Modelling the pumping characteristics of power station ash

Modelling the pumping characteristics of

power station ash in a dense phase

hydraulic conveying system

A thesis submitted for fulfilment of the requirements

for the award of the degree of

Doctor of Philosophy

From

The University of Newcastle, Australia

By

Thomas Francis Bunn ME Mechanical Engineering, Newcastle University

BSc Macquarie University

Faculty of Engineering and Built Environment Centre for Bulk Solids and Particulate Technologies and TUNRA Bulk Solids

March 2015

i

CERTIFICATION

I, Thomas Francis Bunn, declare that this thesis, submitted in fulfilment of the

requirements for the award of Doctor of Philosophy, in the Faculty of Building

Environment and Engineering, The University of Newcastle, contains no material

which has been accepted for the award of any other degree or diploma in any university

or other tertiary institution and, to the best of my knowledge and belief, contains no

material previously published or written by another person, except where due reference

has been made in the text. I give consent to the final version of my thesis being made

available worldwide when deposited in the University’s Digital Repository, subject to

the provisions of the Copyright Act 1968.

(Signed): ……………………………………….……….

Thomas Bunn

ii

ACKNOWLEDGEMENTS

The work for this thesis has been carried out with the Centre for Bulk Solids and

Particulate Technologies at the University of Newcastle. I would like to thank the

directors, Professor Mark Jones and Associate Professor Craig Wheeler, who were also

my co-supervisors, for providing the opportunity to study within the Centre. Over the

course of my studies, both Mark and Craig have been very helpful, and they have

offered kind words of encouragement when needed. .

The technical staff at TUNRA Bulk Solids must also be acknowledged as they have

offered support at different stages of my research. Every member of staff was always

more than happy to offer help when I needed it. In particular, I would like to thank

fellow doctoral student Wei Chen for many enlightening discussions on modelling and

rheology and for proof reading this thesis.

Lastly, and most importantly, I would like to acknowledge my family and friends.

My family though, and in particular my wife Elizabeth and daughter Kate, have been

instrumental through the course of my research in keeping me focused, happy and sane.

To them I offer many thanks.

iii

TABLE OF CONTENTS

CERTIFICATION i

ACKNOWLEDGEMENTS ii

TABLE OF CONTENTS iii

ABSTRACT x

LIST OF PUBLICATIONS xi

NOMENCLATURE xv

CHAPTER 1 INTRODUCTION 1

CHAPTER 2 LITERATURE REVIEW 6

2.1 Introduction 6

2.2 History 7

2.3 Lean Phase Power Station Ash Disposal 11

2.4 High Concentration Power Station Ash Disposal 13

2.5 Bayswater Dense Phase Power Station Ash Disposal System Plant 17

2.5.1 Operating Procedure 18

2.5.1.1 Bayswater Pipeline Rheology 19

2.5.1.2 Ravensworth Ash Disposal Site 20

2.5.1.3 Water Reclamation 20

2.6 Callide B High Concentration Slurry Disposal Plant 21

2.7 Concluding Remarks 23

CHAPTER 3 RHEOMETRY AND RHEOLOGICAL MEASUREMENT 24

3.1 Slurry Rheology Introduction 24

3.2.1 Time-independent slurries 24

3.2.1.1 Viscous Behaviour 24

3.2.1.2 Newtonian Behaviour 25

3.2.1.3 Pseudoplastic Slurries 25

3.2.1.4 Dilatant Slurries 26

3.2.1.5 Plastic Behaviour 27

3.2.1.6 Yield-Pseudoplastic and Yield-Dilatants Slurries 27

iv

3.2.2 Time-Dependent Slurries 27

3.3 Introduction Rheometry and Rheological Measurement 28

3.3.1 Capillary Tube Viscometer 29

3.3.2 Laminar Flow in Cylindrical Tubes 29

3.3.3 Errors in Capillary Viscometry 33

3.3.4 Applications for Capillary Viscometers 33

3.4 Concentric Cylinder Rotational Viscometers 34

3.4.1 Principle of Operation 34

3.4.2 Sources of Errors in Rotary Viscometers 37

3.4.3 Applications for Rotary Viscometers 38

3.5 Slurry Flow 38

3.6 Homogeneous Fluid Models 40

3.7 Rheology Studies of Fly Ash 43

3.8 Flow Cones 44

3.8.1 Flow Cones as Rheological Devices 46

CHAPTER 4 EMPIRICAL APPROACH 50

4.1 Introduction 50

4.1 Estimation of Critical Velocity 50

4.2 Determining Pipeline Pressure Drop – Head Loss 55

CHAPTER 5 PREVIOUS RESEARCH 59

5.1 Introduction 59

5.2 Vales Point Dense Phase Ash Pumping Plant 59

5.2.1 Dense Phase Pumping Plant Pipeline Sizing 61

5.2.3 Dense Phase Pumping Plant Control System 62

5.2.4 Dense Phase Pumping Plant Operations 65

5.2.5 Determination of Pipeline Slurry Settling Velocity 68

5.2.6 Dense Phase Pumping Plant Slurry Transfer 69

5.3 Pipeline Viscometers 70

5.3.1 Mono Pump Test Rig 71

5.3.2 Mixing Technique and Measurements for Mono Pump Test Rig 71

5.3.3 Calculations Mono Pump Test Rig 72

v

5.4. Rotary Ram Slurry Pump Thornton Test Rig 74

5.4.1 Mixing Technique and Measurements for Rotary Ram Slurry Pump 74

5.4.2 Calculations for the Rotary Ram Slurry Pump 75

5.5 Viscometers Results 75

CHAPTER 6 PREVIOUS RESEARCH PAPERS 78

6.1 Introduction 78

6.2 Summary 78

6.3 11th

International Conference Bulk Materials Storage

Handling and Transportation (2013) - Comparative

Rheology of Fly Ash Slurries using Rotary and Pipeline Viscometers 81

6.3.1 Experimental Material and Equipment 84

6.3.2 Slurry Mixing and Measurement 86

6.3.3 Experimental Results and Analysis 88

6.3.4 Conclusions 93

6.4 7th

International Conference for Conveying and Handling of

Particulate Solids - ChoPS (2012) - Comparison between

Flow Cones and a Rotary Viscometer 95

6.4.1 Particle Size Distribution and Density 96

6.4.2 Methodology 97

6.4.3 Results and Discussions 98

6.5 International Freight Pipeline Society Symposium (2011)

- The Pumping Characteristics of Fly Ash Slurry in a Pipeline 102


6.2 Results and Discussions 105

6.6 International Seminar on Paste and Thickened Tailings (2010)

- Pumping Power Station Ash as a High Concentration Slurry 109


6.6.2 Fly Ash Testing with Rotary Viscometry 113

6.6.3 Pilot Pumping Plant 113

6.6.4 Slurry Mixing and Pumping 115


vi

6.7 6th

World Congress on Particle Technology (2010) - Thixotrophic

Behavior of Fly Ash Slurries 121



6.8 The 6th

International Conference for Conveying and Handling

Particulate Solids and 10th

International Conference on Bulk

Materials Storage, Handling and Transportation (2009)

- Are Tailing Dams Viable in the Modern Environment? 126

6.8.1 Why Are Tailing Dams Still Being Built? 129

6.8.2 Alternative Disposal Systems 129

6.8.3 Example of Industries Changing from Slurry to Paste Production 132

6.8.4 Material Handling Solution for Disposal to Underground Mine Voids 134

6.8.5 Conclusion 135

6.9 Innovation in Bulk Materials Handling & Processing (2008) and

Australian Bulk Handling Review, Volume 14 No. 1 (2009)

- The Pumpability of Coal Washery Thickener Underflow 137



6.10 International Symposium of Reliable Flow of Particulate

Solids IV (RELPOWFLOW IV), (2008) – Water Available

for Recycling after the Placement of Dense Phase Fly Ash 142



6.11 9th

International Conference on Bulk Materials Storage,

Handling and Transportation (2007) - The Relationship

between Packing Density and Pumpability of Fly Ash Slurries 149



6.12 5th

International Conference for Conveying and Handling

Particulate Solids (2006) - The Effect of Particle Size

Distribution on the Rheology of Fly Ash Slurries 155


vii


6.13 5th

World Congress on Particle Technology (2006)

- A Model to Determine the Packing Density of Fly Ash Slurries 160

6.13.1 Simulation Model 160

6.13.2 Simulation Model Validation 162



6.13.5 Packing Efficiency Calculation 165

6.13.6 Conclusion 169

6.14 16th

International Conference on Hydrotransport (2004) – What

a change in coal supply can mean to a dense phase handling

and pumping system for a large coal fired power station 170


6.14.3 Conclusions 177

CHAPTER 7 HIGH CONCENTRATION SLURRY TESTING 178

7.1 Introduction 178

7.1.1 Pipeline Viscometer 178

7.1.2 Rotary Viscometer 187

7.1.3 ASTM Flow Cone 188

7.1.4 Calibration of Test Rig Instrumentation 189

7.1.4.1 Calibration of Weigh Hopper 189

7.1.4.2 Calibration of Pressure and Differential Pressure Transmitters 190

7.1.4.3 Calibration of PT 100 Resistance Temperature Detector 193

7.2 Slurry Mixing 194

7.3 Slurry Testing 196

CHAPTER 8 RESULTS AND DISCUSSIONS 198


8.2 Pipeline Viscometers Water Tests 198

8.3 Fly Ash “B” Characteristics 200

8.4 Comparison of Slurry Flows Measurements 201

8.5 Testing Fly Ash “B” Slurry in the Test Facility 203

viii

8.6 Determining Non-Newtonian Fly Ash “B” Slurry Characteristics 211

8.7 Non- Newtonian Slurry Modelling Fly Ash “B” 214

8.8 Site Collected Data 217

8.9 Non- Newtonian Slurry Grout Modelling Fly Ash “B” 218

8.10 Fly Ash “B” Slurries Comparison of 50 mm and 80 Pipeline Viscometers 220

8.11 Fly Ash “E” Characteristics 222

8.12 Testing Fly Ash “E” Slurry in Test Facility 224

8.13 Determining Non-Newtonian Fly Ash “E” Slurry Characteristics 231

8.14 Slurry Modelling Fly Ash “E” 234

8.15 Non- Newtonian Slurry Grout Modelling Fly Ash “E” 238

8.16 Fly Ash “E” Slurries Comparison of 50 mm and 80 Pipeline Viscometers 241

8.17 Fly Ashes “B” and “E” Slurries Comparison of

Pipeline Pressure Drop Models 242

8.18 Fly Ash “E” Determining the Settling Velocity 245

8.19 Fly Ash “B” and “E” Laminar or Turbulent Flow 247

8.20 Fly Ash “B” and “E” Homogeneous or Heterogeneous Slurries 252

8.21 New Definition for Fly Ash Slurries Homogeneous Behaviour 254

8.22 Spread Sheet Program 256

8.23 Determine the Standard Error of the Models 260

CHAPTER 9 CONCLUSIONS


9.2 Pipeline Viscometers Water Tests 262

9.3 Fly Ash “B” Characteristics 262

9.4 Comparison of Slurry Flows Measurements 263

9.5 Testing Fly Ash “B” Slurry in the Test Facility 263

9.6 Non-Newtonian Fly Ash “B” Slurry Characteristics 264

9.7 Non-Newtonian Slurry Modelling Fly Ash “B” 264

9.8 Site Collected Data Comparison 265

9.9 Non- Newtonian Slurry Grout Modelling Fly Ash “B 266

9.10 Fly Ash “B” Slurries Comparison of 50 mm and 90 Pipeline Viscometers 266

9.11 Fly Ash “E” Characteristics 267

9.12 Testing Fly Ash “E” Slurry in Test Facility 267

ix

9.13 Non-Newtonian Fly Ash “E” Slurry Characteristics 268

9.14 Non-Newtonian Slurry Modelling Fly Ash “E” 268

9.15 Non- Newtonian Slurry Grout Modelling Fly Ash “E” 269

9.16 Fly Ash “E” Slurries Comparison of 50 mm and 80 Pipeline Viscometers 270

9.17 Fly Ashes “B” and “E” Slurries Comparison of

Pipeline Pressure Drop Models 270

9.18 Fly Ash “E” Determining the Settling Velocity 271

9.19 Fly Ashes “B” and “E” Laminar or Turbulent Flow 271

9.20 Fly Ash “B” and “E” Homogeneous or Heterogeneous Slurries 272

9.21 Redefining Homogeneous Behaviour for Fly Ash Slurries 272

9.22 Spread Sheet Program 272

9.23 Conclusions 273

9.24 Recommendations 276

BIBLIOGRAPHY 278

APPENDIX A – Data 323

x

ABSTRACT

This study examines the flow of dense phase fly ash slurries in horizontal pipes. It

includes an evaluation of the previous work, a rigorous experimental investigation, a

new and original model for determining pipeline pressure drop characteristics and a new

method of characterising typically homogeneous fluid behaviour based on a particle size

distribution, slope factor and a median particle size.

The experimental investigation was undertaken to obtain data for modelling the flow of

dense phase fly ash slurries. Tests were conducted using fly ashes from different power

stations in a purposely built test facility. The test facility contained 50 mm and 80 mm

bore internal pipeline viscometers in series.

Slurry pump discharge pressure, differential pressure over 5 meters of a 80 mm pipe

section, differential pressure over 5 meters of a 50 mm of pipe section, slurry temperature,

slurry volumetric and mass flowrates were measured. Slurries settling were determined

visually using an 80 mm glass pipe section. The particle size distribution and solids

density of the fly ash were analysed and the solids concentration of the slurries were

determined using the wet weight, drying and dry weight method.

The experimental results were used to develop a new model to determine the pressure

drop characteristics of dense phase fly ash slurry pumping systems and grout pumping

plants, in order to develop a new description of what typical characteristics

homogeneous fluid contain. The model indicated a polynomial relationship between

pipeline differential pressure and solids concentration which has proven to be a much

improved predictor of actual system performance.

A software based design program has been produced that utilises power station physical

and operational details to determine the pumping characteristics of dense phase ash slurries

which will lead to better practical outcomes in the power industry.

xi

LIST OF PUBLICATIONS

The following is a list of publications achieved by the author prior to the submission of

this thesis.

Bunn T. F., Jones M. and Wheeler C. A., (2013), "Comparative rheology of fly ash

slurries using a rotary and pipeline viscometers", 11th


Materials Storage, Handling and Transportation, Newcastle, Australia, 2 – 4 July.

Bunn T. F., Jones M., Wheeler C. A. and Wedmore G., (2012), "Comparison between

Flow Cones and a Rotary Viscometer", 7th

International Conference for Conveying and

Handling of Particulate Solids – ChoPS 12, Friedrichshafen, Germany, 10 – 13

September.

Bunn T. F., Jones M. and Wheeler C. A., (2011),"The Pumping Characteristics of Fly

Ash Slurry in a Pipeline", International Freight Pipeline Society Symposium, Madrid,

Spain, 29 June – 1 July.

Bunn T. F., Jones M. and Wheeler C. A., (2010) "Pumping Power Station Ash as a High

Concentration Slurry", 13th

International Seminar on Paste and Thickened Tailings,

Toronto, Canada, 3 - 6 May.

Bunn T. F., Jones M. and Wheeler C. A., (2010) "Thixotrophic Behavior of Fly Ash

Slurries". 6th

World Congress on Particle Technology, Nuremberg, Germany 26 - 29

April.

Bunn T. F., Jones M. and Wheeler C. A., (2009) "Water Available for Recycling after

Placement of High Concentration Fly Ash Slurries". Australian Bulk Handling Review,

Volume 14 No. 5 September/October.

xii

Bunn T. F., Jones M. and Wheeler C. A., (2009) "The Pumpability of Coal Washery

Thickener Underflow", Australian Bulk Handling Review, Volume 14 No. 1 February.

Bunn T. F., Gilroy T., Jones M. G. and Wheeler C.A., (2009) "Are Tailing Dams Viable

in the Modern Environment? ", The 6th

International Conference for Conveying and

Handling Particulate Solids and 10th

International Conference on Bulk Materials

Storage, Handling and Transportation, Brisbane, Queensland, Australia, 3 – 7 August

pp. 615 - 620.

Bunn T. F., Jones M. and Wheeler C. A., (2008) "The Pumpability of Coal Washery

Thickener Underflow". Innovation in Bulk Materials Handling & Processing, Sydney,

NSW, Australia 26 -27 November.

Bunn T. F., Jones M. and Wheeler C. A., (2008) "Water Available for Recycling after

Placement of Dense Phase Fly Ash Slurries". International Symposium of Reliable Flow

of Particulate Solids IV (RELPOWFLOW IV), Tromso, Norway, 10 – 12 June.

Bunn T. F., Jones M. and Wheeler C. A., (2007) "The Relationship between Packing

Density and Pumpability of Fly Ash Slurries", 9th


Materials Storage, Handling and Transportation, Newcastle, NSW, Australia, 9 – 11

October.

Bunn T. F., Jones M. and Wheeler C. A., (2006), "The Effect of Particle Size

Distribution on the Rheology of Fly Ash Slurries", 5th

International Conference for

Conveying and Handling Particulate Solids, Sorrento, Italy, 27 – 31 August.

Bunn T. F., Jones M., G., Donohue T., J. and Wheeler C.A., (2006), "A Model to

Determine the Packing Density of Fly Ash Slurries", 5th

World Congress on Particle

Technology, Lake Buena Vista, Florida, USA, 23 – 27 April.

Bunn T. F., Jones M. G. and Wiche S., (2004), "What a change in coal supply can mean

to a dense phase handling and pumping system for a large coal fired power station", 16th

International Conference on Hydrotransport, Santiago, Chile, 26 – 28 April.

xiii

Bunn T. F., and Chambers A. J., (1999), "Pressure Loss Calculations for Thickened

Slurries Containing Large Particles", 14th

International Conference on Slurry Handling

and Pipeline Transport, Maastricht, The Netherlands, 8 – 10 September.

Bunn T. F., and Chambers A. J., (1998), "Experiences Pumping Dense Slurries

Containing Large Particles", 46th Japanese National Conference on Rheology, August,

Rakuno-Gakuen University, Sapporo, Japan, pp. 117-118.

Ward A., Bunn T. F. and Chambers A. J., (1998), "Transportation of Fly Ash, The

Bayswater Ash Disposal System", In proceedings International Symposium Upgrading

and Slurrification of Low Rank Coals, September, Faculty of Engineering, Kobe

University, Japan, pp. 102-115.

Ward P. and Bunn T. F., (1997), "The use of High Density Technology for Power

Station Fly Ash Disposal and Mine Rehabilitation", Successful Tailings Management,

Sydney, Australia.

Bunn T. F., (1995), "Progression from Research to Pilot Plant to Full Size Plant – Dense

Phase Ash Slurry Conveying of Power Station Ash", 5th

International Conference on

Bulk Materials Storage, Handling and Transportation, Newcastle, NSW, Australia.

Bunn T. F. and Chambers A. J., (1995), "Pipeline Transport of Power Station Ash as a

High Mass Concentration Slurry", International Journal of Storage, Handling and

Processing Powder, 2/95, pp. 133 - 137.

Bunn T. F. and Chambers A. J., (1992), "Experiences with Dense Phase Hydraulic

Conveying of Vales Point Fly Ash", International Journal of Storage, Handling and

Processing Powder, No. 3, pp. 221 - 226.

Bunn T. F. and Chambers A. J., (1992), "Experiences with Dense Phase Hydraulic

Conveying of Vales Point Fly Ash", 4th

International Conference on Bulk Materials

Storage, Handling and Transportation/7th

International Symposium on Freight

Pipelines, Wollongong, NSW, Australia: 6-8 July, pp. 75-83.

xiv

Bunn T. F. and Chambers A. J., (1991), "Characterisation of Fly Ash Slurries",

International Mechanical Engineering Congress, Sydney, NSW, Australia, 8 - 12 July,

pp. 50 - 61.

xv

NOMENCLATURE

𝐴 Cross sectional area (m2)

a Symounds constant

𝐶𝐷 Drag coefficient

𝐶𝑣 Solids concentration by volume (%)

𝐶𝑤 Solids concentration w/w

𝐶𝑤 Concentration by weight (%)

𝐷 Pipe diameter (m)

𝑑𝑠 PSD slope curve

𝑑10 10th percentile particle diameter (m)



𝐹𝐷𝐿 Durand velocity factor

𝑓 Friction factor

FCT Flow cone time (s)

g Acceleration due to gravity (m s-1

)

𝐻𝑒 Hedstrom number

𝐻𝐹 Total height of the cone portion of the funnel (cm)

ℎ0 Initial height in the funnel (cm)

ℎ𝑓 Head loss (m)

ℎ𝑠𝑠 Steady state height in the funnel (cm)

k′ Consistency index

L Length (m)

𝑁 Index number

n Size of the sample

n′ Flow behaviour index

𝑃 Pressure (kPa)

PSD Particle size distribution

𝑄 Volumetric Flowrate (m3 s

-1)

𝑅 Tube radius (m)

xvi

R2

Comparing the variability of the estimation errors with the

variability of the original values

𝑅𝑎 Outer cylinder radius (m)

𝑅𝑒 Reynolds’ Number

𝑅𝐹 Maximum radius of the funnel (cm)

𝑅𝑖 Cup radius (m)

𝑟 Radius (m)

𝑟 Radial coordinate

𝑟1 Radial coordinate at rotor surface

𝑆𝐸 Standard error

𝑆𝐺 Specific gravity

𝑆𝑓 Slope factor

SSE Sum of squares due to error

SST Total sum of squares

s Sample standard deviation

𝑇 Torque (Nm)

𝑡 Time (s)

𝑡𝑓 Total drainage time (s)

𝑉 Average velocity (m-2

)

𝑉𝑠 Settling velocity (m-2

)

𝑊 Mass flowrate (kg s-1

)

Greek Symbols

∆𝑃 Differential pressure (kPa)

Shear rate (s-1

)

𝛿 Ratio of the radii

𝜇 Newtonian viscosity (Pa s)

μa Apparent viscosity (Pa s)

μc Effective viscosity (Pa s)

μf Carrier Fluid viscosity (Pa s)

υ Kinematic viscosity (m2 s

-1)

https://en.wikipedia.org/wiki/Standard_deviation#Corrected_sample_standard_deviation

xvii

𝜌𝑠 Solids density (kg m-3)

𝜌𝑠𝑙 Slurry density (kg m-3)

𝜌𝑤 Water density (kg m-3)

τb Bingham model shear stress (Pa)

τ Shear stress (Pa)

τc Casson yield stress (Pa)

τw Wall shear Stress (Pa)

τy Bingham yield stress (Pa)

τ1 Shear stress rotor surface (Pa)

τ0 Shear stress (Pa)

Γ𝑤 Apparent shear rate (s-1

)

1

CHAPTER 1: INTRODUCTION

Modern coal fired power stations in New South Wales (Eraring and Bayswater) burn a

large amount of coal up to 7 × 106 t y

-1. The combustion of such a large quantity of coal

results in the production of large quantities of ash that has to be removed from the gas

stream.

The coal is delivered from the crushing mills by hot air and burnt in the furnace. The

coal burnt in the furnace is ground to the fineness of 75 % < 75 µm, 90 % <150 µm and

99.9 % < 300 µm. The products of combustion then pass through superheaters, re-

heaters, economisers, air heaters and into the fly ash collection system. Ash classified

as bottom ash is collected from the bottom of the furnace and from hoppers under the

economisers and or air heaters (grits). The remaining ash (fly ash) is separated from

the gas stream through fabric filters before it passes out of the chimney. Coals burnt

can have an ash content of up to 30 % by weight, therefore a power station burning 7 ×

106 t y

-1 could produce up to 2.1 × 10

6 t y

-1 of ash. It is generally accepted that up to 15

% of the ash produced is bottom ash and the remaining ash is fly ash.

The pumping of power station ash for disposal prior to 1990 was always by lean phase

slurry systems which contained, at most, 10 % by weight of solids. Although many

dense phase hydraulic transport systems have been installed, around the world to

hydraulically convey a variety of materials, no operational systems have been installed to

convey power station ash over long distances, (Sive 1989). Bunn and Gorsuch (1988)

reported that the solids concentration of the Eraring Power Station bottom and fly ash

slurry system at full load was 3 % solids for the bottom ash and 7 % solids for the fly

ash.

At Eraring, which is a zero release station, the lean phase system for fly ash and bottom

ash requires the pumping of 2500 m3

h-1

of ash and water to the disposal site and the

recycling of the same amount of water to the station. A dense phase hydraulic system

would be both economically and environmentally superior. For example, to pump Eraring

fly ash in a dense phase hydraulic system requires a pump with a capacity of 240 m3 h

-1

and a return water system capable of 120 m3 h

-1. Therefore, the cost of pumping the fly ash

2

slurry and return water would be greatly reduced.

The greatest challenge for any designer of a dense phase hydraulic system has to be the

variability of the quality of fly ash received from the power station. As an example,

approximately 400 000 t y-1

of fly ash from Eraring Power Station was sold to the cement

industry. The specification for the supply of the fly ash requires that it has to be processed

so that the loss of ignition products are < 4 % and has a fineness 90 % < 45 µm. The

removal of this quantity of fine material adversely affects the PSD and therefore the

pumpability of the fly ash slurries. This along with changes in coal supply, coal milling

system maintenance and power system load changes lead to large variations in the PSD

of the “run of station” fly ash for disposal.

Bunn et al. (2007) postulated that to pump fly ash slurries requires all the void spaces

between the fly ash particles to be filled with water and extra water added to transport

the slurry through the pipeline. To attain the higher pumping Cw’s required that the void

spaces between particles to be filled with fly ash particles not water. The greater the

range of different sized particles the less the void spaces. The removal of fines < 45 µm

from fly ash for the cement reduces the pumpable Cw of slurry. In the same way the fine

ash that appears in chimneys from power stations with precipitators also reduces the

pumping Cw of the fly ash slurries.

The source and makeup of the coal mean that the components of the fly ash produced

vary considerably. A typical chemical analysis of the fly ash indicates that it contains

substantial amounts of silicon dioxide (SiO2) (between 55 to 75 %) and aluminium

oxide (Al2O3) (between 15 to 30 %). The analysis also reveals that a combination of

these two constituents make up approximately 90 % of the fly ash constituents.

Scanning Electron Microscope analysis indicates that the fly ash particles are

predominantly spherical in shape (Bunn et al. 2004).

During the grouting of disused underground coal mines on the Hunter Freeway Project

over one week in November 2011, (Wedmore 2011) reported that the PSD of the

Bayswater “run of station fly ash” was both variable and unpredictable. This is shown

by the variance in the weight of water required in the batching of a 2-ton mixture of fly

3

ash and cement grout. The weight of water needed to achieve a specified flow cone time

of 20 seconds varied between 800 kg to 1200 kg. Therefore, the Cw of the grout pumped

varied between 62.5 % and 71.5 % at a similar viscosity

Scope

A basic understanding of the underlying phenomena is vital to the design and control of

a dense phase slurry transport system. Literature review reveals that studies concerned

with solid-liquid mixture flows have followed either the rheological or the empirical

approach.

The rheological approach, as the science of flow phenomena, made a significant impact

in the 1950’s. In the context of this study of rheological viscous characteristics of slurry,

specifically the relationship between shear stress and shear rate, are applicable to

slurries of ultra-fine non-colloidal particles.

The empirical approach seems to have received the most attention, perhaps as a

concession to the complexity of slurry flows. Because of its long history and an

increasingly large body of knowledge of empirical studies dealing with slurry transport,

there has been an accumulation of correlations for the prediction of critical velocity,

pressure drop and classification of flow regimes.

The objective of this study is to develop a model, using a rheological approach, which

accurately predicts the behaviour of solids transport in laminar, non-Newtonian, pipe

flow. The added complexity of turbulent flow is beyond the scope of this work and will

not be considered.

Heterogeneous slurries exhibit more complicated flow behaviour when compared to

homogeneous slurries. As a result, concentrations across the flow domain are non-

uniform and distorted.

4

Thesis Outline

Chapter 1

This chapter introduces the idea of the variability in the quality of fly ash in the

combustion of coal in the generation of power. The concepts of a rheological approach

and an empirical approach to predict the behaviour of solids transport in laminar, non-

Newtonian, pipe flow are also investigated. Chapter 1 also outlines the scope of the

thesis as well as its structure.

Chapter 2

The literature review presents an examination of past work, the history of slurry

pipelines and introduces the lean phase system that operates in all power stations prior

to the 1990’s. It also presents an overview of high concentration power station ash

disposal installed in Australia.

Chapter 3

The theory of rheometry and rheological measurement are introduced while discussing

rheological behaviour and the measurement techniques used. It also discusses slurry

flows, homogeneous fluid models, rheology studies of fly ash and flow cones.

Chapter 4

This chapter outlines the empirical approach in determining pipeline critical velocity,

starting with the Durand in 1953, and pipeline pressure drop – head loss by predicting

the friction factor from the Moody diagram.

Chapter 5

This chapter is a description of the work contained in research thesis tilted “The Dense

Phase Hydraulic Conveying of Power Station Ash”, that was submitted by the author in

1991 for a Master Degree of Mechanical Engineering, University of Newcastle.

5

Chapter 6

This chapter contains a description of all the papers that the author has published over the

last ten years, which have been presented at national, international conferences or

published in journals to further his understanding of the transport and disposal of power

station ash.

Chapter 7

This chapter describes the slurry testing facility as well as the comparative testing of

different high concentration fly ash slurries. Comparative rheological analyses were

undertaken using a pipeline viscometer, rotary viscometer and an ASTM flow cone.

Chapter 8

This chapter summarise the findings and presents a prediction model that will

accurately reproduce the pressure drop values experimentally obtained from the test

facility.

Chapter 9

Processing of this data led to a number of valuable correlations which will be of key

importance in the development and assessment of a successful pressure drop prediction

model.

6

CHAPTER 2 LITERATURE REVIEW

2.1 Introduction

With the exception of a few sites, the disposal of ash from power stations invariably

requires hydraulic transport of the solids through pipes from the plant to the disposal

site. If one considers tailings, inter-process transfers and freight products, the amount of

material that is conveyed hydraulically by pipeline each year is staggering. Despite this

ubiquity, pipeline transport is still treated with suspicion and remains a dark art to

many.

(Conventional tailings disposal since that time has typically involved pumping very low

concentrations of solids to large catchments). Here the solids settle, forming a denser

bed, while the conveying water is either drained to the environment, returned to the

plant or simply left to evaporate. This mode of transport is relatively simple, as low

concentrations of small particles are unlikely to block pipes, and the transport

characteristics occur under pipe turbulent flow. This form of tailings disposal, however,

is generally no longer acceptable in the 21st century, where environmental and economic

imperatives prevent the construction of such large tailings dams, or allow such low

concentration suspensions to be pumped to disposal sites.

Flow behaviour is the result of complex interactions between fluid dynamics, rheology

and particles science and can range from the simple laminar flow of homogenous

materials through turbulent suspension flows to granular flows, where the solids are

conveyed as a packed bed. The result of these imperatives has been to increase the

solids concentration of the suspension delivered to the disposal sites or tailings dams.

This increase, however, dramatically changes the behaviour in the pipelines, as particle

to particle interaction starts to dominate the flows.

Since the latter half of the 20th

century, it has been obvious that a paradigm shift was

required and new forms of waste disposal are required. Now, rather than pumping the

material out to the dam at low concentrations and allowing settling and evaporation to

7

concentrate the deposit, it was proposed that the tailings be pumped at a higher

concentrations. The higher concentration discharge systems had the advantage that they

required a much smaller footprint than conventional dams and could be built on flat

planes. Other alternatives are the disposal of tailings in worked out open cut or

underground mines. These high concentration flows can similarly be run under laminar

flow at low velocities, and, providing the high solids concentrations is maintained,

without the fear of blocking the pipe. For the relatively short distances, i.e. tens of

kilometres, required for tailings disposal, the pressure gradient is no longer the

overarching constraint. Instead, minimizing the size of the deposit, minimizing water

consumption, improving deposit stability, increasing drainage and reducing chemical

species mobility are more important criteria.

2.2 History

A slurry pipeline is used to transport solid particles entrained in a fluid flowing in a

pipeline. The earliest mention of transport of slurries was in open channels by (Hoover

and Hoover 1912), translated from the Latin book Georgius Agricola - De Re Metallica

(1566). Figure 2.1 is an illustration from the book showing launders and open channel

flow. Here the solids are mixed with the liquid (usually water) prior to flowing into the

launders, whereas the solids are usually kept in suspension while flowing in the launder

and are separated from the liquid upon exiting from the launder at the destination.

Slurry pipelines have been using water as the carrier fluid since about 1880 to transport

solid material including coal, limestone, copper, iron ore concentrates and phosphate.

For example, (Pullum 2008) cited in London, coal unloaded from barges in the river

Thames was transported via an underwater pipeline to Battersea Power Station, close to

the Houses of Parliament. This lump coal was pumped as a low concentration

heterogeneous suspension in water, an inherently unstable mode of transport, which

eventually blocked the pipeline, and Pullum believes it remains blocked to this day.

It is perhaps this public, less than auspicious, start to a technology and the surprisingly

complex nature of the many flow regimes that has slowed down the acceptance of new

modes of hydraulic transport and ensured that a very conservative approach be adopted.

8

Figure 2.1 Illustration from Georgius Agricola - De Re Metallica (1566).

This is especially true for low value products such as waste streams from coal fired

power stations.

The modern beginning of long distance hydraulic transport commenced in the 1957

with the commissioning of the 173 km, 254 mm diameter Consolidation Copal pipeline

between Cadiz, Ohio and Lake Erie, USA, (Wasp et al. 1977). The pipeline capacity was

1.3 × 106 t y

-1. This pipeline was constructed to pump coal at a rate of 3,700 t d

-1 a

distance of 173 km at a Cw of 50.0 % s. The pumps were Wilson-Snyder 336 kW duplex

double acting piston pumps operating at a pressure of 8.3 MPa.

9

The classic long distance pipeline was the “Black Mesa Pipeline” that carried coal from

the coal mine near Kayenta, Arizona to the Mohave Power Plant in southern Nevada,

(Wasp et al. 1977). From 1970 to 2005 the Black Mesa slurry pipeline carried 4.5 × 106

t y-1

of coal through a 440 kilometre long steel pipeline with an internal diameter of 240

mm. The coal was crushed to < 1 mm and mixed with water to form slurry with a Cw of

50 % before being pumped in the pipeline. The pipeline contains four pumping stations

fitted with four 1700 PT Wilson-Snyder duplex double acting piston pumps in parallel and

the other two pumping stations with three 1700 PT Wilson-Snyder pumps in parallel. The

pipeline starts at Kayenta at an elevation of 1830 meters and ends power station at an

elevation of 230 meters. The pipeline operations were suspended in 2005 due to

shortage of water and the cost of refurbishing the power plant to meet new pollution

standards. The mining process and pumping the coal was using four billion litres of

water per year and towns as far as 80 km away from the mine site were noticing a

substantial loss of groundwater.

In Australia, the first significant slurry pumping plant was at the Savage River in

Tasmania where iron ore concentrate was pumped from the mine at Savage River 85 km

to a pellet plant at Port Latta through a 230 mm steel pipeline, (Wasp el al. 1977). At the

time of commissioning in 1967, this was the longest iron ore pipeline in the world with

a throughput of 2.25 ×106 t y

-1. After restructuring in 1990, the throughput was reduced

to 1.5 × 106 tons per year.

In Gladstone Queensland, a pipeline transports a mixture of limestone, clay overburden,

ironstone, sand and water at Cw = 62.0 % from the East End Mine to Fisherman’s Landing,

a distance of 25 km. The kiln at Fisherman’s Landing produces 500,000 t y-1

of cement

clinker. The high pressure pumping plant consists of two Wilson-Snyder, twin cylinder,

and double-acting, positive displacement piston pumps with inlet and outlet valves. The

pumps each have a capacity of 250 m3 h

-1 at 50 strokes per minute. The pumps are driven

by an 860 kW electric motor via a hydraulic speed control unit. The single welded steel

pipeline has a diameter of 200 mm, a wall thickness of 10 mm and is totally buried

underground.

The use of high capacity positive displacement pumps to transport a stable slurry mixture

10

of coarse and fine coal with water was described by, (Brooks and Snoek 1986). The pumps

were Putzmeister single-acting duplex piston pumps with a rating of 125 m3 h

-1 at a

pressure of 5 MPa with a hydraulically actuated change-over "Delta" outlet valve. The

pumps were used to transport the slurry mixture through several test loops to demonstrate

the advantages of this type of mixture.

The line pressure, the mass flow rate and the abrasively of the material to be transported

are important factors when selecting the pump type for dense phase hydraulic conveying of

solids, (Bhambry and Wallrafen 1987). Higher discharge pressure requirements for Cw >

60.0 % rule out the application of centrifugal pumps. Reciprocating pumps are therefore

used as they are capable of producing discharge pressures up to 35 MPa. These pumps

have the advantage of high volumetric efficiency at the desired flowrate.

In describing the history of positive displacement pumps, (Prudhomme et al. 1970),

indicated that the first pumps were oilfield mud pumps which are capable of pressures up

to 28 MPa.

The longest slurry pipeline in the world is the proposed 550 km Anglo American

Minas-Rio JV mining operation. Here, the iron ore will be turned into slurry and

pumped down the pipeline to the coastal terminal at Port of Acu. The pipeline’s

capacity is 26.6 × 106 t y

-1 (dry base). The pumping system contains 18 Geho positive

displacement pumps will be installed in two pump stations, one at the mine with 8

pumps and one pump station about half way, where 10 pumps are installed. The Geho

Positive Displacement pumps will develop pressures up to 20.6 MPa to transport the

heavy iron ore slurry. The longest pipeline in Australia, the Century Zinc/Lead Slurry

Pipeline in Queensland, is a single pipeline operation. The pipeline simultaneously

transports lead or zinc concentrates a distance of 304 km from the Zinifex Century Mine at

Lawn Hill to the port facility at Karumba on the Gulf of Carpentaria. The 300 mm nominal

bore high-density polyethylene (HDPE) lined slurry pipeline simultaneously transports

both lead or zinc concentrate slurry at a nominal flow rate of 304 m3

h-1

at a pressure of

nominally 11 MPa and velocity of 1.1 m s-1

. The different slurries are separated by 1 hour

pumping of water. Both slurries have a nominal concentration of solids Cw 35% to 37%,

(Hoskins 2002).

11

2.3 Lean Phase Power Station Ash Disposal

The power stations at Eraring and Bayswater burn a large amount of coal up to 7 × 106 t

y-1

. The combustion of such a large quantity of coal results in the production of large

quantities of ash that has to be removed from the gas stream. The coal is delivered from

the crushing mills by hot air and burnt in the furnace. The products of combustion then

pass through superheaters, re-heaters, economisers, air heaters and into the fly ash

collection system. Ash classified as bottom ash is collected from the bottom of the

furnace and from hoppers under the economisers and or air heaters (grits). The

remaining ash (fly ash) is separated from the gas air stream either in precipitators or

fabric filters before it passes out of the chimney. Australian coals have an ash contents

in the range of 15% to 30 %. Therefore for a power station burning 7 × 106 t y

-1

produces between 1 × 106 to 2.1 × 10

6 t y

-1 of ash. It is generally accepted that up to 15

% of the ash produced is bottom ash and the remainder is fly ash.

The bottom ash is usually collected in water filled hoppers located at the bottom of the

furnace. The ash is removed on a routine bases by dumping the contents into a

sluiceway lined with basalt tiles where it is sluiced to an ash plant. The ash then passes

into an ash crusher where it is crushed to nominally < 25 mm and passes into a large

mixing chamber. It is then pumped with centrifugal pumps as lean phase slurry to the

station ash dam. The fly ash is collected in hoppers under the precipitators or fabric

filters. In stations with precipitators, the hoppers are also used as storage and the fly ash

is removed routinely by water ejector and sluiced similarly to the bottom ash to the fly

ash plant. In stations with fabric filters the removal process is continuous using water

ejectors and sluiceways. In the fly ash plant the fly ash is sluiced in a mixing chamber

and pumped to the station ash dam as lean phase slurry using centrifugal pumps.

A station burning 7 × 106 t y

-1would therefore produce up to 315,000 t y

-1 of bottom ash

and up to 1.755 × 106 t y

-1 of fly ash. That is, each hour the power station needs to

dispose of up to 36 tons of bottom ash and up to 200 tons of fly ash.

12

At Eraring Power Station, bottom ash system has a basalt lined pipeline with an internal

diameter of 350 mm through which 1.242 m3

h-1

of slurry is pumped at a velocity of

3.58 m s-1

. Figure 2.2 is a photograph of the bottom ash pipeline at the disposal point.

Figure 2.2 Eraring Lean Phase Bottom Ash Disposal.

The station’s fly ash was pumped through a 450 mm inside diameter ferrocement

pipeline at a flowrate of 864 m3

h-1

at a velocity of 1.5 m s-1

. In normal operation, both

systems remain in service continuously requiring a return water system capable of

returning at least 2500 m3

h-1

of water from the ash dam back to the station. Bunn and

Gorsuch (1988) reported that the solids concentration of the Eraring bottom and fly ash

slurry system at power station full load was 3 % solids for the bottom ash system and 7

% solids for the fly ash system. Figure 2.3 is a photograph of the fly ash pipeline at the

disposal point.

13

Figure 2.3 Eraring Lean phase Fly Ash Disposal.

Singh (1991) reported that a comparison between a lean phase and a high concentration

system to remove similar tonnages of both bottom ash and fly ash. It was stated by the

author that in a traditional lean phase slurry disposal system, Cw < 15 % requires slurry

pumping plant capable of pumping 755 m3

h-1

of slurry at a minimum velocity > 3.5 m

s-1

through basalt lined pipeline where the internal diameter was 275 mm. However, a

high concentration system with a Cw of 67 % would require a slurry flow rate of 51.8 m3

h-1

at a velocity of 1.8 m s-1

through a 100 mm mild steel pipeline.

2.4 High Concentration Power Station Ash Disposal

During the 1980’s, interest was developing throughout the world in alternative disposal

systems for power station ash. In Australia, both the Electricity Commission of New

South Wales and Queensland Electricity Commission began research and development

projects to determine alternatives to the existing lean phase slurry system.

14

Testing to obtain the hydraulic transport characteristics of high concentration fly ash

slurries were conducted in South Africa, (Sive and Lazarus 1987). These tests were carried

out using centrifugal pumps for a range 20.0 % < Cw < 48.0 %. A closed pumping system

was used where the fly ash slurry was continually re-circulated through the system. The

tests used flow rates up to 360 m3 h

-1, with velocities up to 6.4 m s

-1. The authors

concluded that for a power station system using centrifugal pumps, the maximum safe

concentration of a slurry is Cv = 30.0 % (this corresponds to a Cw = 48.0 %).

Work was conducted by, (Verkerk 1987) on the hydraulic transport of ash slurries ranging

from a dilute mixture to a thick paste. The fly ash was obtained from Grootvlei Power

Station in the South Eastern Transvaal and bottom ash was obtained from the Kelvin

Power Station near Johannesburg. Two test facilities were utilised. One was a dilute a

slurry pipe loop test using a centrifugal pump and the other a dense phase pipe loop test

using a positive displacement pump. For these facilities the slurry was pumped around the

test loop a number of times at different concentrations. A major finding was that the

maximum Cw limit for pumping fly ash with a centrifugal pump was about 45.0 %. This

corresponds with the findings of, (Sive and Lazarus 1987), where a Cw = 48.0 % was

suggested. Verkerk (1987) also conducted test on fly ash slurries with a Cw varying from

66.3 % to 73.5 % pumped through a 120 mm internal diameter pipeline 125 meters in

length using a positive displacement pump. From the results, it was shown that there was a

gradual rise in pressure loss with increasing concentration up to Cw of 72.0 % where a

steep rise in pressure occurred.

The author observed that the rise in pressure loss was accompanied by a change in the fluid

properties of the slurry. The slurry changed from a fluid like character to one that tended to

form sliding planes at a Cw > 68.0 %. Above these concentrations the flow changed to a

plug flow with the pumping limit being in the region of a Cw of 74.0 %. Concurrently, a

series of tests was conducted where the slurry consisted of a mixture of fly ash and bottom

ash at differing ratios and varying concentrations. The fly ash to bottom ash mix gave

greatly reduced line pressure drops compared to fly ash slurry at similar Cw. This is a mix

at a ratio of fly ash to bottom ash of 60:40. The ratio of the production of fly ash to bottom

ash from a power station boiler is approximately 85:15. Figure 2.5 indicates the results of

the tests. The test indicated that the slurries follow a Bingham Plastic Rheological Model.

15

Figure 2.4 Comparison of Pressure Drop verses Flow for Fly Ash and Fly Ash/Bottom Ash

Slurries from Verkerk (1987).

Singh (1989) described a pilot plant installed at Queensland Electricity Commission

Bulimba Power Station investigated the feasibility of continuous mixing and pumping of

high density fly ash water slurries for the proposed Stanwell Power Station in Queensland.

The pilot plant has been designed to pump at a flowrate of 6 m3 h

-1 and consisted of a

mixing tank with stirrer, the fly ash was feed from a silo via a rotary feeder, a water

supply, constant speed positive displacement pump and a 140 metre pipeline with an

internal diameter of 38 mm. Plant control is from bubblier level devices in the mixing tank.

The bubblier level tubes of different length fitted to the mixing tank were used to measure

16

both tank level and slurry density. Density measurement was obtained from the head

difference between two tubes a set distance apart, level measurement by the head on the

longest tube modified to account for density. The slurry density measurement was used to

control the fly ash feed rate via rotary feeder and tank level was controlled by the water

control valve. The majority of these tests were conducted on a closed loop basis at

different Cw’s although some of the material was pumped to a disposal site to determine

placement characteristics. The results of the pilot plant investigation showed that it is

possible to mix and pump fly ash slurry continuously up to a Cw = 70.0 % on an

intermittent basis.

Singh (1989) concluded that the rheological properties of the fly ash slurries may be

influenced by the fly ash particle properties such as particle density and particle size

distribution. The pilot plant was moved to Gladstone Power Station in 1988 where the steel

pipeline length was extended to 900 metres of 38 mm internal diameter pipe with the last

50 metres plastic pipe and the pump output reduced to 4.5 m3 h

-1. On the day the plant was

inspected by the author, the fly ash slurry pumped had a density of 1600 kg m-3

which

corresponds to a Cw of 66.5 %. At a Cw of 66.5 %, the pipeline parameters had a pump

discharge pressure of 2 MPa, and a velocity of 1.1 m s-1.

As a result of this research the first high concentration slurry disposal system in Australia

was constructed at Stanwell Power Station, (Singh and Foley 1991). The power station at

Stanwell consists of 4 x 350 MW units burning Curragh coal delivered by rail from the

Bowel Basin 185 km away. The Stanwell high concentration plant consists of a unit

system where a mixture of bottom and fly ash from each boiler was pumped 2 km to the

disposal site through 4 x 100 mm steel pipes at a maximum flow rate of 50 m3 h

-1. The

successful operation of the Stanwell plant saw the installation of high concentration slurry

systems at other Queensland power stations. High concentration plants were installed at

new power plants at the Callide “C’, Tarong North, Millmerran and Kogan Creek and the

retrofitted at Callide “B”.

Concurrently in 1987, The Electricity Commission of New South Wales constructed a

pilot dense phase fly ash slurry system at Vales Point Power Station, (Bunn 1991). It

pumped as a dense phase slurry, a mixture of fly ash and water at Cw > 55.0 % in order to

17

determine the rheological characteristics of fly ash slurries. A complete description of the

pilot plant pumping trial at Vales Point using Vales Point fly ash is given in Chapter 5.

During the research at Vales Point, a pumping study was undertaken to determine the

rheological characteristics of Bayswater fly ash slurries prior to the design and construction

of the Bayswater dense phase ash slurry system, Bunn and Chambers (1995). 1270

tonnes of Bayswater fly ash was transported to Vales Point for the study. The study

indicated that for the Bayswater fly ash slurry pumped at a flow rate of 40 m3 h

-1 over a

distance of 1750 meters in a 150 mm nominal bore pipe the optimal Cw was 75 %. This

equates to a slurry shear rate (𝛾) of 34 s-1 and a shear stress (𝜏) of 25 Pa. The study also

indicated that the slurry pipeline could be shutdown full of slurry, left overnight and

restarted the next day without any problems. Therefore the design recommendation for the

Bayswater dense phase ash slurry system with a 10 km pipeline was for a fly ash flow

rate of 300 t h-1

, a pipeline nominal diameter of 200 mm and pipeline flow rate of 250

m3 h-1

would result in a nominal pipeline pressure drop of 5 MPa.

2.5 Bayswater Dense Phase Power Station Ash Disposal System Plant

Ward et al. (1998) described the Bayswater dense phase ash slurry system as consisting

of two parallel systems, each containing a silo, a mixing system, pumping system, slurry

pipeline, valve station and fully welded discharge pipelines. The philosophy was to

operate using one system at a time, leaving the other as a standby system. Dry fly ash

was discharged from the silo via a rotary feeder supplies into a conditioner at a

controlled rate of approximately 300 t h-1

. An impact weigher, located under the rotary

valve, measured the weight of fly ash. Water was added in the conditioner to produce

slurry with Cw of 85%. The amount of water added was determined by the desired

pumping concentration and the measured fly ash inflow rate. About 50% of the total

water required for pumping was added to the conditioner. The conditioned fly ash was

fed into a 58 m3 paddle mixer where further water was added to bring the slurry to the

desired Cw. A centrifugal slurry pump then supplied the slurry at the necessary pressure

to the Geho slurry pump suction. A sample loop, or consistency meter, ran parallel to

the booster pump. The consistency meter is a short pipe loop for which the differential

pressure was measured along with mass flow rate. This was installed to measure the

18

rheology of slurry for pipeline pressure control as it is being fed to the pipeline. The

slurry pumps were a triplex diaphragm positive displacement pumps capable of 9 MPa

at a flow rate of 250 m3 h-1

and up to a pressure of l4.5 MPa a flow rate of 180 m3 h

-1.

Variable speed 682 kW DC motors drove the pumps. The pressure and flow were

monitored at the inlet and outlet of the pipeline.

The two steel, 200 mm internal diameter slurry pipelines ran approximately 9.5 km to a

valve station at the Ravensworth No. 2 mine site. The pipeline was configured to

discharge into the disused Ravensworth open cut coal mine where temporary plastic

pipelines ran from the discharge valve station for a distance of up to 1 km into the

respective voids. The slurry pumps could be configured to either pipeline. A significant

amount of ground and surface water had collected in void 4 before the ash system was

commissioned. Water seepage from deposited slurry was expected to flow through into

the lowest void. Water was returned from this void to the mixing plant. Station water

from Lake Liddell or the Hunter River may be used as a supplementary or make-up

supply.

2.5.1 Operating Procedure

The operating procedure is as follows:

On establishment of a full silo of fly ash the pump start sequence is initiated;

The mixing process is started with water only, circulating through the booster

pump, sample loop and mixer tank;

The pipeline inlet and outlet valves are opened;

The main pump is started and ramped to a flow of 250 m3

h-1

;

On establishment of required flow the, rotary feeder begins operating to give fly

ash flow of 300 t h-1

of fly ash;

Pumping continues at Cw of 72% until the silo reaches a low level;

The slurry concentration is automatically reduced to Cw of 65% and the pipe

allowed filling with the lower concentration slurry;

The mixing plant and pumps are cleared of solids; and,

The slurry pump stops and the outlet valve closes.

19

The system remains shutdown until the main silo reaches a 'high level'. Should this take

longer than 24 hours the pipeline was to be flushed. Many operating parameters of the

dense phase slurry system are continuously monitored to ensure correct operation of the

plant. A pipeline blockage can be defined as any situation where the installed plant

cannot re-establish flow. This has failed to occur since the plant was commissioned in

1994. The risk of a blockage increases with an increasing Cw. A Cw of 72% has been

established as a safe operating concentration and the system is monitored and controlled

to prevent excursions above 72%. As the material properties of power station fly ash are

known to change with time, coal and operating conditions, the rheology of the slurry

will also change. Hence, a viscosity (or a consistency) meter to detect changes in slurry

rheology is part of the slurry preparation system. The meter consists of a small pipe

loop for which differential pressure and flow is continually monitored. For this

measurement to be of value the flow through the sample loop has to be controlled. The

continual monitoring of pipeline inlet and outlet pressures provides an excellent

indication of slurry rheology and an indication of impending problems. If the inlet

pressure rise above 8.2 MPa, the fly ash feed is restricted and Cw is reduced, above 9

MPa the fly ash flow ceases.

2.5.1.1 Bayswater Pipeline Rheology

It is known that as the apparent viscosity increases as the slurry concentration increases

and rapidly increase with a Cw > 75%. The measured pressure loss and flow rate allows

comparison of the present slurry rheology to that obtained during pilot plant operation.

The velocity and concentration range is restricted to 1.9 m s-1

< V < 2.3 m s-1

and 65% <

Cw < 73%. It takes approximately 1.5 hours for the slurry to travel the pipe length thus

comparisons are only based on a quasi-steady flow. With an average slurry flow of 250

m3 h

-1 and pipeline pressure of 4.5 MPa, the shear rate () is usually between 80 and 90

s-1

and the shear stress (𝜏 ) varies from 20 to 30 Pa giving a nominally pipeline apparent

viscosity ( 𝜇) of 230 m Pas. Physical tests have shown that it is possible to leave the

pipeline full of slurry with a Cw of 65% for up to 24 hours and still successfully restart

the plant. The pump is able to automatically ramp up in speed to 250 m3 h

-1 without any

major pressure increases.

20

On a site visit by the author on the 17th

March 2013, the following parameters were

observed: fly ash flow 260 m3 h

-1; water flow 115 m

3 h

-1; and, slurry pump flow 240 m

3

h-1

with a pipeline pressure of 6.8 MPa. This relates to a pipeline shear rate () of 85 s-1

,

a shear stress (𝜏) of 34 Pa and a pipeline apparent viscosity ( 𝜇) of 230 m Pas. The Cw

was calculated to be 69.3 %.

2.5.1.2 Ravensworth Ash Disposal Site

At the current rate of production in the Ravensworth mine it will take about 30 years to

fill the four voids with fly ash, (Ward and Bunn 1997). The aim is to achieve a final

landform that is free draining and minimises the earth works required to cap the ash.

After filling each void it will be capped with mine spoil pushed in from the side of the

void. For the first void a 400 mm capping layer has been placed over the ash.

Observation of slurry deposition behaviour in the voids indicates that after discharge the

slurry flow across, the ash surface was initially channelled but then fanned out into a

classic delta formation with small meandering flows across the surface. No water has

been observed to pond against the batter. The deposited ash also exhibits a relative fast

strength gain which allows rehabilitation to be commenced within 1-2 weeks of

cessation of ash deposition. A person can safely walk over an ash that was deposited

only some hours earlier.

2.5.1.3 Water Reclamation

The lowest void collects run off and seepage water from the other voids. The volume of

water mixed with the fly ash is approximately constant at 16,000 t w-1

. Water is

recycled back to the power station from the void at a rate of approximately 8,000 t w-1

the difference between water required for pumping and recycled water is supplied from

Lake Liddell. Measurements made on the deposited ash indicate that the Cw of the

deposited ash was 79%, indicating that 52% of the water used to transport the ash slurry

remained bound in the deposited as in the void. The use of a dense phase system to

transport fly ash from the Bayswater Power Station to the Ravensworth mine site, a

21

distance of about 11 km, has proven to be very successful. Thus any feasibility

assessment of a dense phase slurry disposal should not only consider the benefits of the

pumping operation, but also the benefits achieved in the disposal operation as well.

2.6 Callide B High Concentration Slurry Disposal Plant

Philips (2009) provides a description of the Callide B high concentration slurry disposal

(HCSD) plant. The HCSD plant was designed to handle an ash production rate of 100 t

h-1

, produce slurry between with a Cw between 50 % and 75 % and pump it 2.4 km, at

slurry flow rate of 100 m3 h

-1 with a pressure of 3.3 MPa, to the disposal site. Under

these conditions the plant typically cycles on a 50 % pumping to resting duty cycle with

a pumping cycle taking approximately six hours to complete.

The plant is configured into 2 x 100 % duty and standby HCSD trains. Each train

consisted of a nominal 500 m3 storage silo that typically filled to 85 per cent before the

process started and emptied to 5 % before the process shut down. The silo also

contained a series of fluidising nozzles critical to ensuring the process was fed a

consistent flow of product. Directly beneath the silo isolation valve was an ash flow

control valve used to carefully control the flow of ash into the process. The ash flow

entering the process was measured by a mass flow weigher installed after the ash flow

control valve. After the weigher, the ash went through an ash conditioner where it was

conditioned with controlled amounts of water so as to prevent the escape of fugitive

dust and to increase its hydroscopic nature, therefore improving its mixing qualities

with the slurry in the mixing tank. The conditioned product passed from the conditioner

to the mixing tank where the majority of the slurry mixing occurred. In the mixing tank

the slurry properties were controlled by modulation of the mixing tank water flow. The

slurry was finally drawn from the mixing tank into a piston diaphragm positive

displacement pump to be pumped to the disposal site via a Victaulic coupling joined

carbon steel pipe.

The earlier conversion of the Callide B furnace ash plant from sluiced wet impounded

hoppers to a dry ash collection system made it possible to store both furnace and fly ash

22

in the same silo. This resulted in significant capital savings and process design

simplification as conventional HCSD plants either do not handle furnace ash, or store it

wet in a separate silo requiring duplication of all silo infrastructure and control.

The Foley process control philosophy relies on controlling the slurry density to a fixed

set point to achieve the desired slurry viscosity, and for Stanwell Power Station with its

tightly controlled coal quality, this works well. At Callide, however, the coal quality can

vary significantly in the space of an hour, to the extent where this control philosophy is

no longer valid, resulting in unreliable operation and unacceptable slurry behaviour at

the deposit. The realisation that the fixed density set point control would not work for

Callide forced a rethink of how the slurry viscosity could be predicted in real time.

Bunn (2008) specified that the density control at Callide “B” Power Station be replaced

with a differential pressure control that calculated the pressure difference between a

pump discharge pressure transmitter and a pressure transmitter located 400 meters

downstream of the pump discharge transmitter.

Philips (2009) indicate that some effort went into researching feasible of using

viscosity transmitters that would be robust enough to survive the environment inside the

mixing tank, but none were found suitable. This prompted the search for methods to

approximate the slurry viscosity. This could be done relatively accurately and

repeatedly by discharge line pressure, given a constant discharge line velocity. Due to

excessive process lag observed in the measurement of discharge line pressure, it is not

suitable for control. A faster responding measurement was required to ensure process

stability. Discharge line differential pressure (ΔP) was found to be the perfect control

parameter. Process lag is still a minor issue for (ΔP) however; with the use of some

derivative action in the PID controller this was overcome. The process controlled by

discharge line (ΔP) has resulted in considerable increase in process control that has led

to substantial reductions in process down time and costs along with a massive increase

in the station’s ash storage capacity, enabling goals for station life to be met. In

achieving this, the process has become immune to variations in coal burned. Further

reflection on this success led to the realisation that the process is now insensitive to

variations in material properties, and as such, the technology is now transportable to

23

other power stations or similar slurry pumping facilities.

The Callide B process measures the differential pressure across the first 400 meters of

discharge pipe and compares this to a differential pressure set point. The mixing tank

water flow control valve is then modulated according to the error. Prior to this change,

the process could not maintain stability. Numerous other controls have been added to

the control system to maintain the process to within safe operating ranges, and to ensure

that the process starts up and shuts down as designed. It was concluded these

innovations have resulted in significant improvements in process capacity, stability and

reliability along with significant reductions in operating and maintenance costs over

comparable HCSD plants. The result is a maximisation of station ash storage capacity

for a minimum cost. Furthermore, the ability of this plant to deal with consistent

variations in material properties makes it compatible with nearly any ash, making the

technology transportable to other power station ash plants or similar materials handling

facilities.

2.7 Concluding Remarks

This chapter examines beginning of long distance hydraulic transport in the mid -19th

century through to the present day. In the modern power station built towards the end

of the 19th

century the norm was to pump station fly ash and bottom ash as lean phase

slurries. During the 1980’s, interest was developing throughout the world in alternative

disposal systems for power station ash. Numerous researches started investigating

hydraulic transport characteristics of high concentration fly ash slurries. The author was

responsible for the design and commissioning of several systems in New South Wales and

Queensland. As results of this work the author proposed this PHD to investigate the flow

of dense phase fly ash slurries in horizontal pipes and develop a new and original model

for determining pipeline pressure drop characteristics and a new method of

characterising typically homogeneous fluid behaviour based on a particle size

distribution, slope factor and a median particle size.

24

CHAPTER 3 RHEOMETRY AND RHEOLOGICAL MEASUREMENT

3.1 Introduction Slurry Rheology

Jinescu (1974) declared that suspensions of solid particles in a liquid medium exhibit a

wide range of rheological behaviours depending on particle concentration, size distribution

and shape and the characteristics of the suspension medium. As water is the original

Newtonian fluid, (Kambe 1969) stated that at a low Cw, solid particle slurries could behave

as a Newtonian fluid. As the concentration of solid particles increases, an interaction

between the particles and suspension medium can lead to non-Newtonian slurry

behaviours with a variable viscosity and even the existence of a yield stress. Non-

Newtonian slurries are plastic, pseudoplastic or dilatant. The two rheology categories of

slurries are either time-dependent or time-independent. For time-dependent slurries, the

flow properties are a function of time of shear as well as the shear history. Typical time-

dependent characteristics are thixotrophic where the viscosity decreases with time at a

constant shear, and rheopexy where viscosity increases with time at a constant fluid shear.

3.2.1 Time-independent slurries

3.2.1.1 Viscous Behaviour

Slurry is purely viscous if it readily flows like a liquid under the application of a shear

stress (𝜏). The shear stress at any point in slurry is a unique function of the shear rate

(𝛾) at that point. A generic equation to describe viscous slurries, (Bird et al. 1960) is:

𝜏 = 𝑓 ( ) (3-1)

𝜏 = 𝜇 ( ) (3-2)

with (𝜇) the coefficient of viscosity defined as a ratio between shear stress and shear rate.

25

3.2.1.2 Newtonian Behaviour

Newtonian behaviour is characterised by a constant viscosity independent of shear rate:

𝜏 = 𝜇 (3-3)

where the proportionality constant (𝜇) is the Newtonian viscosity. On a shear diagram

with linear coordinates, a plot of a Newtonian fluid would be linear and pass through the

origin as shown in Figure 3.1.

3.2.1.3 Pseudoplastic Slurries

Pseudoplastic slurries are slurries described by decreasing viscosity with increasing shear

rate (shear thinning). The shear curve for pseudoplastic behaviour was non-linear as

delineated in Figure 3.1. Thomas (1963) described pseudoplastic behaviour by means of a

power-law equation:

𝜏 = 𝑘′ 𝑛′ 𝑛′ < 1 (3-4)

The parameter (𝑘′) was defined as the "consistency index" and (𝑛′) as the "flow

behaviour index". For such a system, a higher (𝑘′) value implies that the slurry was more

viscous. The deviation of (𝑛′) from unity indicates pseudoplasticity that is more

pronounced and therefore, (𝑛′) = 1 corresponds to a Newtonian System.

Since (𝜇) was not constant in a pseudoplastic system, the value of (𝜇) was worthless

unless the shear rate was specified. To overcome this problem, the term "apparent

viscosity" (𝜇𝑎) was introduced to distinguish this viscosity from the Newtonian viscosity,

(Skelland 1967).

With the simple power law model, Equation (3-4) the range of shear rate where (𝑘′) and

(𝑛′) are constant is limited. To overcome this constraint has led to the development of

complicated power law models with extra constants. These models have been developed

26

Figure 3.1 Flow Curves for Time-Independent Viscous and Plastic Fluids from, (Skelland

1967).

by (Ellis, DeHaven, Prandtl-Eyring, Powell-Eyring, Reiner-Philippoff, Sisko, Symounds et

al., Spencer-Dillon, Williamson and Reiner-Rivlin to name but a few). A complete list of

models can be found in, (Skelland 1967). The model developed by (Symounds) was one

of the simpler:

𝜏𝑤 = 𝑎 (8𝑣

𝑑)

1−𝑘′

(3-5)

3.2.1.4 Dilatant Slurries

Dilatant slurries exhibit an increase in apparent viscosity with increasing shear rates. Shear

thickening defines this behaviour. Shear thickening is the opposite of pseudoplastic

behaviour. Figure 3.1 indicates a typical flow curve for dilatant slurries.

Sh

aer S

tress

(P

a)

Shear Rate (s-1)

Newtonian Dilatant Pseudoplastic

Bingham Yeild Dilatant Yeild Pseudoplastic

27

3.2.1.5 Plastic Behaviour

Bingham (1922), reported that some slurries exhibit plastic or visco-plastic behaviour, i.e.

they behave as solids at lower shear stresses but behaved like viscous fluids when a critical

shear stress was exceeded. Bingham developed a simple model for this characteristic

described as the Bingham Model Equation:

𝜏 = 𝜏𝑦 + 𝜇 ∶ (𝜏 ≥ 𝜏𝑦) (3-6)

The Bingham model predicts a linear relationship between shear stress and shear rate at

shear stress above( 𝜏𝑦), referred to as the Bingham yield stress. Figure 3.1 indicates a

typical flow curve for Bingham plastic fluids. The intercept of the flow curve at a zero

shear rate determines the yield stress.

3.2.1.6 Yield-Pseudoplastic and Yield-Dilatants Slurries

The majority of slurries observed in the real world do not follow the Bingham model but

possess a yield stress and non-linear behaviour, Wasp et al. (1977). These slurries exhibit

the flow behaviour as illustrated in Figure 3.1. Jinescu (1974) and Kambe (1969)

concluded that yield-pseudoplastic behaviour was more prevalent in real systems.

3.2.2 Time-Dependent Slurries

Skelland (1967) discussed the characteristics of certain slurries where the properties were

not only dependent on the shear history, but also on the period of shear. These were either

slurries where the apparent viscosity increases or decreases depending on the duration of

shear. When the apparent viscosity decreased with time of shear, the slurry was called

thixotrophic slurry. Alternatively, if the apparent viscosity increased with the time of shear

the slurry was called rheopectic slurry. Developing a shear curve for time-dependent slurry

is where the shear rate was constantly increased from zero to a maximum and then

decreased at the same rate to zero. This resulted in a hysteresis loop as exhibited in Figure

28

3.2. The structure of the hysteresis loop was dependent on the rate at which the shear rate

was increased and decreased as well as the shear history of the slurry. Skelland (1967)

suggested the difference between a thixotrophic or rheopectic slurries and pseudoplastic

slurries was the time element in structural breakdown, which was finite and measurable for

thixotrophic slurries and very small and undetectable for pseudoplastic slurries.

Figure 3.2 Flow Curves for Time-Dependent Fluids Demonstrating Hysteresis Loops from,

(Skelland 1967).

3.3 Introduction Rheometry and Rheological Measurement

Rheometry was defined by, (Harris 1972), as the experimental determination of the

mechanical properties of the matter. Although many types of viscometers are available,

most of them are unsuitable for defining practical flow properties. The most suitable types

of viscometers for determining the rheological properties of non-Newtonian slurries are the

capillary tube viscometer and the rotational viscometers, Skelland (1967). These

apparatuses readily determine the relationship between shear stress and shear rate.

Sh

ea

r S

tress

(P

a)

Shear Rate (s-1)

Thixotrophic Rising Thixotrophic Falling Rheopectic Falling Rheopectic Rising

29

Examples of viscometers that fall outside the above categories are the falling-ball

apparatus, rising-bubble viscometers, penetrometers, mobilometers, etc., Goh (1986).

Capillary tube and rotational viscometers are described in the following section.

3.3.1 Capillary Tube Viscometer

The purpose, of a capillary tube viscometer is to measure the frictional pressure drop

associated with a given flow rate of fluid through a long, cylindrical tube of known length

and diameter (Skelland 1967). Figure 3.3 demonstrates a schematic diagram of a typical

capillary tube viscometer.

3.3.2 Laminar Flow in Cylindrical Tubes

The assumptions see (Skelland 1976), made are:

a) The fluid was steady and fully developed in the laminar flow regime,

(b) The fluid was time-independent under the prevailing conditions,

(c) There was no slip between the fluid and the tube wall,

(d) The shear rate = − (𝑑𝑢

𝑑𝑟) = 𝑓(𝜏).

From basic fluid mechanics, the shear stress (𝜏) at any radius (r):

𝜏 = 𝑟

2 (

𝛥𝑃

𝐿) (3-7)

where (ΔP) is the pressure drop through the tube of length (L).

Therefore, (Wasp et al. 1977):

30

𝜏 = 𝜏𝑤 𝑟

𝑅 (3-8)

where (R) is the tube radius.

Figure 3.3 Schematic Diagram of a Typical Capillary Tube Viscometer.

For the shear stress at the wall (𝜏𝑤), (Wilkinson 1960), we have:

𝜏𝑤 = 𝐷𝛥𝑃

4𝐿 (3-9)

The shear rate at the tube wall may be determined for the general case from all the

previous assumption and (D) is the tube inner diameter, (Van Wazer 1963).

The expression for volumetric flow rate (𝑄), (Wasp 1977):

Controlled Air Pressure

L

D

31

𝑄 = ∫ 2 𝜋 𝑟 𝜇𝑅

0 𝑑𝑟 (3-10)

where 𝜇 = 𝜇(𝑟) is the velocity at radius (𝑟).

Integration by parts gives:

𝑄 = 2𝜋 [ 𝑟2 𝜇 𝑟 ]0𝑅 − 𝜋 ∫ 𝑟2 (

𝑑𝜇𝑟

𝑑𝑟) 𝑑𝑟

𝑅

0 (3-11)

Using assumptions (c) i.e. u(R) = 0 and (d), and transferring the variables in conjunction

with Equation (3-10), the variables are eliminated and the Equation (3-11) reduces to:

4𝑄

(𝜋 𝑅3)=

4

𝜏𝑤3 𝑑𝑟 ∫ 𝑟2𝜏𝑤

0𝑓𝜏 𝑑𝜏 (3-12)

In terms of the average velocity (𝑉):

8𝑉

𝐷=

4

𝜏𝑤3 ∫ 𝑟2𝜏𝑤

0 𝑓𝜏 𝑑𝜏 (3-13)

For a Newtonian fluid where 𝑓𝑡 = (𝜏

𝜇), substituting in Equation (3-13) gives:

𝜇 = 𝜏𝑤8𝑉

𝐷

(3-14)

For fluids with unknown rheology, the Weissenberg-Rabinowitsch equation Chambers el

al. (1986) can be used. It shows that the wall shear rate for a non-Newtonian fluid can be

calculated from the value for a Newtonian fluid having the same flow rate in the same

pipe, the correction factor being the quantity in the square brackets in equation (3-15).

This gives 𝑓(𝜏𝑤) = 𝑤 where (𝑛′) the slope of the 𝑙𝑜𝑔(𝜏𝑤) verses log () curve and is

the flow behaviour index:

= [3𝑛′+1

4𝑛′] (

8𝑉

𝐷) (3-15)

32

where 𝑛′ = (𝑑𝑙𝑛 𝜏𝑤)

(𝑑𝑙𝑛(8𝑉

𝐷))

(3-16)

The term (8𝑉

𝐷) is a unique function of wall shear stress (𝜏𝑤) and is a valid expression for

all time - independent fluids irrespective of their rheological model. Relating this term

back to Equation (3-9), (8𝑉

𝐷) is commonly referred to as the "apparent" shear rate and to

distinguish it from the Newtonian wall shear rate (), it is denoted by the symbol (𝛤𝑤),

(Goh 1986). When the values of wall shear stress (𝜏𝑤) and "apparent" shear rate (𝛤𝑤) are

plotted as a flow curve, this flow curve is referred to as pseudo - shear diagrams, (Wasp et

al. 1977).

The "apparent" shear rate (𝛤𝑤) can also be expressed as:

𝛤𝑤 = 32𝑊

(𝜋 𝜌𝑠𝑙 𝐷3) (3-17)

The plot of (𝜏𝑤) and (𝛤𝑤) describes the rheological behaviour of the slurry. The ratio of

(𝜏𝑤) and (𝛤𝑤) is an indication of the "apparent" viscosity 𝜇𝑎 of the slurry, (Skelland

1967). (Steward and Slatter 2009) indicated, in order to make use of standard

Newtonian theory, a value for the viscosity of the fluid is required. Usually the term

viscosity is meaningless once a non-Newtonian approach has been adopted. However,

an apparent viscosity can be defined at the pipe wall as:

𝜇𝑎 = 𝜏𝑤

𝛤𝑤 (3-18)

The Reynolds number may now be calculated using:

𝑅𝑒𝑁𝑒𝑤𝑡 = 𝜌𝑠𝑙 𝑉𝐷

𝜇𝑎 (3-19)

Note that (𝜇𝑎) is not a constant for a given fluid and pipe diameter, but must be

evaluated at a given value for (𝜏𝑤). The transition criterion is ReNewt = 2100.

33

3.3.3 Errors in Capillary Viscometry

When capillary viscometers are used for flow properties measurement the major sources of

error, (Van Wazer 1963), are:

(a) Kinetic energy losses: loss of effective pressure because of the energy in the issuing

stream;

(b) End effects: energy losses due to viscous or elastic behaviour when the slurry

converges or diverges at the ends of the capillary, and;

(c) Wall effects: surface phenomena at the fluid interface.

The kinetic energy losses end and wall effects may result in a dependency of the

experimental data on respective tube length and diameter. Kinetic energy losses can be

practically eliminated if the (𝐿

𝐷) ratio of the tube is greater than 100, Skelland (1967), and

the slurry efflux time as long as possible, (Van Wazer 1963), the convergence losses are

usually greater than the divergence losses, (Van Wazer 1963). The end effects can be

corrected by empirical methods, (Van Wazer 1963), provided data is obtained from several

different tube lengths. Wasp (1977) stated that end effects are rendered practically

negligible if the (𝐿

𝐷) ratio is greater than 100. Wall effects which always result in a

reduction of the apparent viscosity from the true value may be minimised by choosing a

large diameter tube relative to particle size. However, this must be balanced against the

advantage of high shear rates that can be achieved using small tube diameters. Thomas

(1963) and Seshadri (1970) reported that the effect of wall slip could be ignored if the ratio

of tube diameter (D) to the mean particle size (𝑑50), (𝐷

𝑑50) ratio is greater than 50 to 60.

3.3.4 Applications for Capillary Viscometers

Capillary viscometers are the most suitable for measuring the rheological properties of

time - independent slurries, (Van Wazer 1963). Capillary viscometers, on the other hand,

34

are not suitable for the measurement of rheological properties of low viscous liquids and

do not allow measurements of yield point and thixotropic structures, (Schramm 1981).

Plots of shear stress (𝐷 𝑑𝑝

4𝐿) and shear rate (

8𝑉

𝐷) and the corresponding quantities (𝐾′) and

(𝑛′) are, however, most conveniently obtained with capillary tube viscometers, (Skelland

1967). The highest ranges of shear rates are obtainable with capillary tube viscometers.

3.4 Concentric Cylinder Rotational Viscometers

The following sections contain a description of a coaxial cylindrical rotational viscometer.

3.4.1 Principle of Operation

Figure 3.6 is a schematic diagram of a Searle Type Concentric Cylinder Rotational

Viscometer. This type of system preselects the shear rate by the speed of the rotor and

measures the rotor torque enabling the shear stress to be calculated. The basic design

consists of an inner cylinder, rotor or bob with a radius (𝑅𝑖) rotating at a defined speed.

The outer cylinder or cup of radius (𝑅𝑎) is held stationary and filled with the material to be

tested. The rotor is immersed to a depth (𝐿) and the rotation of the rotor forces the slurry in

the annulus to flow.

The resistance of the slurry being sheared between the stationary and rotating boundaries

results in a viscosity related torque on the rotor which counteracts the torque provided by

the drive motor. A torque-sensing element is used as a direct measure of the viscosity for

the sample tested, (Schramm 1981).

35

Figure 3.4 Schematic Diagram of a Searle Type Concentric Cylinder Rotational

Viscometer.

To determine the shear stress (𝜏1) and the shear rate () at the rotor surface, the following

assumptions are required, (Van Wazer 1963):

(a) The fluid was steady and fully developed in the laminar flow region;

(b) There was no slippage at the surfaces of both cylinders;

(c) There are no end effects, shearing only takes place in the annulus, and;

(d) The shear rate = 𝑓(𝜏) only.

For a constant rotor speed, the torque (𝑇) required to rotate the rotor is given by:

𝑇 = 2 𝜋 𝐿 𝑟2 𝜏 (3-20)

Electric Motor

Bob

𝑅𝑖

𝑅𝑎

L

Torque Sensor

Cup

36

In terms of the shear stress at any radial coordinate (r):

r = 𝑇

2 𝜋 𝐿 𝑟2 (3-21)

The shear stress at the rotor surface (𝑟1) is:

𝑟1 = 𝑇

2 𝜋 𝐿 𝑟2 (3-22)

The shear rate () at position (r), where the fluid rotates at an angular velocity, is given by:

= 𝑓(𝑟) = −𝑟 (𝑑𝜔

𝑑𝑟) (3-23)

Integration of Equation (3-24), for r = Ri to Ra and ω = Ω to 0 gives:

Ω = − ∫ γ (dr

r)

Ra

Ri (3-24)

Substituting = 𝑓(𝑟) and replacing (𝑟) by (𝜏) with Equation (3-24), we get:

= (−1

2) ∫ 𝑓

𝜏𝑖

𝜏𝑎 (𝜏) (

𝑑𝑟

𝜏) (3-25)

When 𝛿 = 𝑅𝑎

𝑅𝑖 the shear stress of the cup (𝜏2) is:

𝜏2 = 𝛿−2 𝜏𝑖 (3-26)

The solution to Equation (3-26) depends on the exact form of 𝑓 (𝜏) i.e. the slurry model.

For unknown slurry, calculations of the shear rate at the rotor (1) require certain

approximations, (Skelland 1967). Krieger (1954), showed how to determine the shear

stress at the rotor by using infinite series. Complete descriptions of rotational viscometry

with various relations to determine the shear rate can be found in, (Skelland 1967) and

(Van Wazer 1963).

37

To allow comparison between the flow diagrams (pseudo - shear diagrams) measured for

capillary viscometers and the rheograms obtained for the rotary viscometer, the same

nomenclature will be used. The shear rate () or apparent shear rate (𝛤𝑤) will be

designated as (8𝑉

𝐷) and the shear stress (𝜏) as (

𝐷𝛥𝑃

4𝐿) regardless of the type of viscometer.

3.4.2 Sources of Errors in Rotary Viscometers

The major errors that can occur with a rotary viscometer are at low shear rates where the

whole gap is not sheared and settling of the slurry may occur. In the previous section, the

mathematical derivations assumed characteristics of an infinitely long rotor and no account

was taken of the drag on the ends of the rotor. Van Wazer (1963) discussed various

methods to overcome the errors due to end effects. One method discussed involved the

determination of a fictitious bob length by extrapolating a graphical plot of torque versus

rotor length. This fictitious rotor length can then be added to the actual bob length to obtain

the correct shear stress. In practical viscometry, selecting suitable cylinder geometrics will

minimise the end effects. Schramm (1981), states that if the gap size between the rotor and

cup was very small compared to the distance from the bottom of the cup to the rotor, then

this end effect becomes negligible. He suggested that a ratio of greater than 100 would be

satisfactory. The velocity gradient across the gap between the rotor and cup was

considered linear for the mathematical derivations. If this gap was large, the velocity

distribution is non-linear and this non-linearity will increase, the more non-Newtonian the

material becomes and the higher the shear rate. This non-linearity can be minimised by

keeping the gap between the rotor and the cup small. Introducing the concept of the ratio of

the radii gives:

𝛿 = 𝑅𝑎

𝑅𝑖 (3-27)

The DIN and International Standards have set the limit of this ratio of radii, (Schramm

1981):

1.00 ≤ 𝛿 ≤ 1.10

38

The further this ratio of radii stays below the upper value of 1.1 the more linear the

velocity distribution and the smaller the error.

3.4.3 Applications for Rotary Viscometers

The concentric cylinder rotational viscometers are a most versatile instrument for both

Newtonian and non-Newtonian materials. They can be used to determine both yield points

and thixotrophic structures, (Schramm 1981). Shear heating can be a problem at high shear

rates and therefore they are limited in the maximum shear rate. They can only be used for

laminar flows because of errors caused by Taylor Vortices at non-laminar flows. In

addition, they can only be used for fine material slurries and not for coarse material

because of the small clearances between the rotor and the cup.

3.5 Slurry Flow

Slurries can generally be classified into two categories: homogeneous and

heterogeneous. Homogeneous flow is a symmetric flow characterizing uniform

distribution of solids about the horizontal axis of the pipe. Durand and Condolios (1952)

published a number of studies indicating that homogeneous suspensions are those that

contain all the particles smaller than 40 μm while (Shook et al 2002) suggested that for

suspensions with a mean particle diameter (d50) greater than 50 μm, the slurry will

display heterogeneous properties. He also suggested that fine particles slurries d50 less

than 50 μm typically exhibit homogeneous fluid behaviour. Bunn (2013) indicated that

a fly ash slurry containing particles with a d50 less than 15 μm and a particle size

distribution (PSD) curve slope (𝑑𝑠) in the range of 3.0 to 3.1 will exhibit homogeneous

fluid behaviour. A measure of PSD curves slope (𝑑𝑠) was obtained as follows:

𝑑𝑠 =𝑑90 − 𝑑10

𝑑50 (3-28)

Fine particle slurries typically exhibit homogeneous fluid behaviour. Even though the

mixture consists of two distinct phases, these mixtures are treated as a continuum

39

possessing the density of the mixture. These types of slurries generally deviate from

Newtonian behaviour and exhibit non-Newtonian characteristics. Many models (i.e.

Power-Law, Bingham, and Casson) have been developed for slurries of this type and

their behaviour can be accurately predicted in laminar flows. Models that have been

developed more recently have considered turbulent pipe flow of non-Newtonian

slurries. Based on the experience of the Saskatoon Research Council Pipe Flow

Technology Centre (Gillies, 2006), the models of (Wilson and Thomas 1985) and

(Thomas and Wilson 1987) have been found to accurately represent the turbulent pipe

flow.

Heterogeneous slurries exhibit more complicated flow behaviours when compared to

homogeneous slurries. These slurries are typically a mixture of coarser particles in a

homogeneous carrier fluid. Due to the submerged weight and effects of gravity on the

coarse particles, sedimentation occurs within the flow. As a result, concentration and

velocity profiles across the flow domain are non-uniform and asymmetrical (Shook and

Roco, 1991). The behaviour of heterogeneous slurries in Newtonian carrier fluids in

turbulent flow has been well studied and numerous models and correlations exist to

predict slurry flow behaviour in pipelines (Shook et al., 1986; Gillies et al., 1991; Shook

and Roco, 1991; Gillies, 1993; Gillies and Shook, 1994; Matousek, 1997; Gillies et al.,

2000; Gillies and Shook, 2000; Shook and Sumner, 2001; Matousek, 2004; Gillies et al.

2004a; Sanders et al., 2004). Studies have investigated the minimum velocity required

to suspend all particles in a pipe flow (critical deposition velocity) and other flow

features including concentration and velocity profiles, and axial pressure gradient.

Regardless of the type of slurry, flows can be classified into two regimes: laminar and

turbulent.

Laminar flow is a low Reynolds Number phenomenon characterized by smooth,

streamline flow which is dominated by momentum diffusive effects as opposed to

convection. Turbulence is a state of fluid motion which is characterized by apparently

random and chaotic motions. It is a high Reynolds number phenomenon and it is a

departure from smooth, organized laminar flow to a chaotic, disorganized flow. A

limited amount of research has been conducted on the study of sand transportation in

laminar pipe flow with a Newtonian carrier fluid. Gillies et al. (1999) showed that

40

significant quantities of sand could be transported in laminar flow as long as the axial

pressure gradient was above a minimum value (approximately 2 kPa m-1

). Thomas et al.

(2004) have shown that the minimum axial pressure gradient principle also applies to

laminar flow of non-Newtonian slurries.

Earlier slurry flow theories suggested that turbulence or inertial effects were required to

support particles within heterogeneous flows. However, in recent studies, (Leighton and

Acrivos, 1987), it has been shown that under laminar flow conditions, viscous forces are

capable of resuspending settled particles. Although an understanding of laminar

transport of coarse solids in a non-Newtonian fluid is of importance to industry, few

studies have been conducted in this area. Laminar flows have the added benefits of

reduced fluid friction and pipe wear. However, if not operated correctly, they typically

result in the formation of a settled bed. An added concern is associated with the fact

that small changes in chemical properties can significantly increase the apparent

viscosity of non-Newtonian carrier fluids (Litzenberger and Sumner, 2004). This could

cause a heterogeneous slurry flow, which was initially operating in the turbulent regime,

to transition to the laminar flow regime.

3.6 Homogeneous Fluid Models

A fluid is a substance that undergoes continuous displacement as long as shearing forces

are applied to it, (Shook et al. 2002). Viscosity is a measure of the resistance of a fluid

to deform under shear stress. Viscosity describes a fluid's internal resistance to flow

(friction) and is a material property relating the shear stress (𝜏) and the time rate of

shear strain () in a moving fluid. Equation 3.29 below shows the relationship for these

parameters in a Newtonian fluid.

𝜏 = 𝜇 (3.29)

For Newtonian fluids, a constant, scalar parameter, (the dynamic viscosity) can be used

to relate the shear stress to the applied rate of shear strain. For Newtonian fluids, the

viscosity is independent of both shear stress and shear rate. The shear stress is a linear

41

function of the shear rate, with the slope of the curve being equal to the viscosity.

Figure 3.5 shows a graphical representation of the shear stress versus time rate of shear

strain behaviour plotted as a rheogram for a number of different fluid models.

Rheology is the study of the deformation of matter. The rheological behaviour of

homogeneous fluids can be described by a shear stress versus rate of shear strain

relationship:

𝜏 = 𝜇𝑎 (3.30)

Equation (3.31) represents the basic equation relating the time rate of deformation of a

fluid to an applied shear stress. In Equation (3.31), (𝜇𝑎) is the apparent viscosity of the

fluid.

Figure 3.5 Rheograms of various continuum fluid models.

For Newtonian mixtures, the apparent viscosity is equal to the fluid viscosity (𝜇𝑎 = 𝜇).

This is not the case for non-Newtonian fluids where η is a function of multiple

Shea

r St

ree

(Pa)

Shear Rate (s-1)

Newtonian Power Law Bingham Herschel Bulkley Casson

42

rheological parameters as well as (𝜏) and (𝛾) . For non-Newtonian fluids, more than a

single parameter is required to relate the shear stress to the applied rate of shear strain.

The constitutive model equations for selected non-Newtonian fluids are shown below,

(Bird et al. 1960) and (Shook and Roco 1991):

Power Law 𝜏 = 𝑘′ 𝑛′ (3.31)

Bingham 𝜏 = 𝜏𝑦 + 𝜇𝑎 (3.32)

Herschel Bulkley 𝜏 = 𝜏𝑦 + 𝐾𝑛 (3.33)

Casson 𝜏½ = 𝜏𝑐½ + (𝜇𝑐 )½ (3.34)

Of most interest in slurry flow applications are the behaviour of the fluids following the

models of Equations (3.32) to (3.34). All show the inclusion of a yield stress

term(𝜏𝑦 𝑜𝑟 𝜏𝑐). Fine particle suspensions, colloidal mixtures, and drilling mud typically

exhibit a yield stress, Litzenberger and Sumner (2004). In order for the fluid to flow, the

applied shear stress must exceed the yield stress. Once the applied shear stress exceeds

the yield stress, the rate of deformation of the fluid is determined by the difference

between the applied stress and the yield stress.

In this study, slurries which exhibit a yield stress will be represented with the Bingham

Rheological Model (3.32). The Bingham Model is the simplest of the rheological

models containing a yield stress. Unlike the Casson Model (3.34) and the Herschel-

Bulkley Model (3.33), the Bingham Model represents a linear relationship between the

shear stress and rate of shear strain. It is described by a yield stress (𝜏𝑦) and a plastic

viscosity (𝜇𝑎) which corresponds to the y-intercept and slope on the Bingham

Rheogram shown in Figure 3.5, respectively.

43

3.7 Rheology Studies of Fly Ash

In different parts of the world, fly ash slurries have been tested to determine the

pumpability characteristics. It is extremely difficult to compare the pumpability tests

obtained for a fly ash type from one region to another due to the variation in the fly ash

properties with time from a single power station and between different power stations.

Bunn et al. (1991), conducted rheological studies on fly ash obtained from three different

New South Wales power stations using a tube viscometer. The viscometer was designed

and manufactured at the University of Newcastle. This study was carried out with the fly

ash slurries stabilised before the testing proceeded. Stabilisation was achieved by mixing

the fly ash with water and allowing it to stand for 24 hours before it was re-mixed and

used. The study showed that the fly ash slurries with Cw in the range from 60 % to 70 %

from different power stations behave quite differently and are sensitive to changes in both

pH and temperature. Bunn el al (1991), the conclusion reached was "that the slurries were

non-Newtonian at their equilibrium state, and that the experimental data fitted both the

Bingham and Symonds models reasonably well".

Rheological studies were conducted by, (Lazarus and Sive 1984), on fly ash obtained from

a South African coal fired power station using a Balanced Beam Tube Viscometer

(BBTV). These viscometer results were then scaled up and compared with data obtained

from the University of Cape Town test system pumping the same fly ash. The fly ash used

was well graded with a particle size d50 = 17 μm and a solids relative density of 2.23. The

rheological study was conducted on slurries with a Cv up to a maximum of 30 %. A

relative density of 2.23 with a Cv = 30 % gives a Cw = 48 %. Lazarus and Sive, (1984)

concluded "that the Balanced Beam Tube Viscometer was a useful tool for the prediction

of reproducible rheograms for non-Newtonian slurries at these concentrations".

A study of the rheological characteristics of fly ash was conducted by (Verkerk, 1987). He

used a Brookfield Rotary Viscometer to determine the apparent viscosity at different Cw

with different ash types. He concluded that a Brookfield Rotary Viscometer was not a

good instrument to use to determine rheology of non-Newtonian slurries. The rotary

viscometer speed was used to determine shear rate and the torque used to determine the

44

shear stress. The rotary viscometer was also used to determine the critical diameter size of

the split between the vehicle and homogeneous portions of the slurry. Verkerk concluded

from these viscosity tests "that the particle size distribution and Cw influences apparent

viscosity. At higher Cw, a generalised Bingham fluid behaviour was indicated".

A comparative study was conducted on kaolin clay slurries and uranium tailing slurries of

different Cw by, (Lazarus and Slatter 1986), using the BBTV and a rotary viscometer. A

range of tests were carried out using different Cw and particle size distributions. The

authors concluded "that the rotary viscometer was not suitable for the rheological

characteristics of the slurries tested; whereas the BBTV was capable of correctly

characterising the slurries tested. Lazarus and Slatter reason the rotary viscometer was

declared not suitable was because the torque reading decayed with time in an exponential

fashion. The "correct" reading was therefore ambiguous".

3.8 Flow Cones

A flow cone is a simple device for measuring viscosity of a grout by observing the time

it takes a known volume of liquid to flow from a cone through a short tube. Grout is a

mixture of fly ash, cement and water.

A Marsh Cone has a working volume of 1.5 litres. It consists of a cone 152 mm across

and 305 mm in height to the apex to which is fixed a tube 50.8 mm long and 12.7 mm

internal diameter. A 2 mm screen is fixed near the top across half the cone. An ASTM

Cone has a working volume of 1725 ml. It consists of a cone to which a tube with is

attached. The cone is 178 mm across and 190 mm in height to the apex to which is fixed

to a tube 38.1 mm long and 12.7 mm in diameter. A 5 mm pointed rod is used as a level

indicator. Figure 3.6 is a photograph of a Marsh and an ASTM cone and Figure 3.7 are

dimensioned drawings of a Marsh Funnel and an ASTM Cone. A Marsh Funnel is a

Marsh Cone made according to American Petroleum Institute (API) specification 13 B.

The API specification for the discharge of one litre of water at 21° Celsius is 26.0 ± 0.5

seconds with a tube of 4.75 mm. Calibration of the flow cones is similar except different

volumes of water are added. The test procedure for the ASTM cone is set out in ASTM

C939 – 10 Standard Test Method for Flow of Grout for Preplaced-Aggregate Concrete

(Flow Cone Method).

http://en.wikipedia.org/wiki/Viscosity

45

The method as described in ASTM C939, first the cone is moistened by filling the cone

with water one minute before introducing the test water allow the water to drain from

the cone.

The method as described in ASTM C939, first the cone is moistened by filling the cone

with water one minute before introducing the test water allow the water to drain from

the cone. The calibration procedure of the level indicator is undertaken by blocking the

outlet tube with a figure or stopper and adding 1725 ± 5 ml of water at a temperature 23

± 2.0° Celsius and after all turbulence has stopped check that the gauge point is just

touching the water surface adjust if required. Start the stop watch, and simultaneously

remove the finger or stopper. Stop the watch at the first break in the continuous flow of

water from the discharge tube. The time indicated by the stop watch is the time of efflux

of water. If this time is 8.0 ± 0.2 seconds, the cone may be used for determining the time

of efflux of slurry grout.

Figure 3.6 Photograph of a Marsh and ASTM Cones.

Marsh Cone ASTM Cone

46

Figure 3.7 Marsh Funnel from (Pitt 2000) and ASTM Flow Cone from (ASTM C939).

The procedure for testing grout is the same as testing with water the only difference is

that when the stop watch is stopped, look into the top of the cone; if the grout has

passed sufficiently, such that light is visible through the discharge tube, the time

indicated by the stop watch is the time of efflux of the grout. If light is not visible

through the discharge tube, then the use of the flow cone is not applicable for grout of

this consistency. At least two tests having times of efflux within 1.8 s of their average

shall be made for each grout mixture. The Marsh cone calibration is similar ASTM cone

the only variation is that the water is filled until it touches the mesh screen 1.5 litres

equivalent to a level of 279.4 mm. The efflux time should be 6.4 ± 0.2 seconds.

3.8.1 Flow Cones as Rheological Devices

Pitt (2000) used a Marsh Funnel and a Weissenberg Rheogoniometer, model R 16 to

conduct comparative studies of glycerol as a Newtonian fluid and solutions of two

polymers commonly used in drilling fluids, namely, xanthan cellulose (XC) and

hydroxyethylcellulose (HEC). The results of the comparative study enabling the

47

collection time from a Marsh Funnel to be converted into a value for effective viscosity

of non-Newtonian fluids. Pitt suggested that the for field use, the following equation

relates the effective viscosity (𝜇𝑒 )of a Marsh Funnel and collection time:

𝜇𝑒 = 𝜌𝑠𝑙 (𝑡−26) (3.35)

where 𝜌𝑠𝑙 = fluid density,

𝑡 = efflux of one litre of fluid.

Pitt stated in principle, a pair of funnels could be made to give measurements at two

shear rates, for example, 1,000 and 2,000 s-1

. That is at two different Funnel flow times

to give a Bingham plastic viscosity, but there is little point since multispeed Rheometry

is readily available.

Le Roy and Roussel (2004) investigated the possibility of using the Marsh Cone as a

viscometer. Rheological measurements using a coaxial Haake Viscotester VT 550 along

with digital image recording of Marsh Cone flow on glycerol-water mixes were carried

out. The equations needed to solve the flow problem were for purely Newtonian viscous

fluids. They showed that flow time can be directly linked to the Newtonian viscosity.

Flow time was proportional to viscosity. The Marsh Cone and rotary viscometer was

then used to test several cement pastes and measured flow time was compared to predict

flow time. They stated that the correlation between flow time and cement pastes

apparent viscosity stays valid only for no yield stress cement pastes and for flow time

higher than about 15 s. They also showed that the Marsh Cone test and the associated

proposed calculation are relevant only when the grout behaviour is close to a Newtonian

behaviour.

Roussel and Le Roy (2005) presented a study concerned with the Marsh Cone that is a

workability test used for specification and quality control of cement pastes and grouts.

Roussel and Le Roy indicated that cone standard vary from one country to another, but

its operating principles are the same. That is time needed for a certain amount of

material to flow out of the cone was recorded and the measured flow time was linked

with the so-called “fluidity” of the tested material. The longer the flow time, the lower

48

the fluidity. The flow time depends on the tested fluid and by the cone geometry. It was

demonstrated that, under several consistency and geometry conditions, the flow time

reflecting “fluidity” may be calculated from the plastic viscosity and yield stress in the

case of a Bingham fluid and specific cone geometry.

The study showed that the flow time can be directly linked to the material behaviour,

namely, parameters such as yield stress and plastic viscosity for Bingham fluids.

Equations needed to solve the flow problem can be derived in the case of a Bingham

fluid, which is a common and simple approximation of a fresh cement based material

behaviour. The correlation between the flow time and the rheological behaviour of the

cement pastes was experimentally validated using the flow cone and a rotary

viscometer. Finally, a method using two cones differing by their nozzles size only was

presented. This method allowed the determination of the two behaviour parameters from

the results of the two different Marsh Cone tests. Various empirical and theoretical

models have been used to describe fresh cement pastes behaviour. Among the most

widely used are the Bingham and Herschel-Bulkley models, which take into account the

pseudoplastic behaviour of concentrated suspensions. Roussel and Le Roy stated that if

the yield stress was very small, as it is the case for cement grouts, then a purely viscous

model, which is a particular case of a Bingham model, was often sufficient to describe

correctly the grouts fresh behaviour.

Balhoff et al. (2011) indicated that accurate and simple techniques for measurement of

fluid rheological properties were important for field operations in the oil industry, but

existing methods are relatively expensive and the results can be subjective. This is

particularly true for measurements of fluid yield stress which are notoriously difficult to

obtain. Marsh Funnels are popular quality-control tools used in the field for drilling

fluids and they offer a simple, practical alternative to viscosity measurement. In the

normal measurements, a single point (drainage time) is used to determine an average

viscosity; little additional information is extracted regarding the non-Newtonian

behaviour of the fluid. The authors presented a new model that could be used to

determine the rheological properties of drilling mud and other non-Newtonian fluids

using data of fluid volume collected from a Marsh Funnel as a function of time. The

funnel results for viscosity and shear-thinning index compare favourably to the values

49

obtained from a commonly used Fann 35 rotary viscometer. More importantly, an

objective, static method for determining yield stress was introduced which has several

advantages over dynamic, extrapolation techniques used for rheometry data. A steady

state height is measured and the yield stress (𝜏0) is calculated:

𝜏0 = 𝜌𝑔 (ℎ𝑠𝑠+𝐿)

(2𝐿

𝑅+

2𝐻𝐹𝑅𝐹

) (3.36)

where 𝜌 = fluid density (cm s-2

),

𝑔 = acceleration due to gravity (cm s-2

),

ℎ𝑠𝑠 = steady state height in the funnel (cm)

𝐿 = length of the capillary tube (cm)

𝑅 = radius of capillary tube (cm)

𝐻𝐹 = total height of the cone portion of the funnel (cm)

𝑅𝐹 = maximum radius of the funnel (cm)

The shear stress (𝜏𝑤)at the tube wall,

𝜏𝑤 = [𝜌𝑔(ℎ+𝐿) ]𝑅

2𝐿−

2𝐻𝐹

𝑅𝐹 𝜏0 (3.37)

The viscosity (𝜇) of the fluid can be estimated for laminar flow,

𝜇 = ⌊𝑅4

8𝐿 (

𝐻𝐹

𝑅𝐹)

2

⌋ ⌊𝜌𝑔

𝑙𝐻0− ½ℎ02

⌋ 𝑡𝑓 (3.38)

where ℎ0 = initial height in the funnel (cm)

ℎ𝑠𝑠 = total drainage time (s)

𝑡𝑓 = total drainage time (s)

50

CHAPTER 4 EMPIRICAL APPROACH

4.1 Introduction

Many industrial slurries exhibit non-Newtonian behaviours. The first step in the

empirical approach is to calculate the critical velocity. It is the minimum velocity

required to maintain all solid particles in a suspension condition. However, if the

velocity is less than the critical flow velocity, solid particles will be deposited,

(Abulnaga 2002). Many investigators such as (Durand 1953), (Wasp et al. 1977),

(Kokpinar and Gogus 2001), have proposed equations for the estimation of the critical

flow velocity.

4.1 Estimation of Critical Velocity

Equations for the estimation of the critical velocity (𝑉𝑐) of flow through pipe-line were

derived empirically in the literature in the terms of fluid, flow and solid particle

characteristics. Based on the experiments performed in pipes of diameter, 𝐷 = 0.04 to

0.58 m for coal and sand of diameter, 𝑑50 = 0.44 to 2.04 mm, with volumetric

concentrations (𝐶𝑣) from 5% to 15%, (Durand 1953) proposed the following equation

for the critical velocity:

𝑉𝑐 = 𝐹𝐷𝐿 √2𝑔𝐷 (𝜌𝑠 − 1) (4.1)

where 𝐹𝐷𝐿 = Durand velocity factor;

𝑔 = acceleration due to gravity (m s-1

);

𝐷 = pipe diameter (m);

𝜌𝑠 = specific gravity of solids.

Durand’s velocity factor (𝐹𝐷𝐿 ) depended on the particle diameter and the concentration

of solids. It can be obtained from the graph provided by (Warman 2000). Figure 4.1 is a

copy of the graph.

51

Figure 4.1 Durand limiting Settling Velocity Factor Warman (2000).

Schiller and Herbich proposed the following equation for Durand velocity factor based

on the 𝑑50 of the particles, (Ahmad and Azamathulla 2012);

𝐹𝐷𝐿 = 1.3 𝐶𝑣0.125 [1 − exp(−6.9𝑑50)] (4.2)

Factor FDL

Particle Size (µm)

52

where 𝐶𝑣 = solids concentration by volume (%);

𝑑50 = mean particle size (µm).

Oroskar and Turian (2008), performed regression analysis of the available data and

proposed;

𝑉𝑐 = 𝜐−0.09𝑑0.17(𝜌𝑠)𝑜.55𝐷0.47 (4.3)

where 𝜐 = kinematic viscosity of liquid (m2 s

-1);

𝑑 = particle diameter (mm);

Zandi and Gavatos extended the work of Durand to other solids (1.02 ≤ s ≤ 2.65) and to

different mixtures, as indicated by (Ahmad and Azamathulla 2012). Based on their

analysis of the test data for sand of particle size ranging from 0.0002 m to 0.0254 m, in

pipes with a diameter from 0.0381 m to 0.56 m, and volumetric concentration (𝐶𝑣) up to

22%, they defined an index number (N), (Ahmad and Azamathulla 2012).

𝑁 = 𝑉2 𝐶𝐷

0.5

𝐶𝑣 𝐷𝑔( 𝑆−1) (4.4)

where 𝑉 = average velocity of flow (m2 s

-1);

𝐶𝐷 = drag coefficient;

At the critical value N = 40, the transition between saltation and heterogeneous flow

condition of particles occurs. The critical velocity is;

𝑉𝑐 = ([40𝐶𝑣 𝐷𝑔(𝑠−1)

√𝐶𝐷 ])

0.5

(4.5)

The drag coefficient (𝐶𝐷 ) depends on the Reynolds number and Albertson shape factor

and can be obtained from the table provided by Zandi and Gavatos, (Ahmad and

Azamathulla 2012).

Babcock showed that for finely graded particles, the transition occurred at an index

53

number 10 and proposed the following expression for the critical velocity, (Ahmad and

Azamathulla 2012).

𝑉𝑐 = ([10𝐶𝑣 𝐷𝑔(𝑠−1)

√𝐶𝐷 ])

0.5

(4.6)

Wasp et al. (1977) modified the Durand equation and included the effect of particle

concentration (𝐶𝑣), ratio of the particle diameter (𝑑) and pipeline diameter (𝐷) and

proposed;

𝑉𝑐 = 3.39 𝐶𝑣0.2156 (

𝑑50

𝐷) √2𝑔𝐷(𝑠 − 1) (4.7)

This equation is valid in the range of data: 1.02 ≤ 𝜌𝑠𝑙 ≤ 2.65; and 0.0106 ≤ 𝑑50

𝐷 ≤

0.0356. Turian et al. (1987) showed that for slurries composed of large non-colloidal

particles, the critical velocity (𝑉𝑐) is virtually independent on the particle size.

Further, (Kokpinar and Gogus 2001), conducted seven series of experiments to obtain

the critical velocity by observing solid particle in a glass section of pipeline and

proposed;

𝑉𝑐

√𝑔𝐷= 0.055 (

𝑑50

𝐷)

0.537

𝐶𝑣0.27(𝑠 − 1)0.07 (

𝜌𝑠𝑙𝑑50

𝜇𝑓)

0.3

(4.8)

where 𝜔𝑚 = viscosity of slurry (Pa s);

𝜇𝑓 = viscosity of carrier fluid (Pa s); and,

𝜌 = density of the fluid (kg m-3

);

Hepy et al. found that the effect of some of the parameters on critical velocity was

negligible and reduced the equation, (Ahmad and Azamathulla 2012);

𝑉𝑐

√𝑔𝐷= 𝛷3 [

𝑑50

𝐷(𝜌𝑠 − 1)𝐶𝑣 (

𝜌𝜔𝑚𝑑50

𝜇𝑓)] (4.9)

54

The following Table 5.1 contains a selection of equation for critical velocity.

Table 5.1 Critical Velocity Equations.

Investigator Equation

Durand (1953) 𝑉𝑐 = 𝐹𝐷𝐿 √2𝑔𝐷 (𝜌𝑠 − 1)

Hungmark (1961) 𝑉𝑐

√𝑔𝐷= 𝛷[𝐶𝑣 (𝜌𝑠 − 1)𝐹𝐷]

Zandi and Gavatos (1967) 𝑉𝑐 = 0.6 √𝑔𝐷 (

𝑤2

𝑔𝑑)

0.25

Rose and Duckworth (1969) 𝑉𝑐 = 10.24𝑤2 𝐶𝑤

𝑜.4 (𝐷

𝑑50)

1.2

𝜌𝑠−1.4 (𝑔𝐷)−0.5

Babcock (1971) 𝑉𝑐 = ([

10𝐶𝑣 𝐷𝑔(𝑠 − 1)

√𝐶𝐷 ])

0.5

Wasp et al. (1977) 𝑉𝑐 = 3.39 𝐶𝑣

0.2156 (𝑑50

𝐷) √2𝑔𝐷(𝜌𝑠 − 1)

Gogus and Kokpınar (1993) 𝑉𝑐

√𝑔𝐷= 0.124 (

𝐷

𝑑50)

0.537

𝐶𝑣0.322(𝜌𝑠 − 1)0,121 (

𝑤𝑑50

𝑣𝑤)

0.234

Kokpınar and Gogus (2001) 𝑉𝑐

√𝑔𝐷= 0.055 (

𝑑50

𝐷)

0.537

𝐶𝑣0.27(𝜌𝑠 − 1)0.07 (

𝜌𝑙𝑤𝑠𝑑50

𝜇𝑓)

0.3

Hepy (2008) 𝑉𝑐

√𝑔𝐷= 𝛷3 [

𝑑50

𝐷(𝜌𝑠 − 1)𝐶𝑣

𝜌𝑙 𝜔𝑚 𝑑50

µ𝑓

]

Another common equation for the determination of critical velocity (𝑉𝑐) used by the

engineering world is the method proposed by, (Kokpinar and Gogus 2001); this method

requires determination of rheological properties whereas the method proposed by

Durand does not:

𝑉𝑐

√𝑔𝐷= 0.055 (

𝑑50

𝐷)

0.537

𝐶𝑣0.27(𝜌𝑠 − 1)0.07 (

𝜌𝑙𝑤𝑠𝑑50

𝜇𝑓)

0.3

(4.10)

Even after development of numerous new critical velocity equations, the engineering

world still relies on the Durand method. Using the Durand method the design engineer,

after establishing the desired flowrate of the slurry and the commercially available

pipeline diameter, ensures that the pipeline velocity is 30 % greater than the calculated

critical velocity.

55

4.2 Determining Pipeline Pressure Drop – Head Loss

The continuity equation for discharge flow (𝑄) is;

𝑄 = 𝜋

4 𝐷2 𝑉 (4.11)

The hydraulic design and analysis of flow conditions of non-Newtonian fluids depend

upon predicting the friction factor (𝑓). It is extremely difficult to arrive at an exact

analytical solution to calculate the friction factor associated with the flow of non-

Newtonian fluids, and therefore, explicit approximations are often used. The friction

factor can be determined from a friction factor chart (Moody diagram). A Moody

diagram for laminar and turbulent flow conditions of Bingham plastic fluids is

presented in Figure 4.2 from, Swamee and Aggarwal (2011).

A friction factor is determined, by firstly, calculating the Reynolds Number (𝑅𝑒) for the

particular slurry:

𝑅𝑒 = 𝜌𝑉𝐷

𝜇 (4.12)

Then, the pipe relative roughness (𝜀

𝐷) is calculated from the absolute roughness (𝜖).

The absolute roughness (𝜖) is obtained from a publish data as shown in Table 5.2. Once

the friction factor (𝑓) has been calculated the head loss (ℎ𝑓) can be determined for a

given flow by the Darcy - Weisbach equation, (Swamee and Aggarwal 2011);

ℎ𝑓 = 𝑓𝐿𝑉2

2𝑔𝐷 (4.13)

56

Figure 4.2 Moody Diagram for Bingham Plastic Fluids Swamee and Aggarwal (2011).

The friction slope (𝑆𝑓) can be determined by combining equations (4-11) and (4-13);

𝑆𝑓 = 8𝑓𝑄2

𝜋2𝑔 𝐷5 (4.14)

The head loss can also be determined from;

𝑆𝑓 = ℎ𝑓

𝐿 (4.15)

where 𝐿 = pipeline length (m).

For laminar flow of Bingham plastic fluids, (𝑓) can also be calculated from the

Buckingham–Reiner equation, Swamee and Aggarwal (2011);

𝑓 =64

𝑅𝑒 [1 +

𝐻𝑒

64 𝑅𝑒−

64

3 (

𝐻𝑒4

𝑓3 𝑅𝑒7)] (4.16)

57

Where, (𝑅𝑒) and (𝐻𝑒) are the Reynolds number and the Hedstrom number respectively

defined as;

Table 5.2 Absolute Roughness of Common Materials from

(http://neutrium.net/fluid_flow/absolute-roughness/).

Pipe Material Roughness (mm)

Drawn Tubing, Glass, Plastic

0.0015-0.01

Drawn Brass, Copper, Stainless Steel (New) >0.0015-0.01

Flexible Rubber Tubing - Smooth 0.006-0.07

Flexible Rubber Tubing - Wire Reinforced 0.3-4

Stainless Steel 0.03

Wrought Iron (New) 0.045

Carbon Steel (New) 0.02-0.05

Carbon Steel (Slightly Corroded) 0.05-0.15

Carbon Steel (Moderately Corroded) 0.15-1

Carbon Steel (Badly Corroded) 1 - 3

Asphalted Cast Iron 0.1 - 1

Cast Iron (New) 0.25

Cast Iron (Old, Sandblasted) 1

Sheet Metal Ducts (with smooth joints) 0.02 – 0.1

Galvanized Iron 0.025-0.15

Wood Stave 0.18-0.91

Wood Stave (Used) 0.25-1

Smooth Cement 0.5

Concrete – Very Smooth 0.025-0.2

Concrete – Fine (Floated, Brushed) 0.2 – 0.8

Concrete – Rough, Form Marks 0.8 - 3

Riveted Steel 0.91 – 9.1

Water Mains with Tuberculation’s 1.2

Brickwork, Mature Foul Sewers 3

𝑅𝑒 = 𝐷𝑉

𝜐 (4.17)

𝐻𝑒 = 𝐷2 𝑠𝑜

𝜐2 (4.18)

http://neutrium.net/fluid_flow/absolute-roughness/

58

The parameter (𝑠𝑜) is determined;

𝑠𝑜 = 𝜏𝑜

𝜌𝑠𝑙 (4.19)

where 𝜏0 = yield shear stress (s-1

);

𝜌𝑠𝑙 = density of slurry (kg m-3

);

The yield shear stress (𝜏0) is related to the following rheological equation;

𝜏 = 𝜏𝑜 + 𝜌𝑠𝑙 𝜐∞ (4.20)

A method for determining friction factor directly was proposed by, Swamee and

Aggarwal (2011), for Bingham plastic slurry when combined with the Darcy –

Weisbach equation;

𝑓 =64

𝑅𝑒 +

10.67+0.1414(𝐻𝑒𝑅𝑒

)1.143

[1+0.0149 (𝐻𝑒𝑅𝑒

)1.16

]𝑅𝑒

(𝐻𝑒

𝑅𝑒) (4.21)

No attempt has been made in this thesis to determine critical velocity and pipeline

pressure drops using the empirical approach, as the recent empirical approach has

involved the addition of rheological properties. Therefor if you have to determine

rheological properties why not use a total rheological approach?

59

CHAPTER 5 PREVIOUS RESEARCH

5.1 Introduction

This chapter is a description of my research thesis tilted “The Dense Phase Hydraulic

Conveying of Power Station Ash", submitted in 1991 for a Master of Engineering,

Department of Mechanical Engineering, University of Newcastle.

In 1987, The Electricity Commission s of New South Wales constructed a pilot dense

phase fly ash slurry system at Vales Point Power Station, Bunn (1991). The intention was

to pump, as a dense phase slurry, a mixture of fly ash and water at Cw > 55.0 % to

determine the slurry’s rheological characteristics. It was the intention to mix bottom ash

with the fly ash and water in the proportion of fly ash to bottom ash up a maximum of 9 to

1 to determine the change in the rheological characteristics at similar concentrations. The

author was given a brief by the Electricity Commission of New South Wales in January

1987 to design, construct, commission and obtain operation data from the dense phase ash

pumping plant (DPAS). The constraints were that the pump had to be a positive

displacement one, with no inlet or outlet valves and all the plant equipment had to be

commercially available.

5.2 Vales Point Dense Phase Ash Pumping Plant

The dense phase ash pumping (DPAS) pilot plant was located adjacent to the Amatek

Facility at Vales Point Power Station. The pilot plant received a supply of fly ash that was

extracted from the station precipitators and was for sale commercially to the cement

industry. Figure 5.1 is a photograph of the pilot plant mixing and pumping system prior to

connection to the fly ash feeder and pipeline.

60

Figure 5.1 Pilot Plant Mixing and Pumping System.

The fly ash path to the mixer consisted of an air slide from a silo within the Amatek

Facility, a surge bin, a mass flow-meter and a rotary feeder. On the bottom ash side, the

plant consisted of a loading hopper, a bucket elevator, a storage bin, a rotary feeder, screw

conveyors and a mass flow-meter. Both the fly ash and bottom ash systems fed directly

into the mixer while the mixer outlet connected directly to the Putzmeister pump. The

pump outlet was connected to a steel pipeline. The mixer was a B.H.S., horizontal, twin

shaft, counter rotating, constant flow mixer, and model LFE 520 x 2600 with a maximum

capacity of 130 t h-1

. The mixer shafts were driven at a speed of 100 rpm by an 18.5 kW, 3

phase, and 415 volt electric motor via a worm drive gearbox. Each shaft had 14 pairs of

mixing paddles with each pair of paddles mounted opposite each other and the next pair

mounted at 90 degrees to the previous pair. The orientation of the paddles caused the fly

ash slurry to move towards the mixer outlet. Water was distributed into the mixer from two

headers at the top of the mixer via 24 spray nozzles which had a total capacity of 42 t h-1

.

A by-pass valve was connected directly to the mixer allowing a maximum water flow rate

of 50 t h-1

for mixer flushing.

The pilot plant pump was a Putzmeister, horizontal, twin cylinder, single acting piston

DPAS Pump

BHS Mixer Fly Ash Mass Flow-meter

Mixer Outlet Hopper

61

pump, model KOS 1460 equipped with an electro hydraulic drive. The pump, had no input

or output valves. The output of the twin cylinders was transferred to the discharge via a

hydraulically actuated "S transfer tube". The pump pistons were 200 mm in diameter with

a stroke of 1400 mm and were hydraulically driven. With a stroke time of 4.5 seconds per

stroke this gave a theoretical output of 35 m3 h

-1. In January 1990 the 200 mm diameter

cylinders and pistons were replaced with 230 mm diameter cylinders and pistons thereby

increasing the maximum theoretical output to 46 m3 h

-1. The hydraulic cylinder

changeover was controlled by pilot valves at the end of each stroke with the same pilot

valves used to sequentially changeover the "S transfer tube". The hydraulic oil system

consisted of a swash plate hydraulic oil pump and an "S transfer tube" changeover

accumulator charging pump driven by a 75 kW, 3 phase, and 415 volt electric motor. The

main hydraulic oil pump operated on a closed circuit free flow basis. The pump output was

variable from zero to maximum flow with the maximum pump speed of 13.3 strokes per

minute.

5.2.1 Dense Phase Pumping Plant Pipeline Sizing

The dense phase ash slurry was to be pumped 1737 meters from the DPAS plant to a

disposal site at the power station ash dam. The original design concept of the dense phase

ash slurry plant was that the plant would pump a dense phase paste; since the design

incorporated the concept of paste pumping the velocity of the paste was considered

immaterial. The major design criterion was the pipeline pressure drop (ΔP). Data obtained

from Verkerk (1986) was used in sizing the pipeline. The tender specified a pump with a

theoretical volumetric flow-rate of 30 m3 h

-1 at 9.0 MPa.

From Verkerk (1986), for a Cw of 65 % and a flow-rate 30 t h-1, the pressure drop was

assumed to be 250 kPa in a 125 mm pipeline 120 meters in length.

Calculating the shear stress (𝜏𝑤) from, Verkerk (1986);

𝜏𝑤 = (𝐷∆𝑃

4𝐿) 𝑃𝑎 (5-1)

62

𝜏𝑤 = 0.125 𝑥 250 000

4 𝑥 120 = 65.1 𝑃𝑎 (5-2)

The design operating pressure for the DPAS pipeline was set at 3.0 MPa.

Calculating the pipeline diameter (𝐷) for Vales Point applying Equation (5-1);

𝐷 = 4𝐿

∆𝑃 =

65.1 𝑥 4 𝑥 1737

3 000 000 = 0.151 𝑚 (5-3)

The nearest commercially available was pipe with a diameter of 150 mm. A 1737 metre

long pipeline was constructed from 12 metre lengths of 150 mm diameter black steel pipe

with a wall thickness of 5.22 mm to API5L, X46. The lengths of pipe were fitted with

welded "Victaulic" shouldered ends, and were joined by high pressure "Victaulic" flexible

couplings with a pressure rating of 7 MPa. The pipeline route followed the existing station

ash pipes and then continued cross country to the disposal site. The pipeline rose to a

maximum elevation of 12.6 metres and then fell so that the discharge was 3.6 metres above

the pump. A 3 metre flexible rubber hose with a pressure rating of 13.4 MPa connected the

pump to the pipeline. The flushing water systems obtained water from the power station

high pressure fire system.

5.2.3 Dense Phase Pumping Plant Control System

Figure 5.2 is a diagrammatic layout of the DPAS plant

The dense phase ash slurry pressure was measured at 5 different locations: the first at the

pump discharge after the flexible connection; and, at the remaining at 32 metres, 93

metres, 229 metres, and 1231 metres from the discharge transmitter.

These transmitters were designated discharge pressure and line pressures A, B, C, & D.

The transmitters were Druck flush diaphragm screw-in types with a pressure rating of 10

MPa.

The hopper level between the mixer and pump was measured using a Druck pressure

transmitter similar to the pipeline transmitters but with a pressure range of 0 to 50 kPa.

63

Figure 5.2 DPAS Mixing and Pumping System from Bunn (1991).

The temperature of the slurry pipeline was measured adjacent to the pipeline inlet with an

ADM RTD strapped to the outside of the pipeline with a range 0 to 100 C.

A Rosemount differential pressure transmitter was used to measure the pressure drop over

10.6 metres of the discharge pipeline. The pressure transmitters range was 0 to 50 kPa

The dry fly ash mass flow was measured with a Sankyo impactline flow-meter model

ILH-37 with a flow range of 0 - 60 t h-1

mounted between the rotary feeder and the mixer.

The bottom ash mass flow was measured with a Flowtek 30/45 impact flow-meter with a

dual flow range of 0 - 15 t h-1

or 0 - 30 t h-1 mounted between the screw conveyor and the

mixer.

The water mass flow was measured using a Processautomatic turbine flow-meter model

PAH/50/66 with a flow range of 6.6 - 66.0 t h-1.

Both the Putzmeister pump motor and the mixer motor were fitted with Universal Paton

64

power transducers, 0 to 80 kW for the pump and 0 to 20 kW for the mixer.

All pumping plant inputs were connected both a Texas Instrument Control Vision Unit

5000 (CVU 5000) and a 5TI Programmable Controller for process control, monitoring,

data processing as well as to two six channel analog recorders. A control loop to control

the slurry concentration required a calculation in the PC to determine the Cw. This

calculation was carried out in the ladder relay logic and was the ratio of fly ash mass flow

to total mass flow. The total mass flow was the sum of fly ash mass flow as measured by

the Sankyo impactline flow meter, plus water mass flow as measured by the

Processautomatic turbine flow-meter.

Test on samples taken from the pump suction hopper during pumping verified the accuracy

of the calculated Cw measurement. Three automatic control loops were installed. The first

used the calculated Cw to control the water mass flow rate. The second controlled the pump

output to maintain the pump suction hopper level. The third also used hopper level to

control the fly ash flow rate. Only the first two loops were commissioned. Because of

fluctuations in the fly ash mass flow due to problems with the air slide, the fly ash rotary

feeder was used to manually control the fly ash flow. For the same reason, the pump output

control loop was left on manual and the pump speed was manually adjusted to the required

(usually maximum) output. The pump hopper level was controlled by manually varying

the fly ash flow by adjusting the speed of the fly ash rotary feeder. A modification to the

Cw control loop was installed. It used differential pressure measured over a 10.6 m section

of pipeline adjacent to the pump discharge to trim the Cw set point. Trimming the Cw set

point with the DP input allows the control of the pipeline total pressure drop irrespective of

the changes in the pumpability of the ash slurry.

Normal DPAS plant operation was that the pump volume flow-rate was held constant, i.e.

at a constant shear rate and the Cw controller varied water flow to control the shear stress.

Varying the Cw thus maintains the shear stress constant, and therefore the pipeline

pressure, accounts for changes in the rheological properties of the slurry. These changes in

rheology were due to changes in the properties of the fly ash as received from the station

over which the DPAS plant had no control.

65

5.2.4 Dense Phase Pumping Plant Operations

The normal plant start-up procedure was that the fly ash air slide vent valve to "J" Silo was

opened and the Roots blower in the Amatek Facility was placed in-service to aerate the

surge bin and air slide. Initially, the slurry pipeline was filled with water from the flushing

water system. However, later operating experience has shown that this was not necessary

and therefore this step has now been deleted. At the same time, the Putzmeister pump

suction hopper was filled with water via the spray water header using the water control

valve on manual control. After the suction hopper was full, the pump was placed

in-service. The pump output was manually adjusted to the required output and the hopper

level was reduced to approximately 1200 mm. The rotary feeder was started and brought

up to approximately the required speed. The air slide inlet regulating valve was then

opened to give the required fly ash mass flow. The rotary feeder speed and air slide inlet

valve were then adjusted to try and stabilise the fly ash flow as much as possible and the

Cw control was placed on automatic.

The DPAS plant was operated for as long as Amatek could supply fly ash or until 1400

hrs. A Cw was selected and the pumping continued at this Cw until the DP stabilised. It

required approximately 50 minutes for priming of the pipeline with slurry with the 200

mm cylinders and 40 minutes with the 230 mm cylinders, pump at maximum speed. The

pipeline was operated up to a maximum of 3.0 MPa.

The plant was shut down by first placing the water control valve on manual and then

closing the air slide inlet regulating valve. The rotary feeder and air slide were purged of

fly ash and taken out of service. The Putzmeister pump was taken out of service and the

pipeline was flushed by slowly placing the flushing water system in-service. With the

pump out of service, the mixer and pump suction hopper was flushed by using the spray

water system and removing a drainage cover in the bottom of the pump suction hopper.

After the pipeline was flushed, the flushing water system was taken out of service. The

pipe between the flushing water connection and the pump was flushed by back pumping

with the Putzmeister pump. Back pumping was a facility available on the Putzmeister

pump that allows the pump to be run so as to pump water from the pipeline back into the

pump suction hopper.

66

The flushing water system, with a maximum flow of 30 l s-1

gave a pipeline velocity of

approximately 2.0 m s-1

which was adequate to re-entrain any deposited fly ash slurry at

the bottom of the pipeline. After the pump was stopped, the flushing water supply valve

was slowly opened over a period of several minutes to maximum flow, after this, it

required approximately another 20 minutes to flush the pipeline. If the pipeline was

blocked during pumping with high Cw slurry, the flushing water system was used to

unblock the pipeline. This was achieved by placing the flushing water system in-service

with a very small flow of water. The small flow of water finds its way through the slurry

thus effectively reducing the slurry Cw. The pump was then placed back in-service and the

test resumed.

Over the period of operation, 110 pumping tests were conducted on dense phase fly ash

slurries utilising Vales Point Units No. 5 & 6 fly ash at differing Cw’s. In a typical

pumping run on the 15/2/1990, the pipeline reached a pressure of 2.4 MPa, with a Cw set

point of 59.0 %, a fly ash flow of 36 t h-1, a water flow of 24 t h

-1, a pump power of 45 kW

and a mixer power of 6.5 kW. When the slurry changed from a fluid flow to a plug flow

situation, there is a large step change in pipeline pressure. When large fly ash fluctuations

occurred, the Cw control loop was normally placed on manual and excess water was added

to the mixer until the fly ash stabilised. When the Cw set point was reduced by

approximately 1.0 % to 2.0 % the plant was operated with frequent small fluctuations in

fly ash flow. This caused drier and wetter plugs of slurry to be scattered along the pipeline

during pumping while still maintaining the desired pipeline pressure drop. If the Cw control

loop was left on automatic at the higher set point during the large fluctuations in fly ash

flow, there were two consequences:

the mixer motor tripped on thermal over-load when a large decrease in fly ash flow

was followed by a rapid increase as there was insufficient water to mix the ash and

the mixer blocked; and,

drier plugs of slurry were pumped into the pipeline causing the pump to stall on

high pressure.

Pumping occurred with both the 230 mm diameter cylinders and the 200 mm diameter

cylinders. The shear stress was calculated using Equation (5.1). The (∆𝑃) used for the

67

calculations was the maximum reached on the particular pumping run at that Cw. Figures

5.3 and 5.4 are diagrams for the pumping for the DPAS plant at Vales Point, using the 200

mm and 230 mm diameter cylinders. This corresponds to a shear rate of 29.1 s-1 for the

200 mm diameter cylinders and 39.0 s-1 for the 230 mm diameter cylinders.

Figure 5.3 Pilot Plant Pumping Results from Bunn (1991).

The diagrams demonstrate that there was a wide range of shear stress at the same Cw. The

Cw ranges from approximately 58.0 % to 64.0 % with the shear stress ranging from 16.6 Pa

to 67.1 Pa. The large variation in shear stress at similar Cw is due to the changes in

pumpability of the slurry, which were due to changes in the particle size distribution of the

68

fly ash. The DPAS has no control on the uniformity of the fly ash received from the

different precipitator hoppers.

Figure 5.4 Pilot Plant Pumping Results from Bunn (1991).

5.2.5 Determination of Pipeline Slurry Settling Velocity

During later pumping, it was determined that sedimentation occurred in the pipeline. To

determine the degree of this sedimentation, a 1 meter × 150 mm glass pipe section was

installed at the ash dam end of the pipeline. On inspection of the glass section, it was

observed that sedimentation occurred to a depth of approximately 50 mm. Sedimentation

in the pipeline occurred when the average superficial velocity of the slurry was both 0.55

m s-1

and 0.73 m s-1

. These two flow rates corresponded to the maximum pump output

with the 200 mm diameter cylinders and the 230 mm diameter cylinders, respectively. To

determine the superficial average velocity where no sedimentation was present, a 30

metres section of 100 mm pipe was installed at the ash dam end of the pipeline. Included in

69

this 30 metre section of pipe was a 1 meter × 100 mm glass pipe. No sedimentation was

observed in the 100 mm glass section with the pump stroke at 100 % (230 mm cylinders)

corresponding to an average superficial velocity was 1.65 m s-1.

The settling velocity of the pipeline was determined indirectly because there was no

instrumentation to directly measure the velocity. The theoretical volumetric capacity of a

pump stroke was calculated. The pump output was then calculated at different speeds by

timing the pump stroke. The velocity of the slurry was then calculated at the various speed

settings. The plant was placed into service and the pipeline was filled with slurry at Cw of

60.0 % with the pump stroke at 100 %. With the pipeline full of slurry, the pump stroke

was reduced in 5 % steps with a 15 minute waiting period at each step. Observations were

then made at the 100 mm glass viewing section. If, after 15-minutes no settling was

observed, the speed was increased to 100 % for 10 minutes and the test was repeated at the

next lower step. It was observed that settling occurred at 80 % stroke but not at 85 %

stroke. When the pump stroke was increased from 80 % to 85 %, the slurry re-suspended

and no settled slurry was left. Therefore it was concluded that settling occurred when the

velocity was reduced to less than 1.2 m s-1

and that re-suspension occurred above 1.3 m s-1.

5.2.6 Dense Phase Pumping Plant Slurry Transfer

Approximately 9,300 tonnes of dense phase fly ash slurry with 58.0 % < Cw < 64.0 % was

transferred to the disposal site. The normal pumping range of the dense phase slurry was

59.0 % < Cw < 63.0 %. This gave the highest possible Cw while maintaining a desired

pipeline pressure. The fly ash slurry was transferred at a rate of approximately 40 t h-1 with

the 200 mm cylinder and at 60 t h-1

with the 230 mm cylinders. The slurry was deposited

on the shore of the existing ash dam. An area of 10 metres by 10 metres was marked out

with the centre of the square directly under the initial pipeline discharge. At the sides of the

square, deposited fly ash slurry had accumulated to a height of 1150 mm. The angle of

inclination of the deposited fly ash was 2 to 3 degrees. The deposited fly ash slurry had

packed down hard and had dried with a capping (approximately 5 mm thick). It had been

observed that later deposited material remained sticky and wet when deposited on top of

the existing slurry, and it required several days to dry out. The heavy rainfall during the

70

pumping program produced no erosion at the site. Erosion however did occur below the

discharge pipe during water flushing of the pipeline.

Figure 5.5 is a photograph of slurry flowing from the DPAS pipeline at the disposal area.

Figure 5.5 Pilot Plant Pumping Pipeline Discharge from Bunn (1991).

5.3 Pipeline Viscometers

Two other pipeline viscometers were used for testing. One was a small scale pumping rig

consisting of a 32 mm Mono positive displacement pump and a steel pipeline. The Mono

pump test rig was designed, and constructed by the author. The other was the ABB

designed and manufactured Rotary Ram Slurry Pump (RRSP) and pipeline. ABB installed

the RRSP test rig in a facility at Thornton, west of Newcastle to hydraulically convey

dense phase coal slurries. The facility at Thornton was made available to the Electricity

Commission to allow co-operative research to obtain data on the characteristics of dense

phase fly ash slurries.

71

5.3.1 Mono Pump Test Rig

The Mono pump was a positive displacement type driven by a 3-phase alternating current

415 volt electric motor. A variable frequency power supply enables the Mono Pump output

to be regulated from zero to maximum. A 19.2-metre steel pipeline with an inside diameter

of 32 mm (±1 mm) was connected to the pump discharge. The pipeline discharged to

waste. A conical hopper was attached to the pump suction with a capacity of 0.046 m3. A

15 mm socket welded into the pipeline adjacent to the pump discharge was used to connect

a pressure transmitter. A Fisher Porter diaphragm type with a range of 0 to 500 kPa and

output of 4 to 20 mA was used. A chart recorder continuously recorded the pressure

transmitter output.

Figure 5.6 Mono Pump Test Rig from Bunn (1991).

5.3.2 Mixing Technique and Measurements for Mono Pump Test Rig

The Mono pump slurries were mixed in the following way. The mixing was achieved

using a small commercially available electric driven concrete mixer. Four 20-litre

containers of freshly collected fly ash were added to the concrete mixer. 200 grams of dry

fly ash was removed from the fly ash containers for particle size distribution analysis.

Domestic tap water from the Hunter District Water Board (HDWB) was added to hopper

72

to obtain the desired Cw and mixed for approximately thirty minutes. Before the slurry was

added to the mono pump suction hopper the hopper was filled with water and pumped dry.

The contents of the mixer were then tipped into the suction hopper and the pump speed

increased to maximum. When a good slurry flow was visible at the end of the pipeline,

approximately 3 kg of slurry was collected in a plastic bag, over approximately 30

seconds, to determine the mass flow rate and the pressure (ΔP) on the recorder noted. Tests

were conducted at different mass flow rates until the hopper was emptied. The Cw was

determined by collecting 2 samples of slurry in glass bottles of known weight, one near the

beginning and the other near the end of the test. The glass bottles filled with slurry were

then weighed and placed in a drying oven with a temperature of approximately 105C. The

bottles were then re-weighed and the Cw was calculated. To ensure that the pipeline was

completely flushed, the pressure transmitter connection was removed and the pipeline was

flushed with a high flow rate of water. The mixer was then cleaned and new fly ash slurry

was mixed at the required Cw.

Comparative tests were conducted with a rotary viscometer and the mono pump test rig. A

pseudo shear diagram for Vales Point fly ash slurries using the Mono pump test results and

rotary viscometer results are shown in Figure 5.7, which indicates the relationship between

shear rate (8𝑉

𝐷) and the shear stress (

𝐷∆𝑃

4𝐿) drawn on linear scales. The pseudo-shear

diagram shows the relationship for fly ash slurries at several different Cw.

5.3.3 Calculations Mono Pump Test Rig

The slurry density (𝜌𝑠𝑙 )was calculated using the relationship:

𝜌𝑠𝑙 =100

⌊((𝐶𝑤𝜌𝑠

) + (100−𝐶𝑤

𝜌𝑤))⌋

(5-4)

where 𝐶𝑤 = Slurry concentration by weight (%)

𝜌𝑠 = Solids density (kg m3)

𝜌𝑤 = density of water (kg m3)

73

The mass flowrate (𝑄) of the slurry was calculated by weighing the plastic bag and using

the following relationship:

𝑄 =(𝑤𝑒𝑖𝑔ℎ𝑡 𝑜𝑓 𝑠𝑙𝑢𝑟𝑟𝑦 𝑠𝑎𝑚𝑝𝑙𝑒)

𝑡 (5-5)

Where 𝑄 = slurry mass flowrate (t h-1

)

𝑡= time

Figure 5.7 Pseudo -shear Diagram for Mono Pump and Rotary Viscometer for Vales Point

Fly Ash from Bunn (1991).

The wall shear stress was calculated using the relationship:

𝜏𝑤 =𝐷𝛥𝑝

4𝐿 (5-6)

74

The wall shear rate was calculated using the relationship:

𝛤𝑤 =32 𝑊

𝜋𝜌𝑠𝑙 𝐷3 (5-7)

5.4. Rotary Ram Slurry Pump Thornton Test Rig

The test rig consisted of a variable speed Rotary Ram Slurry Pump (RRSP) and associated

equipment, twin mixing tanks and a 50 mm inside diameter steel pipeline. The pipeline

consisted of several loops of differing length and discharged back to the mixing tanks or to

waste. The pipeline instrumentation consisted of a magnetic flow-meter and pump

discharge pressure transmitters.

5.4.1 Mixing Technique and Measurements for RRSP

Three 200-litre drums of fly ash were collected from the station being tested and

transferred to Thornton. Approximately 200 grams of dry fly ash was removed from the fly

ash containers for particle size analysis. ABB workmen flushed the mixer, pump and

pipeline with water prior to each test. After flushing, the mixer was left with enough water

to cover the blades, the stirrer was started and fly ash was added to the mixer until the

designated Cw was obtained. Water was then added to reduce the Cw for subsequent tests.

Flow and pressure measurements for the RRSP test rig were carried out as follows:

the pump was started and the slurry recirculated through the system;

measurement commenced after several minutes of slurry recirculation;

approximately 200 grams of slurry was placed in a pre-weighed plastic container;

the wet sample was dried in an oven at 60C;

the dry slurry was then re-weighed and the Cw was calculated; and,

the slurry mass flow was manually checked.

75

In parallel with the pumping test using the RRSP, comparative testing was carried out

using the rotary viscometer for the Bayswater and Eraring Slurries.

Figure 5.8 RRSP Layout Diagram from Bunn (1991).

5.4.2 Calculations for the RRSP

The slurry density, mass flowrate, wall shear stress and wall shear rate were calculated as

in section 5.3.2. A pseudo-shear diagram for Vales Point, Eraring and Bayswater fly ash

slurries using the RRSP and rotary viscometer (RV) results are shown in Figures 5.9, 5.10

and 5.11. They indicate the relationship between shear rate (8𝑉

𝐷) and the shear stress

(𝐷∆𝑃

4𝐿) drawn on linear scales.

5.5 Viscometers Results

The pipeline viscometers results indicate that the fly ash slurries mixed with Hunter

District Water Board domestic water could be pumped up to a maximum Cw = 61.4 % for

Vales Point fly ash, a Cw = 69.5 % for Eraring fly ash and a Cw = 78.6 % for the Bayswater

fly ash.

76

Figure 5.9 Pseudo-shear Diagram for RRSP and RV for Vales Point Fly Ash from Bunn

(1991).

Figure 5.10 Pseudo-shear Diagram for RRSP and RV for Eraring Fly Ash from Bunn

(1991).

77

Figure 5.11 Pseudo Shear Diagram for RRP and RV Bayswater Fly Ash from Bunn

(1991).

Initially all testing of the Vales Point fly ash slurries was carried out using either the Mono

pump test rig or the rotary viscometer. However, because of the disparity between the

results comparative tests were conducted. The calibration of the rotary viscometer was

checked using the Contra 150 oil of known viscosity. The comparative rheometry

conducted on Eraring and Vales Point fly ashes indicated that there was a significant

difference between the results using the Mono pump test rig, RRSP and the RV at the same

Cw, whereas the Bayswater fly ash indicates a certain similarity.

The assumption is that the Bayswater fly ash slurries tested are homogeneous slurries,

whereas, the Vales Point and Eraring fly ash slurries are heterogeneous.

78

CHAPTER 6 PREVIOUS RESEARCH PAPERS

6.1 Introduction

This chapter is a collection of all the papers since 2004 that the author has published and

presented at national, international conference or published in journals to further my

understanding of the transport and disposal of power station ash.

6.2 Summary

The conclusions reached from these papers were:

difference in the flow times between the ASTM Flow Cone and the Marsh

Funnel are small but not insignificant;

fly ash grout strength can vary significantly depending on the source of the ash;

for a 20 seconds flow cone the Cw of the Bayswater power station “Run of

Station Ash” grout varied between 71.5 % and 62.5 % at a similar viscosity.

the question arises as to what effect this extra water would have on the strength

of the grout?

pumping characteristics of high concentration fly ash slurry pipeline change due

to shearing in the pipeline depends on the type of coal the power station burns

and the properties of the process water;

the change in pumping characteristics due to the shearing in the pipeline is

insignificant compared with the changes in pumping characteristics due to

changes in particle size distribution of the fly ash;

79

a slurry pipeline pressure characteristics of piston pumps and Peristaltic hose

pumps are similar;

comparison of the rheological results for a rotary viscometer and a pipeline

viscometer for ash from a Queensland black coal power station indicate that the

rotary viscometer overestimated the pumping Cw by 7%, i.e. the pumping Cw

would be 64 % not 71 %;

a laboratory trial to simulate high concentration fly ash slurry pipeline pumping

characteristics indicated there would be little change in the pipeline pressure drop

per unit length due to thixotrophic behaviour;

a review of low concentration tailing dam failures indicted that the following are

viable paste alternatives:

o tailings as a paste can be placed without binder on a temporary surface

stacked stockpile, and then placed in an open cut void after mixing with

a binder;

o tailings as a paste can be placed with a binder on a surface emplacement

of a desired shape of bound stabilized fill;

o tailings as a paste can be placed without binder as a mine backfill into

old voids either open cut or underground, and:

o tailings as a paste can be placed with a binder as a mine backfill into old

stopes to improve mining extraction ratios.

a rheological study with a rotary viscometer indicated that coal washeries thickener

underflow could be pumped as a high concentrate slurry;

a paper that examined the maximum amount of water that is available for

recycling from a range of dense phase fly ash slurries found:

80

o the percentage of water available for recycling varies depending on the

pumped Cw and the PSD;

o for all the slurries tested the percentage of water that was available for

recycling varied between 25 to 60 % of the water mixed with the fly ash;

o there was no relationship between the surface deformation of the

deposited slurry and the Cw at which the slurry was pumped;

o there is no relationship between the Cw of the pumped slurry and

percentage of return water, PSD and packing density, and:

o the deposited slurry placement density showed an increase when the

slurry could be pumped above a Cw of 65%.

the relationship between the packing density of slurry obtained by assisted

compaction and the pumpability as determined by rheology testing and found

that if you add 15 % extra water to the results of the compaction tests you can

assume that this will give you a reliable indication of the pumpability of fly ash

slurries;

fly ash slurries from different power stations show a great variation in rheology

which can be related to the differences in PSD. However, the variation in

rheology cannot be equated directly to the d50 of the fly ash particles but the

variation in the distribution of the particles across the PSD range;

packing efficiency as predicted by the computer model and the packing

efficiency as determined from the assisted settling tests results for the fly ash

tested show good correlation, and:

a change in coal supply can adversely affect the operation of a dense phase ash

handling and pumping system.

81

6.3 11th

International Conference Bulk Materials Storage Handling and

Transportation (2013) - Comparative Rheology of Fly Ash Slurries using Rotary

and Pipeline Viscometers

This paper describes a pilot plant investigation of Bayswater fly ash slurries where the

flow behaviour of the slurries was investigated in respect to the effect of solid

concentration, pipe size and flow velocity. A re-circulation pipe loop configuration

with instrumented pipes of nominal bore of 80 mm and 53 mm was used for measuring

the slurry flow parameters. Fly ash from Bayswater power station with an average d50 of

23.8 μm was used as solid phase, with water as the carrier liquid. Slurry mass

concentration reached 67.9 % and shear rates up to 200 s-1

.

Over the last two decades there has been an increase in the quantity of power station fly

ash pumped to disposal sites using dense phase slurry systems. These dense phase

systems have either been new plant or retrofitted to an existing plant. The

determination of the pumping characteristics of these fly ash slurries requires a

combination of either bench top or pilot plant studies.

It is important to first determine if these slurries are homogeneous or heterogeneous

suspensions. Homogeneous flow is a symmetric flow characterizing uniform

distribution of solids about the horizontal axis of the pipe. (Durand and Condolios 1952)

published a number of studies indicating that homogeneous suspensions are those that

contain all particles smaller than 40 μm while (Shook et al 2002) suggests that for

suspensions with a mean particle diameter (d50) greater than 50 μm the slurry will

display heterogeneous properties. He also indicates that fine particles slurries d50 less

than 50 μm typically exhibit homogeneous fluid behaviour. Fly ash particles from

modern coal fired power stations are nominally spherical with a d50 ranging from 8 to

45 µm, therefore some of the fly ashes with low d50 should be classified as homogenous,

but which ones?

(Thomas 1976) outlined that if pipe loop tests are performed on slurry at the desired Cw

in a number of different diameter pipes and the pressure gradient verses velocity is

plotted on log-log plot, if the results are a straight line, then the slurry is homogeneous.

82

(Bunn 1991) conducted comparative rheometry using a rotary viscometer and the

Rotary Ram Slurry Pump Thornton Test Rig (RRSP) with Bayswater fly ash slurry in May

1990. Figure 6.1 is a graph of the results. The results showed non- Newtonian curves

fitting as a Bingham plastic model and clearly indicate that the slurries were homogeneous

because of the similarity between the rotary viscometer results and the RRSP results at the

same concentrations by weight. The fly ash particle size distribution had a d10 of 3 µm, d50

of 15 µm and a d90 of 52 µm as shown in Figure 6.2.

All non-Newtonian fluids show shear rate dependent viscosity (Khan 1992). A flow

curve of shear rate verses shear stress is used to characterise a non-Newtonian flow.

The shear stress at the wall of a pipe of diameter D and length L is related to the

pressure drop (ΔP) by,

𝜏𝑤 = 𝐷ΔP

4𝐿 (6.1)

Figure 6.1 Bayswater Fly Ash Slurry Comparative Rheometry Results from (Bunn

1991)

83

The wall shear stress for a Newtonian fluid is determined by,

𝜏𝑤 = 𝜇 [𝑑𝑣

𝑑𝑟] (6.2)

where µ is the fluid viscosity and 𝑑𝑣

𝑑𝑟 is the velocity gradient at the wall. The shear rate

is determined by,

[𝑑𝑣

𝑑𝑟]

𝑤 = =

8𝑉

𝐷 (6.3)

where V is the mean linear velocity. Then the wall shear stress is,

𝜏𝑤 = 𝐷ΔP

4𝐿 = 𝜇 (

8𝑉

𝐷) (6.4)

𝜇 = [

𝐷ΔP

4𝐿]

[8𝑣

𝐷]

(6.5)

For Newtonian fluids µ is constant for all values of shear rate in the laminar flow regime.

In the case of non-Newtonian fluids µ is not constant and is however dependent on the

shear rate.

(Sieve and Lazarus 1986), (Verkerk 1985), (Singh 1989), Bunn and (Chambers1992), and

(Chandel el al 2009), observed that fly ash slurries were Non-Newtonian and behaved as a

Bingham Plastic. (Bingham 1922) reported that some slurries exhibit plastic or visco-

plastic behaviour, i.e. they behaved as solids at lower shear stresses but behaved like

viscous fluids when a critical shear stress was exceeded. Bingham developed a simple

model for this characteristic described as;

𝜏𝑤 = 𝜏𝑦 + 𝜇 ∶ (𝜏 > 𝜏𝑦) (6.6)

The Bingham model predicts a linear relationship between shear stress and shear rate at a

84

shear stress above 𝜏𝑦. This is referred to as the Bingham Yield Stress. A typical flow

curve for Bingham Plastic Fluids is linear and the intercept of the flow curve at a zero

shear rate determines the yield stress.

(Bird el al 1960) and (Skelland 1967) indicated that laminar flow conditions in a tube

viscometer could be verified by showing that the Reynolds Number 𝑅𝑒is less then 2100

using the relationship:

𝑅𝑒 = 𝐷𝑉 𝜌𝑠𝑙

𝜇 (6.7)

where 𝜌𝑠𝑙 is slurry density.

A concentric cylinder rotational viscometer is a suitable instrument for measuring both

Newtonian and non-Newtonian slurries. It can be used to determine both yield points and

thixotrophic structures (Schramm1981). Shear heating can be a problem at high shear rates

and therefore the rotary viscometer is limited by the maximum shear rate. It can only be

used for laminar flows because of errors caused by Taylor Vortices. In addition, it can only

be used for fine material slurries and not for coarse material because of the small

clearances between the rotor and cup.

6.3.1 Experimental Material and Equipment

The fly ash used in the experiment was “run of station ash” supplied in 2 x 200 litre

steel drums supplied by Bulk Flyash Grout Pty Ltd. A Malvern Particle Size Analyser (a

laser diffraction technique instrument) was used to determine the fly ash particle size

distribution (PSD). These initial dry fly ash samples were designated to be PSD 1 to

PSD 4, and are shown in Figure 6.2.

85

Figure 6.2 PSD’s for Bayswater Fly Ash and data PSD May 1990 from (Bunn (1991)

The slurry flow parameters were measured on an experimental re-recirculation slurry

pump pilot plant, shown diagrammatically in Figure 3. The pipeline was constructed of

80 mm mild steel schedule 40 pipe with actual inside diameter of 77.92 mm. On the

return leg, a 6.5 m length of 50 mm mild steel schedule 40 pipe with actual inside

diameter of 52.5 mm, replaced some of the 80 mm pipe. The pipeline also included a 80

mm inside diameter glass pipe which allowed for visual inspection of the slurry flow.

The slurry was pumped with a Hidrostal Screw Centrifugal Impeller Pump from an

agitated open storage tank to the pipeline. Attached to the Hidrostal Pump Motor was a

Zener MSC-3 Variable Speed Drive which allowed for variable slurry flow. The pump

discharge pressure was measured with a Impress Pressure Transmitter and the

differential pressure over 5 m of both the 80 mm and 50 mm pipe was measured with

individual Yokogawa Diaphragm Sealed Differential Pressure Transmitters. The slurry

flow-rate was measured with a Foxboro Magnetic Flow Meter and the slurry mass flow-

rate was measured with an automatic weigh hopper. The slurry weigh hopper was

mounted on load cells. When the measured hopper weight was 25 kg an automatic valve

would open and dump the slurry into the mixing hopper.

0

10

20

30

40

50

60

70

80

90

100

0.1 1 10 100 1000 10000

Per

cen

tage

Pas

sin

g (%

)

Particle Size (µm)

PSD 1 PSD 2 PSD 3 PSD 1A

PSD 2A PSD 3A PSD 4A PSD 5A

PSD 6A PSD 7A PSD 4 PSD May 1990

86

All the pressure, differential pressure, temperature, volumetric flow and differential

weight data was collected by a DataTaker DT 800. The data from the DataTaker was

collected in real time by a laptop computer.

6.3.2 Slurry Mixing and Measurement

Initially the pilot plant slurry hoppers and pipeline were flushed with water. The mixing

tank isolating valve was then closed leaving the pump and pipeline full of water. The

slurry was then prepared in an external mixer by adding 130 kg of water then slowly

adding 200 kg of fly ash.

Figure 6.3 Schematic of Slurry Test Rig (1 – Slurry Mixing Hopper, 2 – Mixer Hopper

Isolating Valve, 3 - Hidrostal Screw Centrifugal Impeller Pump, 4 – Pipeline Isolating

Valve, 5 – Pressure Transmitter, 6 - Differential Pressure Transmitter 80 mm Pipe, 7 -

Reducers 80 to 50 mm, 8 - Differential Pressure Transmitter 50 mm Pipe, 9 – Pipeline

RTD, 10 - 80 mm Glass Viewing Section, 11- Magnetic Flow Meter, 12 – Weight

Hopper, 13 – Weigh Hopper Control Valve.)

87

The mixed slurry was then added to the pilot plant mixing hopper and agitated for

several minutes to ensure a homogeneous mixture. The isolating valve was opened and

the slurry pump started on low speed. The slurry in the pipeline flushed water from the

pipe and this continued till slurry appeared at the pipeline discharge point. The system

was placed on full recirculation flow for several minutes at a pump frequency of 30 Hz.

The pump speed was then reduced to minimum and the data collection system started.

The slurry pump speed was run for several minutes and then the pump frequency was

increased in 5 Hz steps with a pause of several minutes at each step.

When the pump frequency reached 35 Hz, a sample of slurry was collected and sheared

in a Contraves model Rheomat RM - 30 rotary viscometer. While pumping, 7 wet

samples of slurry were taken and analysed to determine Cw, density and PSD. These

samples were designated PSD 1A to PSD 7A and were added to Figure 6.2, along with

the PSD of the fly ash used in Figure 6.1.

After collecting the pumping data at maximum pump speed the pump frequency was

reduced to 30 Hz. At this stage an extra 25 kg of dry fly ash was added slowly to the

mixing hopper and the pump run for several minutes at 30 Hz to ensure complete

mixing. The speed was then reduced to minimum and data collected. The process of

adding fly ash was repeated twice more.

The fly ash solids density was tested with a Micromeritics AccuPyc Pycnmoter 1330.

The results are displayed in Table 6 1. The average fly ash solids density was calculated

at 2047 kg m-3

.

Table 6.1 Bayswater Fly Ash Density

Date

Density

1A

kg m-3

Density

2A

kg m-3

Density

3A

kg m-3

Density

4A

kg m-3

Density

5A

kg m-3

Density

6A

kg m-3

Density

7A

kg m-3

Density

1

kg m-3

27/11/12 2004.4 2006.1 2005.5 2077.1 2065.5 2070.3 2077.4 2069.0 Average

2003.7 2005.6 2005.5 2076.9 2065.0 2070.7 2076.5 2069.4

2003.2 2005.4 2005.2 2076.3 2064.8 2. 69.2 2076.2 2068.3

Average 2003.8 2005.7 2005.4 2076.8 2065.1 2070.1 2076.7 2068.9 2046.6

88

6.3.3 Experimental Results and Analysis

Comparison of the measured volumetric flowrate and the volumetric flowrate calculated

from the mass flowrate and slurry density is shown in Figure 6.4. These results show

that the measured volumetric flow-rate was accurate and therefore this measurement

was used for all calculations.

The slurry concentration by weight (Cw) was calculated by weighing the wet slurry, then

drying it in an oven and reweighing the dry sample. The measured Cw’s of the slurries

were 59.7%, 61.8 %, 65.1% and 67.9%. Experimental values of the slurry velocity 𝑉𝑠 ,

pressure gradient 𝑃𝑔, shear stress 𝜏𝑤 and shear rate were determined from the

measured values of flowrate and differential pressure.

The flow behaviour of the fly ash slurries was obtained by plotting the pressure gradient

verses velocity for fly ash slurries at different Cw and pipe sizes, see Figures 6.5 and 6.6.

Figure 6.7 is a log–log plot of pressure gradient verses velocity for the fly ash slurries at

selected Cw’s and water for both pipe sizes. From Figure 6.7 it can be determined that

the water curves are straight lines therefore are consistent with homogeneous flow. The

graphs for the fly ash slurries are not straight lines indicating heterogeneous flow.

Figure 6.8 is a pseudo shear diagram for the fly ash slurries at different Cw’s to which is

added the rotary viscometer results. Inspection of Figure 6.8 indicates that the rotary

viscometer results at each Cw are significantly less that the corresponding results

obtained from the pipeline viscometers.

89

Figure 6.4 Measured and Calculated Volumetric Flowrate at Cw 65.1 %

Table 6.2 is a comparison of the results for the pipeline viscometers and rotary

viscometer at a shear rate of 100 s-1

. On average the rotary viscometer readings are 50 to

60 % less than the pipeline viscometers.

The Reynolds Number, 𝑅𝑒 , was calculated for the pipeline viscometer results using

equation (7) and the results are displayed in Table 6.3. Row 9 is data from (Ward el al

1998) for the 26th April 1996 and row 10 is from data collect from Bayswater by the

author on the 19th March 2013. In all cases 𝑅𝑒 < 2100 indicating that the results from

the test loop and the full scale operating plant at Bayswater Power Station under normal

operating conditions operate in laminar flow.

Figure 6.2 indicates that the PSD distribution of the present day fly ash from Bayswater

is coarser than the fly ash burnt in 1990. The sourcing of some of the power station coal

supplies from the Western Coal Fields instead from the Hunter Valley as indicated by

(Bunn el al 2004) would explain the change in the d50 values from 1990 to 2013.

15

20

25

30

35

40

0 2 4 6 8 10 12 14 16 18 20

Pu

mp

Fre

qu

ency

( Hz)

Flowrate (m3 h-1)

Flowrate Measured by 80 mm Magnetic FlowmeterFlowrate Calculated from Slurry Weigh Hopper

90

Figure 9 is a column graph displaying the d10, d50 and d90 of the Bayswater fly ash used

for pumping and data from Bunn (1991). A measure of the slope,ds, of the PSD curves

was obtained as follows,

𝑑𝑠 =𝑑90 − 𝑑10

𝑑50 (6.8)

The average slope of the 11 data points for the Bayswater fly ash was 3.21 and for the

1990 Data from Bunn (1991) was 3.26.

Figure 6.5 Pressure Gradient Verses Velocity for 80 mm Pipe at Different Cw

0

0.5

1

1.5

2

2.5

3

0 0.5 1 1.5

Pre

ssu

re G

rad

ien

t (k

Pa

m-1

)

Velocity (m s-1)

Cw 67.9 %" Cw 65.1 % Cw 61.8 %'" Cw 59.7 % Water

91

Figure 6.6 Pressure Gradient Verses Velocity for 50 mm Pipe at Different Cw

Figure 6.7 Pressure Gradient Verses Velocity of Water and Two Different Cw Slurries

0

1

2

3

4

5

6

7

8

9

0 0.5 1 1.5 2 2.5 3

Pre

ssu

re G

rad

ien

t (k

Pa

m-1

)

Velocity (m s-1)

Cw 67.9 %" Cw 65.1 % Cw 61.8 %'" Cw 59.7 % Water

0.001

0.01

0.1

1

10

100

0.01 0.1 1 10

Pre

ssu

re G

rad

ien

t (k

Pa

m-1

)

Velocity (m s-1)

80 mm Cw 67.9 % 50 mm Cw 67.9 % 80 mm Cw 59.7 %

50 mm Cw 59.7 % 80 mm Water 53 mm Water

92

Figure 6.8. Pseudo Shear Diagram

Figure 6.9 Bayswater Fly Ash PSD d10, d50 and d90 and data PSD 5 May 1990 from Bunn

(1991)

0

10

20

30

40

50

60

70

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

80 mm Cw 67.9 % 50 mm Cw 67.9 % 80 mm Cw 65.1 % 50 mm Cw 65.1 %80 mm Cw 61.8 % 50 mm Cw 61.8 % 80 mm Cw 67.9 % 50 mm Cw 59.7 %RV Cw 67.9 % RV Cw 65.1 % RV Cw 61.8 % RV Cw 59.7 %

4.7 4.8 4.8 4.8 4.2 4.4 4.0 4.3 4.1 4.1 4.3 3.0

25.2 22.7 23.4 25.2 22.9 25.5

21.2 24.2

22.3 24.3 25.5

15.0

89.5

83.0 77.8

88.0

75.9

86.5

67.1

80.9

72.5

82.8 86.2

52.0

0.0

10.0

20.0

30.0

40.0

50.0

60.0

70.0

80.0

90.0

100.0

1 2 3 4 5 6 7 8 9 10 11 12

Par

ticl

e Si

ze µ

m

Fly Ash d10, d50 & d90 1 to 11, 12 - 5th May 1990

d10 d50 d90

93

Table 6.2 Comparison of Pipeline and Rotary Viscometer at Same Shear Rate and Different Cw

Pipeline Viscometer Rotary Viscometer

(%)

Cw

(s-1)

(Pa)

𝜏

(s-1)

(Pa)

𝜏

59.7 100 8.0 100 5.4

61.8 100 14.8 100 7.8

65.1 100 25.0 100 13.0

67.9 100 46.0 100 27.0

Table 6.3 Calculated Values of Reynolds Number for Different Cw

No. (m)

D

(%)

Cw

V

(m s-1

)

𝜌𝑠𝑙

(kg m-1

)

(s-1)

𝜏

(Pa)

𝜇

(Pas)

𝑅𝑒

1 0.0525 59.7 0.65 1426 100 8.0 0.08 616

2 0.0525 61.8 0.65 1450 100 14.8 0.15 330

3 0.0525 65.1 0.65 1480 100 24.0 0.24 210

4 0.0525 67.9 0.65 1520 100 45.0 0.45 115

5 0.0779 59.7 0.97 1426 100 8.0 0.08 1364

6 0.0779 61.8 0.97 1450 100 14.9 0.15 736

7 0.0779 65.1 0.97 1480 100 26.0 0.26 430

8 0.0779 67.9 0.97 1520 100 47.5 0.48 242

9 0.2 73.0 2.2 1666 76 32.0 0.42 1745

10 0.2 69.0 2.2 1530 88 32.5 0.37 1824

6.3.4 Conclusions

The Bayswater fly ash tested in the experimental re-recirculation slurry pump pilot plant

behaved as a heterogeneous slurry as indicated by the curved lines on the log-log plot.

The significant difference in the results between the pipeline viscometer and the rotary

viscometer also indicated that the slurries were heterogeneous. Furthermore all testing

was carried out in the laminar flow region so the models used were suitable.

94

This study indicates that the assumptions of (Durand and Condolios 1952), that

homogeneous suspensions are only those that contain all the particles smaller than 40

μm particles may need to be reviewed. (Durand and Condolios 1952) did not take into

account the shape of the PSD curve as the particle measurement technology was not

available at that time.

The assumption of (Shook et al 2002) that suspensions contain particles with d50 less

than 50 μm typically exhibit homogeneous fluid behaviour conflicts with results of this

study. Results presented indicate that a suspension containing particles with a d50 less

than 15 μm and a PSD curve slope 𝑑𝑠 of > 3.26 will exhibit homogeneous fluid

behaviour.

95

6.4 7th

International Conference for Conveying and Handling of Particulate

Solids-ChoPS (2012) - Comparison between Flow Cones and a Rotary Viscometer

The building of modern high speed motorways requires construction on stable

foundations and when old mine working are located under new road construction sites.

It is common practice to fill the mine working with cemented fly ash grout. The coal

from these mines was removed using the “Bord and Pillar” extraction technique leaving

open underground roadways and pillars to support the roof. Most of these mines are

now flooded and there is no access underground, therefore holes are drilled from the

surface into the old roadways where the voids can be filled with grout consisting of

power station fly ash, cement and water. The mixture of fly ash, cement and water has

to give strength greater than 1 MPa after 28 days curing. A grout with this strength is

enough to provide stable foundation for the construction of the motorway. Prior to the

grout being batch pumped underground a flow cone was employed to determine the

flow properties.

This paper investigated the flow properties of fly ash grouts using an ASTM Flow

Cone, a Marsh Funnel (Flow Cone) and a Rotary Viscometer. A transportable grout

plant was established on the surface adjacent to old underground workings. It consisted

of separate silos for the fly ash and cement, a batch mixing plant and a high pressure

pump.

The fly ash and cement were delivered to site in separate sealed trucks and

pneumatically unloaded into the respective silos. The maximum rate that fly ash tankers

can deliver and unload into the silo was approximately 40 m3

h-1

. The grout was a

mixture of fly ash, water and 7 % cement. Prior to placing it underground it was

manually tested, as specified, using a flow cone with a nominal flow cone time greater

than 20 seconds. However, this specification does not account for variations in the

particle size distribution of the different fly ashes and, therefore the percentage of water

to achieve the 20 second flow cone or the type of flow cone to be used.

Wedmore (2011) indicated that the PSD of the Bayswater “run of station fly ash” as

variable and unpredictable. As an example one week in November 2011 the weight of

96

water required to batch a 2-ton mixture of fly ash and cement varied from 800 kg to

1200 kg to achieve a flow cone time of 20 seconds. Therefore, the Cw of the grout

pumped varied between 71.5 % and 62.5 % at a similar viscosity.

6.4.1 Particle Size Distribution and Density

Fly ash which was collected from four different Australian power stations was tested for

particle size distribution using a laser diffraction technique (Malvern Particle Size

Analyser). The results are shown in Figure 6.10. The solids density of the fly ashes was

tested with a Micromeritics AccuPyc Pycnmoter 1330. Table 6.4 displays the results.

Figure 6.10 Particle Size Distributions of the Four Fly Ashes.

Table 6.4 Density of the Four Fly Ashes.

Density BW CL SB SW

t m-3

2.1129 2.1583 2.3961 2.0532

t m-3

2.1282 2.1585 2.3945 2.0511

t m-3

2.1278 2.1581 2.3941 2.0497

Average t m-3

2.1230 2.1583 2.3949 2.0513

0

10

20

30

40

50

60

70

80

90

100

0.1 1 10 100 1000

Per

cen

tage

Pas

sin

g (%

)

Particle Size (µm)

BW CL SW SB

97

6.4.2 Methodology

Grout tests were carried out using an ASTM Flow Cone, a Marsh Funnel and a Rotary

Viscometer. The ASTM Cone was 178 mm across the top and 190 mm to the apex to

which is fitted a 30.8 mm long tube with a diameter of 12.7 mm. At the top of the cone

was a cylinder of internal diameter 178 mm and 75 mm high. The 1725 ml level was at

a point indicated by the adjustable marker.

The Marsh Funnel consisted of a cone 152 mm across the top and 305 mm to the apex

to which was fixed to a tube 50.8 mm long with an internal diameter of 12.7 mm. A

mesh was fixed 25.4 mm from the top across half the cone and the grout to be measured

was poured through the mesh. When the liquid reached the mesh it gave a volume of

1500 ml.

The ASTM Flow Cone and Marsh Funnel were initially tested to determine the flow

time with water. The ASTM Flow Cone was tested as per ASTM C939 – 10 Standard

Test Method for Flow of Grout for Preplaced-Aggregate Concrete (Flow Cone Method).

The Marsh Funnel was tested using the same procedure but only passing 1500 ml of

water. The average time for the ASTM Flow Cone was 8.02 seconds was within the

specification outline in ASTM C939. No such standard exists for the Marsh Funnel but

the average time recorded was 6.39 seconds. A photograph of the flow cones is shown

in Figure 6.11.

A grout was prepared, by adding 3000 grams of the designated fly ash to a 3000 ml

beaker of water, to the consistency of thick honey using a variable speed laboratory

mixer for approximately 10 minutes. To determine the flow time for the ASTM Flow

Cone, 1725 ml of the grout was passed through and the time recorded. The collected

grout was returned to the mixing beaker and mixed for another 5 minutes. The same

procedure was followed using the Marsh Funnel but only 1500 ml of the grout was

passed through the Marsh Funnel then the collected grout was the returned to the

mixing beaker and mixed for 5 minutes.

98

A sample of approximately 200 grams of grout was then placed in a Contraves Rotary

Viscometer for shearing, placed in a Petri Dish, weighed and dried and re-weighed to

determine the sample Cw. Water was then added to the mixing beaker to reduce the

solids concentration by weight (Cw) for the next test. The slurry test was repeated

several times at different Cw’s.

Figure 6.11 Marsh Funnel and ASTM Flow Cone.

6.4.3 Results and Discussions

Figures 6.12 to 6.15 are the Rheograms for rotary viscometer results for different fly ash

grouts at different Cw’s. Figure 6.16 is the graph of the ASTM Flow Cone and the

Marsh Funnel flow cone times for different fly ash grouts at different Cw’s.

The Rheograms indicate there is significant difference in the rheological properties of

the grout mixed from different fly ashes. If the fly ash from power station CL was used

Marsh Funnel ASTM Flow Cone

99

with a 20 second ASTM Flow Cone with a Cw of approximately 73 %, this equates to

730 kg of fly ash to 270 kg of water per ton of grout. Whereas if the fly ash was sourced

from power station SB to provide a 20 second ASTM Flow Cone would require a grout

with a Cw of 53 % which equates to 530 kg of fly ash to 470 kg of water. The question

arises as to what effect this extra water would have on the strength of the grout.

The results as shown in Figure 6.16 indicate that there is a difference in the flow times

between the ASTM Flow Cone and the Marsh Funnel. However, it is only marginal

compared to the possible difference in grout strength which is influence by the source of

the fly ash.

The implication of these results is that if the incorrect fly ash or flow cone was used it

would impact on the economic cost of providing stable foundations for motorway

construction.

Figure 6.12 Rheogram CL Fly Ash.

0

20

40

60

80

100

120

140

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1 )

Cw = 75.4 % Cw = 73.9 % Cw = 72.8 %

Cw = 72.6 % Cw = 71.7 % Cw = 69.5 %

100

Figure 6.13 Rheogram BW Fly Ash.

Figure 6.14 Rheogram SW Fly Ash.

0

20

40

60

80

100

120

140

0 20 40 60 80 100 120 140 160 180 200

Shar

t St

ress

( P

a)

Shear Rate (s-1 )

Cw = 70.6 % Cw = 70.1 % Cw = 68.8 % Cw = 67.6 % Cw = 66.4 % Cw = 65 %

0

20

40

60

80

100

120

140

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

Cw = 62.8 % Cw = 62.4 % Cw = 60.6 % Cw = 58.5 %

101

Figure 6.15 Rheogram SB Fly Ash.

Figure 6.16 Flow Cone Results.

0

20

40

60

80

100

120

140

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

Cw = 53.8 % Cw = 52.8 % Cw = 52.3 % Cw = 51.3 % Cw = 50.4 %

0

5

10

15

20

25

30

35

40

45

50 55 60 65 70 75

Flo

w T

imes

(se

con

ds)

Cw (%)

ASTM BW Marsh Funnel BW ASTM CL Marsh Funnel CL

ASTM SW Marsh Funnel SW ASTM SB Marsh Funnel SB

ASTM Water 8.02

Marsh Funnel Water 6.39 seconds

102

6.5 International Freight Pipeline Society Symposium (2011) - The Pumping

Characteristics of Fly Ash Slurry in a Pipeline

Does the pumping characteristics of a pipeline pumping high concentration fly ash

slurry change due to shearing in the pipeline? This paper examines the pumping

characteristics of fly ash slurry in a 10 km pipeline by simulating the pipeline

characteristics in the laboratory using a vertical mixer and a Rotary Viscometer.

The pumping of power station ashes as high concentration slurry has become common

place in Australian black coal fired power stations (Ward et al., 1999). At Bayswater

Power Station high concentration fly ash slurry, which is a combination of station

process water and fly ash mixed in a pug mill and mixing tank, and pumped at a rate of

240 m3

h-1

through a 200 mm pipeline using a triplex diagram pump to a disposal site 10

km from the station. At a flowrate of 240 m3

h-1

the pseudo shear rate (8𝑉

𝐷) in the

pipeline is 85 s-1

and a single particle of fly ash takes approximately 2 hours to pass

from the fly ash silo to the disposal site. Changes in coal properties can lead to

significant variations in the pumping characteristics of fly ash slurries (Bunn et al.

2004). Senapati et al. (2010) also indicated that changes in particle size distribution of

Indian fly ash slurries can lead to more than doubling of the shear stress at similar shear

rates at the same Cw.

A change in coal properties can alter the pumping characteristics by an order of

magnitude, or a reduction of up to 10 % in the pumping Cw at similar pipeline pressures.

These changes were attributed to the differences in particle size distribution of the fly

ash from the different coal seams. However, by maintaining the fly ash particle size

distribution of the slurry, any changes that occur can be attributed to pipeline shearing

or chemical reaction. To observe if changes were due to chemical reactions different

water samples were used.

To simulate the shearing in the pipeline the slurry was continuously sheared using a

laboratory mixer. When the fly ash was mixed with water at a high concentration and

sheared in a rotary viscometer, the resulting Rheogram indicated classical pseudoplastic

(shear thinning) behaviour (Senapati et al., 2010 and Bunn et al., 1991).

103

Fly ash “A” was from a coal seam that was classified as acid coal, whereas fly ash “B”

was from a coal seam classified as a basic coal. Bunn (1991) examined several different

fly ashes from coal types classified as acid and basic coal to determine the equilibrium

PH. If was found that after 2 hours the acid coals had a PH in the range of 3.5 to 4,

whereas the basic coal had a PH in the range of 11 to 11.5.

6.5.1 Methodology

Fly ash samples were procured from two Australian black coal power stations

designated as (A & B) and process water was obtained from three different Australian

power stations. The process waters were designated R, W & K. Process water A and K

were from fresh water cooling stations whereas process water W was from a salt water

cooled station. Demineralised water D and local tap water T were also used. No

chemical analysis of the different process waters was obtained.

The particle size distributions (PSD) were determined for the fly ash samples using a

laser diffraction technique (Malvern Particle Size Analyser). The PSD’s of the fly ashes

are shown in Figure 6.17.

Slurries with a Cw of 68 % were mixed by adding 808 grams of water to a 2 litre

container and then progressively adding 1712 grams of fly ash while mixing with a

vertical mixer. After 10 minutes of mixing a sample of approximately 200 grams of the

slurry was the placed in the Rotary Viscometer cup. The slurry was then sheared in the

rotary viscometer (Figure 6.18). After shearing the slurry from the rotary viscometer

was removed and the cup and bob cleaned. The remaining slurry was left shearing in the

mixer until the next test as shown in Figure 6.19. Samples of the slurry were sheared in

the rotary viscometer every half hour for two hour. All mixing and testing was carried

out at 21 degrees Celsius.

104

Figure 6.17 Particle Size Distribution of the Fly Ash.

Figure 6.18 Shearing Fly Ash Slurry in a Rotary Viscometer.

0

10

20

30

40

50

60

70

80

90

100

1 10 100 1000

Per

cen

tage

pas

sin

g (%

)

Particle Size (um)

Fly Ash A Fly Ash B

105

Figure 6.19 Fly Ash Slurry shearing in Mixer.


The d50 of fly ash “A” was 14.9 µm and fly ash B was 37.8 µm. Figure 6.20 is a typical

Rheogram of fly ash slurries when mixed a Cw of 68 %.

Table 6.5 shows the shear stress at a shear rate of 85 s-1

for the fly ash slurries at a Cw of

68 %. Table 6.6 is the calculated pipeline pressure drops at a shear rate of 85 s-1

for the

fly ash slurries at a Cw of 68 %.

The results of the Rheograms show that all the slurries indicate a classical pseudoplastic

(shear thinning) behaviour. The pipeline pressures calculated at a shear rate of 85 s-1

are

shown in Table 6.6 and Figure 6.21. The results for all the slurries mixed with the

different waters and fly ash B (basic coal fly ash) showed very little variation in

calculated pipeline pressure at a constant shear rate of 85 s-1

. The results for slurries

106

mixed with demineralised water, tap water and process water R and fly ash B (acid coal

fly ash) also showed very little variation in shear stress at a constant shear rate of 85 s-1

.

Figure 6.20 Typical Rheogram of Fly Ash Slurry.

Table 6.5 Shear Stress at a Shear Rate of 85 s-1

and Cw of 68 %.

Time hours 0 0.5 1 1.5 2 0 0.5 1 1.5 2

Fly Ash A Fly Ash B

Shear Stress Pa

Demineralised

Water D 14.9 15 16 17 17.5 10.5 10.5 10.5 10.5 11.5

Tap Water T 14.2 15 15 17.4 18.5 10.5 10 10.5 10 10.8

Process Water R 14.2 14.9 15.5 17.4 18.5 12.5 9.8 9.8 10.2 11.5

Process Water W 24.5 22.5 21.5 20.5 20.5 9 9 9.2 10 10.2

Process Water K 19 18.2 19 20 21.5 10.2 9 10.8 10.9 10.9

0

5

10

15

20

25

30

35

40

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

Time 0 hr Time 0.5 hr Time 1hr Time 1.5 hr Time 2 hr

107

The results for the slurries mixed with process waters K and W and fly ash B (acid coal

fly ash) showed a greater variation than the slurries mixed with the other waters. The

slurry from the mixing of process water K and fly ash A showed the same

characteristics as the other 3 slurries, i.e. a rising pipeline pressure over time, whereas

the slurry mixed with process water W and fly ash B showed a falling pipeline pressure

over time. The pumping characteristics of high concentration fly ash slurry pipeline

change due to shearing in the pipeline depends on the type of coal the power station

burns and the properties of the process water mixed with it. The change in pumping

characteristics due to the shearing in the pipeline is insignificant compared with the

changes in pumping characteristics due to changes in particle size distribution.

Table 6.6 Calculate Pipeline Pressure at a Shear Rate of 85 s-1

and Cw of (68) %.

Time hours 0 0.5 1.0 1.5 2.0 0 0.5 1.0 1.5 2.0

Pipeline Pressure (MPa)

Demineralised Water

D 3.0 3.0 3.2 3.4 3.5 2.1 2.1 2.1 2.1 2.3

Tap Water T 2.8 3.0 3.0 3.5 3.7 2.1 12 2.1 2.0 2.2

Process Water R 2.8 3.0 3.1 3.5 3.7 2.5 2.0 2.0 2.0 2.3

Process Water W 4.9 4.5. 4.3 4.1 4.1 1.8 1.8 1.8 2.0 2.0

Process Water K 3.8 3.6 3.8 4.0 4.3 2.0 1.8 2.2 2..2 2.2

108

Figure 6.21 Calculate Pipeline Pressure at a Shear Rate of 85 s-1

for Fly Ashes A and B

Slurries.

0

1

2

3

4

5

6

0 0.5 1 1.5 2

Pip

elin

e P

ress

ure

(M

Pa)

Time (hr) Fly Ash A & Water D Fly Ash A & Water T Fly Ash A & Water R

Fly Ash A & Water W Fly Ash A & Water K Fly Ash B & Water D

Fly Ash B & Water T Fly Ash B & Water R Fly Ash B & Water W

Fly Ash B & Water K

Fly Ash A

Fly Ash B

109

6.6 International Seminar on Paste and Thickened Tailings (2010) - Pumping

Power Station Ash as a High Concentration Slurry

Determining the pumping characteristics of a full scale power station ash disposal

system is fraught with difficulties. This paper examines the pumping characterises of

high concentration slurry comprising of a mixture of a power station ash and process

water as determined using a rotary viscometer and small scale pumping test rig.

The pumping of power station ashes as high concentration slurry has become common

place in Australian black coal fired power stations (Ward et al., 1999). The fly ash was

collected from either the precipitator or fabric filter hopper and was conveyed either by

chain scraper and bucket elevator or pneumatically to a fly ash silo. The bottom ash was

collected by either a wet or dry scraper conveyor. The bottom ash that was collected

from a wet scraper conveyor was transferred to a bottom ash silo after being crushed.

Alternatively, the bottom ash was collected by a dry scraper conveyor, crushed and

pneumatically conveyed directly to the fly ash silo. The fly ash, bottom ash and water

were then mixed together in a pug mill and transferred to a mixing tank before being

pumped to a disposal site using high pressure positive displacement pumps. The

pipeline to the disposal site was usually constructed of steel and was either welded or

joined with high pressure couplings. However, at some disposal sites, polyethylene

pipes are used to transfer the slurry to multi-disposal points.

This study looked at the comparison of results of shearing high concentration fly ash

slurry in a rotary viscometer with the pumping of slurry consisting of fly ash and bottom

ash using a 50 mm pilot pumping plant. After a rotary viscometer study was undertaken

it was later followed by the pumping trial. The studies were undertaken to determine the

pressure drop characteristics of pumping a mixture of fly ash and bottom ash using a

twin-cylinder piston pump fitted with “S” transfer tube output system in a 150 mm

pipeline with a flowrate of 100 m3

h-1

at a maximum pressure of 6 MPa to disposal 5 km

from the power station. The velocity of the proposed pipeline was 1.57 m s-1

. The

calculated pseudo-shear rate of proposed high concentrated slurry disposal was 84 s-1

.

To reflect the pulse pressure characteristics of the piston pumps a hose pump was used

in the pilot plant.

110

Agreement between shear rate (measured using a rotary viscometer) and the pseudo-

shear rate (8𝑉

𝐷) in a pipeline was noted by Bunn (1991) while conducting a pumping

trial using a rotary ram pump. The trials consisted of pumping fly ash slurries from

three different Hunter Valley Power Stations with the rotary ram pump and

simultaneously shearing the slurries in a rotary viscometer. Only the fly ash slurry from

Bayswater Power Station showed a similarity between the shear rate (measured by a

rotary viscometer) and the pseudo-shear rate measured in the pipeline. In the case of the

other fly ashes the rotary viscometer measurements underestimated the shear stress in

the range of 50% to 400 %. This however depended on the source of the fly ash and the

Cw of the slurry as the higher the Cw the greater the underestimation. This led to the

assumption that the fly ash slurries from Bayswater Power Station were homogeneous

whereas the slurries from the other power stations were heterogeneous.

6.6.1 Methodology

Several 25-litre containers were used to collect samples of fly ash, bottom ash and

process water from an Australian black coal power station. It was assumed that the

samples were typical of the station ash and process water. A sample of fly ash was sent

for Scanning Electron Microscope (SEM) analysis. Elemental and Atomic Analysis are

shown in Tables 6.7 and 6.8. SEM photographs at different magnifications are seen in

Figures 6.22 and 6.23.

Table 6.7 Elemental Analysis of Fly Ashes.

Ash

No.

No.

No.

Elemental Analysis %

Na Mg Al Si S K Ca Ti Fe

1 0.98 0.36 33.03 58.62 1.71 0.94 2.47 1.58 1.02

2 0.87 0.43 29.32 53.14 1.48 1.29 4.28 3.21 5.99

3 0.87 0.37 29.29 52.48 1.17 1.58 4.26 3.22 7.5

4 0.93 0.36 29.23 53.02 1.68 1.14 4.25 2.98 6.41

111

The PSD of the fly ash sample was determined using a laser diffraction technique

(Malvern Particle Size Analyser). A 2 kg sample of bottom ash was then dried and

sieved. This PSD of the fly ash and bottom ash is shown in Figure 6.24.

Table 6.8 Atomic Analysis of Fly Ashes.

Ash

No.

Atomic Analysis %

Na Mg Al Si S K Ca Ti Fe

1 1.21 0.42 34.68 59.13 1.51 0.68 1.74 0.94 0.52

2 1.11 0.52 32.01 55.74 1.36 0.97 3.15 1.98 3.16

3 1.12 0.45 32.3 55.59 1.09 1.20 3.16 2.0 4.0

4 1.19 0.44 31.94 55.67 1.54 0.86 3.12 1.84 3.39

\

Figure 6.22 SEM Photograph of Fly Ash.

112

The solids density of the fly ash and bottom ash from several different buckets was

tested with a Micromeritics AccuPyc Pycnmoter 1330. The data is displayed in Table

6.6. The d50 of the fly ash was 40 µm and of the bottom ash was 300 µm. The average

solids density of the fly ash was 1813.8 kg m-3

while the average solids density of the

bottom ash was 1859.8 kg m-3

.

Figure 6.23 SEM Photograph of Fly Ash.

Table 6.8 Density of Fly Ash and Bottom Ash.

Test Fly Ash

1

kg m-3

Fly Ash

2

kg m-3

Fly Ash

3

kg m-3

Fly Ash

4

kg m-3

Bottom Ash

Sample A

kg m-3

Bottom Ash

Sample B

kg m-3

1 1768.0 1798.0 1848.8 1840.7 1863.3 1852.7

2 1767.3 1797.7 1848.5 1840.2 1867.1 1854.0

3 1767.1 1797.4 1846.2 1839.4 1867.8 1853.9

Average 1767.5 1797.7 1847.8 1840.1 1866.0 1853.5

113

Figure 6.24 Particle Size Distributions of the Fly Ash and Bottom Ash.

6.6.2 Fly Ash Testing with Rotary Viscometry

Rheology tests were conducted using a Contraves Rotary Viscometer as described in

(Bunn 1991).

6.6.3 Pilot Pumping Plant

A layout drawing of the pilot plant used is shown in Figure 6.24 and a picture of the

pump is shown in Figure 6.25.

The pilot plant consisted of a variable speed Bredel SPX100 hose pump with a

maximum flowrate of 24 m3

h-1

at a pressure of 500 kPa and a horizontal 50 mm

pipeline. The hose pump was calibrated using water with the classical stopwatch and

bucket technique. It compared favourably with the manufacturer’s performance chart.

The capacity versus flow rate calibration is shown in Table 6.9.

0

10

20

30

40

50

60

70

80

90

100

1 10 100 1000 10000 100000

Per

cen

tage

pas

sin

g (%

)

Particle Size (µm) Fly ash Bottom ash

114

The pilot plant instrumentation consisted of four pressure transmitters and a pipeline

temperature transmitter. A data collector and laptop computer were used.

Figure 6.24 Diagram of the Pilot Pumping Plant.

Figure 6.25 Bredel SPX100 Hose Pump.

50 mm ID Steel Pipe Stirrer

Pump Flexible Suction &

Discharge Hoses

Bredel SPX100 Hose Pump

Variable

Speed Drive

Variable

Speed Drive

Supports

Glass Tube

Pressure Transmitters

5350 mm apart

Valve

P1 P2

P3 P4

115

Table 6.9 Comparison of Speed Verses Flow for the Bredel SPX100 Hose Pump.

Potentiometer

Positions

Digital

Readout

Flow

m3

h-1

1 5.0 2.2

2 9.1 4

3 13.5 6

4 18 8

5 25.2 10

6.6.4 Slurry Mixing and Pumping

In modern coal fired boiler units, the typical collection ratio of fly ash to bottom ash is

85:15. Initially, pumping tests were conducted on slurry at this ratio and at different

Cw’s. Additional bottom ash was then added to change the ratio to 80:20 and the slurry

was also tested at different Cw’s. Prior to pumping the pipeline and test rig hopper was

filled with water, the pump started and the pipeline flushed. On completion of flushing,

the hopper was emptied to a level just below the hopper outlet valve and the pump and

pipeline was left full of water. The hopper outlet valve was closed. The test rig hopper

electrical stirrer was started and the slurry mixed in external mixer was added to the

hopper. When approximately 110 litres of slurry was added to the hopper, the hopper

outlet valve was opened and the hose pump was started on low flow. A temporary

flexible hose was used to discharge water in the pipeline into a 200-litre drum. When

slurry appeared the hose pump was stopped and the flexible hose removed. The hose

pump was then restarted and set to full speed. The slurry was circulated until uniform

data was obtained. The data was then was recorded for 2 to 3 minutes. At this time a

sample of the slurry was removed for Cw verification. The same procedure was followed

for the remaining four potentiometer positions. Slurry discharging into the hopper is

shown in Figures 6.26. On completion of the final pumping run, the pump was stopped,

the hopper was filled with water and the pipeline discharge was again diverted to the

200-litre drum for environmental friendly disposal.

116

The pilot plant pressure transducers with a range of 0 to 250 kPa with an accuracy of

0.25 % FSO. The Cw was determined using oven drying technique and scales with and

accuracy to 0.001 grams.

Figure 6.26 Slurry Pumping.


Figure 6.27 shows the pressure difference between pairs of pressure transmitters in the

pipeline. Figure 6.28 and 6.29 are graphs at position A and B on Figure 6.27. The

fluctuation in pressure demonstrated the pumping characteristics of a Bredel SPX hose

pump. The dashed lines indicate the pressure used in the calculation for the shear

stresses at shear rates between 45.63 to 194.37 s-1

.

117

Table 6.10 is a table of the flow diagram results. Figure 6.30 is a Rheogram of the

results fly ash slurry tested with the rotary viscometer. Figure 6.31 is a pseudo-shear

diagram of the pumping trials conducted with fly ash and bottom ash slurry. The

rheogram of the fly ash slurry as determined by the rotary viscometer and the flow

diagram of the fly ash and bottom ash slurry obtained from the pilot pumping plant

indicates the pumpability of the slurry at different Cw’s. The SEM photographs indicate

the lack of diversity of fly ash particles sizes, that is, there was lack of intermediate

sized particles between the larger and smaller particles. To enable the pumping of

slurries at high concentration, the void spaces between different particles needs to filled

with particles, rather than water. Therefore, a diversity of different size particles is

required.

Figure 6.27 Pressures between the Pairs of Pressure Transmitters.

0

5

10

15

20

25

30

35

40

45

50

Dif

fere

nti

al P

ress

ure

(kP

a)

Pressure Drop Between P3 and P4 Pressure Drop Between P1 and P2

A

B

118

Figure 6.28 Pressure at point “B” on Figure 6.15.

Figure 6.29 Pressure at point “A” on Figure 6.15.

0

5

10

15

20

25

30

35

40

45

50

Dif

fere

nti

al P

ress

ure

(kP

a)

Pressure Drop Between P3 and P4

0

1

2

3

4

5

6

7

Dif

fere

nti

al P

ress

ure

(kP

a)

Pressure Drop Between P1 and P2

119

Table 6.10 Flow Diagram Results.

Shear Rate

s-1

Shear Stress

Pa

46 19.66 15.98 14.75 13.52 13.52

78 41.79 33.18 27.04 26.55 24.58

117 73.74 54.07 46.70 41.79 39.33

156 120.44 82.34 66.36 61.45 56.53

194 196.64 130.27 100.78 89.72 86.03

FA:BA Ratio 85:15 85:15 85:15 80:20 80:20

Cw % 71 68 68 65 64

Figure 6.30 Shear Diagram of Fly Ash Slurry from Rotary Viscometer.

0

5

10

15

20

25

30

35

40

45

50

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

Cw = 71.0 % Cw = 70.7 % Cw = 69.5 % Cw = 68.9 % Cw= 67.6 % Cw = 66.4 %

120

Figure 6.31 Pseudo-shear Diagram for Fly Ash and Bottom Ash Slurry.

In the proposed pumping system with a maximum pipeline operating pressure of 3.5

MPa, the shear stress was calculated to be 26 Pa. On examining the rheogram developed

from the rotary viscometer testing at a shear rate of 84 s-1

and a shear stress of 26 Pa the

rheogram indicated that fly ash slurry could be pumped with a Cw in excess of 71%.

Examining the pseudo-shear diagram developed from the pilot pumping plant at a shear

rate of 84 s-1

and a pseudo shear stress of 26 Pa, the flow diagram indicated that a the

maximum Cw of a mixture ratio of fly ash to bottom ash of both 85:15 or 80:20 slurry

could only be pumped at a Cw of 64 %. The explanation for this is that the fly ash and

fly ash and bottom ash slurries are heterogeneous slurries; therefore the rotary

viscometer underestimated the pumpability of the slurry as rotary viscometers are only

accurate when the slurry is homogenous.

This again leads to the discussion: how do we determine if fly ash slurry is homogenous

or heterogeneous?

0

20

40

60

80

100

120

140

160

180

200

0 20 40 60 80 100 120 140 160 180 200

She

ar S

tre

ss (

Pa)

Psuedo-shear Rate (s-1)

FA/BA 85/15 Cw = 70.5 % FA/BA 85/15 Cw = 68 % FA/BA 85/15 Cw = 67.7 %FA/BA 80/20 Cw= 64.4 % FA/BA 80/20 Cw = 65.1 %

121

6.7 6th

World Congress on Particle Technology (2010) - Thixotrophic Behavior

of Fly Ash Slurries

Fly ash slurries have been labelled as both thixotrophic and shear thinning (Naik el al.

2009, Senapati el al. 2010). Wikipedia (2010) describes thixotrophy as “the property of

fluids that are viscous under normal conditions, but change due to flow, when agitated

or otherwise stressed”. A thixotrophic fluid is also a fluid which takes a finite time to

attain equilibrium viscosity when introduced to a step change in shear rate.

The distinction between a thixotrophic fluid and a shear thinning fluid is that:

a thixotrophic fluid displays a decrease in viscosity over time at a constant shear

rate while;

a shear thinning fluid displays decreasing viscosity with increasing shear rate.

Some non-Newtonian pseudoplastic (shear thinning) fluids show a time-dependent

change in viscosity. That is, the longer the fluid undergoes shear stress the lower its

viscosity.

Mixing fly ash slurries at a high concentration and shearing in a rotary viscometer

resulted in a Rheogram indicating a classical pseudoplastic behaviour (Bunn el al.,

1990). The pumping of power station ashes as high concentration slurry has become

common place in Australian black coal fired power stations (Ward el al., 1998).

At Bayswater Power Station, high concentrated fly ash slurry was pumped at a rate of

240 m3 h

-1 through a 200 mm pipeline to a disposal site 10 km from the station.

At 240 m3 h

-1 flowrate the shear rate in the pipeline is 85 s

-1 and a single particle of fly

ash takes less than 2 hours to pass from fly ash silo to the disposal site. The slurry is

mixed in a pug mill and mixing tank before being pumped in a triplex diagram pump to

the disposal site.

http://en.wikipedia.org/wiki/Shear_thinning


http://en.wikipedia.org/wiki/Non-Newtonian_fluid

http://en.wikipedia.org/wiki/Power-law_fluid


http://en.wikipedia.org/wiki/Viscosity

http://en.wikipedia.org/wiki/Shear_stress


122

6.7.1 Methodology

Two fly ash samples (designated A and B) and process water from three different

Australian power stations (designated R, W & K), as well as demineralised water (D)

and local tap water (T) were procured.

The particle size distributions (PSD) were determined for the fly ash samples using laser

diffraction technique (Malvern particle size analyser). The d50 of fly ash A was 14.9 µm

and fly ash B was 37.8 µm. Figure 6.32 shows the particle size distribution for fly ash A

and B.

Figure 6.32 PSD of Fly Ash A and B.


Figure 6.33 is a typical rheograms of the fly ash slurries. Figures 6.34 and 6.35 are

graphs of viscosity at different shear rates of slurries from mixtures of fly ash A & B

mixed with demineralised water (D).

0

10

20

30

40

50

60

70

80

90

100

1 10 100 1000

Per

cen

tage

Pas

sin

g (%

)

Particle Size (µm)

Fly Ash A Fly Ash B

123

Figure 6.33 Typical Rheogram.

Tables 6.11 and 6.12 show the shear stress of the slurry at half hour time intervals with

a shear rate of 85 s-1

. The shear stress verses time data from tables 1 and 2 are plotted in

Figure 6.36.

Examination of the Rheograms and viscosity charts indicate that these fly ash slurries

exhibit classic pseudoplastic shear thinning behaviour, that is, as the shear rate

increased, the viscosity decreased. At a constant shear rate of 85 s-1

, 9 out of 10 slurries

tended to exhibit some rheopectic behaviour, that is, there was a slight increase in shear

stress at a constant shear rate. The 10th slurry consisting of a mixture of fly ash B and

water W exhibited slight thixotrophic behaviour. For pumping of these high Cw fly ash

slurries, there would be little change in the pipeline pressure drop per unit length due to

thixotrophic behaviour.

0

5

10

15

20

25

30

35

40

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(Pa)

Shear Rate (s-1)

Time 0 hr Time 0.5 hr Time 1hr Time 1.5 hr Time 2 hr


124

Figure 6.34 Fly Ash “A” Viscosity verses Shear Rate.

Figure 6.35 Fly Ash B Viscosity verses Shear Rate.

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

0 20 40 60 80 100 120 140 160 180 200

Vis

cosity (P

as)

Shear Rate (s-1)

Time 0 hr Time ½ hr Time 1 hr Time 1½ hrs Time 2 hrs

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 20 40 60 80 100 120 140 160 180 200

Vis

cosity (P

as)

Shear Rate (s-1)

Time 0 hr Time ½ hr Time 1 hr Time 1½ hrs Time 2 hrs

125

Table 6.11 Fly Ash “A” Shear Stress. Table 6.12 Fly Ash “B” Shear Stress.

Time hours 0 1 1½ 2 Time hours 0 1 1½ 2

H2O Shear Stress Pa H2O Shear Stress Pa

Dem 14.9 16.0 17.0 17.5 Dem 10.5 10.5 10.5 11.5

Tap 14.2 15.0 17.4 18.5 Tap 10.5 10.5 10.5 10.8

R 14.2 15 17.4 18.5 R 12.5 9.8 10.2 11.5

W 24.5 21.5 20.5 20.5 W 9.0 9.2 10.0 10.2

K 19.0 19.0 20.0 21.5 K 10.2 10.8 10.9 10.9

Figure 6.36 Fly Ashes A & B and Different Waters Shear Rate over Time.

0

5

10

15

20

25

30

0 0.5 1 1.5 2

Shea

r St

ress

(P

a)

Time (hr)

Fly Ash A & Water D Fly Ash A & Water T Fly Ash A & Water R Fly Ash A & Water W Fly Ash A & Water K

Fly Ash B & Water D Fly Ash B & Water T Fly Ash B & Water R Fly Ash B & Water W Fly Ash B & Water K

Fly Ash A

Fly Ash B

126

6.8 The 6th

International Conference for Conveying and Handling Particulate

Solids and 10th

International Conference on Bulk Materials Storage, Handling and

Transportation (2009) - Are Tailing Dams Viable in the Modern Environment?

Tailing dams are normally built to contain rejects from mining, mineral processing and

power generation and have been an essential part of the minerals extraction process.

However, history shows that serious environmental and safety issues are associated with

tailing dams. There are options to the conventional tailing dams that may provide

solutions to the problems experienced with the dams and, although, economically more

costly in the short term, may be economically viable in the longer term.

Tailings dams are an economic solution to the management of refuse. The economic

cost of tailings dams is somewhere in between $1 and $5 a tonne of tailings deposited.

However, it could be argued that the cost is actually between $2 and $10 a tonne

depending on the circumstances and estimated indirect costs. Indirect costs include

amenity (physical and visual), ongoing insurance, monitoring, groundwater

contamination, dust contamination and loss of real estate value in areas on and around

the tailings dams (e.g. reluctance to build an agricultural industry in the valley below a

tailings dam).

The main hazard the dams present is an unacceptably high historical rate of failure

which typically cause substantial losses, including loss of lives. The failures occur due

to:

Inadequate design and or construction;

Rainfall events in excess of the design allowances;

Seismic activity causing re-liquefaction.

The ICOLD Committee on Tailing Dams and Waste Lagoons (1995-2001) has

developed guidelines for the safe design, construction and closure of tailings dams. To

reduce the degree of dam failure guidelines for a dam’s can be found in publications

such as ICOLD Bulletins Nos. 45 (1982), 74 (1989), 97 (1994), 98 (1995), 101 (1995),

103 (1996), 104 (1996), 106 (1996) and ANCOLD (1999).

127

It is of major concern that tailing dam’s failures continue at a high rate. Unfortunately,

the number of major incidents continues at an average of more than one a year. In the

whole during the last 6 years doubled. While tailings dams are considered permanent

fixtures in the environment, past experience shows that minor and major spills pose a

serious environmental threat that stay behind long after mine closure. A number of

characteristics make tailings dams more vulnerable than other types of retention

structures (e.g. water retention dams), namely:

embankments formed by locally collected fills (soil, coarse waste, overburden

from mining and tailings);

dams subsequently raised with an increase in solid material and therefore

effluent;

lack of regulations on specific design criteria;

lack of dam stability requirements regarding continuous monitoring and control

during emplacement, construction and operation;

high cost of maintenance works for tailings dams after closure of mining

activities;

mining industry changes mean the rates of refuse vary with market conditions

(due to changes in yields from process plant and capacity of process plant), this

means the planning of dam raisings is often lacking during a cyclical mining

boom;

changes in mining and processing techniques are always occurring, and again

the planning of dam raisings is often lacking due to unexpected capacity

changes.

A major historic factor in dam failures was rainfall followed by occurrences associated

with seismic liquefaction. Over 90% of dam failure incidents occurred in active mine

tailings dams while only 10% are associated with abandoned dams. The number of

reported incidents throughout the world involving tailing dams was 221, ICOLD (2001),

resulting in 147 tailing dam failures occurring Rico (2008).

Due to the nature of mining and mineral processing, the volumes of mining wastes are

significantly larger than those of both domestic and industrial wastes. The material

128

stored in tailings dams is usually very fine and loose, placed there hydraulically, and is

at, or above, saturation. Any major movement of the retaining boundaries of the

impoundment can induce shearing strains that disturb the structure of the tailings mass,

inducing a rapid rise of pore water pressures and liquefaction of a section of the

impoundment. An event like this can cause even greater pressures to be applied to the

retaining boundaries. Failure of the tailings dam can release liquefied tailings that can

travel for great distances, and, because of its specific weight, destroys everything in its

path. Unlike water that will flow through and around buildings, liquefied tailings can

destroy structures. Historically the tendency is for tailing dams to become ever higher

and impoundments even larger. Table 6.13 list some of the dam failure over recent

years.

Table 6.13 Examples of Tailing Dam Failures.

Date Location Material Results

May 2009 Huayuan County, China Manganese tailings 3 killed

December 2008 Kingston fossil plant,

Harriman, Tennessee, USA

Coal ash 4.1 m3 released covering an

area of 1.6 km2 to a deep 1.83

meters

September 2008 Taoshi, Linfen City, China Iron ore tailing 245 killed, 43 injured

April 2006 Miliang, China Gold mine tailings Toxic potassium cyanide

released into the Huashui river

August 2002 Dizon Copper Silver Mines,

Zambales, Philippines

Copper & Silver tailings 1,000 families evacuated

June 2001 Mineração Rio Verde Brazil Iron ore tailings 5 killed

October 2000 Martin Country Coal

Corporation, Kentucky, USA

Coal waste slurry 950,000 m3 released killing fish

in Tug River and drinking water

intakes closed.

April 1999 Placer, Surigao del Norte,

Philippines

Cyanide tailings 700,000 tons released burying

17 homes

September

1995

Placer, Philippines Copper & Gold tailings 50,000 m3 released 12 killed

February 1994 Merriespruit, South Africa Gold mine tailings 6000,000 m3 released 17 killed

July 1985 Stava, Italy Fluoride tailings 190 000 m3 released 269 killed

January 1978 Arcturus, Zimbabwe Gold Mine tailings 20,000 m3 released 1 killed

November 1974 Bafokeng, South Africa Platinum mine tailings 3 million m3 flowed 45 km 12

killed

February 1972 Buffalo Creek, USA Coal tailings 500 000 m3 released 125 killed,

500 homes destroyed

September 1970 Mufilira, Zambia Cyanide tailings 68,000 m3 released 89 killed

129

6.8.1 Why Are Tailing Dams Still Being Built?

Economically they are still the cheapest option to build, maintain and rehabilitate.

While governments demand a bond for tailing dam rehabilitation after closure, it means

that disposal alternatives are less viable as there is less capital available.

Most decisions about tailings dams use a probability factorised cost for various potential

failure events. However, society should be wary of this, by learning from the loss of the

Challenger Space Shuttle. This space vehicle was designed for a failure rate lower than

1 in a 100,000 event according to all the experts before the disastrous event involving

the loss of the shuttle and its crew, but after investigation it turned out to be 1 in a 100.

This was an error in the failure rate estimate of 1000 times, and it was not due to poor

science which was very detailed, but to the variability of human behaviour, from the

designer to the operators Feynman (2009). Many tailings dams around the world today

claim to have catastrophic failure rates lower than one in a million, yet the actual

statistics indicate this is overly optimistic by a few orders of magnitude. What if we

reconsider the indefinite time cost of a rehabilitated tailings dam, whether the

rehabilitation was entirely successful or not (as defined by negligible leachates escaping

or maximum consolidation)? Statistically at some point in the next 10,000 years, an

earthquake, volcano, 1 in 10,000 year flood, tsunami or any major event will occur at

every tailings dam site. The risk of a catastrophic failure of a tailings dam, which is

currently estimated as a one in a million event has a one in a hundred chance of

occurring in this timeframe. And if the 1 in a million evaluated risk was in error by a

1000 times, like the Challenger Space Shuttle disaster Feynman (2009), then this

catastrophic failure will occur 10 times, and society will have to clean it up 10 times.

So in conclusion, nearly every rehabilitated dam is likely to have some impact at some

stage on the people and environment due to leachates, liquefaction, groundwater

contamination or surface water contamination.

6.8.2 Alternative Disposal Systems

Tailings are a mixture of particles, water and chemicals left over from the processing

130

plant. If it is “chemically bound” it makes a solid. This “bound” solid can be quite

useful in construction and landfill as the noxious chemicals are locked in the solids

matrix. These tailings can then be used in the construction of useful manmade

landforms. For example, a steep valley could be made less steep to prevent erosion or an

old mine pit could be filled, making the land more suitable urban development.

As the most common binder is cement, suitable placement characteristics can be

achieved with the addition of only 2% cement by weight. At $285 a tonne the cement

represents an additional cost to the tailing disposal system of less than $6 a tonne. The

binding of particles in an inert matrix can occur through different chemical reactions.

For this assessment we will assume that this binding occurs through the use of standard

grade cement.

Most binders are sensitive to the presence of water, especially where the binding

reaction requires a specific concentration of water such as mixing concrete using

cement. If dewatering is not required then the only additional cost will be the $6/t as

mentioned above. However, if dewatering is required the following additional cost will

occur:

thermal drying (which is expensive at $30 a tonne);

mechanical drying using belt press vacuum filters (which is less than $5 a

tonne), or;

adding dry material such as fly-ash or ground blast furnace reject material. The

addition of this material at 25% concentration may attract a cost of $5 a tonne.

The next cost after binding is materials handling. In normal tailings dam systems a

slurry pipeline provides low cost transport with centrifugal pumps and the flexibility of

a short pipeline to get to the emplacement sites. For a typical paste system with a binder

and delivery designed to create useful landforms, paste pumping or trucking is required.

Pumping the tailings as a paste would add an extra cost of between of $2 to $5 a tonne.

So in summary the costs are shown in Table 6.14.

131

Table 6.14 Summary of Costs Comparison.

Costs Tailings dam

$ a tonne

Bound stabilized fill

$ a tonne

Direct $1 to $5 $6+$5+$2 to $5=$13 to $16

Indirect $1 to $5 No indirect costs

TOTAL $2 to $10 $13 to $16

Clearly a bound stabilized fill is more than twice the cost of a tailings dam. However, it

could increase the value of the land. If we look at a hypothetical location in rural

Australia, for a dam or filled area which is say 10 m deep, this can be converted to a real

estate cost. Table 15 indicates the real estate costs.

Table 6.15 Real Estate Costs.

Tailings dam m2 Bound stabilized fill m

2

Land cost $20 to $100 $130 to $160

Typical unimproved land

value (rural)

Less than $0.5 More than $0.5

The results demonstrate there is little commercial viability for good bound stabilized

sites. Chemically binding the tailings into a solid does however mitigate the perpetual

risks outlined earlier.

Can society afford to chemically bind mine tailings? The cost of tailings disposal is a

small cost component of everything that is mined, equating to only a few percent of the

total cost. To keep a level playing field for our mining companies this would need to an

act of legislation from all governments around the world.

132

6.8.3 Example of Industries Changing from Slurry to Paste Production

The disposal of power station ash in Australia has been undergoing a significant shift in

emphasis during the past ten years. In older power stations, fly ash and bottom ash were

transported to a tailing dam in two purpose built systems:

The first system was for fly ash (dust). The dust removed from the boiler gas

passes by either fabric filters or precipitators collection systems. It was

hydraulically evacuated from the fabric filters or precipitators storage hoppers

on either an intermittent or continuous base and sluiced to the dust plant. In the

dust plant the sluiced dust was mixed with large quantities of water and pumped

using centrifugal pumps as lean phase slurry with a Cw <10 %;

The second system was for bottom ash, which was intermittently dumped from

the wet bottom ash hopper into a sluiceway and sluiced to the ash plant. In the

ash plant the sluiced bottom ash was first crushed to < 25 mm, then mixed with

large volumes of water and pumped using centrifugal pumps as lean phase slurry

Cw <10 %.

The slurry pipelines discharge into a tailings dam simply called the ash dam. The water

from the ash dam was recycled to the power station for reuse. The water used for ash

disposal systems could either be fresh or salt water depending on the power station

location.

For newer power stations and as a retrofit to existing stations, an alternative ash disposal

system was one where both the bottom ash and fly ash were mixed together and pumped

as high concentration slurry to a disposal site. The fly ash was removed from the

precipitators or fabric filters by a pneumatic conveying system and conveyed to a

HCSD (High Concentration Slurry Disposal) storage silo. The bottom ash was removed

from the boilers by a dry removal system and after passing through a hammer mill,

where the size was reduced to < 8 mm, was also pneumatically conveyed to the HCSD

storage silo. The ash from the HCSD storage silo was mixed as high concentration

slurry Cw of 63% in a mixing plant and pumped using diaphragm pumps at a flow rate

133

of 100 m3

h-1

to the disposal site in a 150 mm diameter pipeline with a pressure of 3

MPa.

In an another power station, fly ash slurry was pumped as a high concentration slurry at

a Cw of 72 % at flow-rates up to a maximum of 240 m3

h-1

in a 200 mm diameter

pipeline with a pressure of 6 MPa a distance of 10 km to a disposal site. The disposal

site is a disused open cut coal mine. Figure 6.37 is a photograph of the disposal site

Figure 6.37 Disused Mine Site.

134

While at another power station with a HCSD system there is no tailing dam, only bung

walls, and the disposal site is progressively rehabilitated. Figure 6.38 is a photograph of

the disposal site.

Figure 6.38 Bund Wall Ash Disposal Site.

6.8.4 Material Handling Solution for Disposal to Underground Mine Voids

Using mineral process tailings to produce paste backfill with a binder is well proven and

documented in specific engineering publications, such as the Australian Centre for

Geomechanics. This field of extensive and proven commercially viable research is

primarily aimed at increasing mining extraction ratios with structurally competent

backfill. An important way in which paste backfilling is beneficial is through reduction

of adverse environmental effects of tailings dams.

There are numerous underground mine voids being filled with tailings in Europe,

135

Australia, The Americas and South Africa. It is not always possible to put all tailings

underground due to insufficient underground voids. If a tailings dam was required its

size will be significantly reduced. Chemically bound and stabilised tailings are already

status quo in metaliferous mining where improved mining efficiencies have justified the

additional cost as a backfill.

In the coal industry in Europe, Deutsche Montan Technologie (DMT) developed a coal

mine backfilling system that was installed in the 1990’s at the Walsum Mine (Mez el

al., 1999). This mine was backfilled with residual material from processing and

combustion of coal, from incineration of domestic refuse and sewage sludge. This

system had a mixing and pumping station on the surface which delivered a 100 m3 h

-1 at

12 MPa of paste according to specific criteria to match both desired high solids content

and a low pressure loss. The paste was pumped through pipes to the coal face using a

powerful piston pump with a total power consumption of 480 kW. This system

successfully pumped the paste up to 12 km through a 200 mm pipeline to the working

face at a depth of 800 m. The paste was deposited in the long wall goaf through trailing

pipes which were 15 to 20 meters in length and were attached to the miner. The paste

deposited stayed in the goaf area and does not flow to other areas of the mine. Unlike

the conventional hydraulic stowing methods, it was not necessity to capture the residual

water and pump it back to the surface. A paste for backfill was prepared from refuse

material from the coal washery. The paste which was prepared from thickener

underflow material and ground rejects. Bunn (2009) conducted paste pumping trials at

the University of Newcastle indicated that the paste comprised finely ground reject

mixed with thickener underflow material that was be pumped at a Cw up to 80 %. This

paste was left in the pipeline for long periods and allowed the pumping system to be

restarted.

6.8.5 Conclusion

Mines should be considering alternatives to tailings dams and these should be

considered by any mine from all possible angles. Serious consideration should be given

to the acceptance of one in a million failure rate with numerous failures of tailing dams

136

throughout the world resulting in of loss of life, destruction of homes and infrastructure

and environmental pollution. Although current practices attempt to mitigate these risks,

claiming that catastrophic events are reduced to the level of one in a million or less,

there is still some argument that this may not be enough, and may not be achievable

when considering the longevity of the dam and human factors involved in design,

building and maintenance.

The principle of returning the refuse to the place of origins as a backfill is a logical

solution that should be pursued where possible. The principal of using dewatering,

binder and paste pumping for dry-stacking new dams or landforms should be pursued to

eliminate risks of tailings dams. Technologies to implement alternative methods exist

and are proven. The additional cost of this could be justified by closely examining the

true indirect costs.

The conceptual options which would replace traditional tailings dams include:

Tailings as a paste can be placed in an open cut void;

Tailings as a paste can be placed with a binder on a surface emplacement of a

desired shape of bound stabilized fill;

Tailings as a paste can be placed without binder as a mine backfill into old

voids;

Tailings as a paste can be placed with a binder as a mine backfill into old stopes

to improve mining extraction ratios.

137

6.9 Innovation in Bulk Materials Handling & Processing (2008) and Australian

Bulk Handling Review, Volume 14 No. 1 (2009) - The Pumpability of Coal

Washery Thickener Underflow

The process of disposing coal washery thickener underflow material is usually by lean

phase disposal systems where it was pumped to a disposal site either above ground, in a

tailing pond, or underground. After the material settled, the excess water was either

reused or allowed to return to the environment. The disposal of thickener underflow as a

high concentration slurry makes economic and environmental sense in the areas of

energy efficiency, water usage and land utilisation. This paper looks at the Rheology of

thickener underflows as a precursor to the disposal of the material as high concentrated

slurry.

The coal washery fines were taken to a thickener and were referred to as “thickener

underflow”. The coarse rejected material was referred to as “rejects”. High

concentration slurry was assumed to be slurry with a Cw above 45%. The design of the

disposal plant was to dispose of 100 m3

h-1

of thickener underflow slurry to a tailing

pond in a 10 km horizontal steel pipeline with a nominal bore of 150 mm and a

nominated pipeline pressure drop of 6.0 MPa.

6.9.1 Methodology

Two samples of thickener underflow materials were collected over several days from

the two washeries, one on the New South Wales South Coast and, the other from the

New South Wales Hunter Valley. The thickener underflow samples were allowed to

settle for several days and the excess water was decanted off and saved. Approximately

200 grams of the settled material was collected from each sample. The PSD of each

sample was determined using a laser diffraction technique (Malvern Particle Size

Analyser). The PSD results are shown in Figure 6.39. The samples from the South

Coast were designated A1 and A 2 and the samples from the Hunter Valley were

designated B 1 & B 2. The medium particle (d50) of the material was obtained from the

Malvern Particle Size Analyser results and is displayed in Table 6.16. The solids density

138

of the four dried samples were tested three times with a Micromeritics AccuPyc

Pycnmoter 1330 and then averaged. These results are displayed in Table 6.17.

Figure 6.39 Particle Size Distributions.

A Contraves Rheomat 30 Rotary Viscometer is a concentric cylinder rotational

viscometer working on the Searle system of a rotating inner cylinder with a stationary

outer cylinder. Rheology tests were conducted on the four thickener underflow samples.

Approximately 2 kg of the settled material was added to a 3-litre container then mixed

with the decanted water while stirring with a vertical stirrer until the consistency of

thick honey was obtained, see Figure 6.40. A sample of approximately 200 grams was

then placed in a Contraves Rotary Viscometer Cup for shearing. After shearing, the

contents of the rotary viscometer cup was placed in a petrie dish and weighed, dried and

reweighed to calculate the solids concentration (Cw). Decanted water was then added to

the mixing container to reduce the Cw for the next test. The shearing was repeated at

several different Cw’s.

Re-suspension tests indicate that a pipeline containing the different thickener underflow

slurries could be re-suspended in a pipeline for periods greater than 7 days after

0

10

20

30

40

50

60

70

80

90

100

1 10 100 1000

Per

cen

tage

Pas

sin

g (%

)

Particle Size (µm)

A1 A2 B1 B2

139

shutdown. This is provided that no water is lost from the slurry during the shut-down.

Figure 6.40 Mixing of Thickener Underflow Slurry.


As can be seen from Figure 6.41 there are significant differences in the d50 of the

thickener underflow materials between not only the samples collected from the different

coal washeries but between the samples collected from the same washery. Table 6.16

displays the different d50. The pipeline velocity calculated for a nominal flow of 100 m3

h-1

was 1.57 m s-1

.

140

Table 6.16 Particle Size Distribution and Solids Density.

South Coast d50

µm

Solids Density

kg m-3

Hunter Valley d50

µm

Solids Density

kg m-3

A1 47 1642.4 B1 17 1918.8

A2 70 1943.4 B2 33 2154.3

To enable a comparison of the pumping characteristics of the samples the shear stress

was calculated to be 22.5 Pa with a calculated shear rate of 84 s-1.

These results are

shown on Figure 6.41. The pipeline viscosity was calculated to be 0.268 Pa s. By

applying the shear stress of 22.5 Pa, shear rate of 84 s-1

to all the rheograms the

approximate Cw’s were obtained. These are shown in Table 5.17.

Figure 6.41 Rheogram for Typical Thickener Underflow Slurry.

Table 5.17 Cw and Slurry Density.

South Coast Cw

%

Slurry Density

kg m-3

Hunter Valley Cw

%

Slurry Density

kg m-3

A1 51.3 1251 B1 54.7 1355

A2 58 1391 B2 47.1 1337

0

5

10

15

20

25

30

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1) Cw = 51.5 % Cw = 50.9 % Cw = 50 % Cw = 49.4 %Cw = 48.7 % Cw = 47.8 % Cw = 42.6 %

22.5

84

141

With the pipeline pressure at a nominal 6.0 MPa and a flowrate of 100 m3

h-1

, the

pumping Cw of the South Coast samples varied between 51% and 58%, whereas, the

Hunter Valley samples varied between 47.1% and 54.7%. The variation in slurry

density was 1251 to 1391 kgm-3

for the South Coast washery and between 1337 to 1355

kgm-3

for the Hunter Valley washery. It can be concluded from the data that thickener

underflow slurries from both washeries can be pumped as high concentration slurries.

142

6.10 International Symposium of Reliable Flow of Particulate Solids IV

(RELPOWFLOW IV), (2008) – Water Available for Recycling after the Placement

of Dense Phase Fly Ash

This paper examines the maximum amount of water that is available for recycling from

a range of dense phase fly ash slurries. The amount of water that is available for

recycling after the placement of dense phase fly ash slurries is dependent on the Cw of

the placed slurry, the packing density of the fly ash particles and the PSD of the original

fly ash.

Lean phase fly ash slurry disposal systems have been operating for many years in power

stations throughout the world. Recent advances in dense phase fly ash slurry pumping

systems offer advantages for ash disposal systems by virtue of reducing land and water

utilisation and by reducing capital and operating costs. When pumping with lean phase

systems all the water is pumped back to the power station while in a dense phase slurry

pumping system some of the water that is mixed with the fly ash remains in the

deposited ash. These tests do not include water loss to evaporation, rainfall infiltration

or disposal site leakage.

6.10.1 Methodology

Bunn et al (2006) and (2007) used a rotary viscometer to determine the rheology of 14

different fly ashes. Figure 6.42 is a typical rheogram of one of these ashes. From the 14

different rheograms the pumping Cw was determined using a shear stress of 100 s-1

and

a shear rate of 10 Pa. The determined pumping Cw’s are shown in Table 6.18. Figure

6.43 is a graph showing the PSD for 14 different fly ashes. The test apparatus consisted

of fourteen 2.5 litre plastic tapered containers with an inside bottom diameter of 127.5

mm, an inner top diameter of 143 mm and a height of approximately 170 mm. The

containers neatly fitted inside each other about two thirds the way down.

The seven containers designated the inner containers, had a volume of 2.5 litres. These

containers had 25 x 4.5 mm holes drilled in the base as shown in Fig. 6.44. The inner

143

containers had a 125 mm diameter Number 54 hardened filter paper placed in them,

weighed and the weights recorder on the outside.

Figure 6.42 Rheogram of Typical Fly Ash.

The outer containers were weighed dry and the weights recorded on the outside of the

containers. The inner and outer containers were fitted together as shown schematically

in Fig. 6.44.

The slurry for the placement tests was mixed as follows:

between 1 and 1.5 kg of dry fly ash was placed in a 2-litre mixing container;

normal tap water was added to the fly ash to match the required slurry Cw;

the water and fly ash were mixed for about 10 minutes or until a homogeneous

mixture was obtained;

the mixed slurry was then carefully poured into the test apparatus so that no fly

ash leaked from under the filter paper:

the weight of the placed slurry was recorded and the exact weight of fly ash and

0

10

20

30

40

50

60

70

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

Cw 76.8 % Cw 75.2 % Cw 74.8 %

Cw 74.2 % Cw 72.7 % Cw 72.3 %

144

water calculated;

a lid was placed on top of the inner container and the test apparatus was left for

approximately 1 month so the excess water passed through the slurry and was

collected in the outer container;

the mixing was repeated for all the fly ash samples at the required Cw; and,

the weight of the outer container was checked every couple of days and weight

observed and when the weight of the bottom container reached equilibrium this

weight was recorded.

Table 6.18 Collected Data.

No. Cw

%

PSD

d50

um

Fly

Ash

grams

Water

grams

Water

Colle-

cted

grams

Water

Retained

in Slurry

grams

Cw of

Placed

Ash

%

Volume

above

Slurry

l

Volume

of

Slurry

l

Placed

Slurry

Density

tm-3

Surface

A 73 10 1455.3 538.3 113.2 300.4 82.9 1329.7 1170.3 1.24 Soft

B 71 36 1475.1 602.5 199.9 282.3 83.9 1315.4 1184.6 1.25 Medium

C 67 11 1474.9 726.4 238.1 376.1 79.7 1283.4 1216.6 1.21 Hard

D 65 11 1460.8 786.5 258.7 427.7 77.4 1210.8 1289.2 1.13 Hard

E 65 45 1453.1 782.4 216.6 470.7 75.5 1232.8 1267.2 1.15 Soft

F 67 10 1152.5 567.6 237.1 330.6 77.7 1629.4 870.6 1.32 Hard

G 57 15 984.7 742.9 228.5 390.0 71.6 1398.5 1101.5 0.89 Soft

H 64 21 1564.5 880.1 262.2 496.6 75.9 860.0 1640 0.95 Soft

I 58 16 1055.8 764.6 158.7 585.4 64.3 1395.4 1104.6 0.96 Soft

J 56 6.4 1167.7 917.5 221.0 571.6 67.1 1149.0 1351.1 0.86 Hard

K 53 17 1169.5 1037.1 194.6 748.5 61.0 1122.8 1377.2 0.85 Soft

L 62 7 1226.5 748.1 130.3 660.4 65.0 1356.2 1143.8 1.07 Soft

M 59 16 1200.7 834.4 151.7 561.7 68.1 1160.5 1339.5 0.90 Medium

N 62 21 1176.4 721.0 241.6 355.4 76.8 1244.0 1256.0 0.94 Hard

145

Figure 6.43 Particle Size Distributions of Fly Ashes.

Section A-A

Figure 6.44 Drawing of Placement Test Apparatus.

0

10

20

30

40

50

60

70

80

90

100

1 10 100 1000

Per

cen

tage

Pas

sin

g (%

)

Particle Size (µm) Fly Ashes

A B C D E F G H I J K L M N

Inner Container

Outer Container

Filter Paper

Slurry

Drain Holes Drilled in

Inner Container

A A Water Drained

from Slurry

146

To determine the packing density of the placed slurry a piece of cling wrap covered the

placed ash in the inner container and the test apparatus was weighed. Water was then

added on top of the cling wrap until the inner container was full. The weight of the

added water was recorded then carefully removed from the top of the cling wrap, the

cling wrap removed and the surface hardness of the placed ash tested.


The results of the placement tests are included in Table 6.18. The percentage of water

available for recycling in a power station dense phase system using the fly ash in

Sample A, at a Cw of 73%, requires the mixing of 730 tons of fly ash with 270 tons of

water to produce 1000 tons of slurry. When this 1000 tons of slurry is deposited at the

disposal site only 120 tons of water is available for recycling. Therefore, 150 tons of

water is captured in the deposited ash. Table 6.18 indicates the amount of water returned

compared to the water pumped. From Table 6.19, it can be seen the percentage of water

available for recycling varies depending on the pumped Cw and the PSD.

The characteristic of each placed slurry was determined qualitatively by the surface

deformation when a finger was pushed onto the surface. The slurries varied from hard,

having no surface deformation, to soft where the finger penetrated into the slurry

approximately 5 mm. Medium deformation was nominally 2.5 mm of penetration.

Figure 6.46 is a graph of the relationship between the percentage return water,

placement density and the Cw of the slurry. Fig. 6.47 is a graph of the relationship

between the Cw, placement density and PSD. The volume of return water varies

between 12.4 % and 59.8 % of the water required to mix the slurry. There was no

relationship between the surface deformation of the deposited slurry and the Cw at

which the slurry was pumped. From Figures 6.46 & 6.47 it can be seen that there is no

relationship between the Cw of the pumped slurry and percentage of return water, PSD

and packing density. The only relationship that was observed was that the deposited

slurry placement density showed an increase when the slurry could be pumped above a

Cw of 65%. When designing a dense phase pumping and return water system assurance

147

needs to be given that not only the pumpability of the slurry is tested but also the

placement characteristics of the slurry PSD is not an indicator of the pumpability or

placement characteristics.

Figure 6.46 Graph of Percentage Return Water and Placement Density Vs. Cw.

0.8

0.9

1

1.1

1.2

1.3

1.4

0

5

10

15

20

25

30

35

40

45

50

50 55 60 65 70 75

Pla

cem

ent

Den

sity

(t m

-3)

Per

cen

tage

Ret

urn

Wat

er (

%)

Cw (%)

Return Water Placement Density

148

Figure 6.46 Graph of Percentage Cw Vs Particle Size d50 and Placement Density.

Table 6.19 Results of Calculations from Deposition 1000 Ton of Slurry at Different Cw.

No.

Slurry

Cw

%

Weight

of Ash

t

Weight

of Water

t

Placed

Cw

%

Water Stored

in Ash

t

Water Available

for Return

t

Returned

Water

%

A 73 730 270 82.9 150 120 44.4

B 71 710 290 83.9 136 154 53.1

C 67 670 330 79.6 171 159 48.0

F 67 670 330 77.7 192 138 41.8

E 65 650 350 75.5 211 139 39.7

D 65 650 350 77.4 190 160 45.8

H 64 640 360 75.9 203 157 43.6

L 62 620 380 65.0 333 47 12.4

N 62 620 380 76.8 187 193 50.7

M 59 590 410 68.1 165 245 59.8

I 58 580 420 64.3 322 98 23.3

G 57 570 430 71.6 226 204 47.4

J 56 560 440 67.1 275 165 37.5

K 53 530 470 61.0 339 131 27.9

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0

5

10

15

20

25

30

35

40

45

50

73 71 67 67 65 65 64 62 62 59 58 57 56 53

Pac

kin

g D

ensi

ty (t

m-3

)

Par

ticl

e Si

ze d

50

(µm

)

Cw (%)

Particle Size Packing Density

149

6.11 9th

International Conference on Bulk Materials Storage, Handling and

Transportation (2007) - The Relationship between Packing Density and

Pumpability of Fly Ash Slurries

The pumping of dense phase fly ash slurries required that the void spaces between the

fly ash particles be filled with water and additional water added to allow transport of the

slurry in the pipeline. This paper looks at the relationship between the packing density

of slurry obtained by assisted compaction and the pumpability as determined by

rheology testing. The theory postulated is that it is possible to use this assisted packing

density to determine the pumpability of fly ash slurries.

The disposal of power station ash in dense phase systems has increased over the last two

decades. These dense phase systems have either been new plant or retrofitted to an

existing plant. The dense phase systems have involved the pumping of a mixture of

bottom ash and fly ash or, fly ash only slurries. The determination of the pumping

characteristics of these ash slurries requires a combination of either bench top or pilot

plant studies.

The pumping of dense phase fly ash slurries is a two-part process. The first part requires

the filling of the void spaces between the fly ash particles with water, while the second

part requires the addition of sufficient extra water to allow transport of the slurry in the

pipeline. Figure 6.47 is a graphical representation of the packing of fly ash slurry in a

pipeline.

Figure 6.47 Representation of the Packing of Fly Ash Slurry in Pipeline.

Pipeline Walls

Fly Ash Particles

with Void Spaces

filled with water

Extra Water added to

Transport Slurry

Pipeline Walls

150

6.11.1 Methodology

Fly ash samples were collected from twelve different hoppers from operating power

stations. PSD was analysed using a Malvern Laser Particle Analyser. The results are

shown in Figure 6.48. Two of the samples were analysed using a Scanning Electron

Microscope. The results are displayed in Figures 6.49 and 6.50.

Figure 6.48 Fly Ashes Particle Size Distributions.

Rheology tests were conducted on the fly ashes as described Bunn et al (1990). Slurry

pumpability is defined as the slurry flow-rate that can be achieved at a certain Cw with a

desired pipeline pressure drop i.e. at a similar shear rate and shear stress.

To determine the packing density of the different fly ashes 400 grams of fly ash was

mixed with water at the Cw indicated in Table 6.20. The Cw’s in Table 6.20 were

extracted from the rheograms for the different slurries at similar values of shear rate and

shear stress. The mixed slurries were added to a 500 ml measuring cylinder.

Compaction of slurries in the 500 ml measuring cylinders occurred in a Branson Series

7000 Ultrasonic Generator. The slurries were compacted for 24 hours See Figure 6.51

Bunn et al (2006).

0

10

20

30

40

50

60

70

80

90

100

1 10 100 1000

Per

cen

tage

Pas

sin

g (%

)

Particle Size (µm) Fly Ashes

A B C D E F G H I J K L

151

Figure 6.49 Scanning Electron Microscope Photograph of Fly Ash “A” Particles.

Figure 6.50 Scanning Electron Microscope Photograph of Fly Ash “J” Particles.

152

Table 6.20 Results from the Rheograms of the 12 Different Fly Ash.

Sample No. Shear Stress Pa Shear Rate s-1

Viscosity mPas-1

Cw %

A 10 100 100 71

B 10 100 100 73

C 10 100 100 71

D 10 100 100 67

E 10 100 100 65

F 10 100 100 65

G 10 100 100 57

H 10 100 100 59

I 10 100 100 56

J 10 100 100 59

K 10 100 100 62

L 10 100 100 64


Figure 6.52 is a photograph showing the consolidated fly ash slurries A, B, C, D, E and

F. The results of the rheology studies are shown in Table 6.21 and also indicate the

results of the rheology study and the compaction tests for all the fly ash slurries. From

the results indicated in Table 6.21 it can be seen that the differences between the Cw’s

from the rheology tests and the compacted fly ash slurries varied from 12.05 % for fly

ash H to 14.94% for fly ash J. The variation in the d50 of the fly ash ranged from 8 μm

for fly ash A to 45 µm for fly ash F. The assumption is that if you add 15 % extra water

to the results of the compaction tests you can assume that this will give you a reliable

indication of the pumpability of fly ash slurries.

153

Figure 6.51 Photograph of Fly Ash Slurries after Compacted in Ultrasonic Generator.

Table 6.21 Summary of the Results for 12 Different Fly Ash Slurries.

No. Cw

%

D50

µm

Initial

Ash

Weight

grams

Initial

Water

Weight

grams

Slurry

Level in

Measuring

Cylinder

mm

Ash Level

after

Compaction

mm

Water

Above

Compacted

Ash

mm

Weight

of

Water in

Ash

grams

Cw

of

Ash

%

Cw

Difference

%

A 71 8 400 163.4 345 260 85 78.38 83.62 12.62

B 73 10 400 147.9 325 240 85 62.95 86.40 13.40

C 71 36 400 163.4 344 258 86 77.38 83.79 12.79

D 67 11 400 197.0 395 300 95 102.01 79.68 12.68

E 65 11 400 215.4 420 320 100 115.38 77.61 12.61

F 65 45 400 215.4 420 315 105 110.38 78.37 13.37

G 57 15 400 301.8 500 365 135 166.75 70.58 13.58

H 59 29 400 278.0 475 360 115 162.97 71.05 12.05

I 58 16 400 289.7 480 340 140 174.29 69.65 13.65

J 59 17 400 278.0 470 333 137 140.97 73.94 14.94

K 62 20 400 245.2 450 330 120 125.16 76.17 14.17

L 64 31 400 225.0 420 310 110 115.00 77.67 13.67

154

Initial Slurry Level Compacted Ash Level

Figure 6.52 Photograph of Six Different Fly Ash Slurries after Assisted Settling.

155

6.12 5th

International Conference for Conveying and Handling Particulate Solids

(2006) - The Effect of Particle Size Distribution on the Rheology of Fly Ash

Slurries

Fly ash slurries from different power stations show a great variation in rheology which

can be related to the differences in particle size distribution. This paper examines the

variation in rheology of different power station fly ash slurries as measured by a coaxial

cylinder rotary viscometer and relates this difference in rheology to differences in fly

ash PSD as measured by a Malvern Particle Size Analyser.

Fly ash is a product of combustion from coal fired power station and is collected either

from electrostatic precipitators or fabric filters. The PSD of the fly ash is however

dependent on numerous factors including the coal type, milling plant, combustion

chamber design and boiler load.

The disposal of power station ash in dense phase systems, either been as new plant or

retrofitted to existing plant, has increased over the last decades. These dense phase

systems involved the pumping of slurries as a mixture of bottom ash and fly ash or fly

ash only. The determination of the pumping characteristics of these ash slurries requires

a combination of either bench top or pilot plant studies. The determination of the

pumping characteristics of material from existing power plants is usually

straightforward.

6.12.1 Methodology

Fly ashes from different power stations in Australia and South East Asia were collected

from power station boilers where either fabric filters or electrostatic precipitators were

installed. Fly ashes “A”, “B”, “E” and “F” were collected from power station fitted with

electrostatic precipitators whereas fly ashes “C”, “D” and “G” were collected from

power stations fitted with fabric filters. Fly ashes “C” and “G” were collected from the

same power station. Fly ash “C” was collected from the combustion of the normal coal

seams whereas fly ash “D” was from a test burn of coal from a different coal seam. Fly

156

ash “A” was collected from a rear hopper of a power station electrostatic precipitator.

Fly ash “B” was a blended sample collected from the same electrostatic precipitator as

sample A. The blended sample contained 10 % material from the rear precipitator

hopper. All the fly ashes PSD were analysed with a Malvern Laser Particle Size

Analyser and ashes C and G were examined and photographed by a Scanning Electron

Microscope. Bunn el al. (2006) conducted rheology tests on the seven fly ashes. The ash

samples were mixed with Hunter District Water and placed in Contraves Rotary

Viscometer for shearing at different Cw’s.


Figures 6.53 and 6.54 are Scanning Electron Microscope photographs of Fly Ashes “C”

and “G”. Both these photographs were taken at the same magnification. The fly ashes

are from the same power station burning coal but from different coal seams. The

Scanning Electron Microscope photograph reveals the major differences between fly

ashes. Fly ash “G” has a higher percentage of larger particles and a lack of smaller

particles. Fly ash “C” largest particles are approximately 35 µm whereas the largest

particles in fly ash “G” are approximately 110 µm. Fly ash “G” shows a lack of

variation in different particles sizes unlike fly ash “C”.

Figure 6.55 is a rheogram of the fly ash “C” slurries at Cw’s ranging from 74.4 % to

79.7 %. PSD was analysed using a Malvern Laser Particle Analyser. The results are

shown in Figure 6.46. The PSD curves, shown in Figure 6.56 for the seven fly ashes,

all have different shapes especially the curve for fly ash “C” which provides the most

efficient particle distribution with the highest Cw. The distribution of the different

particles sizes in fly ash “C” provides the minimum inter-particle void spaces between

the particles and therefore the highest Cw.

Table 6.22 shows a comparison of the d50 and Cw’s of the seven fly ashes at a shear rate

of 100 s-1

and a shear stress of 20 Pa. The variation in d50 of fly ashes was from 2.4 µm

for fly ash “A” to 78 µm for fly ash “G” and the Cw of the seven fly ashes varied from

66 % for fly ash “A” to 77 % for fly ash “C”. The Cw of the fly ashes “B” and “E” are

157

similar at the same shear rate and shear stress but the d50 varied from 13.2 µm for fly

ash “B” and 31 µm for fly ash “E”. The Cw of the seven fly ashes did not vary as the d50

increased as the lowest and highest d50 had similar Cw’s while two of the mid-range d50

has similar Cw’s. The fly ash with the greatest Cw was fly ash “C”.

Figure 6.53 Scanning Electron Microscope Photograph of Fly Ashes C.

Table 6.22 Cw and d50 of the Seven Fly Ashes.

Fly Ash A Fly Ash B Fly Ash C Fly Ash D Fly Ash E Fly Ash F Fly Ash G

d50 2.4 µm 13.2 µm 17.1 µm 22 µm 31 µm 70 µm 78 µm

Cw % 66 75 77 71 73 67 67

The conclusion reached was that fly ash slurries from different power stations show a

great variation in rheology which can be related to the differences in PSD. However, the

variation in rheology cannot be equated directly to the d50 of the fly ash particles but the

variation in the distribution of the particles across the PSD range. Fly ash “C”

approaches the ideal PSD Curve which allows for pumping at the maximum Cw.

158

Figure 6.54 Scanning Electron Microscope Photograph of Fly Ashes G.

Figure 6.55 Rheogram of Fly Ash Slurry “C”.

0

10

20

30

40

50

60

70

80

90

100

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

Cw 79.7 % Cw 78.6 % Cw 78.1 %

Cw 77.5 % Cw 76.8 % Cw 74.4 %

159

Figure 6.56 Particle Size Distributions of the Seven Fly Ashes.

0

10

20

30

40

50

60

70

80

90

100

1 10 100 1000

Per

cen

tage

Pas

sin

g (%

)

Particle Size (mm) Fly Ashes

A B C D E F G

160

6.13 5th

World Congress on Particle Technology (2006) - A Model to Determine

the Packing Density of Fly Ash Slurries

The pumping of dense phase fly ash slurries requires the void spaces between the fly

ash particles to be filled with water and approximately 5% to 8 % extra water is needed

to allow transport of the slurry in a pipeline. A model has been written to determine the

volume of the void spaces in fly ash and thus enable the pumpability of slurry to be

determined. This paper reports on a comparison of the data from the model and the

pumpability of fly ash slurries as determined by a rotary viscometer.

The disposal of power station ash in dense phase systems has increased over the last

decades and have either been installed as a new plants or retrofitted to an existing

plants. They involved the pumping of a mixture of bottom ash and fly ash or fly ash

only slurries. The determination of the pumping characteristics of these ash slurries

requires a combination of either bench top or pilot plant studies. The determination of

the pumping characteristics of material from existing power plants is usually

straightforward whereas the determination of the slurry characteristics of ash from new

plants is fraught with difficulty, especially when new grinding, combustion or collection

techniques are employed.

This research is aimed at reducing the cost of determining the pumping characteristics

of fly ash slurries from existing power plants using reduced testing. It requires the

collection of only a couple of kilograms of a representative sample of fly ash which was

analysed for both chemical properties and PSD. If the fly ash chemical properties are

inert, the model can be used.

6.13.1 Simulation Model

The model is based on the optimisation of an energy potential, with the energy potential

being given by:

𝑈 = 𝑈𝐺 + 𝑈𝐻 (6 - 9)

161

where:

𝑈𝐺 = gravitational component of the form 𝑈𝐺 = 𝑚𝑔ℎ; and,

𝑈𝐻 = the Hertzian contact strain energy of the form.

𝑈𝐻 = 𝐻(𝑥𝑖𝑥𝑛) is given by:

𝑈𝐻 = 4

15

𝐸

1 − 𝜐2 (𝑑

2−

𝑟𝑎

2) 2.5 (

𝑑

4)

0.5

𝑓𝑜𝑟 ‖𝑥𝑖 − 𝑥𝑛‖ < 𝑟𝑖 + 𝑟𝑛 (6 -10 )

𝑈𝐻 = 0 𝑓𝑜𝑟 ‖𝑥𝑖 − 𝑥𝑛‖ ≥ 𝑟𝑖 + 𝑟𝑛 (6 - 11)

where:

𝐸 = Young’s modulus;

𝜐 = Poisson’s ratio;

𝑑 = diameter of the sphere;

𝑟𝑎 = distance between the centres of the spheres; and

𝑈𝐻 = strain energy stored in the deformations at the points of contact between

particles.

The Hertz Contact Strain between particles (𝑖) and (𝑛) is given by, (𝐻(𝑥𝑖,𝑥𝑛)) and is

calculated from the work done in overcoming the Hertz Contact Force by Coste and

Gilles (1999). Obviously, (𝐻(𝑥𝑖,𝑥𝑛)) is only non-zero when particles (𝑖) and (𝑛) are

in contact. The particles are ‘dropped’ from a height above the container, by giving each

particle a much higher “y” value than the highest existing one. To simulate a physical

container boundaries are placed in the “x” and “z” directions with a steep penalty

function to prevent particles going outside the container. A radial distribution function

was used to measure the packing efficiency of the assembly once a simulation had been

completed. This radial distribution function was defined by the chance of finding a

particle any given distance from another particle.

162

6.13.2 Simulation Model Validation

This model Donohue and Wensrich (2006) has been used in the past to simulate single

size mixtures, binary mixtures and mixtures of varying distributions. Experiments have

been carried out in the laboratory to validate these results from the model against actual

experiments. Five different sizes of spherical lead shot were used to make single size

packings, binary packings and irregularly distributed packings. Three sizes of spherical

glass beads, each with their own normal distribution, were also studied to find the

packing efficiency. The model was able to produce dense random packing (~64%) for

the single size packings and produced comparable results to those measured in the lab

for the other packings. A full list of the results can be seen in Table 6.23.

Table 6.23 Packing Efficiency Results.

Packing Efficiency (Model) Packing Efficiency (Measured)

Single Size 64.0 % 63.9 %

Binary (50/50) 65.8 % 65.8 %

Lead Shot Mix 1 65.9 % 64.4 %

Lead Shot Mix 2 65.5 % 65.0 %

Lead Shot Mix 3 66.0 % 65.4 %

Lead Shot Mix 4 65.4 % 65.7 %

Lead Shot Mix 5 65.7 % 65.2 %

Glass Bead Set 1 64.0 % 64.9 %



6.13.3 Methodology

To model the packing of fly ash that has a given distribution, the cumulative distribution

found from the Malvern Particle Size Analyser, see Figure 6.57, was inputted into the

simulation program. To reproduce the same PSD within the model a random number

between zero and one was generated with the introduction of each new particle, with

this number corresponding to a point on the distribution curve. This point was then used

to interpolate a value for the particle radius. This method of simulating the size

distribution produced excellent results with the distribution of the model matching

163

almost exactly with the fly ash distribution. d50 = 14.34 µm, d90 = 53.31 µm & d10 =

4.19 µm.

Figure 6.57 Fly Ash PSD as Analysed by the Malvern Laser Particle Size Analyser.

Rheology tests (Bunn 1991) were conducted on the fly ash. Slurry pumpability is

defined as the slurry flow-rate that can be achieved at a certain Cw with a desired

pipeline pressure drop i.e. at a similar shear rate and shear stress. To determine the

packing of the fly ash a volume of 400 ml of fly ash weighing 375.716 grams was

mixed with water at Cw’s of 60%, 65% and 70%. The mixed slurries as well as a sample

of the dry fly ash were added to a 500 ml measuring cylinder which was then allowed to

settle for 96 hours. The cylinders were inspected after 24 and 48 hours yet the material

was fully settled after 24 hrs. The above tests were repeated using dry fly ash and

slurries with a Cw of 65% and 70%. The samples were then placed in measuring

cylinders which were placed in a water filled Branson Series 7000 Ultrasonic Generator

for 24 hours to increase compaction.

0

10

20

30

40

50

60

70

80

90

100

1 10 100 1000

Per

ecn

tage

Pas

sin

g (%

)

Particle Size (um)

Percentage Passing (%) Volume (%)

164


The results for the radial distribution function can be seen in Figure 6.58. It shows that

the model predicts the packing efficiency of the fly ash with the given distribution to be

approximately 67.5 %. Figures 6.59 and 6.60 are computer generated graphical

representations of the packing efficiency.

Figure 6.58 Packing Efficiency vs. Radial Distribution for Fly Ash.

Figure 6.59 Computer Generated Graphical Representation of the Model (Bottom

View).

Packing Efficiency vs Radial Distribution for Fly Ash

0

10

20

30

40

50

60

70

80

90

100

0 10 20 30 40 50

Radial Distribution (d50)

Pa

ck

ing

Eff

icie

nc

y (

%)

165

Figure 6.48 is a rheogram of the fly ash slurries with Cw’s from 65.3% to 72.4%. From

the rheogram it can be seen that with an increase in Cw the shear rate increases at a

constant shear stress.

The fly ash tested in this paper is from an operating power station in the process of

converting their existing lean phase fly ash disposal system to a dense phase system.

The dense phase fly ash slurry is to be pumped a distance of 3 km at a flow rate of 300

m3

h-1

. This corresponds to a shear rate of 55 s-1

and a shear stress of 25 Pa. Figure 6.61

indicates a Cw of approximately 70 %.

Figure 6.60 Computer Generated Graphical Representation of the Model (Side

View).

6.12.5 Packing Efficiency Calculation

𝑉 =𝑚𝑎

𝜌 (6 - 12)

𝜃 =𝑉

𝑉𝑚 (6 - 13)

where:

𝜃 = packing efficiency;

𝑉 = theoretical volume of fly ash;

𝑉𝑚 = measured volume;

𝑚𝑎 = mass of fly ash; and,

𝜌 = density of fly ash.

166

Figure 6.61 Fly Ash Rheology.

Figures 6.62 and 6.63 are representations of the settled and assisted of the slurries in the

measuring cylinders. Figures 6.64 and 6.65 are photographs of settled and assisted

settled ash slurries at a Cw of 70 %. A comparison of Figures 6.62, 6.63, 6.64 and 6.65

show the effect of assisted settling.

Key

420 ml

400 ml 400 ml Initial Level

340 ml 325 ml 300 ml

290 ml

290 ml After 24 hr.

Dry Fly Ash Cw 60% Cw 65% Cw 70%

Figure 6.62 Settling Tests of Dry Fly Ash and Fly Ash Slurries.

0

10

20

30

40

50

60

70

80

90

100

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

Cw 72.4% Cw 71.2% Cw 70.6% Cw 69.7% Cw 68.9%

Cw 68.2% Cw 67.7% Cw 66.7% Cw 65.8% Cw 65.3%

167

400 ml 360 ml

350 ml

260 ml 300 ml

242 ml Key

After 24 hours Vibration

Initial Level

Dry Fly Ash Cw 65% Cw 70%

Figure 6.63 Assisted Settling Tests of Dry Fly Ash and Fly Ash Slurries.

Table 5.23 Settling Tests Results.

Dry Fly Ash Slurry Cw 60% Slurry Cw 65% Slurry Cw 70%

Ash Weight m 375.716 grams 375.716 grams 375.716 grams 375.716 grams

Initial Water Weight Nil 250.477 grams 202.309 grams 161.021

Initial Water Level Nil 420 ml 360 ml 300 ml

Final Water Level Nil 340 ml 325 ml 290 ml

Initial Ash Volume 400 ml 420 ml 360 ml 300 ml

Final Ash Volume 400 ml 340 ml 325 ml 290 ml

Change in Water Level Nil 80 ml 35 ml 10 ml

Final Volume of Water in Ash

Nil 170.477 ml 167.309 ml 151.021

Ash Particle Density 2300 kg m-3 2300 kg m-3 2300 kg m-3 2300 kg m-3

Theoretical Volume of Fly Ash

1.634x10-4 m-3 1.634x10-4 m-3 1.634x10-4 m-3 1.634x10-4 m-3

Volume of Settled Ash 4.00x10-4 m-3 3.40x10-4 m-3 3.25x10-4 m-3 2.90x10-4 m-3

Cw of Settled Slurry Nil 68.79 % 69.19 % 71.33 %

Packing Efficiency 40.80 % 48.04 % 50.26 % 56.33

168

Table 5.24 Assisted Settling Tests Results.

Dry Fly Ash Cw 65% Cw 70%

Ash Weight m 375.716 grams 375.716 grams 375.716 grams

Initial Water Weight nil 202.309 grams 161.021

Initial Water Level nil 360 ml 300 ml

Final Water Level nil 260 ml 242 ml

Initial Ash Volume 400 ml 360 ml 300 ml

Final Ash Volume 350 ml 260 ml 242 ml

Change in Water Level nil 100 ml 58 ml

Final Volume of Water in Ash nil 102.477 ml 103.021

Ash Particle Density 2300 kg m-3

2300 kg m-3

2300 kg m-3

Theoretical Volume of Fly Ash 1.634x10-4

m-3

1.634x10-4

m-3

1.634x10-4

m-3

Volume of Settled Ash 4.00x10-4

m-3

2.60x10-4

m-3

2.42x10-4

m-3

Cw of Settled Slurry Nil 78.57 % 78.48 %

Packing Efficiency 46.67 % 62.83 % 67.50 %

Figure 5.62 Settled Ash Figure 5.63 Assisted Settled Ash

Figure 5.64 Settled Ash. Figure 5.65 Assisted Settled Ash.

169

6.13.6 Conclusion

The packing efficiency as predicted by the computer model and the packing efficiency

as determined from the assisted settling tests results for the fly ash tested show good

correlation. The Cw of the maximum packing density was 78.48 % corresponding to the

Cw 70 % so the hypothesis that the 5% to 8 % extra water needed to allow transport of

the slurry in the pipeline is proven correct. To improve the maximum placement density

of a dense phase slurry system some a vibrating system installed at the deposition site

would improve the placement density.

170

6.14 16th

International Conference on Hydrotransport (2004) - What a change in

coal supply can mean to a dense phase handling and pumping system for a large

coal fired power station

In a large coal fired power station, the greatest operating cost is the supply of fuel. In

this modern age of cost reduction the decision to source cheaper coal is purely an

economic one, but consideration has to be given to the impact of the cheaper coal on

plant operations.

The power station ash handling and pumping system removes dry fly ash from the

power station boilers fabric filter hoppers and pneumatically conveys it 800 meters to a

dense phase pumping plant were it is mixed with water and hydraulically conveyed 10

km to a disposal site. This paper examines the difference in mixing and pumping

properties of the present coal seam fly ash and the new coal seam fly ash.

Macquarie Generation is a power producer located in the Hunter Valley of New South

Wales operating Bayswater (4 x 660 MW) and Liddell (4 x 500 MW) Power Stations

where combined coal consumption is approximately 10 million tonnes per year. At

present coal is supplied from several coal mines in the vicinity of the stations though in

order to reduce operating costs Macquarie Generation is considering sourcing coal from

Ulan, approximately 100 km west of the stations. In order to evaluate the Ulan Coal a

preliminary trial burn was conducted at Bayswater. This trial burn suggested that the

Ulan coal fly ash could have significant different handling properties to the existing

regional coal. Another trial burn of Ulan coal was organised and fly ash samples were

collected for both Bayswater and Ulan coals.

6.14.1 Methodology

Fly Ash samples were obtained from Bayswater Power Station consisting of normal

Bayswater ash from the ash plant 2000 tonne storage bin (100 % Bayswater) and Ulan

coal fly ash from the trial burn (100 % Ulan).

Ash plant water was obtained from Bayswater Power Station. Bayswater and Ulan fly

171

ashes were blended in the following ratios:

i) 75 % Bayswater to 25 % Ulan;

ii) 50 % Bayswater to 50 % Ulan; and,

iii) 25 % Bayswater to 75 % Ulan.

Constituent analysis of normal Bayswater and Ulan coals were conducted (see Table

6.25) with the results showing a significant difference in the amount of silicon and

aluminium between the ashes. Bayswater coal had 57% silicon and 28% aluminium

while Ulan coal had 74% silicon and 17% aluminium. Other properties were similar.

Table 6.25 Analysis of Bayswater and Ulan Fly Ash.

Ash Constituents Percentage % 100 % Bayswater 100 % Ulan

Silicon As SiO2 56.9 73.4

Aluminium As AL2O3 27.9 16.6

Iron As Fe2O3 7.8 5.6

Calcium As CaO 2.5 0.85

Magnesium As MgO 1.2 0.42

Sodium As Na2O3 0.30 0.11

Potassium As K2O 1.2 0.88

Titanium As TiO2 1.4 0.76

Manganese As Mn3O4 0.11 0.10

Sulphur As SO3 0.20 0.07

Phosphorus AS P2O5 0.24 0.08

Strontium As SrO 0.07 0.02

Barium As BaO 0.07 <0.05

Zinc As ZnO <0.02 <0.02

Ultimate Analysis Carbon % 54.12 60.4

Calorific Value DAF MJ kg-1 32.56 34.4

Hardgrove Grindability Index 49.8 47

Ash % 26.7 19.3

Volatile Matter % 25.7 27.8

The PSD was determined for both the normal Bayswater fly ash and Ulan fly ash using

172

a laser diffraction technique (Malvern Particle Size Analyser). The results are shown in

Figure 6.66 which clearly indicates that the Ulan fly ash contains a lower proportion of

fines. The Bayswater ash has a d50 of 14 µm while the Ulan ash has a d50 of 60 µm.

Figure 6.66 PSD of Bayswater and Ulan Fly Ash

Rheology tests were conducted on normal Bayswater and Ulan fly ashes as well as the

blended samples. The fly ash samples were mixed with Bayswater ash plant water by

progressively adding water to about 1 kg of dry ash in a 3-litre container while mixing

with a vertical stirrer until the consistency of thick honey was obtained. A sample of

approximately 200 grams was then placed in a Contraves Rotary Viscometer for

shearing. Ash plant water was then added to the mixing container to reduce the solids

concentration by weight ‘Cw’ for the next test. The slurry test was repeated six times at

different Cw’s.

Figure 6.67 is a comparison of the Cw’s at similar shear rates and shear stresses taken

from the rheograms of the different ash mixtures. Figures 6.68 and 6.69 are rheograms

of the 100 % Bayswater and 100% Ulan fly ash slurries.

0

10

20

30

40

50

60

70

80

90

100

1 10 100 1000

Per

cen

tage

pas

sin

g (%

)

Particle Size (µm)

Bayswater Fly Ash Ulan Fly Ash

173

6.67 Comparison of Cw’s at Similar Shear Rates and Shear Stress.

Figure 6.68 Rheogram of 100 % Bayswater Fly Ash Slurry.

64 65 66 67 68 69 70 71 72 73 74 75 76 77

1

2

3

4

5

Cw (%)

Bayswater Ash 100 %

BW/Ulan 75/25 %

BW/Ulan 50/50 %

BW/Ulan 25/75 %

Ulan Ash 100 %

0

10

20

30

40

50

60

70

80

90

100

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

Cw 79.7 % Cw 78.6 % Cw 78.1 %

Cw 77.5 % Cw 76.8 % Cw 74.4 %

174

Figure 6.69 Rheogram of 100 % Ulan Fly Ash Slurry.

Comparative re-suspension tests were carried out with Bayswater and Ulan fly ash

slurries in the laboratory. The Bayswater Cw’s were 72% and 65% and Ulan Cw’s were

64 % and 58 %. All the slurries re-suspended after 48 hours indicating that if the

Bayswater Dense Phase Ash Plant Pipeline was shut down while pumping with any of

the previous slurries, it could be restarted within the 24 hours with no problems.

PH tests were conducted on 6:1 water and ash mixtures for both the normal Bayswater

and Ulan ashes. 100 grams of fly ash was mixed with 600 ml of tap water at a pH of 7.5

and pH measurements were taken at 30 minute intervals. The results are presented

graphically in Figure 6.70 and indicate that after 2.5 hours the normal Bayswater fly ash

stabilises to a pH of 11.2 while the Ulan fly ash stabilises to a pH of 10.3.

Figure 6.71 to 6.74 are photographs from a scanning electron microscope of the

Bayswater and Ulan fly ash samples are shown at two different magnifications. The

photographs indicate that the particles of both ashes are mainly spherical. However, the

particle sizes of the Ulan fly ash are significantly larger than the Bayswater fly ash.

0

5

10

15

20

25

30

35

40

45

50

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

Cw = 68.20 % Cw = 67.80% Cw = 65.30%

Cw = 64.5% Cw = 62.10% Cw = 60.50%

175

8

8.5

9

9.5

10

10.5

11

11.5

0 0.5 1 1.5 2 2.5 3 3.5 4

Normal

Ulan

pH

Time (hours)

Figure 6.70 pH Bayswater (Normal) and Ulan Fly Ash Slurries.

Figure 6.71 Bayswater Fly Ash 100 µm Scale.

176

Figure 6.72 Ulan Fly Ash 100 µm Scale.

Figure

6.73 Bayswater Fly Ash 10 µm Scale.

177

Figure 6.74 Ulan Fly Ash 10 µm Scale.

6.14.3 Conclusions

The test work conducted shows that the Ulan ash is coarser than the Bayswater normal

ash and this is reflected its pumpability. Figure 6.65 shows that at similar pumpability,

the Cw difference between normal Bayswater and Ulan fly ash slurries is in the order of

10 %. This means that if the nominal Cw for pumping with the Bayswater Dense Phase

Ash Plant with 100 % Bayswater Ash was Cw 72 % the with 100 % Ulan Ash the plant

would be pumping with a Cw in the order of 62 %.

This answers the question that a change in coal supply can adversely affect the

operation of a dense phase ash handling and pumping system.

Re-suspension tests indicated that both ashes could be restarted in the Bayswater

pipeline within 24 hours with no problems.

178

CHAPTER 7 HIGH CONCENTRATION SLURRY TESTING

7.1 Introduction

This chapter describes the slurry testing facility as well as the comparative testing of

different high concentration fly ash slurries. The fly ash was analysed to determine the

particle size distribution (PSD) and particle solids density. The fly ash particle shape was

analysed using a Scanning Electron Microscope.

Comparative rheological analyses were undertaken using a pipeline viscometer, rotary

viscometer and an ASTM flow cone.

7.1.1 Pipeline Viscometer

A pipeline viscometer test facility was constructed to determine the rheological behaviour

of fly ash slurries using two different size pipelines in series. Figure 7.1 is a schematic

diagram of the high concentration slurry pipeline test rig which includes a pipeline,

pump, and a mixing tank with stirrer, Figure 7.2 is a photograph of the test rig.

The mixing tank was attached to a Hidrostal Screw Centrifugal Impeller Pump driven

by 415 volt 7.5 kW electric motor as shown in Figure 7.3. The pipeline in Figure 7.4

was constructed of 80 mm nominal bore mild steel pipe (red). The 80 mm pipe had a

wall thickness of 5.49 mm and an inside diameter of 77.92 mm. On the return leg a 6.5

meter section of 80 mm pipe was replaced with a length of 50 mm nominal steel pipe

(yellow). The 50 mm pipe had a bore wall thickness of 3.91 mm and an inside diameter

of 52.5 mm.

Figure 7.5 shows the 80 mm inside diameter glass pipe to allow for visual inspection of

the slurry flow. Rubber hoses were installed on the suction and discharge sides of the

Hidrostal Pump.

179

Figure 7.1 Schematic Diagram of Slurry Test Rig (1 – Slurry Mixing Hopper, 2 –

Mixing Hopper Isolating Valve, 3 – Hidrostal Screw Centrifugal Impeller Pump, 4 –

Pipeline Isolating Valve, 5 – Pressure Transmitter, 6 – Differential Pressure Transmitter

80 mm Pipe, 7 – Reducers 80 to 50 mm, 8 - Differential Pressure Transmitter 50 mm

Pipe, 9 – Pipeline RTD, 10 80 mm Glass Viewing Section, 11 – Magnetic Flow Meter,

12 – Weigh Hopper, 13 – Weight Hopper Control Valve).

The slurry was pumped with a Hidrostal Screw Centrifugal Impeller Pump from an

agitated open storage tank to the pipeline.

Attached to the Hidrostal Pump Motor was a Zener MSC-3 Variable Speed Drive which

allowed for changeable slurry flow, which is shown in Figure 7.6.

The pump discharge pressure was measured with an Impress Pressure Transmitter and

the differential pressure over 5 meters of both the 80 mm and 50 mm pipes were

measured with individual Yokogawa Diaphragm Sealed Differential Pressure

Transmitter. These can be seen in Figures 7.7, 7.8 and 7.9.

180

Figure 7.2 Photograph of Mixing Tank and Stirrer.

The slurry pipeline temperature was measured with an R 100 resistance temperature

detector (RTD) attached to the 50 mm pipe adjacent to the glass section as shown in

Figure 7.10.

Test Facility Mixing Tank

Weigh Hopper

Hidrostal Pump

External Mixing Tank

Mixing Tank Stirrer Motor

and Speed Control

External Mixing Tank

Transfer Valve

181

Figure 7.3 Photograph of the Hidrostal Screw Centrifugal Impeller Pump.

Figure 7.4 Photograph of the 50 mm (Yellow) and 80 mm (Red) Pipeline Sections.

Hidrostal Pump

Hidrostal Pump

Electric Motor

182

Figure 7.5 Photograph of the 80 mm Glass Section.

Figure 7.6 Photograph of the Zener MSC-3 Variable Speed Drive.

183

Figure 7.7 Photograph of the Pump Discharge Impress Pressure Transmitter.

Figure 7.8 Photograph of a Yokogawa Differential Pressure Transmitters.

184

Figure 7.9 Photograph of the Yokogawa Transmitter Diaphragm.

Figure 7.10 Photograph of Pipeline Temperature RTD.

185

The slurry flow-rate was measured with an 80 mm Foxboro Magnetic Flow Meter and

the slurry mass flow-rate was measured with an automatic weigh hopper which was

mounted on load cells. When the measured hopper weight exceeded 25 kg, an automatic

valve opened and returned the slurry into the mixing hopper, the valve closed and the

cycle was repeated. Both instruments are shown in Figure 7.11.

All the pressure, differential pressure, temperature, volumetric flow and differential

weight data were collected by a DataTaker DT 800 (Figure 7.12). The data from the

DataTaker was collected in real time by a laptop computer. The parameters measured

and collected were:

Date and time;

Pump discharge pressure;

80 mm pipe differential pressure over 5 meters;

50 mm pipe differential pressure over 5 meters;

Slurry temperature;

Slurry flow; and,

Weight of slurry.

The data from the DataTaker 800 was collected at the maximum scan rate of 14 points

per second.

186

Figure 7.11 Photograph Magnetic Flow Meter and Weigh Hopper.

Weigh Hopper

Weigh Hopper

Control Valve Foxboro Magnetic

Flow Meter

187

Figure 7.12 Photograph DataTaker.

7.1.2 Rotary Viscometer

Rheological data was concurrently collected using a Contraves Rheomat 30 Rotary

Viscometer as shown in Figure 7.13. The Rheomat 30 was a concentric cylinder

rotational viscometer working on the Searle System of rotating inner cylinder which was

driven by an electric motor which rotated in the material to be measured and a stationary

outer cylinder. The torque exerted on the rotor by the material was measured and indicated.

The shear rate of the material was a function of rotational speed, and the resulting torque

was a function of the shear stress. The measuring rotor was driven by an electric motor

through a gear train. The control unit produced a variable frequency which was pre-set by

the 30-position switch. Therefore, the speed steps and the rate of shear follow a

geometrical progression over the entire speed range. The rotation was transmitted to the

measuring system by a cardanic chuck to prevent horizontal forces from affecting the

measuring system. The instruments rotating system was held by a torsion bar.

The rotating system caused a deflection on the torsion bar that was measured inductively

and displayed on the control panel indicator. The sensor system used on the RM-30 was a

188

rotating inner cylinder with a stationary outer cylinder designed to meet DIN standard

53109. The measuring system calibration factor was 0.504. Chen (2013) checked the

calibration of the rotary viscometer using Shell Omala S2 G320 oil.

Figure 7.13 Photograph Contraves Rheomat 30 Rotary Viscometer.

7.1.3 ASTM Flow Cone

Comparative tests were carried out using an ASTM Flow Cone. The ASTM Cone was

178 mm across the top and 190 mm to the apex to which was fitted a 30.8 mm long tube

with a diameter of 12.7 mm. The top of the cone was a cylinder of internal diameter 178

mm and 75 mm high. The 1725 ml level was the top of the cone.

The ASTM Flow Cone calibration was checked with water, as per ASTM C 939 – 10

Standard Test Method for Flow of Grout for Preplaced-Aggregate Concrete (Flow Cone

Method), which was 8 seconds. This was within the tolerance outlined in ASTM C 939

– 10. A photograph of the ASTM Flow Cone is shown in Figure 7.14.

189

Figure 7.14 Photograph ASTM Flow Cone.

7.1.4 Calibration of Test Rig Instrumentation

]

7.1.4.1 Calibration of Weigh Hopper

The weigh hopper was mounted on a frame that was supported by four PT4000 – 50 kg

load cells connected in parallel via a digital indicator as shown in Figure 7.15.

Calibration was undertaken by applying a series of known weights to the weigh

hopper and recording the readings on the digital indicator. Initially the digital indicator

zero was set accounting for the weight of the empty weigh hopper and frame. Five 10 kg

calibrated weights were used to calibrate the Weigh Hopper at 10 kg intervals to 50 kg.

A calibration curve as depicted in Figure 7.16 was derived from the calibration data.

190

Figure 7.15 Photograph Weigh Hopper

7.1.4.2 Calibration of Pressure and Differential Pressure Transmitters

The pump discharge pressure transmitter was calibrated using the of Barnett “Dead

Weigh” Tester see Figure 7.16 and the calibration sheet for the pressure transmitter is

shown in Figure 7.17

The differential pressure transmitters were factory calibrated with a maximum error of

0.004 % of full range. The calibration sheets are shown in Figures 7.18 and 7.19.

Weigh Hopper

Weigh Hopper

Control Valve

Weigh Hopper

Digital Indicator

Weigh Hopper

Load Cell

191

Figure 7.16 Calibration Curve Weigh Hopper

Figure 7.17 Photograph of Barnett “Dead Weigh” Tester

y = 0.9959x + 0.0156

R² = 1

0

5

10

15

20

25

30

35

40

45

50

0 5 10 15 20 25 30 35 40 45 50

Wei

ght

Mas

s (k

g)

Weigh Hopper Indicator Reading (kg)

192

Figure 7.18 Pump Discharge Pressure Transmitter Calibrations

Figure 7.19 80 mm Pipeline Differential Pressure DP1 Calibration Curve

y = 6.4099x + 403.97

R² = 1

0

200

400

600

800

1000

1200

1400

1600

1800

2000

0 25 50 75 100 125 150 175 200 225 250

mV

Pressure (kPa)

y = 0.9962x - 0.0991

R² = 1

0

10

20

30

40

50

60

70

80

90

100

0 10 20 30 40 50 60 70 80 90 100

Test

Pre

ssu

re (k

Pa)

Differential Pressure Transmitter Reading (kPa)

193

Figure 7.20 53 mm Pipeline Differential Pressure DP2 Calibration Curve

7.1.4.3 Calibration of PT 100 Resistance Temperature Detector

The PT100 Resistance Temperature Detector (RTD) was not calibrated as it was

purchases to DIN Standard 43760. The calibration accuracy to Din Standard 43760 is

shown in Table 7.

Table 7.2 RTD Calibration Accuracy

PT100 DIN 43760

Accuracy ± °C

0 °C 0.03

10 °C 0.04

20 °C 0.04

30 °C 0.05

40 °C 0.06

50 °C 0.07

60 °C 0.08

70 °C 0.09

80 °C 0.1

90 °C 0.11

100 °C 0.12

y = 0.9962x - 0.0991

R² = 1

0

10

20

30

40

50

60

70

80

90

100

0 10 20 30 40 50 60 70 80 90 100

Test

Pre

ssu

re (k

Pa)

Differential Pressure Transmitter Reading (kPa)

194

7.2 Slurry Mixing

Prior to the testing fly ash slurries, the slurry plant was tested using main water. The

pressure drop, flow and velocity were measured and water curves generated of pressure

drop verses flow and velocity was generated for both the 80 mm and 50 mm pipes. The

generated pressure drop verses flow water curves were compared to established water

curves using the Hazen-Williams equations.

Bulk samples of fly ash were obtained from Bayswater Power Station located 100 km

North West of Newcastle New South Wales in the Hunter Valley and Eraring Power

Station located 31 km south west of Newcastle and designated as sample “B” and “E”

respectively. Several sub-samples of the fly ash were collected and a Malvern Particle

Size Analyser (a laser diffraction technique instrument) was used to determine the fly

ash PSD. The solids density was tested with a Micromeritics AccuPyc Pycnmoter 1330.

Also during pumping samples of fly ash slurries were collected for Cw verification and

after being oven dried were also tested for PSD and solids density.

The fly ash slurries were first mixed in an external slurry mixing tank. Figure 7.20 is a

photograph of the external mixer.

Mixing was achieved by first adding approximately 100 litres of water to the external

mixer. The stirrer was started then enough fly ash was added slowly to fill the mixer.

Figure 7.21 is a photograph of the external mixer filled with fly ash “E” slurry. No attempt

was made to control the Cw of the fly ash slurry in the external mixer.

Prior to the transfer of fly ash slurry from the external mixer to the test facility, the mixing

hopper and pipeline was flushed with water via a temporary plastic pipe. After flushing the

pipeline the water level in the mixing hopper was lowed and the mixing hopper outlet

valve closed and the pump stopped. The pump and pipeline was left full of water. The fly

ash slurry in the external mixer was then poured into the mixer hopper and the stirrer

started. When all the fly ash from the external mixer had been transferred, the hopper

outlet vale was opened and the Hidrostal pump started.

195

Figure 7.21 Photograph of the External Mixer.

Figure 7.22 Photograph of the External Mixer Fly Ash “E”.

196

The water in the pipeline was flushed to waste. When all the water had been flushed as

indicated by slurry flow from the temporary pipework, the Hidrostal pump was stopped

and the temporary flushing pipe was removed. The pumped speed was increased to 30

Hz and the slurry circulated in the pipeline for 10 minutes to ensure complete mixing.

7.3 Slurry Testing

The determination of slurry rheology was achieved by circulating the fly ash slurry at the

initial Cw. The pump speed was reduced to 15 Hz and data collection commenced. The

slurry was circulated for a couple of minutes and then measurement over a one minute

time interval was recorded. The speed was incremented in 5 Hz steps until the flow rate

was approximately 20 m3 h

-1 for the fly ash “B” slurries and (after Hidrostal pump suction

pipe modification) to 27 m3 h

-1 for the fly ash “E” slurries. The same data collection

procedure was repeated at each speed step. At 30 Hz, a 3 litre sample of slurry was

procured and tested using the rotary viscometer and the ASTM flow cone. After shearing

the sample of slurry from the rotary viscometer cup was collected for Cw verification, PSD

and density analysis.

After completion of a test run, the mixer speed was reduced to 30 Hz and approximately

20 kg of dry fly ash was added progressively to the mixing hopper over 5 minutes to

increase the slurry Cw. After all the dry fly ash had been added, the slurry circulated in the

pipeline for 10 minutes to ensure complete mixing. The previously undertaken data

collection and testing was repeated with the slurry tested a several different Cw’s.

During testing of fly ash “E” at each speed point, the 80 mm glass pipe was observed for

the migration of fly ash particles on the internal glass surface. The speed at which all the

particles were migrating over the total internal circumference of the glass tube was

recorded. Figure 7.23 is a photograph of the 80 mm glass section during pumping.

At completion of all the pumping runs, the pump was stopped and the temporary plastic

pipe installed that led to an empty 200-litre drum. The pump was started and the slurry was

transferred to the 200-litre drum. When the mixer hopper level was at a low level, the

pump was stopped and the mixing hopper was filled to the top with water, the pump

197

started and the remaining slurry transferred to the 200-litre drum. After filling one 200-litre

drum the pump was again stopped and the drum was replaced with an empty drum. The

pump restarted and when the level in the mixer hopper was low, the pump was stopped and

the remaining water drained to waste by removing the pump discharge rubber pipe and

opening a drain valve at the bottom of the pump casing.

Figure 7.23 Photograph of the 80 mm Glass Pipeline Full of Slurry.

198

CHAPTER 8 RESULTS AND DISCUSSIONS

8.1 Introduction

In the previous chapter, hydraulic conveying trials were described, resulting in the

collection of a large quantity of data. Processing of this data led to a number of valuable

correlations which are key importance in the development and assessment of a

successful pressure drop prediction model. The data points in this chapter were

averaged over one minute or approximately 800 data points.

Comparative bench scale tests were conducted in parallel with the pipeline viscometer

using the rotary viscometer and ASTM flow cone. The values obtained from the rotary

viscometer tests were used to assess their feasibility for application in the pressure drop

prediction procedure.

This chapter will summarise the findings and present a prediction model that aims to

accurately reproduce the pressure drop values experimentally obtained from the test

facility.

8.2 Pipeline Viscometers Water Tests

Prior to testing of fly ash slurries in the pipeline viscometer test facility, water tests were

conducted to confirm reliable operation of the facility. The water curves generated were

then compared to established water curves using the Hazen Williams equation, Streeter

and Wylie (1986).

Figure 8.1 is the graph of the pressure gradient per meter of both the 50 and 80 mm

pipeline viscometer compared with pipeline velocity. Figure 8.2 is a graph of the pressure

gradient per meter of both pipeline viscometers compared to the system volumetric flow.

Added to this graph are the same parameters as calculated from the Hazen Williams

equation. The results of the water test indicate that the system was operating as expected.

199

Figure 8.1 Water Curve to the 50 and 80 mm Pipeline Viscometer.

Figure 8.2 Water Curve to the 50 and 80 mm Pipeline Viscometer Including the Calculated

Hazen Williams Curve.

0

0.5

1

1.5

2

2.5

3

3.5

4

0 0.5 1 1.5 2 2.5 3 3.5 4 4.5

Pre

ssu

re G

rad

ien

t (k

Pa

m-1

)

Velocity (ms-1)

80 mm Pipe 50 mm Pipe

0

1

2

3

4

5

6

0 5 10 15 20 25 30 35 40 45

Pre

ssu

re G

rad

ien

t (k

Pa

m-1

)

Flow (m3h-1) 80 mm Pipe 50 mm Pipe

80 mm Pipe Hazen-Williams Calculated 50 mm Pipe Hazen-Williams Calculated

200

8.3 Fly Ash “B” Characteristics

Ten samples of fly ash “B” were tested for PSD and particle density using a Malvern

Particle Size Analyser and a Micromeritics AccuPyc Pycnmoter 1330. The PSD results

are displayed in Figure 8.3 and the particle densities in Table 8.1.

Figure 8.3 Fly Ash “B” PSD.

Table 8.1 Fly Ash “B” Particle Density.

1

(kg m-3

)

2

(kg m-3

)

3

(kg m-3

)

Average

(kg m-3

)

1 2006.1 2005.6 2005.4 2005.7

2 2005.8 2005.5 2005.0 2005.4

3 1896.0 1896.1 1895.9 1896.0

4 2077.1 2076.9 2076.3 2076.7

5 2006.5 2006.5 2006.4 2006.4

6 2078.1 2076.9 2077.3 2077.4

7 2070.3 2070.7 2069.2 2070.0

8 2077.4 2076.5 2076.2 2076.6

9 2069.0 2069.4 2068.3 2068.9

10 2077.3 2076.7 2077.0 2077.0

Average 2036.0

0

10

20

30

40

50

60

70

80

90

100

0.1 1 10 100 1000

Per

cen

tage

Pas

sin

g (%

)

Particle Size (µm)

PSD 1 PSD 2 PSD 3 PSD 4 PSD 1A

PSD 2A PSD 3A PSD 4A PSD 5A PSD 6A

201

Sample 1 to 4 were dry sub samples removed from the bulk supply and samples 5 to 10

were the dried fly ash from the Cw verification samples. Figure 8.4 is a graph indicating

the d10, d50 and d90 of the 10 fly ash “B” samples, Figures 8.5 and 8.6 are Scanning

Electron Microscope photographs of fly ash “B” at different magnifications.

Figure 8.4 Fly Ash “B” PSD.

8.4 Comparison of Slurry Flows Measurements

To determine if the 80 mm Foxboro magnetic flow meter operated successfully on fly ash

slurries at high Cw’s? A comparison was conducted between the measured fly ash “B”

slurry volumetric flow rate as measure Foxboro magnetic flow meter and the volumetric

flow rate calculated from the mass flow measured by the weight hopper. The Cw used in

the calculation was 65.1 % with a slurry density of 1495 kg m-3. Figure 8.7 is a graph of

this comparison.

4.7 4.8 4.8 4.8 4.2 4.4 4.0 4.3 4.1 4.1 4.3

25.2 22.7 23.4 25.2

22.9 25.5

21.2 24.2 22.3 24.3 25.5

89.5

83.0 77.8

88.0

75.9

86.5

67.1

80.9

72.5

82.8 86.2

0

10

20

30

40

50

60

70

80

90

100

1 2 3 4 5 6 7 8 9 10 11

Par

ticl

e Si

ze (

µm

)

d10 d50 d90

202

Figure 8.5 Fly Ash “B” Scanning Electron Microscope Photograph.

Figure 8.6 Fly Ash “B” Scanning Electron Microscope Photograph.

203

Figure 8.7 Fly Ash “B” Comparison of Measured and Calculated Slurry Flow Rates.

The results of this comparison indicated that the Foxboro magnetic flow meter is an ideal

instrument for measuring slurry flow rate. This measurement was therefore used in all

calculations.

8.5 Testing Fly Ash “B” Slurry in Test Facility

To understand the subsequent tables, this nomenclature was required:

P1 – Slurry pump discharge pressure;

𝛥𝑃1 – Differential Pressure over 5 meter 80 mm pipeline viscometer;

𝛥𝑃2 – Differential Pressure over 5 meter 50 mm pipeline viscometer:

Table 8.2 was the average data over approximately 800 lines of data at different

frequency set-points. The pipeline velocity was calculated using the expression:

𝑉 = 4 𝑄

(𝜋 𝐷2) (8.1)

15

20

25

30

35

40

0 2 4 6 8 10 12 14 16 18 20

Pu

mp

Fre

qu

ency

( H

z)

Flowrate (m3h-1)

Flowrate Measured by 80 mm Magnetic Flowmeter Flowrate Calculated from Slurry Weigh Hopper

204

Table 8.2 Fly Ash “B” Averaged Data Cw 59.7%.

Frequency

(Hz)

P1

(kPa)

𝛥𝑃1

80 mm

(kPa)

𝛥𝑃2

50 mm

(kPa)

Temperature

(°C)

Flow

Q

(m3 h-1)

17.5 36.44457 0.50401 1.891301 24.73702 0.841137

20 41.15838 1.187134 3.722807 25.13378 7.568681

22.5 45.20752 1.605883 5.48014 25.21652 12.08708

25 50.29029 1.939746 7.447226 25.47468 14.81664

27.5 52.07372 1.998367 8.239223 25.64339 15.73381

30 55.28636 2.105168 9.699283 25.78348 16.94076

32.5 61.71037 2.393827 12.74105 26.13648 19.3874

35 63.4291 2.421023 13.65922 26.35294 19.92248

37.5 66.17433 2.5029423 14.80084 26.62649 20.80788

The slurry concentration by weight (Cw) was calculated by weighing a sample of the

wet slurry then drying it in an oven and reweighing the dry sample using the formula;

𝐶𝑤 = 𝑆𝑙𝑢𝑟𝑟𝑦 𝐷𝑟𝑦 𝑊𝑒𝑖𝑔ℎ𝑡

𝑆𝑙𝑢𝑟𝑟𝑦 𝑊𝑒𝑡 𝑊𝑒𝑖𝑔ℎ𝑡 (8.2)

The concentration by weight (Cw) was expressed as a percentage or as a fraction on a

weigh to weight basis (w/w).

The Newtonian shear rate (𝛾) was calculated from the expression:

() = ( 8 𝑉

𝐷) (8.3)

And the shear stress at the wall (𝜏𝑤) was calculated from the expression:

𝜏𝑤 = (𝐷 𝛥𝑃

4𝐿) (8.4)

Tables 8.3 and 8.4 are the calculated pipeline Newtonian viscometer data for the 50 mm

and 80 mm pipeline viscometers at a Cw of 59.7 %.

205

Table 8.3 Fly Ash “B” 50 mm Pipeline Viscometer Data Cw 59.7 %.

Frequency

(Hz)

Flow Q

(m3 h-1)

Flow Q

(m3 s-1)

Pipe Dia. (m)

Pipe Area

(m2)

Velocity V

(m s-1)

Pressure

(kPa m-1)

Pressure

(Pa m-1)

Shear Rate (s-1)

Shear Stress (Pa)

17.5 0.84113 0.00023 0.0525 0.00216 0.10793 0.17826 178.260 16.447 2.33967

20 7.56868 0.00210 0.0525 0.00216 0.97120 0.74456 744.561 147.993 9.77237

22.5 12.0870 0.00335 0.0525 0.00216 1.55099 1.09603 1096.03 236.342 14.3854

25 14.8166 0.00411 0.0525 0.00216 1.90124 1.48944 1489.45 289.714 19.5490

27.5 15.7338 0.00437 0.0525 0.00216 2.01893 1.64784 1647.85 307.648 21.6280

30 16.9407 0.00470 0.0525 0.00216 2.17381 1.93985 1939.86 331.247 25.4606

32.5 19.3873 0.00538 0.0525 0.00216 2.48776 2.54820 2548.21 379.087 33.4453

35 19.9224 0.00553 0.0525 0.00216 2.55642 2.73184 2731.84 389.550 35.8555

37.5 20.8078 0.00578 0.0525 0.00216 2.67003 2.96016 2960.17 406.862 38.8522

Table 8.4 Fly Ash “B” 80 mm Pipeline Viscometer Data Cw 59.7%.

Frequency

(Hz)

Flow Q

(m3 h-1)

Flow Q

(m3 s-1)

Pipe Dia. (m)

Pipe Area (m2)

Velocity V

(m s-1)

Pressure

(kPa m-1)

Pressure

(Pa m-1)

Shear Rate (s-1)

Shear Stress (Pa)

17.5 0.84113 0.00023 0.0779 0.00476 0.04899 0.10080 100.801 5.03057 1.93636

20 7.56868 0.00210 0.0779 0.00476 0.44089 0.23742 237.426 45.2658 4.62507

22.5 12.0870 0.00335 0.0779 0.00476 0.70409 0.32118 321.177 72.2889 6.25652

25 14.8166 0.00411 0.0779 0.00476 0.86309 0.38620 386.195 88.6136 7.52307

27.5 15.7338 0.00437 0.0779 0.00476 0.91652 0.39967 399.673 94.0989 7.78563

30 16.9407 0.00470 0.0779 0.00476 0.98683 0.42103 421.033 101.317 8.20173

32.5 19.3873 0.00538 0.0779 0.00476 1.12935 0.47876 478.765 115.949 9.32634

35 19.9224 0.00553 0.0779 0.00476 1.16052 0.48420 484.204 119.15 9.43230

37.5 20.8078 0.00578 0.0779 0.00476 1.21209 0.50059 500.589 124.445 9.75146

Figures 8.8 and 8.9 are graphs of slurry velocity in the 50 mm and 80 mm pipeline

viscometers against pipeline pressure gradient. Figures 8.10 and 8.11 are graphs of

slurry flow in 50 mm and 80 mm pipeline viscometers against pipeline pressure

gradient. A curve for water at the same conditions is also shown. One can see that the

frictional losses for water are much less than those for the slurries.

206

Figure 8.8 Fly Ash “B” Slurry Velocity verses Pressure Gradient at Different Cw in the

50 mm Pipeline Viscometer.

Figure 8.9 Fly Ash “B” Slurry Velocity verses Pressure Gradient at Different Cw in the


0

2

4

6

8

10

12

0 0.5 1 1.5 2 2.5 3 3.5 4

Pre

ssu

re G

rad

ien

t (k

Pa

m-1

)

Slurry Velocity (m s-1) 50 mm Pipe Cw = 59.7 % 50 mm Pipe Cw = 61.8 % 50 mm Pipe Cw = 65.1 %

50 mm Pipe Cw = 67.9 % 50 mm Pipe Water

0

0.5

1

1.5

2

2.5

3

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4 1.5

Pre

ssu

re G

rad

ien

t (k

Pa

m-1

)

Slurry Velocity (m s-1)

80 mm Pipe Cw = 59.7 % 80 mm Pipe Cw = 61.8 % 80 mm Pipe Cw = 65.1 %


207

Figure 8.10 Fly Ash “B” Slurry Flow verses Pressure Gradient at Different Cw in the 50

mm Pipeline Viscometer.

Figure 8.11 Fly Ash “B” Slurry Flow verses Pressure Gradient at Different Cw in the 80


0

1

2

3

4

5

6

7

8

9

10

0 5 10 15 20 25

Pre

ssu

re G

rad

ien

t (k

Pa

m-1

)

Slurry Flow (m3 h-1)



0

0.5

1

1.5

2

2.5

3

0 5 10 15 20 25

Pre

ssu

re G

rad

ien

t (kP

a m

-1)




208

Tables 8.5 and 8.6 are the results for the ASTM flow cone and rotary viscometer at a Cw

of 59.7 %. In the rotary viscometer MS was the measuring system constant and the

indicator reading was the output from the measuring system.

Table 8.7 contains the fly ash “B” slurries Cw, average particle density, water density

and the calculated slurry densities for both pipeline viscometers.

Figure 8.12 is a graph of the ASTM flow cone flow times at different Cw’s.

Table 8.5 Fly Ash “B” ASTM Flow Cone Times at Different Cw’s.

Cw

(%)

Cw

(w/w)

Flow Cone Time

(s)

59.67 0.597 10.41

61.80 0.618 10.72

65.09 0.651 12.60

67.91 0.679 18.91

Table 8.6 Fly Ash “B” Rotary Viscometer Result at Cw 59.7%.

Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 2 1.008

2.65 0.504 2.2 1.109

3.6 0.504 2.6 1.310

4.89 0.504 3 1.512

6.64 0.504 3.5 1.764

9.03 0.504 3.8 1.915

12.3 0.504 4 2.016

16.7 0.504 4.5 2.268

22.7 0.504 5 2.52

30.8 0.504 6.1 3.074

41.9 0.504 6.8 3.427

57 0.504 7.8 3.931

77.5 0.504 9.2 4.637

105 0.504 11 5.544

143 0.504 13 6.552

195 0.504 17 8.568

209

Table 8.5 Fly Ash “B” Slurries Cw’s and Densities.

Cw

(%)

Density Solids

ρs

(kg m-3

)

Density Water

ρw

(kg m-3

)

Density Slurry

ρsl

(kg m-3

)

59.7 2036 1000 1436.0

61.8 2036 1000 1458.7

65.1 2036 1000 1495.3

67.9 2036 1000 1528.0

Figure 8.12 Fly Ash “B” graph of ASTM Flow Cone Results.

All the data not shown in this chapter appears in Appendix A.

Figure 8.13 is a shear diagram of the rotary viscometer results compared to a Newtonian

pseudo-shear diagram of the 50 mm and 80 mm pipeline viscometers at different Cw’s.

From Figure 8.13 it can be seen that the curves a typical Bingham plastic curve with the

shear stress with increasing shear rate. Also the slope of the curve increase with

increasing Cw.

10

11

12

13

14

15

16

17

18

19

20

59 60 61 62 63 64 65 66 67 68 69

Flo

w C

on

e Ti

me

(s)

Cw (%)

Fly Ash B

210

Figure 8.13 Shear Diagram of the Rotary Viscometer Results compared to a Newtonian

Pseudo-Shear Diagram of the 50 mm and 80 mm Pipeline Viscometer at different Cw’s.

Table 8.8 compares the rotary viscometer results with the Newtonian results for 50 mm

and 80 mm pipeline viscometers at a shear rate of 100 s-1

at different Cw’s.

Table 8.8 Fly Ash “B” Comparison of the Rotary Viscometer Results with those from

the Newtonian 50 mm and 80 mm Pipeline Viscometers at a Shear Rate of 100 s-1

at

Different Cw’s.

50 mm Pipeline

Viscometer

80 mm Pipeline

Viscometer

Rotary Viscometer

Cw

(%)

(s-1)

𝜏

(Pa)

(s-1)

𝜏

(Pa)

(s-1)

𝜏

(Pa)

59.7 100 7.0 100 8.0 100 5.4

61.8 100 15.0 100 14.8 100 8.0

65.1 100 23.2 100 23.2 100 13.1

67.9 100 41.8 100 42.0 100 27.0

0

10

20

30

40

50

60

70

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

80 mm Pipe Cw = 57.9 % 50 mm Pipe Cw = 59.7 % RV Cw = 59.7 %




211

To align the rotary viscometer and pipeline viscometer results, the scale used on the

shear diagrams and the pseudo-shear diagrams was 0 to 200 seconds. These values

reflect the maximum operation range for power station dense phase fly ash pumping

system and grout pumping systems.

8.6 Determining Non-Newtonian Fly Ash “B” Slurry Characteristics

For the fly ash slurries with unknown rheology, the Weissenberg-Rabinowitsch equation

Chambers el al. (1986) was used to determine the wall shear rate (𝛤𝑤) for the non-

Newtonian fly ash slurries. The correction factor was calculated for the shear rate by

applying Equation (3-15), and the flow behaviour index (𝑛′) was determined using

linear regression in Excel. The values for the correction factor and the flow behaviour

index for the 50 mm pipeline and the 80 mm pipeline viscometer at different Cw’s are

shown in Tables 8.9 and 8.10.

Table 8.9 Fly Ash “B” 50 mm Pipeline Viscometer.

Cw

(%)

Cw

(w/w)

Flow Behaviour Index (𝑛′)

Correction

Factor

59.7 0.597 0.093 2.975

61.7 0.617 0.103 2.727

65.1 0.651 0.134 2.266

67.9 0.679 0.258 1.743

Table 8.10 Fly Ash “B” 80 mm Pipeline Viscometer.

Cw

(%)

Cw

(w/w)

Flow Behaviour

Index

(𝑛′)

Correction

Factor

59.7 0.597 0.0656 4.009

61.7 0.617 0.100 2.808

65.1 0.651 0.1336 2.274

67.9 0.679 0.260 1.741

212

The values for the Newtonian and non-Newtonian shear stress and shear rate along with

the flow behaviour index and correction factor for the 50 mm pipeline and the 80 mm

pipeline viscometer at a Cw of 59.7 % are shown in Tables 8.11 and 8.12.


Newtonian Non- Newtonian

Shear

Rate

(s-1)

Shear

Stress

(Pa)

Flow Behaviour

Index

(n')

Correction

Factor

Shear

Rate

(s-1)

Shear

Stress

(Pa)

16.447 2.340 0.0927 2.975 48.93 2.34

147.993 9.7720 0.0927 2.975 440.262 9.772

236.3423 14.3850 0.0927 2.975 703.092

289.714 19.549 0.0927 2.975 861.867 19.549

307.648 21.628 0.0927 2.975 915.218 21.628

331.247 25.460 0.0927 2.975 985.424 25.461

379.087 33.445 0.0927 2.975 1127.743 33.445

389.550 35.855 0.0927 2.975 1158.868 35.855

406.862 38.852 0.0927 2.975 1210.371 38.852



Shear

Rate

(s-1)

Shear

Stress

(Pa)

Flow Behaviour

Index

(n')

Correction

Factor

Shear

Rate

(s-1)

Shear

Stress

(Pa)

5.031 1.964 0.0656 4.009 20.168 1.9642

45.266 4.625 0.0656 4.009 181.470 4.625

72.289 6.257 0.0656 4.009 289.806 6.257

88.614 7.5231 0.0656 4.009 355.251 7.523

94.099 7.7856 0.0656 4.009 377.242 7.786

101.317 8.2017 0.0656 4.009 406.180 8.202

115.950 9.3263 0.0656 4.009 464.841 9.326

119.150 9.4323 0.0656 4.009 477.671 9.432

124.445 9.7515 0.0656 4.009 498.900 9.751

Table 8.13 compares the rotary viscometer results with the non-Newtonian results for

50 mm and 80 mm pipeline viscometers at a shear rate of 100 s-1


Figure 8.14 is a shear diagram of the rotary viscometer results comparing them to the

213

non-Newtonian pseudo-shear diagram of the 50 mm and 80 mm pipeline viscometers at

different Cw’s.

Figure 8.14 Shear Diagram of the Rotary Viscometer Results compared to a Non-

Newtonian Pseudo-Shear Diagram of the 50 mm and 80 mm Pipeline Viscometer at

different Cw’s.

Table 8.13 Fly Ash “B” Comparison of the Rotary Viscometer Results with those for the

Non-Newtonian 50 mm and 80 mm Pipeline Viscometers at a Shear Rate of 100 s-1

at

Different Cw’s.

50 mm Pipeline

Viscometer

80 mm Pipeline

Viscometer

Rotary Viscometer

Cw

(%)

(s-1)

𝜏

(Pa)

(s-1)

𝜏

(Pa)

(s-1)

𝜏

(Pa)

59.7 100 3.2 100 3.2 100 5.4

61.8 100 10.0 100 9.8 100 8.8

65.1 100 16.9 100 17 100 13.2

67.9 100 30.8 100 30.9 100 27

0

5

10

15

20

25

30

35

40

45

50

0 50 100 150 200

Shea

r St

ress

(P

a)

Shear Rate (s-1) 80 mm Pipe Cw =57.9 % 50 mm Pipe Cw = 59.7 % RV Cw = 59.7 %80 mm Pipe Cw = 61.8 % 50 mm Pipe Cw = 61.8 % RV Cw = 61.8 %80 mm Pipe Cw =65.1 % 50 mm Pipe Cw = 65.1 % RV Cw = 65.1 %80 mm Pipe Cw = 67.9 % 50 mm Pipe Cw = 67.9 % RV Cw = 67.9 %

214

8.7 Non-Newtonian Slurry Modelling Fly Ash “B”

Bingham plastic models were fitted to the non-Newtonian 50 mm and 80 mm pipeline

viscometer curves as indicated in Figure 8.15. The Bingham models contain both yield

stress and viscosity at the different Cw’s. The Bingham models are displayed in Table 8.14.

Figure 8.15 Fly Ash “B” Non-Newtonian Pseudo-Shear Diagram of the 50 mm and 80

mm Pipeline Viscometers at Different Cw’s with fitted Bingham Plastic Models.

Table 8.14 Fly Ash “B” Bingham Plastic Models fitted to the 50 mm and 80 mm

Pipeline Viscometer Curves.

Cw

(%)

Cw

(w/w)

Model Yield Stress

(𝜏𝑦)

(Pa)

Viscosity

(𝜂𝑏) (Pas)

59.7 0.597 𝜏𝑏 = 2.001 + 0.017 2. 001 0.017

61.8 0.618 𝜏𝑏 = 7.020 + 0.029 7. 020 0.029

65.1 0.651 𝜏𝑏 = 11.133 + 0.057 11.133 0.057

67.9 0.679 𝜏𝑏 = 15.700 + 0.149 15.700 0.149

0

10

20

30

40

50

60

70

80

0 50 100 150 200 250 300 350 400

Shea

r St

ress

(P

a)

Shear Rate (s-1)

80 mm Pipe Cw = 59.7 % 50 mm Pipe Cw = 59.7 % 80 mm Pipe Cw = 61.8 %50 mm Pipe Cw = 61.8 % 80 mm Pipe Cw = 65.1 % 50 mm Pipe Cw = 65.1 %80 mm Pipe Cw = 67.9 % 50 mm Pipe Cw = 67.9 % Model Cw = 59.7 %Model Cw = 61.8 % Model Cw = 65.1 % Model Cw = 67.9 %

215

To develop a model that predicts the pipeline pressure drop of high Cw fly ash slurries

involves determining the relationships between Cw and the yield stress and viscosity.

By fitting a model for yield stress ( 𝜏𝑏𝑦) from the data shown in Table 8.14, the

following relationship was obtained:

𝑌𝑖𝑒𝑙𝑑 𝑆𝑡𝑟𝑒𝑠𝑠 (𝜏𝑏𝑦) = 33519𝐶𝑤3 − 64665𝐶𝑤

2 + 41692𝐶𝑊 − 8973.2 (8.5)

And by fitting a model for viscosity (𝜂𝑏) from the data shown in Table 8.11 the


𝑉𝑖𝑠𝑐𝑜𝑠𝑖𝑡𝑦 (𝜂𝑏) = 429.14𝐶𝑤

3 − 3796.29𝐶𝑤2 + 492.93 𝐶𝑊 − 101.77 (8.6)

Figure 8.16 Fly Ash “B” Relationships between Cw and Bingham Yield Stress.

τby = 33519 × (Cw)3 - 64665 × (Cw)2 + 41692 × (Cw)- 8973.2

R² = 0.999

0

20

40

60

80

100

120

140

160

180

0.59 0.62 0.65 0.68 0.71 0.74 0.77 0.8

Yeild

Str

ess

(Pa)

Cw

Yeid Stress Vs Cw

216

Figure 8.17 Fly Ash “B” Relationships between Cw and Bingham Viscosity.

Therefore 𝑆ℎ𝑒𝑎𝑟 𝑆𝑡𝑟𝑒𝑠𝑠 (𝜏0) = 𝜏𝑏𝑦 + 𝜂𝑏 𝛤𝑤 (8.7)

Integrating equation 8.5 and 8.6 into equation 8.7 gives:

𝑆ℎ𝑒𝑎𝑟 𝑆𝑡𝑟𝑒𝑠𝑠 (𝜏0) = (33519𝐶𝑤3 − 64665𝐶𝑤

2 + 41692𝐶𝑊 − 8973.2) +

(429.14𝐶𝑤3 − 3796.29𝐶𝑤

2 + 492.93𝐶𝑊 − 101.77) 𝛤𝑤 (8.8)

By calculating the shear rate (𝛤𝑤) from the pipeline and flow parameters and

substituting in Equation 8.8, the pipeline shear stress (𝜏0) was determined. The pipeline

pressure (∆𝑃) was calculated by substituting Shear Stress (𝜏0) in Equation 8.9.

𝑃𝑖𝑝𝑒𝑙𝑖𝑛𝑒 (∆𝑃) = 𝜏0 4 𝐿

𝐷 (8.9)

ηb = 429.14 × (Cw)3 - 796.29 × (Cw)2 + 492.93 × (Cw) - 101.77

R² = 0.999

0

0.5

1

1.5

2

2.5

3

0.55 0.6 0.65 0.7 0.75 0.8

Vis

cosi

ty (P

a s)

Cw

Viscosity Vs Cw Model

217

Figure 8.18 is a graph of calculated pipeline pressure (ΔP) at different Cw’s for a 10 km

long slurry pipeline (L) with a nominal diameter of 200 mm (D) at a flowrate (Q) of 240

m3 h-1

. At a calculated pipeline shear rate (𝛤𝑤) of 84.88 s-1

.

Figure 8.18 Fly Ash “B” Calculated ΔP compared to Cw.

Therefore, a model to determine pipeline pressure drop was:

𝛥𝑃 = 11036𝐶𝑤3 − 20959𝐶𝑤

2 + 13299𝐶𝑊 − 2817.2 (8.10)

8.8 Site Collected Data

Data was collected from the high concentration pumping system by the author on the 19th

March 2013 from where fly ash “B” was procured. The data was for pumping on the

previous day where the pipeline flow was 240 m3 h

-1 with a pipeline inlet pressure of 6.8

MPa, a fly ash flow of 260 t h-1 and a water flow of 115 t h

-1. Using these figures, the

slurry Cw was calculated to be 69.3 %.

∆P = 11036 × Cw3 - 20959 × Cw

2 + 13299 × Cw - 2817.2

R² = 0.999

0

5

10

15

20

25

30

35

0.59 0.61 0.63 0.65 0.67 0.69 0.71 0.73 0.75

Pip

elin

e P

ress

ure

∆P

(Mp

a)

Cw

Fly Ash 'B' Slurry Pipeline Non-Newtonian Pressure Vs Cw Model

218

From Figure 8.18 the pipeline pressure according to the model at a Cw of 69.3 % was

7.9 MPa compared to the site data collected data pipeline pressure of 6.8 MPa. This

difference was probably due to a difference in the fly ash PSD.

8.9 Non- Newtonian Slurry Grout Modelling Fly Ash “B”

When operating a batch grouting plant in the field to maintain the consistency of the

grout, the plant operator is required to manually test the grout every batch using a flow

cone. To overcome this repetitive testing, it was proposed that this testing could be

replaced by using pipeline pressure drop of high Cw fly ash grouts when compared to

flow cone time, which requires determining the relationships between Cw, flow cone

time, yield stress and viscosity. These relationships are displayed in Figures 8.16, 8.17

and 8.19.

Figure 8.19 Fly Ash “B” Grout Flow Cone Time Compared to Cw.

By fitting a model for flow cone time (𝐹𝐶𝑇) from the data shown in Table 8.5 the


0.59

0.61

0.63

0.65

0.67

0.69

0.71

0.73

0.75

0.77

0 10 20 30 40 50 60 70 80

Cw

Flow Cone Time (s)

Cw Vs Flow Cone Time

Cw = -0.689 × (FCT)-0.583 + 0.805

219

𝐶𝑤 = −0.689 × (𝐹𝐶𝑇)−0.583 + 0.805 (8.11)

Inserting the relationship from Equation 8.11 into Equation 8.8, the following

relationship was obtained that allows the determination of shear stress (𝜏0) based on the

ASTM Flow Cone time.

𝑆ℎ𝑒𝑎𝑟 𝑆𝑡𝑟𝑒𝑠𝑠 (𝜏0) = 33519 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805)3

− 64665 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805)2

+41692 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805) − 8973.2

+ 429.14 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805)3

− 3796.29 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805)2

+492.93 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805) − 101.77 × (𝛤𝑤) (8.12)

By calculating the apparent shear rate (𝛤𝑤) from the pipeline and flow parameters and

substituting in equation 8.12, the pipeline shear stress (𝜏0) was calculated. The pipeline

pressure (∆𝑃) was calculated by substituting (𝜏0) in Equation 8.9. Normal practice

when pumping grout into worked out underground mines is to place the grout plant

close to the area that requires rehabilitation. The process that limits the pumping rate

from the grouting plant is the rate at which the dry fly ash is transported and unloaded at

the grouting plant. This limits the pumping flowrate (Q) to nominally 30 m3 h

-1through

an 80 mm (D) nominal bore pipe. Figure 8.20 is a graph of calculated pipeline pressure

(ΔP) at different Flow cone time (FCT) for a 100 m long slurry pipeline (L). From the

previous parameter the calculated pipeline apparent shear rate (𝛤𝑤) was 179 s-1

.

A model to determine fly ash “B” grout pipeline pressure drop was:

𝛥𝑃 = −0.0986 × (𝐹𝐶𝑇)2 + 27.923 × (𝐹𝐶𝑇) − 238.27 (8.13)

220

Figure 8.20 Fly Ash “B” Calculated ΔP compared to Flow Cone Time.

8.10 Fly Ash “B” Slurries Comparison of 50 mm and 80 Pipeline Viscometers

To validate the measuring technique of the 50 mm and 80 mm pipeline viscometers, the

Newtonian and non-Newtonian data was plotted on a pseudo-shear diagram. Figure

8.21 was a pseudo-shear diagram for the Newtonian 50 mm and 80 mm pipeline

viscometer and Figure 8.22 is the non-Newtonian results plotted on the a pseudo-shear

diagram.

∆P = - 0.0986 × (FCT)2 + 27.923 × (FCT) - 238.27

R² = 0.998

0

200

400

600

800

1000

1200

1400

0 10 20 30 40 50 60 70

Pip

elin

e P

ress

ure

(kP

a 10

0m-1

)

Flow Cone Time (s)

Flow Cone Time Vs DP

221

Figure 8.21 Fly Ash “B” Newtonian Pseudo-Shear Diagram of the 50 mm and 80 mm

Pipeline Viscometer at different Cw’s.

Figure 8.22 Fly Ash “B” Non-Newtonian Pseudo-Shear Diagram of the 50 mm and 80

mm Pipeline Viscometer at different Cw’s.

0

10

20

30

40

50

60

70

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)



80 mm Pipe Cw = 67.9 % 50 mm Pipe Cw = 67.9 %

0

5

10

15

20

25

30

35

40

45

50

0 50 100 150 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

80 mm Pipe Cw =57.9 % 50 mm Pipe Cw = 59.7 % 80 mm Pipe Cw = 61.8 %

50 mm Pipe Cw = 61.8 % 80 mm Pipe Cw =65.1 % 50 mm Pipe Cw = 65.1 %


222

8.11 Fly Ash “E” Characteristics

Nine samples of fly ash “E” were tested for PSD and particle density using a Malvern

Particle Size Analyser and a Micromeritics AccuPyc Pycnmoter 1330. The PSD results

are displayed in Figure 8.23 and the densities in Table 8.15.

Sample 1 was a dry sub sample removed from the bulk supply and samples 2 to 9 were

the dried fly ash from the Cw verification samples. Figure 8.24 is a graph indicating the

d10, d50 and d90 of the 10 fly ash “E” samples.

Figure 8.23 Fly Ash “E” PSD.

0

10

20

30

40

50

60

70

80

90

100

0.1 1 10 100 1000

Per

cen

tage

Psi

ing

(%)

Particle Size (µm)

PSD 1 PSD 2 PSD 3 PSD 4 PSD 5

PSD 6 PSD 7 PSD 8 PSD 9

223

Table 8.15 Fly Ash “E” Particle Density.

Figure 8.24 Fly Ash “E” PSD.

6.6 6.5 6.0 6.2 6.0 5.6 5.3 5.6 5.3

34.4 33.1 36.8 37.6 37.8

33.6 34.2 34.3 31.9

125.9 127.9

145.2 143.4 141.6

126.7 131.3

128.5

119.9

0

20

40

60

80

100

120

140

160

1 2 3 4 5 6 7 8 9

Par

ticl

e Si

ze (

µm

)

d10 d50 d90

1

(kg m-3

)

2

(kg m-3

)

3

(kg m-3

)

Average

(kg m-3

)

1 2.0909 2.0904 2.0908 2.0907

2 2.0879 2.0885 2.0749 2.0837

3 2.0885 2.0856 2.0871 2.0870

4 2.0854 2.0872 2.0847 2.0857

5 2.0875 2.0808 2.0872 2.0851

6 2.0826 2.083 2.0826 2.0827

7 2.0842 2.0855 2.0843 2.0846

8 2.0851 2.084 2.0853 2.0848

9 2.091 2.0905 2.0906 2.0907

Average 2.0862

224

8.12 Testing Fly Ash “E” Slurry in Test Facility

Fly ash “E” was pumped through the test facility and the data analysed the same as fly ash

“B”.

Table 8.16 was the average data over approximately 800 lines of data at different

frequency set-points for a Cw = 66.6 %.

The 50 mm and 80 mm pipeline viscometers velocity, Cw, Newtonian shear rate and

shear stress were calculated from Equation 8.1, 8.2, 8.3 and 8.4 respectively and are

displayed in Tables 8.17 and 8.18 for a Cw of 66.6 %.

Table 8.16 Fly Ash “E” Averaged Data Cw 66.6 %.

Frequency

(Hz)

P1

(kPa)

𝛥𝑃1

80 mm

(kPa)

𝛥𝑃2

50 mm

(kPa)

Temperature

(°C)

Flow

Q

(m3 h-1)

20 64.383 2.661 7.8061 30.980 1.692

25 71.326 4.175 11.156 30.951 3.294

30 86.165 6.371 16.895 31.058 7.071

35 101.507 8.084 24.495 31.048 10.909

40 117.482 9.277 32.390 31.111 15.151

45 136.354 10.184 41.602 31.304 19.803

50 156.624 11.280 50.822 31.468 24.160

55 165.108 11.561 54.620 31.624 25.980

Table 8.17 Fly Ash “E” 50 mm Pipeline Viscometer Data Cw 66.6 %.

Frequency

(Hz)

Flow Q

(m3 h-1)

Flow Q

(m3 s-1)

Pipe Dia. (m)

Pipe Area (m2)

Velocity V

(m s-1)

Pressure

(kPa m-1)

Pressure

(Pa m-1)

20 1.692 0.000470 0.0525 0.002165 0.217 1.648 1647.893

25 3.294 0.000915 0.0525 0.002165 0.423 2.354 2354.053

30 7.071 0.001964 0.0525 0.002165 0.908 3.567 3566.824

35 10.909 0.00303 0.0525 0.002165 1.400 5.1711 5171.141

40 15.152 0.004209 0.0525 0.002165 1.944 6.838 6837.910

45 19.803 0.005501 0.0525 0.002165 2.541 8.783 8782.739

50 24.160 0.006711 0.0525 0.002165 3.100 10.729 10729.020

55 25.980 0.007217 0.0525 0.002165 3.334 11.531 11530.94

225

Figures 8.24 and 8.25 are graphs of slurry velocity in the 50 mm and 80 mm pipeline

viscometers against pipeline pressure gradient. Figure 8.26 and 8.27 are graph of slurry

flow in the 50 mm and 80 mm pipeline viscometers against pipeline pressure gradient.

A curve for water at the same conditions is also shown. One can see that the frictional

losses for water are much less than those for the slurries.

Table 8.18 Fly Ash “E” 80 mm Pipeline Viscometer Data Cw 66.6 %.

Frequency

(Hz)

Flow Q

(m3 h-1)

Flow Q

(m3 s-1)

Pipe Dia.

(m)

Pipe Area

(m2)

Velocity V

(m s-1)

Pressure

(kPa m-1)

Pressure

(Pa m-1)

20 1.692 0.000470 0.07792 0.004769 0.099 0.732 732.220

25 3.294 0.000915 0.07792 0.004769 0.192 0.875 874.901

30 7.071 0.001964 0.07792 0.004769 0.412 1.274 1274.155

35 10.909 0.00303 0.07792 0.004769 0.635 1.619 1616.833

40 15.152 0.004209 0.07792 0.004769 0.883 1.875 1875.464

45 19.803 0.005501 0.07792 0.004769 1.154 2.204 2203.671

50 24.160 0.006711 0.07792 0.004769 1.407 2.494 2493.597

55 25.980 0.007217 0.07792 0.004769 1.513 2.652 2652.223

226

Figure 8.24 Fly Ash “E” Slurry Velocity verses Pressure Gradient at Different Cw 50


Figure 8.25 Fly Ash “E” Slurry Velocity verses Pressure Gradient at Different Cw in the


0

2

4

6

8

10

12

0 0.5 1 1.5 2 2.5 3 3.5 4

Pre

ssu

re G

rad

ien

t (k

Pa)


50 mm Pipe Cw 58.1 % 50 mm Pipe Cw 59.2 % 50 mm Pipe Cw 62.3 % 50 mm Pipe Cw 63.8 %

50 mm Pipe Cw 65.4 % 50 mm Pipe Cw 65.9 % 50 mm Pipe Cw 66.6 % 50 mm Pipe Water

0

0.5

1

1.5

2

2.5

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2

Pre

ssu

re G

rad

ien

t (kP

a)



80 mm Pipe Cw 65.4 % 80 mm Pipe Cw 65.9 % 80 mm Pipe Cw 66.6 % 80 mm Pipe Water

227

Figure 8.26 Fly Ash “E” Slurry Flow verses Pressure Gradient at Different Cw in the 50


Figure 8.27 Fly Ash “E” Slurry Flow verses Pressure Gradient at Different Cw in the 80


0

2

4

6

8

10

12

0 5 10 15 20 25 30 35

Pip

elin

e P

ress

ure

(kP

a m

-1)



50 mm Pipe Cw 65.3 % 50 mm Pipe Cw 65.9 % 50 mm Pipe Cw 66.6 % 50 mm Water

0

0.5

1

1.5

2

2.5

0 5 10 15 20 25 30 35

Pip

elin

e P

ress

ure

(kP

a m

-1)

Slurry Flow (m3 h-1) 80 mm Pipe Cw 58.1 % 80 mm Pipe Cw 59.2 % 80 mm Pipe Cw 62.3 % 80 mm Pipe Cw 63.8 %

80 mm Pipe Cw 65.3 % 80 mm Pipe Cw 65.9 % 80 mm Pipe Cw 66.6 % 80 mm Water

228

Tables 8.19 and 8.20 are the results for the ASTM flow cone and rotary viscometer at a

Cw of 66.6 %

Table 8.19 Fly Ash “E” ASTM Flow Cone Times at Different Cw’s.

Cw

(%)

Cw

(w/w)

Flow Cone Time

(s)

58.06 0.581 10.37

59.25 0.592 10.59

62.34 0.624 10.87

63.83 0.683 11.75

65.34 0.653 13.54

65.93 0.659 14.75

66.63 0.666 16.44

Table 8.20 Fly Ash “E” Rotary Viscometer Result at Cw 66.6%.

Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 5.2 2.621

2.65 0.504 6.9 3.478

3.6 0.504 8.2 4.133

4.89 0.504 9.5 4.788

6.64 0.504 11.2 5.645

9.03 0.504 13 6.552

12.3 0.504 15.5 7.812

16.7 0.504 18 9.072

22.7 0.504 21 10.584

30.8 0.504 26 13.104

41.9 0.504 29 14.616

57 0.504 34 17.136

77.5 0.504 40 20.160

105 0.504 48 24.192

143 0.504 58 29.232

195 0.504 73 36.792

Figure 8.28 is a graph of the ASTM flow cone flow times at different Cw’s.

229

Figure 8.28 Fly Ash “E” Graph of ASTM Flow Cone Results.

Table 8.21 contains the fly ash “E” slurries Cw, average particle density, water density

and the calculated slurry densities for both pipeline viscometers.

Figure 8.29 is a shear diagram of the rotary viscometer results comparing then to a

Newtonian pseudo-shear diagram of the 50 mm and 80 mm pipeline viscometers at

different Cw’s. From Figure 8.29 it can be seen that the curves a typical Bingham plastic

curve with the shear stress with increasing shear rate. Also the slope of the curve

increase with increasing Cw.

Table 8.22 compares the rotary viscometer results, with those from the 50 mm and 80

mm pipeline viscometers at a shear rate of 100 s-1


The slurry concentration by weight (Cw) was calculated by weighing the wet slurry, then

drying it in an oven and reweighing the dry sample using the equation 8.4. To align the

rotary viscometer and pipeline viscometer results, the scale used on the shear diagrams

and the pseudo-shear diagrams was 0 to 200 seconds. These values reflect the maximum

10

11

12

13

14

15

16

17

57 58 59 60 61 62 63 64 65 66 67 68

Tim

e (s

)

Cw (%)

Fly Ash E

230

operating range for power station dense phase fly ash pumping system and grout

pumping systems.

Table 8.21 Fly Ash “E” Slurries Cw’s and Densities

Cw

(%)

Density Solids

ρs

(kg m-3

)

Density Water

ρw

(kg m-3

)

Density Slurry

ρsl

(kg m-3

)

58.1 2086 1000 1433.33

59.2 2086 1000 1446.0

62.3 2086 1000 1480.6

63.8 2086 1000 1497.8

65.4 2086 1000 1515.8

65.9 2086 1000 1522.6

66.6 2086 1000 1531.1

Figure 8.29 Fly Ash “E” Shear Diagram of the Rotary Viscometer Results compared to a

Newtonian Pseudo-Shear Diagram for the 50 mm and 80 mm Pipeline Viscometer at

different Cw’s.

0

10

20

30

40

50

60

70

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1) 50 mm Pipe Cw = 58.1 % 80 mm Pipe Cw = 58.1 % RV Cw = 58.1 %50 mm Pipe Cw = 59.2 % 80 mm Pipe Cw = 59.2 % RV Cw = 59.2 %50 mm Pipe Cw = 62.3 % 80 mm Pipe Cw = 62.3 % RV Cw = 62.3 %50 mm Pipe Cw = 63.8 % 80 mm Pipe Cw = 63.8 % RV Cw = 63.8 %50 mm Pipe Cw = 65.4 % 80 mm Pipe Cw= 65.4 % RV Cw = 65.4 %50 mm Pipe Cw = 65.9 % 80 mm Pipe Cw = 65.9 % Rv Cw = 65.9 %50 mm Pipe Cw = 66.6 % 80 mm Pipe Cw = 66.6 % RV Cw = 66.6 %

231

Table 8.22 Fly Ash “E” Comparison of the Rotary Viscometer Results with those from

the Newtonian 50 mm and 80 mm Pipeline Viscometers at a Shear Rate of 100 s-1

at

Different Cw’s.

50 mm Pipeline

Viscometer

80 mm Pipeline

Viscometer

Rotary Viscometer

Cw

(%)

(s-1)

𝜏

(Pa)

(s-1)

𝜏

(Pa)

(s-1)

𝜏

(Pa)

58.1 100 7.2 100 7.6 100 4.6

59.2 100 6.8 100 7.6 100 4.9

62.3 100 14.0 100 14.4 100 8.7

63.8 100 19.6 100 19.6 100 12.0

65.4 100 26.3 100 26.3 100 15.5

65.9 100 34.5 100 34.5 100 19.5

66.6 100 38.4 100 38.5 100 23.5

8.13 Determining Non-Newtonian Fly Ash “E” Slurry Characteristics

For the fly ash slurries with unknown rheology, the Weissenberg-Rabinowitsch equation

Chambers el al. (1986) was used to determine the wall shear rate (𝛤𝑤)for the non-

Newtonian fly ash slurries. The correction factor was calculated for the shear rate by

applying Equation (3-15), and the flow behaviour index (𝑛′) was determined using

linear regression in Excel. The values for the correction factor and the flow behaviour

index for the 50 mm pipeline and the 80 mm pipeline viscometer at different Cw’s are

shown in Tables 8.23 and 8.24.

The values for the Newtonian and non-Newtonian shear stress and shear rate along with

the flow behaviour index and correction factor for the 50 mm pipeline and the 80 mm

pipeline viscometer at a Cw of 66.6 % are shown in Tables 8.25 and 8.26.

232

Table 8.23 Fly Ash “E” 50 mm Pipeline Viscometer.

Cw

(%)

Cw

(w/w)

Flow Behaviour

Index

(𝑛′)

Correction

Factor

58.1 0.581 0.067 3.911

59.2 0.592 0.070 3.777

62.3 0.623 0.121 2.425

63.8 0.638 0.140 2.202

65.4 0.654 0.180 1.929

65.9 0.659 0.200 1.850

66.6 0.666 0.272 1.735

Table 8.24 Fly Ash “E” 80 mm Pipeline Viscometer.

Cw

(%)

Cw

(w/w)

Flow Behaviour

Index

(𝑛′)

Correction

Factor

58.1 0.581 0.062 4.206

59.2 0.592 0.0628 4.172

62.3 0.623 0.115 2.522

63.8 0.638 0.139 2.2121

65.4 0.654 0.191 1.882

65.9 0.659 0.219 1.800

66.6 0.666 0.252 1.748

Table 8.25 Fly Ash “E” 50 mm Pipeline Viscometer Data Cw 66.6%.


Shear

Rate

(s-1)

Shear

Stress

(Pa)

Flow Behaviour

Index

(n')

Correction

Factor

Shear

Rate

(s-1)

Shear

Stress

(Pa)

33.089 21.627 0.272 1.736 57.410 21.629

64.401 30.897 0.272 1.736

111.738 30.897

138.262 46.815 0.272 1.736

239.890 46.815

213.298 67.871 0.272 1.736

370.081 67.871

296.263 89.748 0.272 1.736

514.028 89.748

387.213 115.274 0.272 1.736

671.828 115.274

472.406 140.818 0.272 1.736

819.642 140.818

507.990 151.344 0.272 1.736

881.381 151.343

233

Table 8.26 Fly Ash “E” 80 mm Pipeline Viscometer Data Cw 66.6%.


Shear

Rate (s-1)

Shear

Stress (Pa)

Flow Behaviour

Index (n')

Correction

Factor

Shear

Rate (s-1)

Shear

Stress (Pa)

10.121 14.264 0.252 1.748 17.6964 14.264

19.698 17.043 0.252 1.748 34.442 17.043

42.290 24.821 0.252 1.748 73.943 24.821

65.241 31.496 0.252 1.748 114.072 31.496

90.617 36.534 0.252 1.748 158.441 36.534

118.435 42.928 0.252 1.748 207.081 42.928

144.493 48.575 0.252 1.748 252.642 48.575

155.377 51.665 0.252 1.748 271.673 51.665

Figure 8.30 is a shear diagram of the rotary viscometer results comparing them to the

non-Newtonian pseudo-shear diagram of the 50 mm and 80 mm pipeline viscometers at

different Cw’s.

Figure 8.30 Shear Diagram of the Rotary Viscometer Results compared to a Non-

Newtonian Pseudo-Shear Diagram for the 50 mm and 80 mm Pipeline Viscometer at

different Cw’s.

0

5

10

15

20

25

30

35

40

45

50

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

50 mm Pipe Cw = 58.1 % 80 mm Pipe Cw = 58.1 % RV Cw = 58.1 %50 mm Pipe Cw = 59.2 % 80 mm Pipe Cw = 59.2 % RV Cw = 59.2 %50 mm Pipe Cw = 62.3 % 80 mm Pipe Cw = 62.3 % RV Cw = 62.3 %50 mm Pipe Cw = 63.8 % 80 mm Pipe Cw = 63.8 % RV Cw = 63.8 %50 mm Pipe Cw = 65.4 % 80 mm Pipe Cw = 65.4 % RV Cw = 65.4 %50 mm Pipe Cw = 65.9 % 80 mm Pipe Cw = 65.9 % Rv Cw = 65.9 %50 mm Pipe Cw = 66.6 % 80 mm Pipe Cw = 66.6 % RV Cw = 66.6 %

234

To align the rotary viscometer and pipeline viscometer results, the scale used on the

shear diagrams and the pseudo-shear diagrams was 0 to 200 seconds. These values

reflect the maximum operation range for power station dense phase fly ash pumping

system and grout pumping systems.

Table 8.27 compares the rotary viscometer results with the non-Newtonian results for

50 mm and 80 mm pipeline viscometers at a shear rate of 100 s-1


Table 8.27 Fly Ash “E” Comparison of the Rotary Viscometer Results with those from

the non-Newtonian 50 mm and 80 mm Pipeline Viscometers at a Shear Rate of 100 s-1

at Different Cw’s.

50 mm Pipeline

Viscometer

80 mm Pipeline

Viscometer

Rotary Viscometer

Cw

(%)

(s-1)

𝜏

(Pa)

(s-1)

𝜏

(Pa)

(s-1)

𝜏

(Pa)

58.1 100 7.2 100 7.6 100 4.6

59.2 100 6.8 100 7.6 100 4.9

62.3 100 14.0 100 14.4 100 8.7

63.8 100 19.6 100 19.6 100 12.0

65.4 100 26.3 100 26.3 100 15.5

65.9 100 34.5 100 34.5 100 19.5

66.6 100 38.4 100 38.5 100 23.5

8.14 Slurry Modelling Fly Ash “E”

Bingham plastic models were fitted to the non-Newtonian 50 mm and 80 mm pipeline

viscometer curves as indicated in Figure 8.31. The Bingham models contain the both yield

stress and viscosity at different Cw’s. The Bingham models are displayed in Table 8.28.

To develop a model that predicts the pipeline pressure drop of high Cw fly ash slurries

involves determining the relationships between Cw and the yield stress and viscosity.

These relationships are displayed in Figures 8.32 and 8.33.

235

Figure 8.31 Fly Ash “E” Non-Newtonian Pseudo-Shear Diagram of the 80 mm and 50

mm Pipeline Viscometers at Different Cw’s with fitted Bingham Plastic Models.

Table 8.28 Fly Ash “E” Bingham Plastic Models fitted to 50 mm Pipeline Viscometer

Curves.

Cw

(%)

Cw

(w/w)

Model Yield Stress

(𝜏𝑦)

(Pa)

Viscosity

(𝜂𝑏)

(Pas)

58.1 0.581 τb = 1.881 + 0.054 1.881 0.054

59.2 0.592 τb = 2.020 + 0.052 2.020 0.052

62.5 0.625 τb = 3.788 + 0.103 3.788 0.103

63.8 0.638 τb = 5.762 + 0.128 5.762 0.128

65.4 0.654 τb = 7.263 + 0.168 7.263 0.168

65.9 0.659 τb = 9.573 + 0.208 9.573 0.208

66.6 0.666 τb = 13.606 + 0.237 13.606 0.237

0

10

20

30

40

50

60

70

80

0 50 100 150 200 250 300 350 400

Shea

r St

ress

(P

a)

Shear Rate (s-1)

50 mm Pipe Cw = 58.1 % 80 mm Pipe Cw = 58.1 % Model Cw = 58.1 %50 mm Pipe Cw = 59.2 % 80 mm Pipe Cw = 59.2 % Model Cw = 59.2 %50 mm Pipe Cw = 62.3 % 80 mm Pipe Cw = 62.3 % Model Cw = 62.3 %50 mm Pipe Cw = 63.8 % 80 mm Pipe Cw = 63.8 % Model Cw = 63.8 %50 mm Pipe Cw = 65.4 % 80 mm Pipe Cw = 65.4 % Model Cw = 65.4 %50 mm Pipe Cw = 65.9 % 80 mm Pipe Cw = 65.9 % Model Cw = 65.9 %50 mm Pipe Cw = 66.6 % 80 mm Pipe Cw = 66.6 % Model Cw = 66.6 %

236

Figure 8.32 Fly Ash “E” Relationship between Cw and Bingham Yield Stress.

By fitting a model for yield stress( 𝜏𝑏𝑦) to Figure 8.30, the following relationship was

obtained:

𝑌𝑖𝑒𝑙𝑑 𝑆𝑡𝑟𝑒𝑠𝑠 𝜏𝑏𝑦 = 49495 𝐶𝑤3 − 90386 𝐶𝑤

2 + 55033 𝐶𝑊 − 11169 (8.14)

And by fitting a model for viscosity (𝜂𝑏) to Figure 8.31, the following relationship was

obtained:

𝑉𝑖𝑠𝑐𝑜𝑠𝑖𝑡𝑦 𝜂𝑏

= 258.89 𝐶𝑤3 − 458.46 𝐶𝑤

2 + 271.44 𝐶𝑊 − 53.668 8.15)

Therefore 𝑆ℎ𝑒𝑎𝑟 𝑆𝑡𝑟𝑒𝑠𝑠 𝜏0 = 𝜏𝑦 + 𝜂𝑏 𝛤𝑤 (8.16)

τby = 49495Cw3 - 90386Cw

2 + 55033Cw - 11169

R² = 0.9887

0

50

100

150

200

250

300

350

400

0.58 0.60 0.62 0.64 0.66 0.68 0.70 0.72 0.74 0.76 0.78 0.80

Yeild

Str

ess

(Pa)

Cw

Yeild Stress Vs Cw Model

237

Figure 8.33 Fly Ash “E” Relationship between Cw and Bingham Viscosity.

Integrating equation 8.9 and 8.10 into equation 8.11 gives:

𝑆ℎ𝑒𝑎𝑟 𝑆𝑡𝑟𝑒𝑠𝑠 𝜏0 = 49495 𝐶𝑤3 − 90386 𝐶𝑤

2 + 55033 𝐶𝑊 − 11169 +

(258.89 𝐶𝑤3 − 458.46 𝐶𝑤

2 + 271.44 𝐶𝑊 − 53.668) 𝛤𝑤 (8.17)


substituting in equation 8.16, the pipeline shear stress (𝜏0) was determined. The

pipeline pressure (∆𝑃) was calculated by substituting (𝜏0) in Equation 8.17.

𝑃𝑖𝑝𝑒𝑙𝑖𝑛𝑒 ∆𝑃 = 𝜏0 4 𝐿

𝐷 (8.18)

Figure 8.34 is a graph of calculated pipeline pressure (ΔP) at different Cw’s for a 10 km

long slurry pipeline (L) with a nominal diameter of 200 mm (D) at a flowrate (Q) of 240

m3 h

-1.

ηb = 258.89Cw3 - 458.46Cw

2 + 271.44Cw - 53.668

R² = 0.9987

0

0.5

1

1.5

2

2.5

3

0.58 0.60 0.62 0.64 0.66 0.68 0.70 0.72 0.74 0.76 0.78 0.80

Vis

cosi

ty (P

as)

Cw

Viscosity Vs Cw Model

238

Figure 8.34 Fly Ash “E” Calculated ΔP compared to Cw.

Therefore, a model to determine pipeline pressure drop was:

𝛥𝑃 = 11112 𝐶𝑤3 − 20469 𝐶𝑤

2 + 12591 𝐶𝑊 − 2585.1 (8.19)

8.15 Non-Newtonian Slurry Grout Modelling Fly Ash “E”

When operating a batch grouting plant in the field to maintain the consistency of the

grout the plant operator is required to manually test the grout every batch using a flow

cone. To overcome this repetitive testing, it was proposed that this testing could be

replaced by using pipeline pressure drop of high Cw fly ash grouts when compared to

flow cone time, requires determining the relationships between Cw, flow cone time,

yield stress and viscosity. These relationships are displayed in Figures 8.32, 8.33 and

8.35.

∆P = 11112 × Cw3 - 20469 × Cw

2 + 12591 × Cw - 2585.1

0

5

10

15

20

25

30

35

0.57 0.59 0.61 0.63 0.65 0.67 0.69 0.71 0.73 0.75

Pip

elin

e P

ress

ure

∆P

(Mp

a)

Cw

Fly Ash "E" Slurry Pipeline Non-Newtonian Pressure Vs Cw Model

239

By fitting a model for flow cone time (𝐹𝐶𝑇) from the data shown in Table 8.14, the


𝐶𝑤 = −0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518 (8.20)

Figure 8.35 Fly Ash “E” Grout Flow Cone Time Compared to Cw.

Integrating the relationship from Equation 8.19 into Equation 8.16 the following

relationship was obtained that allow the determination of shear stress (𝜏0) based on the

ASTM Flow Cone time.

𝑆ℎ𝑒𝑎𝑟 𝑆𝑡𝑟𝑒𝑠𝑠 𝜏0 = 49495 (−0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518)3

− 90386 (−0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518)2

+55033 (−0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518) − 11169

+(258.89 (−0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518)3

− 458.46 (−0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518)2

+271.44 (−0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518) − 53.668) 𝛤𝑤 (8.21)

0.58

0.6

0.62

0.64

0.66

0.68

0.7

0.72

0.74

0.76

0 10 20 30 40 50 60 70 80

Cw

Flow Cone Time (s)

Cw Vs Flow Cone Time Model

Cw = -0.5681 × FCT-0.398 + 0.8518

240


substituting in equation 8.20 the pipeline shear stress (𝜏0) was calculated. The pipeline

pressure (∆𝑃) was calculated by substituting (𝜏0) in Equation 8.17. Normal practice

when pumping grout into worked out underground mines is to place the grout plant

close to the area that requires rehabilitation. The process limits the pumping rate from

the grouting plant is the rate at which the dry fly ash is transported and unloaded at the

grouting plant. This limits the pumping flowrate (Q) to nominally 30 m3 h

-1 through an

80 mm (D) nominal bore pipe. Figure 8.36 is a graph of calculated pipeline pressure

(ΔP) at different Flow cone time (FCT) for a 100 m long slurry pipeline (L). At a

calculated pipeline shear stress (𝛤𝑤) of 179 s-1

.

A model to determine fly ash “E” grout pipeline pressure drop was:

𝛥𝑃 = −1301 × (𝐹𝐶𝑇)2 + 39.225 × (𝐹𝐶𝑇) − 302.86 (8.22)

Figure 8.36 Fly Ash “E” Grout Calculated ΔP Compared to Flow Cone Time.

∆P = - 0.1301 × (FCT)2 + 39.225 × (FCT) - 302.86

R² = 0.9988

0

200

400

600

800

1000

1200

1400

1600

1800

2000

0 10 20 30 40 50 60 70

Pip

elin

e P

ress

ure

(kP

a 1

00

m-1

)

Flow Cone Time (s)

Flow Cone Time Vs DP

241

8.16 Fly Ash “E” Slurries Comparison of 50 mm and 80 Pipeline Viscometers

To validate the measuring technique of the 50 mm and 80 mm pipeline viscometers, the

Newtonian and non-Newtonian data was plotted on a pseudo-shear diagram. Figure

8.37 was a pseudo-shear diagram for the Newtonian 50 mm and 80 mm pipeline

viscometer and Figure 8.38 is the non-Newtonian results plotted on the a pseudo-shear

diagram.

Figure 8.37 Fly Ash “E” Newtonian Pseudo-Shear Diagram of the 50 mm and 80 mm

Pipeline Viscometer at different Cw’s.

0

10

20

30

40

50

60

70

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)

50 mm Pipe Cw = 58.1 % 80 mm Pipe Cw = 58.1 % 50 mm Pipe Cw = 59.2 % 80 mm Pipe Cw = 59.2 %


50 mm Pipe Cw = 65.4 % 80 mm Pipe Cw= 65.4 % 50 mm Pipe Cw = 65.9 % 80 mm Pipe Cw = 65.9 %


242

Figure 8.38 Fly Ash “E” Non-Newtonian Pseudo-Shear Diagram of the 50 mm and 80

mm Pipeline Viscometer at different Cw’s.

8.17 Fly Ashes “B” and “E” Slurries Comparison of Non-Newtonian Pipeline

Pressure Drop Models

Comparison of the pressure drop (∆𝑃) at different Cw’s for slurries mixed with fly ash

“B” and “E” for a 10 km long 200 mm pipeline pumped at a flowrate of 240 m3 h

-1 as

indicated in Figures 8.18 and 8.34 are shown on Figure 8.39.

Examination of Figure 8.39 revealed a similarity between fly ash “B” and “E” pressure

drop models characteristics up to a Cw of 75 %. A model was produced that would

predict the pressure drop characteristics.

𝑃𝑟𝑒𝑠𝑠𝑢𝑟𝑒 𝐷𝑟𝑜𝑝 ∆𝑃 = 12141 𝐶𝑤3 − 22619 𝐶𝑤

2 + 14075 𝐶𝑊 − 2923.4 (8.23)

The model is displayed in Figure 8.40. It can be concluded that this model can be used

to calculate the pressure drop in a high concentration slurry pipeline up to a Cw of 75 %

provided the fly ash PSD falls within the PSD envelope indicated in Figure 8.41.

0

5

10

15

20

25

30

35

40

45

50

0 20 40 60 80 100 120 140 160 180 200

Shea

r St

ress

(P

a)

Shear Rate (s-1)





243

Table 8.29 Fly Ashes “B” and “E” d10, d50, d90 and Density.

Fly Ash “B”

Average

Fly Ash “E”

Average

d10 d50 d90 Density

t m-3

d10 d50 d90

Density

t m-3

4.4 23.8 80.9 2.0360 5.9 34.9 132.3 2.0862

Figure 8.39 Fly Ashes “B” &“E” Slurries Pipeline Pressure Models.

0

5

10

15

20

25

30

35

0.59 0.61 0.63 0.65 0.67 0.69 0.71 0.73 0.75

Pip

elin

e P

ress

ure

∆P

(MP

a)

Cw

Fly Ash "E" Slurry Pipeline Pressure Vs Cw Fly Ash "B" Slurry Pipeline Pressure Vs Cw

244

Figure 8.40 Fly Ashes “B” &“E” Slurries Pipeline Pressure Model.

Figure 8.41 Fly Ashes “B” & “E” PSD.

∆P = 12141 × Cw3 - 22619 × Cw

2 + 14075 × Cw - 2923.4

0

5

10

15

20

25

30

35

0.59 0.61 0.63 0.65 0.67 0.69 0.71 0.73 0.75

Pip

elin

e P

ress

ure

∆P

(MP

a)

Cw

Fly Ash "E" Slurry Pipeline Pressure Vs Cw Fly Ash "B" Slurry Pipeline Pressure Vs CwFly Ash Slurry PSD Evenlope Model

0

10

20

30

40

50

60

70

80

90

100

0.1 1 10 100 1000 10000

Pe

rce

nta

ge P

siin

g (%

)

Particle Size (µm)

E PSD 1 E PSD 2 E PSD 3 E PSD 4 E PSD 5 E PSD 6 E PSD 7

E PSD 8 E PSD 9 B PSD 1 B PSD 2 B PSD 3 B PSD 4 B PSD 1A

B PSD 2A B PSD 3A B PSD 4A B PSD 5A B PSD 6A

Purposed PSD Envelope

Proposed PSD Envelope

245

8.18 Fly Ash “E” Determining the Settling Velocity

On examination of the flow of the fly ash “E” slurries through the 80 mm glass at each

speed step during the pumping cycle at the higher Cw’s it was observed that fly ash

particles were moving over the total inside surface of the glass. The speed step at which

this occurred was noted. The velocity at this speed step was then calculated using

equation 8.1. This data was presented in Table 8.28 and Figure 8.42. Figure 8.43 is a

graph of the settling velocity compared to the slurry viscosity and slows a linear

relationship. The model for the velocity compared to the fly ash slurry viscosity:

𝑉𝑒𝑙𝑜𝑐𝑖𝑡𝑦 (𝑉) = 0.0041(μa) + 1.7613 (8.24)

Table 8.28 Fly Ash “E” Observed Full Pipe Flow.

Cw

(%)

Observed Full Pipe Flow

V

(m s-1

)

Calculated Viscosity

(m Pas)

60.2 1.28 119.6

62.3 1.24 124.4

63.8 1.13 150.2

65.4 0.92 193.9

65.9 0.80 245.0

66.6 0.41 320.8

246

Figure 8.42 Fly Ash “E” Slurries Settling Velocity.

Figure 8.43 Fly Ash “E” Slurries Viscosity.

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

59 60 61 62 63 64 65 66 67

Vel

oci

ty (m

s-1

)

Cw (%)

Suspension Velocity Vs Cw

V = -0.0041x + 1.7613

R² = 0.9907

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

0 50 100 150 200 250 300 350

Vel

oci

ty (m

s-1

)

Viscosity µa (mPas)

Velocity Vs Viscosity

247

8.19 Fly Ash “B” and “E” Laminar or Turbulent Flow

The laminar flow conditions of non-Newtonian slurry in a tube viscometer can be verified

by showing that the Reynolds Number (𝑅𝑒𝑛𝑒𝑤𝑡) was less than 2100 using the relationship:

𝑅𝑒𝑁𝑒𝑤𝑡 = 𝜌𝑠𝑙 𝑉𝐷

𝜇𝑎 (8.25)

The calculated Reynolds Number (𝑅𝑒𝑁𝑒𝑤𝑡) for fly ashes “B” and “E” at a shear rate of

100 s-1

are displayed in Tables 8.30, 8.31 and 8.32.

For comparison with an operating power station, dense phase plant Test 5 was the data

from Ward el al (1998) for the 26th April 1996 and Test 6 was data collect from

Bayswater by the author on the 19th March 2013. The Reynolds Number for 50 mm and

80 mm pipeline viscometers was extracted using a shear rate (𝛤𝑤) of 100 s-1

. The

Reynolds Number for the Bayswater high concentration pumping system Test 5 and 6

using a shear rate (𝛤𝑤) of 88 s-1 calculated for a nominal plant slurry flow of 240 m

3 h

-1

through a 200 mm diameter pipeline.

Table 8.30 Fly Ashes “B” Reynolds Numbers Shear Rate 100 s-1.

Test

No.

(m)

D

(%)

Cw

V

(m s-1

)

𝜌𝑠𝑙

(kg m-1

)

𝛤𝑤

(s-1)

𝜏0

(Pa)

𝜇𝑎

(Pas)

𝑅𝑒𝑁𝑒𝑤𝑡

1 0.0525 59.7 0.65 1436 100 3.3 0.033 1485

2 0.0525 61.8 0.65 1459 100 10 0.1 498

3 0.0525 65.1 0.65 1495 100 17 0.17 300

4 0.0525 67.9 0.65 1528 100 30.6 0.306 170

1 0.0779 59.7 0.97 1436 100 3.3 0.033 3288

2 0.0779 61.8 0.97 1459 100 9.8 0.098 1124

3 0.0779 65.1 0.97 1495 100 17.2 0.172 657

4 0.0779 67.9 0.97 1528 100 31 0.31 372

5 0.2 73.0 2.2 1666 88 32 0.364 2016

6 0.2 69.0 2.2 1530 88 32.5 0.369 1823

248

Table 8.31 Fly Ashes “E” Reynolds Numbers for 50 mm Pipeline Viscometer Shear Rate

100 s-1

.

Test

No.

D

(m)

Cw

(%)

V

(m s-1

)

𝜌𝑠𝑙

(kg m-1

)

𝛤𝑤

(s-1)

𝜏0

(Pa)

𝜇

(Pas)


1 0.0525 58.1 0.65 1433 100 3.2 0.032 1528

2 0.0525 59.2 0.65 1446 100 3.3 0.033 1495

3 0.0525 62.3 0.65 1481 100 8 0.08 632

4 0.0525 63.8 0.65 1498 100 12 0.12 426

5 0.0525 65.3 0.65 1516 100 16.4 0.164 315

6 0.0525 65.9 0.65 1522 100 21.8 0.218 238

7 0.0525 66.6 0.65 1531 100 29.2 0.292 179


100 s-1

.

Test

No.

D

(m)

Cw

(%)

V

(m s-1

)

𝜌𝑠𝑙

(kg m-1

)

𝛤𝑤

(s-1)

𝜏0

(Pa)

𝜇

(Pas)


1 0.0779 58.1 0.97 1433 100 3.2 0.032 3384

2 0.0779 59.2 0.97 1446 100 3.3 0.033 3311

3 0.0779 62.3 0.97 1481 100 7.5 0.075 1492

4 0.0779 63.8 0.97 1498 100 12 0.12 943

5 0.0779 65.3 0.97 1516 100 16.7 0.167 689

6 0.0779 65.9 0.97 1522 100 23 0.23 500

7 0.0779 66.6 0.97 1531 100 29.3 0.293 395


150 s-1


249


Test

No.

(m)

D

(%)

Cw

V

(m s-1

)

𝜌𝑠𝑙

(kg m-1

)

𝛤𝑤

(s-1)

𝜏0

(Pa)

𝜇𝑎

(Pas)


1 0.0525 59.7 0.98 1436 150 4.3 0.029 2578

2 0.0525 61.8 0.98 1459 150 11.5 0.077 979

3 0.0525 65.1 0.98 1495 150 19.7 0.131 586

4 0.0525 67.9 0.98 1528 150 38 0.253 310

1 0.0779 59.7 1.46 1436 150 4.1 0.027 5975

2 0.0779 61.8 1.46 1459 150 11.2 0.075 2222

3 0.0779 65.1 1.46 1495 150 19.7 0.131 1294

4 0.0779 67.9 1.46 1528 150 38 0.253 689

5 0.2 73.0 2.2 1666 88 32 0.364 2016

6 0.2 69.0 2.2 1530 88 32.5 0.369 1823


150 s-1

.

Test

No.

(m)

D

(%)

Cw

V

(m s-1

)

𝜌𝑠𝑙

(kg m-1

)

𝛤𝑤

(s-1)

𝜏0

(Pa)

𝜇

(Pas)


1 0.0525 58.1 0.98 1433 150 3.9 0.026 2836

2 0.0525 59.2 0.98 1446 150 3.9 0.026 2861

3 0.0525 62.3 0.98 1481 150 10.2 0.068 1121

4 0.0525 63.8 0.98 1498 150 15.2 0.101 761

5 0.0525 65.3 0.98 1516 150 22 0.147 532

6 0.0525 65.9 0.98 1522 150 28 0.187 420

7 0.0525 66.6 0.98 1531 150 35 0.233 338

250


150 s-1

.

Test

No.

(m)

D

(%)

Cw

V

(m s-1

)

𝜌𝑠𝑙

(kg m-1

)

𝛤𝑤

(s-1)

𝜏0

(Pa)

𝜇

(Pas)


1 0.0779 58.1 1.46 1433 150 3.9 0.026 6268

2 0.0779 59.2 1.46 1446 150 3.95 0.0263 6245

3 0.0779 62.3 1.46 1481 150 10.2 0.068 2477

4 0.0779 63.8 1.46 1498 150 14.8 0.099 1727

5 0.0779 65.3 1.46 1516 150 22.5 0.15 1149

6 0.0779 65.9 1.46 1522 150 29 0.193 895

7 0.0779 66.6 1.46 1531 150 35.5 0.237 736


200 s-1



Test

No.

(m)

D

(%)

Cw

V

(m s-1

)

𝜌𝑠𝑙

(kg m-1

)

𝛤𝑤

(s-1)

𝜏0

(Pa)

𝜇𝑎

(Pas)


1 0.0525 59.7 1.31 1436 200 5.2 0.026 3798

2 0.0525 61.8 1.31 1459 200 13 0.065 1544

3 0.0525 65.1 1.31 1495 200 23.6 0.118 871

4 0.0525 67.9 1.31 1528 200 45.5 0.228 462

1 0.0779 59.7 1.95 1436 200 4.8 0.024 9089

2 0.0779 61.8 1.95 1459 200 12.8 0.064 3463

3 0.0779 65.1 1.95 1495 200 23.6 0.118 1925

4 0.0779 67.9 1.95 1528 200 45.5 0.228 1020

5 0.2 73.0 2.2 1666 88 32 0.364 2016

6 0.2 69.0 2.2 1530 88 32.5 0.369 1823

251


200 s-1

.

Test

No.

(m)

D

(%)

Cw

V

(m s-1

)

𝜌𝑠𝑙

(kg m-1

)

𝛤𝑤

(s-1)

𝜏0

(Pa)

𝜇

(Pas)


1 0.0525 58.1 1.31 1433 200 4.5 0.0225 4380

2 0.0525 59.2 1.31 1446 200 4.5 0.0225 4419

3 0.0525 62.3 1.31 1481 200 12.6 0.063 1616

4 0.0525 63.8 1.31 1498 200 19 0.095 10848

5 0.0525 65.3 1.31 1516 200 27 0.135 772

6 0.0525 65.9 1.31 1522 200 34 0.17 616

7 0.0525 66.6 1.31 1531 200 41.5 0.2075 507


200 s-1

.

Test

No.

(m)

D

(%)

Cw

V

(m s-1

)

𝜌𝑠𝑙

(kg m-1

)

𝛤𝑤

(s-1)

𝜏0

(Pa)

𝜇

(Pas)


1 0.0779 58.1 1.95 1433 200 4.7 0.0235 92629

2 0.0779 59.2 1.95 1446 200 4.2 0.021 10460

3 0.0779 62.3 1.95 1481 200 12.6 0.063 3571

4 0.0779 63.8 1.95 1498 200 18.2 0.091 2500

5 0.0779 65.3 1.95 1516 200 27.3 0.1365 1687

6 0.0779 65.9 1.95 1522 200 34.7 0.1735 13321

7 0.0779 66.6 1.95 1531 200 42 0.21 11076

252

8.20 Fly Ash “B” and “E” Homogeneous or Heterogeneous Slurries

It is important to determine whether these slurries are homogeneous or heterogeneous

suspensions. Homogeneous flow is a symmetric flow characterizing uniform

distribution of solids about the horizontal axis of the pipe. Durand and Condolios

(1952) published a number of studies indicating that homogeneous suspensions were

those that contained all particles smaller than 40 μm while Shook et al. (2002)

suggested that for suspensions with a mean particle diameter (d50) greater than 50 μm

the slurry would display heterogeneous properties. He also indicated that fine part icle

slurries d50 less than 50 μm typically exhibit homogeneous fluid behaviour.

Fly ash particles from modern coal fired power stations are nominally spherical with a

d50 ranging from 8 to 45 µm, therefore, some of the fly ashes with low d50 should be

classified as homogenous, but which ones?

Thomas (1967) outlined that if pipe loop tests are performed on slurries at the desired

Cw in a number of different diameter pipes, and the pressure gradient verses velocity is

plotted on a log-log plot, if the results are a straight line, then the slurry is

homogeneous. Figure 8.44 was a log–log plot of pressure gradient verses velocity for

the fly ash “B” slurries at selected Cw’s and water for both pipe sizes. Figure 8.45 was a

log–log plot of pressure gradient verses velocity for the fly ash “E” slurries at different

Cw’s and water for both pipe sizes. From Figures 8.44 and 8.45, it can be determined

that the water curves are straight lines, therefore are consistent with homogeneous flow.

The graphs for the fly ash slurries are not straight lines indicating heterogeneous flow.

253

Figure 8.44 Fly Ash “B” Slurries and Water.

Another measure of homogeneous flow is when comparative rheometry tests were

conducted using a rotary viscometer and a pipeline viscometer. When the results of the

rotary viscometer tests were plotted on a shear diagram and the pipeline viscometer

results were plotted on a pseudo-shear diagram, the results for the fly ash slurry at the

same Cw were similar.

Inspection of Figure 8.14 and Table 8.13 for fly ash “B” slurries and Figure 8.30 and

Table 8.27 for fly ash “E” slurries reveals that the rotary viscometer shear stresses at a

shear rate of 100 s-1

are approximately 15 % lower at all Cw’s. Therefore, by definition

all the slurries tested are clearly heterogeneous.

0.01

0.1

1

10

0.01 0.1 1 10

Pre

ssu

re G

rad

ien

t (kP

a m

-1)

Velocity (m s-1)

80 mm Pipe Cw 59.7 % 50 mm Pipe Cw 59.7 % 80 mm Pipe Cw 61.8 %50 mm Pipe Cw 61.8 % 80 mm Pipe Cw 65.1 % 50 mm Pipe Cw 65.1 %80 mm Pipe Cw 67.9 % 50 mm Pipe Cw 67.9 % 80 mm Pipe Water50 mm Pipe Water

254

Figure 8.45 Fly Ash “E” Slurries and Water.

8.21 New Definition for Fly Ash Slurries Homogeneous Behaviour

Bunn (1991) reported that on one occasion when conducting comparative rheometry

tests on fly ash slurries from Bayswater Power Station that the rotary viscometer results

coincided with a pipeline viscometer results. Refer to Figure 5.9. The PSD for 5th

May

1990 was plotted on the PSD Data for fly ash “B”. Refer to Figure 8.46. Table 8.39

compares PSD for fly ashes “B” and “E” with PSD data from Bayswater fly ash from

1990 Bunn (1991).

A model was proposed to measure the slope, 𝑑𝑠 of the PSD curves:

𝑑𝑠 =𝑑90 − 𝑑10

𝑑50 (8.26)

The slope of the average of the 10 data PSD points for fly ash “B” was 3.21 while the

slope of the average of the 9 data PSD points for fly ash “E” was 3.44. The slope of the

PSD data from Bunn (1991) for the 5th May was 3.13.

0.001

0.01

0.1

1

10

0.01 0.1 1

Pre

ssu

re G

rad

ien

t (k

Pa

m-1

)





80 mm Pipe Cw 66.6 % 50 mm Pipe Cw 66.6 % 80 mm Pipe Water 50 mm Pipe Water

255

From the results of this research and the author’s previous research, it can now be

proposed that homogeneous fluid behaviour occurs in fly ash slurries containing

particles with a d50 less than 15 μm and a PSD curve slope 𝑑𝑠 of less than 3.13.

Figure 8.46 Fly Ash “B” PSD and PSD for Bayswater Fly Ash from Bunn (1991).

Table 8.39 PSD of Fly Ashes “B” and “E” Bayswater Fly Ash from Bunn (1991).

Fly Ash d10

µm

d50

µm

d90

µm

ds

µm

B 4.4 23.8 80.9 3.21

E 5.9 34.9 132.3 3.62

BW (1991) 3 15 50 3.13

Figure 8.47 is a column graph displaying the d10, d50 and d90 of the Bayswater fly ash

used for pumping and Bayswater fly ash data from Bunn (1991).

0

10

20

30

40

50

60

70

80

90

100

0.1 1 10 100 1000

Per

cen

tage

Pas

sin

g (%

)

Particle Size (µm)

PSD 1 PSD 2 PSD 3PSD 4 PSD 1A PSD 2APSD 3A PSD 4A PSD 5APSD 6A BW PSD 5th May 1990

256

Figure 8.47 Fly Ash “B” PSD d10, d50 and d90 and PSD for Bayswater Fly Ash from Bunn

(1991).

8.22 Spread Sheet Program

A spread sheet program has been developed that allows the input of a power station

operating parameters to ascertain the pressure drop characteristics of a dense phase ash

pumping system.

The program is divided into three sections;

Stage 1 determine the operating system parameters,

Stage 2 calculating pipeline size,

Stage 3 pipeline calculations.

Program Stage 1. The program requests that you insert what is the unit configuration of

the proposed dense phase ash pumping system, the number of unit, and if the bottom

and fly ash unit system integrated or separate. This initial stage of the program requires

a yes or no answer. There should only be two yes’s in section 1. After that the program

requires inputting the tons of coal burnt per hour per unit at the maximum overload rate

4.7 4.8 4.8 4.8 4.2 4.4 4.0 4.3 4.1 4.1 4.3 3.0

25.2 22.7 23.4 25.2 22.9 25.5 21.2

24.2 22.3

24.3 25.5

15.0

89.5

83.0 77.8

88.0

75.9

86.5

67.1

80.9

72.5

82.8 86.2

52.0

0

10

20

30

40

50

60

70

80

90

100

1 2 3 4 5 6 7 8 9 10 11 12

Par

ticl

e Si

ze (µ

m)

Fly Ash d10 d50 & d90 1 to 11, 12 - 5th May 1990

d10 d50 d90

257

per unit and the ash content of the coal. The program calculates the amount of ash to be

pumped of per hour. The split between the percentage fly ash to bottom ash is assumed.

Stage 1 determine the operating system parameters

Inputs System

Coal

System

Ash

System

Fly Ash

t h-1

t h-1

t h-1

Bottom and fly ash plants disposal

system to operate as a unit plant? Y/N No


system to operate as a multi-unit plant?

Y/N

No

If multi-unit plant how many units? 2

Bottom and fly ash unit system

integrated? Y/N No

Bottom and fly ash unit system

separate? Y/N No

Bottom and fly ash multi-unit system

integrated? Y/N No

Bottom and fly ash multi-unit system

separate? Y/N No


system to operate as a station plant?

Y/N

Yes 1600

Bottom and fly ash system integrated?

Y/N No

Bottom and fly ash system separate?

Y/N Yes

280.5

How many units? 4

How many tons of coal burnt per hour

per unit at maximum overload rate per

unit?

330

Maximum coal ash content (%)? 25

Percentage of bottom ash (%) 15

Percentage of fly ash (%) 85

For example: - A power station burning bituminous coal for every 1000 MWe per day

produced burns approximately 11,314 t/day of coal (Wikipedia 20104. For Station with

4 x 660 MW units (2640 MW) with an overload capability of 4 x 700 MW could burn

up to 31 680 t/day of coal Macgen, (2014).

258

Burning 31 680 t/day of coal with an ash content of 25 %, with a nominal split between

bottom ash and fly ash of 15% bottom and 85 % fly ash, produces 7 920 t/day (330 t/h)

of ash consisting of 6732 t/day (280.5 t/h) of fly ash and 1 430 t/day (49.5 t/h) of bottom

ash.

Program Stage 2. Thesis section of the program calculates the pipeline size based on

the mass of ash to be transported from section 1. The calculated pipe diameter is

compared to commercially available pipes and the closest pipe diameter was selected.

The program then calculates the pipeline flow and pump capacity.

Stage 2 Calculating pipeline size Inputs

Mass of ash to be pumped. 280.5

Determine volume flow rate of pipeline (m3 h

-1). 280.5

Assume approximate volume flow rate

(m3 h

-1) = (t h

-1) of ash.

Assumed Velocity (m s-1

). 1.5

Calculate pipeline area. A = Q/V (m2). 0.052

Calculate pipeline inside diameter from D = √ (4A/π) (m). 0.257

Determine near commercial pipe diameter (m). 0.254

Pipe Nominal Bore (mm). 250

Calculate new flow using pipe diameter and velocity of

(1.5 m s-1

) (m3 h

-1).

274

Pump Capacity (m3 h

-1) 274

Assume approximate volume flow rate

(m3 h

-1) = (t h

-1) of ash.

Determine new fly ash flow (t h-1

). 274

Program Stage 3. The first section of the program calculates the pipeline shear rate

from inputs of pipeline length and the maximum pipeline design pressure. The

operating pressure is set at half the design pressure. The values of yield stress and

viscosity from the thesis for fly ash “E” along with the calculated shear stress are used

to calculate the pipeline pressure at different Cw’s. Figure 8.48 is a graph of output from

the program indicating the relationship between Cw and pipeline pressure. The

calculated Cw corresponding to the pipeline operating pressure of 6 MPa was 69.7 %.

259

Stage 3 Pipeline calculations Inputs

What is the pipeline length? (m). 10000

Maximum pipeline pressure (MPa). 12

Maximum operating pipeline pressure (MPa). 6

Calculate pipeline shear rate (s-1

). 47.24

Calculate Yield Stress with Cw 0.55 (w/w) to 0.75 (w/w).

Calculate Viscosity with Cw 0.55 (w/w) to 0.75 (w/w).

Calculation pipeline shear stress Cw 0.55 (w/w) to 0.75 (w/w).

Calculation pipeline pressure drops Vs Cw.

Output Graph of Cw Vs DP.

Figure

8.48

Cw at 6 MPa (%). 69.7

Figure 8.48 Calculated Pipeline Pressure Compared at Different Cw’s.

0

2

4

6

8

10

12

14

16

18

20

0.6 0.61 0.62 0.63 0.64 0.65 0.66 0.67 0.68 0.69 0.7 0.71 0.72 0.73 0.74 0.75

Pip

elin

e P

ress

ure

(M

Pa)

Cw (w/w)

Cw Vs Pipeline Pressure

260

8.23 Determine the Standard Error of the Models

The standard error of the mean (SE of the mean) estimates the variability between

sample means from multiple samples from the same population. The standard error of

the mean estimates the variability between samples whereas the standard deviation

measures the variability within a single sample.

The standard error of the sample is an estimate of how far the sample mean is likely to

be from the population mean.

𝑆𝐸 = 𝑠

√𝑛 (27)

s is the sample standard deviation n is the size of the sample.

However, to validate a model, the goodness of fit R2 was subsequently calculated.

Generally, R2 is the indicator for the fitting quality, and can be computed by the

following form:

𝑅2 = 1 − 𝑆𝑆𝐸

𝑆𝑆𝑇 (28)

where SSE is the sum of squares due to error, and SST is the total sum of squares.

R2

can take on any value between 0 and 1, with a value closer to 1 indicating that a

greater proportion of variance is accounted for by the model.

Figures 8.49 and 8.50 are the graphs of the models for fly ash “B” and “E” with the

error bar included.

https://en.wikipedia.org/wiki/Standard_deviation#Corrected_sample_standard_deviation

261

Figure 8.49 Fly Ash “B” Calculated ΔP compared to Cw Including Error Bars

Figure 8.50 Fly Ash “E” Calculated ΔP compared to Cw Including Error Bars

∆P = 11036Cw3 - 20959Cw

2 + 13299Cw - 2817.2

R² = 0.997

0

5

10

15

20

25

30

35

0.59 0.61 0.63 0.65 0.67 0.69 0.71 0.73 0.75 0.77

Pip

elin

e P

ress

ure

∆P

(Mp

a)

Cw

∆P = 11112 × Cw3 - 20469 × Cw

2 + 12591 × Cw - 2585.1

R2 = 0.999

0

5

10

15

20

25

30

35

0.57 0.62 0.67 0.72 0.77

Pip

elin

e P

ress

ure

∆P

(Mp

a)

Cw

262

CHAPTER 9 CONCLUSIONS

9.1 Introduction

This body of work has provided a new insight to the characteristics of high

concentration slurry pumping of power station ash. This is a significant problem to the

power industry and this work has produced a new and original approach for the

determination of slurry pipeline pressure drop: the fundamental basis of design. A

software based design tool has been developed that encompasses the outputs of the

work which will provide a practical outcome for the benefit of the power industry. A

model has been presented that used a new method of characterising homogeneous

behaviour of power station ash slurry based on particle size distribution, slope factor

and the median particle size

In the previous chapter, hydraulic conveying trials were conducted which resulted in the

collection of a large quantity of data. Processing of this data led to a number of valuable

correlations which will be of key importance in the development and assessment of a

successful pressure drop prediction model.

9.2 Pipeline Viscometers Water Tests

The generated water curves show extremely good correlation to the established

water curves calculated from the Hazen Williams equation.

The results of the water test indicate that the system was operating as expected.

9.3 Fly Ash “B” Characteristics

The PSD curves and the graph of the d10, d50 and d90 show that there was little if

any attrition of the fly ash particle during pumping of “B” fly ash slurry.

263

The d50 (23.8 µm) and density (2036 kg m-3

) are similar to previous results of

PSD and density for fly ash “B”.

Scanning electron microscope photographs indicated that fly ash “B” particles

are predominantly spherical in shape.

9.4 Comparison of Slurry Flows Measurements

Comparison of the fly ash “B” slurry volumetric flowrate as measured by the

Foxboro Magnetic Flow Meter and the volumetric flowrate calculated from the

mass flow measured by the weight hopper, Cw and density showed excellent

correlations.

9.5 Testing Fly Ash “B” Slurry in Test Facility

The graphs of slurry flow and velocity indicated that:

As the pipe diameter increased, the pressure gradient decreased;

As the flowrate increased, the pressure gradient increased;

As the Cw increased, the pressure gradient increased;

Slurries with a Cw up to 61.8 % as indicated by the upwards sloping curve

behaved as a Newtonian fluid;

Slurries with a Cw greater than 61.8 % behaved as Non-Newtonian fluid;

and,

The frictional losses for water are much less than those for the slurries.

For an ASTM flow cone, as the Cw increased, the flow times increased.

The comparison of the shear diagram for the rotary viscometer and the

Newtonian pseudo-shear diagram for the 50 mm and 80 mm pipeline

viscometers indicated that:

As the shear rate increased, the shear stress increased;

264

As the Cw increased the shear stress increased, at similar shear rates;

At a similar shear rate and Cw the shear stress for the 50 mm and 80 mm

pipeline viscometer showed excellent correlation;

At the same Cw and shear rate, the rotary viscometer shear stresses were

approximately 40 % lower than the pipeline viscometers; and,

In a full scale system, a rotary viscometer would grossly under estimate a

slurry system pipeline pressure drop.

9.6 Non-Newtonian Fly Ash “B” Slurry Characteristics

The Weissenberg-Rabinowitsch equation was applied to the Newtonian data to

convert to non-Newtonian characteristics.

The comparison of the non-Newtonian shear diagram for the rotary viscometer

and the pseudo-shear diagram for the 50 mm and 80 mm pipeline viscometers

indicated that:


As the Cw increased, the shear stress increased at similar shear rates;

At a similar shear rate and Cw, the shear stress for the 50 mm and 80 mm



approximately 15 % lower than the pipeline viscometers,

In a full scale system, a rotary viscometer would under estimate a slurry

system pipeline pressure drop; and,

The slurries density varied from 1436 kg m-3

for slurry with a Cw of 59.7

% to 1528 kg m-3

for slurry with a Cw of 67.9 %.

9.7 Non-Newtonian Slurry Modelling Fly Ash “B”

Bingham visco-plastic models provided positive correlation to the pipeline

viscometers data.

265

A model was produced that enabled the calculation of a pipeline pressure drop

for a power station slurry pipeline at different Cw’s.

To apply the model, the shear rate (𝛤𝑤) has to be determined from the pipeline

and flow parameters.

The shear stress was calculated from the Bingham yield stress and viscosity by

applying the model at different Cw’s.

𝑌𝑖𝑒𝑙𝑑 𝑆𝑡𝑟𝑒𝑠𝑠 (𝜏𝑏𝑦) = 33519𝐶𝑤3 − 64665𝐶𝑤

2 + 41692𝐶𝑊 − 8973.2

𝑉𝑖𝑠𝑐𝑜𝑠𝑖𝑡𝑦 (𝜂𝑏) = 429.14𝐶𝑤

3 − 3796.29𝐶𝑤2 + 492.93 𝐶𝑊 − 101.77

𝑆ℎ𝑒𝑎𝑟 𝑆𝑡𝑟𝑒𝑠𝑠 (𝜏0) = (33519𝐶𝑤3 − 64665𝐶𝑤

2 + 41692𝐶𝑊 − 8973.2) +

(429.14𝐶𝑤3 − 3796.29𝐶𝑤

2 + 492.93𝐶𝑊 − 101.77) 𝛤𝑤.

The pressure drop was calculated by using the pipeline parameter of a 10 km

long slurry pipeline (L) with a nominal diameter of 200 mm (D) at a flowrate

(Q) of 240 m3 h-1


The pressure drop model for fly ash ”B” was;

𝛥𝑃 = 11036𝐶𝑤3 − 20959𝐶𝑤

2 + 13299𝐶𝑊 − 2817.2.

9.8 Site Collected Data Comparison

The site data was collected from Bayswater; for a pipeline flow of 240 m3 h

-1 with

a pipeline pressure drop of 6.8 MPa, a fly ash flow of 260 t h-1

and a water flow of

115 t h-1

. Using these figures the slurry Cw was calculated to be 69.3 %.

The estimated pressure drop from the model was 7.9 MPa.

266

9.9 Non- Newtonian Slurry Grout Modelling Fly Ash “B

A model was produced comparing flow cone time (FCT) to Cw.

The model was 𝐶𝑤 = −0.689 × (𝐹𝐶𝑇)−0.583 + 0.805.

The determination of shear stress (𝜏0) based on the ASTM Flow Cone time:

𝑆ℎ𝑒𝑎𝑟 𝑆𝑡𝑟𝑒𝑠𝑠 (𝜏0) = 33519 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805)3

− 64665 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805)2

+41692 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805) − 8973.2

+ 429.14 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805)3

− 3796.29 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805)2

+492.93 × (−0.689 × 𝐹𝐶𝑇−0.583 + 0.805) − 101.77 × (𝛤𝑤)

The grout pressure drop was calculated by using the pipeline parameter for a 100

m long slurry pipeline (L) with a nominal diameter of 80 mm (D) at a flowrate

(Q) of 30 m3 h

-1 at different Cw’s.

The pressure drop grout model for fly ash ”B” was;

𝛥𝑃 = −0.0986 × (𝐹𝐶𝑇)2 + 27.923 × (𝐹𝐶𝑇) − 238.27

9.10 Fly Ash “B” Slurries Comparison of 50 mm and 80 Pipeline Viscometers

There was a positive correlation between Newtonian data collected from the 50

mm pipeline viscometer when compared to the 80 mm pipeline viscometer.

There was a positive correlation between non-Newtonian data collected from the

50 mm pipeline viscometer when compared to the 80 mm pipeline viscometer.

267

9.11 Fly Ash “E” Characteristics

The PSD curves and the graph of the d10, d50 and d90 show that there was little

attrition the fly ash particle during pumping of the “E” fly ash slurry.

The d50 (34.9 µm) and density (2.0862 kg m-3

) are similar to previous results of

PSD and density for fly ash “E”.

9.12 Testing Fly Ash “E” Slurry in Test Facility

The graphs of slurry flow and velocity indicated that:

As the pipe diameter increased, the pressure gradient decreased;

As the flowrate increased, the pressure gradient increased;

As the Cw increased, the pressure gradient increased;

The changeover point from Newtonian fluid to a Non-Newtonian fluid was

similar to fly ash “B”; and,

The frictional losses for water are much less than those for the slurries.

For an ASTM flow cone, as the Cw increased, the flow times increased;

The comparison of shear diagram for the rotary viscometer and the Newtonian

pseudo-shear diagram for the 50 mm and 80 mm pipeline viscometers indicated

that:

As the shear rate increased, the shear stress increased:

As the Cw increase, the shear stress increased at similar shear rates;




approximately 40 % lower than the pipeline viscometer; and,

In a full scale system, a rotary viscometer will grossly under estimate a

slurry system pipeline pressure drop.

268

9.13 Non-Newtonian Fly Ash “E” Slurry Characteristics

The Weissenberg-Rabinowitsch equation was applied to the Newtonian data to

convert to non-Newtonian characteristics.

The comparison of the non-Newtonian shear diagram for the rotary viscometer

and the pseudo-shear diagram for the 50 mm and 80 mm pipeline viscometers

indicated that:


As the Cw increased, the shear stress increased at similar shear rates;




approximately 15 % lower than the pipeline viscometers;

In a full scale system, a rotary viscometer would under estimate a slurry

system pipeline pressure drop; and,

The slurries density varied from 1443 kg m-3

for slurry with a Cw of 58.1

% to 1531 kg m-3

for slurry with a Cw of 66.6 %.

9.14 Non-Newtonian Slurry Modelling Fly Ash “E”

Bingham visco-plastic models provided positive correlation to the pipeline

viscometers data.

A model was produced that enabled the calculation of a pipeline pressure drop

for a power station slurry pipeline at different Cw’s.

To apply the model, the shear rate (𝛤𝑤) has to be determined from the pipeline

and flow parameters.

The shear stress was calculated from the Bingham yield stress and viscosity by

applying the model at different Cw’s.

269

𝑌𝑖𝑒𝑙𝑑 𝑆𝑡𝑟𝑒𝑠𝑠 𝜏𝑦 = 49495 𝐶𝑤3 − 90386 𝐶𝑤

2 + 55033 𝐶𝑊 − 11169

𝑉𝑖𝑠𝑐𝑜𝑠𝑖𝑡𝑦 𝜂𝑏

= 258.89 𝐶𝑤3 − 458.46 𝐶𝑤

2 + 271.44 𝐶𝑊 − 53.668

𝑆ℎ𝑒𝑎𝑟 𝑆𝑡𝑟𝑒𝑠𝑠 𝜏0 = 49495 𝐶𝑤3 − 90386 𝐶𝑤

2 + 55033 𝐶𝑊 − 11169 +

(258.89 𝐶𝑤3 − 458.46 𝐶𝑤

2 + 271.44 𝐶𝑊 − 53.668) 𝛤𝑤.



(Q) of 240 m3 h-1 at different Cw’s.




The pressure drop model for fly ash ”E” was;

𝛥𝑃 = 11112 𝐶𝑤3 − 20469 𝐶𝑤

2 + 12591 𝐶𝑊 − 2585.1

9.15 Non- Newtonian Slurry Grout Modelling Fly Ash “E”

A model was produced comparing flow cone time (FCT) to Cw.

The model was 𝐶𝑤 = −0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518.

The determination of shear stress (𝜏0) based on the ASTM Flow Cone time:

Shear Stress τ0 = 49495 (-0.5681 × ((𝐹𝐶𝑇)−0.398 + 0.8518)

- 90386 (-0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518)2

+ 55033 (-0.5681 × (𝐹𝐶𝑇)−0.398 +0.8518) -11169

+ 258.89 (-0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518)3

- 458.46 (-0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518)2

+ 271.44 (-0.5681 × (𝐹𝐶𝑇)−0.398 + 0.8518) - 53.668) Γw

270

The grout pressure drop was calculated by using the pipeline parameter for a 100

m long slurry pipeline (L) with a nominal diameter of 80 mm (D) at a flowrate


The pressure drop grout model for fly ash ”E” was;

𝛥𝑃 = −1301 × (𝐹𝐶𝑇)2 + 39.225 × (𝐹𝐶𝑇) − 302.86

9.16 Fly Ash “E” Slurries Comparison of 50 mm and 80 Pipeline Viscometers

There was a positive correlation between Newtonian data collected from the 50

mm pipeline viscometer when compared to the 80 mm pipeline viscometer.

There was a positive correlation between non-Newtonian data collected from the

50 mm pipeline viscometer when compared to the 80 mm pipeline viscometer.

9.17 Fly Ashes “B” and “E” Slurries Comparison of Pipeline Pressure Drop

Models

Plotting of a graph of a pipeline pressure drop for a power station slurry pipeline

at different Cw’s for fly ashes “B” and “E” indicated a strong correlation for Cw

up to 75 %.

A pipeline pressure drop model was produced for a power station slurry pipeline


𝑃𝑟𝑒𝑠𝑠𝑢𝑟𝑒 𝐷𝑟𝑜𝑝 ∆𝑃 = 12141 𝐶𝑤3 − 22619 𝐶𝑤

2 + 14075 𝐶𝑊 − 2923.4

It is proposed that this model will be used to determine the pipeline pressure

drop for fly ash that fall with the envelope of PSD as indicated on Figure 8.41

for Cw up to 75 %.

271

The R2 value for both models fly ash “B” = 0.997 and fly ash “E” = 0.999

indicate that a greater proportion of variance is accounted for by the model.

9.18 Fly Ash “E” Determining the Settling Velocity

Observation of fly ash “E” slurries revealed that as the Cw increased the setting

velocity decreased. At a Cw of 60 % the slurry was in suspension at a velocity of

approximately 1.23 m s-1

and at a Cw of approximately 67 % the suspension

velocity was < 0.4 m s-1

.

These results were consistent with the results of the testing by Bunn (1991) of

the settling velocity of Vales Point Power Station fly ash slurry at a Cw of 60 %,

where settling occurred when the velocity was reduced to less than 1.2 m s-1

and

that re-suspension occurred above 1.3 m s-1

.

The change in settling velocity followed a linear relationship with increasing

viscosity.

This pipe flow scenario vindicates the procedure followed in operating power

station dense phase slurry systems and grout pumping plants of reducing the

systems flow to control unexpected increases in pipeline viscosity.

9.19 Fly Ashes “B” and “E” Laminar or Turbulent Flow

Calculating the Reynolds Number indicated that the testing of fly ash “B” and “E”

slurries above a Cw of 61.8 % occurred under laminar flow conditions.

For fly ash “B” and “E” with a re-suspension velocity > 1.3 m s-1

the optimal slurry

pumping Cw would have to be above 63 %.

272

Calculation also concluded that the high concentrate slurry plant at Bayswater

Power Station operates under laminar flow.

9.20 Fly Ash “B” and “E” Homogeneous or Heterogeneous Slurries

Using the Thomas (1967) criteria, the fly ash slurries “B” and “E” are

heterogeneous slurries.

Comparative rheometry between a rotary viscometer and pipeline viscometers

also indicates that in the Cw ranges tested both fly ash “B” and “E” are

heterogeneous slurries.

Using the Thomas (1967) criteria at the higher Cw the fly ash “B” slurry flow

appears to be tending towards homogeneity. However, this is not reflected using

the rotary viscometry pipeline viscometer comparison.

9.21 Redefining Homogeneous Behaviour for Fly Ash Slurries

From the results of this research and the author’s previous research, it can now

be proposed that homogeneous fluid behaviour occurs in fly ash slurries

containing particles with a d50 less than 15 μm and a PSD curve slope 𝑑𝑠 of

3.13.

9.22 Spread Sheet Program

A spread sheet program has been developed that has three output stages.

Stage 1 determines the weight of ash to be pumped per hour when provide with

information regarding the physical configuration of the proposed dense phase

273

ash pumping system, the number of unit, and if the bottom and fly ash unit

system integrated or separate. Also required are the operational details of how

many tons of coal is burnt per hour per unit at the maximum overload rate per

unit and the ash content of the coal.

Stage 2 establishes the pipeline size and volumetric flowrate.

Stages 3 produces a output the relationship between Cw and pipeline pressure.

The calculated Cw corresponding to the pipeline operating pressure of 6 MPa

was 69.7 %.

9.23 Conclusions

This work has made some valuable contributions that give a greater understanding in

the pumping and placement characteristics of high concentration fly ash slurries.

Experimental findings and data from a power station high concentration slurry plant

proposes that pumping of fly ash slurries always occurs in the laminar flow region

within the range of rheology tested and the operational range of all high concentration

ash slurry plants.

Models have been presented for fly ash slurries from two similar sized power stations,

100 km apart, burning coal from different coal seams. The power stations were fitted

with similar boiler and ash collection systems that calculated pipeline pressure drop as a

function of the pumping Cw. From an inspection of the two fly ash models, a singular

model was developed to determine the pipeline pressure drop for fly ash that falls with a

particular PSD envelope for Cw up to 66 %.

From this work and author’s previous research, it is recommended minimum design

velocity for fly ash slurry pipelines be 1.5 m s-1

. This will be more economical in terms

of a decrease in pipeline wear and pumping power.

Other findings of work presented in this thesis:

274

fly ash slurries with a Cw up to 61.8 % behaved as Newtonian fluids, whereas at

a Cw greater than 61.8 % behaved as Non-Newtonian fluids;

comparison of the fly ash “B” slurry volumetric flowrate as measured by the

Foxboro Magnetic Flow Meter and the volumetric flowrate calculated from the

mass flow measured by the weight hopper, Cw and density showed excellent

correlations;

a Bingham visco-plastic model provided positive correlation to the pipeline

viscometers data;

there was a positive correlation between the Bingham plastic models obtained

between the 50 mm pipeline viscometer and the rheological data obtained from

the 80 mm pipeline viscometer. This indicates that the design engineer can

accurately predict slurry pipeline parameters when scaling up rheometry data

obtained in smaller diameter pipeline viscometers to larger diameter slurry

pipeline systems;

full pipe flow was observed in a glass section of pipe for fly ash slurries with at

a velocity of approximately 1.23 m s-1

at a corresponding Cw of 60 % with the

velocity then dipping down to < 0.4 m s-1

when the Cw was increased to

approximately 67 %. These results were consistent with the results of the testing

by Bunn (1991) of the settling velocity of Vales Point Power Station fly ash

slurry at a Cw of 60 %, where settling occurred when the velocity was reduced to

less than 1.2 m s-1

and that re-suspension occurred above 1.3 m s-1

;

changes in settling velocity followed a linear relationship with increasing viscosity.

This pipe flow scenario vindicates the procedure followed in operating power

station high concentration slurry systems and grout pumping plants of reducing the

systems flow to control unexpected increases in pipeline viscosity;

275

comparative rheometry between a rotary viscometer and pipeline viscometers

indicated that in the Cw ranges tested, both power station fly ashes are

heterogeneous slurries;

for the fly ash slurries tested at the generally accepted flow time (20 seconds) for

grout, the pipeline viscosity could be an excellent cost effective alternative to

manual testing. The relationship between flow cone time and viscosity was a

linear relationship at flow cone times greater than 15 seconds or viscosities

above 150 mPa s;

differences in the flow times between a ASTM Flow Cone and a Marsh Funnel

are small but not insignificant;

fly ash grout strength could vary significantly depending on the source of the ash

because of the extra water required to achieve an acceptable flow cone time (20

seconds);

the pumping characteristics of high concentration fly ash slurry pipeline changes

due to shearing in the pipeline and depends on the type of coal the power station

burns and the properties of the process water. However, these changes are

insignificant compared with the changes in pumping characteristics due to

changes in particle size distribution of the fly ash;

a laboratory trial to simulate high concentration fly ash slurry pipeline pumping

characteristics indicated there would be little change in the pipeline pressure drop

per unit length due to thixotrophic behaviour;

a review of low concentration tailing dam failures indicted that there are viable

alternatives deposition sites;

276

the maximum amount of water that is available for recycling from a range of

dense phase fly ash slurries indicated that the percentage of water available for

recycling varies depending on the pumped Cw and the PSD. The amount of

water varied between 25 to 60 % of the water mixed with the fly ash. The

deposited slurry placement density showed an increase when the slurry could be

pumped above a Cw of 65%;

fly ash slurries from different power stations show a great variation in rheology

which can be related to the differences in PSD. However, the variation in

rheology cannot be equated directly to the d50 of the fly ash particles but the

variation in the distribution of the particles across the PSD range: and,

changes in coal supply can adversely affect the operation of a dense phase ash

handling and pumping system.

9.24 Recommendations

There are currently a number of areas that can be addressed as future areas of work

directly related to the subject matter in this thesis. They have been listed below in point

form.

Development of a pilot plant that would allow for rheology studies of the same

slurry mix to be tested in pipeline with small, intermediate and large diameter

pipes simultaneously. This research tested slurries in 50 mm and 80 mm

pipeline but simultaneously testing in a 150 or 200 mm pipe would allow for

more precise analysis.

Research is required to see if a model could be developed based on the physical

properties of ash to determine the pumpability of slurries. Adapting of the void

volume model produced by Donohue and Wensrich (2006) to interface with ash

particle size distribution measured using laser diffraction technique. If these void

277

volumes are filled with water and small quantity of water added to correspond

lubricate the pipe wall is added. Could the Cw determined by this method be

correlated to Cw obtained by rheological measurement? Eliminating the need for

rheological testing.

278

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APPENDIX A

Data

The following appendices contain the raw experimental data from all conveying and

bench-scale experiments in a summarised form.

324

Table A. 1 Fly Ash “B” Averaged Recorded Data Cw 59.7 %

Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h-1

17.5 36.44457 0.50401 1.891301 24.73702 0.841137

20 41.15838 1.187134 3.722807 25.13378 7.568681

22.5 45.20752 1.605883 5.48014 25.21652 12.08708

25 50.29029 1.930975 7.447226 25.47468 14.81664

27.5 52.07372 1.998367 8.239223 25.64339 15.73381

30 55.28636 2.105168 9.699283 25.78348 16.94076

32.5 61.71037 2.393827 12.74105 26.13648 19.3874

35 63.4291 2.421023 13.65922 26.35294 19.92248

37.5 66.17433 2.502942 14.80084 26.62649 20.80788

Table A. 2 Fly Ash “B” 80 mm Pipeline Viscometer Calculated Newtonian Data Cw

59.7 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 0.8411 0.0002 0.0779 0.0047 0.04899 0.10080 100.801 5.0305 1.9363

20 7.5686 0.0021 0.0779 0.0047 0.44089 0.23742 237.426 45.265 4.6250

22.5 12.087 0.0033 0.0779 0.0047 0.70409 0.32118 321.177 72.288 6.2565

25 14.816 0.0041 0.0779 0.0047 0.86309 0.38620 386.195 88.613 7.5230

27.5 15.733 0.0043 0.0779 0.0047 0.91652 0.39967 399.673 94.098 7.7856

30 16.940 0.0047 0.0779 0.0047 0.98683 0.42103 421.033 101.31 8.2017

32.5 19.387 0.0053 0.0779 0.0047 1.12935 0.47876 478.765 115.94 9.3263

35 19.922 0.0055 0.0779 0.0047 1.16052 0.48420 484.204 119.15 9.4323

37.5 20.807 0.0057 0.0779 0.0047 1.21209 0.50059 500.589 124.44 9.7514

325


59.7 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 0.84113 0.00023 0.0525 0.00216 0.10793 0.17826 178.260 16.447 2.33966

20 7.56868 0.00210 0.0525 0.00216 0.97120 0.74456 744.561 147.993 9.77237

22.5 12.0870 0.00335 0.0525 0.00216 1.55099 1.09603 1096.03 236.342 14.3854

25 14.8166 0.00411 0.0525 0.00216 1.90124 1.48944 1489.45 289.714 19.5490

27.5 15.7338 0.00437 0.0525 0.00216 2.01893 1.64784 1647.85 307.648 21.6280

30 16.9407 0.00470 0.0525 0.00216 2.17381 1.93985 1939.86 331.247 25.4606

32.5 19.3873 0.00538 0.0525 0.00216 2.48776 2.54820 2548.21 379.087 33.4453

35 19.9224 0.00553 0.0525 0.00216 2.55642 2.73184 2731.84 389.55 35.8555

37.5 20.8078 0.00578 0.0525 0.00216 2.67003 2.96016 2960.17 406.862 38.8522

Table A. 4 Fly Ash “B” 80 mm Pipeline Viscometer Calculated Non-Newtonian Data

Cw 59.7 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 0.8411 0.0002 0.0779 0.0047 0.04899 0.10080 100.801 20.167 1.9363

20 7.5686 0.0021 0.0779 0.0047 0.44089 0.23742 237.426 181.47 4.6250

22.5 12.087 0.0033 0.0779 0.0047 0.70409 0.32118 321.177 289.81 6.2565

25 14.816 0.0041 0.0779 0.0047 0.86309 0.38620 386.195 355.25 7.5230

27.5 15.733 0.0043 0.0779 0.0047 0.91652 0.39967 399.673 377.24 7.7856

30 16.940 0.0047 0.0779 0.0047 0.98683 0.42103 421.033 406.18 8.2017

32.5 19.387 0.0053 0.0779 0.0047 1.12935 0.47876 478.765 464.84 9.3263

35 19.922 0.0055 0.0779 0.0047 1.16052 0.48420 484.204 477.67 9.4323

37.5 20.807 0.0057 0.0779 0.0047 1.21209 0.50059 500.589 498.9 9.7514

326


Cw 59.7 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 0.84113 0.00023 0.0525 0.00216 0.10793 0.17826 178.260 48.9279 2.33966

20 7.56868 0.00210 0.0525 0.00216 0.97120 0.74456 744.561 440.261 9.77237

22.5 12.0870 0.00335 0.0525 0.00216 1.55099 1.09603 1096.03 703.092 14.3854

25 14.8166 0.00411 0.0525 0.00216 1.90124 1.48944 1489.45 861.867 19.5490

27.5 15.7338 0.00437 0.0525 0.00216 2.01893 1.64784 1647.85 915.218 21.6280

30 16.9407 0.00470 0.0525 0.00216 2.17381 1.93985 1939.86 985.424 25.4606

32.5 19.3873 0.00538 0.0525 0.00216 2.48776 2.54820 2548.21 1127.74 33.4453

35 19.9224 0.00553 0.0525 0.00216 2.55642 2.73184 2731.84 1158.88 35.8555

37.5 20.8078 0.00578 0.0525 0.00216 2.67003 2.96016 2960.17 1210.37 38.8522

Table A. 6 Fly Ash “B” Rotary Viscometer Result Sheet Cw 59.7 %

Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 2 1.008

2.65 0.504 2.2 1.1088

3.6 0.504 2.6 1.3104

4.89 0.504 3 1.512

6.64 0.504 3.5 1.764

9.03 0.504 3.8 1.9152

12.3 0.504 4 2.016

16.7 0.504 4.5 2.268

22.7 0.504 5 2.52

30.8 0.504 6.1 3.0744

41.9 0.504 6.8 3.4272

57 0.504 7.8 3.9312

77.5 0.504 9.2 4.6368

105 0.504 11 5.544

143 0.504 13 6.552

195 0.504 17 8.568

327


Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h-1

17.5 36.32569 1.997223 1.710416 27.30398 1.460594

20 41.70581 2.377115 3.684906 27.38223 4.883154

22.5 47.27416 2.817194 5.666474 27.43925 8.62242

25 52.07418 3.323192 7.258343 27.53221 12.38487

27.5 54.56541 3.430663 8.171378 27.75855 14.64203

30 59.04873 3.817789 9.78536 27.96452 16.97145

32.5 62.47432 4.031289 11.1658 28.0768 18.24263

35 64.75276 4.132852 12.16659 28.196 19.14488

37.5 73.05824 4.464846 16.12134 28.34936 21.39434


61.8 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 1.46159 0.00041 0.0779 0.00477 0.04899 0.39946 399.446 8.73537 7.78118

20 4.88315 0.00136 0.0779 0.00477 0.44089 0.47542 475.423 29.2046 9.26124

22.5 8.62242 0.00240 0.0779 0.00477 0.70409 0.56349 563.439 51.5680 10.9759

25 12.3849 0.00344 0.0779 0.00477 0.86309 0.66468 664.638 74.0700 12.9476

27.5 14.6420 0.00407 0.0779 0.00477 0.91652 0.68613 686.137 87.5693 13.3659

30 16.9715 0.00471 0.0779 0.00477 0.98683 0.76358 763.558 101.509 14.8741

32.5 18.2426 0.00507 0.0779 0.00477 1.12935 0.80628 806.258 109.103 15.7059

35 19.1449 0.00538 0.0779 0.00477 1.16052 0.82657 826.570 114.499 16.1016

37.5 21.3943 0.00594 0.0779 0.00477 1.21209 0.89297 892.969 127.958 17.3950

328


61.8 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 1.46059 0.00041 0.0525 0.00216 0.18742 0.71802 718.028 28.5594 9.42404

20 4.88315 0.00136 0.0525 0.00216 0.62660 1.12421 1124.21 95.4817 14.7552

22.5 8.62242 0.00239 0.0525 0.00216 1.10642 1.58412 1584.15 168.597 20.7919

25 12.3849 0.00344 0.0525 0.00216 1.58921 1.99155 1991.55 242.165 26.1391

27.5 14.6420 0.00407 0.0525 0.00216 1.87884 2.20442 2204.42 286.300 28.9331

30 16.9715 0.00471 0.0525 0.00216 2.17775 2.83263 2832.63 331.848 37.1783

32.5 18.2426 0.00507 0.0525 0.00216 2.34087 3.25953 3259.53 356.703 42.7813

35 19.1449 0.00532 0.0525 0.00216 2.45664 3.56244 3562.44 374.345 46.7570

37.5 21.3943 0.00594 0.0525 0.00216 2.74529 4.68452 4684.52 418.330 61.4843


Cw 61.8 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 1.46159 0.00041 0.0779 0.0047 0.04899 0.39946 399.446 24.5275 7.78118

20 4.88315 0.00136 0.0779 0.0047 0.44089 0.47542 475.423 82.0020 9.26124

22.5 8.62242 0.00240 0.0779 0.0047 0.70409 0.56349 563.439 144.795 10.9759

25 12.3849 0.00344 0.0779 0.0047 0.86309 0.66468 664.638 207.977 12.9476

27.5 14.6420 0.00407 0.0779 0.0047 0.91652 0.68613 686.137 245.881 13.3659

30 16.9715 0.00471 0.0779 0.0047 0.98683 0.76358 763.558 284.999 14.8741

32.5 18.2426 0.00507 0.0779 0.0047 1.12935 0.80628 806.258 306.345 15.7059

35 19.1449 0.00538 0.0779 0.0047 1.16052 0.82657 826.570 321.497 16.1016

37.5 21.3943 0.00594 0.0779 0.0047 1.21209 0.89297 892.969 359.272 17.3950

329


Cw 61.8 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 1.46059 0.00041 0.0525 0.00216 0.18742 0.71802 718.028 77.8676 9.42404

20 4.88315 0.00136 0.0525 0.00216 0.62660 1.12421 1124.21 260.332 14.7552

22.5 8.62242 0.00239 0.0525 0.00216 1.10642 1.58412 1584.15 459.681 20.7919

25 12.3849 0.00344 0.0525 0.00216 1.58921 1.99155 1991.55 660.266 26.1391

27.5 14.6420 0.00407 0.0525 0.00216 1.87884 2.20442 2204.42 780.600 28.9331

30 16.9715 0.00471 0.0525 0.00216 2.17775 2.83263 2832.63 904.787 37.1783

32.5 18.2426 0.00507 0.0525 0.00216 2.34087 3.25953 3259.53 972.556 42.7813

35 19.1449 0.00532 0.0525 0.00216 2.45664 3.56244 3562.44 1020.66 46.7570

37.5 21.3943 0.00594 0.0525 0.00216 2.74529 4.68452 4684.52 1140.58 61.4843


Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 3 1.512

2.65 0.504 3.2 1.6128

3.6 0.504 3.8 1.9152

4.89 0.504 4.1 2.0664

6.64 0.504 4.8 2.4192

9.03 0.504 5.1 2.5704

12.3 0.504 5.9 2.9736

16.7 0.504 6.7 3.3768

22.7 0.504 7.8 3.9312

30.8 0.504 8.8 4.4352

41.9 0.504 10 5.04

57 0.504 11.5 5.796

77.5 0.504 13 6.552

105 0.504 16 8.064

143 0.504 19.5 9.828

195 0.504 25 12.6

330


Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h-1

17.5 38.13036 2.911012 2.055298 30.19147 0.392856

20 43.6474 3.306015 4.164969 30.16152 2.85198

22.5 50.60625 3.904303 6.773737 30.19457 5.173308

25 57.49134 4.676493 9.275187 30.24672 8.691905

27.5 61.92857 5.019408 10.67066 30.50665 10.88703

30 66.03733 5.481571 12.16163 30.60996 12.89793

32.5 72.3758 6.016003 14.64945 30.68253 15.72699

35 78.57294 6.486526 17.0286 30.737 17.94871

37.5 82.60108 6.880186 18.68098 30.83215 19.92789


65.1 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 0.39286 0.00011 0.0779 0.0047 0.02289 0.58220 582.202 2.34955 11.3413

20 2.85198 0.00079 0.0779 0.0047 0.16613 0.66120 661.203 17.0568 12.8802

22.5 5.17331 0.00144 0.0779 0.0047 0.30136 0.78086 780.861 30.9399 15.2112

25 8.69190 0.00241 0.0779 0.0047 0.50632 0.93530 935.299 51.9835 18.2196

27.5 10.8870 0.00302 0.0779 0.0047 0.63419 1.00388 1003.88 65.1119 19.5556

30 12.8979 0.00358 0.0779 0.0047 0.75133 1.09631 1096.31 77.1384 21.3562

32.5 15.7270 0.00437 0.0779 0.0047 0.91613 1.20320 1203.20 94.0581 23.4384

35 17.9487 0.00450 0.0779 0.0047 1.04555 1.29731 1297.31 107.346 25.2715

37.5 19.9279 0.00554 0.0779 0.0047 1.16084 1.37604 1376.04 119.182 26.8052

331


65.1 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 0.39282 0.00011 0.0525 0.00216 0.05041 0.93949 939.494

7.68163 12.3309

5

20 2.85198 0.00079 0.0525 0.00216 0.36596 1.40275 1402.75

55.7656 18.4111

22.5 5.17331 0.00144 0.0525 0.00216 0.66383 1.85909 1859.09

101.155 24.4005

25 8.69190 0.00241 0.0525 0.00216 1.11533 2.52045 2520.45

169.955 33.0809

27.5 10.8870 0.00302 0.0525 0.00216 1.39701 3.00057 3000.57

212.877 39.3825

30 12.8979 0.00358 0.0525 0.00216 1.65504 3.36069 3360.69

252.197 44.1091

32.5 15.7270 0.00439 0.0525 0.00216 2.01806 3.88622 3886.22

307.514 51.0067

35 17.9487 0.00498

6 0.0525 0.00216 2.30315

4.39582 4395.82 350.956 57.6951

37.5 19.9279 0.00554 0.0525 0.00216 2.55712 4.91199 4911.99

389.656 64.4699


Cw 65.1%

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 0.39286 0.00011 0.0779 0.0047 0.02289 0.58220 582.202 5.34264 11.3413

20 2.85198 0.00079 0.0779 0.0047 0.16613 0.66120 661.203 38.7855 12.8802

22.5 5.17331 0.00144 0.0779 0.0047 0.30136 0.78086 780.861 70.3543 15.2112

25 8.69190 0.00241 0.0779 0.0047 0.50632 0.93530 935.299 118.205 18.2196

27.5 10.8870 0.00302 0.0779 0.0047 0.63419 1.00388 1003.88 148.058 19.5556

30 12.8979 0.00358 0.0779 0.0047 0.75133 1.09631 1096.31 175.405 21.3562

32.5 15.7270 0.00437 0.0779 0.0047 0.91613 1.20320 1203.20 213.879 23.4384

35 17.9487 0.00450 0.0779 0.0047 1.04555 1.29731 1297.31 244.093 25.2715

37.5 19.9279 0.00554 0.0779 0.0047 1.16084 1.37604 1376.04 271.009 26.8052

332


Cw 65.1 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 0.39282 0.00011 0.0525 0.00216 0.05041 0.93949 939.494 17.4091 12.3309

5

20 2.85198 0.00079 0.0525 0.00216 0.36596 1.40275 1402.75 126.383 18.4111

22.5 5.17331 0.00144 0.0525 0.00216 0.66383 1.85909 1859.09 229.251 24.4005

25 8.69190 0.00241 0.0525 0.00216 1.11533 2.52045 2520.45 385.175 33.0809

27.5 10.8870 0.00302 0.0525 0.00216 1.39701 3.00057 3000.57 482.451 39.3825

30 12.8979 0.00358 0.0525 0.00216 1.65504 3.36069 3360.69 571.562 44.1091

32.5 15.7270 0.00439 0.0525 0.00216 2.01806 3.88622 3886.22 696.930 51.0067

35 17.9487 0.00499 0.0525 0.00216 2.30315 4.39582 4395.82 795.384 57.6951

37.5 19.9279 0.00554 0.0525 0.00216 2.55712 4.91199 4911.99 883.090 64.4699


Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 3.8 1.9152

2.65 0.504 5 2.52

3.6 0.504 5.5 2.772

4.89 0.504 6.5 3.276

6.64 0.504 7.2 3.6288

9.03 0.504 8 4.032

12.3 0.504 9.2 4.6368

16.7 0.504 10.6 5.3424

22.7 0.504 12.2 6.1488

30.8 0.504 14.5 7.308

41.9 0.504 17 8.568

57 0.504 19.5 9.828

77.5 0.504 23 11.592

105 0.504 27 13.608

143 0.504 33.5 16.884

195 0.504 43 21.672

333


Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h-1

20 45.37864 4.584231 4.661189 34.84758 1.034985

25 59.582 5.427427 10.39384 34.82204 3.522628

30 75.05502 7.068751 16.49036 34.90192 7.298148

35 91.18428 8.427503 23.65488 34.9042 10.95192

40 107.2807 10.12091 30.30439 35.02356 15.11443

45 125.7329 11.93073 38.05548 35.17529 19.5352

50 134.1062 12.73276 41.63162 35.46462 21.56849


67.9 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 1.03505 0.00029 0.0779 0.0047 0.06029 0.91685 916.846 6.18992 17.8602

25 3.52263 0.00098 0.0779 0.0047 0.2052 1.08549 1085.49 21.0677 21.1453

30 7.29815 0.00203 0.0779 0.0047 0.42513 1.41375 1413.75 43.6479 27.5398

35 10.9519 0.00304 0.0779 0.0047 0.63797 1.68550 1685.50 65.5000

32.8336

40 15.1144 0.00420 0.0779 0.0047 0.88044 2.02418 2024.18 90.3946 39.4311

45 19.5352 0.00543 0.0779 0.0047 1.13796 2.38615 2386.15 116.834 46.4821

50 21.5685 0.00599 0.0779 0.0047 1.25641 2.54655 2546.55 128.994 49.6068

334


67.9 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 1.03499 0.00098 0.0525 0.00216 0.13281 1.57402 1574.02 20.2374 20.6590

25 3.52263 0.00203 0.0525 0.00216 0.45202 2.57151 2571.51 68.8789 33.7511

30 7.29815 0.00304 0.0525 0.00216 0.93649 4.02000 4020.00 142.703 52.7624

35 10.9519 0.00420 0.0525 0.00216 1.40533 5.41275 5412.75 214.146 71.0423

40 15.1144 0.00543 0.0525 0.00216 1.93946 6.81505 6815.06 295.537 89.4475

45 19.5352 0.00599 0.0525 0.00216 2.50673 8.65896 8658.96 381.977 113.649

50 21.5685 0.00098 0.0525 0.00216 2.76764 9.61175 9611.75 421.735 126.154


Cw 67.9 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 1.03505 0.00029 0.0779 0.0047 0.06029 0.91685 916.846 10.7788 17.8602

25 3.52263 0.00098 0.0779 0.0047 0.2052 1.08549 1085.49 36.6862 21.1453

30 7.29815 0.00203 0.0779 0.0047 0.42513 1.41375 1413.75 76.0062 27.5398

35 10.9519 0.00304 0.0779 0.0047 0.63797 1.68550 1685.50 114.058 32.8336

40 15.1144 0.00420 0.0779 0.0047 0.88044 2.02418 2024.18 157.408 39.4311

45 19.5352 0.00543 0.0779 0.0047 1.13796 2.38615 2386.15 203.448 46.4821

50 21.5685 0.00599 0.0779 0.0047 1.25641 2.54655 2546.55 224.624 49.6068

335


Cw 67.9 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 1.03450 0.00098 0.0525 0.00216 0.13281 1.57402 1574.02 35.2695 20.6590

25 3.52263 0.00203 0.0525 0.00216 0.45202 2.57151 2571.51 120.042 33.7511

30 7.29815 0.00304 0.0525 0.00216 0.93649 4.02000 4020.00 248.702 52.7624

35 10.9519 0.00420 0.0525 0.00216 1.40533 5.41275 5412.75 373.213 71.0423

40 15.1144 0.00543 0.0525 0.00216 1.93946 6.81505 6815.06 515.060 89.4475

45 19.5352 0.00599 0.0525 0.00216 2.50673 8.65896 8658.96 665.708 113.649

50 21.5685 0.00098 0.0525 0.00216 2.76764 9.61175 9611.75 734.997 126.154


Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 6 3.024

2.65 0.504 7.5 3.78

3.6 0.504 8.9 4.4856

4.89 0.504 10.1 5.0904

6.64 0.504 12 6.048

9.03 0.504 14.5 7.308

12.3 0.504 17.5 8.82

16.7 0.504 21 10.584

22.7 0.504 24 12.096

30.8 0.504 28 14.112

41.9 0.504 32.5 16.38

57 0.504 38 19.152

77.5 0.504 46 23.184

105 0.504 55 27.72

143 0.504 66 33.264

195 0.504 80 40.32

336

Table A. 25 Fly Ash “E” Averaged Recorded Data Cw 58.1 %

Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h-1

17.5 48.53604 0.622503 1.579577 22.2528 2.24401

20 53.54777 1.104381 3.427829 22.28225 6.730524

25 65.39604 1.837936 7.553345 22.30273 15.59955

30 76.48767 2.142943 12.76981 22.34512 19.58118

35 84.87483 2.41647 16.41033 22.44069 22.34837

40 92.01968 2.538793 19.43629 22.54775 24.48234

45 98.37462 2.662875 22.13866 22.69262 26.34433

Table A. 26 Fly Ash “E” 80 mm Pipeline Viscometer Calculated Newtonian Data Cw

58.1 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 2.24401 0.00062 0.0779 0.0047 0.13073 0.12450 124.501 13.4207 2.42527

20 6.73052 0.00187 0.0779 0.0047 0.39207 0.22088 220.876 40.2531 4.30267

25 15.5996 0.00433 0.0779 0.0047 0.90870 0.36759 367.587 93.2960 7.16060

30 19.5812 0.00544 0.0779 0.0047 1.14064 0.42859 428.589 117.109 8.34891

35 22.3484 0.00621 0.0779 0.0047 1.30183 0.48329 483.294 133.659 9.41457

40 24.4823 0.00680 0.0779 0.0047 1.42614 0.50776 507.759 146.421 9.89114

45 26.3443 0.00732 0.0779 0.0047 1.53461 0.53258 532.575 157.557 10.3746

337


58.1 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 2.24401 0.00062 0.0525 0.00216 0.28795 0.31592 315.915 43.8778 4.14639

20 6.73052 0.00187 0.0525 0.00216 0.86365 0.68557 685.566 131.604 8.99805

25 15.5996 0.00433 0.0525 0.00216 2.00171 1.51070 1510.67 305.022 19.8275

30 19.5812 0.00544 0.0525 0.00216 2.51263 2.55396 2553.96 382.876 33.5207

35 22.3484 0.00621 0.0525 0.00216 2.86771 3.28207 3282.07 436.984 43.0771

40 24.4823 0.00681 0.0525 0.00216 3.14154 3.88726 3887.26 478.710 51.0203

45 26.3443 0.00732 0.0525 0.00216 3.38046 4.42773 4427.73 515.118 58.1140

Table A. 28 Fly Ash “E” 80 mm Pipeline Viscometer Calculated Non-Newtonian Data

Cw 58.1 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 2.24401 0.00062 0.0779 0.0047 0.13073 0.12450 124.501 13.4207 2.42527

20 6.73052 0.00187 0.0779 0.0047 0.39207 0.22088 220.876 40.2531 4.30267

25 15.5996 0.00433 0.0779 0.0047 0.90870 0.36759 367.587 93.2960 7.16060

30 19.5812 0.00544 0.0779 0.0047 1.14064 0.42859 428.589 117.109 8.34891

35 22.3484 0.00621 0.0779 0.0047 1.30183 0.48329 483.294 133.659 9.41457

40 24.4823 0.00680 0.0779 0.0047 1.42614 0.50776 507.759 146.421 9.89114

45 26.3443 0.00732 0.0779 0.0047 1.53461 0.53258 532.575 157.557 10.3746

338


Cw 58.1 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 2.24401 0.00062 0.0525 0.00216 0.28795 0.31592 315.915 43.8778 4.14639

20 6.73052 0.00187 0.0525 0.00216 0.86365 0.68557 685.566 131.604 8.99805

25 15.5996 0.00433 0.0525 0.00216 2.00171 1.51070 1510.67 305.022 19.8275

30 19.5812 0.00544 0.0525 0.00216 2.51263 2.55396 2553.96 382.876 33.5207

35 22.3484 0.00621 0.0525 0.00216 2.86771 3.28207 3282.07 436.984 43.0771

40 24.4823 0.00681 0.0525 0.00216 3.14154 3.88726 3887.26 478.710 51.0203

45 26.3443 0.00732 0.0525 0.00216 3.38046 4.42773 4427.73 515.118 58.1140

Table A. 30 Fly Ash “E” Rotary Viscometer Result Sheet Cw 58.1 %

Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 1.8 0.9072

2.65 0.504 2 1.008

3.6 0.504 2.2 1.1088

4.89 0.504 2.7 1.3608

6.64 0.504 3 1.512

9.03 0.504 3.2 1.6128

12.3 0.504 3.5 1.764

16.7 0.504 4.1 2.0664

22.7 0.504 4.5 2.268

30.8 0.504 5.2 2.6208

41.9 0.504 5.8 2.9232

57 0.504 7 3.528

77.5 0.504 8.2 4.1328

105 0.504 9.5 4.788

143 0.504 11.5 5.796

195 0.504 14.5 7.308

339


Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h

-1

17.5 48.53939 0.428102 1.393883 23.88648 1.753623

20 53.38095 0.917251 3.129861 23.93897 6.402774

25 65.06508 1.784409 7.200578 23.97468 15.06585

30 77.98649 2.095168 13.07111 24.01179 19.75144

35 86.47936 2.366901 16.7816 24.07324 22.43264

40 92.89239 2.555614 19.45739 24.18926 24.28735

45 99.72059 2.757024 22.48312 24.31149 26.15987


59.2 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 1.75362 0.00049 0.0779 0.0047 0.10215 0.08562 85.6204 10.4879 1.66788

20 6.40277 0.00178 0.0779 0.0047 0.37297 0.18345 183.450 38.2930 3.57361

25 15.0659 0.00419 0.0779 0.0047 0.87761 0.35688 356.882 90.1041 6.95206

30 19.7514 0.00549 0.0779 0.0047 1.15056 0.41903 419.034 118.127 8.16278

35 22.4326 0.00623 0.0779 0.0047 1.30674 0.47338 473.383 134.163 9.22145

40 24.2874 0.00675 0.0779 0.0047 1.41478 0.51112 511.123 145.255 9.95667

45 26.1599 0.00727 0.0779 0.0047 1.52386 0.55141 551.405 156.454 10.7414

340


59.2 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 1.75362 0.00049 0.0525 0.00216 0.22502 0.27878 278.777 34.2891 3.65894

20 6.40277 0.00178 0.0525 0.00216 0.82159 0.62597 625.972 125.195 8.21589

25 15.0659 0.00419 0.0525 0.00216 1.93323 1.44012 1440.12 294.587 18.9015

30 19.7514 0.00549 0.0525 0.00216 2.53447 2.61422 2614.22 386.206 34.3117

35 22.4326 0.00623 0.0525 0.00216 2.87852 3.35632 3356.32 438.632 44.0517

40 24.2874 0.00675 0.0525 0.00216 3.11652 3.89148 3891.48 474.898 51.0757

45 26.1599 0.00727 0.0525 0.00216 3.35679 4.49662 4496.62 511.511 59.0182


Cw 59.2 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 1.75362 0.00049 0.0779 0.0047 0.10215 0.08562 85.6204 43.7586 1.66788

20 6.40277 0.00178 0.0779 0.0047 0.37297 0.18345 183.450 159.770 3.57361

25 15.0659 0.00419 0.0779 0.0047 0.87761 0.35688 356.882 375.942 6.95206

30 19.7514 0.00549 0.0779 0.0047 1.15056 0.41903 419.034 492.862 8.16278

35 22.4326 0.00623 0.0779 0.0047 1.30674 0.47338 473.383 559.767 9.22145

40 24.2874 0.00675 0.0779 0.0047 1.41478 0.51112 511.123 606.048 9.95667

45 26.1599 0.00727 0.0779 0.0047 1.52386 0.55141 551.405 652.773 10.7414

341


Cw 59.2 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

17.5 1.75362 0.00049 0.0525 0.00216 0.22502 0.27878 278.777 129.497 3.65894

20 6.40277 0.00178 0.0525 0.00216 0.82159 0.62597 625.972 472.817 8.21589

25 15.0659 0.00419 0.0525 0.00216 1.93323 1.44012 1440.12 1112.55 18.9015

30 19.7514 0.00549 0.0525 0.00216 2.53447 2.61422 2614.22 1458.56 34.3117

35 22.4326 0.00623 0.0525 0.00216 2.87852 3.35632 3356.32 1656.55 44.0517

40 24.2874 0.00675 0.0525 0.00216 3.11652 3.89148 3891.48 1793.51 51.0757

45 26.1599 0.00727 0.0525 0.00216 3.35679 4.49662 4496.62 1931.79 59.0182


Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 2 1.008

2.65 0.504 2.2 1.1088

3.6 0.504 2.5 1.26

4.89 0.504 2.8 1.4112

6.64 0.504 3 1.512

9.03 0.504 3.5 1.764

12.3 0.504 3.9 1.9656

16.7 0.504 4.2 2.1168

22.7 0.504 5 2.52

30.8 0.504 5.8 2.9232

41.9 0.504 6.5 3.276

57 0.504 7.5 3.78

77.5 0.504 8.5 4.284

105 0.504 10 5.04

143 0.504 12.5 6.3

195 0.504 16 8.064

342


Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h

-1

18.5 51.785 0.719358 2.638151 26.48765 1.523656

20 55.14509 1.292735 4.222695 26.50161 3.649295

25 67.34732 2.738866 9.653296 26.52994 10.71418

30 79.41974 3.746028 14.76745 26.56271 17.1741

35 90.05099 4.437847 19.68392 26.6228 21.23604

40 99.3498 4.781852 24.26544 26.71775 23.85365

45 109.1784 5.159924 28.8683 26.85675 26.39956


62.3 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 1.52366 0.00042 0.0779 0.0047 0.08876 0.14387 143.872 9.11250 2.80262

20 3.64930 0.00101 0.0779 0.0047 0.21258 0.25855 258.547 21.8253 5.03650

25 10.7142 0.00298 0.0779 0.0047 0.62412 0.54777 547.773 64.0781 10.6706

30 17.1741 0.00477 0.0779 0.0047 1.00042 0.74921 749.206 102.713 14.5945

35 21.2360 0.00590 0.0779 0.0047 1.23704 0.88757 887.570 127.006 17.2899

40 23.8537 0.00663 0.0779 0.0047 1.38952 0.95637 956.370 142.661 18.6301

45 26.3996 0.00733 0.0779 0.0047 1.53782 1.03199 1031.99 157.887 20.1031

343


62.3 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 1.52366 0.00042 0.0525 0.00216 0.19553 0.52763 527.630 29.7925 6.92515

20 3.64930 0.00101 0.0525 0.00216 0.46827 0.84454 844.539 71.3557 11.0846

25 10.7142 0.00298 0.0525 0.00216 1.37483 1.93066 1930.66 209.497 25.3399

30 17.1741 0.00477 0.0525 0.00216 2.20375 2.95349 2953.49 335.81 38.7646

35 21.2360 0.00590 0.0525 0.00216 2.72498 3.93678 3936.78 415.234 51.6703

40 23.8537 0.00663 0.0525 0.00216 3.06086 4.85309 4853.09 466.417 63.6968

45 26.3996 0.00733 0.0525 0.00216 3.38755 5.77366 5773.66 516.198 75.7793


Cw 62.3 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 1.52366 0.00042 0.0779 0.0047 0.08876 0.14387 143.872 22.9832 2.80262

20 3.64930 0.00101 0.0779 0.0047 0.21258 0.25855 258.547 55.0468 5.03650

25 10.7142 0.00298 0.0779 0.0047 0.62412 0.54777 547.773 161.615 10.6706

30 17.1741 0.00477 0.0779 0.0047 1.00042 0.74921 749.206 259.058 14.5945

35 21.2360 0.00590 0.0779 0.0047 1.23704 0.88757 887.570 320.329 17.2899

40 23.8537 0.00663 0.0779 0.0047 1.38952 0.95637 956.370 359.814 18.6301

45 26.3996 0.00733 0.0779 0.0047 1.53782 1.03199 1031.99 398.217 20.1031

344


Cw 62.3 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 1.52366 0.00042 0.0525 0.00216 0.19553 0.52763 527.630 72.2500 6.92515

20 3.64930 0.00101 0.0525 0.00216 0.46827 0.84454 844.539 173.045 11.0846

25 10.7142 0.00298 0.0525 0.00216 1.37483 1.93066 1930.66 508.054 25.3399

30 17.1741 0.00477 0.0525 0.00216 2.20375 2.95349 2953.49 814.375 38.7646

35 21.2360 0.00590 0.0525 0.00216 2.72498 3.93678 3936.78 1006.99 51.6703

40 23.8537 0.00663 0.0525 0.00216 3.06086 4.85309 4853.09 1131.11 63.6968

45 26.3996 0.00733 0.0525 0.00216 3.38755 5.77366 5773.66 1251.84 75.7793


Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 3 1.512

2.65 0.504 3.8 1.9152

3.6 0.504 4.2 2.1168

4.89 0.504 4.8 2.4192

6.64 0.504 5.1 2.5704

9.03 0.504 5.9 2.9736

12.3 0.504 6.5 3.276

16.7 0.504 7.2 3.6288

22.7 0.504 8.2 4.1328

30.8 0.504 9.5 4.788

41.9 0.504 11 5.544

57 0.504 13.8 6.9552

77.5 0.504 15 7.56

105 0.504 17.8 8.9712

143 0.504 22 11.088

195 0.504 27 13.608

345


Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h

-1

18.5 52.23401 1.463693 2.872873 27.61681 0.873736

20 55.49459 1.817273 4.718671 27.6226 2.404939

25 68.44046 3.274891 10.93124 27.62689 8.507337

30 80.0368 4.436524 16.75929 27.70532 14.1961

35 93.53986 5.644618 23.32147 27.75524 19.45113

40 106.9215 6.44629 29.8073 27.8285 23.7547

45 113.3347 6.833905 33.26992 27.93252 25.93533

50 120.8144 7.241576 37.43986 28.1906 27.87516


63.8 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 0.87374 0.00024 0.0779 0.0047 0.05090 0.29274 292.739 5.22554 2.48147

20 2.40494 0.00067 0.0779 0.0047 0.14009 0.36346 363.455 14.3832 5.13210

25 8.50734 0.00236 0.0779 0.0047 0.49557 0.65498 654.978 50.8797 12.7590

30 14.1961 0.00394 0.0779 0.0047 0.82695 0.88731 887.305 84.9024 17.2847

35 19.4511 0.00540 0.0779 0.0047 1.13306 1.12892 1128.92 116.331 21.9914

40 23.7547 0.00660 0.0779 0.0047 1.38376 1.28926 1289.26 142.069 25.1148

45 25.9353 0.00720 0.0779 0.0047 1.51078 1.36678 1366.78 155.111 25.8457

50 27.8751 0.00774 0.0779 0.0047 1.62378 1.44832 1448.32 166.713 28.2132

346


63.8 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 0.87374 0.00024 0.0525 0.00216 0.11212 0.57458 574.575 17.0844 7.03854

20 2.40494 0.00067 0.0525 0.00216 0.30860 0.94373 943.734 47.0245 11.5608

25 8.50734 0.00236 0.0525 0.00216 1.09165 2.18625 2186.25 166.346 26.7815

30 14.1961 0.00394 0.0525 0.00216 1.82162 3.35186 3351.86 277.580 41.0603

35 19.4511 0.00540 0.0525 0.00216 2.49594 4.66430 4664.30 380.333 57.1376

40 23.7547 0.00660 0.0525 0.00216 3.04817 5.96146 5961.46 464.482 73.0279

45 25.9353 0.00720 0.0525 0.00216 3.32798 6.65398 6653.98 507.121 81.5113

50 27.8752 0.00774 0.0525 0.00216 3.57690 7.48797 7487.97 545.051 91.727


Cw 63.8 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 0.87374 0.00024 0.0779 0.0047 0.05090 0.29274 292.739 11.5596 2.48147

20 2.40494 0.00067 0.0779 0.0047 0.14009 0.36346 363.455 31.8175 5.13210

25 8.50734 0.00236 0.0779 0.0047 0.49557 0.65498 654.978 112.553 12.7590

30 14.1961 0.00394 0.0779 0.0047 0.82695 0.88731 887.305 187.815 17.2847

35 19.4511 0.00540 0.0779 0.0047 1.13306 1.12892 1128.92 257.339 21.9914

40 23.7547 0.00660 0.0779 0.0047 1.38376 1.28926 1289.26 314.276 25.1148

45 25.9353 0.00720 0.0779 0.0047 1.51078 1.36678 1366.78 343.126 25.8457

50 27.8751 0.00774 0.0779 0.0047 1.62378 1.44832 1448.32 368.790 28.2132

347


Cw 63.8 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 0.87374 0.00024 0.0525 0.00216 0.11212 0.57458 574.575 37.6214 7.03854

20 2.40494 0.00067 0.0525 0.00216 0.30860 0.94373 943.734 103.552 11.5608

25 8.50734 0.00236 0.0525 0.00216 1.09165 2.18625 2186.25 366.309 26.7815

30 14.1961 0.00394 0.0525 0.00216 1.82162 3.35186 3351.86 611.256 41.0603

35 19.4511 0.00540 0.0525 0.00216 2.49594 4.66430 4664.30 837.527 57.1376

40 23.7547 0.00660 0.0525 0.00216 3.04817 5.96146 5961.46 1022.83 73.0279

45 25.9353 0.00720 0.0525 0.00216 3.32798 6.65398 6653.98 1116.72 81.5113

50 27.8752 0.00774 0.0525 0.00216 3.57690 7.48797 7487.97 1200.25 91.727


Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 3.9 1.9656

2.65 0.504 4.9 2.4696

3.6 0.504 5.2 2.6208

4.89 0.504 6.3 3.1752

6.64 0.504 7 3.528

9.03 0.504 8 4.032

12.3 0.504 8.9 4.4856

16.7 0.504 10 5.04

22.7 0.504 11.2 5.6448

30.8 0.504 13 6.552

41.9 0.504 15 7.56

57 0.504 17.5 8.82

77.5 0.504 21 10.584

105 0.504 24.5 12.348

143 0.504 29.5 14.868

195 0.504 37 18.648

348


Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h

-1

18.5 53.12272 1.099421 3.297754 28.80239 0.418351

20 56.40981 2.389538 4.813396 28.83354 1.517837

25 70.08288 3.732966 11.67809 28.86464 6.318416

30 82.87721 5.257038 18.28901 28.91708 11.52752

35 95.1445 6.533381 24.47972 28.9191 15.81961

40 112.8821 7.954038 33.42978 29.01716 21.64352

45 124.6641 8.853004 39.17699 29.14556 24.97334

50 130.073 9.289238 42.00767 29.37977 26.57054


65.4 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 0.41835 0.00012 0.0779 0.0047 0.02437 0.21988 219.884 2.50203 3.87343

20 1.51784 0.00042 0.0779 0.0047 0.08842 0.47791 477.908 9.07770 5.41364

25 6.31842 0.00176 0.0779 0.0047 0.36806 0.74659 746.593 37.7884 14.5436

30 11.5275 0.00320 0.0779 0.0047 0.67150 1.05141 1051.41 68.9424 20.4814

35 15.8196 0.00439 0.0779 0.0047 0.92152 1.30668 1306.68 94.6121 25.4541

40 21.6435 0.00601 0.0779 0.0047 1.26078 1.59081 1590.81 129.443 30.1565

45 24.9733 0.00694 0.0779 0.0047 1.45474 1.77060 1770.60 149.358 31.2850

50 26.5705 0.00738 0.0779 0.0047 1.54778 1.85785 1857.85 158.91 30.8713

349


65.4 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 0.41835 0.00012 0.0525 0.00216 0.05368 0.65955 659.551

8 8.18014 7.04997

20 1.51784 0.00042 0.0525 0.00216 0.19477 0.96268 962.679 29.6787 11.7928

25 6.31842 0.00176 0.0525 0.00216 0.81077 2.33562 2335.62 123.546 28.6113

30 11.5275 0.00320 0.0525 0.00216 1.47919 3.65780 3657.80 225.401 44.8081

35 15.8196 0.00439 0.0525 0.00216 2.02995 4.89595 4895.95 309.325 59.9753

40 21.6435 0.00601 0.0525 0.00216 2.77726 6.68596 6685.96 423.202 81.9030

45 24.9733 0.00694 0.0525 0.00216 3.20454 7.83540 7835.40 488.311 95.9836

50 26.5705 0.00738 0.0525 0.00216 3.40949 8.40154 8401.54 519.541 102.919


Cw 65.4 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 0.41835 0.00012 0.0779 0.0047 0.02437 0.21988 219.884 4.70808 3.87343

20 1.51784 0.00042 0.0779 0.0047 0.08842 0.47791 477.908 17.0816 5.41364

25 6.31842 0.00176 0.0779 0.0047 0.36806 0.74659 746.593 71.1068 14.5436

30 11.5275 0.00320 0.0779 0.0047 0.67150 1.05141 1051.41 129.730 20.4814

35 15.8196 0.00439 0.0779 0.0047 0.92152 1.30668 1306.68 178.032 25.4541

40 21.6435 0.00601 0.0779 0.0047 1.26078 1.59081 1590.81 243.574 30.1565

45 24.9733 0.00694 0.0779 0.0047 1.45474 1.77060 1770.60 281.047 31.2850

50 26.5705 0.00738 0.0779 0.0047 1.54778 1.85785 1857.85 299.022 30.8713

350


Cw 65.4 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

18.5 0.41835 0.00012 0.0525 0.00216 0.05368 0.65955 659.552 15.7802 7.04997

20 1.51784 0.00042 0.0525 0.00216 0.19477 0.96268 962.679 57.2528 11.7928

25 6.31842 0.00176 0.0525 0.00216 0.81077 2.33562 2335.62 238.331 28.6113

30 11.5275 0.00320 0.0525 0.00216 1.47919 3.65780 3657.80 434.818 44.8081

35 15.8196 0.00439 0.0525 0.00216 2.02995 4.89595 4895.95 596.716 59.9753

40 21.6435 0.00601 0.0525 0.00216 2.77726 6.68596 6685.96 816.394 81.9030

45 24.9733 0.00694 0.0525 0.00216 3.20454 7.83540 7835.40 941.995 95.9836

50 26.5705 0.00738 0.0525 0.00216 3.40949 8.40154 8401.54 1002.24 102.919


Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 5 2.52

2.65 0.504 5.7 2.8728

3.6 0.504 6.5 3.276

4.89 0.504 7.5 3.78

6.64 0.504 8.5 4.284

9.03 0.504 9.5 4.788

12.3 0.504 11 5.544

16.7 0.504 12.5 6.3

22.7 0.504 14.5 7.308

30.8 0.504 17 8.568

41.9 0.504 19 9.576

57 0.504 22.5 11.34

77.5 0.504 26.5 13.356

105 0.504 31.5 15.876

143 0.504 38 19.152

195 0.504 48 24.192

351


Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h

-1

20 57.1883 2.812775 4.760991 29.86101 0.776546

25 71.09452 3.949975 11.53262 29.87278 4.642472

30 86.60956 5.902066 19.37864 29.88568 9.400545

35 100.691 7.383966 26.46584 29.97597 13.73973

40 117.5438 8.863896 34.97882 30.03338 18.48393

45 142.0503 10.44072 46.32229 31.12372 23.89391

50 156.8372 11.67123 54.19213 30.26491 27.66873


65.9 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 0.77655 0.00022 0.0779 0.0047 0.04524 0.56256 562.555 4.64428 5.95211

25 4.64247 0.00129 0.0779 0.0047 0.27043 0.79000 789.995 27.7652 15.3891

30 9.40055 0.00261 0.0779 0.0047 0.54760 1.18041 1180.41 56.2217 22.9945

35 13.7397 0.00382 0.0779 0.0047 0.80037 1.47679 1476.79 82.1730 28.7680

40 18.4833 0.00513 0.0779 0.0047 1.07672 1.77278 1772.78 110.547 34.5337

45 23.8939 0.00665 0.0779 0.0047 1.39187 2.08814 2088.14 142.902 40.6771

50 27.6687 0.00769 0.0779 0.0047 1.61176 2.33425 2334.25 165.478 43.5231

352


65.9 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 0.77655 0.00022 0.0525 0.00216 0.09965 0.95220 952.198 15.1840 11.3425

25 4.64247 0.00129 0.0525 0.00216 0.59572 2.30652 2306.52 90.7756 28.6798

30 9.40055 0.00261 0.0525 0.00216 1.20626 3.87573 3875.73 183.812 48.1916

35 13.7397 0.00382 0.0525 0.00216 1.76306 5.29317 5293.17 268.657 65.8164

40 18.4839 0.00513 0.0525 0.00216 2.37183 6.99577 6995.77 361.422 86.9868

45 23.8939 0.00664 0.0525 0.00216 3.06603 9.26446 9264.46 467.204 115.196

50 27.6687 0.00769 0.0525 0.00216 3.55041 10.8384 10838.4 541.015 134.768


Cw 65.9 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 0.77655 0.00022 0.0779 0.0047 0.04524 0.56256 562.555 8.35779 5.95211

25 4.64247 0.00129 0.0779 0.0047 0.27043 0.79000 789.995 49.9658 15.3891

30 9.40055 0.00261 0.0779 0.0047 0.54760 1.18041 1180.41 101.176 22.9945

35 13.7397 0.00382 0.0779 0.0047 0.80037 1.47679 1476.79 147.878 28.7680

40 18.4833 0.00513 0.0779 0.0047 1.07672 1.77278 1772.78 198.938 34.5337

45 23.8939 0.00665 0.0779 0.0047 1.39187 2.08814 2088.14 257.165 40.6771

50 27.6687 0.00769 0.0779 0.0047 1.61176 2.33425 2334.25 297.792 43.5231

353


Cw 65.9 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 0.77655 0.00022 0.0525 0.00216 0.09965 0.95220 952.198 28.0907 11.3425

25 4.64247 0.00129 0.0525 0.00216 0.59572 2.30652 2306.52 167.936 28.6798

30 9.40055 0.00261 0.0525 0.00216 1.20626 3.87573 3875.73 340.054 48.1916

35 13.7397 0.00382 0.0525 0.00216 1.76306 5.29317 5293.17 497.019 65.8164

40 18.4839 0.00513 0.0525 0.00216 2.37183 6.99577 6995.77 668.635 86.9868

45 23.8939 0.00664 0.0525 0.00216 3.06603 9.26446 9264.46 864.335 115.196

50 27.6687 0.00769 0.0525 0.00216 3.55041 10.8384 10838.4 1000.88 134.768


Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 5.1 2.5704

2.65 0.504 6.1 3.0744

3.6 0.504 7.1 3.5784

4.89 0.504 8.2 4.1328

6.64 0.504 10 5.04

9.03 0.504 11.5 5.796

12.3 0.504 13.5 6.804

16.7 0.504 15.5 7.812

22.7 0.504 17.9 9.0216

30.8 0.504 20.5 10.332

41.9 0.504 24 12.096

57 0.504 28 14.112

77.5 0.504 33 16.632

105 0.504 40 20.16

143 0.504 48 24.192

195 0.504 60 30.24

354


Speed

Hz

P1

kPa

𝛥𝑃1

80 mm

kPa

𝛥𝑃2

50 mm

kPa

Temperature

°C

Flow

Q

m3 h

-1

20 64.38282 3.661099 8.239466 30.97937 1.692234

25 71.32618 4.374523 11.77027 30.95108 3.293618

30 86.16474 6.370775 17.83412 31.05845 7.071055

35 101.5071 8.084167 25.85571 31.04811 10.90857

40 117.4817 9.377321 34.18955 31.11131 15.15158

45 136.354 11.01835 43.9137 31.30402 19.80295

50 156.6241 12.46799 53.64508 31.46812 24.15993

55 165.1077 13.26112 57.65472 31.62423 25.97977


66.6 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 1.69223 0.00047 0.0779 0.0047 0.09858 0.73222 732.220 10.1207 14.2636

25 3.29362 0.00092 0.0779 0.0047 0.19186 0.87491 874.905 19.6981 17.0431

30 7.07156 0.00196 0.0779 0.0047 0.41190 1.27416 1274.16 42.2897 24.8205

35 10.9088 0.00303 0.0779 0.0047 0.63545 1.61683 1616.83 65.2407 31.4959

40 15.1516 0.00421 0.0779 0.0047 0.88262 1.87546 1875.46 90.6168 36.5340

45 19.8030 0.00550 0.0779 0.0047 1.15356 2.20367 2203.67 118.435 42.9275

50 24.1599 0.00671 0.0779 0.0047 1.40736 2.49360 2493.60 144.493 48.5753

55 25.9798 0.00722 0.0779 0.0047 1.51337 2.65222 2652.22 155.377 51.6653

355


66.6 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 1.69223 0.00047 0.0525 0.00216 0.21715 1.64789 1647.89 33.0887 21.6286

25 3.29362 0.00092 0.0525 0.00216 0.42263 2.35405 2354.05 64.4011 30.8970

30 7.07156 0.00196 0.0525 0.00216 0.90738 3.56682 3566.82 138.262 46.8146

35 10.9088 0.00303 0.0525 0.00216 1.39977 5.17114 5171.14 213.298 67.8712

40 15.1516 0.00421 0.0525 0.00216 1.94423 6.83791 6837.91 296.263 89.7476

45 19.8030 0.00550 0.0525 0.00216 2.54108 8.78274 8782.74 387.213 115.274

50 24.1599 0.00671 0.0525 0.00216 3.10016 10.7290 10729.0 472.406 140.818

55 25.9798 0.00722 0.0525 0.00216 3.33368 11.5309 11530.9 507.990 151.344


Cw 66.6 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 1.69223 0.00047 0.0779 0.0047 0.09858 0.73222 732.220 17.6958 14.2636

25 3.29362 0.00092 0.0779 0.0047 0.19186 0.87491 874.905 34.4417 17.0431

30 7.07156 0.00196 0.0779 0.0047 0.41190 1.27416 1274.16 73.9426 24.8205

35 10.9088 0.00303 0.0779 0.0047 0.63545 1.61683 1616.83 114.072 31.4959

40 15.1516 0.00421 0.0779 0.0047 0.88262 1.87546 1875.46 158.441 36.5340

45 19.8030 0.00550 0.0779 0.0047 1.15356 2.20367 2203.67 207.081 42.9275

50 24.1599 0.00671 0.0779 0.0047 1.40736 2.49360 2493.60 252.642 48.5753

55 25.9798 0.00722 0.0779 0.0047 1.51337 2.65222 2652.22 271.673 51.6653

356


66.6 %

Speed

Hz

Flow

Q

m3 h-1

Flow

Q

m3 s-1

Pipe

Dia.

m

Pipe

Area

m2

Velocity

V

m s-1

Pressure

kPa m-1

Pressure

Pa m-1

Shear

Rate

s-1

Shear

Stress

Pa

20 1.69223 0.00047 0.0525 0.00216 0.21715 1.64789 1647.89 57.4102 21.6286

25 3.29362 0.00092 0.0525 0.00216 0.42263 2.35405 2354.05 111.738 30.8970

30 7.07156 0.00196 0.0525 0.00216 0.90738 3.56682 3566.82 239.890 46.8146

35 10.9088 0.00303 0.0525 0.00216 1.39977 5.17114 5171.14 370.081 67.8712

40 15.1516 0.00421 0.0525 0.00216 1.94423 6.83791 6837.91 514.028 89.7476

45 19.8030 0.00550 0.0525 0.00216 2.54108 8.78274 8782.74 671.828 115.274

50 24.1599 0.00671 0.0525 0.00216 3.10016 10.7290 10729.0 819.642 140.818

55 25.9798 0.00722 0.0525 0.00216 3.33368 11.5309 11530.9 881.381 151.344


Shear Rate

(s-1)

MS

Factor

Indicator

Reading

Shear Stress

(Pa)

1.95 0.504 5.2 2.6208

2.65 0.504 6.9 3.4776

3.6 0.504 8.2 4.1328

4.89 0.504 9.5 4.788

6.64 0.504 11.2 5.6448

9.03 0.504 13 6.552

12.3 0.504 15.5 7.812

16.7 0.504 18 9.072

22.7 0.504 21 10.584

30.8 0.504 26 13.104

41.9 0.504 29 14.616

57 0.504 34 17.136

77.5 0.504 40 20.16

105 0.504 48 24.192

143 0.504 58 29.232

195 0.504 73 36.792

Documents

Modelling the pumping characteristics of power station ash