsum pastes, carpet bottom ring) adhesives, granulates

Downward Flow Upward Flow (Spirals 1, 2, 3) (Spiral 4) Figure 6.11: Intended Flow Direction for Spiral Impellers (From Manufacturer’s Packaging)

Qualitative observations and quantitative measurements were made with the four impellers in three concrete mixtures, the properties of which are indicated in Table 6.13. All testing was conducted in the 16-inch diameter container. Concrete was remixed with a shovel between each test.

165

Table 6.13: Concrete Properties for Series III Impeller Type Tests Slump before Slump after Final Elapsed Mix Testing Testing Time (Inches) (Inches) (Hr:Min) Mix 1 5 3 1:00 Mix 2 3.5 3 0:53 Mix 4 3.75 3.5 1:00

6.3.3.2 Qualitative Descriptions of Observed Flow Behavior To assess qualitatively the flow behavior around the potential spirals, each spiral was rotated at two speeds, 1.5 rpm and 40 rpm, and the flow behavior was observed for Mix 1 and 4. For all tests, the bottom cover was 10 inches and the container diameter was 16 inches. The top of each spiral was allowed to protrude 1 inch above the concrete surface. The qualitative observations of the flow behavior for each spiral impeller are listed in Table 6.14. In the forward direction, the spiral impellers were turned clockwise in accordance with the manufacturer’s recommendations. In the reverse direction, the spiral impellers were turned counter-clockwise. Photographs of the resulting concrete flow are shown in Figure 6.12 and Figure 6.13 for the forward and reverse directions, respectively.

166 Table 6.14: Observations for Spiral Impellers Spiral 1 Forward Direction The rotation of the spiral creates a channel behind the blades, which isolates a plug of concrete. At low speeds, some concrete is able to fall back in to the channel. At high speeds, a distinct channel is created quickly and maintained throughout the test. Reverse Direction The rotation of the spiral clearly moves concrete upward and outward, although a plug is still created inside the spiral. The plug of concrete stays relatively stationary as the spiral rotates around it. A portion of the concrete immediately outside the spiral does slump in towards the center of the spiral. A mortar-rich region is created at the top of the concrete surface as mortar is ejected from the upward-moving blades. If the spiral were fully covered, this mortar-rich region would likely remain at the top of the spiral and may not move to the top of the concrete surface. Spiral 2 Forward Direction A plug is created inside the spiral; however, some material within the plug attempts to flow outward. Horizontal movement of the surrounding concrete is visible. Reverse Direction Rotation of the spiral clearly results in a pushing of material upward and outward. Further, the rotation of the spiral appears to pull more aggregates through the impeller as compared to Spiral 1, which pulls more mortar through the impeller. Spiral 3 Forward Direction Due to the angle of the blades, rotation of the spiral does pull some concrete into the center of the spiral. A plug is still created within the spiral. A limited amount of elastic engagement of surrounding concrete is visible; however, very little overall flow is evident. Reverse Direction Rotation of the spiral clearly results in pushing of material upward and outward. A plug is created. Some mortar is pulled out of the concrete and pushed upward by the rotation of the spiral; however, the amount of mortar moving through the spiral is less than the amount for Spiral 1 but more than the amount for Spiral 2. Spiral 4 Forward Direction Rotation of the spiral clearly results in a pushing of the concrete upward and outward. A plug is created. Reverse Direction Rotation of the spiral creates an open void inside the spiral blades as no material is available from above to replace concrete displaced by the spiral.

167 Spiral 1 Spiral 2

Spiral 3 Spiral 4 Figure 6.12: Flow of Concrete around Spiral Impellers – Mix 1 (5-Inch Slump)

168 Spiral 1 Spiral 2

Spiral 3 Spiral 4 Figure 6.13: Flow of Concrete around Spiral Impellers Operated in Reverse – Mix 4 (3.5-Inch Slump)

The flow pattern created by Spiral 1 was distinctively different for each direction of rotation. When Spiral 1 was operated in the reverse direction, mortar was pulled from the concrete and pushed upward to the top of the spiral. Given the fact that the impeller was uncovered, much less resistance to upward flow existed than if the impeller had been covered. If the spiral were larger relative to the maximum aggregate size, it is likely that some coarse aggregates would also

169 flow with the mortar upwards through the spiral. In fact, when the bottom ring was removed for Spiral 2, which eliminated the narrow gap at the bottom of the spiral, more coarse aggregates were moved upward by the rotation of the spiral. Neither Spiral 1 nor 2 appeared to be effective when operated in the manufacturer’s recommended forward direction, although in Spiral 2, some horizontal movement of surrounding concrete was visible. Likewise, Spiral 3 was more effective when operated in the reverse direction; however, it did pull mainly mortar upward when operated in reverse. Spiral 4 was more effective when operated in the forward direction rather than the reverse direction. The forward direction for Spiral 4 resulted in an upward movement of material. With all of the spirals, solid plugs were still created when they were rotated in such a direction as to result in an upward movement of material. Most of the upward movement of material occurred in the vicinity of the outer spiral blades, while minimal movement occurred inside the outer spiral blades, resulting in the creation of the plug.

6.3.3.3 Quantitative Measurements In order to judge the degree of structural breakdown for each spiral impeller, the percentage change in torque from the peak torque achieved during the first few seconds to the steady torque achieved after 90 seconds of shearing was measured. The average reductions in torque for each spiral are shown in Table 6.15. These results can be compared to the first series of tests using the six impellers (vane, egg-beater, half-egg beater, offset egg-beater, joint compound paddle, and Spiral 1), which were tested in similar concrete mixtures. It can be concluded that the reduction in torque was generally less for the spiral impellers than for the other impellers.

170 Table 6.15: Average Percentage Reduction in Torque for Each Spiral Impeller Mix 1 Mix 2 Mix 4 Average Initial Slump (In.) 5 3.5 3.75 Spiral 1 54.0% 52.7% 66.8% 57.8% Spiral 2 51.0% 50.6% 58.2% 53.2% Spiral 3 49.9% 50.0% 61.0% 53.6% Spiral 4 58.7% 55.0% 69.0% 60.9% Average 53.4% 52.1% 63.7%

Relative flow curves where measured with each impeller in each concrete mixture. The average relative parameters are shown in Table 6.16. It is important to consider the magnitudes of the yield values and viscosity values. Higher values of these parameters are desirable to ensure sufficient torque resolution. The lowest values or R2 were obtained in tests with Spiral 2 and Spiral 4. These spirals both differ from the other two spirals in that their bottom rings were removed. The bottom ring assists in promoting a certain flow path; however, the narrow gap between the ring and the spiral blades could prevent coarse aggregate from moving within the vane. In appears that, on balance, the removal of the bottom ring was disadvantageous. The average viscosity values were also lower for Spiral 2 and Spiral 4 than for the other spirals. The relative parameters for Mix 4 (3.5-inch slump) are indicted in Table 6.17 for the case where the direction of rotation of the spiral was reversed from the manufacturer’s recommendation. These data indicates that operating the spirals in reverse direction is of no advantage.

171 Table 6.16: Average Relative Parameters for Each Impeller (Forward Direction) Yield Viscosity Test R2 Value Value (Nm) (Nm.s) Spiral 1 5 sec, descending 2.209 1.485 0.686 90 sec, descending 2.053 0.613 0.885 90 sec, ascending 1.930 0.704 0.830 All 2.064 0.934 0.800 Spiral 2 5 sec, descending 2.370 0.806 0.903 90 sec, descending 2.255 0.277 0.486 90 sec, ascending 2.278 0.145 0.278 all 2.301 0.409 0.556 Spiral 3 5 sec, descending 2.016 1.302 0.938 90 sec, descending 1.931 0.564 0.843 90 sec, ascending 1.872 0.505 0.773 all 1.940 0.791 0.851 Spiral 4 5 sec, descending 3.896 0.673 0.485 90 sec, descending 3.590 0.755 0.480 90 sec, ascending 4.454 -0.654 0.328 all 3.980 0.258 0.431 Average 5 sec, descending 2.623 1.067 0.753 90 sec, descending 2.457 0.552 0.674 90 sec, ascending 2.633 0.175 0.552 all 2.571 0.598 0.660

172 Table 6.17: Average Relative Parameters for Each Impeller (Mix 4, Reverse Direction) Yield Viscosity Test R2 Value Value (Nm) (Nm.s) Spiral 1 -Reverse 5 sec, descending 2.183 0.779 0.932 90 sec, descending 2.325 0.558 0.547 90 sec, ascending 1.985 0.671 0.564 All 2.164 0.669 0.681 Spiral 2 - Reverse 5 sec, descending 1.723 1.062 0.891 90 sec, descending 1.435 0.673 0.942 90 sec, ascending 1.804 -0.052 0.055 All 1.654 0.561 0.629 Spiral 3 - Reverse 5 sec, descending 2.967 0.986 0.744 90 sec, descending 3.026 0.444 0.436 90 sec, ascending 2.761 0.592 0.800 All 2.918 0.674 0.660 Spiral 4 - Reverse 5 sec, descending 3.873 0.759 0.351 90 sec, descending 2.079 0.965 0.788 90 sec, ascending 2.436 0.141 0.165 All 2.796 0.622 0.435 Average - Reverse 5 sec, descending 2.687 0.897 0.730 90 sec, descending 2.216 0.660 0.678 90 sec, ascending 2.247 0.338 0.396 All 2.383 0.632 0.601

6.3.3.4 Conclusions All four spirals were unsuitable for measuring concrete rheology. First, the intended flow patterns for each spiral impeller were invalid for concrete due to the high viscosities of concrete and the presence of large, coarse aggregates. As a result, a solid plug formed within the blades of each spiral. Although the plugs formed in the spirals were not as distinct as in other impellers, the formation of

173 plugs was still problematic. Second, the rotation of the spirals—especially Spiral 4—pulled mortar out of the concrete and left behind the coarse aggregate. Third, the yield values and viscosity values measured by the spiral impellers were generally lower than those for other impellers, possibly resulting in inadequate torque resolution. In most cases for the spiral impellers, the difference in torque between 0.2 and 1.0 rev/sec was less than 1 Nm. Fourth, the variability of flow curve points was too high for all spiral impellers, as evidenced by the low values of R2. Finally, when the spiral impellers pushed material downward, the resistance provided by the concrete beneath each of the spirals was too large to allow concrete to move downward as intended. When the spiral impellers pushed material upward, the concrete was not sufficiently fluid to fill the void left at the bottom of each of the spirals; therefore, some mortar was sucked from the coarse aggregate network, resulting in a mortar-rich region at the top of the impeller.

6.3.4 Final Selection Based on the three series of tests of the candidate impellers, the vane exhibited the best performance and was selected for use in the ICAR rheometer. The qualitative observations indicated that the vane clearly engaged flow of surrounding material and produced favorable flow patterns. The wall effect was confined to the inside of the cut cylinder where its impact on results was minimal. When used to measure a flow curve, the vane consistently produced R2 terms close to unity. The degree of structural breakdown was comparable to other impellers. The amount of torque generated by the impeller was acceptable for providing adequate torque resolution while not exceeding the maximum torque capacity of the rheometer during the breakdown stage. Although the vane provided better results in the 16-inch container than in the 10-inch container, other container diameters could be acceptable. The selection of the vane does not

174 preclude the use of alternate impellers, which, if properly designed, could provide comparable or improved performance.

6.4 OPTIMUM SPEED FOR STRESS GROWTH TESTS Tests were conducted with the vane to determine the optimum speed for use in measuring yield stress in stress growth tests. If the speed is too slow, the structure of the concrete may reform before the yield stress is reached, resulting in an artificially high yield stress reading. If the speed is too fast, viscous forces and dynamic forces can amplify the static torque needed to initiate flow. By determining the yield stress at a range of different speeds, the minimum value of yield stress should indicate the optimum speed for stress growth tests. The impeller was rotated in four different concrete mixtures at a range of constant speeds ranging from 0.00833 rev/sec (0.5 rpm) to 0.1833 rev/sec (11 rpm). The minimum speed was limited by the capacity of the motor. The properties of the concrete mixtures that were tested are listed in Table 6.18. During the testing of each concrete mixture, the concrete remained in the container and was remixed with a shovel between each test. The 5-inch by 5-inch vane was tested in the 16-inch diameter container, with 5-inch gap sizes above and below the vane.

Table 6.18: Concrete Properties for Tests of Optimum Stress Growth Test Speed Slump before Slump after Final Elapsed Mix w/c Testing Testing Time (Inches) (Inches) (Hr:Min) Mix 1 0.416 4.5 3.75 0:36 Mix 2 0.387 2.5 2.5 0:28 Mix 3 0.484 8 7 0:40 Mix 4 0.469 3.5 4 0:36

175 Typical stress growth plots for a range of speeds are shown in Figure 6.14. The test was stopped once it was clear that the yield stress had been exceeded and the material around the vane was beginning to flow.

35 0.5 rpm 30 2.0 rpm 10.0 rpm 25

15 Torque, Nm

0 0 10203040506070 Time, Seconds Figure 6.14: Typical Stress Growth Plots at Various Rotation Speeds

The results of the yield stress speed testing are listed in Table 6.19 and plotted in Figure 6.15. A clear downward trend in peak torque was evident for the first three speeds of 0.5 rpm to 1.5 rpm in each mix. The rotation speeds of 10 and 11 rpm dis not allow for an elastic build-up of stress and were too fast for determining yield stress. Based on the plot in Figure 6.15, a rotation speed of 1.5 rpm (0.025 rev/sec) was selected as the optimum speed for stress growth tests.

176 Table 6.19: Results of Tests for Optimum Stress Growth Test Speed Peak Rotation Speed Mix Torque (rpm) (rev/sec) (Nm) 0.5 0.00833 30.3 1.0 0.01667 27.1 1.5 0.02500 24.1 Mix 1 2.0 0.03333 25.7 10 0.16667 22.1 11 0.18333 25.6 0.5 0.00833 53.4 1.0 0.01667 54.4 Mix 2 1.0 0.01667 54.9 2.0 0.03333 47.7 0.5 0.00833 13.5 1.0 0.01667 13.1 1.5 0.02500 7.5 Mix 3 2.0 0.03333 8.5 10 0.16667 9.2 11 0.18333 9.0 0.5 0.00833 43.6 1.0 0.01667 32.6 1.5 0.02500 36.1 Mix 4 2.0 0.03333 33.3 10 0.16667 24.6 11 0.18333 24.6

177 60.0 Mix 1 Mix 2 50.0 Mix 3 Mix 4 40.0

30.0

20.0 Peak Torque,Nm

10.0

0.0 0.001 0.010 0.100 1.000 Yield Stress Speed, rev/sec

Figure 6.15: Results of Tests for Optimum Stress Growth Test Speed

6.5 GAP SIZES The effect of the gap sizes between the vane and the container and between the vane and the free surface of the concrete was examined to determine the proper container size. Testing was performed with the 5-inch by 5-inch vane at the stress growth test speed of 0.025 rev/sec and at a speed of 0.667 rev/sec, which is representative of flow curve measurements. The vane was first rotated at the stress growth test speed up to the yield stress. The vane was then rotated at a constant speed of 0.667 rev/sec until structural breakdown occurred and an average steady-state torque had been reached. The properties of the four standard concrete mixtures tested are shown in Table 6.20.

178 Table 6.20: Concrete Properties for Gap Size Tests Slump before Slump after Final Elapsed Mix w/c Testing Testing Time (Inches) (Inches) (Hr:Min) Mix 1 0.472 5.25 4 0:59 Mix 2 0.470 5 4.5 0:53 Mix 3 0.465 7 6.5 0:49 Mix 4 0.402 4 3.75 0:51

To test the horizontal gap size, the vane was placed in the center of the 10- inch diameter container and the 16-inch diameter container, resulting in side gap sizes of 2.5 inches and 5.5 inches, respectively. These gap sizes represented likely maximum and minimum values. The 2.5-inch gap size would provide a gap size to maximum aggregate size ratio of 2.5 for a concrete with a 1-inch maximum aggregate size. This ratio of 2.5 has been found to be suitable by other authors (cf. Section 5.3.1). The 5.5-inch gap size produced a gap size to maximum aggregate size ratio of 5.5 for a concrete with a 1-inch maximum aggregate size. This higher ratio of 5.5 is slightly more than the generally accepted value of 3 to 5 for concrete. The test data for each of the four mixes are presented in Table 6.21 As summarized in Table 6.22, the decrease in gap size from 5.5 inches to 2.5 inches resulted in an increase in peak torque at 0.025 rev/sec of between 48 and 89% and an increase in steady torque at 0.667 rev/sec of between 30 and 64%. Aggregate interlocking effects were presumably the main contributing factor for the increase in torque in the smaller container. Accordingly, the increase in torque at the yield stress speed was less for the rounded river gravel than for the angular crushed limestone. If aggregate interlocking is present to a sufficient degree and the mortar is sufficiently cohesive, the vane may try to turn the entire volume of concrete within the container as a single plug. Further, if material tends to move

179 radially outward, the horizontal resistance provided by the smaller container can increase the torque.

Table 6.21: Testing of Horizontal Side Gap Peak Steady Side Torque Torque Mix Gap (0.025rev/s) (0.67rev/s) (Inches) (Nm) (Nm) 5.5 24.5 6.3 Mix 1 2.5 36.9 10.3 5.5 37.9 9.7 Mix 2 2.5 56.0 12.5 5.5 16.4 4.4 Mix 3 2.5 29.9 6.3 5.5 27.5 7.1 Mix 4 2.5 51.8 --

Table 6.22: Effect of Horizontal Side Gap (10-Inch Container Torque)/ (16-Inch Container Torque) Mix 1 0.025 rev/sec 1.51 0.667 rev/sec 1.64 Mix 2 0.025 rev/sec 1.48 0.667 rev/sec 1.30 Mix 3 0.025 rev/sec 1.83 0.667 rev/sec 1.42 Mix 4 0.025 rev/sec 1.89 0.667 rev/sec --

To test for the effect of bottom gap and top cover, the 5-inch by 5-inch vane was tested in the 16-inch diameter container with varying distances between the vane and the top of the concrete and bottom of the container. The vane was

180 rotated at a constant speed of either 0.025 or 0.667 rev/sec. The effects of the bottom gap and top cover distances are shown in Table 6.23.

Table 6.23: Effects of Top Cover and Bottom Gap Peak Steady Bottom Top Torque Torque Mix Gap Cover (0.025rev/s) (0.67rev/s) (Inches) (Inches) (Nm) (Nm) 5 0 18.5 5.5 5 2.5 14.5 5.9 5 5 24.5 6.3 Mix 1 5 10 25.5 6.6 2.5 5 41.7 6.8 10 5 28.5 7.4 5 0 20.2 7.4 5 2.5 26.0 8.7 5 5 37.9 9.7 Mix 2 5 10 33.3 10.3 2.5 5 53.2 10.2 10 5 38.3 9.8 5 0 13.7 4.7 5 2.5 11.5 4.7 5 5 16.4 4.4 Mix 3 5 10 15.2 4.8 2.5 5 17.7 5.9 10 5 19.5 4.8 5 0 18.1 5.7 5 2.5 16.9 7.5 5 5 27.5 7.1 Mix 4 5 10 27.8 7.5 2.5 5 35.6 8.9 10 5 30.2 8.0

At the stress growth test speed, a clear trend in torque was evident, as shown in Figure 6.16. Whereas the peak torque dropped slightly for three of the four mixes when the top cover was increased from no cover to 2.5 inches, the peak torque for all four mixes increased sharply when the cover was further 181 increased to 5 inches. At the low value of top cover, the stress acting on top of the cylindrical volume cut by the vane was negligible. Once a sufficient amount of material covered the vane, the torque due to top cover reached an approximately constant level. The effect of the bottom gap at the stress growth test speed, shown in Figure 6.17, was pronounced at a distance of 2.5 inches. Above a distance of 5 inches, or 5 times the maximum aggregate size of 1 inch, the torque was constant, suggesting that a bottom gap distance of 5 inches was suitable.

25 Mix 1 Mix 2 20 Mix 3 15 Mix 4 Peak Torque, Nm Peak Torque, 10

0 02.557.510 Top Cover, Inches

Figure 6.16: Effect of Top Cover at 0.025 rev/sec (Bottom Gap = 5 Inches)

182 60

40 Mix 1 Mix 2 30 Mix 3 Mix 4 20 Peak Torque, Nm Peak Torque,

0 02.557.510 Bottom Gap, Inches

Figure 6.17: Effect of Bottom Gap at 0.025 rev/sec (Top Cover = 5 Inches)

At the higher speed of 0.667 rev/sec, the effects of bottom gap and top cover were slightly different than at stress growth test speed. Except for Mix 1, the increase in top cover resulted in an increase in torque, as indicated in Figure 6.18. For Mix 3, the effect of top cover appeared to be negligible. For the bottom gap distance, Figure 6.19 indicates that increasing the gap size from 2.5 inches to 5 inches resulted in a decrease in torque for all four mixes. When the gap size was increased to 10 inches, the torque increased only modestly. Based on these data, it appears that a bottom gap distance of 5 inches was suitable for the flow curve speed.

183 12

8 Mix 1 Mix 2 6 Mix 3 Mix 4 4 Steady Torque, Nm Steady Torque,

0 02.557.510 Top Cover, Inches

Figure 6.18: Effect of Top Cover at 0.667 rev/sec (Bottom Gap = 5 Inches)

8 Mix 1 Mix 2 6 Mix 3 Mix 4 4 Steady Torque, Nm Steady Torque,

0 02.557.510 Bottom Gap, Inches

Figure 6.19: Effect of Bottom Gap at 0.667 rev/sec (Top Cover = 5 Inches) 184 Based on the data presented above, it is evident that for bottom gap, top cover, and side gap, there are minimum distances above which the gap or cover distance does not have a significant influence on the torque measurement at either 0.025 rev/sec or 0.667 rev/sec. The data indicate that bottom gap and top cover distances of 5 inches are suitable; however, additional measurements may indicate that smaller distances are acceptable. Further, the 16-inch container appears to be suitable, although a smaller container diameter could be acceptable.

6.6 END EFFECTS For flow curve measurements, it is necessary to know the amount of shear stress acting on the ends of the vane in order to determine analytically yield stress in plastic viscosity. Likewise, with direct yield stress measurements, the shear stress at the ends of the vane must be known. The magnitude of the shear stress acting on the top and bottom of the cylindrical volume cut by the vane was assessed at the stress growth test speed (0.025 rev/sec) and a speed representative of flow curve measurements (0.667 rev/sec). The properties of the concrete tested are shown in Table 6.24.

Table 6.24: Concrete Properties for End Effects Testing Slump before Slump after Final Elapsed Mix w/c Testing Testing Time (Inches) (Inches) (Hr:Min) Mix 1 0.496 6.5 5.5 0:38 Mix 2 0.444 3 2.5 0:47 Mix 3 0.447 6 4.75 0:41 Mix 4 0.401 3.25 3 0:51

Vanes of three heights—5, 7.5, and 10 inches—and a constant diameter of 5 inches were tested at speeds of 0.025 rev/sec and 0.667 rev/sec. The peak torque was determined for the stress growth test speed and continuous steady-state 185 torque was determined for the representative flow curve speed. The torque measurements were plotted versus height to determine the contribution of end shear stresses. The minimum vane height was five inches in order to maintain a minimum H/D ratio greater than one. For comparison purposes, the tests were conducted in both the 10-inch diameter and 16-inch diameter containers. The test data for end effect testing are shown in Table 6.25.

Table 6.25: Test Data for End Effects Testing Peak Steady Container Vane Torque Torque Mix Diameter Height (0.025rev/s) (0.67rev/s) (Inches) (Inches) (Nm) (Nm) 5 14.8 3.6 16 7.5 14.3 5.3 10 17.3 7.1 Mix 1 5 15.8 4.8 10 7.5 22.5 6.7 10 23.0 10.0 5 28.9 7.4 16 7.5 36.0 10.9 10 47.6 14.2 Mix 2 5 45.4 12.1 10 7.5 44.3 16.9 10 51.1 20.4 5 17.4 4.3 16 7.5 15.3 5.3 10 26.0 8.7 Mix 3 5 17.0 5.3 10 7.5 26.5 8.0 10 18.9 10.7 5 28.6 6.0 16 7.5 30.4 9.7 10 35.6 12.0 Mix 4 5 35.5 8.6 10 7.5 40.1 12.7 10 49.6 17.9

186 The end effects for stress growth tests and for flow curve measurements must be considered separately. The test data for stress growth tests are plotted in Figure 6.20 and Figure 6.21. Based on these data, the end effects for each mix can be expressed in two ways. First, a straight line fit of the torque versus height data can be extrapolated to a vane height of zero to determine a constant quantity of torque attrituable to the ends of the vane. Second, the straight line fit can be extrapolated back further to the height axis to determine the marginal effective height of the vane. Both of these values are presented in Table 6.26.

40 Mix 1 Mix 2 30 Mix 3 Mix 4

Peak Torque, Nm Peak Torque, 20

0 024681012 Vane Height, Inches

Figure 6.20: Determination of End Effects: 0.025 rev/sec, 16-Inch Container

187 60

40 Mix 1 Mix 2 30 Mix 3 Mix 4

Peak Torque, Nm Peak Torque, 20

0 024681012 Vane Height, Inches

Figure 6.21: Determination of End Effects: 0.025 rev/sec, 10-Inch Container

Table 6.26: Determination of End Effects for Stress Growth Test Marginal Container Zero Height Effective Mix Diameter Torque Height (Inches) (Nm) (Inches) 16 11.75 23.88 Mix 1 10 9.56 6.58 16 9.46 2.53 Mix 2 10 38.46 33.97 16 6.63 3.85 Mix 3 10 18.01 48.78 16 21.00 14.97 Mix 4 10 20.51 7.25

In order to use the end effects data in calculating yield stress, the data was used in the three methods described previous in Chapter 2 for determining the yield stress. In Method 1, the shear stress at the peak torque is assumed to be

188 evenly distributed across the cylindrical volume cut by the vane and equal to the yield stress. While the determination of yield stress by this first method only requires the use of one vane, the yield stress should be equal for any vane geometry, assuming the geometry is not such that it distorts test results. The yield stresses determined for each vane height based on the first method are shown in Table 6.27 for both the 10-inch and 16-inch containers. The scatter in yield stress was not unexpected given the scatter of data in Figure 6.20 and Figure 6.21.

Table 6.27: Yield Stress Values Determined from Method 1 Vane Yield Stress (Pa) Mix Height 16-Inch 10-Inch (Inches) Container Container 5 3447 3677 7.5 2424 3822 Mix 1 10 2298 3068 Average 2723 3522 CoV 23.2% 11.4% 5 6746 10590 7.5 6095 7514 Mix 2 10 6346 6805 Average 6396 8303 CoV 5.1% 24.2% 5 4050 3964 7.5 2591 4489 Mix 3 10 3462 2511 Average 3367 3655 CoV 21.8% 28.0% 5 6660 8266 7.5 5157 6805 Mix 4 10 4740 6609 Average 5519 7227 CoV 18.3% 12.5%

In the second method, the end shear stress is assumed to be distributed based on a power-law relationship, with the shear stress equal to zero at the center

189 of the vane and equal to the yield stress at the outer tip of the vane. If the exponent, m, in the equation is equal to unity, then a linear relationship exists between the cylinder radius and shear stress. The results, shown in Table 6.28, indicated that for all 8 determinations, the value of the exponent, m, was negative. Such a negative value is not reasonable, as it suggests a zero shear stress at the outer end of the vane and an infinite shear stress at the center of the vane.

Table 6.28: Yield Stress Values Determined from Method II 16-Inch Container 10-Inch Container Mix Yield Stress Yield Stress m m (Pa) (Pa) Mix 1 765 -2.79 2257 -2.24 Mix 2 5813 -1.02 1759 -2.85 Mix 3 2677 -1.70 574 -2.90 Mix 4 2180 -2.67 4398 -2.31

In Method 3, no assumption is made about the distribution of shear stress on the ends of the cylinder. For the data generated, this method results in calculated yield stress values that are equal to the yield stress values computed in the second method. The results of Method 3 are unreasonable given the relative amounts of torque attributable to the side and ends of the vane. The percentages of torque attributable to the side of the effective cylinder cut by the vane are shown in Table 6.29. Not only is the variation between each mix wide, some of the values are impractical. For instance, it is not reasonable to assume that 17% of the torque is attributable to the side of the effective cylinder when the side comprises half the area and is located at the greatest moment arm distance from the center of the vane.

190 Table 6.29: Torque Attributable to Side of Effective Cylinder Based on Method III for 5-Inch Vane Torque Attributable to Side Mix 16-Inch 10-Inch Container Container Mix 1 17% 46% Mix 2 65% 12% Mix 3 50% 11% Mix 4 25% 40%

The use of the second and third methods is heavily dependent on the fit of the straight line in the torque versus vane height plots. Given the scatter evident in these plots, the fact that the second and third methods did not provide reasonable results was not surprising. Had more data points been available with less scatter, the second or third methods may have been appropriate. The first method is based on reasonable assumptions—namely that the shear stress is evenly distributed and equal to the yield stress. In view of the results of Method 2 and Method 3, the use of the first method was determined to be appropriate for determining yield stress. For flow curve measurements, the end effect was determined to be negligible. When the continuous torque was plotted versus vane height for the 16- inch and 10-inch containers—shown in Figure 6.22 and Figure 6.23, respectively—the intercept of the fitted straight line was near the origin for all but one mixture. As indicated in Table 6.30, the average intercept for the 16-inch container was 0.14 Nm, resulting in an average marginal effective height of 0.09 inches. These test results do not necessarily imply that the only material engaged by the vane is the material horizontally adjacent to the vane, namely, in the annulus. Indeed, visual observations of the concrete in the rheometer indicated deformation of the concrete at the surface. The limited data do, however, suggest that shear stress is distributed throughout the concrete—even for different concrete mixtures—in such a way that the effective height of the vane can be assumed to be equal to the actual height of the vane.

191

Table 6.30: Determination of End Effects for Flow Curve Measurements Marginal Container Zero Height Effective Mix Diameter Torque Height (Inches) (Nm) (Inches) 16 0.16 0.23 Mix 1 10 -0.52 -0.51 16 0.67 0.50 Mix 2 10 3.98 2.40 16 -0.51 -0.57 Mix 3 10 -0.07 -0.07 16 0.26 0.21 Mix 4 10 -0.82 -0.44 16 0.14 0.09 Average 10 0.64 0.35 All 0.39 0.22

10 Mix 1 Mix 2 Mix 3 5 Mix 4

Steady-State Torque,Steady-State Nm 0

-5 024681012 Vane Height, Inches Figure 6.22: Determination of End Effects: 0.667 rev/sec, 16-Inch Container

192 25

15 Mix 1 Mix 2 10 Mix 3 Mix 4 5 Steady-State Torque, Nm 0

-5 024681012 Vane Height, Inches

Figure 6.23: Determination of End Effects: 0.667 rev/sec, 10-Inch Container

Based on the data for end effects testing, it can be concluded that the first method for determining yield stress—namely, where shear stress is assumed to be evenly distributed on the side and ends of the effective cylinder cut by the vane and equal to the yield stress—is the most appropriate method. For flow curve measurements, the effective marginal height can be assumed to be equal to zero, although it cannot be assumed that no stresses act on the ends of the vane.

6.7 CONCLUSIONS The testing described in this chapter enabled the design of the first generation prototype of the ICAR rheometer to be further finalized. The results are summarized in Table 6.31. Additional testing is needed to define in more detail all of the operating characteristics described in this chapter. The testing indicated that the 5-inch by 5-inch vane impeller was preferable to the other eight

193 candidate impellers. When used to perform a stress growth test to determine yield stress, it was determined that the vane should be rotated at a speed of 0.025 rev/sec (1.5 rpm). For both stress growth tests and flow curve measurements, it was determined that minimum bottom gap, side gap, and top cover distances existed above which larger distances had minimal impact on the measured torque. Based on the test data obtained, bottom gap and top cover distances of 5 inches and side gap distances of 5.5 inches were found to be acceptable. For stress growth tests, it was determined that the yield stress should be calculated assuming that the shear stress acting on the cylinder cut by the vane is evenly distributed and equal to the yield stress. For flow curve measurements, it was found that the effective height due to end effects can be assumed to be equal the actual height of the vane.

Table 6.31: Summary of Results from Chapter 6 Parameter Conclusion Impeller Vane (5-inch by 5-inch) Optimum Stress Growth Test Speed 0.025 rev/sec Gap Sizes Vane to Side of Container Minimum of 5.5 Inches Vane to Bottom of Container Minimum of 5 Inches Vane to Top of Concrete Minimum of 5 Inches End Effects Shear stress evenly distributed Speed = 0.025 rev/sec on ends and equal to yield stress Effective marginal height of Speed = 0.667 rev/sec vane negligible

194

CHAPTER 7: LABORATORY TESTING PROGRAM

7.1 INTRODUCTION A series of laboratory tests was conducted to verify the ability of the ICAR rheometer to detect changes in workability and rheology. Over 100 concrete mixtures, with workability ranging from a slump of 2 inches (50 mm) to self-consolidating concrete, were tested with the rheometer. The mixtures included a broad range of materials commonly incorporated into concrete, including fly ash, silica fume, ground granulated blast furnace slag, air-entraining agent, viscosity modifying admixture, and three water-reducing admixtures. Additionally, different aggregates and aggregate gradations were considered. By comparing the changes in rheology measured by the ICAR rheometer with changes reported in the literature and with qualitative observations of workability, it was possible to evaluate the effectiveness of the ICAR rheometer. This chapter describes the laboratory test program in detail. First, the materials, mixture proportions, and test procedures are presented. Then, the results for the conventional concrete mixtures (with slumps from 2 to 9 inches) are presented followed by the results for the self-consolidating concrete mixtures. Finally, field testing conducted to demonstrate the portability of the rheometer is described.

7.2 MATERIALS The materials used in the testing were selected to be broadly representative of those typically used in construction. The cement used in all

195 mixtures was an ASTM C 150 Type I portland cement from Capitol Aggregates in San Antonio, TX. Its chemical composition and physical properties, as provided by the manufacturer, are shown in Table 7.1.

Table 7.1: Chemical and Physical Properties of Cement Chemical Composition Percent Sulfur Trioxide (SO3) 3.30 Silica Dioxide (SiO2) 20.70 Ferric Oxide (Fe2O3) 1.21 Magnesium Oxide (MgO) 1.21 Aluminum Oxide (Al2O3) 5.11 Equivalent Alkalies (as Na2O) 0.56 Calcium Oxide (CaO) 64.43 Tricalcium Silicate (C3S) 58.60 Tricalcium Aluminate (C3A) 10.35 Property Value Loss on Ignition (%) 1.59 Insoluble Residue (%) 0.47 Blaine Fineness (m2/kg) 358 Vicat Initial (minutes) 106 Vicat Final (minutes) 147

Seven different aggregates were used throughout the testing. A river gravel and natural sand were obtained from Texas Industries in Austin, TX. A crushed limestone coarse aggregate, intermediate aggregate, and manufactured sand were obtained from Capitol Aggregates in Georgetown, TX. An additional granite manufactured sand, which was the same material tested by Quiroga (2003), was used. Finally, a blast furnace slag coarse aggregate was obtained from the Edward C. Levy Corporation in Detriot, MI. The basic properties of these aggregates are shown in Table 7.2. The specific gravities (BSGSSD) and absorption capacities (AC) were determined in accordance with ASTM C 127 for coarse aggregates and ASTM C 128 for fine aggregates. The particle size distributions and fineness moduli (FM) of the as-received material, as determined

196 in accordance with ASTM C 136 and ASTM C 117, are shown in Table 7.3 and plotted in Figure 7.1. The particle size distribution of the granite sand is not shown because its original gradation was not used.

Table 7.2: Aggregate Properties

Aggregate Designation Source Location BSGSSD AC (%) Coarse Crushed Capitol Georgetown, LS 2.47 5.42 Limestone Aggregates TX River Gravel RG TXI Austin, TX 2.61 1.02 Slag SlagAgg E.C. Levy Detroit, MI 2.34 2.49 Intermediate Crushed Capitol Georgetown, LSI 2.47 5.45 Limestone Aggregates TX Sand Natural River NS TXI Austin, TX 2.56 0.55 Sand Limestone Capitol Georgetown, Manufactured MS 2.45 4.55 Aggregates TX Sand Granite Manufactured GR Unknown Oklahoma 2.73 0.25 Sand

197 Table 7.3: Aggregate Particle Size Distributions US Mesh Percent Passing Standard Size Sieve (mm) LS RG SlagAgg LSI NS MS 1 1/4 in. 31.5 100.0 99.6 100.0 1 in. 25.0 100.0 83.4 98.7 ¾ in. 19.0 82.4 44.1 81.1 ½ in. 12.5 10.7 17.0 64.8 99.9 3/8 in 9.5 0.6 5.1 45.0 81.3 #4 4.75 0.3 0.5 24.1 11.9 97.3 100.0 #8 2.36 13.6 1.8 86.2 92.3 #16 1.18 11.6 1.2 70.5 67.8 #30 0.600 49.0 46.2 #50 0.300 27.3 29.2 #100 0.150 6.8 16.1 #200 0.075 0.6 11.8 FM ------2.63 2.55

100%

90% LS RG 80% SlagAgg 70% LSI NS 60% MS 50%

40%

Percent Passing Percent 30%

20%

10%

0% 0.01 0.1 1 10 100 Size, cm

Figure 7.1: Aggregate Particle Size Distributions

198 Except for the mixtures where the gradation was the variable being changed, the aggregates were used as-received—that is, they were not sieved and recombined but taken directly from stockpiles. The process of sieving and reblending fine aggregates can result in significant errors in the achieved gradation because of inaccuracies in the sieving operation. Some material that should pass through to smaller sieves is retained erroneously on larger sieves. The majority of material that that does not pass to the correct sieve is finer than the #200 sieve. These microfines, which mainly remain as dust on larger particles, can be removed from the larger size fractions by shaking the sieves for a longer period of time or by running less material through the sieves at one time. Regardless of how long the sieves are operated or how little material is added to the sieves, some microfines will remain on larger sieves. This error must be addressed when reblending the size fractions for fine aggregate. At least two approaches for producing accurate sand gradations are possible. First, the remaining microfines on larger sieves could be removed by washing the material over a #200 sieve until the water flowing through the sieve is clear. Since microfines can vary widely in size, shape, and texture, it is important that the material washed out of larger size fractions be recovered in order to avoid a systematic elimination of microfines with certain characteristics. The microfines left on larger size fractions may be different from those that pass through to the pan. Indeed, Quiroga (2003) found that microfines removed by dry sieving had a lower methylene blue value than those removed by washing. Further, the process of recovering the washed material is complicated by the fact that the microfines suspended in the wash water settle at different rates, depending on their size. Therefore, the particles remaining in the wash water will be segregated once the water has dried. A second approach would be to determine the amount of microfines present in each size fraction after a standard sieving procedure and account for these microfines when reblending the size fractions.

199 Neither option can exactly replicate the distribution of material achieved when aggregate is taken directly from a quarry operation; however, both approaches are preferable to making no corrections. In this project, the second approach was used. During the sieving process, an approximately constant quantity of sand was added to the sieves, which were then operated for a fixed period of 10 minutes. Then, a random sample of material was taken from each aggregate size fraction and the amount of microfines in the size fraction was determined by washing based on the procedure described in ASTM C 117. The quantities of microfines present in each size fraction for the limestone and granite manufactured sands after the standard sieving operation are shown in Figure 7.2. These quantities were then used when the aggregate size fractions were reblended to ensure an accurate reporting of the total amount of microfines in the sand. If this precaution had not been made, the amount of microfines actually in the material would have been significantly understated. All chemical admixtures used in the lab testing were supplied by Grace Construction Products, Cambridge, MA. The air-entraining agent was Daravair® 1000, which meets the requirements of ASTM C 260. Three different water- reducing admixtures were used. The high-range water reducer was ADVA® Flow, a polycarboxylate-based admixture meeting the requirements of ASTM C 494 for a Type F admixture. The low-range water reducer was WRDA® 82, a lignosulfonate-based admixture meeting the requirements of ASTM C 494 for a Type A admixture. The self-consolidating concrete mixtures incorporated ADVA® Cast 530, a polycarboxylate-based admixture intended specifically for the production of self-consolidating concrete in precast concrete applications. ADVA® Cast 530 meets the requirements of ASTM C 494 for a type Type F admixture and the requirements of ASTM C 1017. The retarder was Daratard® 17, which meets the requirements of ASTM C 494 for Type B and Type D admixtures. The viscosity modifying admixture was V-MAR® 3.

200

50%

45% 43.3% Limestone 40% Granite 35%

30% 25.0% 25%

20% 17.5% Percent < #200 15% 11.4% 10%

4.4% 5% 2.2% 2.4% 1.5% 1.3% 0.3% 0.7% 0.8% 0% #8 #16 #30 #50 #100 #200 US Standard Sieve Size Figure 7.2: Amount of Microfines Remaining in Each Size Fraction after Standard Sieving Operation for Limestone and Granite Manufactured Sands

Three supplementary cementitious materials were used. The class F fly ash, meeting the requirements of ASTM C 618, was obtained from a Boral Material Technologies facility in Rockdale, TX. The densified silica fume, Force 10,000® D, was obtained from Grace Construction Products. The ground granulated blast furnace slag was obtained from a Holcim (US) Incorporated facility in Chicago, IL.

7.3 MIXTURE PROPORTIONS Two base concrete mixtures, which were designed in accordance with ACI 211.1, were systematically altered to produce changes in workability. One change

201 to the mixture proportions was made at a time. The first base mixture, designated RG_NS, consisted of river gravel and natural sand while the second base mixture, designated LS_NS, consisted of crushed limestone coarse aggregate and natural sand. The proportions of base mixtures, which are shown in Table 7.4, vary by more than just the aggregate type; therefore, a direct comparison of river gravel to crushed limestone coarse aggregate should not be made for any of the data presented in this chapter. The water contents indicated for each mixture correspond to slumps of 100 mm (4 inches). Based on anticipated changes in workability, the water content was adjusted for each mixture series so that the range of slumps would be between 75 and 175 mm (3 to 7 inches) for all mixtures. For instance, when high-range water reducer was added to the mixtures, the water contents of the control mixtures were reduced. When silica fume was added, the water contents were increased for the control mixtures. Therefore, control mixtures for different mixture series should not be directly compared.

Table 7.4: Proportions of Base Mixtures Mass, kg/m3 (lb/yd3) Mix Coarse Fine Aggregate Designation Cement Water Aggregate (SSD) (SSD) RG_NS 1114.4 (1878.3) 687.7 (1159.2) 365.0 (615.3) 173.4 (292.3) LS_NS 1023.4 (1724.9) 696.2 (1173.5) 388.2 (654.4) 174.7 (294.5)

The self-consolidating concrete mixtures utilized different mixture proportions, which are described later in Section 7.8.

7.4 MIXING AND TESTING PROCEDURES The concrete mixtures were prepared in general accordance with ASTM C 192. Batching, mixing, and testing operations were conducted in an air- conditioned room, where the temperature was maintained between 21.1º and 23.3º

202 C (70º and 74º F). All materials—including aggregates, cement, supplementary cementitious materials, and admixtures—were placed in the mixing room at least 12 hours prior to the start of mixing to allow the materials to reach a constant temperature. The aggregates, cementitious materials, and water were batched by weight while the chemical admixtures were batched by volume. With the exception of the high-microfines mixtures, which were batched in an oven-dry condition, all other aggregates were batched in a moist condition, close to their saturated surface dry condition. The moisture content of each aggregate was determined and batch quantities were adjusted accordingly. The batched materials were stored in sealed containers between the time of batching and mixing. The concrete was mixed in a rotating drum mixer. For consistency, the standard batch size was 0.0623 m3 (2.2 ft3) for all mixtures except for the self- consolidating concrete mixtures. To begin mixing, all aggregates were placed in the mixer along with a portion of the mixing water. The mixer was started to blend the materials. Next, with the mixer stopped, the cement and any supplementary cementitious materials were added. Finally, the mixer was restarted and the remaining mixing water and any chemical admixtures were added. The concrete was then mixed for 3 minutes, allowed to rest for 3 minutes, mixed for an additional 2 minutes, and then immediately discharged into a wheelbarrow. Fresh concrete tests were performed on a consistent, predetermined schedule starting from the time of contact between water and cement. For mixes containing an air entrainment agent, the air content was determined by the pressure method in accordance with ASTM C 237. The air content test was started at 9 minutes after the contact of water and cement. Concrete was transferred by shovel from the wheelbarrow into the rheometer container at 12 minutes. The concrete was dropped into the container from a consistent height; no additional effort was made to achieve further consolidation of the concrete. The slump test

203 was performed in accordance with ASTM C 143 at 16 minutes. Finally, the rheometer was started at 18 minutes. After this series of testing, the concrete was discarded. The dimensions of the container and vane of first generation prototype of the ICAR rheometer, which was used in all testing, are shown in Figure 7.3. The ICAR rheometer was used to perform a stress growth test and to measure a flow curve. The stress growth test was started as soon as the vane was immersed into the concrete. The vane was rotated at a constant rotation speed of 0.025 rev/sec while the build-up in torque was monitored on the computer screen. Once the peak torque had been reached, the flow curve measurement was started. Five equally spaced points were measured on the flow curve in descending order from 1.0 rev/sec to 0.2 rev/sec. Upon completion of the flow curve, the vane was removed and the concrete was remixed with a shovel. The vane was then reinserted into the concrete and the same test procedure was repeated a second time. The stress growth test results were reported in terms of maximum torque in Nm. The flow curve results were expressed in two forms: in terms of yield value (Nm) and viscosity value (Nm.s) and in terms of yield stress (Pa) and plastic viscosity (Pa.s). Yield value and viscosity value were determined from a straight line fit of the torque versus rotation speed data. Yield stress and plastic viscosity were determined with the Effective Annulus Method.

204

Figure 7.3: ICAR Rheometer Dimensions

In addition to the quantitative fresh concrete measurements, a qualitative assessment was made of each mixture by rating six aspects of workability on a scale of 1 to 5. These six aspects of workability along with a key to their ratings are shown in Table 7.5. For brevity, only ratings of 1, 3, and 5 are described in the table. All observations and ratings were made by the author.

205 Table 7.5: Visual Observations of Fresh Concrete Factor Ratings 1 = The concrete clumps and is extremely dry. 3 = The concrete does not clump but is cohesive and difficult to move Flowability or shovel. 5 = The concrete is highly fluid and flows with little assistance. 1 = The concrete is extremely harsh due to a deficiency of mortar. When the slump cone is removed, large voids between coarse aggregates are clearly visible. Richness 3 = The concrete exhibits a moderate mortar content. 5 = The concrete exhibits a smooth consistency due to an abundance of mortar. 1 = The concrete exhibits severe segregation in the mixer and upon being discharged into the wheelbarrow. Segregation 3 = The concrete appears stable when at rest but segregates when Resistance subjected to shearing. 5 = The concrete is uniform throughout even after being subjected to shearing. 1 = Bleeding is evident when concrete is in the wheelbarrow and becomes severe when concrete is agitated. 3 = The concrete exhibits moderate bleeding after being agitated. Bleeding Water rises to the surface when the concrete is sheared in the Resistance rheometer. 5 = The concrete does not bleed, even after being undergoing shearing in the rheometer. 1 = The concrete requires excessive troweling—including the application of high vertical pressure—to achieve a smooth finished surface. Finishability 3 = The concrete requires a moderate number of passes with a trowel to achieve a smooth finished surface. 5 = The concrete forms a smooth finished surface after a minimal number of passes with a trowel. 1 = The concrete exhibits extremely poor workability and could not be properly placed in the field regardless of the amount of effort. Overall 3 = The concrete could be properly placed in the field but only with an Workability above average amount of effort. 5 = The concrete could be properly placed in the field with a minimal amount of effort.

206 7.5 TEST RESULTS FOR CONVENTIONAL CONCRETE The results for each change in mixture proportions are presented in the following subsections. Six graphs in each section report the results in terms of yield value, viscosity value, yield stress, plastic viscosity, yield torque, and slump. Full test data for all mixtures are located in Appendix A.

7.5.1 Fly Ash Fly ash was used as a mass replacement for cement at rates of 10, 20, 35, and 55%. The 55% replacement level was selected to represent high-volume fly ash concrete. The test results are shown in Figure 7.4. As expected, the addition of increasing levels of fly ash resulted in a reduction in yield value, yield stress and yield torque in both the river gravel and crushed limestone mixtures. Whereas the viscosity value was reduced with increasing replacement levels of fly ash, the effect of fly ash on the plastic viscosity was variable. Notably, the general trend for plastic viscosity was similar for both the river gravel and limestone mixtures. The addition of fly ash resulted in an increase in slump.

207 9 2.5 RG_NS RG_NS 8 LS_NS LS_NS 2 7

6 m

N 1.5 5 ue, l a V 4 ld

e 1

Yi 3 Viscosity Value, Nm.s 2 0.5 1

0 0 0 102030405060 0 102030405060 Fly Ash Replacement Level (%) Fly Ash Replacement Level (%) Yield Value Viscosity Value 2500 35 RG_NS LS_NS 30 2000 25

1500 20

15 1000 Yield Stress, Pa

Plastic Viscosity, Pa.s 10 500 5 RG_NS LS_NS 0 0 0 1020304050600 102030405060 Fly Ash Replacement Level (%) Fly Ash Replacement Level (%) Yield Stress Plastic Viscosity 45 9 RG_NS 40 8 LS_NS

35 7 , Nm )

30 6 rowth G 25 5 tress tress

(S 20 4 Slump, Inches Slump, 15 3

10 2

Yield Torque RG_NS 5 1 LS_NS 0 0 0 1020304050600 102030405060 Fly Ash Replacement Level (%) Fly Ash Replacement Level (%) Yield Torque Slump Figure 7.4: Influence of Fly Ash on Rheology

208 7.5.2 Ground Granulated Blast Furnace Slag Ground granulated blast furnace slag (GGBFS) was used to replace cement on a mass basis at rates of 20, 35, and 50%. The test results are plotted in Figure 7.5. At the 20% replacement rate, the yield value, yield stress, and yield torque all increased for both aggregates. Additional replacement of cement with slag up to 50% resulted in decreases in these values. The trend in the viscosity value generally matched that of the yield value. Plastic viscosity increased slightly at the 20% replacement rate for the river gravel mixture, but otherwise decreased with increasing levels of GGBFS.

209 7 3 RG_NS RG_NS LS_NS 6 LS_NS 2.5

5 2

4 1.5 3

Yield Value, Nm Value, Yield 1

2 Nm.s Value, Viscosity

0.5 1

0 0 0 1020304050600 102030405060 GGBFS Replacement Level (%) GGBFS Replacement Level (%) Yield Value Viscosity Value 1400 35 RG_NS RG_NS LS_NS LS_NS 1200 30

1000 25

800 20 tress, Pa S 600 15 Yield Yield

400 Plastic Viscosity, Pa.s 10

200 5

0 0 0 1020304050600 102030405060 GGBFS Replacement Level (%) GGBFS Replacement Level (%) Yield Stress Plastic Viscosity 35 7

30 6 , Nm ) 25 5 rowth G 20 4 tress tress

(S 15 3 lump, Inches S

10 2

Yield Torque 5 RG_NS 1 RG_NS LS_NS LS_NS 0 0 0 1020304050600 102030405060 GGBFS Replacement Level (%) GGBFS Replacement Level (%) Yield Torque Slump Figure 7.5: Influence of GGBFS on Rheology

210 7.5.3 Silica Fume Silica fume was used as a mass replacement of cement at rates of 3, 5, and 8%. The test results are shown in Figure 7.6. For the river gravel concrete mixture, the addition of silica fume resulted in increases in yield value, yield stress, and yield torque at all addition rates. The yield value, yield stress and yield torque remained approximately constant for the crushed limestone concrete mixture. The plastic viscosity decreased for the low replacement rates of 3% and 5% but began increasing for the 8% replacement rate. Only in the river gravel concrete mixture did the plastic viscosity exceed the control mixture at the 8% replacement rate. While the slump decreased for the river gravel concrete mixture with increasing silica fume contents, it remained constant for the crushed limestone concrete mixture at 3, 5, and 8% replacement rates.

211 7 3 RG_NS 6 LS_NS 2.5

5 2

4 1.5 3

Yield Value, Nm 1

2 Viscosity Value,Nm.s RG_NS 0.5 LS_NS 1

0 0 02468100246810 Silica Fume Replacement Level (%) Silica Fume Replacement Level (%) Yield Value Viscosity Value 1400 60 RG_NS RG_NS LS_NS 1200 LS_NS 50

1000 40

800

tress, Pa tress, 30 S 600

Yield 20

400 Plastic Viscosity, Pa.s

10 200

0 0 02468100246810 Silica Fume Replacement Level (%) Silica Fume Replacement Level (%) Yield Stress Plastic Viscosity 45 5 RG_NS RG_NS 4.5 40 LS_NS LS_NS 4 35 , Nm ) 3.5 30 rowth

G 3 25 2.5 tress tress

(S 20

lump, Inches 2 15 S 1.5 10 1 Yield Torque 5 0.5

0 0 02468100246810 Silica Fume Replacement Level (%) Silica Fume Replacement Level (%) Yield Torque Slump Figure 7.6: Influence of Silica Fume on Rheology

212 7.5.4 Water-to-Cement Ratio The water-to-cement ratio was varied to four different values for each of the concrete mixtures. As shown in Figure 7.7, the addition of water resulted in a decrease in all rheological parameters–namely, yield value, yield stress, yield torque, viscosity value, and plastic viscosity. Likewise, the slump increased with increasing water-to-cement ratio.

213 4.5 3.5 RG_NS RG_NS 4 LS_NS LS_NS 3 3.5 2.5 3

2.5 2

2 1.5

Yield Value, Nm Value, Yield 1.5

Viscosity Value, Nm.s Value, Viscosity 1 1

0.5 0.5

0 0 0.4 0.45 0.5 0.55 0.40.450.50.55 w/c Ratio w/c Ratio Yield Value Viscosity Value 1200 50 RG_NS RG_NS LS_NS 45 LS_NS 1000 40

35 800 30

600 25

Yield Stress,Pa 400 15 Plastic Viscosity, Viscosity, Pa.s Plastic

10 200 5

0 0 0.4 0.45 0.5 0.55 0.40.450.50.55 w/c Ratio w/c Ratio Yield Stress Plastic Viscosity 60 8 RG_NS RG_NS LS_NS 7 LS_NS 50 6 40 5

30 4

3 Slump, Inches Slump,

Yield Torque, Torque, Nm Yield 20 2 10 1

0 0 0.4 0.45 0.5 0.55 0.40.450.50.55 w/c Ratio w/cm Ratio Yield Torque Slump Figure 7.7: Influence of Water-to-Cement Ratio on Rheology

214 7.5.5 Water-Reducing Admixtures The lignosulfonate-based low-range water reducer and polycarboxylate- based high-range water reducer were added based on the range of the supplier’s recommendations. The test results are shown in Figure 7.8. The dosages are expressed in fluid ounces per 100 pounds of cement (oz/cwt). Although the dosages are plotted on the same scale, the solids concentrations of the two water reducers may be different. The high-range water reducer resulted in an approximately linear decrease in yield value and yield stress. In contrast, the effect of the low-range water reducer on yield value and yield stress was most pronounced at the low dosage. The trends in plastic viscosity varied with dosage. At the low dosage of 3 oz/cwt, the plastic viscosity decreased for both aggregates and both admixtures. At 6 oz/cwt, the high-range water reducer began to increase the plastic viscosity while the effect of the low-range water reducer was mixed. When 9 oz/cwt of high-range water reducer were added to the river gravel concrete mixture, the plastic viscosity decreased. The slump increased for both aggregates; however, the high-range water reducer resulted in a greater increase in slump than the low-range water reducer.

215 4.5 3 RG_NS_HRWR 4 LS_NS_HRWR RG_NS_WRA 2.5 3.5 LS_NS_WRA

3 2

2.5 1.5 2

Yield Value, Nm 1.5 1 Viscosity Value,Nm.s 1 RG_NS_HRWR 0.5 LS_NS_HRWR 0.5 RG_NS_WRA LS_NS_WRA 0 0 0246810 0246810 Dosage (oz/cwt) Dosage (oz/cwt) Yield Value Viscosity Value 900 45 RG_NS_HRWR 800 LS_NS_HRWR 40 RG_NS_WRA 700 LS_NS_WRA 35

600 30

500 25

400 20

Yield Stress,Pa 300 15 Plastic Viscosity, Pa.s Viscosity, Plastic 200 10 RG_NS_HRWR LS_NS_HRWR 100 5 RG_NS_WRA LS_NS_WRA 0 0 0246810 0246810 Dosage (oz/cwt) Dosage (oz/cwt) Yield Stress Plastic Viscosity 45 12 RG_NS_HRWR RG_NS_HRWR 40 LS_NS_HRWR LS_NS_HRWR RG_NS_WRA 10 RG_NS_WRA 35 LS_NS_WRA LS_NS_WRA

30 8

25 6 20 lump, Inches S

Yield Torque, Torque, Nm Yield 15 4

10 2 5

0 0 02468100246810 Dosage (oz/cwt) Dosage (oz/cwt) Yield Torque Slump Figure 7.8: Influence of Water Reducers on Rheology

216 7.5.6 Air-Entraining Agent Air-entraining agent was added to produce a range of air contents typical for concrete. The effect of entrained air content depended on the concrete mixture, as shown in Figure 7.9. The yield value and yield stress increased for the river gravel concrete mixture but decreased for the crushed limestone concrete mixture. The yield torque decreased with increasing air content for both concrete mixtures. The plastic viscosity was reduced at low dosages for both aggregates. When the air content was increased to about 7%, the plastic viscosities for both aggregates began to increase, although the values still remained below the control values.

217 4.5 2.5 4 2 3.5

3 1.5 2.5

2 1

Yield Value, Nm Value, Yield 1.5 Viscosity Value, Nm.s 1 0.5 RG_NS 0.5 RG_NS LS_NS LS_NS 0 0 02468 02468 Air Content (%) Air Content (%) Yield Value Viscosity Value 1200 35

30 1000

25 800

20 600 15

Yield Stress, Pa 400

Plastic Viscosity, Pa.s 10

200 5 RG_NS RG_NS LS_NS LS_NS 0 0 02468 02468 Air Content (%) Air Content (%) Yield Stress Plastic Viscosity 40 7

35 6

30 5

25 4 20 3

15 Inches Slump,

2 10 Yield Torque Growth),(Stress Nm 5 RG_NS 1 RG_NS LS_NS LS_NS 0 0 02468 02468 Air Content (%) Air Content (%) Yield Torque Slump

Figure 7.9: Influence of Air Content on Rheology

218 7.5.7 Blends of Natural and Manufactured Sand The natural river sand (NS) was blended with the limestone manufactured sand (MS) from 100% natural sand to 100% manufactured sand. The results are shown in Figure 7.10. The addition of manufactured sand resulted in an increase in yield value, yield stress, and yield torque for both coarse aggregates. For the crushed limestone coarse aggregate concrete mixture, the plastic viscosity remained approximately unchanged with increasing manufactured sand content; however, the use of manufactured sand in the river gravel concrete mixture resulted in increases in plastic viscosity. This divergence in plastic viscosity trends was not reflected in the slump test.

219 4 2.5 RG_NS RG_NS 3.5 LS_NS LS_NS 2 3

2.5 1.5

1.5 1 Yield Value, Nm Value, Yield Viscosity Value, Nm.s 1 0.5 0.5

0 0 0 20406080100 0 20406080100 Manufactured Sand (%) Manufactured Sand (%) Yield Value Viscosity Value 900 30 RG_NS 800 RG_NS LS_NS LS_NS 25 700

600 20

500 15 400

Yield Stress,Pa 300 10 Plastic Viscosity, Pa.s 200 5 100

0 0 0 20406080100 0 20406080100 Manufactured Sand (%) Manufactured Sand (%) Yield Stress Plastic Viscosity 35 8 RG_NS LS_NS 30 7

6 25

5 20 4 15 lump, Inches

S 3 10 2

Yield Torque (Stress Growth), Nm Growth), (Stress Torque Yield 5 1 RG_NS LS_NS 0 0 0 20406080100 0 20406080100 Manufactured Sand (%) Manufactured Sand (%) Yield Torque Slump Figure 7.10: Influence of Blended Sands on Rheology

220 7.5.8 Sand-to-Aggregate Ratio The sand-to-aggregate ratio (S/A) was varied from 0.3 to 0.55. The test results are shown in Figure 7.11. Three of the mixtures—the crushed limestone mixtures with S/A ratios 0.3 and 0.55 and the river gravel mixture with an S/A ratio of 0.55—were too stiff to be tested in the rheometer. The yield value and yield stress both increased with the S/A ratio. It is possible that points with lower S/A ratios, if tested, would have produced higher yield values or yield stresses, resulting in the identification of an optimum S/A ratio. The yield torque; however, did appear to indicate an optimum S/A ratio, at least for the river gravel concrete mix. For plastic viscosity, the optimum S/A ratio for the limestone mixture appeared to be 0.45. The plastic viscosity of the river gravel decreased with increasing S/A ratio. If the S/A ratio of 0.5 is an outlier, then the optimum value for the river gravel mixture could have been similar to the crushed limestone concrete mixture. If the base mixtures had been set to higher initial slumps, the variability would likely have been reduced and more points for high and low S/A ratios could have been measured. The slump clearly indicated an optimum value of S/A for the crushed limestone concrete mixture.

221 8 3.5 RG_NS 7 LS_NS 3

6 2.5 5 2 4 1.5 3 Yield Value, Nm

Viscosity Value,Nm.s 1 2

1 0.5 RG_NS LS_NS 0 0 0.25 0.3 0.35 0.4 0.45 0.5 0.55 0.25 0.3 0.35 0.4 0.45 0.5 0.55 S/A Ratio S/A Ratio Yield Value Viscosity Value 2000 50 RG_NS 1800 45 LS_NS 1600 40

1400 35

1200 30

1000 25

800 20 Yield Stress, Pa Stress, Yield 600 15 Plastic Viscosity, Pa.s Viscosity, Plastic

400 10 RG_NS 200 5 LS_NS 0 0 0.25 0.3 0.35 0.4 0.45 0.5 0.55 0.25 0.3 0.35 0.4 0.45 0.5 0.55 S/A Ratio S/A Ratio Yield Stress Plastic Viscosity 30 4.5 RG_NS 4 LS_NS 25 3.5

20 3

2.5 15 2 Slump, Inches 10 1.5

Yield Torque (StressGrowth), Nm 5 RG_NS 0.5 LS_NS 0 0 0.25 0.3 0.35 0.4 0.45 0.5 0.55 0.25 0.3 0.35 0.4 0.45 0.5 0.55 0.6 S/A Ratio S/A Ratio Yield Torque Slump Figure 7.11: Influence of Sand-to-Aggregate Ratio of Rheology

222 7.5.9 Aggregate Microfines Limestone and granite manufactured sands, each with varying contents of microfines, were tested. Both sands were sieved and reblended to the as-received gradation of the limestone manufactured sand, which conforms to ASTM C 33 requirements when the microfines content is kept below 7%. The microfines contents were then adjusted—no other changes to the gradation were made. The LS_NS base mixture was used, with the sand replaced with either the limestone manufactured sand or granite manufactured sand on a volume basis. The initial water content was adjusted such that the slump at 20% microfines was 7.5 to 10 cm (3 to 4 inches). The granite sand required a water-to-cement ratio of 0.603 to reach a slump of 9 cm (3.5 inches) while the limestone required a much lower water-to-cement ratio of 0.517 to reach a slump of 7.5 cm (3 inches). The effect of microfines content on rheology is shown in Figure 7.12. The addition of microfines resulted in an increase in yield value, yield stress, and yield torque for both aggregates. When the microfines content of the limestone concrete mixture was increased to 25% from 20%, the yield value, yield stress and yield torque decreased. The addition of microfines generally increased plastic viscosity, although the trend was not uniform. The addition of microfines resulted in decreases in slump for both sands.

223 5 3.5 Limestone 4.5 Limestone Granite 3 Granite 4

3.5 2.5

3 2 2.5 1.5 2 Yield Value, Nm 1.5 Viscosity Value,Nm.s 1 1 0.5 0.5

0 0 0 5 10 15 20 25 30 0 5 10 15 20 25 30 Microfines (%) Microfines (%) Yield Value Viscosity Value 1200 50 Limestone Limestone Granite 45 Granite 1000 40

35 800 30

tress, Pa tress, 600 25 S 20

Yield Yield 400 15 Plastic Viscosity, Pa.s Viscosity, Plastic

10 200 5

0 0 0 5 10 15 20 25 30 0 5 10 15 20 25 30 Microfines (%) Microfines (%) Yield Stress Plastic Viscosity 60 9 Limestone Limestone 8 Granite Granite 50 7 , Nm )

40 6 rowth G 5 30 tress tress

(S 4 lump, Inches 20 S 3

Yield Torque Yield Torque 10 1

0 0 0 5 10 15 20 25 30 0 5 10 15 20 25 30 Microfines (%) Microfines (%) Yield Torque Slump Figure 7.12: Influence of Microfines Content on Rheology

224 In addition to varying the microfines content, the gradation of the limestone manufactured sand was changed from the as-received gradation to a new gradation on the 0.45 power curve. Again, the sand was used in the LS_NS mixture on a direct volume replacement of the natural sand—the rest of the gradation remained unchanged. The test results, shown in Figure 7.13, indicate that improving the gradation resulted in a dramatic reduction in yield stress and a slight increase in plastic viscosity. This significant reduction in yield stress was not reflected in the slump or yield torque test results, both of which increased. This dichotomy between slump, yield torque and the flow curve yield stress was a key finding. The yield stress from the flow curve is a dynamic measurement while slump and yield torque are static measurements. The dynamic test is better suited to measuring high-microfines concrete than the yield torque or the slump test.

1.6 15% Microfines, Original Gradation 1.4 10% Microfines, Original Gradation

) 15.5% Microfines, 0.45 Power Curve

15%mf 1.2 /Mix i 1

0.8

0.6

0.4 Normalized Result: (Mix Result: Normalized 0.2

0 Slump Yield Value Viscosity Yield Stress Plastic Yield Value Viscosity Torque Figure 7.13: Influence of Changing Sand Particle Size Distribution in High- Microfines Mixtures

225 7.5.10 Slag Aggregate Five mixtures were prepared with the slag aggregate. First the LS_NS base mixture was produced with the slag aggregate by replacing all of the crushed limestone coarse aggregate with slag aggregate on a volume basis. Further changes were then made to the slag aggregate concrete mixture. The results of the testing are presented in Figure 7.14 by normalizing each slag aggregate mixture against the original LS_NS base mix. At a constant water-to-cement ratio of 0.457, the slump decreased from 13 cm (5 inches) in the LS_NS mix (Mix 1) to 4 cm (1.5 inches) in the slag aggregate mix (Mix 2). Although this first slag aggregate mix was too stiff to be tested in the rheometer, it appeared rich and did not segregate. In order to achieve a similar slump as in the original LS_NS base mixture, the water-to-cement ratio was increased to 0.525, which resulted in a 14- cm (5.5-inch) slump (Mix 3). At this higher water-to-cement ratio, the slag aggregate mixture was harsh and experienced segregation. Despite having similar slump, the mix exhibited lower yield stress and plastic viscosity. Next, the S/A ratio was changed from 0.40 (Mix 3) to 0.50 (Mix 4) and 0.35 (Mix 5) At the higher S/A ratio of 0.50, the mix appeared to exhibit better workability—it was fluid and resisted segregation—despite being sandy. This increase in sand content resulted in an increase in yield stress but a slight reduction in plastic viscosity. When the S/A ratio was reduced to 0.35, the increased quantity of coarse aggregate resulted in a very harsh mix with poor segregation resistance. Although the yield stress was reduced, the plastic viscosity was increased substantially and the slump was reduced. Finally, Mix 3 was modified by using fly ash at a 20% replacement rate (mix 6). This final mix was still harsh, although it was flowable and exhibited improved segregation resistance. Compared to Mix 3, the yield stress was reduced, while the plastic viscosity was increased. The improved segregation

226 resistance and increased average qualitative rating were likely due to this increased plastic viscosity.

1.4 1) LS_NS, w/c=0.457 2) SlagAgg, w/c=0.457 3) SlagAgg, w/c=0.525 4) SlagAgg, w/c=0.525, S/A=0.5 5) SlagAgg, w/c=0.525, S/A=0.35 6) SlagAgg, w/c=0.525, 20% Fly Ash 1.2

1 )/(Mix LS_NS) i 0.8

0.6

0.4

0.2 Normalized Result: (Mix Result: Normalized

0 Slump Yield Value Viscosity t0 m Avg. Qual. τ0 µ Value Rating

Figure 7.14: Influence of Slag Aggregate

Although the use of slag aggregate did reduce the workability when replaced on a direct volume basis, the use of the ICAR rheometer provided insights into ways to improve the workability. In this case, increasing the sand content or adding fly ash improved the overall workability.

227 7.6 RHEOMETER PERFORMANCE In addition to comparing the effects of various mixture changes, the data generated in the testing of the conventional concrete mixtures were used to assess the performance of the ICAR rheometer.

7.6.1 Repeatability of Test Results In any testing, three main sources of variance contribute to the overall variance of test results: sampling, testing, and material variance. In an ideal test, all variance should be attributable to material variance, that is, changes in material properties. An advantage of rheology is that it considers multiple points to plot one flow curve, instead of a single point for other common field tests such as slump and air content. High values of the coefficient of determination of the straight line fit between torque and rotation speed can indicate not only that the material exhibits linear behavior, but that the testing variance is low. The R2 values for the straight line fit of torque versus rotation speed are summarized in Table 7.6. The value of R2 was above 0.90 for 85% of all tests. The values of R2 are comparable for the first and second tests. Therefore, the ICAR rheometer does accurately measure the rheology of the material in the container; however, this material can change due to sampling errors or unintentional material variance. In particular, the shear history of the sample can have a significant influence on test results.

Table 7.6: Coefficients of Determination from Flow Curve Measurements Test 1 Test 2 All Tests Maximum 0.999 0.999 0.999 Minimum 0.574 0.654 0.574 Average 0.937 0.942 0.940 Median 0.958 0.961 0.958 Number of Tests 84 83 167

228 To consider further the effect of shear history, the rheological parameters were compared from the first test to the second test. Although the same concrete was present in the container for the first and second tests, the material was remixed with a shovel between tests. Figure 7.15 presents a comparison of rheometer results from the first test to the second test for yield value, viscosity value, yield stress, plastic viscosity, and yield torque. The plots indicate that the test results were strongly dependent on the shear history of the sample, especially for measurements of viscosity and yield torque. Several features of the plots are noteworthy. First, the measurements of all parameters generally decreased from the first test to the second test, suggesting that additional structural breakdown, segregation, or both occurred between tests. Second, the magnitude of decrease from the first test to the second test was greatest for the yield torque, which is the most sensitive to structural breakdown. The first yield torque test was performed with no structural breakdown, while the second test took place after the intensive shearing of the first test and the remixing. Third, the variability in measurements from the first test to the second test was greatest for the viscosity value and plastic viscosity. This greater variability is not surprising, given the fact that rotational rheometers are typically less accurate than capillary viscometers in measuring viscosity (Whorlow 1992). Since the mixtures were manually remixed, as opposed to being remixed by mechanical means, some variability is to be expected. Fourth, the variance from the line of equality was greatest at high values of the rheological parameters. This trend indicates that the rheometer is better suited for low yield stress concrete materials, which behave more like homogenous fluids and typically exhibit less structural breakdown and segregation during shearing.

229 8 3.5

7 3

6 2.5

5 2 4 1.5 3

1 2 Yield Value-Second Test, Nm Test, Yield Value-Second Viscosity Value-Second Test, Nm.s 1 0.5

0 0 0246800.511.522.533.5 Yield Value-First Test, Nm Viscosity Value-First Test, Nm.s Yield Value Viscosity Value 2500 60

50 2000

40 1500

1000 20

Yield Stress-Second Test, Pa Test, Yield Stress-Second 500 10 Plastic Viscosity-Second Test, Pa.s

0 0 0 500 1000 1500 2000 2500 0 102030405060 Yield Stress-First Test, Pa Plastic Viscosity-First Test, Pa.s Yield Stress Plastic Viscosity 60

20 Yield Torque-Second Test, Nm Test, Torque-Second Yield 10

0 0 102030405060 Yield Torque-First Test, Nm Yield Torque Figure 7.15: Comparison of Rheometer Results from First and Second Tests

230 Since shear history is an important consideration, particularly for high yield stress and high viscosity concrete mixtures, it is important to consider ways to achieve a uniform shear history. In all of the testing described herein, the concrete was loaded into the rheometer container in a consistent manner and the structural breakdown was performed consistently before the first flow curve measurement. Such a testing regime decreased any variability due to shear history for the first test. The same level of variability was not assured for the second or any subsequent tests. One option to reduce variability in shear history would be to consolidate the concrete fully by rodding or vibration. Of the existing concrete rheometers, only the BTRHEOM rheometer includes a vibrator to consolidate concrete. While the lack of full consolidation does likely lead to additional variability in test results, the measurement of fully consolidated concrete is disadvantageous for three reasons. First, concrete on a jobsite flows in an unconsolidated state and is not fully consolidated until it reaches its final location. Therefore, the flow of fully consolidated concrete is not directly relevant to construction operations. Second, the torque required from the rheometer to measure fully consolidated concrete would be substantially higher than that required for partially consolidated concrete. Third, the rotation of the vane would reduce the degree of consolidation in fully consolidated concrete, thereby resulting in a zone of loose concrete surrounded by denser concrete.

7.6.2 Calculation of Test Results The data presented in Section 7.5 indicate that trends in rheology can vary depending on whether relative parameters of yield value and viscosity value or absolute parameters of yield stress and plastic viscosity are considered. Figure 7.16 shows that yield value and yield stress are linearly correlated, with the largest deviations from the linear fit at high values of yield stress and yield value. In contrast, the relationship between viscosity value and plastic 231 viscosity, shown in Figure 7.17, is approximately linear but shows wider scatter. The presence of a dead zone, the dimensions of which are a function of the yield stress and the rotation speed, means that the relationship between rotation speed and shear rate is not linear, as it would be in a traditional coaxial cylinders rheometer where all material flows. The nonlinearity in the relationship between rotation speed and shear rate mainly influences the slope of the flow curve and not the intercept with the shear stress axis. Hence, the scatter is larger between viscosity value and plastic viscosity than between yield value and yield stress.

2500

2000

1500

1000

Yield Stress, Pa Stress, Yield y = 255.59x - 82.132 R2 = 0.9685 500

0 0246810 Yield Value, Nm Figure 7.16: Comparison of Yield Value and Yield Stress (First Test)

232 60

50 y = 14.785x - 1.7327 R2 = 0.6805

20 Plastic Viscosity, Pa.s Viscosity, Plastic

0 00.511.522.533.5 Viscosity Value, Nm.s Figure 7.17: Comparison of Viscosity Value and Plastic Viscosity (First Test)

The use of the Effective Annulus Method to calculate yield stress and plastic viscosity is based on the assumptions that the marginal effective height of the vane due to end effects is negligible and, by extension, that only flow within the annulus is relevant. Visual observations of the concrete in the rheometer during and after the test indicated that flow likely occurred outside the annulus. It was difficult, however, to assess what material flowed and what material deformed elastically visually. The assumption that the distinct radius at which flow ceases can be determined precisely in any coarse-grained material, especially a material with the sizes of aggregates present in concrete, is questionable. In reality, the flow pattern is not precisely as assumed, although the deviation is unknown. The data presented herein are insufficient to conclude which of the two approaches—relative parameters or absolute parameters—provides the better indication of rheological properties. Without further experimental testing or

233 computer modeling, it is not possible to describe the true flow pattern. Therefore, both relative and absolute parameters have been presented. While both calculation methods have clear limitations, neither method appears to be unreasonable. Figure 7.18 indicates that the yield stress measured from the flow curve is a different entity than the yield stress measured in the stress growth test. In general, the yield value was approximately an order of magnitude less than the yield torque. The yield torque is a static test conducted without any prior structural breakdown and is influenced by the degree of thixotropy of the mixture. In contrast, additional stress due to structural breakdown and thixotropy are minimized for the yield stress determined from the flow curve, which is a dynamic measurement. Though different, both values are useful. The yield torque indicates the amount of force needed to initiate concrete from rest, while the yield value and flow curve yield stress indicate the amount of stress that must be continuously applied to keep the concrete moving after structural breakdown.

8 y = 0.0718x + 1.0474 R2 = 0.2624 7

4 Yield Value, Nm Yield Value, 3

0 0 102030405060 Yield Torque, Nm Figure 7.18: Comparison of Yield Torque and Yield Stress (First Test)

234

7.7 RELATIONSHIPS BETWEEN RHEOLOGY AND WORKABILITY In order to relate the measurements from the ICAR rheometer to practical workability requirements, the measured rheological parameters were related to the visual observations of each concrete mixture. Figure 7.19 indicates that no correlation existed between yield stress and plastic viscosity for the concrete tested. Therefore, it is necessary to consider both parameters in assessing workability.

20 Plastic Viscosity, Pa.s Viscosity, Plastic

0 0 500 1000 1500 2000 2500 Yield Stress, Pa Figure 7.19: Yield Stress and Plastic Viscosity Measurements for All Conventional Concrete Tested

If workability is to be assessed with only two variables, yield stress and plastic viscosity, then it should be possible to define ranges of yield stress and plastic viscosity that produce good workability. The concept of a workability box, which was discussed earlier as a possible approach, was implemented using the qualitative observations of each concrete mixture. Instead of qualitative 235 observations, it would also have been possible to use other quantitative tests— such as for slump, bleeding, or segregation—to define the workability boxes. The workability boxes were constructed for combinations of yield value and viscosity value and of yield stress and plastic viscosity. Figure 7.20 shows the workability boxes based on the segregation resistance rating. A segregation rating of five is considered good; a rating of 3-4, acceptable; and a rating less than 3, bad. It is possible to draw a rectangle around a zone of good or acceptable points. Although each box does include one bad mix and omits several good mixes, it is possible to conclude that mixtures falling within the workability boxes for segregation resistance will likely be resistant to segregation.

3.5 60 Good (=5) Acceptable (=3, 4) 3 50 Bad (=1, 2)

2.5 40

2 30 1.5

20 Viscosity Value, Nm.s Value, Viscosity 1 Pa.s Viscosity, Plastic

Good (=5) 10 0.5 Acceptable (=3, 4) Bad (=1, 2) 0 0 02468100 200 400 600 800 1000 1200 1400 1600 Yield Value, Nm Yield Stress, Pa Figure 7.20: Workability Boxes for Segregation Resistance Rating

Figure 7.21 shows the workability boxes for the overall workability rating given to each concrete mixture. For the plot of relative parameters, it was necessary to draw two boxes covering adjacent regions. Like the workability boxes for segregation, the workability boxes for the overall rating omit some good mixes and include some bad mixes; however, they do provide an acceptable definition of good or acceptable workability. The criteria are slightly different for

236 the overall rating workability boxes—a rating of four or higher constitutes a good rating.

3.5 60 Good (>=4) Acceptable (=3) 3 50 Bad (<=2)

2.5 40

2 30 1.5

20 Viscosity Value, Nm.s Value, Viscosity 1 Pa.s Viscosity, Plastic

Good (=5) 10 0.5 Acceptable (=3, 4) Bad (=1, 2) 0 0 02468100 200 400 600 800 1000 1200 1400 1600 Yield Value, Nm Yield Stress, Pa Figure 7.21: Workability Boxes for Overall Workability Rating

Figure 7.22 shows the workability boxes plotted for the average of the six qualitative ratings. The zones of good or acceptable workability are not as clear as in the previous workability boxes for segregation and overall rating.

3.5 60 Good (=5) Acceptable (=3, 4) 3 50 Bad (=1, 2)

2.5 40

2 30 1.5

20 Viscosity Value, Nm.s Value, Viscosity 1 Pa.s Viscosity, Plastic

Good (>=4) 10 0.5 Acceptable (3-4) Bad (<=3) 0 0 02468100 200 400 600 800 1000 1200 1400 1600 Yield Value, Nm Yield Stress, Pa Figure 7.22: Workability Boxes for Average of All Ratings

237 The workability boxes developed above cover similar areas. It would be possible to combine workability boxes from a variety of criteria to improve the definition. Several important caveats apply to the implementation of the workability boxes shown above. First, the definitions of desirable workability were general. For a particular application, the requirements for workability will be much more specific. Second, an extremely wide range of concrete mixtures was considered. In practice, the workability box should be defined over much narrower ranges of mixture proportions and materials. Third, the qualitative ratings were based on observations, which were subject to human error and likely varied from day to day. To utilize a workability box successfully, qualitative observations, if used, must be well defined in terms of thresholds for various levels of workability. Fourth, it is beneficial to use several workability boxes for different features of the mixture. It is desirable to know more about the mix than just whether it is good or bad. For example, it would be helpful to know such information as whether a concrete mixture is too stiff to flow readily or highly fluid and prone to segregation. Multiple workability boxes could ultimately be combined into a single workability box for all criteria. Fifth, the workability box does not need to be a rectangle. Other shapes may be more accurate in defining zones of desirable workability; however, the rectangle is simple to understand and easy to use in specifications. The use of a workability box is appealing for its apparent simplicity. If used carefully as shown above, the workability box can be effective in enabling the utilization of only two parameters to describe concrete workability. While rheology is largely a new concept to the concrete industry, the concept of a workability box provides an initial framework for the use of the ICAR rheometer in developing mixture designs and assessing field performance.

238 7.8 SELF-CONSOLIDATING CONCRETE A separate series of self-consolidating concrete mixtures was tested. The ICAR rheometer was used to measure the influence of the following factors: initial water content, high-range water reducer dosage, viscosity modifying admixture dosage, fly ash content, total cementitious materials content, and aggregate gradation.

7.8.1 Mixture Proportions and Test Procedures Thirteen self-consolidating concrete mixtures were prepared. Each mixture included an initial dosage of admixture. After all fresh concrete tests were performed on this first mixture, the tested concrete was placed back into the mixer, the admixture dosage was increased, and the concrete was remixed and then retested. As a result, up to three or four test tests were performed for each mixture. The mixture proportions for the base concrete mixtures are shown in Table 7.7. In order to ensure consistent mixing energy, all concrete mixtures were produced in 0.113 m3 (4 ft3) batches. The batching and mixing procedures were identical to those used for the conventional concrete mixtures, with several modifications. Only the retarder, which did not vary between mixes, was added with the mixing water. The concrete with the retarder but no other admixtures was mixed for three minutes so that all constituents would be blended together. The mixer was then stopped and the initial dosages of high-range water reducer and, if needed, viscosity modifying admixture were added. The concrete was then mixed for three minutes, allowed to rest for three minutes, and mixed for another two minutes. A sufficient quantity of concrete for the fresh concrete tests was discharged from the mixer. After the tested concrete was placed back into the mixer and subsequent dosages of admixtures were added, the concrete was mixed for three minutes, allowed to rest

239 for three minutes, mixed for two minutes, and discharged from the mixer for testing.

Table 7.7: Self-Consolidating Concrete Mixture Proportions Aggregate (SSD) Initial Mix Change Cement Fly Ash w/cm Retarder Coarse Int Fine HRWR kg/m3 kg/m3 kg/m3 kg/m3 kg/m3 mL/cwt mL/cwt

(lb/yd3) (lb/yd3) (lb/yd3) (lb/yd3) (lb/yd3) (oz/cwt) (oz/cwt) 564.2 303.8 861.0 385.6 326 130 1 Reduce cement 0.40 (951.1) (512.1) (1451.3) (650.0) (5) (2.0) 677.8 274.8 673.0 445.0 391 130 2 Coarser gradation 0.40 (1142.4) (463.1) (1134.3) (750.0) (6) (2.0) 523.3 281.8 798.6 439.3 261 130 3 w/cm = 0.43 0.43 (882.1) (474.9) (1346.0) (740.5) (4) (2.0) 530.1 285.4 808.9 445.0 391 130 4 w/cm = 0.40 0.40 (893.5) (481.1) (1363.4) (750.0) (6) (2.0) 542.1 291.9 827.3 455.1 456 130 5 w/cm = 0.35 0.35 (913.8) (492.0) (1394.5) (767.1) (7) (2.0) 495.9 267.0 756.8 504.3 261 130 6 Increase cement 0.40 (835.9) (450.1) (1275.5) (850.0) (4) (2.0) 530.1 285.4 808.9 356.0 89.0 326 130 7 20% fly ash 0.40 (893.5) (481.1) (1363.4) (600.0) (150.0) (5) (2.0) 530.1 285.4 808.9 267.0 178.0 196 130 8 40% fly ash 0.40 (893.5) (481.1) (1363.4) (450.0) (300.0) (3) (2.0) 530.1 285.4 808.9 445.0 522 130 9 VMA 0.40 (893.5) (481.1) (1363.4) (750.0) (8) (2.0) 20% fly ash, 542.8 292.3 828.3 364.5 91.1 326 130 10 0.32 w/cm = 0.32 (914.9) (492.6) (1396.1) (614.4) (153.6) (5) (2.0) 537.3 289.3 819.8 451.0 391 130 11 w/cm = 0.37 0.37 (905.6) (487.6) (1381.9) (760.2) (6) (2.0) Manufactured 530.1 285.4 845.2 445.0 522 130 12 0.40 sand (MS) (893.5) (481.1) (1424.6) (750.0) (8) (2.0) 537.3 289.3 819.8 451.0 717 130 13 Repeat mix 11 0.37 (905.6) (487.6) (1381.9) (760.2) (11) (2.0) (mL/cwt = mL/100 kg cementitious materials, oz/cwt = oz/100 lb cementitious materials) Coarse aggregate = crushed limestone (LS) Intermediate aggregate = crushed limestone (LSI) Fine Aggregate = natural sand (NS), except Mix 12

Rheological measurements were made with the ICAR rheometer and the BTRHEOM rheometer. The concrete was discharged directly from the mixer into the ICAR rheometer container, while concrete was loaded into the BTRHEOM with a scoop. The test regime for the ICAR rheometer consisted of a 25-second

240 breakdown period at 1.0 rev/sec. Seven flow curve points were measured in descending order from 1.0 rev/sec to 0.05 rev/sec. Each point was measured for 5 seconds. For the BTRHEOM rheometer, five flow curve points were made in descending order from 1.0 rev/sec to 0.2 rev/sec. A malfunction of the torque transducer in the BTRHEOM resulted in errors in the yield stress measurements. The voltage offset for zero torque shifted slightly for each test. As the voltage offset shifted upwards or downwards, the torque versus rotation speed plot—and thus the yield stress—likewise shifted. Whereas the calibration coefficient for incremental torque (in volts/Nm) did not fluctuate, any shift in the zero torque offset during a test could have influenced the slope of the torque versus rotation speed plot—and thus the viscosity. The error in viscosity is believed to have been inconsequential; however, the error in yield stress cannot be neglected. Some fluctuation in the zero torque offset is to be expected. In fact, the BTRHEOM operation software includes a feature to correct for any shift in the zero offset calibration. Even with this feature, the fluctuations were excessive. The effect of any fluctuation is most significant for low-yield stress concrete, where the offset in yield stress is greatest in percentage terms. The rheometer results were compared to the slump flow test. For the slump flow test, concrete was poured into the slump cone, which was centered on a level, plastic plate. No rodding of the concrete was needed. The slump cone was lifted and three measurements were made: the time for the concrete to flow to a diameter of 50 cm (T50), the final horizontal spread, and the visual stability index (VSI). The VSI ratings, which are based on the definition of Daczko (2002), are shown in Table 7.8. The concrete was considered to be SCC if it exhibited a

minimum slump flow of 25 inches, a VSI of 1 or less, and a T50 between 2 and 7 seconds.

241 Table 7.8: Visual Stability Index Ratings (Daczko 2002) VSI Criteria No evidence of segregation in slump flow patty or in mixer drum or 0 wheelbarrow. No mortar halo or aggregate pile in the slump flow patty but some slight 1 bleed or air popping on the surface of the concrete in the mixer drum or wheelbarrow. A slight mortar halo (< 10 mm) and/or aggregate pile in the slump flow 2 patty and highly noticeable bleeding in the mixer drum and wheelbarrow. Clearly segregating by evidence of a large mortar halo (>10 mm) and/or a large aggregate pile in the center of the concrete patty and a thick layer of 3 paste on the surface of the resting concrete in the mixer drum or wheelbarrow.

7.8.2 Test Results For each mixture change—including water-to-cement ratio, cement content, fly ash content, and sand type—the effects on fresh concrete properties are shown in terms of high-range water reducer dosage. The test data for all mixtures are located in Appendix B. Figure 7.23 shows the influence of the water-to-cement ratio on SCC rheology. Whereas the results for yield stress exhibited considerable variability, the trends for plastic viscosity were clear and consistent between the ICAR rheometer and the BTRHEOM rheometer. The variability in yield stress measurements was due to the relatively low values of yield stress. As long the yield stress was kept sufficiently low, plastic viscosity was the more important parameter. Each increase in the water-to-cement ratio not only reduced the viscosity, it also decreased the potency of the high-range water reducer in terms of viscosity.

242 1.6 6 w/c=0.35 w/c=0.35 1.4 w/c=0.37 w/c=0.37 w/c=0.40 5 w/c=0.40 1.2 w/c=0.43 w/c=0.43

1 4

0.8 3 0.6

0.4 2 Yield Value, Nm 0.2 Nm.s Value, Viscosity

0 1 024681012 -0.2 0 -0.4 024681012 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt ICAR - Yield Value ICAR - Viscosity Value 350 120 w/c=0.35 w/c=0.35 w/c=0.37 w/c=0.37 300 w/c=0.40 100 w/c=0.40 w/c=0.43 w/c=0.43 250 80

200 60 150

Yield Stress, Pa Stress, Yield 40

100 Pa.s Viscosity, Plastic

20 50

0 0 0 2 4 6 8 10 12 024681012 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt ICAR - Yield Stress ICAR - Plastic Viscosity 800 250 w/c=0.35 w/c=0.35 w/c=0.37 w/c=0.37 600 w/c=0.40 w/c=0.40 200 w/c=0.43 w/c=0.43

400 150

200 100 Yield Stress, Pa Stress, Yield

0 Plastic Viscosity, Pa.s 024681012 50 -200

0 -400 024681012 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt BTRHEOM - Yield Stress BTRHEOM – Plastic Viscosity Figure 7.23: Influence of Water-to-Cement Ratio on SCC Rheology

243 The influence of the water-to-cement ratio on slump flow is indicated in Figure 7.24. In general, as the water-to-cement ratio was decreased, the dosage of high-range water reducer required to achieve a slump flow of 25 inches was increased. The trend between water-to-cement ratio and high-range water reducer dosage for a 25-inch slump flow was not consistent—the high-range water reducer dosage was approximately 7 oz/cwt for a w/c of 0.40 or 0.43 but then jumped to approximately 9.5-10 oz/cwt for a w/c of 0.35 or 0.37. This variability complicates the selection of a high-range water-reducing admixture dosage for achieving a consistent slump flow.

40 w/cm=0.35 35 w/cm=0.37 w/cm=0.40 30 w/cm=0.43

15 Slump Flow,Inches 10

0 024681012 HRWR Dosage, oz/cwt Figure 7.24: Influence of Water-to-Cement Ratio on Slump Flow

Figure 7.25 demonstrates that increasing the cement content reduced yield stress and plastic viscosity. The cement content of 386 kg/m3 (650 lb/yd3) is below the typical range of cement contents for SCC while the value of 504 kg/m3 (850 lb/yd3) is above the typical range cement contents. The potency of the high- range water reducer was lower at the higher cement contents.

244 3 2.5 386 kg (650 lb) 386 kg (650 lb) 445 kg (750 lb) 445 kg (750 lb) 2.5 504 kg (850 lb) 504 kg (850 lb) 2

2 1.5

1.5

Yield Value, Nm 1 Viscosity Value, Nm.s Value, Viscosity

0.5 0.5

0 0 024681012024681012 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt ICAR - Yield Value ICAR - Viscosity Value 700 35 386 kg (650 lb) 386 kg (650 lb) 445 kg (750 lb) 445 kg (750 lb) 600 30 504 kg (850 lb) 504 kg (850 lb)

500 25

400 20

300 15 Yield Stress, Pa Stress, Yield

200 Pa.s Viscosity, Plastic 10

100 5

0 0 0 2 4 6 8 10 12 024681012 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt ICAR - Yield Stress ICAR - Plastic Viscosity 2000 120 386 kg (650 lb) 386 kg (650 lb) 445 kg (750 lb) 445 kg (750 lb) 504 kg (850 lb) 100 504 kg (850 lb) 1500

1000 60

500 40 Yield Stress, Pa.s Stress, Yield Plastic Viscosity, Pa.s Viscosity, Plastic

0 20 024681012

0 -500 024681012 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt BTRHEOM - Yield Stress BTRHEOM – Plastic Viscosity Figure 7.25: Influence of Cement Content on SCC Rheology

245 The effect of cement content on slump flow is shown in Figure 7.26. The dosage of HRWR to achieve a 25-inch slump flow increased from 4 oz/cwt at the high cement content to 11 oz/cwt at the low cement content. When enough HRWR was added to bring each mixture to a 25-inch slump flow, the yield stresses were similar; however, the viscosities varied widely, with the high cement content exhibited a much lower viscosity. Figures 7.25 and 7.26 can be used to consider the tradeoffs between adding cement and adding HRWR.

40 386 kg (650 lb) 35 445 kg (750 lb) 504 kg (850 lb) 30

15 Slump Flow, inches 10

0 024681012 HRWR Dosage, oz/cwt Figure 7.26: Influence of Cement Content on Slump Flow

Figure 7.27 indicates that increasing the fly ash content from 0 to 20 or 40% resulted in decreased plastic viscosity. The influence of fly ash content on yield stress was variable because most values of yield stress were close to zero. Given the high fluidity of the fly ash mixtures at a w/cm of 0.35, the w/cm was decreased to 0.32, which increased both Bingham parameters such that the plastic viscosity was above and the yield stress was below the cement-only control mix for comparable HRWR dosages. This trend suggests that fly ash can be used to lower yield stress while avoiding undesirably large decreases in viscosity.

246 3.5 9 PC only, w/cm=0.35 PC only, w/cm=0.35 20% FA, w/cm=0.35 8 20% FA, w/cm=0.35 3 40% FA, w/cm=0.35 40% FA, w/cm=0.35 20% FA, w/cm=0.32 7 20% FA, w/cm=0.32 2.5 6 2 5

1.5 4

1 3 Yield Value, Nm Viscosity Value, Nm.s Value, Viscosity

0.5 2

1 0 0 2 4 6 8 10 12 0 -0.5 024681012 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt ICAR - Yield Value ICAR - Viscosity Value 500 180 PC only, w/cm=0.35 PC only, w/cm=0.35 450 20% FA, w/cm=0.35 160 20% FA, w/cm=0.35 40% FA, w/cm=0.35 40% FA, w/cm=0.35 400 20% FA, w/cm=0.32 140 20% FA, w/cm=0.32 350 120 300 100 250 80 200

Yield Stress, Pa Stress, Yield 60 150 Plastic Viscosity, Pa.s Viscosity, Plastic 40 100

50 20

0 0 0 2 4 6 8 10 12 024681012 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt ICAR - Yield Stress ICAR - Plastic Viscosity 2500 250 PC only, w/cm=0.35 PC only, w/cm=0.35 20% FA, w/cm=0.35 20% FA, w/cm=0.35 2000 40% FA, w/cm=0.35 40% FA, w/cm=0.35 200 20% FA, w/cm=0.32 20% FA, w/cm=0.32

1500 150

1000 100 Yield Stress, Pa Stress, Yield

500 Pa.s Viscosity, Plastic

50 0 024681012 0 -500 024681012 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt BTRHEOM - Yield Stress BTRHEOM – Plastic Viscosity Figure 7.27: Influence of Fly Ash on SCC Rheology

247 Figure 7.28 indicates that the addition of fly ash reduced the HRWR dosage required to achieve a 25-inch slump flow. The change was most notable from the cement-only control mixture to the mixture with a 20% fly ash replacement. The 20% fly ash mix with a w/cm of 0.32 produced a greater slump flow for the same HRWR dosage than the cement-only control mix, further indicating the advantage of using fly ash.

Slump Flow,Inches 10 PC only, w/cm=0.35 20% FA, w/cm=0.35 5 40% FA, w/cm=0.35 20% FA, w/cm=0.32 0 024681012 HRWR Dosage, oz/cwt Figure 7.28: Influence of Fly Ash on Slump Flow

Figure 7.29 shows the effect of replacing the natural sand (NS) with a limestone manufactured sand (MS) on a direct volume basis. The use of manufactured sand reduced both yield stress and plastic viscosity. Although SCC can still be achieved, the HRWR dosage to achieve a slump flow of 25 inches was increased from 7 oz/cwt to approximately 12 oz/cwt, as indicated in Figure 7.30.

248 1.2 3.5 NS, w/cm=0.40 NS, w/cm=0.40 MS, w/cm=0.40 MS, w/cm=0.40 3 1

2.5 0.8

2 0.6 1.5

Yield Value, Nm 0.4 Viscosity Value, Nm 1

0.2 0.5

0 0 0 2 4 6 8 10 12 14 02468101214 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt ICAR - Yield Value ICAR - Viscosity Value 180 70 NS, w/cm=0.40 NS, w/cm=0.40 160 MS, w/cm=0.40 MS, w/cm=0.40 60 140 50 120

100 40

80 30

Yield Stress, Pa Stress, Yield 60

Plastic Viscosity, Pa.s Viscosity, Plastic 20 40 10 20

0 0 02468101214 02468101214 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt ICAR - Yield Stress ICAR - Plastic Viscosity 800 140 NS, w/cm=0.40 NS, w/cm=0.40 MS, w/cm=0.40 120 MS, w/cm=0.40 600

100 400 80

200 60 Yield Stress, Pa Stress, Yield

0 Pa.s Viscosity, Plastic 40 0 2 4 6 8 10 12 14

20 -200

0 -400 02468101214 HRWR Dosage, oz/cwt HRWR Dosage, oz/cwt BTRHEOM - Yield Stress BTRHEOM – Plastic Viscosity Figure 7.29: Influence of Sand Type on SCC Rheology

249

40 NS, w/cm=0.40 35 MS, w/cm=0.40

15 Slump Flow, Inches 10

0 02468101214 HRWR Dosage, oz/cwt Figure 7.30: Influence of Sand Type on Slump Flow

7.8.3 Relationships between Rheology and Workability As with conventional concrete, the rheological measurements must be related to practical field applications. Figure 7.30 indicates that no correlation existed between yield stress and plastic viscosity as measured by the ICAR rheometer; therefore, both of these independent parameters should be considered.

250 180

160

140

120

100

60 Plastic Viscosity, Pa.s 40

0 0 100 200 300 400 500 600 700 Yield Stress, Pa Figure 7.31: Yield Stress and Plastic Viscosity from ICAR Rheometer for All SCC Measurements

It is first important to consider what constitutes an acceptable SCC mixture. Figure 7.32 shows the results of a slump flow test for an excellent SCC mixture. The aggregates are uniformly distributed across the concrete sample. The lack of segregation is evident in the fact that no halo exists at the edge of the slump flow test specimen, as indicated in Figure 7.33. In contrast, a mortar halo was evident in the concrete mixture shown in Figure 7.34, resulting in a VSI rating of 2. The mix shown in Figure 7.35 exhibited severe segregation, and was rated with a VSI of 3. The mixture had a large mortar halo. Aggregate piling was evident at the center of the slump flow specimen and in a ring near the perimeter of the specimen. The excellent self-consolidating concrete mixture shown in Figure 7.32 and Figure 7.33 had a yield stress of 17.8 Pa and a plastic viscosity of 56.0 Pa.s. In contrast, the mix with the mortar halo in Figure 7.34 exhibited a higher yield stress of 49.2 Pa and a lower plastic viscosity of 18.5 Pa.s. For the poor mix shown in Figure 7.35, the yield stress was 3.56 Pa and the plastic viscosity was 34.0 Pa.s. In selecting an SCC mix, both the viscosity of the

251 concrete and the mortar should be considered. Segregation of coarse aggregates is a function of the yield stress and plastic viscosity of the mortar. At lower concrete yield stresses, the viscosity of the mortar must be increased to reduce segregation. This increased mortar viscosity should result in higher concrete viscosity.

Figure 7.32: Slump Flow Test for Excellent SCC Mixture (Mix 5, Test 4, Slump Flow = 27 Inches, VSI = 0, T50 = 5 sec)

252

Figure 7.33: No Segregation Evident in Slump Flow Test (Mix 5, Test 4, Slump Flow = 27 Inches, VSI=0)

Figure 7.34: Slight Mortar Halo Evident in Slump Flow Test (Mix 8, Test 2, Slump Flow = 30.5 Inches, VSI=2)

253

Figure 7.35: Severe Segregation (Mix 13, Test 1, Slump Flow = 31 Inches, VSI = 3)

Figure 7.36 indicates that a correlation existed between yield stress and slump flow for the ICAR rheometer, while no correlation existed between viscosity and slump flow. The large scatter in the plot of slump flow versus yield stress for the BTRHEOM was likely due to the torque transducer offset error in the BTRHEOM. Based on the experimental data, to achieve a slump flow of 25 to 30 inches, the yield stress, recorded by the ICAR rheometer, should be below 50 Pa. Simply achieving a yield stress below 50 Pa, however, does not ensure that the mix will be SCC.

Figure 7.37 shows the relationship between T50 and the rheological

parameters. Since T50 indicates the speed that the concrete flows, it is reasonable

to assume that T50 should be related to viscosity. While there was no correlation

between T50 and yield stress, a correlation does appear to exist between T50 and viscosity, although the scatter is large.

254 40 40

35 35

30 30

25 25

20 20

15 15 Slump Flow, Inches Slump Flow, Inches 10 10

5 5

0 0 -0.5 0 0.5 1 1.5 2 2.5 3 02468 Yield Value, Nm Viscosity Value, Nm.s ICAR - Yield Value ICAR - Viscosity Value 40 40

35 35

30 30

25 25

20 20

15 15 Slump Flow, Inches Slump Flow, Inches 10 10

5 5

0 0 0 100 200 300 400 500 0 50 100 150 200 Yield Stress, Pa Plastic Viscosity, Pa.s ICAR - Yield Stress ICAR - Plastic Viscosity 40 40

35 35

30 30

25 25

20 20

15 15 Slump Flow, Inches Slump Flow, Inches 10 10

5 5

0 0 -500 0 500 1000 1500 0 50 100 150 200 250 Yield Stress, Pa Plastic Viscosity, Pa.s BTRHEOM - Yield Stress BTRHEOM – Plastic Viscosity Figure 7.36: Relationships between Slump Flow and Rheological Parameters

255 7 7

6 6

5 5

4 4

3 3 T50, Seconds T50, Seconds T50,

2 2

1 1

0 0 -0.4 -0.2 0 0.2 0.4 0.6 0.8 1 0123456 Yield Value, Nm Viscosity Value, Nm.s ICAR - Yield Value ICAR - Viscosity Value 7 7

6 6

5 5

4 4

3 3 T50, Seconds T50, Seconds T50,

2 2

1 1

0 0 0 50 100 150 050100150 Yield Stress, Pa Plastic Viscosity, Pa.s ICAR - Yield Stress ICAR - Plastic Viscosity 7 7

6 6

5 5

4 4

3 3 T50, Seconds T50, Seconds

2 2

1 1

0 0 -350 -250 -150 -50 50 150 0 50 100 150 200 250 Yield Stress, Pa Plastic Viscosity, Pa.s BTRHEOM - Yield Stress BTRHEOM – Plastic Viscosity

Figure 7.37: Relationships between T50 and Rheological Parameters

256 Based on the SCC test data, it is possible to state that an SCC mixture will exhibit a certain range of yield stress and plastic viscosity. It is not possible, however, to state that a concrete within that same range of yield stress and plastic viscosity will be SCC. In general, as yield stress is decreased, plastic viscosity should be increased to prevent stability problems. The yield stress and plastic viscosity measurements can be distorted by segregation. Since segregation cannot be detected with the ICAR rheometer, a separate test for segregation is needed to define further the fresh concrete properties that constitute excellent SCC mixtures.

7.8.4 Rheometer Performance As with the conventional concrete mixtures, it is important to assess the performance of the ICAR rheometer. The first question is the choice of relative or absolute results from the ICAR rheometer. As shown in Section 7.8.2, the trends in viscosity value and plastic viscosity were generally very similar but did vary in some instances. Figure 7.38 indicates that linear relationships existed both between yield value and yield stress and between viscosity value and plastic viscosity. The coefficient of determination for viscosity value versus plastic viscosity was much higher for the SCC mixtures than for the conventional concrete mixtures. Given the low yield stresses of the SCC mixtures, a much greater portion of the material in the annulus flowed. The reduction in size or complete elimination of the dead zone for low-yield stress mixtures resulted in less variability between the viscosity value and plastic viscosity. When all of the material in the annulus flows, the relative and absolute parameters are related simply by constants. The fact that more material was flowing is evident in picture shown in Figure 7.39. Flow lines were visible on the top of the concrete. These flow lines extended to the container, suggesting that all material in the container was flowing. However, the fact that the flow lines were evident on the surface also suggests that it is inappropriate to consider only material in the annulus—

257 material above and below the annulus flows and should also be considered. Like the testing of conventional concrete, the data from the testing of SCC are insufficient to conclude whether relative or absolute parameters provide a more accurate representation of the true rheological properties of concrete.

700 200

180 600 y = 210.13x - 7.3626 160 y = 21.681x - 2.5235 R2 = 0.9413 R2 = 0.9718 500 140

120 400 100 300 80 Yield Stress, Pa Stress, Yield 60 200 Plastic Viscosity, Pa.s 40 100 20

0 0 -1 0 1 2 3 4 0246810 Yield Value, Nm Viscosity Value, Nm Yield Value – Yield Stress Viscosity Value – Plastic Viscosity Figure 7.38: Comparison of Parameters from ICAR Rheometer

Figure 7.39: Evidence of Flow throughout ICAR Rheometer Container for SCC Mix

258 Figure 7.40 shows a comparison of the results between the ICAR rheometer and the BTRHEOM rheometer. A portion of the scatter for yield value and yield stress was likely attributable to the zero torque offset error in the BTRHEOM rheometer. The scatter for viscosity value and plastic viscosity was high even though the trends for changes in concrete proportions were generally similar. This dichotomy in performance was also identified by Ferraris and Brower (2001) for the BTRHEOM and four other concrete rheometers.

2500 250

2000 200

1500 150

1000 100

500

BTRHEOM Yield Stress, Pa Stress, Yield BTRHEOM 50 BTRHEOM Plastic Viscosity, Pa.s Viscosity, Plastic BTRHEOM 0 -1 0 1 2 3 4 0 -500 0246810 ICAR Rheometer Yield Value, Nm ICAR Rheometer Viscosity Value, Nm Yield Value Viscosity Value 2500 250

2000 200

1500 150

1000 100

500

BTRHEOM Yield Stress, Pa Stress, Yield BTRHEOM 50 BTRHEOM Plastic Viscosity, Pa.s Viscosity, Plastic BTRHEOM 0 0 100 200 300 400 500 600 700 0 -500 0 50 100 150 200 ICAR Rheometer Yield Stress, Pa ICAR Rheometer Plastic Viscosity, Pa.s Yield Stress Plastic Viscosity Figure 7.40: Comparison of Rheological Parameters from BTRHEOM and ICAR Rheometer

259 Figure 7.41 shows the variation of rheological parameters versus time for Mix 10. At the beginning of the plot, the concrete had a slump flow 28 inches and a VSI of 0. In just 18 minutes, the yield stress more than doubled from 15.6 Pa to 36.4 Pa. The decrease in viscosity was less significant—it dropped from 71.2 Pa.s to 59.6 Pa.s. Although the rate of variation in rheological properties likely varied over time and with different concrete mixtures, the changes in properties can be significant and should be taken into careful consideration when examining the results of SCC testing. The changes in rheological parameters also present a challenge in the field for maintaining proper rheology from the time the concrete is batched until the concrete is in its final location. Figure 7.42 shows similar results for Mix 12, which started with a slump flow of 29 inches and a VSI of 2. While the plastic viscosity remained approximately unchanged, the yield stress increased form 49.6 Pa to 79.2 Pa in the 11 minutes of testing. While the changes were large in relative terms for both mixes, the changes were not as significant in absolute terms, given the low yield stresses of the self-consolidating concrete mixtures. The results in Figure 7.41 and Figure 7.42 also give an indication of the repeatability of the test results. For each test, the concrete was left in the rheometer container and was not remixed between tests.

260 80

General Units 30

10 Yield Stress, Pa Plastic Viscosity, Pa.s 0 1:04 1:12 1:19 1:26 1:33 Time, h:mm Figure 7.41: Evolution of Yield Stress and Plastic Viscosity over Time (Mix 10)

40 General Units 30

10 Yield Stress, Pa Plastic Viscosity, Pa.s 0 1:12 1:19 1:26 1:33 Time, h:mm Figure 7.42: Evolution of Yield Stress and Plastic Viscosity over Time (Mix 12)

261 7.9 FIELD TESTING Field testing at a concrete batch plant was conducted to demonstrate the portability of the ICAR rheometer. The field testing was organized by the National Institute of Standards and Technology and was part of a larger study of field measurements of fresh concrete rheology (Amziane, Ferraris, and Koehler 2004). Nine concrete mixtures were sampled from a concrete truck and tested with the slump test and the ICAR rheometer. The concrete was mixed in a central batch plant in two separate batches and discharged into the concrete truck mixer, where it was subsequently modified further by adding high-range water reducer, viscosity modifying admixture, or water. The first batch of concrete from central batch plant contained enough material to fill the concrete truck mixer to half capacity. Concrete was sampled directly from the end of the mixer truck discharge chute. After five different mixtures had been tested (Series I), additional concrete was added from the central batch plant to bring the truck to full capacity. Four subsequent tests were conducted until the concrete began to set (Series II). A retarding admixture was added to the original concrete mixtures from the central batch plant to allow a maximum number of tests to be conducted. The mixture proportions are shown in Table 7.9. The admixtures were supplied from Sika Corporation, Lyndhurst, NJ. The admixtures used were Sika ViscoCrete® 6100, a polycarboxylate-based high- range water reducer meeting the requirements of ASTM C 494 for Types A and F admixtures; Sika® Stabilizer VMA, a viscosity modifying admixture; and Sika Plastiment®, a water-reducing and -retarding admixture meeting the requirements of ASTM C 494 for Types B and D admixtures. Both concretes were composed of a natural river sand and a #67 gravel. The mixtures also incorporated Type I portland cement and ground granulated blast furnace slag.

262

Table 7.9: Mixture Proportions for Field Testing Mass, kg/m3 (lb/yd3) Admixtures, L/m3 (oz/cwt) Mix Coarse Fine Cement Slag Water Retarder HRWR VMA 1182.0 879.7 174.9 174.3 91.0 0.92 0 0 1-1 (1992.3) (1482.8) (294.9) (293.8) (153.3) (4.0) 1182.0 879.7 174.9 174.3 91.0 0.92 1.96 0 1-2 (1992.3) (1482.8) (294.9) (293.8) (153.3) (4.0) (8.6) 1182.0 879.7 174.9 174.3 91.0 0.92 3.91 0 1-3 (1992.3) (1482.8) (294.9) (293.8) (153.3) (4.0) (17.2) 1182.0 879.7 174.9 174.3 91.0 0.92 4.75 0 1-4 (1992.3) (1482.8) (294.9) (293.8) (153.3) (4.0) (20.9) 1182.0 879.7 174.9 174.3 91.0 0.92 4.75 0.28 1-5 (1992.3) (1482.8) (294.9) (293.8) (153.3) (4.0) (20.9) (1.2) 951.5 1048.0 209.5 209.2 91.9 1.4 0 0 2-1 (1603.8) (1766.4) (353.1) (352.6) (154.9) (5.2) 951.5 1048.0 209.5 209.2 91.9 1.4 2.8 0 2-2 (1603.8) (1766.4) (353.1) (352.6) (154.9) (5.2) (10.4) 951.5 1048.0 209.5 209.2 91.9 1.4 5.6 0 2-3 (1603.8) (1766.4) (353.1) (352.6) (154.9) (5.2) (20.8) 946.1 1042.0 208.3 208.0 97.0 1.4 5.6 0 2-4 (1594.7) (1756.4) (351.1) (350.6) (163.5) (5.2) (20.8)

The field testing demonstrated the ease with which the ICAR rheometer could be used routinely on a jobsite. Pictures from the testing are shown in Figure 7.43. The concrete could be sampled directly from the discharge chute and then tested at the same location in less than 60 seconds. A mass-production version of the ICAR rheometer, which would likely include smaller components and streamlined electronics, would be even simpler to operate. The rapid availability of results means that the results can be used immediately to confirm that the concrete meets specifications or to adjust the concrete mixture as needed. For instance, if it is found that the viscosity of the concrete is too low, the rheometer software could suggest the needed dosages of admixtures to bring the concrete to the correct rheology. The ability to obtain

263 objective measurements quickly could allow chemical admixtures to be dispensed accurately and effectively on site.

Concrete Truck Sampling Concrete Testing Concrete

Figure 7.43: Field Testing with Rheometer

The results of the field testing are shown in Table 7.10, Figure 7.44, and Figure 7.45. In general, the results were consistent with expectations. The three additions of the high-range water-reducing admixture in Series I resulted in reductions in yield stress and plastic viscosity. The viscosity modifying admixture increased both yield stress and plastic viscosity. For Series II, the addition of high-range water-reducing admixture resulted in decreases in yield stress and plastic viscosity from the first mix to the second mix. In the final two mixtures, the setting of the concrete began to dominate the rheology of the concrete.

264

Table 7.10: Field Testing Results Rheometer Results Yield Viscosity Yield Plastic Yield Mix Slump R2 mse Value Value Stress Viscosity Torque (cm) (Nm) (Nm.s) (Pa) (Pa.s) (Nm) 1-1 7.0 4.38 2.01 0.995 1092.8 19.00 0.054 24.3 1-2 11.5 3.04 0.97 0.991 818.0 6.51 0.020 10.0 1-3 17.2 2.65 0.73 0.872 667.2 6.81 0.485 9.3 1-4 24.1 1.43 0.54 0.986 372.7 4.22 0.047 4.2 1-5 15.2 1.97 0.95 0.979 478.3 9.77 0.124 9.4 2-1 6.4 4.66 0.97 0.946 1264.4 6.07 0.253 13.7 2-2 12.1 2.73 0.66 0.989 760.7 3.49 0.010 8.8 2-3 6.4 2.34 0.79 0.969 595.6 6.84 0.183 12.3 2-4 3.8 4.11 0.76 0.954 1134.2 4.26 0.227 20.8

1.2

1 base /Mix i 0.8 Mix 1-1: Base #1 Mix 1-2: Add HRWR 0.6 Mix 1-3: Add HRWR Mix 1-4: Add HRWR 0.4 Mix 1-5: Add VMA

Rheological Parameter, Mix Parameter, Rheological 0.2

0 Yield Value Viscosity Yield Stress Plastic Yield Value Viscosity Torque

Figure 7.44: Results of Field Testing– Series I

265 1.6

1.4 base 1.2 /Mix i

1 Mix 2-1: Base #2

0.8 Mix 2-2: Add HRWR Mix 2-3: Add HRWR 0.6 Mix 2-4: Add Water

0.4 Rheological Parameter, Mix Parameter, Rheological 0.2

0 Yield Value Viscosity Yield Stress Plastic Yield Value Viscosity Torque

Figure 7.45: Results of Field Testing – Series II

7.10 WORKABILITY RANGE While it is desirable to measure the broadest range of concrete workability possible, it is not practical to measure the full range of workability. The testing of the first generation prototype of the ICAR rheometer, as presented in Chapters 6 and 7, indicated that the ICAR rheometer was capable of measuring concretes with slumps greater than approximately 50 to 75 mm (2 to 3 inches). The two main issues that must be addressed are the maximum torque required to rotate the impeller and the Deborah number associated with the deformation process. The maximum torque capacity of 50 Nm for the first generation prototype was typically achieved when slumps of less than 2 inches were tested. The torque capacity could always be increased by using a larger motor and gearbox; however, the characteristic time of the deformation process must also be considered. To this end, the Deborah number was assessed qualitatively. The fact

266 that concrete did not flow to fill the voids created behind the vane blades for slumps less than 2 to 3 inches indicates that such concretes are generally too stiff for testing in the rheometer. A typical example for a zero slump concrete is shown in Figure 7.46. In general, better results with less scatter were obtained from higher slump mixtures, indicating that the characteristic time of the deformation process better matched the characteristic time of higher slump materials for liquid-like behavior.

Figure 7.46: Zero Slump Concrete Tested with Vane Impeller (Concrete Filled to One Inch below Top of Vane)

In addition to the fact that low-slump concretes do not readily flow within the rheometer, other practical problems increase the difficulty of testing low- slump concrete. These problems are described as follows: • A given low-slump concrete will exhibit a different response depending on whether it is in a loose, uncompacted state; a consolidated, cohesive state; or somewhere in between.

267 • The maximum torque capacity of the rheometer is often exceeded in the initial torque spike prior to structural breakdown, even if the steady-state torque is below the maximum torque capacity. • The vane cannot be inserted downward into low-slump concrete as it can in more fluid mixes. Instead, concrete must be placed by hand around the vane. • Large lateral and axial loads can be transmitted to the vane shaft when testing low-slump mixtures. Whenever the vane strikes an aggregate, the supporting material around the aggregate is much stiffer than in more fluid mixes. The increased resistance of the aggregates results in larger forces imparted in the vane shaft. The slots in the container used to connect the rheometer frame to the container are designed to accommodate such forces. The rheometer frame can twist and can move laterally about 1/16-inch within in the container slot, reducing the maximum lateral force in the vane shaft. Once a zone of loose concrete is created in the sample, the vane can tend to “walk” due to the lack of a solid connection between the rheometer frame and the container. Although the rheometer is designed to prevent damage to the torque transducer or motor, the vane shaft can be bent when large lateral forces are generated. • The container is very difficult to empty of dry, cohesive concrete upon completion of a test.

7.11 CONCLUSIONS A series of tests on a wide range of concrete mixtures has shown that the ICAR rheometer is capable of successfully determining the rheology and

268 workability of fresh concrete. The following conclusions can be reached about the use of the ICAR rheometer: • The ICAR rheometer is able to detect systematic changes in workability successfully. • The ICAR rheometer is capable of measuring concrete mixtures with slumps greater than 50 to 75 mm (2 to 3 inches) up to self-consolidating concrete. • The operation of the ICAR rheometer is simple, fast, and well suited for use in the field. • The ICAR rheometer is most effective on high-fluidity mixes because these materials most closely represent homogeneous “fluids.” • The ICAR rheometer is a dynamic test and is, therefore, well suited for measuring high-microfines concrete and other highly thixotropic materials. The slump test, in contrast, is unsuitable for such concretes. • The results from the ICAR rheometer are strongly dependent on the shear history of sample, particularly for concrete mixtures with high yield stresses and high plastic viscosities. • Concrete rheology can be related to workability. The use of workability boxes is one effective way or relating rheology to workability. • Additional research is needed to determine the best way to compute yield stress and plastic viscosity. The use of either relative parameters based on a fit of torque versus rotation speed data or absolute parameters based on the Effective Annulus Method appears reasonable.

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CHAPTER 8: SUMMARY, CONCLUSIONS, AND RECOMMENDATIONS

8.1 SUMMARY In 1922, Duff Abrams wrote that the search for a method of measuring and controlling concrete flow properties “is giving engineers more concern today than probably any other one thing.” More than eight decades later, the concrete industry has yet to find a test to characterize workability adequately, although progress has been made. Current techniques for workability characterizations are inadequate for many concrete mixtures, such as those incorporating aggregate microfines, fiber reinforcement, ground granulated blast furnace slag, and new classes of chemical admixtures. The slump test, which is still used predominantly throughout the world, can provide incomplete, inaccurate, or misleading results. The purpose of the research described in this report was to identify an effective field test method for characterizing the workability of concrete in general and of high-microfines concrete in particular. A literature search conducted at the beginning of the research identified 61 existing test methods for concrete workability (Koehler and Fowler 2003). Based on the literature search and feedback from industry, government, and academia, criteria for a new workability test device were developed. Any new device should be portable, simple, rapid, and capable of measuring values equal to or related to fundamental flow properties. Therefore, it was decided to develop the ICAR rheometer, a low-cost, accurate, and portable rheometer for fresh concrete. The ICAR rheometer is a controlled-rate rheometer that is capable of measuring a flow curve or performing a stress growth test. It features a four- bladed vane, measuring 5 inches in diameter and 5 inches in height, that is

271 immersed into a concrete sample. The rheometer can be operated by hand or secured to a fixed frame above a standard container. Experimental testing was conducted to determine certain operating characteristics of the rheometer, including the optimum rotation speed for determining yield stress, the effects of container dimensions, and the magnitude and distribution of shear stress acting on the ends of the vane. The ICAR rheometer was next tested on a wide range of concrete mixtures ranging from a slump of 2 inches to self-consolidating concrete. The purpose of the testing was to verify the ability of the ICAR rheometer to detect changes in workability. For the self-consolidating concrete mixtures, the ICAR rheometer results were compared to the BTRHEOM rheometer. Finally, field testing was conducted to demonstrate the portability of the ICAR rheometer.

8.2 CONCLUSIONS Based on the literature review; feedback from government, industry, and academia; development of the ICAR rheometer; and experimental testing, the following conclusions can be reached:

8.2.1 Concrete Workability Characterization • Concrete workability is a broadly defined term. No single test method is capable of measuring all aspects of workability. Workability encompasses many interrelated properties, such as flowability, consistency, mobility, pumpability, plasticity, compactability, stability, cohesiveness, and finishability. These properties are often assessed by qualitative and subjective means. • The characterization of concrete workability is complicated by the fact that concrete is a complex material with time-dependent properties and

272 a wide range of particle sizes. Concrete is essentially a concentrated suspension of aggregates in cement paste. Cement paste is a concentrated suspension of cement grains in water. • A myriad of test methods has been developed for measuring workability. Most of these test methods are empirical; that is, they attempt to simulate a field condition and measure a value—such as a distance or time—that is related to some aspect of workability. In contrast, fundamental tests measure the scientific flow properties of concrete. A second distinction can be made between static and dynamic tests. In static tests, only the force of gravity acts on the sample. In contrast, dynamic tests add external energy to the concrete and can typically provide information on the flow properties after the yield stress has been exceeded. Static tests, such as the slump test, are inappropriate for highly thixotropic concrete mixtures. • The slump test is inadequate for a growing portion of mixtures. Despite being simple and inexpensive, the slump test only measures consistency and can provide misleading results. The slump test should not be used to determine workability but can be used as a quality control device to detect changes in mixture proportions, material characteristics, or mixing operations. • A viable, accurate, and relevant test method for concrete workability would foster an improved understanding of concrete workability, reduce concrete construction costs, improve hardened concrete quality, and promote the use of new and underutilized materials. • The prospect of characterizing concrete workability by measuring concrete rheological properties is promising. If properly designed, fluid rheometers are able to characterize the scientific flow properties of concrete. Fluid rheometers cannot be used for dry-consistency

273 concrete because the Deborah numbers for the associated deformation processes are too high, resulting in solid-like behavior. • The rheology and workability of concrete are influenced by nearly every aspect of the mixture proportions, material characteristics, and construction conditions. • The concept of a portable concrete rheometer fulfills the criteria established for a new workability test method. Such a device can quickly and accurately measure the relevant fundamental flow properties of concrete. By being portable and low-cost, such a device can enable laboratory-type measurements of rheological parameters to be made in the field on a routine basis.

8.2.2 Development of the ICAR Rheometer • Most conventional, commercially available fluid rheometers are not suitable for testing concrete due, in part, to the large maximum aggregate sizes in concrete. Although fluid rheometers have been developed specifically for concrete, such devices are too large and expensive for routine use and include artifacts that limit the accuracy of their test results. • The ICAR rheometer addresses many of the limitations of existing rheometers. It is a low-cost, fully portable rheometer capable of measuring nearly the full range of workability from a slump of about 2 inches to self-consolidating concrete. It can be used to measure a flow curve or perform a stress growth test. The ICAR rheometer meets the original criteria for a new workability test device. It is fully portable and can be easily transported on a jobsite. The test procedure is comprehensible and requires no advanced training. The test can be

274 performed quickly—a single flow curve test can be completed in less than 60 seconds. The rheometer can be used for research and development, mixture proportioning, and field testing. • Of the nine impellers evaluated, the vane was selected as the best impeller for concrete. Other impellers, if properly designed, may also be suitable. • Although the ICAR rheometer can be operated by hand, the best results are obtained when the ICAR rheometer is positioned above a standard container. The test results should not be influenced by container dimensions provided each dimension is larger than a certain minimum size.

8.2.3 Testing of ICAR Rheometer • The ICAR rheometer was able to detect changes in concrete rheology and workability successfully. The effects of making various systematic changes to mixture proportions on the rheological parameters measured by the ICAR rheometer were generally consistent with expectations. The ICAR rheometer provided information the slump test could not provide. • The rheometer is capable of measuring concretes with slumps greater than 50 to 75 mm (2 to 3 inches) up to self-consolidating concrete. The rheometer is most effective on highly fluid mixes because these materials most closely represent homogeneous “fluids.” The results from the ICAR rheometer are strongly dependent on the shear history of sample, particularly for concretes with high yield stresses and high plastic viscosities.

275 • The ICAR rheometer is effective for measuring concretes with high microfines contents and other thixotropic concrete mixtures because the test method adds energy to the concrete and, therefore, provides a dynamic measurement of workability. Consequently, the test is a significant improvement over the slump test, which is unable to measure high-microfines concrete properly. • The ICAR rheometer can be used successfully to measure self- consolidating concrete. As long as yield stress is kept sufficiently close to zero, plastic viscosity is the more important parameter in predicting how the concrete will flow. Segregation can distort test results; therefore, a separate test is needed for segregation. • The relative parameters of yield value and viscosity value can indicate slightly different trends in workability than the absolute parameters of yield stress and plastic viscosity. The test data were inconclusive on the best approach for calculating the correct flow curve parameters. Given the granular nature of concrete and the limitations on building a “true” rheometer with flow throughout the annulus, it is not possible to determine the stress and velocity profiles precisely in any concrete rheometer. • The yield stress measured from a stress growth test is a different physical property than the yield stress measured from the flow curve. The stress growth test is a static test that is influenced by thixotropy while the flow curve is based on a dynamic measurement. Both values are useful in predicting workability. • Concrete rheology, as measured with the ICAR rheometer, can be related to workability. Instead of using an assortment of highly subjective terms to describe concrete workability—as is still commonly done—it is desirable to measure just two rheological

276 parameters—yield stress and plastic viscosity—to provide a standardized, scientific description of concrete workability. The use of workability boxes is one effective way of relating rheology to workability.

8.3 RECOMMENDATIONS FOR FUTURE WORK As with any product, further development of the ICAR rheometer is always possible. Additional work is needed in the following areas: • Create a streamlined, mass-production version of the ICAR rheometer. The first generation prototype of the ICAR rheometer can be made smaller and more lightweight. The off-the-shelf components in the first generation prototype can be replaced with smaller custom components, which can be enclosed in a plastic case. The software can be operated from a handheld computer (PDA) or similar embedded electronics. • Develop a better understanding of the flow generated by the ICAR rheometer. Computer modeling or additional experimental work is needed to determine more precisely the distribution of stress and flow velocity within the rheometer. The use of relative parameters likely understates viscosity; however, more information is needed to improve the accuracy of the calculations of yield stress and plastic viscosity from the torque versus rotation speed data. Although the flow in the rheometer is complex, it may be possible to develop simple, reliable relationships that could be programmed into the computer to determine the Bingham parameters. • Calibrate the ICAR rheometer with standard fluids of known flow properties. The use of standard fluids of known flow properties can be

277 used to verify the accuracy of the ICAR rheometer further. Currently, no standard reference material exists for calibrating concrete rheometers. Newtonian fluids with viscosities similar to concrete have been used in the past to compare the results from different concrete rheometers. While a standard reference material for concrete rheometers would be preferable, a high-viscosity Newtonian fluid would be an acceptable first step. • Conduct additional experimental testing to the supplement the findings of Chapter 6. The testing in Chapter 6 was based on four concrete mixtures and a limited number of data points. The operating characteristics considered in Chapter 6 should be evaluated on a greater number of concrete mixtures covering a wider range of workability. Further testing may demonstrate that a smaller container size is acceptable. • Relate rheology to workability. In order for the results of the ICAR rheometer to be useful in the field, the rheological parameters measured by the ICAR rheometer must be related to practical workability requirements. A better understanding of what constitutes acceptable workability is needed. Experienced field workers should be surveyed. A better understanding of the connection between rheology and workability will be developed over time as a range of parties in the concrete industry use the ICAR rheometer in a wide array of applications. • Develop guidelines and standards for the use of rheology in concrete. Knowledge of rheology within the concrete industry is limited. The lack of experience with rheology is due, in part, to the fact that existing concrete rheometers are too expensive for routine use. The availability of the ICAR rheometer should make the

278 characterization of rheology economically viable. Guidelines and standards are needed to demonstrate the use of rheology and to ensure that rheology is correctly applied to concrete. • Extend the use of the ICAR rheometer to materials other than concrete. Although the ICAR rheometer was developed for concrete, it should be suitable for other cementitious and non-cementitious coarse-grained materials. If necessary, the vane can be replaced with a different-sized vane or with an impeller of a different geometry. Mortar could be tested by using a smaller vane and container.

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289

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APPENDIX A: LABORATORY TEST DATA FOR CONVENTIONAL CONCRETE MIXTURES

291

Average 3.5 3.0 2.8 3.8 2.3 4.3 4.2 4.3 4.2 2.5

Overall 3 3 3 4 2 4 4 4 4 2

Finishability 4 3 3 4 2 4 4 5 5 3

Bleeding 3 2 2 3 2 5 5 5 4 2

Segregation 4 3 2 4 1 5 4 4 4 2

Richness 3 3 3 4 3 4 4 4 4 3

Flowability 4 4 4 4 4 4 4 4 4 3

0.65 0.19 0.37 0.20 0.02 0.11 0.26 0.17 0.12 0.29 mse µ 6.6 8.0 30.8 16.1 16.0 33.1 23.9 31.1 23.1 14.9 (Pa.s)

0 τ (Pa) 195.0 631.4 448.7 360.3 294.4 155.8 1781.8 1166.8 1439.7 1171.2

R^2 0.806 0.956 0.874 0.965 0.986 0.981 0.932 0.948 0.966 0.929 Test 2 V 2.53 1.03 1.19 1.73 1.10 2.38 1.61 1.94 1.52 0.84 (Nm.s) Y 7.61 4.32 5.36 4.68 0.92 2.94 2.16 1.83 1.45 0.85 (Nm) 4.8 3.9 31.4 12.0 15.1 11.4 22.5 11.3 11.7 11.5 (Nm) Yld T

0.3 0.37 0.21 1.34 0.45 0.12 0.07 0.02 0.09 0.02 mse µ 15.8 15.6 17.8 10.2 20.2 25.7 26.3 32.4 24.5 24.0 (Pa.s)

0 τ (Pa) 50.6 152.4 753.3 530.1 358.1 269.7 2071.3 1201.0 1118.5 1346.5

Table A.1: Fly Ash Test Data R^2 Test 1 0.881 0.961 0.574 0.865 0.972 0.983 0.996 0.985 0.902 0.995 V 2.19 1.70 1.08 1.28 1.15 2.19 2.03 2.02 1.49 1.15 (Nm.s) Y 7.81 4.78 4.99 5.17 0.84 3.24 2.38 1.82 1.39 0.35 (Nm) 4.3 9.2 7.9 42.1 14.2 19.5 13.6 32.9 17.8 12.4 (Nm) Yld T 8 3 7 2.5 5.5 6.5 5.5 5.5 7.5 6.25 (inch) Slump

w/cm 0.460 0.460 0.460 0.460 0.460 0.432 0.432 0.432 0.432 0.432 0 0 10 20 35 55 10 20 35 55 Fly (%) Ash FA NS NS NS NS NS NS NS NS NS NS LS LS LS LS LS CA RG RG RG RG RG Mix # 01-01 01-02 01-03 01-04 01-05 01-06 01-07 01-08 01-09 01-10

292 Average 4.5 4.3 4.2 3.8 3.0 4.0 3.0 3.0

Overall 4 4 4 4 3 4 3 3

Finishability 4 4 4 4 3 4 3 3

Bleeding 5 5 4 4 3 5 3 3

Segregation 5 5 4 3 2 4 2 2

Richness 5 5 4 3 3 4 3 3

Flowability 4 3 5 5 4 3 4 4

0.24 0.11 0.16 0.07 0.22 0.09 0.12 0.17 mse µ 14.1 24.7 12.4 14.8 18.1 17.2 21.1 19.7 (Pa.s)

0 τ (Pa) 502.7 996.1 658.3 515.0 932.3 380.4 405.5 1251.0

R^2 0.94 0.98 0.97 0.95 0.93 0.97 0.98 0.95 Test 2 V 1.17 2.47 1.41 1.58 1.49 1.84 1.48 1.46 (Nm.s) Y 2.18 5.10 3.93 2.60 2.24 3.72 1.80 1.86 (Nm) 9.6 9.9 29.0 13.0 14.7 17.8 13.7 12.7 (Nm) Yld T

0.32 0.41 0.17 0.26 0.13 0.05 0.12 0.21 mse µ 30.0 31.6 20.7 19.3 30.4 23.4 25.6 16.7 (Pa.s)

0 τ (Pa) 383.3 919.5 635.0 456.7 939.9 425.7 526.2 1315.2

R^2 0.92 0.88 0.96 0.91 0.97 0.99 0.98 0.96 Test 1 V 1.76 2.58 1.91 1.58 2.04 2.20 1.76 1.33 (Nm.s) Y 2.02 5.72 3.85 2.77 2.21 3.90 2.03 2.33 (Nm) 14.1 32.3 18.7 16.0 28.0 30.3 27.5 31.8 (Nm) Yld T Table A.2: Ground Granulated Blast Furnace Slag Test Data 5 4.5 2.5 4.5 4.5 3.5 5.5 5.75 (inch) Slump

w/cm 0.483 0.483 0.483 0.483 0.465 0.465 0.465 0.465 0 0 20 35 50 20 35 50 (%) Slag FA NS NS NS NS NS NS NS NS LS LS LS LS CA RG RG RG RG Mix # 02-01 02-05 02-06 02-07 02-08 02-12 02-13 02-14

293 Average 4.5 4.3 4.8 3.2 3.0 3.0 4.0 4.2

Overall 4 4 5 2 3 3 4 4

Finishability 4 4 5 3 3 3 4 4

Bleeding 5 5 5 5 3 3 4 5

Segregation 5 4 5 4 2 2 4 4

Richness 5 5 5 3 3 3 4 4

Flowability 4 4 4 2 4 4 4 4

-- 0.5 0.24 0.17 0.22 0.03 0.13 0.13 mse µ -- 14.1 15.5 14.6 18.1 24.7 17.6 15.9 (Pa.s)

0 -- τ (Pa) 502.7 735.0 754.0 515.0 403.9 475.1 621.3

-- R^2 0.94 0.85 0.96 0.93 0.99 0.97 0.97 Test 2 -- V 1.17 1.27 1.43 1.49 1.80 1.41 1.43 (Nm.s) -- Y 2.18 3.17 3.09 2.24 1.86 2.10 2.63 (Nm) -- 9.6 12.6 14.6 14.7 16.2 12.0 15.3 (Nm) Yld T

0.1 0.32 0.41 0.93 0.13 0.03 0.02 0.07 mse µ 30.0 23.1 22.7 51.1 30.4 29.5 20.6 26.1 (Pa.s)

0 τ (Pa) 383.3 607.4 687.7 456.7 496.1 520.5 448.0 1145.5

1 0.9 R^2 0.92 0.98 0.72 0.97 0.99 0.99 Test 1 Table A.3: Silica Fume Test Data Table A.3: Silica Fume V 1.76 1.84 1.63 2.72 2.04 2.14 1.73 1.87 (Nm.s) Y 2.02 2.69 3.16 5.91 2.21 2.30 2.24 2.09 (Nm) 14.1 21.3 23.6 40.7 28.0 32.4 27.6 34.0 (Nm) Yld T 4 2 4.5 3.5 3.5 3.5 4.5 3.75 (inch) Slump

w/cm 0.483 0.483 0.483 0.483 0.465 0.465 0.465 0.465 0 3 5 8 0 3 5 8 SF (%) FA NS NS NS NS NS NS NS NS LS LS LS LS CA RG RG RG RG Mix # 02-01 02-02 02-03 02-04 02-08 02-09 02-10 02-11

294 Average 3.6 4.2 3.5 4.2 3.3 3.8 3.8 3.0

Overall 4 4 3 4 3 4 4 3

Finishability 4 4 3 5 3 4 4 3

Bleeding 4 4 3 4 3 3 3 3

Segregation 3 4 3 5 4 4 4 2

Richness 3 4 4 4 4 4 4 3

Flowability 4 5 5 3 3 4 4 4

0.1 0.22 0.08 0.32 0.11 0.22 0.16 0.06 mse µ 19.4 23.4 15.9 33.5 45.6 17.7 24.0 11.2 (Pa.s)

0 τ (Pa) 378.6 541.0 122.4 695.6 764.0 948.5 312.6 356.0

R^2 0.94 0.99 0.92 0.98 0.94 0.96 0.98 0.99 Test 2 V 1.35 1.81 0.82 2.48 3.08 1.77 1.56 0.97 (Nm.s) Y 1.79 2.43 0.73 3.19 3.68 3.87 1.55 1.53 (Nm) 4.9 8.6 11.4 11.0 15.7 20.8 18.1 13.3 (Nm) Yld T

0.12 0.12 0.55 0.23 0.43 0.27 0.02 0.16 mse µ 29.1 29.9 18.5 45.2 45.0 22.4 25.2 23.6 (Pa.s)

0 τ (Pa) 351.0 508.3 113.2 672.5 831.1 975.7 340.4 289.9

1 R^2 0.96 0.98 0.81 0.94 0.87 0.95 0.97 Test 1 V 1.93 2.07 0.94 2.96 2.84 1.94 1.74 1.48 (Nm.s) Table A.4: Water-to-Cement Ratio Test Data Table A.4: Water-to-Cement Y 1.70 2.42 0.69 3.30 4.11 4.17 1.62 1.47 (Nm) 7.7 18.4 20.6 27.6 50.3 35.2 20.7 13.6 (Nm) Yld T 5 7 3 3 5.5 4.5 5.5 4.25 (inch) Slump

w/cm 0.500 0.472 0.539 0.454 0.430 0.457 0.485 0.525 FA NS NS NS NS NS NS NS NS LS LS LS LS CA RG RG RG RG Mix # 04-01 04-02 04-03 04-04 04-08 04-09 04-10 04-11

295 Average 3.5 4.0 3.8 3.5 4.0 4.0 3.2 4.5 3.0 2.8 3.2 3.5

Overall 3 4 4 3 4 4 3 5 2 2 3 3

Finishability 3 4 4 3 4 4 4 5 3 3 3 4

Bleeding 4 4 4 3 4 4 2 4 3 3 3 4

Segregation 4 4 3 3 4 4 5 5 3 2 3 3

Richness 4 4 3 4 4 4 2 4 3 2 3 3

Flowability 3 4 5 5 4 4 3 4 4 5 4 4

0.2 0.07 0.17 0.06 0.01 0.19 0.31 0.11 0.12 0.12 0.06 0.46 mse µ 22.3 21.1 20.0 19.7 29.1 26.7 33.7 31.0 29.6 27.8 25.7 18.0 (Pa.s)

0 τ (Pa) 914.1 579.1 367.7 123.2 438.4 402.6 497.6 383.6 238.0 143.1 384.2 373.4

R^2 0.95 0.96 0.99 0.99 0.95 0.92 0.95 0.98 0.97 0.98 0.99 0.88 Test 2 V 2.28 1.66 1.48 1.17 1.96 1.65 2.21 1.98 1.72 1.45 1.77 1.14 (Nm.s) Y 3.67 2.58 1.68 0.65 2.12 2.04 2.44 1.93 1.29 0.87 1.83 1.83 (Nm) 8.2 3.9 9.8 7.9 9.0 5.5 22.9 13.1 14.7 22.9 10.3 10.3 (Nm) Yld T

0.2 0.31 0.02 0.08 0.24 0.21 0.13 0.14 0.05 0.18 0.26 0.32 mse µ 30.2 21.4 38.2 20.1 22.8 33.8 40.8 33.2 24.8 35.3 24.8 21.0 (Pa.s)

0 τ (Pa) 851.0 644.4 297.4 160.7 548.0 395.5 639.2 428.2 328.2 132.9 404.4 386.7

1 R^2 0.96 0.88 0.98 0.95 0.94 0.97 0.97 0.97 0.95 0.94 0.93 Test 1 V 2.34 1.72 2.29 1.21 1.65 2.08 2.78 2.10 1.74 1.72 1.59 1.35 (Nm.s) Y 3.81 2.83 1.56 0.84 2.54 2.03 3.07 2.15 1.53 0.88 2.01 1.90 (Nm) 5.2 26.0 18.8 15.6 19.1 15.7 39.6 22.0 16.7 10.5 27.5 20.7 (Nm) Yld T Table A.5: Water-Reducing Admixtures Test Data Table A.5: Water-Reducing 6 4 5 8 11 4.5 6.5 4.5 2.5 5.5 2.75 3.75 (inch) Slump

w/cm 0.459 0.459 0.459 0.459 0.459 0.459 0.413 0.413 0.413 0.413 0.413 0.413

0 3 6 9 0 3 5 6 (oz cwt) Dose HRWR

0 3 6 0 3 6 (oz cwt) WRA Dose FA NS NS NS NS NS NS NS NS NS NS NS NS LS LS LS LS LS LS CA RG RG RG RG RG RG Mix # 06-01 06-02 06-03 06-04 06-05 06-06 06-07 06-08 06-09 06-10 06-11 06-12

296 Average 4.2 4.0 4.3 4.8 3.8 3.0 4.0 4.3

Overall 4 4 4 4 4 3 4 4

Finishability 4 4 5 5 4 3 4 4

Bleeding 4 3 4 5 3 3 4 4

Segregation 4 4 4 5 4 2 4 4

Richness 4 4 4 5 4 3 4 5

Flowability 5 5 5 5 4 4 4 5

1.1 0.08 1.08 0.07 0.16 0.05 0.11 0.14 mse µ 7.6 23.4 19.0 16.8 17.7 20.5 14.8 14.2 (Pa.s)

0 τ (Pa) 541.0 653.6 816.8 916.0 948.5 611.6 448.6 530.3

R^2 0.99 0.65 0.66 0.98 0.96 0.99 0.96 0.95 Test 2 V 1.81 1.17 0.65 1.78 1.77 1.76 1.29 1.29 (Nm.s) Y 2.43 3.09 3.34 3.67 3.87 2.62 1.91 2.23 (Nm) 9.4 8.5 11.0 12.4 11.0 10.2 18.1 11.1 (Nm) Yld T

0.1 0.12 0.15 0.27 0.02 0.27 0.22 0.06 mse µ 1.8 29.9 24.5 15.4 17.9 22.4 22.4 11.6 (Pa.s)

0 τ (Pa) 508.3 622.1 788.0 963.4 975.7 663.9 836.3 632.0

R^2 0.98 0.95 0.94 0.99 0.95 0.93 0.75 0.98 Test 1 V 2.07 1.96 1.40 1.92 1.94 1.80 1.04 1.21 (Nm.s) Table A.6: Air-Entraining Agent Test Data Table A.6: Air-Entraining Y 2.42 2.74 3.30 3.84 4.17 2.92 2.46 2.55 (Nm) 20.6 20.2 15.3 16.8 35.2 23.0 18.0 12.4 (Nm) Yld T 5 4 4 5 6 5 4.5 4.3 (inch) Slump

w/cm 0.472 0.472 0.472 0.472 0.457 0.457 0.457 0.457 Air 1.2 3.6 4.5 7.5 1.3 3.7 6.0 6.9 (%) FA NS NS NS NS NS NS NS NS LS LS LS LS CA RG RG RG RG 04-02 04-05 04-06 04-07 04-09 04-12 04-13 04-14 Mix #

297 Average 3.7 4.2 4.0 4.3 4.0 2.2 2.5 2.2 3.3 2.7

Overall 4 4 4 4 3 1 3 2 3 2

Finishability 3 4 4 4 4 2 2 2 3 3

Bleeding 3 4 4 4 4 2 2 2 3 2

Segregation 3 4 4 5 5 2 2 2 3 4

Richness 4 4 4 5 5 2 3 2 3 3

Flowability 5 5 4 4 3 4 3 3 5 2

0.09 0.15 0.27 0.12 0.34 0.46 0.16 0.07 0.07 0.25 mse µ 15.8 14.5 18.3 14.5 19.4 18.7 13.6 11.8 17.3 14.9 (Pa.s)

0 τ (Pa) 255.8 336.1 414.7 597.6 805.1 175.2 269.4 390.6 455.7 608.4

R^2 0.99 0.96 0.93 0.98 0.91 0.87 0.95 0.98 0.99 0.92 Test 2 V 1.09 1.11 1.29 1.32 1.65 0.97 0.99 1.04 1.41 1.32 (Nm.s) Y 1.22 1.51 1.94 2.52 3.45 1.01 1.25 1.65 2.00 2.58 (Nm) 5.8 7.4 9.6 6.0 7.2 11.9 16.5 11.2 13.5 20.2 (Nm) Yld T

0.34 0.07 0.09 0.29 0.19 0.02 0.27 0.29 0.27 0.39 mse µ 12.5 15.1 22.0 28.3 23.3 18.0 17.7 15.1 14.5 20.1 (Pa.s)

0 τ (Pa) 353.6 420.9 486.5 491.8 851.3 212.8 265.0 408.0 562.0 650.7

0.9 R^2 0.91 0.97 0.98 0.93 0.92 0.98 0.91 0.93 0.86 Test 1 V 0.92 1.31 1.68 1.83 2.14 1.24 1.14 1.20 1.25 1.63 (Nm.s) Y 1.62 1.79 2.20 2.42 3.55 1.00 1.31 1.80 2.41 2.84 (Nm) 8.8 11.9 15.3 15.1 24.6 12.9 12.2 20.8 25.7 30.5 (Nm) Yld T Table A.7: Natural Sand/Manufactured Sand Blends Test Data 7 7 5 5 4 7 5 4 5.5 4.5 (inch) Slump

w/cm 0.526 0.526 0.526 0.526 0.526 0.492 0.492 0.492 0.492 0.492 0 0 25 50 75 25 50 75 MS (%) 100 100 FA N/M N/M N/M N/M N/M N/M N/M N/M N/M N/M LS LS LS LS LS CA RG RG RG RG RG Mix # 05-01 05-02 05-03 05-04 05-05 05-06 05-07 05-08 05-09 05-10

298 Average 2.7 3.0 4.2 4.5 4.5 3.0 1.7 2.5 3.5 4.5 4.5 2.3

Overall 2 3 4 4 4 2 1 2 3 4 4 1

Finishability 3 3 5 5 5 4 1 3 3 5 5 2

Bleeding 2 2 4 4 4 5 3 2 4 4 5 4

Segregation 2 3 3 5 5 3 1 2 3 5 5 3

Richness 3 3 4 5 5 3 2 2 3 5 5 3

Flowability 4 4 5 4 4 1 2 4 5 4 3 1

------0.7 0.7 0.2 0.12 0.15 0.21 0.43 0.01 0.05 mse µ ------27.8 17.8 17.4 13.2 14.1 31.4 21.6 17.5 31.9 (Pa.s)

0 ------τ (Pa) 434.3 445.1 741.9 313.4 929.5 682.8 742.1 1180.2 1555.7

1 ------0.8 R^2 0.78 0.98 0.95 0.95 0.87 0.96 0.99 Test 2 ------V 1.59 1.38 1.25 1.64 1.81 1.68 1.92 1.67 2.53 (Nm.s) ------Y 2.27 1.99 3.33 4.55 5.96 1.77 3.93 2.82 3.29 (Nm) ------9.3 27.1 13.6 14.5 27.4 30.0 17.3 12.4 22.3 (Nm) Yld T

------0.01 0.43 0.46 0.11 0.62 0.13 0.28 0.38 0.04 mse µ ------8.9 39.0 28.3 23.9 22.7 45.7 25.6 20.5 36.2 (Pa.s)

0 ------τ (Pa) 422.7 465.9 682.7 813.9 265.1 925.2 614.1 863.6 1799.3

------R^2 0.99 0.89 0.89 0.96 0.79 0.97 0.94 0.89 0.99 Test 1 V ------2.60 1.73 1.65 2.06 1.30 2.44 2.07 1.52 2.97 (Nm.s) Table A.8: Sand-to-Aggregate Ratio Test Data ------Y 2.04 2.36 3.19 3.42 6.69 1.56 4.06 2.79 3.76 (Nm) ------26.4 22.1 22.4 19.7 26.2 20.2 25.8 19.7 25.4 (Nm) Yld T 4 4 4 2 3 3.5 1.5 0.5 3.5 0.5 2.75 3.25 (inch) Slump

w/cm 0.481 0.481 0.481 0.481 0.481 0.481 0.440 0.440 0.440 0.440 0.440 0.440 0.3 0.4 0.4 0.5 0.5 0.6 0.3 0.4 0.4 0.5 0.5 0.6 S/A FA NS NS NS NS NS NS NS NS NS NS NS NS LS LS LS LS LS LS CA RG RG RG RG RG RG Mix # 03-01 03-02 03-03 03-04 03-05 03-06 03-07 03-08 03-09 03-10 03-11 03-12

299

Average 4.0 3.5 3.3 4.0 3.7 2.0 3.0 4.0 4.0 3.8 3.7

Overall 4 3 3 3 3 2 3 4 4 4 3

Finishability 4 4 4 4 4 2 3 4 4 4 4

Bleeding 4 3 2 5 4 2 2 3 3 3 3

Segregation 4 4 5 5 4 3 3 4 4 4 5

Richness 3 4 4 5 4 2 3 4 4 4 5

Flowability 5 3 2 2 3 1 4 5 5 4 2

-- 0.4 0.06 0.01 0.28 0.16 0.04 0.42 0.31 0.14 0.08 mse µ -- 24.6 22.7 28.2 26.7 31.4 55.7 10.5 11.1 17.6 11.1 (Pa.s)

0 -- τ (Pa) 335.1 527.6 828.9 956.0 602.0 104.8 304.5 480.6 550.1 1205.9

1 -- R^2 0.99 0.99 0.94 0.97 0.88 0.88 0.92 0.96 0.98 Test 2 -- V 1.65 1.88 2.14 2.27 2.35 2.22 0.80 0.99 1.49 1.50 (Nm.s) -- Y 1.62 2.28 3.74 4.12 2.74 1.00 1.37 2.03 2.37 4.58 (Nm) -- 6.7 8.7 16.5 14.9 34.7 25.4 14.9 34.1 11.8 16.8 (Nm) Yld T

-- 0.3 0.25 0.16 0.17 0.12 0.14 0.11 0.05 0.03 0.03 mse µ -- 29.7 33.5 22.9 45.2 26.1 34.8 15.5 11.4 28.8 21.2 (Pa.s)

0 -- τ (Pa) 302.7 437.3 979.6 892.2 687.7 237.0 250.5 567.2 461.9 1111.1

1 -- R^2 0.93 0.94 0.97 0.98 0.97 0.97 0.99 0.99 0.92 Test 1 V -- 1.72 2.24 2.06 3.27 2.05 1.98 1.10 1.18 2.08 1.99 (Nm.s) Table A.9: Aggregate Microfines Test Data -- Y 1.62 2.11 4.14 4.14 3.07 1.30 1.18 2.29 2.14 4.62 (Nm) -- 8.8 23.9 34.5 53.4 37.7 27.0 40.6 18.8 18.7 33.3 (Nm) Yld T 5 3 3 3 4 8 5 3.5 3.5 5.5 3.5 (inch) Slump

w/cm 0.517 0.517 0.517 0.517 0.517 0.517 0.517 0.603 0.603 0.603 0.603 5 10 15 20 25 15 15 16 10 15 20 (%) MF FA MS MS MS MS MS MS MS GR GR GR GR LS LS LS LS LS LS LS LS LS LS LS CA Mix # 08-01 08-02 08-03 08-04 08-05 08-06 08-07 08-08 08-09 08-10 08-11

300

Average 3.5 3.3 1.8 4.0 2.2 2.8

Overall 4 3 2 4 2 2

Finishability 4 3 2 4 2 3

Bleeding 2 4 2 4 2 3

Segregation 4 4 1 4 2 3

Richness 3 4 1 4 1 2

Flowability 4 2 3 4 4 4

-- 0.2 0.08 0.09 0.11 0.11 mse µ -- 26.4 22.5 12.7 28.9 36.6 (Pa.s)

0 -- τ (Pa) 320.3 249.3 399.9 216.6 142.7

-- R^2 0.96 0.99 0.97 0.96 0.97 Test 2 -- V 1.61 1.42 1.12 1.70 1.83 (Nm.s) -- Y 1.65 1.26 1.69 1.16 0.91 (Nm) -- 7.8 8.5 10.3 15.8 19.2 (Nm) Yld T

-- 0.3 0.11 0.22 0.05 0.15 mse µ -- 26.0 21.1 20.4 30.1 24.0 (Pa.s)

0 -- τ (Pa) 377.3 262.8 355.4 213.5 235.7

-- R^2 0.97 0.92 0.93 0.98 0.95 Test 1 V -- Table A.10: Slag Aggregate Test Data Table A.10: Slag Aggregate Test Data 1.78 1.38 1.31 1.80 1.48 (Nm.s) -- Y 1.80 1.29 1.76 1.12 1.21 (Nm) -- 23.7 26.7 11.7 33.2 13.7 (Nm) Yld T 5 6 4 1.5 5.5 4.5 (inch) Slump

w/cm 0.457 0.457 0.525 0.525 0.525 0.525

0.4 0.5 (%) S/A 0.35 FA NS NS NS NS NS NS LS CA SGA SGA SGA SGA SGA SGA Mix # 07-01 07-02 07-03 07-04 07-05 07-06

301

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APPENDIX B: LABORATORY TEST DATA FOR SELF- CONSOLIDATING CONCRETE MIXTURES

303

µ 93 95 66 72 29 92 78 50 88 44 59 36 12 85 42 108 105 113 215 162 102 124 (Pa.s)

0 τ -1 11 69 -16 -73 -62 -77 313 293 202 311 170 740 273 360 153 (Pa) -130 -105 -203 -195 -149 1711

1 R^2 0.996 0.993 0.992 0.985 0.997 0.986 0.934 0.991 0.976 0.998 0.957 0.993 0.998 0.999 0.999 0.999 0.997 0.976 0.999 0.999 0.995 BTRHEOM V 0.66 1.92 2.113 2.453 2.166 1.489 1.638 2.393 2.083 1.783 1.133 2.577 1.988 0.998 4.882 3.683 2.324 1.344 0.829 0.264 2.818 0.958 (Nm.s) Y 0.25 (Nm) -0.38 1.296 0.038 6.1645 1.1262 1.0568 0.7262 1.1198 0.6116 2.6649 0.9831 0.5527 -0.4695 -0.0583 -0.2639 -0.0047 -0.2226 -0.7302 -0.2769 -0.7018 -0.5366

-- 0.10 0.02 0.02 0.05 0.01 0.57 0.33 0.16 0.26 0.02 0.03 0.15 0.09 0.02 0.02 0.03 0.47 0.25 0.03 0.03 0.01 mse µ -- 27.8 32.8 31.8 29.5 17.5 12.5 15.2 14.1 15.5 32.2 30.6 20.0 79.8 56.0 21.1 15.7 12.8 40.1 33.8 28.7 104.4 (Pa.s)

0 -- τ 5.8 0.0 1.8 0.1 0.1 (Pa) 33.6 27.3 56.9 36.0 44.2 15.4 13.8 82.5 17.8 38.5 26.5 18.7 641.5 185.8 141.3 317.5

-- R^2 0.990 0.994 0.999 0.990 0.995 0.857 0.942 0.933 0.933 0.988 0.992 0.964 0.990 0.992 0.995 0.977 0.891 0.938 0.994 0.993 0.995 ICAR Rheometer V -- 2.27 1.95 1.78 1.35 0.85 0.47 1.10 0.83 0.75 1.57 1.36 0.86 4.95 3.73 2.51 1.10 0.65 0.54 1.81 1.51 1.32 (Nm.s) crete Test Data (Section 1 of 3) crete Test -- Y 2.78 0.95 0.78 0.25 0.17 0.09 1.46 0.27 0.24 0.27 0.13 0.14 0.56 0.15 0.20 0.24 0.03 0.00 0.13 (Nm) -0.21 -0.01

2 2 0 1 1 3 0 0 0 0 0 3 2 2 2 VSI

50 5 2 1 1 T 2.5 4.4 1.4 6.1 4.5 6.5 (Sec)

16 17 24 20 29 36 25 16 27 32 30 27.5 12.5 20.5 24.5 22.5 23.5 25.5 26.5 27.5 31.5 Flow Slump (Inches)

(oz/ 1.57 4.14 cwt) VMA Dose 5 8 9 6 6 8 4 6 7 6 7 9 7 8 4 5 6 5 6 7 11 10 (oz/ cwt) Dose HRWR

0:22 0:37 0:54 1:12 0:13 0:33 0:56 0:17 0:48 1:06 0:13 0:34 0:58 0:14 0:35 0:53 0:15 0:38 1:00 Time Elapsed (Hr:Min) Table B.1: Self-Consolidating Con

Change w/cm=0.43 w/cm=0.40 w/cm=0.35 20% fly ash Reduce Cement Increase Cement Coarser Gradation # 1 2 3 4 1 2 3 1 2 3 1 2 3 1 2 3 1 2 3 1 2 3 Test

# 1 2 3 4 5 6 7 Mix

304

µ 74 43 98 74 61 104 131 207 161 (Pa.s)

τ 88 138 304 472 462 (Pa) -296 -197 -196 1980

R^2 0.98 0.998 0.999 0.999 0.999 0.995 0.996 0.993 0.997 BTRHEOM

V 1.678 0.979 2.235 2.361 1.674 1.388 2.976 4.705 3.664 (Nm.s)

Y (Nm) 0.3179 0.4958 1.0943 1.6999 1.6656 7.1319 -1.0655 -0.7115 -0.7045

0.07 0.01 0.02 0.02 0.05 0.34 0.04 0.11 0.06 0.02 0.01 0.02 0.02 0.01 0.01 0.01 0.01 0.01 0.01 0.01 0.01 0.02 0.02 mse µ 9.8 6.6 8.2 9.0 26.3 18.5 36.8 31.3 18.3 84.5 71.9 68.6 63.3 65.0 63.9 60.9 62.1 59.8 64.4 58.5 59.6 165.1 112.7 (Pa.s)

0 τ 3.6 0.0 0.0 (Pa) 49.2 99.7 15.7 12.8 16.0 10.6 22.2 26.5 25.9 33.7 27.2 41.2 36.4 182.7 245.7 397.6 416.2 403.3 384.7 471.9

R^2 0.987 0.995 0.995 0.996 0.995 0.914 0.986 0.984 0.958 0.998 0.996 0.997 0.994 0.998 0.998 0.997 0.996 0.997 0.999 0.998 0.998 0.991 0.991 ICAR Rheometer V 1.58 0.97 1.64 1.65 1.30 0.90 0.82 0.87 1.01 8.22 5.09 3.94 3.20 3.06 2.84 2.95 2.91 2.73 2.79 2.68 2.90 2.77 2.84 (Nm.s) crete Test Data (Section 2 of 3) crete Test Y 0.94 0.28 0.03 0.58 1.14 1.65 1.59 1.60 1.48 3.05 0.14 0.11 0.13 0.06 0.15 0.22 0.21 0.28 0.22 0.23 0.18 (Nm) -0.03 -0.11

2 0 0 0 VSI

50 T 3.4 6.3 4.8 (Sec)

15 18 14 12 28 28 30.5 25.5 Flow Slump (Inches)

10 25 40 (oz/ cwt) VMA Dose

3 5 8 8 8 8 5 8 9.5 (oz/ cwt) Dose HRWR 0:14 0:30 0:23 0:42 0:59 1:18 1:20 1:21 1:22 0:15 0:37 1:09 1:12 1:14 1:15 1:16 1:19 1:20 1:22 1:24 1:28 1:29 1:30 Time Elapsed (Hr:Min) Table B.2: Self-Consolidating Con

VMA Change 40% fly ash 20% fly ash, w/cm=0.32 20% fly w/cm=0.32 ash, # 1 2 1 2 3 4 1 2 3 4_2 4_3 4_4 3_2 3_3 3_4 3_5 3_6 3_7 3_8 3_9 Test 3_10 3_11 3_12

# 8 9 10 Mix

305

µ 94 45 96 21 69 143 134 109 132 (Pa.s)

τ -37 -68 102 619 (Pa) -166 -239 -204 -254 -296

1 1 R^2 0.995 0.997 0.983 0.998 0.999 0.999 0.931 BTRHEOM

V 3.252 3.054 2.142 1.012 2.483 2.989 2.178 0.487 1.564 (Nm.s)

Y (Nm) 0.367 -0.861 -1.066 -0.246 2.2315 -0.1333 -0.5967 -0.7351 -0.9149

0.02 0.03 0.01 0.03 0.02 0.03 0.01 0.10 0.05 0.05 0.06 0.10 0.12 0.03 0.02 mse µ 82.1 46.3 34.4 31.6 57.8 39.9 30.3 24.0 23.9 24.6 25.1 27.0 25.5 23.2 34.0 (Pa.s)

0 τ 3.5 (Pa) 84.0 40.6 27.5 22.5 11.5 49.6 57.6 61.4 65.1 72.3 67.1 79.2 148.4 169.6

R^2 0.998 0.997 0.998 0.996 0.999 0.992 0.996 0.947 0.990 0.988 0.981 0.967 0.959 0.980 0.994 ICAR Rheometer ) V 3.74 2.42 1.59 1.42 2.98 1.79 1.37 1.32 1.20 1.25 1.29 1.41 1.34 1.29 1.52 (Nm.s crete Test Data (Section 3 of 3) crete Test Y 0.64 0.88 0.30 0.23 1.03 0.18 0.09 0.21 0.36 0.38 0.39 0.41 0.38 0.41 0.03 (Nm)

0 0 0 2 3 VSI

50 5 T 2.5 4.8 4.5 2.5 (Sec)

16 23 12 22 24 29 31 28.5 Flow Slump (Inches)

(oz/ cwt) VMA Dose

8 6 11 10 7.5 9.5 (oz/ 11.5 12.5 10.5 cwt) Dose HRWR 0:14 0:36 1:00 1:15 0:17 0:37 0:58 1:17 1:19 1:21 1:22 1:26 1:27 1:28 0:20 Time Elapsed (Hr:Min) Table B.3: Self-Consolidating Con

Change w/cm=0.37 Repeat mix 11 Manufactured sand # 1 2 3 4 1 2 3 4 1 4_2 4_3 4_4 4_5 4_6 4_7 Test

# 11 12 13 Mix

306