University of Wollongong Thesis Collection University of ... · quite extensively in the design of pneumatic conveying systems, is investigated and found to be inadequate in particular

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University of Wollongong Thesis Collection University of Wollongong Thesis Collections

1989

Pneumatic conveying of bulk solidsP. W. WypychUniversity of Wollongong

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Recommended CitationWypych, P. W., Pneumatic conveying of bulk solids, Doctor of Philosophy thesis, Department of Mechanical Engineering, Universityof Wollongong, 1989. http://ro.uow.edu.au/theses/1590

http://ro.uow.edu.au/


http://ro.uow.edu.au

http://ro.uow.edu.au/theses

http://ro.uow.edu.au/thesesuow



PNEUMATIC CONVEYING OF BULK SOLIDS

A thesis submitted in fulfilment of the

requirements for the award of the degree of

DOCTOR OF PHILOSOPHY

from , UNIVERSITY OF llum 'WOLLONGONG

LIBRARY

THE UNIVERSITY OF WOLLONGONG

by

P. W. WYPYCH, BE, MIEAust.

Department of Mechanical Engineering

1989

This is to certify that this work has not been submitted for a degree to any other university or institution

Peter W. Wypych

Dedicated to my wife, Linda, and my children, David, Emma and Amanda

for their love, support and patience.

i

SUMMARY

The pneumatic conveying of bulk solids through pipelines has been used in industry for several decades. With the introduction in recent years of new techniques and more efficient hardware, there has been a considerable increase in the use of this method of transport (e.g. dense-phase, low-velocity and longdistance conveying). Unfortunately, the technology available to assess the relative merits of the large number of commercial systems that now compete for a particular application is lacking sadly, especially when efficient and reliable dense-phase or long-distance transportation is required. The main objective of this thesis is to provide industry with some of this technology in relation to fine powders (e.g. pulverised coal, fly ash, P V C powder, fly ash/cement mix) and some coarser products (e.g. screened coke, crushed bath, granulated aluminate). A convenient method for presenting the variation of major steady-state conveying parameters is needed for efficient design, system evaluation and optimisation. One technique based on other work and extended to include saltation and minimum transport behaviour is established. A standardised-test procedure comprising three different types of pneumatic conveying experiment also is developed to generate efficiently the data required for this purpose. The method of scaling up test rig data to full-scale installations, previously used quite extensively in the design of pneumatic conveying systems, is investigated and found to be inadequate in particular applications. T w o popular forms of definition and three existing empirical correlations for the solids pressure drop are modified to demonstrate the possible extent of this inadequacy. Steady-state pipeline conveying characteristics of three products are used in the development of an improved scale-up procedure. Methods to predict the air-only pressure drop for both single- and stepped-diameter pipelines and to generalise the conveying characteristics of a particular material (applicable to other combinations of length and diameter) also are formulated and verified. Pulverised coal conveyed over 25 m and fly ash over 943 and 293 m (utilising three different configurations of blow tank) are used to investigate the effect of blow tank air injection on the performance of a pneumatic conveying system. The addition of supplementary conveying-air to a blow tank incorporating a top-air supply and transporting a good dense-phase material (pulverised coal) is shown to achieve higher values of mass flow ratio and/or conveying rate and also provide smoother and more consistent transportation. The installation of a fluidising discharge cone to the outlet of a blow tank conveying a cohesive fly ash is found to improve the discharge characteristics of the blow tank, as well as decrease pressure and flow rate fluctuations. The method of air injection also is found to have a significant impact on the plug-phase mode of conveying. Experiments on three different products are carried out to demonstrate the advantages of this method of transport (i.e. to handle conventionally difficult dense-phase materials, such as crushed bath) but also its sensitivity to changes in material property (viz. particle size). However, it is shown further that this may be compensated to some extent by selecting a different method of air injection.

ii

T w o powder classification techniques based on physical properties are evaluated and found useful in explaining and indicating the minimum transport (dense-phase) behaviour for a wide range of materials. The steady-state pipeline conveying characteristics (dilute- and dense-phase) and the fluidisation behaviour of ten products are compared for this purpose. Various mathematical models utilising numerical integration and analytical approximations are formulated to predict blow tank performance characteristics. Despite the lack of good accurate data for the experimental verification of these models (i.e. due to certain difficulties in measurement technique), preliminary results still are obtained and presented in graphical format. Five existing pipeline theories also are investigated and reviewed. One particular model is found useful in predicting the dense-phase conveying parameters of fine powders, and a worked example is presented. The applicability of generalised solids friction factor correlations to the design of pneumatic conveying systems is reviewed. The resulting degree of uncertainty is considered too great for applications involving relatively high operating pressures (e.g. long-distance and/or large-throughput conveying). Test rig data obtained from pulverised coal, a fly ash/cement mix and various fly ash samples are used to identify certain areas of improvement. Based on this work, a test-design procedure is developed to determine an accurate solids friction factor correlation (i.e. for a given material and a wide range of diameters). Results from recent investigations into the long-distance pneumatic conveying of pulverised coal are used to demonstrate the good accuracy and reliability of this improved approach.

iii

A C K N O W L E D G E M E N T S

The author gratefully acknowledges the guidance, continuous support and encouragement of his supervisor Professor P. C. Arnold throughout the course of this work.

The support provided by the following colleagues during the various stages of this work also are acknowledged sincerely by the author.

Mr O. C. Kennedy for his assistance with the laboratory test work and the processing of some of the experimental results and figures.

Mr D. M. Cook for his patience and assistance with the pneumatic conveying test work, construction and installation of the experimental apparatus.

The author particularly acknowledges the assistance provided by the staff of the Maintenance Workshop for the construction and installation of the various test rigs and equipment.

The financial support provided by the National Energy Research Development and Demonstration Council, the Australian Electrical Research Board and The University of Wollongong is acknowledged gratefully by the author.

The contributions made by Ramsey Engineering and Keystone Valve (A/Asia) Pty. Ltd. for the donation/supply of Clarkson knife-gate and butterfly valves respectively for the various test rigs also are acknowledged.

iv

TABLE OF CONTENTS

SUMMARY

ACKNOWLEDGEMENTS TABLE OF CONTENTS

LIST OF FIGURES LIST OF TABLES NOMENCLATURE

P a g e

i

iii

iv

vii

xiii

xvii

CHAPTER 1 INTRODUCTION

CHAPTER 2 PNEUMATIC CONVEYING TEST RIGS

2.1

2.2

2.3

2.4

2.5

2.6

2.7

2.8

TEST RIG A

TEST RIG B

TEST RIG C

TEST RIG D

TEST RIG E

TEST RIG F

AIR SUPPLY AND FLOW RATE MEASUREMENT

DATA ACQUISITION

6

7

14

15

18

21

24

26

CHAPTER 3 PNEUMATIC CONVEYING CHARACTERISTICS 27

3.1 PULVERISED COAL

3.2 DEFINITION OF DENSE-PHASE

3.3 FLY ASH

3.3.1 Introduction

3.3.2 Test Rig Description

3.3.3 Test Results

3.4 STANDARDISED-TEST PROCEDURE

3.4.1 Experiments

3.4.1.1 Test 1 - Standard Batch Cycle

3.4.1.2 Test 2 - Increase of Apt for Approximately Constant mf

3.4.1.3 Test 3 - Decrease of mf at Steady-State Conditions

3.4.2 Results

3.4.3 Minimum Transport Behaviour

3.4.4 Test Procedure Applications and Limitations

31

35

36

36

37

39

45

46

46

48

48

48

53

55

CHAPTER 4 BLOW TANK CONFIGURATION & AIR INJECTION 60

4.1

4.2

4.2.1

4.2.2

4.3

4.3.1

4.3.2

4.3.3

4.3.4

PULVERISED COAL

FLY ASH

Introduction

Test Results

PLUG-PHASE CONVEYING

62

66

66

68

72 Screened & Unscreened Granulated Aluminate (SGA & UGA) 72 Bone Char 77

Crushed Bath 80

Summary 82

CHAPTER 5 POWDER CHARACTERISATION 86

5.1 INTRODUCTION

5.2 PHYSICAL PROPERTIES

5.2.1 Definitions of Particle Size

5.3 FLUIDISATION

5.3.1 Experimental Apparatus

5.4 PIPELINE CONVEYING CHARACTERISTICS

5.5 POWDER CLASSIFICATION TECHNIQUES

5.5.1 Fluidisation

5.5.2 Slugging

5.5.2.1 Slugging Diagram Modifications

5.5.2.2 Results

87

87

87

90

90

95

95

97

100

100

102

CHAPTER 6 SCALE-UP CONVEYING CHARACTERISTICS 109

6.1 INTRODUCTION

6.2 SCALING RELATIONSHIPS 6.2.1 Definitions for Aps

6.2.2 Empirical Relationships

6.3 EXPERIMENTAL INVESTIGATIONS

6.3.1 Fly Ash / Cement Mix

6.3.2 Screened Coke

6.3.3 PVC Powder

6.4 SCALE-UP OF Apt

6.5 SUMMARY

110

113

113

115

117

117

120

122

126

1

VI

6.6 GENERALISED PIPELINE CONV. CHARACTERISTICS 130

CHAPTER 7 THEORETICAL INVESTIGATIONS 133

7.1 INTRODUCTION

7.2 BLOW TANK DISCHARGE CHARACTERISTICS

7.2.1 Approximate Analytical Solution

7.2.1.1 Results

7.2.1.2 Discussion

7.2.2 Numerical Analysis

7.2.2.1 Results

7.3 DENSE-PHASE PIPELINE CONV. CHARACTERISTICS

7.3.1 Pressure Loss Predictions by Muschelknautz &

Krambrock [59]

7.3.1.1 Theory

7.3.1.2 Calculation Procedure

7.3.1.3 Worked Example

7.4 CORRELATION ANALYSIS AND STEPPED-DIAMETER

PIPELINES

7.4.1 Generalised Correlation for Solids Friction Factor

7.4.2 Design of Stepped-Diameter Pipelines

7.4.2.1 Stepping Pipe Criteria

7.4.3 Test-Design Procedure

134

134

136

137

138

139

142

144

144

145

146

149

151

152

159

160

162

CHAPTER 8 CONCLUSIONS 174

8.1 FURTHER WORK 177

CHAPTER 9 REFERENCES 179

APPENDIX A

APPENDIX B

APPENDIX C

APPENDIX D

Compilation of Particle Size Data (Samples 1 to 11,

Table 5.1) 186

Modified Slugging Diagram based on Dixon [23,39] and

Cliftefa/. [41] 191

Compilation of Operating Conditions for Correlation Analysis

(Samples 1 to 11, 12 and 13, Table 7.2) 197

Summary of Solids Friction Factor Calculations for Pulverised

Brown Coal (Test-Design Procedure, Section 7.4.3) 203

Vll

LIST O F FIGURES

Chapter 2 Page

Figure 2.1 Configuration of the original 0.425 m3 blow tank (Test Rig A). 6

Figure 2.2 General arrangement of the original pneumatic conveying

Test Rig A. 8

Figure 2.3 Configuration of the final 0.425 m 3 blow tank (Test Rig B). 9

Figure 2.4 Exploded view of a typical pipeline air pressure tapping

location. 11

Figure 2.5 Full-sectional view of a 50 m m N.B. 90° blinded-tee bend. 12

Figure 2.6 General arrangement of Test Rig B. 13

Figure 2.7 Configuration of the original 0.9 m 3 blow tank (Test Rig C). 14

Figure 2.8 General arrangement of Test Rig C. 16

Figure 2.9 Configuration of the original tandem 0.9 m 3 blow tank feeding

system (Test Rig D). 17

Figure 2.10 Configuration of the final tandem 0.9 m 3 blow tank feeding

system (Test Rig E). 19

Figure 2.11 General arrangement of Test Rig E1 (refer to Figure 2.8 for

arrangement of pipe loops). 20

Figure 2.12 Configuration of the 0.113 m 3 plug-phase blow tank. 22

Figure 2.13 General arrangement of Test Rig F. 23

Figure 2.14 General arrangement of compressed air supply. 25

Figure 2.15 HP-85B plot of a typical uncalibrated pipeline air pressure

transducer response. 26

Chapter 3

Figure 3.1 General form of steady-state pneumatic conveying

characteristics for a given material and pipeline configuration. 29

Figure 3.2 Alternative form of pneumatic conveying characteristics. 29

Figure 3.3 The Rizk [7] two-phase flow diagram for pneumatic conveying

in horizontal pipes. 30

Figure 3.4 Pneumatic conveying characteristics of pulverised coal for

Test Rig A1 (L = 25 m & D = 52 m m ) , displaying lines of

constant Apj. 31

VIII


Test Rig A1 (L = 25 m & D = 52 m m ) , displaying lines of constant ms. 32


Test Rig A1 (L = 25 m & D = 52 m m ) displaying lines of constant, steady-state m s (Apj ordinate). 33


Test Rig A1 (L = 25 m & D = 52 m m ) , displaying lines of constant, steady-state m s (Apt ordinate). 34

Figure 3.8 Schematic layout of the pneumatic conveying Test Rig B1

used during the fly ash test program. 38

Figure 3.9 Pipeline conveying characteristics of Eraring fly ash for L = 71

m & D = 52 m m (Test RigB1). 41


m & D = 52 m m (Test Rig B1). 42

Figure 3.11 Pipeline conveying characteristics of Tallawarra fly ash for L =

71 m & D = 52 m m (Test Rig B1). 42

Figure 3.12 Pipeline conveying characteristics of Munmorah fly ash for L =

71 m & D = 52 m m (Test Rig B1). 43

Figure 3.13 Pipeline conveying characteristics of Vales Point fly ash for L =

71 m & D = 52 m m (Test Rig B1). 43

Figure 3.14 Pipeline conveying characteristics of Gladstone fly ash for L =

71 m & D = 52 m m (Test Rig B1). 44

Figure 3.15 Pipeline conveying characteristics of Wallerawang fly ash for L

= 71 m & D = 52 m m (Test Rig B1). 44

Figure 3.16 Pipeline conveying characteristics of Liddell fly ash for L = 71

m & D = 52 m m (Test Rig B1). 45

Figure 3.17 Transient plots of major conveying parameters for Eraring fly

ash demonstrating Test 1 (Test Rig B1, Exp. No. 236). 47





Figure 3.20 Pipeline air pressure drop (Test Rig B1, Exp. No. 236). 51


m & D = 52 m m (Test Rig B1) demonstrating Tests 1, 2 and 3. 53

IX


ash demonstrating blockage condition using Test 3 (Test Rig

B1, Exp. No. 232). 54

Figure 3.23 Pipeline conveying characteristics of P V C powder [21] for L =

71 m & D = 52 m m (Test Rig B1). 56

Figure 3.24 Transient plots of major conveying parameters for P V C

powder demonstrating plugging condition using Test 2 (Test

Rig B1, Exp. No. 387). 57

Figure 3.25 Transient plots of major conveying parameters for P V C

powder demonstrating plugging condition using Test 3 (Test

Rig B1, Exp. No. 414). 58

Chapter 4

Figure 4.1 0.425 m3 Sturtevant™ blow tank and air supply arrangement. 62

Figure 4.2 Transient plots of major conveying parameters from Exp. Nos.

21, 23 and 35 for pulverised coal conveyed over 25 m (Test

RigA1). 63

Figure 4.3 Transient plots of major conveying parameters from Exp. Nos.

61 and 62 for pulverised coal conveyed over 25 m (Test Rig

A1). 65

Figure 4.4 Configuration of bottom-discharge blow tank demonstrating

incomplete discharge of material due to rat-holing. 67

Figure 4.5 Configuration of top-discharge blow tank demonstrating

incomplete discharge of material due bad channelling and rat-

holing. 68

Figure 4.6 Blow tank comparison using fly ash and Test Rig D2 (L = 940

m & D = 60/69/81/105 m m ) . 70

Figure 4.7 Blow tank comparison using transient plots of major conveying

parameters for fly ash conveyed over 293 m (Test Rig D1). 71

Figure 4.8 Transient plots of major conveying parameters for S G A (Exp.

No. 1274, Test Rig F2). 74

Figure 4.9 Particle size distributions of S G A and UGA. 75

Figure 4.10 Transient plots of major conveying parameters for U G A (Exp.

No. 1356, Test Rig F2). 76

Figure 4.11 Transient plots of blow tank and pipeline air pressure for bone

char (Exp. Nos. 1227 & 1235, Test Rig F2). 79

x

Figure 4.12 Transient plots of blow tank and pipeline air pressure for bone

char (Exp. Nos. 1237 & 1245, Test Rig F2). 81

Figure 4.13 Transient plots of major conveying parameters for crushed

bath (Exp. No. 108-12, orifice-air only, Test Rig F3). 83

Figure 4.14 Transient plots of major conveying parameters for crushed

bath (Exp. No. 108-16, orifice-, ring- and supplementary-air,

Test Rig F3). 84

Chapter 5

Figure 5.1 Schematic layout of the fluidisation test facility. 92

Figure 5.2 Comparison of fluidisation curves for pulverised coal (Sample

1) and fly ash (Samples 2 to 8). 93

Figure 5.3 Fluidisation curves of P V C powder (Sample 9) and screened

coke (Sample 10). 94

Figure 5.4 Comparison of pipeline conveying characteristics for fly ash

(Samples 2 to 8, Test Rig B1). 96

Figure 5.5 The Geldart [24] fluidisation diagram. 97

Figure 5.6 The Geldart [24] fluidisation diagram showing the location of

Samples 1 to 11. 99

Figure 5.7 The Dixon [23] slugging diagram for a 50 m m pipe diameter

system. 101

Figure 5.8 The Dixon [23] slugging diagram for a 100 m m pipe diameter

system. 101

Figure 5.9 The modified Dixon [23] slugging diagram for a 50 m m pipe

diameter system showing the classification of Samples 1 to 11

listed in Table 5.1. 103

Figure 5.10 Transient plots of major conveying parameters demonstrating

flow irregularities for Sample 6 (Exp. No. 662, Test Rig B1). 104

Figure 5.11 Pipeline conveying characteristics of screened coke [14,16,26]

for L = 25 m & D = 52 m m (Test Rig A1). 106

Chapter 6

Figure 6.1 Pipeline conveying characteristics of fly/ash cement mix for L|

= 162 m & Di = 0.060 m (Test Rig C1). 118

Figure 6.2 Pipeline conveying characteristics of fly/ash cement mix for L|

= 1 6 2 m & D i =0.105 m (Test Rig C3). 118

xi

Figure 6.3 Scale-up conveying characteristics of fly/ash cement mix for L2

= 162 m & D2 = 0.105 m based Figure 6.1 and Equation (6.6). 119

Figure 6.4 Scale-up conveying characteristics of fly/ash cement mix for L2

= 162 m & D2 = 0.105 m based on Figure 6.1 and Equation (6.28) with rj = 2.8. 120

Figure 6.5 Scale-up conveying characteristics of screened coke for L2 =

71 m & D2 = 0.052 m (based on Figure 5.11 and Equations

(6.5) to (6.7)) with four experimental data points from Test Rig

A 2 (L-i = 71 m & Di = 0.052 m ) . 121

Figure 6.6 Pipeline conveying characteristics of P V C powder for Li = 162

m & Di = 0.105 m (Test Rig C3). 122

Figure 6.7 Scale-up conveying characteristics of P V C powder for L2 =

162 m & D2 = 0.105 m, based on Figure 3.23 and Equation

(6.29). 123

Figure 6.8 Scale-up conveying characteristics of P V C powder for L2 =

162 m & D2 = 0.105 m, based on Figure 3.23 and Equation

(6.30). 124 Figure 6.9 Variation of Apt according to Equation (6.37) with experimental

data points obtained from six different pipeline configurations. 127

Figure 6.10 Generalised pipeline conveying characteristics of fly

ash/cement mix based on Test Rig C1 results (L|' = 162 m &

Di= 0.060 m). 131

Figure 6.11 Generalised pipeline conveying characteristics of fly

ash/cement mix based on Test Rig C3 results (L-|' = 162 m &

Di= 0.105 m). 131

Chapter 7

Figure 7.1 The Enstad [62] element of a converging flow channel. 135

Figure 7.2 Example of blow tank model results (approximate analytical

solution). 140

Figure 7.3 Example of blow tank model results (numerical solution). 143

Figure 7.4 Full-bore plug transport system. 145

Figure 7.5 Variation of velocity ratio [59]. 147

Figure 7.6 Variation of particle free settling velocity based on the Clift et

al. [41] drag correlations. 148

Figure 7.7 Correlation of pipe friction coefficient due to solids according

to Stegmaier [68]. 153

xii

Figure 7.8 Comparison between experimental data and the Stegmaier

[68] correlation. 1 5 5

Figure 7.9 Improved correlation of pipe friction coefficient due to solids. 157

Figure 7.10 Examples of air pressure drop for the Dj = 0.060 m section of

pipe, showing the location of the three 1 m radius x 90°

bends. 164 Figure 7.11 Relationship between Xs and Frm showing actual values of

m*. 166

Figure 7.12 Relationship between Xs and X = Frm p f m0 2 showing

experimental values of m* and predicted curves, based on

Equation (7.48). 168

Figure 7.13 Relationship between Y and X, where Y = Xs (m*)0-5 and X =

Frmpfm0-2. 169

Figure 7.14 Comparison between actual and predicted values of Xs, based

on Equation (7.48). 170

Figure 7.15 Pipeline conveying characteristics of pulverised coal for L =

947 m and D = .060/.069/.081/.105 m (Test Rig E1), showing

experimental data points and predicted curves, based on

Equation (7.48). 171

Appendix B

Figure B.1 The modified Dixon [39] slugging diagram for a 52 mm pipe

diameter system. 192

Figure B.2 T h e modified Dixon [39] slugging diagram for a 7 8 m m pipe


Figure B.3 T h e modified Dixon [39] slugging diagram for a 102 m m pipe

diameter system. 1 9 4

Figure B.4 T h e modified Dixon [39] slugging diagram for a 154 m m pipe


Figure B.5 T h e modified Dixon [39] slugging diagram for a 2 0 3 m m pipe


xiii

LIST O F T A B L E S

Chapter 2 Page

Table 2.1 Pipeline details for Test Rigs A1 & A2.

Table 2.2 Pipeline details for Test Rig B1.

Table 2.3 Pipeline details for Test Rigs C1, C2, C3 & C4.

Table 2.4 Pipeline details for Test Rigs D1 & D2.

Table 2.5 Pipeline details for Test Rigs E1.

Table 2.6 Pipeline details for Test Rigs F1.

Table 2.7 Orifice plate details.

7

10

15

18

19

21

24

Chapter 3

Table 3.1 List of power station fly ash samples. 36

Table 3.2 Chronology of the fly ash test program. 39

Table 3.3 Summary of experiments and data points for Eraring fly ash. 52

Table 3.4 Steady-state operating conditions obtained from Exp. Nos.

236, 240 and 249. 52

Chapter 4

Table 4.1 Physical properties of test materials. 61

Table 4.2 Set-up conditions for the blow tank air injection experiments. 64

Table 4.3 Conveying parameters of fly ash for L = 293 m & D = 69 m m

(Test Rig D1). 69

Table 4.4 Cumulative % mass passing through sieve size (for orifice-

and ring-air). 78

Table 4.5 Cumulative % mass passing through sieve size (for orifice-,

ring- and supplementary-air). 80

Table 4.6 Summary of plug-phase conveying parameters for crushed

bath (Test Rig F3, L = 160 m & D = 105 mm). 82

Chapter 5

Table 5.1 List of samples and physical properties. 90

Chapter 6

Table 6.1 Physical properties of test materials. 117

Table 6.2 Comparison of predicted and actual values of ms for mf > 0.3

kg s'1 (i.e. based on Figures 6.2 and 6.3). 119

Table 6.3 Summary of screened coke results for Test Rig A2 (Li = 71 m

& D 2 = 0.052 m). 121

Table 6.4 Empirical expressions for Apt. 128

Table 6.5 Long-distance pneumatic conveying pipeline (Test Rig D2). 129

Table 6.6 Comparison of experimental and theoretical values of Apt for

the long-distance pneumatic conveying stepped diameter

pipeline (Test Rig D2). 129

Chapter 7

Table 7.1 Summary of results obtained from Steps 4, 5, 6 and 7.

Table 7.2 Summary of products and experimental data for correlation

analyses.

Table 7.3 Pipeline configuration for Test Rig E1.

Table 7.4 Steady-state operating conditions for the 947 m Test Rig E1

pipeline.

Table 7.5 Predicted values of pressure drop for the test rig pipeline,

based on Equation (7.48).

Table 7.6 Comparison between experiment and predicted values of Apt.

Table 7.7 Suggested pipeline configurations and predicted operating

conditions for pulverised brown coal conveyed at 241 rr1 over

L = 1800 m.

150

154

163

164

170 171

172

Appendix A

Table A.1 Mass percentage frequency distribution for Tallawarra

pulverised coal (Sample 1), using the Coulter Counter. 187

Table A.2 Mass percentage frequency distribution for Tallawarra fly ash

(Sample 2), using the Coulter Counter. 187

Table A.3 Mass percentage frequency distribution for Eraring fly ash


Table A.4 Mass percentage frequency distribution for Munmorah fly ash,


XV

Table A.5

Table A.6

Table A.7

Table A.8

Table A.9

Table A. 10

Table A. 11

Mass percentage frequency distribution for Vales Point fly ash


Mass percentage frequency distribution for Gladstone fly ash


Mass percentage frequency distribution for Wallerawang fly

ash (Sample 7), using the Coulter Counter. 189

Mass percentage frequency distribution for Liddell fly ash


Mass percentage frequency distribution for PVC powder

(Sample 9), using the sieve test. 189

Mass percentage frequency distribution for screened coke

(Sample 10), using the sieve test. 1 go

Mass percentage frequency distribution for coarse fly ash

(Sample 11), using the Malvern analyser. 190

Appendix C

Table C.1 Steady-state operating conditions of pulverised coal (Sample

1) for Test Rigs A1 (L = 25 m & D = .052 m) and A3 (L = 96 m &

D = .052m). 198

Table C.2 Steady-state operating conditions of Tallawarra fly ash

(Sample 2) for Test Rig B1 (L = 71 m & D = .052 m). 198

Table C.3 Steady-state operating conditions of Eraring fly ash (Sample

3) for Test Rig B1 (L = 71 m & D = .052 m). 199

Table C.4 Steady-state operating conditions of Munmorah fly ash


Table C.5 Steady-state operating conditions of Vales Point fly ash


Table C.6 Steady-state operating conditions of Gladstone fly ash

(Sample 6) for Test Rigs B1 (L = 71 m & D = .052 m) and C3 (L

= 162 m & D = .105 m). 200

Table C.7 Steady-state operating conditions of Wallerawang fly ash


Table C.8 Steady-state operating conditions of Liddell fly ash (Sample 8)

for Test Rig B1 (L = 71 m & D = .052 m). 201

xvi

Table C.9 Steady-state operating conditions of fly ash/cement mix

(Sample 12) for Test Rigs C1 (L = 162 m & D = .060 m) and C3

(L = 162 m & D = .105 m). 202

Table C.10 Steady-state operating conditions of fly ash [59] (Sample 13)

for L = 1200 m & D = .200 m. 202

Appendix D

Table D.1 Solids friction factor calculations for pipe section No. 1 (Di =

0.105 m & ALi = 150.0 m). 204

Table D.2 Solids friction factor calculations for pipe section No. 2 (D2 =

0.081 m & AL2 = 261.0 m). 204


0.069 m & AL3 = 390.0 m). 205


0.060 m & AL4 = 146.0 m). 205

NOMENCLATURE

a A A1.A2.A3

As

b

c

Co

Cp

CV

c

d

°50 dp

dp50

dpm

dpwm

dsv

dsvm

dv

dV50

dvm

dvwm

dpg/dL

D Dj

DP

Do DT e E Ek»Ev f

Fr

Exponent in permeability Equation (7.10)

Cross-sectional area of pipe, A = 0.25 TC D'2

Variables in velocity Equation (7.7)

Surface area of Enstad [62] element

Exponent in compressibility Equations (7.9) and (7.18)

Permeability coefficient of bulk solid, Equation (7.4)

Value of c when 01 = a-|0 Value of c when 01 = oip

Volumetric concentration of solids, Equation (6.13)

Constant relating o 0 with dynamic head at outlet, Equation (7.13) Drag coefficient, Equation (7.26)

Particle diameter

Median particle diameter

Arithmetic mean of adjacent sieve sizes

Value of dso based on a sieve size distribution

Mean particle size from a standard sieve analysis, Equation (5.1)

Weighted mean diameter based on a sieve analysis, Equation (5.2)

Diameter of a sphere with the same surface area to volume ratio as the particle

Mean surface volume diameter, Equation (5.3)

Diameter of a sphere with the same volume as the particle

Value of dso based on a volume diameter distribution

Mean equivalent volume diameter, Equation (5.4)

Volume weighted mean diameter, Equation (5.5)

Pipeline air pressure gradient due to solids

Internal diameter of pipe

Value of D for pipe section No. i

Differential pressure

Outlet diameter of blow tank

Diameter of blow tank at transition

Exponent used in the equation X = Frm (pfm)e, Section 7.4.3

Constant in Equation (7.39)

Variables used in the Ergun [64] Equation (7.17)

Exponent used in the equation Y = kg (m*)f, Section 7.4.3

Froude No., Equation (7.24)

http://A1.A2.A3

xviii

Fr-i, Fr2 Value of Fr at upstream and downstream end of test or pipe section

Fr m M e a n value of Fr based on Equation (7.38)

Frmin Minimum reliable value of Fr

Frs Particle Froude No., Equation (7.25)

g Acceleration due to gravity, g « 9.81 m s_1

Gi Constant used in Equations (7.11) and (7.12) and to define A2

h D Height of bed of material in a fluidisation test chamber

i Numbering system used to designate different sections of pipe or

different ranges of particle size

k Constant in Equation (6.26)

K Ratio of vertical to horizontal pipeline air pressure gradient,

Equation (6.33)

K1 Constant in Equation (7.45), K-| = 10 - <°

L Total effective [9] length of pipeline L' Value of L modified to allow for bends

Lh Value of L for all horizontal pipe sections

Lv Value of L for all vertical pipe sections

mt Air mass flow rate

rrifm Maximum value of mf

m s Solids mass flow rate m s' Value of m s modified to allow for bends

m s o Value of m s at outlet of blow tank

m * Material to mass flow rate ratio, m* = m s m f1

Mbt Mass of solids loaded into blow tank M s Mass of conveyed solids

n Maximum number of different sections of pipe in a stepped-diameter pipeline

Nb Number of bends

N B Nominal bore of pipe

p Gauge air pressure inside blow tank

Pbt Blow tank top-air pressure gauge

Pbt.i Initial value of pbt

PA1 Blow tank top-air pressure gauge (transducer location A1, Figure

3.8)

PAi.i Initial value of P A I

PC4 Pipeline air pressure gauge (transducer location C 4 , Test Rig D1,

» 78.3 m from blow tank outlet)

xix

PG1 Pipeline air pressure gauge (transducer location G1, Test Rig B1,

« 15.7 m from blow tank outlet, Figure 3.8)

PG2 Pipeline air pressure gauge (transducer location G 2 , Test Rig B1,

= 17.8 m from blow tank outlet, Figure 3.8)

Po Value of p at blow tank outlet

P T Value of p at blow tank transition

P Absolute air pressure inside blow tank, P= p + P a t m

Patm Atmospheric air pressure, Patm » 1010 hPa or 101 kPa abs

Pfi Initial (or upstream) absolute air pressure of a pipeline section

Pf2 Final (or downstream) absolute air pressure of a pipeline section

Pfm Mean absolute air pressure of test or pipe section Qf Volumetric flow rate of air

r Radial distance from vertex of flow channel

r0 Value of r to blow tank outlet

rr Value of r to blow tank transition

r' Radius of Enstad [62] element, Figure 7.1

R Gas constant for air, R = 287.1 N m kg-1 K*1

Re s Particle Reynolds No. defined by Equation (7.27)

t Cycle time

tc Conveying cycle time

T Absolute air temperature

"h, T2, T3, T4 Variables used to define A-(, A 2 and A 3

u Interstitial air velocity inside blow tank

vs Solids velocity

vso Value of vs at blow tank outlet

vsj Value of vs at blow tank transition

vTO Terminal velocity or free settling velocity of particle Vf Superficial air velocity

Vfi, Vf2 Value of Vf at upstream and downstream end of test or pipe section

Vfm Mean value of Vf based on pfm, Vf m = 4 mf (% ptm D2)"1

VfS Value of Vf that almost produces saltation of a material under load

conditions

VfS0 Value of Vf that almost produces saltation of a single particle

Vf.min Minimum superficial conveying air velocity (at minimum or reliable

transport limit)

V p Velocity of a full-bore plug, Figure 7.4

XX

W i Variable used to define A-j

x Constant in Equation (6.9)

X Abscissa variable used in correlation analysis, X = Frm (pfm)e

Xi Variable used in Enstad [62] theory and to define Ai, A2 and A3

y Exponent in Equation (6.9)

Y Ordinate variable used in correlation analysis, Y = Xs (m*)f

Yi, YY1 Variables used in Enstad [62] theory and to define A 3 and G1

Z-| Variable used to define A 2 and A 3

a Half angle of converging flow channel

p Angle between major principal stress and normal to hopper wall for

flow conditions

y Material coefficient, Equation (7.21)

8 Effective angle of internal friction of a bulk solid

e Power index for air density ratio in Equation (6.20)

£ Power index for pipeline length ratio in Equation (6.20)

H Power index for pipe diameter ratio in Equations (6.20) and (6.28)

0 Exponent in Equation (6.31)

X\ Overall friction factor for test section No. i, X\ = Xi\ + m* Xs\

Xf Air-only friction factor

Xu Value of Xf for test section No. i

Xs Pipe friction coefficient due to solids

Xs\ Value of ̂ s for test section No. i

Lif Absolute or dynamic air viscosity

£ Variable defined by Equation (7.23)

Pb Bulk density of bulk solid

Pbc Value of pb when 01 = aic

Pbi Loose-poured bulk density of bulk solid

pb0 Value of pb when 01 = 010

Pt Air density

pf.atm Value of pf at atmospheric conditions

pf.max Value of pf at maximum operating pressure

pfm Mean value of air density based on Equation (7.37)

ps Solids density

p* Density ratio defined by Equation (7.22)

a Mean consolidation stress

o 0 Value of o at blow tank outlet

XXI

a-| Major consolidation stress

o-jc Reference value of 01 to define pb, Equation (7.9)

010 Value of 01 at blow tank outlet

aip Reference value of G-\ to define c, Equation (7.10)

oj Value of o at blow tank transition

x , v Exponents in Equation (7.39)

<j> Kinematic angle of friction between a bulk solid and a hopper wall

X , co Exponents in Equation (7.45)

\|r Particle sphericity, Equation (5.6)

r Porosity or voidage of a bulk solid, r = 1 - (pb ps"1)

AdPi Size range No. i for sieve size distribution

AdVj Size range No. i for volume diameter distribution

AL Length of test section

ALj Value of AL for test section No. i

A M Mass percent of material contained in a given size range

AMj Value of A M for size range No. i

Ap Air pressure drop for test section of length AL

Api Value of Ap for Dj and ALj (i.e. test section No. i)

Apb Air pressure drop across bed of material in a fluidisation chamber

Apbt Air pressure drop across material in a blow tank

Apt Air-only pipeline pressure drop

Ap s Pipeline air pressure drop due to solids

Aps* Value of A p s modified according to Equation (6.34)

ApF Air pressure drop across receiving hopper filter

Apt Total pipeline air pressure drop, Apt = Apt + A p s

Apu Value of Apt for Dj and ALj (i.e. for pipe section No. i)

A p j Total system pressure loss (including feeder)

Subscripts

1 Experimental data pertaining to a test rig

2 Scale-up data pertaining to an actual or proposed system

f Fluid or air

h Horizontal

i Pipe or test section number

s Solids or particles

v Vertical

1

CHAPTER 1

2

1. INTRODUCTION

The pneumatic transportation of bulk solids is continuing to gain popularity for a wide range of applications, especially as more efficient hardware and techniques are introduced onto the market (e.g. long-distance [1] and low-velocity [2] conveying). Subsequently, there has been a substantial increase in the number of commercial systems available to industry. The main features that make this method of transport attractive to the designer of materials handling plants are listed below. The relative ease of routing the conveying pipeline (e.g. verticals, bends,

inclines) adds flexibility to the design or upgrading of a plant.

The physical size of a pneumatic conveying pipeline is small compared to an equivalent conveyor-belt/bucket-elevator system (especially for the dense-phase mode of transport which requires usually relatively small sizes of pipe).

Atmospheric contamination is avoided due to the completely enclosed nature of the transport system (e.g. dusty, hygroscopic and even toxic products are able to be conveyed safely and hygienically).

New technology allows friable products to be transported at low-velocity and with either extremely low or undetectable levels of product degradation or damage [3]. As a consequence, air consumption and hence running costs are reduced significantly. Also, erosion of the system (e.g. bends, conveying pipeline) is minimised.

The use of a pipeline can offer increased security as opposed to an open-belt conveyor system (e.g. for diamond recovery plants).

With improved hardware (e.g. blow tanks) and techniques (e.g. solids metering [1], air injection [4], stepped-diameter pipelines [5,6]), several materials such as pulverised coal, cement and fly ash are able to be transported efficiently at large conveying rates (e.g. 100 to 200 t Ir1) and/or over long distances (e.g. 1 to 3 km).

Unfortunately, the technology available to assess for a given application the relative merits of the competing systems is lacking sadly, particularly when dense-phase [7] or long-distance conveying is considered. Although Flain [4] and more recently Klintworth and Marcus [2] have provided a general overview of several of the more c o m m o n types of commercial system and also have indicated their fields of application, the potential user of such equipment usually is faced still with the difficult problem of selecting the most appropriate configuration (i.e. in terms of cost and, more importantly, operational efficiency and reliability). Furthermore, when attempting to design or optimise a pneumatic conveying system, the following additional difficulties need to be overcome.

3

(a) Establishing a standardised-test procedure to determine sufficient information on the material (e.g. from a test rig) and also deciding on what data are relevant for a particular requirement (e.g. plant specification). Also, it is necessary to present this information in an efficient and workable form.

(b) Scaling-up the test rig data to the full-scale system.

(c) Determining minimum transport behaviour to optimise the operating conditions (e.g. for the dilute- [7] or dense-phase mode of conveying).

(d) Choosing between dilute-, dense-, pulse-phase [8] or low-velocity conveying as the most suitable method of transport for a given material and specification. Also, the most efficient feeder (e.g. blow tank, rotary valve, screw feeder) and method of air injection [4] need to be selected in terms of reliability, running costs, product conditioning requirements and maintaining a constant and reliable feed rate of product into the pipeline.

(e) Determining an optimal size of pipe for a proposed pipeline route (over short and long distances) and also predicting operating conditions (e.g. pressure drop for a given air flow and product conveying rate). Also, a stepped-diameter pipeline [5] m a y need to be considered for longdistance conveying applications.

(f) Predicting operating conditions for existing or working installations (e.g. for the requirements of troubleshooting or uprating system capacity).

(g) Establishing the feasibility of transporting a certain material in the dense-phase mode or over long distances (e.g. up to 3 km).

(h) Minimising hardware problems and improving the reliability of system instrumentation and control (e.g. level indicators, discharge and vent valves, bend/pipe erosion).

The main aim of this thesis is to provide industry with some of the technology that is needed in relation to Items (a) to (g) above. Particular research objectives include

determining and presenting pneumatic conveying characteristics for the purposes of system comparison, optimisation of operating conditions and general design,

investigating the effect of blow tank configuration and method of air injection on pneumatic conveying performance,

assessing the influence of material properties on conveying characteristics and minimum transport behaviour,

evaluating existing and developing improved techniques to scale-up test rig data to a full scale installation,

improving/developing mathematical models and computer software to predict system design parameters (viz. for the blow tank and pipeline) and verifying these predictions by experiment.

4

This thesis investigates many aspects of pneumatic conveying and does result in significant improvements to the procedures for designing and/or selecting pneumatic transport systems. Note that due to the wide ranging nature of this thesis, the literature appropriate to each chapter is reviewed in the relevant section(s) of work.

A total of fourteen different test rigs comprising six configurations of blow tank were employed throughout the test work program and are described in the following sections (i.e. Chapter 2). Note that the majority of the experimental work and investigations undertaken in this thesis were limited mainly to

bottom-discharge blow tank feeders, which were available in the Bulk Solids Handling Laboratory at The University of Wollongong,

fine powders (e.g. pulverised coal, fly ash, fly ash/cement mix, PVC powder) and some coarser products (e.g. crushed bath, bone char, screened coke).

Chapter 3 is concerned with the development of a technique to represent test rig data efficiently for the purpose of general design. Referred to as pipeline conveying characteristics, this representation of data includes the dilute- and dense-phase regimes, minimum transport boundaries and the air-only component of pressure drop.

Investigations into the effect of blow tank air injection on the performance of pneumatic conveying systems are considered in Chapter 4. This is followed by an evaluation of powder classification techniques in Chapter 5 to determine for a given material the suitability of dense-phase transportation (i.e. based on bench-type experiments).

Chapter 6 is concerned with the scale-up of test rig data and the development of an improved technique based on experimental data. The generalisation of pipeline conveying characteristics also is included as an alternative method.

The problem of estimating blow tank discharge characteristics and pipeline operating conditions is considered in Chapter 7. Also, the development of a new technique (based on correlation analysis) to predict total pipeline air pressure drop (for single- or stepped-diameter pipelines, as well as short- or long-distance conveying) is included.

Finally, concluding remarks and suggestions for further work based on the investigations and results presented in this thesis are contained in Chapter 8.

5

CHAPTER 2

6

2. PNEUMATIC CONVEYING TEST RIGS

A total of ten test rigs incorporating six different configurations of blow tank and/or method of air injection were used to obtain all the data necessary for the various aspects of this thesis project. The test rigs were developed at different stages over a period of approximately six years. A general description of the overall test facility and five case studies to emphasise the need for large-scale product testing prior to design, have been presented recently by Wypych and Arnold [10]. The main purpose of this section is to provide a description of each test rig configuration and, where necessary, a brief explanation, as appropriate, of some of the more important features, modifications and/or improvements. Note that a system of letters and numbers is employed to label each test rig configuration. A letter is used to refer to a particular blow tank feeding system (including its method of air injection) and numbers are used to designate different pipeline layouts fed by the same blow tank. For example, Test Rigs A1 and A2 refer to different conveying pipelines (viz. L = 25 and 96 m, respectively) fed by the same blow tank. 2.1 Test Rig A In 1980, the Electricity Commission of N.S.W. provided funds to the University of Wollongong for the purchase of a Sturtevant™ Pulse-Phase Powder Conveyor. This original test rig was installed as part of an undergraduate thesis project during 1980 and consisted of the following major components.

Material Inlet

Top Air Vent Filter

Fluidising Ring Air

Original S Discharge Valve

Figure 2.1 Configuration of the original 0.425 m 3 blow tank (Test Rig A).

7

A 0.425 m 3 capacity blow tank with a maximum safe working pressure (S.W.P.) of 350 kPag (see Figure 2.1) and fitted with a 50 m m N.B. discharge valve (viz. a stainless steel ball valve). An electro-pneumatic control cabinet housing all the necessary control equipment for conveyor operation (e.g. pressure regulators). A 0.5 m 3 receiving hopper supported by tension load cells to monitor the delivered mass of solids. A D C E Model U M A 70V venting dust control unit mounted on top of the 0.5 m 3 receiving hopper and fitted with polypropylene filter bags. Four horizontal loops of mild steel pipeline with change-over sections to provide effective [9] conveying distances of 25, 48, 71 or 96 m. A general arrangement of the original test rig and the 25 m pipe loop is presented in Figure 2.2. Only three different configurations of pipeline were used in this thesis project and are summarised in Table 2.1. Test Rig

A1

A2

A3

L (m)

25

71

96

D (m)

.052

.052

.052

Lv(m)

3.6

3.6

3.6

Lh (m)

21.4

67.4

92.4

No. & Type of Bends

5 x 1 m radius 90° bends



Table 2.1 Pipeline details for Test Rigs A1 & A2.

2.2 Test Rig B

Operating principles of the original Sturtevant™ blow tank (i.e. as shown in Figure 2.1) were based on the pulse-phase concept [8] and were found to have considerable limitations. Several modifications to the blow tank and its operating sequence were necessary to allow sufficient versatility for testing purposes (e.g. extending the range of air flow) and to fulfil the requirements of the thesis project in general (e.g. investigating different methods of air injection, determining pipeline conveying characteristics of various products). The following list summarises the major modifications and improvements that were carried out to Test Rig A. The vent filter, which is shown in Figure 2.1 and is used to remove the

displaced air during the filling cycle of the blow tank, was found to be ineffective for fine powders such as pulverised coal and was replaced by a 25 m m N.B. ball valve and pipe connected directly to the 0.5 m 3

receiving hopper. The blow tank outlet was modified to provide additional air for transportation (viz. conveying-air). Also, the existing discharge valve, which proved unsuitable for fly ash (e.g. the stainless steel ball valve seized frequently), was replaced by a Figure 990 Keystone butterfly valve and positioner. Note that this valve was bolted directly to the outlet flange of the blow tank. Refer to Figure 2.3 for the final configuration of the 0.425 m 3 blow tank.

8

25m of 52mm I.D. Mild Steel Pipeline

Ring Air

Original Discharge-Valve

\ Knife Air

Figure 2.2 General arrangement of the original pneumatic conveying Test Rig A.

Material Inlet

Vent Air

Top Air

Fluidising Ring Air

Discharge Valve

Conveying Air

Figure 2.3 Configuration of the final 0.425 m 3

blow tank (Test Rig B).

10

Shear-beam-type load cells were installed on the blow tank to monitor the supplied mass of solids. An accurate weighing-scale system was introduced to calibrate the load cells mounted on both the receiving hopper and blow tank. Three orifice plate assemblies with D and D/2 pressure tappings and designed according to B.S. 1042: Part 1:1964 were installed to measure the amount of air being used in various sections of the test rig (viz. conveying-air, blow tank top- and ring-air). Numerous pressure tappings were installed along the pipeline, so that air pressure gradients could be recorded. Refer to Figure 2.4 for an exploded view of a typical pressure tapping location. An efficient pipeline unblocking technique (using an in-line backpressure valve) was installed at the end of the pipeline to minimise the amount of stoppage time due to blockages. This valve was used also to pressurise the pipeline and blow tank, so that all pressure transducers could be calibrated accurately at selected pressures. The two 1 m radius 90° bends, which were connected to the vertical pipe in Test Rig A, were replaced by two 90° blinded-tee bends (see Figure 2.5) and connecting spool pieces. This was carried out for the main purpose of increasing the actual length of vertical pipe to provide more accurate measurements of the vertical pipeline air pressure gradient. Note that despite these modifications, the effective conveying distances of Test Rig A essentially remained unchanged. Figure 2.6 presents a general arrangement of Test Rig B showing the four horizontal pipe loops, which provided a total effective conveying distance of 96 m. Table 2.2 provides details on the 71 m pipeline, which was the only configuration used in this thesis project. Test Rig

B1

L(m)

71

D (m)

.052

Lv (m)

3.6

Lh (m)

67.4

No. & Type of Bends

11 x 1 m radius 90° bends and

2 x 90° blinded-tee bends

Table 2.2 Pipeline details for Test Rig B1.

The air knife (see Figure 2.1) was removed from the pipeline (mainly due to its ineffectiveness on materials such as pulverised coal, see Section 4.1). All other components used on Test Rig A (e.g. the control cabinet, receiving hopper and venting dust control unit) essentially remained unchanged for Test Rig B. However, during the test program on fly ash, which is discussed later in Section 3.3, the polypropylene filter bags of the dust control unit were replaced with epitropic Goretex™ bags (due to an excessive build up of material, see Section 3.3.3).

11

Pressure Transducer"

j^L

Retaining Screw

Porex Disc

Quick-Connect Coupling

1/4" BSPT Thread

TM O-Ring

1/4" BSP Socket

52mm I.D. Pipeline

Figure 2.4 Exploded view of a typical pipeline air pressure tapping location.

12

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L. ^

/ / / /

-£. / / / /

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50mm N.B. Table E Flanges

V /////////. >; ;/;;//;///;;//;;// r-ry

52mm I.D. Mild Steel Pipeline ~

3

Direction — of

Flow

Figure 2.5 Full-sectional view of a 50 m m N.B. 90° blinded-tee bend.

13

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14

2.3 Test Rig C

In 1982, NEI John Thompson (Aust) formed a consortium with Kloeckner-Becorit Industrietechnik-KBI G m b H and then negotiated with the University of Wollongong to install a pneumatic conveying test rig in the Bulk Solids Handling Laboratory. A blow tank fitted with a cone dosing valve, which is a solids metering device designed primarily for long distance transportation, was imported from West Germany and installed initially with 162 m of 65 m m N.B. Schedule 80 (i.e. 60 m m I.D.) mild steel pipe. However, after preliminary test work, the length and size of the pipeline was found to be inadequate (i.e. for the needs of Australian industry) and additional pipework was installed. The following list summarises the major components of the test rig, which was used in this thesis project. A 0.9 m3 capacity blow tank with a maximum S.W.P. of 700 kPag and

fitted with a cone dosing valve and a 100 m m N.B. discharge Argus™ ball valve (see Figure 2.7). Also, the blow tank is supported by shear-beam-type load cells. A pneumatic PI controller with an adjustable set point (i.e. maximum operating pressure) is used to control the stroke and oscillation frequency of the cone dosing valve. The measured air pressure signal is taken upstream of the blow tank discharge valve. However, during the major part of the test program for this thesis project, the cone dosing valve was .not used and was either removed physically or raised in a position not to interfere with the normal operation of the blow tank. Note, the cone dosing valve was required only for recent investigations into the long-distance pneumatic conveying of pulverised brown coal (refer to Section 7.4.3).

Conveying Air

Vent ' Line

Pipeline

Figure 2.7 Configuration of the original 0.9 m 3 blow tank (Test Rig C).

15

A 5 m 3 receiving silo supported by two load blocks to monitor the delivered mass of solids. However, due to possible eccentric loading effects, the conveying rate during an experiment is determined from the response of the blow tank load cells. A D C E Model D L M V8/7B reverse jet insertable vent filter mounted on top of the 5 m 3 receiving silo and fitted with polypropylene filter bags. An NEI John Thompson (Aust) standard mini-pot [3] to transfer conveyed material to the blow tank (see Figure 2.8). Horizontal loops of mild steel pipeline with change-over sections to provide effective [9] conveying distances from 42 to 940 m. Figure 2.8 presents a general arrangement of Test Rig C and the 940 m pipeline (containing 60, 69, 81 and 105 m m I.D. sections of pipe). However, only four different pipeline configurations were used in this thesis project and these are summarised in Table 2.3.

Test Rig

C1

C2

C3

C4

L(m)

162

59

162

553

D (m)

.060

.105

.105

.069

Lv(m)

4.4

4.5

4.5

4.4

Lh (m)

157.6

54.5

157.5

548.6

No. & Type of Bends





Table 2.3 Pipeline details for Test Rigs C1, C2, C 3 & C4.

2.4 Test Rig D

During investigations into long-distance pneumatic transportation on Test Rig C, the capacity of the 0.9 m 3 blow tank (i.e. refer to Figure 2.7) was found to be insufficient for the establishment of steady-state flow conditions (i.e. during the conveying cycle). To overcome this deficiency and other problems (e.g. overpressurisation of the 5 m 3 silo due to material build up on the polypropylene filter bags - similar to the problem described in Section 2.2), the following improvements and modifications were carried out to Test Rig C.

A second 0.9 m3 blow tank with a max. S.W.P. of 700 kPag was manufactured by NEI John Thompson (Aust) and installed alongside the original KBI G m b H blow tank (see Figure 2.9). Due to occasional feeding problems from the original blow tank (e.g. rat-holing which is discussed later in Section 4.2), the new blow tank was fitted with a fluidising discharge cone. Also, an improved cone dosing valve and actuator was installed.

16

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17

Material Inlet

Material Inlet

Aeration Air

Fluidising Discharge Cone

Conveying Air

Figure 2.9 Configuration of the original tandem 0.9 m 3 blow tank feeding system (Test Rig D).

The 0.7 m long polypropylene bags of the D L M V8/7B insertable vent filter were replaced by 1.5 m long epitropic Goretex™ bags. This effectively converted the filter unit to a Model D L M V8/15B.

Apart from the above-mentioned modifications to the filter, the rest of the 5 m 3 silo, as described previously for Test Rig C, remained unchanged.

The pipeline installations for Test Rig C were not modified. Throughout the thesis project, only two pipeline configurations were fed by the blow tank system shown in Figure 2.9, and these are summarised in Table 2.4. Note that Test Rig D 2 employs a stepped-diameter pipeline, which is used primarily to minimise pressure drop and conveying velocity (i.e. mainly for long-distance pneumatic conveying applications [4,5]).

18

Test Rig

D1

D2

L(m)

293

146 940 390

261 143

D (m)

.069

.060

.069

.081

.105

Lv (m)

4.4

4.5

Lh (m)

288.6

146.0 390.0 261.0 138.5

No. & Type of Bends


3 x 1 m radius 90° bends 13 x 1 m radius 90° bends 8 x 1 m radius 90° bends 5 x 1 m radius 90° bends

Table 2.4 Pipeline details for Test Rigs D1 & D2.

2.5 Test Rig E

After carrying out several experiments on Test Rig D and also analysing the transient plots of major conveying parameters, the following disadvantages were realised.

(a) The method of air injection used on Blow Tank No. 1 (see Figure 2.9) was inferior to the one used on Blow Tank No. 2 (e.g. refer to Section 4.2 for some typical results on fly ash). This prevented proper tandem operation of the blow tanks (i.e. discharging one blow tank after another and achieving the same steady-state flow conditions).

(b) Despite the problems referred to in (a), the duration of discharge of the two 0.9 m 3 blow tanks, during the high-pressure (i.e. large conveying rate) experiments, was insufficient to maintain steady-state operating conditions.

(c) The mini-pot [3] filling system for the blow tanks was slow and tedious and allowed only a limited number of experiments to be carried out during one day of testing.

As a result of these deficiencies, the following modifications and improvements were carried out to Test Rig D.

Blow Tank No. 1 was replaced with one of similar design to Blow Tank No. 2 (Le. see Figure 2.10).

A 3 m3 receiving silo was designed, manufactured and installed on top of the tandem blow tank system to provide gravity discharge and also enable continuous transportation (viz. in conjunction with a Programable Logic Controller, which was contained inside the original control panel of Test Rig C). Refer to Figure 2.11 for a general arrangement of Test Rig E and the pipe loops providing a total possible conveying distance of 947 m (see Table 2.5). Note that a D C E Model D L M V12/7B reverse jet insertable vent filter (complete with epitropic Goretex™ filter bags, antistatic provisions, horizontal upstand, polypropylene explosion panels and a pressure relief flap valve) was mounted on top of the receiving silo.

19

The pipeline configurations of Test Rig D were extended slightly to connect the existing pipework to the new 3 m 3 silo. However, during this thesis project, only one configuration of pipeline was used (i.e. refer to Table 2.5).

Test Rig

E1

L(m)

146 947 390

261 150

D (m)

.060

.069

.081

.105

Lv (m)

7.0

Lh (m)

146.0 390.0 261.0 143.0

No. & Type of Bends

3 x 1 m radius 90° bends 13 x 1 m radius 90° bends 8 x 1 m radius 90° bends 5 x 1 m radius 90° bends

Table 2.5 Pipeline details for Test Rigs E1.

Material Inlet

Vent

Conveying Air ».

Figure 2.10 Configuration of the final tandem 0.9 m 3 blow tank feeding system (Test Rig E).

20

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21

2.6 Test Rig F

Some coarse materials such as crushed coal and petroleum coke can be conveyed more efficiently in the plug-phase mode (i.e. where a limited amount of material, usually in the form of a plug, is conveyed through the pipeline per cycle). Note this method of transport is similar to the pulse-phase mode [8], except that instead of conveying numerous plugs or slugs of material through the pipeline, only one plug of material is transferred during each cycle. Also, note that a discharge valve is not required usually for a plug-phase blow tank. NEI John Thompson (Aust), as a part of their development program to design and market a plug-phase conveying system, supplied to the University of Wollongong a 0.113 m 3 blow tank, which was connected to the existing 105 m m I.D. pipeline of Test Rig C. Additional changeover sections were installed to provide intermediate conveying distances of 41, 58, 80 and 100 m. This test rig consisted of the following major components. A 0.113 m3 capacity plug-phase blow tank with a maximum S.W.P. of 700

kPag (see Figure 2.12).

A 5 m3 receiving silo supported by two load blocks to monitor the delivered mass of solids. However, due to possible eccentric loading effects (see Section 2.3), the batch conveying rate during an experiment is determined by dividing the actual mass of conveyed solids (removed from the silo and weighed on a load platform) by the conveying time (i.e. m s = M S V

1 ) - The overall conveying rate is determined by allowing for transient effects such as blow tank fill time, tf, and total valve switching time ,tv (i.e. overall m s = M s (tc + tf + tv)"

1). Note that this 5 m 3 silo and its vent filter are identical to those described for Test Rig D (i.e. in Section 2.4).

Of the five different possible configurations of pipeline, only three were used for this thesis project. Table 2.6 provides a summary of the relevant details. Also, refer to Figure 2.13 for a general arrangement of Test Rig F.

Test Rig

F1

F2

F3

L(m)

41

58

161

D (m)

.105

.105

.105

Lv (m)

4.5

4.5

4.5

Lh (m)

36.5

53.5

156.5

No. & Type of Bends




Table 2.6 Pipeline details for Test Rigs F1.

22

Material Inlet

Orifice Air

Figure 2.12 Configuration of the 0.113 m 3 plug-phase blow tank.

23

Figure 2.13 General arrangement of Test Rig F.

24

2.7 Air Supply and Flow Rate Measurement

Air at a maximum pressure head of« 800 kPag is supplied from any combination of the three following rotary screw compressors.

Atlas Copco electric-powered Model GA-308, 3.1 m3 min-1 free air delivery.

Ingersoll Rand diesel-powered Model P375-WP, 10.6 m3 min"1 free air delivery.

Ingersoll Rand diesel-powered Model P850-WGM, 24.1 m3 min-1 free air delivery.

The compressors are connected to an aftercooler, two refrigerated air dryers and two air receivers (1.75 and 6.0 m 3 capacity). Various filters and separators are installed in series with these compressors to ensure a dry and oil-free air supply. Figure 2.14 provides a general arrangement of the air supply system.

Depending on the test rig and desired rangeability of flow rate, one of the five orifice plates listed in Table 2.7 (with D and D/2 pressure tappings and designed according to B.S. 1042 : Part 1 : 1964 ) is selected to monitor the conveying air usage.

Orifice Plate

+ ++

No.

1 2 4 5 6

Orifice Dia. (mm)

14.73 9.98 20.65 33.08 44.55

Pipe Dia. (mm)

26.64 25.30 78.10 78.10 78.10

Min. mf+ (kg s-1)

.027

.012

.050

.130

.250

Based on a DP of = 381 m m H2O @ 600 kPag & 20° C. Based an a DP of - 3810 mm H2O @ 600 kPag & 20° C.

Max. mf+ +

(kg s-1)

.085

.037

.155

.410

.775

Test

Rig

A&B A&B CtoF CtoF CtoF

Table 2.7 Orifice plate details.

-{§—04-P850-WGM 24.1 m '/min

Ingersoll Rand Compressors

To Test Rig

Pressure Regulator

Filters'

Aflercooler 4>-V-4>

o Atlas Copco Compressor

Dryer

HXr-B) <J>-

D and 0/2 tapping*

t/iiiiMYSir>/i'/V. Z%4m

OHfle. plol..

uzzzzzzzzzzzzzzzi

>?>>>>i//rrm

V. fr T7ZZZZZZZZZZZZZ2ZZZZ

* ORIFICE PLATE DETAIL

Figure 2.14 General arrangement of compressed air supply.

26

2.8 Data Acquisition

During the early stages of this thesis project, priority was given to developing the necessary software to capture the voltages of up to 20 analogue channels using a portable Hewlett-Packard 3054A Data Acquisition System. Major components of the system included a HP-85B desk-top computer, a HP-3497A scanning control unit and a transducer signal-conditioning unit.

Typical transducer channels, which were recorded with respect to cycle time, included : blow tank top-air pressure; pipeline air pressure; upstream pipeline and differential air pressures for the orifice plate assemblies; the mass of material entering the receiving hopper and/or leaving the blow tank. After storing these responses on either the HP-85B computer or a Tektronix 4923 digital tape recorder, the data are then transferred to the University's Univac™ mainframe computer for final processing and graphical output. On-site graphics also were developed on the HP-85B computer, so that plots of raw data also could be achieved easily after the completion of any experiment. An example of a typical pipeline air pressure response copied from the HP-85B C R T screen is presented in Figure 2.15. Throughout the course of this project, the software of all major programs was updated and improved continually as required. For example, memory capacity and scanning speed were increased recently to accommodate a maximum number of 64 channels for the investigations into long-distance pneumatic conveying.

10

>

c o Q. VI

m

A

Plot of C h.

\

14

CD

<\J

Scan Number

Figure 2.15 HP-85B plot of a typical uncalibrated pipeline air pressure transducer response.

27

CHAPTER 3

28

3. PNEUMATIC CONVEYING CHARACTERISTICS

If a pneumatic conveying system is to be designed or upgraded to ensure satisfactory and efficient operation, it is suggested that as much information as possible be obtained on the bulk solid to be handled (e.g. physical properties, conveying performance). Also, any possible operational problems should be investigated (e.g. material cohesion causing incomplete discharge of a blow tank, or unusual physical properties producing unforeseen blockage phenomena). It is important for efficient and reliable design, that any conveying performance data be summarised in a convenient and workable form. Several methods were investigated and the technique [11] that was selected finally is shown in Figure 3.1, which displays the variation of steady-state m s (the solids mass flow rate, kg s

_1) with respect to mt (the supplied air mass flow rate, kgs-1) and Apt (the total pipeline air pressure drop, kPa). An alternative form is presented in Figure 3.2, on which straight lines of constant mass flow ratio (m* = m s mf

1) are superimposed easily. Note that m* is adopted frequently by researchers to define the dense-phase mode of transport and compare the efficiencies of commercial systems. However, there are certain inadequacies with this form of definition and these are discussed further in Section 3.2. These two methods of data presentation were selected for the following reasons. (a) The data represent steady-state operating conditions and are accurate

for the selected configuration of pipeline. Hence, they usually are referred to as steady-state pipeline conveying characteristics of a given product. This information also can be applied to other pipelines of similar configuration and with different types of feeder (e.g. blow tanks, rotary valves, screw feeders) as long as the feed rate is consistent and steady-state.

(b) The use of air velocity instead of mf to represent operating conditions can be confusing and often leads to calculation error due to

the frequent lack of proper definition to distinguish between pick-up, average and exit velocity, as well as actual and superficial air velocity,

the functional dependency of air velocity on pressure and hence, pipeline length.

(c) Under steady-state conditions, mt is constant at any point along the pipeline (for both positive and negative pressure systems). Hence, the superficial velocity Vf may be calculated easily using the continuity equation Vf = mf (pf A)*1.

(d) the use of pressure gradient instead of actual pipeline air pressure drop can be misleading and often tempts the user to apply the data to vastly extrapolated lengths of pipeline. It is Important to be aware of the particular pipeline configuration that actually was used to generate the data.

APt (kPa)

\

1

1

1

I

I

1

1

^ > ^

•

mf (kgs~:)

Figure 3.1 General form of steady-state pneumatic conveying characteristics for a given material and pipeline configuration.

m< (kgs-1)

mf (kgs-1)

Figure 3.2 Alternative form of pneumatic conveying characteristics.

Logarithmic scales often are used on the abscissa and ordinate axes to represent the variation in conveying rate. However, the generation or utilisation of such graphs is far more tedious than the simpler linear scale representations, as shown in Figures 3.1 and 3.2. Also, note that the main reason for selecting log scales is to linearise the m s curves with respect to say velocity and pressure gradient, as shown in Figure 3.3. Unfortunately, such simplifications do not always occur especially for materials of wide particle size range and realistic pipelines containing bends and vertical sections.

dense phase dilute phase

1

< i

S

pressurized flow vacuum and pressurized flow steady stale unsteady statel steady state

; plug-dune | layer-disperse

Vmin

lag (average air velocity)

State Diagram for Horizontal Conveying

Product

Particle size

Density Pipe diameter

Pipe material

Styropor -3

dp = 2,385mm pp =1050 kg/m3

d =52,6mm

stainless steel Average pipe watt roughness R ~ 6 r 10 \im

Figure 3.3 The Rizk [7] two-phase flow diagram for pneumatic conveying in horizontal pipes.

31

3.1 Pulverised Coal

During the first two years of this thesis project, pulverised coal (obtained through the Electricity Commission of N.S.W. from the Tallawarra Power Station) was used as the test material. The first attempt to determine the pneumatic conveying characteristics of this material resulted in a series of experiments being carried out on the original test rig (i.e. Test Rig A1) with the following specification.

0.425 m3 blow tank with top-air only (refer to Figure 2.1). 25 m of 52 m m I.D. pipeline (refer to Figure 2.2). Five 1 m radius, 90° bends. 270, 240, 200, 165 and 135 kPag initial blow tank air pressures.

From the conveying parameters recorded for each experiment, values of mf, ms and A p j (total system pressure loss) were extracted at selected increments of the conveying cycle and plotted on a graph similar to that shown in Figure 3.2 (except for the use of Apj instead of Apt). Lines of constant Apjwere then drawn through the data to provide a family of curves at intervals of 10 kPa, as shown in Figure 3.4. With mf representing the abscissa axis and m s the ordinate axis, straight lines of material to air mass flow rate ratio, m*, also were drawn on Figure 3.4. The alternative method of presenting this information is shown in Figure 3.5 (adopting the form given in Figure 3.1). These methods of data presentation are similar to those presented by Mason etal. [11].

.004 .006 .008 .010 .012 .014

mf (kgs-1)

Figure 3.4 Pneumatic conveying characteristics of pulverised coal for Test Rig A1 (L = 25 m & D = 52 m m ) , displaying lines of constant Apj.

140

120 -

ApT

(kPa)

100 -

80

60 .004 .006 .012

32

.014 .008 .010

mf (kgs-1)

Figure 3.5 Pneumatic conveying characteristics of pulverised coal for Test Rig A1 (L = 25 m & D = 52 m m ) , displaying lines of constant ms.

It must be emphasised that such information is relevant only to the

conveyed material (viz. pulverised coal), blow tank (Sturtevant™, 0.425 m 3 capacity), pipeline/bend configuration (L = 25 m, D = 52 m m , five 1 m radius 90° bends), method of air injection (viz. top-air)

that were used for this particular set of experiments. However, if one of these conditions were to be varied with respect to the other three, then the technique would provide a very useful design tool. For example, the relative conveyability of different materials, the effects of pipeline length on conveying performance and a comparison of the various conveying modes may be summarised and evaluated easily on such plots.

Note that the above values of mf, ms and Apr were extracted at certain increments of the conveying cycle and, hence, actually represented instantaneous values. Furthermore, note that

where

and

Apj = Apbt + Apt + ApF (3.1)

Apj is the total system pressure loss (kPa),

Apbt is the pressure drop across the blow tank (kPa),

Apt is the total pipeline air pressure drop (kPa),

ApF is the pressure drop across the receiving hopper filter unit (kPa).

33

Assuming that A p F - 0 and noting that the final pressure of the system essentially is atmospheric, the value of A p T was taken to be numerically equal to the air pressure on top of the material in the blow tank. That is,

APT = (Pbt + Patm) - Patm = Pbt (3.2)

where pbt is the blow tank top-air pressure (usually transducer location A1), Patm is atmospheric pressure (usually = 101000 Pa abs).

During later work on fly ash (and especially in relation to the development of the standardised-test procedure, described in Section 3.4), it was decided to consider only steady-state conveying parameters and plot Apt instead of ApT. The reasons were : some doubt existed over the accuracy of the instantaneous curves drawn in Figures 3.4 and 3.5 (due to the pipeline creating a time-delay in the system); a graph displaying lines of constant Apt would be more applicable to other pipelines (of similar configuration) fed by different types of feeder (e.g. rotary valve) and blow tank configuration (e.g. top-discharge). To explore these matters further, steady-state values of mt, m s, Apt and A p j were obtained from the original conveying parameter plots and the resulting family of m s curves were plotted, as shown in Figures 3.6 and 3.7 (the former representing Apj and the latter Apt).

After comparing the trends displayed in Figures 3.5, 3.6 and 3.7, it can be seen that, although some similarities do exist, the m s contour lines displayed in Figures 3.6 and 3.7 are significantly flatter. W h e n these discrepancies were realised during the latter stages of the thesis (viz. during the fly ash test program), it was intended to apply the standardised-test procedure (described in Section 3.4) to the same pulverised coal sample (i.e. to determine more accurate conveying characteristics).

140

120

(kPa)

100

80

60

.004 .006 .008 .010 .012 .014

mf (kgs-1)

Figure 3.6 Pneumatic conveying characteristics of pulverised coal for Test Rig A1 (L = 25 m & D = 52 mm) displaying lines of constant, steady-state ms (Apj ordinate).

140

120 -

Apt

(kPa)

100 -

80 _

60 .004 .006 .008 .010

mf (kgs-1)

.012 .014

Figure 3.7 Pneumatic conveying characteristics of pulverised coal for Test Rig A1 (L = 25 m & D = 52 m m ) , displaying lines of constant, steady-state m s (Apt ordinate).

Unfortunately, due to the following reasons, this work was not able to be completed during the normal course of this project.

The original four 200 litre drum samples were no longer available. The d e m a n d s of the fly ash test program (including powder characterisation, long-distance conveying) reduced the emphasis on the investigations involving pulverised coal. During the initial experiments on fly ash, significant sparking was observed along a perspex sight tube installed at the end of the pipeline (i.e. on the original Test Rig A1). This was considered as a possible problem for pulverised coal, and the acquisition and testing of further material had to be postponed.

In fact, the latter prompted further investigations into possible sources of ignition (i.e. in all test rigs) and methods of explosion prevention and relief. The most applicable was found to be sparking caused by electrostatic charging of the conveying pipeline (due to inadequate earthing) and the filter bag surfaces (i.e. inside the venting dust control unit). Earthing of the perspex surface was attempted, but the tube was replaced eventually by a section of mild steel pipe to provide improved earthing of the pipeline and hence, eliminate any possibility of sparking. Electrostatic charging of the filter bag surfaces was still considered to be a problem and an extensive investigation resulted in the request to purchase (via University funds) a reverse-jet air filter comprising epitropic Goretex™ filter bags earthed to the filter housing,

a horizontal upstand fitted with an explosion relief panel, an exhaust fan to provide additional vacuum in the receiving hopper and the filter housing, and

35

a discharge duct to provide venting of any explosion directly to atmosphere.

The funding for this equipment was forthcoming and a DCE Model DLM V7/7F1 was purchased in April, 1985. The installation was postponed until July/August, 1985 due to the priority of completing the fly ash testing program. Note that the polypropylene filter bags which were used in the original D C E Model U M A 70V venting dust control unit were replaced with Goretex™ bags in August, 1984 to eliminate the dust emission problems that were being experienced with Vales Point fly ash. This aspect is discussed further in Section 3.3.3. Although accurate conveying characteristics were not able to be determined for the pulverised coal, the results that were obtained from the initial test work still were considered sufficient for the other requirements of this thesis (viz. blow tank air injection, powder characterisation and mathematical model verification). One other important aspect which stemmed from this initial work on pulverised coal was the need to clarify the definition for dense-phase. 3.2 Definition of Dense-Phase

Several researchers have adopted material to air mass flow rate ratio (viz. m*) as the basis of definition for dense-phase. For example, Mason et al. [12] have suggested that dense-phase conveyors operate normally with m* > 40; Duckworth [13] has indicated that for dense-phase suspensions, m* typically is greater than 100. However, as reported by Wypych and Arnold [14], this form of definition seems to be inadequate. The main reasons are m* is dependent on pipeline length for a given condition of flow, as

indicated by the scale-up criteria used by Mills et al. [15], and a material may display dense-phase pneumatic conveying characteristics at relatively low values of mass flow ratio (or dilute-phase performance at relatively high values of m*).

Such aspects particularly are important if long-distance transport is considered or if a theoretical prediction of pressure drop is required. S o m e researchers, for example, have compared discrete experimental pressure drop data with dilute-phase mathematical model predictions without verifying whether dilute- or dense-phase conditions actually are prevalent.

It would be more convenient to use a definition based on actual flow characteristics, and it is suggested that dense-phase should refer generally to the condition of non-suspension flow, whether it occurs in the form of unstable dunes [16], discrete full-bore plugs [17], sliding beds [18] or as an extruded condition [19]. Saltation is defined therefore, as the transition from dilute- to dense-phase (or suspension to non-suspension) pneumatic transport in horizontal pipes. Rizk [7] has provided a similar form of definition by using a phase diagram for horizontal pneumatic conveying. This diagram, which is similar to the Zenz [20] two-phase flow diagram, was presented earlier in Figure 3.3. Note the similarity with the pneumatic conveying characteristics given in Figure 3.1. A comparison between Figure 3.3 and Figure 3.7, indicate that the pulverised coal was transported in the dense-phase mode (i.e. for the range of mf values considered). Furthermore, this is supported by the relatively large values of m*. which were obtained from these experiments (e.g. 260< m* < 420).

36

3.3 Fly Ash

3.3.1 Introduction

During the first two years of this project, considerable interest was expressed by industry to extend the research work to investigate the pneumatic conveying of power station fly ash. Preliminary investigations on Vales Point fly ash in 1982, demonstrated that the handling problems significantly were greater than those associated with pulverised coal and emphasised the importance of continuing work in this area. Therefore, fly ash was included in the scope of work of this thesis and resulted in priorities being placed on comparing the conveyability of fly ash samples collected from six power stations (viz. Tallawarra, Eraring, Munmorah, Vales Point, Wallerawang and Liddell. Details of each sample and the type of boiler used are provided in Table 3.1.

Power Station Fly Ash Sample

Tallawarra

Eraring

Munmorah

Vales Point

Wallerawang

Liddell

Unit No.

6

1

2&3

5

8

4

Unit Load Op./Max. (MW)

90/100

400/660

190/350 (U2) 290/350 (U3) 630/660

460/500

390/500

Type of Boller+

A

B

C

D

D

D

Method of Dust Collection (Location)

Fabric Filter (Pozzolanic Tank)

Fabric Filter (Cell 23)

Electrostatic Precipitator (Zone 2)




+ Type of Boiler: A Tangentially-fired B Opposed-wall-fired conventional type conventional type

C Tangentially-fired D Tangentially-fired down-flow type tower type

Table 3.1 List of power station fly ash samples

Note that all the above samples were collected and supplied by the Electricity Commission of N.S.W. Particular objectives of the fly ash work included the determination of

complete pneumatic conveying characteristics displaying minimum transport behaviour (i.e. dense-phase performance),

relevant physical properties (e.g. particle diameter, solids density, loose-poured bulk density), and

fluidisation performance.

37

This information was intended to provide a data base, from which the pneumatic conveying performance of a given fly ash sample could be estimated qualitatively (i.e. after ascertaining the physical properties and fluidisation behaviour of the sample, as described later in Section 5.1).

As indicated in Section 3.1, in order to compare the conveyability of different materials, it was found necessary to employ the same blow tank, pipeline/bend configuration and operating modes. A detailed description of the test rig that was used in these investigations is given in the following section. Section 3.3.3 presents the pipeline conveying characteristics that were obtained for each fly ash sample.

3.3.2 Test Rig Description

A schematic layout of the test rig that was used extensively during the fly ash test program is presented in Figure 3.8, which includes the location of all recorded transducers. The following list provides a description of the major components of the test rig (i.e. Test Rig B1) and the relevant operating modes.

0.425 m3 Sturtevant™ blow tank with a maximum safe working pressure of 350 kPag (refer to Figure 2.3). Blow tank top-air, fluidising ring-air and conveying-air (originally referred to as supplementary-air) all were used in various amounts throughout the test program. 71 m total effective length of 50 m m Schedule 40 (52 m m I.D.) conveying pipeline. 3.6 m vertical lift (located 5.4 m from the blow tank outlet). Eleven 1 m radius, 90° bends. T w o blinded-tee bends connected to the vertical pipe (refer to Figure 2.5). A D C E Model U M A 70V venting dust control unit mounted on top of the receiving hopper.

Note that the effective length [9] of a 1 m radius 90° bend is 2.0 m, whereas the actual length is 1.57 m. Therefore, the total actual length of the pipeline is "71 m -11 x (2.0 - 1.57) « 66 m". However, for the purpose of scale-up (discussed later in Chapter 5), it was decided to use effective length (instead of actual length) to represent the total length of pipelines. The air supply consists of the Atlas Copco Model GA308 rotary screw compressor (see Section 2.7), an S M C Model LDP-150S refrigerated air dryer and two air receivers totalling 7.8 m 3 volumetric capacity. The capacity of the compressor is 3.1 Nm3min-1 (free air delivery) with a maximum pressure head of 800 kPag. Orifice Plate No. 1 (see Table 2.7) with D and D/2 pressure tappings and designed according to B.S.1042 : Part 1 : 1964 is selected to record the amount of air used during the conveying cycle.

All the important conveying parameters such as blow tank top-air pressure, pipeline air pressure (refer to Figure 2.4 for details of a typical pressure tapping location), supplied/delivered mass of solids and supplied air mass flow rate are recorded with respect to cycle time using the portable HP3054A Data Acquisition System (described earlier in Section 2.8). A chart recorder is used in conjunction with the Data Acquisition System to ensure that steady-state conditions occur during the conveying cycle (a requirement for the standardised-test procedure described in Section 3.4).

38

O)

cu • c

39

3.3.3 Test Results

The procedure for determining the pneumatic conveying characteristics of each fly ash sample consisted of the following six major steps.

1. Apply the standardised-test procedure (refer to Section 3.4).

2. Obtain steady-state values of mf, ms and Apt from the transient plots of the major conveying parameters recorded for each experiment.

3. Present this information in tabular form.

4. Plot values of ms against mf (abscissa axis) and Apt (ordinate axis), as depicted by the generalised form given in Figure 3.1. Note that Apt is the preferred primary variable since it is a limitation on the plant by the compressor used.

5. Using interpolation, draw lines of constant ms through the data at appropriate intervals.

6. Delineate any regions of unreliable transport displayed by the transient plots of both the pipeline air pressure and the delivered mass of solids. This supplementary information will determine the limits of reliable transport and the suitability of conveying the material in the dense-phase mode. The two main types of flow irregularity, which predominated during the fly ash test program, were unstable duning and blockage conditions. Examples are presented in Section 3.4.3, which also includes results from P V C powder, which was found to display unstable plugging or blockage conditions in the conventional dense-phase (or non-suspension) mode of transport.

The chronological order in which the fly ash samples were tested is given in Table 3.2, which also includes a summary of the experiments carried out for each sample.

Fly Ash Sample

Tallawarra Eraring

Munmorah Vales Point Gladstone Wallerawang

Liddell

Duration of Test Work

25-10-83 to 31-01-84 03-02-84 to 15-02-84 13-03-84 to 02-04-84 09-08-84 to 19-09-84 03-10-84 to 11-10-84 26-11-84 to 18-01-85 26-03-85 to 22-07-85

Exp. NOS.

160 to 193 220 to 251 321 to 384 520 to 546 661 to 685 740 to 766 830 to 865

Total No.of Useful Tests

32 30 61 27 22 24 35

NO. Of Data Pts.

58 70 88 52 41 44 57

Table 3.2 Chronology of the fly ash test program.

40

Important information relevant to the fly ash test program is summarised below.

(a) The Gladstone fly ash referred to in this table was tested for the Queensland Electricity Generating Board (now the Queensland Electricity Commission) and has been included for an additional comparison.

(b) The total number of useful tests listed against each sample refers to the actual total number of pneumatic conveying experiments, from which data were extracted (i.e. mt, m s and Apt). Therefore, the experiments which either had to be rejected due to faulty equipment (e.g. transducers) or were concerned only with air flow/pressure drop measurements (e.g. orifice plate checks, empty pipeline air pressure drop data) have not been included in this total.

(c) No serious operational problems occurred during the fly ash program except for the Vales Point fly ash sample, which was found to create excessive blinding of the original polypropylene filter bag surfaces. The shaking/cleaning operation of the filter unit was not able to remove the caked-on fly ash material. Subsequent over-pressurisation occurred in the receiving hopper and caused large amounts of dust to be released to the atmosphere (from between the joining surfaces of the filter housing and the access panel). The polypropylene bags were replaced with epitropic Goretex™ bags and the over-pressurisation problem was eliminated. Although such problems were not repeated for the Gladstone, Wallerawang and Liddell samples, it is possible that they still could have occurred with the original polypropylene bags.

The pneumatic conveying characteristics of the Eraring fly ash are presented in Figure 3.9, which includes the actual values of m s obtained from Exp. Nos. 220 to 251. Note that for reasons of clarity not all data points have been shown on this plot. This form of representation is similar to the one presented by Mason et al. [11], except for the following two extensions. 1. The range of mf is selected to span the regions of both the dilute- and

dense-phase modes (i.e. suspension and non-suspension flow regimes). 2. Regions of unreliable transport (e.g. unstable duning, plugging,

blockages) are investigated and delineated on the conveying characteristics. Certain experiments are designed actually to approach the minimum transport condition. This task is made feasible with the unblocking technique described in Section 2.2.

However, for the purpose of providing clarity in reports and publications, the ms values are omitted usually from the conveying characteristics. Hence, the final form of presentation is obtained by reproducing the lines of constant ms, as shown in Figure 3.10. Refer to Figures 3.11 to 3.16 for the pneumatic conveying characteristics of the Tallawarra, Munmorah, Vales Point, Gladstone, Wallerawang and Liddell fly ash samples.

41

o

O

in

o

o

i to CJ1

co o

CM O

o in

o o

o m

•<-> CO

D. Q.

200

Apt

(kPa)

100

T +

Blockage Condition

Unstable /

0 .02 .04

raf (kgs-1)

.06 .08

Figure 3.10 Pipeline conveying characteristics of Eraring fly ash for L = 71 m & D = 52 m m (Test Rig B1).

200

Apt

(kPa)

100

0 L 0

OO-v i

I i

©• I

Blockage Condition

Unstable Duning

(kgs"1)

Air OnV

.02 .04 .06

m.f (kgs"1)

.08

Figure 3.11 Pipeline conveying characteristics of Tallawarra fly ash for L = 71 m & D = 52 m m (Test Rig B1).

200

Apt

(kPa)

100

0

1

i Blockage Condition

Unstable Duning

Air Only.

.02 .04

mf (kgs-1)

.06 .08

Figure 3.12 Pipeline conveying characteristics of Munmorah fly ash for L = 71 m & D = 52 m m (Test Rig B1).

200

Apt

(kPa)

iBlockage Condition

Ûnstable Duning

100

0 A Air OnT:

.02 .04

mf (kgs-1)

.06 .08

Figure 3.13 Pipeline conveying characteristics of Vales Point fly ash for L = 71 m & D = 52 m m (Test Rig B1).

200

Apt

(kPa)

100

o Unstable

0 .02 .04

mf (kgs-1)

.06 .08

Figure 3.14 Pipeline conveying characteristics of Gladstone fly ash for L = 71 m & D = 52 m m (Test Rig B1).

200

Apt

(kPa)

100 -

.02 .04

mf (kgs-1)

.06 .08

Figure 3.15 Pipeline conveying characteristics of Wallerawang fly ash for L = 71 m & D = 52 m m (Test Rig B1).

45

200

Apt

(kPa)

100

0 0 .02 .04 .06 .08

mf (kgs"1)

Figure 3.16 Pipeline conveying characteristics of Liddell fly ash for L = 71 m & D = 52 m m (Test Rig B1).

3.4 Standardised-Test Procedure

During the initial stages of the fly ash testing program, it was realised that it would be advantageous to standardise experimental procedures for the purpose of

minimising the total number of necessary pneumatic conveying experiments, providing sufficient data for the representation of complete conveying characteristics, investigating minimum transport behaviour (to determine the extent of dense-phase suitability), and defining undesirable operational problems such as plugging, blockages or unstable transport.

A standardised-test procedure was developed and also generalised to apply to different types of materials, as reported by Wypych and Arnold [21], In conjunction with suggested scale-up procedures, which are discussed in Section 3.5, this procedure provides a basis for reliable pneumatic conveying design.

The results presented previously on Eraring fly ash (i.e. in Section 3.3) and also PVC powder [21] are employed to demonstrate the standardised-test procedure and the determination of minimum transport boundaries. The 71 m x 52 m m test rig (i.e. Test Rig B1) described in Sections 2.2 and 3.3.2 is applicable to all results presented in the following sections.

46

3.4.1 Experiments

Three different types of pneumatic conveying experiment (referred to as Test 1, 2 and 3) were developed and designed to provide efficiently the data necessary for the representation of conveying characteristics. Note that the transient plots of the following major conveying parameters are presented to demonstrate the differences between each test.

Blow tank top-air pressure, pDt (referred to as PM , for transducer location A1, as shown in Figure 3.8). Pipeline air pressure, PG2 (« 17.8 m from the blow tank outlet). Mass of solids conveyed, Ms. Supplied air mass flow rate, mf.

3.4.1.1 Tesf 1 - Standard Batch Cycle

The blow tank is pressurised to an initial steady-state value. The conveying-air and discharge valves are opened, and the conveying parameters are recorded for the duration of the cycle. This experiment obtains usually one steady-state operating condition, which represents one conveying characteristic coordinate location. A typical example of a standard batch cycle is presented in Figure 3.17 (Eraring fly ash, Test Rig B1, Exp. No. 236). Note that for this experiment

steady-state m* = ms mf-1

2.42 kg s_1

0.045 kg s"1

= 54 kg kg"1

whereas average m* = Total mass of solids conveyed

Total mass of air used

390 kg 10.2 kg

= 38 kg kg-1,

which is 30% lower than the steady-state value. Note that the 10.2 kg of air was determined by calculating the amount of air used during the conveying cycle (viz. 9.4 kg, which is equal to the area under the mf curve) and adding on the amount of air required for initial pressurisation of the blow tank (viz. ~ 0.8 kg of air to provide Pbt.i = 170 kPag). Note that such differences between the steady-state and average values of m* are typical of the standard batch cycle and should be allowed for when designing a pneumatic conveying system.

47

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3.4.1.2 Test 2 - Increase of Apt for Approximately Constant mf

The blow tank is pressurised initially to a relatively low steady-state value. The discharge and conveying-air isolation valves are opened, and the conveying parameters are allowed to reach a steady-state condition. The blow tank air pressure then is increased by opening-up the regulators to a predetermined value. The operating conditions once again are allowed to reach a steady-state condition and then the procedure is repeated until the completion of the cycle. For certain materials it is possible to obtain up to six coordinate locations. Refer to Figure 3.18 (Eraring fly ash, Exp. No. 240) for typical transient plots of the major conveying parameters obtained using this type of test. 3.4.1.3 Test 3 - Decrease of mf at Steady-State Conditions

The blow tank is pressurised to an initial steady-state value (similar to Test 1). The discharge and conveying-air valves are opened, and the operating conditions are allowed to reach a steady-state value. The conveying air mass flow rate is reduced gradually by a predetermined amount using a calibrated flow control valve. The conveying parameters are allowed to reach a steady-state condition before the air flow rate is reduced any further. This type of test usually produces up to three coordinate locations and an example is presented in Figure 3.19 (Eraring fly ash, Exp. No. 249). 3.4.2 Results

The following technique then is used to obtain the three steady-state conveying parameters required for the presentation of the pipeline conveying characteristics.

1. Establish a steady-state region during the conveying cycle taking into account all parameters (i.e. blow tank and pipeline air pressure, air flow rate and supplied/delivered mass flow rate of solids).

1. mf: read the value directly from the transient plot, as shown in Figures 3.17(d), 3.18(d) and 3.19(d).

2. ms : determine the slope of the delivered mass of solids curve, as shown in Figures 3.17(c), 3.18(c) and 3.19(c).

3. Apt : extrapolate the pipeline air pressure drop curve back to the blow tank outlet (see Figure 3.20). However, in some blow tank installations, where the conveying air pipe has the same internal diameter as the actual conveying pipeline, Apt may be approximated by the gauge pressure measured upstream of the blow tank outlet (referred to as the conveying-air pressure or back-pressure). All Apt results presented in this section were determined using the extrapolation technique.

The test procedure described in Section 3.4.1 was employed to obtain sufficient data for the representation of pipeline conveying characteristics (Eraring fly ash and P V C powder). The resulting groups of experiments and data points (noting that one data point represents one value of mf, ms, and Apt, as determined from the technique described above) are summarised in Table 3.3.

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No. of Test 1

14 4

No. of Test 2

6 13

No. of Test 3

10 15

Total No. Exps.

30 32

Total No. of Data Points

67 63

Table 3.3 Summary of experiments and data points for Eraring fly ash.

To demonstrate graphically the standardised-test procedure (and the differences between Tests 1, 2 and 3), the mf, ms, and Apt results which were obtained from Exp. Nos. 236, 240 and 249 ( refer to Figures 3.18, 3.19 and 3.20) have been reproduced in Table 3.4 and superimposed on the pipeline conveying characteristics, repeated in Figure 3.21.

Note that in Exp. No. 249, unstable duning occurred at mf » 0.0130 kgs~1 and that the corresponding values of Apt and m s are approximations averaged over the steady-state period.

Experiment Number

236

240

249

mf (kg s-1)

0.0450

0.0395 0.0407 0.0415 0.0425 0.0435

0.0252 0.0130

Apt (kPa)

121

32 56 64 78 87

109 =106

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2.42

0.34 0.76 0.94 1.31 1.53

2.30 =2.10

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54

9 19 23 31 35

91 =162

Table 3.4 Steady-state operating conditions obtained from Exp. Nos. 236, 240 and 249.

53

Blockage Condition

0 .02 .04 .06 .08

mf (kgs"1)

Figure 3.21 Pipeline conveying characteristics of Eraring fly ash for L = 71 m & D = 52 m m (Test Rig B1) demonstrating Tests 1, 2 and 3..

3.4.3 Minimum Transport Behaviour

The large differences in flow performance and efficiency, that were found to occur in a variety of materials by Lohrmann and Marcus [22] and Wypych and Arnold [16], emphasise the need to investigate minimum transport behaviour. For the case of Eraring fly a£h (and the other fly ash samples considered in this thesis), considerable difficulty was experienced in obtaining a well-defined locus of blockage conditions. However, a broad region of unstable duning was observed and recorded. Transient plots of the major conveying parameters obtained from Exp. No. 232 (Test 3), which produced the only well-defined blockage condition, are presented in Figure 3.22. Note the increase in pipeline air pressure fluctuation as mt was decreased (i.e. as the reliable transport limit was approached). However, the steady-state section of the M s curve was considered still as a stable response and, hence, the corresponding operating conditions were not recorded as an unstable duning coordinate. A total of three experiments (including Exp. No. 249) provided unstable duning phenomena similar to that shown in Figure 3.19. From these data, an approximate reliable transport boundary was estimated and plotted on the pipeline conveying characteristics of Figures 3.9, 3.10 and 3.21.

54

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In contrast, P V C powder (dp50 = 135 Lim, ps = 1400 kg m-3, pbi = 575 kg m-3), which was conveyed on the same test rig (i.e. Test Rig B1), displayed a well-defined minimum transport condition [21], as shown on the pipeline conveying characteristics presented in Figure 3.23. At the onset of blockage, severe pipe vibrations occurred. However, most of the blockages only were temporary and a restart in the conveying cycle usually was possible. Both Tests 2 and 3 were used to estimate the locus of blockage/plugging conditions, and examples are presented in Figures 3.24 and 3.25 (PVC powder, Exp. Nos. 387 and 414). Note that any evidence of imminent blockage (e.g. unstable duning for fly ash) was not observed for the P V C powder. 3.4.4 Test Procedure Applications and Limitations

Although the total number of experiments and corresponding data points required for the presentation on conveying characteristics were similar for Eraring fly ash and P V C powder, considerable differences occurred between the number of actual experiments (i.e. Tests 1, 2 and 3). This was due mainly to the limitations and particular applications of each test to provide data over the (maximum) regions

0<mf < 0.10 kg s"1

and 0 < Apt < 200 kPa.

Note these values are relevant only to the test rig that was employed (i.e. Test Rig B1) and will vary for other blow tank/pipeline configurations and air supply compressors. The necessary number of individual experiments (i.e. for Tests 1, 2 and 3), as indicated in Table 3.3, will depend also on the physical properties and the minimum transport behaviour of the material.

Test 1 is used usually to achieve one steady-state operating condition at the upper end of the pressure, drop scale (typically, for Apt > 100 kPa).

Test 2 is employed mainly to minimise the total number of experiments required to generate sufficient data for the presentation of conveying characteristics. Up to ~ six steady-state operating conditions are possible (i.e. for Test Rig B1). The actual number depends on the response rate of the material subjected to a given change in operating conditions, the batch size of the blow tank and the pipeline configuration. Test 2 is used also to investigate minimum transport behaviour and locate plugging/blockage boundaries (e.g. refer to the results presented for P V C powder). W h e n approaching this type of boundary for the purpose of delineation, it is easier and more accurate to use Test 2 than Test 1. The latter requires pbt.i values in the blockage region to produce the necessary steady-state coordinate location just outside the unstable zone. In fact, when attempting to use Test 1 to locate the blockage conditions of the P V C powder, severe plugging occurred immediately after the commencement of the conveying cycle, preventing any accurate estimation of the corresponding operating condition. However, for some materials, Test 2 will be restricted to an upper limit of pressure drop during the conveying cycle. This will depend mainly on the physical nature of the material being conveyed (e.g. permeability, compressibility) and the flow restrictions imposed by the blow tank air supply lines (viz. top- and fluidising ring-air). For the test rig, the Eraring fly ash sample and the P V C powder considered in this section, the following Apt limits were observed (i.e. for Test 2).

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100

50

0

1 f

Blockage ' Conditions

Blockage Boundary N.

,02 .04 JL

06 m.

.08

(kgs-1)

(kgs-1)

.10 12

Figure 3.23 Pipeline conveying characteristics of PVC powder [21] for L = 71 m & D = 52 m m (Test Rig B1).

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Eraring fly ash : Maximum Apt« 100 kPa. P V C powder : Maximum Apt « 200 kPa.

The 100 kPa limitation for Eraring fly ash also was observed for the other fly ash samples (i.e. considered in Section 3.3). Note that for the P V C powder and Apt > 150 kPa, only three steady-state operating conditions were able to achieved.

Test 3 is used also to minimise the total number of experiments, although not as many operating conditions can be achieved (i.e. when compared with Test 2). However, its main use is to supplement the unreliable transport information that cannot be obtained by Test 2 due to the pressure drop limitations mentioned previously. Furthermore, for unstable transport boundaries, which are either difficult to locate or relatively insensitive to changes in mf (e.g. compare Figures 3.10 and 3.23), Test 3 is found to be very useful. In fact, the ten experiments of Test 3 conducted on the Eraring fly ash were used mainly for this purpose. The final selection of the actual number of experiments (i.e. Test 1, 2 or 3) required for the standardised-test procedure depends on the test rig configuration and the physical nature and minimum transport behaviour of the material in question. However, upon the commencement of experimentation, it soon becomes evident which tests will predominate the test program. For example, the Eraring fly ash and the P V C powder displayed different conveying characteristics and minimum transport behaviour, which was reflected by the different number of Tests 1, 2 and 3. With the assistance of the powder classification results reported by Wypych and Arnold [16], it may be suggested generally and qualitatively that for Dixon [23] Group A and some Group C powders (e.g. fly ash), the

individual number of tests will be similar to that of the Eraring fly ash, although certain unpredictable phenomena may require additional experimentation (e.g. refer to the unusual shape of the reliable transport limit displayed in Figure 3.12),

Group B and some Group D materials, the number will be similar to that of P V C powder [21], where four experiments of Test 1, thirteen of Test 2 and fifteen of Test 3 were required.

These matters of powder classification are discussed further in Chapter 5.

60

CHAPTER 4

61

4. BLOW TANK CONFIGURATION & AIR INJECTION

There are many interacting considerations involved in the design of a pneumatic conveying system. For example, the estimation of a suitable pipeline configuration and corresponding air supply is required usually to fulfil a given specification. An equally important consideration is the selection of the most suitable mode of conveying (e.g. dilute-, dense- or plug-phase flow), feeder configuration (e.g. top- or bottom-discharge blow tank) and the method of air injection (e.g. aeration nozzles, fluidising membrane). This will depend on the given specification and the behavioural properties of the material. Only a limited number of publications have considered such matters and unfortunately, these have been quite general and lacking useful detail or results (i.e. for the designer or potential user of pneumatic conveying equipment). For example, Flain [4] has provided a general overview of several commercial blow tank systems and also has indicated their field of application. More recently, Klintworth and Marcus [2] have reviewed various commercial low-velocity pneumatic conveying systems, with particular emphasis on discontinuous dense-phase and specialised designs of pipeline. This section of work investigates the effect of blow tank design and the method of air injection on the pneumatic conveying performance of bulk solids and is aimed at emphasising the importance of matching both the method of air injection and the design of blow tank to the material (and its behavioural properties), as well as demonstrating how the overall performance of a pneumatic conveying system may be improved by incorporating a more suitable method of air injection. Three case studies involving the dense-phase transport of pulverised coal, the long-distance conveying of fly ash and the plug-phase transportation of screened and unscreened granulated aluminate (SGA and U G A ) , bone char and crushed bath are discussed in some detail for this purpose. Table 4.1 provides a list of the relevant physical properties for each material and the test rigs that were used to obtain the experimental data. Materia!

Pulverised Coal Fly Ash SGA UGA

Bone Char Crushed Bath

dso (Jim)

30 10

2500 2150 600 3900

ps (kg m"3)

1600 2240 2320 2350 3500 3080

Pbi (kg nr3)

760 522 1000 1520 1090 645

Test Rlg(s)

A D1&D2 F2 F2 F1 F3

Table 4.1 Physical properties of test materials.

Note that a Malvern Model 2600C laser diffraction particle sizer was employed to obtain the particle size distributions of the pulverised coal and fly ash, whereas a sieve analysis was used for the other products. The solids density, ps, was determined by a Beckman Model 930 air pycnometer and the loose-poured bulk density, pDi, by pouring gently a known mass of material into a measuring cylinder.

62

4.1 Pulverised Coal

Several groups of experiments were carried out on the original Test Rig A1 (see Figures 2.1 and 2.2) to investigate the relative effect of blow tank top-air and supplementary conveying-air on the dense-phase pneumatic conveying characteristics of the pulverised coal. Each test group consisted of three experiments, which were subjected to the same set-up conditions (e.g. air flow, pbt.i) but with one of the following different combinations of air injection. (a) Blow tank top-air only.

(b) Blow tank top-air and supplementary-air (solenoid valve removed). (c) Blow tank top-air and pulsed supplementary-air (using a solenoid

valve controlled by an electrical timer). A more detailed arrangement of the various air supply lines is shown schematically in Figure 4.1. Note both probe- and ring-air were not used for these experiments

Material Inlet

Conveying *" Pipeline

Solenoid Valve

Discharge Valve

Figure 4.1 0.425 m 3 Sturtevant™ blow tank and air supply arrangement.

and the supplementary-air line simply consisted of a 12 mm O.D. nylon tube connecting the top-air line to the blow tank outlet via a timer-operated solenoid control valve (for combination (c) only) and a non-return valve. For combination (b), the solenoid valve was removed completely from the supplementary-air line. By operating the isolation valves shown in Figure 4.1, the above three methods of air injection were compared for five different values of pt>t,i (viz. 115,165, 200, 240 and 270 kPag). Transient responses of the blow tank top-air pressure (i.e. pDt = PA1, transducer location A1, see Figure 3.8), a typical pipeline air pressure ( P G L see Figure 3.8) the delivered mass of solids (Ms) and the supplied air mass flow rate (mf) were obtained from each experiment.

63

A typical set of results from three experiments are superimposed in Figure 4.2, and the relevant set-up conditions are summarised in Table 4.2. Note that

in Exp. No. 35, a 0.5 s ON / 0.5 s OFF setting was used for the timer-operated solenoid valve,

additional experiments, which were carried out to determine the effectiveness of knife-air (see Figures 2.1, 2.2 and [8]), have been included in Table 4.2,

combination (d) consisted of blow tank top-air and knife air (with a timer setting of 0.5 s O N / 0.5 s OFF),

the results obtained from Exp. Nos. 61 & 62 are presented in Figure 4.3.

Exp. No.

21 23 35

61 62

Method of Air Injection

(a) (b) (c)

la) (d)

PA1,I (kPag)

115 115 115

185 185

Blow Tank Top-Air

Yes Yes Yes

Yes Yes

Supplementary-Air

No Yes

Yes (pulsed)

No No

Knife-Air

No No No

No Yes

Table 4.2 Set-up conditions for the blow tank air injection experiments.

The following steady-state conveying parameters also were determined from Exp. No. 23 (i.e. the combination (b) results).

ms = 2.19kgs'1=7.9th-"'.

mf = 0.0076 kg s'1 = 22.8 m 3 h"*" free air (i.e. @ 20°C & 1010 hPa).

Vf = 1.74 m s"1 @ pipe inlet & 2.98 m s-1 @ pipe exit (where D = 52 mm).

m* = (2.19) (0.0076)"1 = 288 kg kg-1.

Apt = 72kPa.

From these results, the following observations are made for pulverised coal.

For low air flows, top-air only seems to compact the material inside the blow tank resulting in a more unstable conveying mode (e.g. a temporary blockage condition occurred during Exp. No. 21, whereas reliable transport was achieved in Exp. No. 61).

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^ TRANSIENT CYCLE TIME (SECS) Figure 4.3 Transient plots of major conveying parameters from Exp. Nos.

61 and 62 for pulverised coal conveyed over 25 m (Test Rig A1).

66

The introduction of supplementary-air has an overall smoothing effect on conveying characteristics and, hence, allows the selection of larger m* values. For example, from Figure 4.2 and the previous set of calculations, it can be seen that top-air only is unable to achieve reliably the combination (b) result of m* = 288. One reason for this improvement in performance may be due to the equalising effect that combination (b) has on the air pressure acting across the bed of material inside the blow tank, hence tending to minimise the compaction effect described in the previous observation.

The addition of a timer-operated solenoid control valve to the supplementary-air line (i.e. combination (c)) produces conveying results between those of combination (a) and (b). In fact, additional experiments were carried out to determine the effect of varying the O N / O F F timer setting (i.e. for the solenoid control valve). The results obtained with Min. O N / Max. O F F were found to be similar to combination (a) and those with Max. O N / Min. O F F similar to combination (b). Note that the minimum timer setting represented « 0.02 s and the maximum « 1.0 s.

The introduction of knife-air has a relatively insignificant effect upon conveying performance (e.g. the average values of m s for Exp. Nos. 61 and 62 are 3.7 and 3.9 kg s_1, respectively) and also tends to create a non-linear M s response curve (see Figure 4.3(b)).

The supplementary-air line assists in the fluidisation of the product prior to transportation (i.e. during the initial pressurisation cycle of the blow tank). This is considered to be an important requirement for the dense-phase conveying of fine powders. Also, the presence of a supplementary-air supply at the blow tank outlet seems to provide an additional feature of mixing the product during transportation (i.e. a region of turbulence at the entrance of the pipe).

The above results further indicate that accurate control of both pressure and flow rate between the blow tank top- and supplementary-air supply lines, would provide a fine-tuning effect for solids flow rate control. For example, relatively large turndown ratios of m s are possible by simply varying the blow tank top-air pressure, as shown by Test 2 of the standardised-test procedure (refer to Section 3.4.1.2 and Figure 3.18). In conjunction with a solids metering device (such as the KBI G m b H blow tank cone-dosing-valve used on Test Rig E1, as shown in Figure 2.10), such innovations could provide an effective and efficient technique for the control of solids flow rate. In fact, KBI G m b H recently has marketed a proportioning valve, which monitors and controls the amount of air being distributed to both the conveying pipeline and blow tank. 4.2 Fly Ash

4.2.1 Introduction

Certain materials including fly ash, which were tested initially on the original Test Rig C, have been found to cause serious rat-holing problems inside the blow tank, similar to that shown in Figure 4.4. The following possible contributing factors have been established.

67

Aeration Air

Dead Region

Discharge

Figure 4.4 Configuration of bottom-discharge blow tank demonstrating incomplete discharge of material due to rat-holing.

Inappropriate blow tank geometry (e.g. critical hopper angle and outlet diameter) for a particular material. Material cohesion (as well as wet or sticky properties) producing poor fluidisation and, hence, bad-channelling (or rat-holing) during the pressurisation and conveying cycles. Strong adhesion between the material and blow tank wall. The method of air injection causing localised penetration, as shown in Figure 4.4 (i.e. the aeration air will follow the line(s) of least flow resistance).

Similar problems also could occur easily in top-discharge blow tanks, as indicated in Figure 4.5 (especially for products which are fine and heavy). From a practical viewpoint, the consequences of such performance are

the effective working capacity of the blow tank is reduced due to incomplete discharge of the contents, thereby lowering the overall solids throughput of the system, inconsistent and, possibly, unstable transport could result from the nonuniform fluidisation of material in the vicinity of the conveying pipe entrance (for top-discharge blow tanks).

To overcome some of the above problems in relation to the original KBI blow tank (Test Rig C, see Figure 2.7), the design of the second blow tank unit (i.e. for Test Rig D) was improved in collaboration with N.E.I. John Thompson (Aust.).

68 Discharge

Porous Membrane Plenum Chamber

Dead Region

Air Inlet

Figure 4.5 Configuration of top-discharge blow tank demonstrating incomplete discharge of material due bad channelling and rat-holing.

That is, a fluidising discharge cone was fitted to the bottom of the vessel and high-flow evassers (instead of sintered bronze nozzles) were installed at the end of each remaining aeration line. Note the evassers make use of a rubber boot that tightly covers a series of aeration holes drilled into a hollow, spherical steel bulb, which is welded to a short section of threaded pipe (for connection to the aeration line). In addition to fluidising the product and pressurising the blow tank, these evassers are intended to promote material discharge from the vessel by washing down its inside surface and also, to function as a non-return valve when the air supply is turned off. The following section presents results obtained from investigations into comparing these two different methods of blow tank air injection (i.e. refer to Test Rig D, Figure 2.9). 4.2.2 Test Results

Test Rig D 2 (L = 940 m & D = 60/69/81/105 m m ) was set up initially with the intention of simulating continuous long-distance transportation and hence, obtaining conveying characteristics of the fly ash sample (refer to Table 4.1) under these conditions. This required the establishment of reliable steady-state conditions during the consecutive discharge of the two blow tanks. Note that as the receiving silo in Test Rig D 2 was not located directly above the tandem blow tanks (to enable gravity filling), this test facility was unable to provide continuous conveying. To provide a direct comparison, the top- and aeration-air supplies to each blow tank were set up in the same manner (i.e. in terms of distribution, pressure and flow rate) and the conveying-air was not altered at any stage during the experiment.

69

Figure 4.6 presents typical solids mass flow rate data obtained from one of the several tests carried out for this purpose. Note for this experiment, Blow Tank No. 2 (with the fluidising cone) was discharged first, followed by Blow Tank No. 1 (with the aeration nozzles only) and the air flow was set at mf» 0.2 kg s-1. It can be seen that there is a considerable difference in solids discharge performance between the two blow tanks. This is reflected also in the mass flow rate of solids into the silo. Note that similar differences in discharge characteristics were observed when Blow Tank No. 1 was discharged before Blow Tank No. 2 (i.e. when the operating sequence was reversed). As a result of observing this behaviour, a further series of tests were conducted on the same fly ash sample to examine in more detail the difference between the two methods of air injection. Test Rig D1 (L = 293 m & D =69 m m ) was used for this purpose, so that steady-state conveying conditions could be established with the discharge of only one 0.9 m 3 blow tank. In this way, the two blow tanks could be run independently and exposed to the same set-up conditions. A series of tests were carried out on Test Rig D1 covering a wide range of mf values. Figure 4.7 superimposes typical transient plots of major conveying parameters obtained from two consecutive tests where Blow Tanks No. 1 and 2 were used alternately. The steady-state and average conveying parameters listed in Table 4.3 have been determined from these experiments.

Blow Tank

1 2

m s

(kg s-1)

1.85 1.85

mf

(kg s-1)

0.25 0.25

m* (kg kg-1)

7.4 7.4

Apt (kPa)

242 240

Disch. Time (s)

365 288

Avg. m s

(kg s-1)

1.47 1.77

Table 4.3 Conveying parameters of fly ash for L 293 m & D = 69 m m (Test Rig D1).

For these experiments, the supply air (aeration and conveying air) was held constant with the top-air not being used. Note that the total amount of air injected into the blow tanks (via evasser nozzles for Blow Tank No. 1 and the combined fluidising discharge cone and evasser nozzles for Blow Tank No. 2) effectively was equal for both tests. It is apparent from these results that the steady-state conveying parameters were very similar for these two experiments. However, from Figure 4.7, the m s response for Blow Tank No. 1 is seen to decline at a relatively early stage of the conveying cycle. For example, this occurred over the final 40 % of the discharge cycle of Blow Tank No. 1 (affecting « 210 kg of the initial batch size of 530 kg), whereas for Blow Tank No. 2 this occurred over the final 8 % (affecting only the final 40 kg of the initial 510 kg batch). An additional comparison is made by determining the average conveying rate for each blow tank. For example, for Blow Tank No. 1, average m s = (530 kg) (365 s)"1 = 1.47 kg S'1 which is « 17 % lower than the 1.77 kg S'1

calculated for Blow Tank No. 2. The following observations are made as a result of these experiments.

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BLOH THNK S

Figure 4.6 Blow tank comparison using fly ash and Test Rig D2 (L = 940 m & D = 60/69/81/105 m m ) .

For a fly ash of this type (i.e. fine and cohesive, refer to the Geldart [24,25] Group C classification), the fluidising discharge cone is a significant improvement over the aeration nozzles in terms of maintaining a steady discharge of material from a bottom-discharge blow tank into the conveying pipeline. This is evident particularly over the final portion of the conveying cycle and is believed to be achieved by providing a more even fluidisation of the material inside the

blow tank, having an effect similar to flow promotion aids in silos (i.e. effectively expanding the active outlet diameter of the blow tank).

This improved method of aeration has the potential of reducing significantly the discharge time required by a blow tank (e.g. 15 to 25 % in the experiments carried out on Test Rig D1) and hence, is important in terms of the overall conveying capacity as well as the air/energy requirements of the system.

The results presented here on fly ash and previously on pulverised coal (i.e. in Section 4.1) also suggest that the fluidising discharge cone may assist in attaining conveying conditions with a higher m* value than may be achieved reliably with an aeration nozzle type of blow tank (especially for fine, heavy and/or cohesive products, which are difficult to fluidise).

71

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4.3 Plug-Phase Conveying

Over the past two decades, dense-phase [2,7,14,17], low-velocity [2,3], longdistance [1,5,6] and also dilute-phase [7] pneumatic conveying have received considerable attention from both researchers and commercial suppliers of equipment. In contrast, discrete plug-phase conveying (i.e. the transportation of a limited plug length of material per cycle) has received little attention except for a few general papers/reports published or presented by vendors of such equipment. Yet this method of transport is able to handle efficiently a large range of conventionally difficult materials (e.g. coarse, heavy, wide size range, abrasive and even friable), that otherwise could be considered only for dilute-phase transport. For example, refer to the case study on screened coke presented by Wypych and Arnold [26]. In fact, even specialised techniques such as low-velocity and/or by-pass technology [2] may not be suitable for such materials. This aspect of powder classification (i.e. in relation to determining the most suitable mode of conveying for a given product) is considered further in Chapter 5. The results obtained from four materials (SGA, UGA, bone char and crushed bath, refer to Table 4.1 for relevant physical properties) are presented to demonstrate (a) the general features and full potential of this method of transport, (b) the effect of blow tank air injection on plug-phase conveying

characteristics, and (c) how considerable improvements in performance (e.g. reliability, reduced

degradation) can be obtained simply by modifying the method of air injection.

4.3.1 Screened & Unscreened Granulated Aluminate (SGA & UGA) Using Test Rig C2 (0.9 m3 blow tank, L = 59 m & D = 105 mm) several experiments were carried out initially on S G A to establish reliable and optimal operating conditions for the dilute- and/or dense-phase mode of transportation. However, severe pipeline blockages occurred for air flows below mf« 0.3 kg s_1 (e.g. for pbtj * 140 kPag) and it was found that conventional dense-phase was not possible for this material. In fact, the following steady-state operating conditions were found reliable for dilute-phase conveying only. mf = 0.325 kg s"1, ms = 6.8 kg S'

1 = 24.51 rr1, Apt = 85 kPa, Vf = 17 m S'1

to 31 m s_1 (i.e. from start to end of the pipeline). On inspection of the material, which was conveyed under these conditions, an excessive amount of fines was noticed in the sample and considered unacceptable for the proposed application (i.e. for a liquor burning impurity removal process, which could not tolerate large amounts of dust). For this reason, additional tests were carried out on Test Rig F2 (0.113 m 3 plug-phase blow tank, L = 58 m & D = 105 m m ) to determine whether lower transport velocities could be achieved (i.e. to minimise degradation). After carrying out 14 experiments at different values of mf, the following conveying parameters were found most reliable for the plug-phase mode. Orifice-air only (see Figure 2.12), mf« 0.079 kg S"1, Avg. ms = 2.5 kg s"

1 = 9.01 h"1, Max. pbt - 200 kPag, Avg. vs « 1.8 m s"

1 (where the time taken for the material to reach the end of the pipeline was « 33 s).

73

Note that

steady-state operating conditions are not applicable to this method of transport, these conveying parameters were based on Exp. No. 1274 (refer to Figure 4.8 for transient plots of the major parameters), the average conveying rate (i.e. avg. m s ) is determined by dividing the mass of product conveyed (viz. 102 kg) by the conveying cycle time (viz 41 s), the mass of solids conveyed was obtained by actually removing the material from the 5 m 3 receiving silo and weighing it on a load platform (refer to the eccentric loading problems described in Section 2.6), the material which was conveyed after Exp. No. 1274 was retained for later inspection and found to contain only minor levels of dust (acceptable for the liquor burning impurity removal process), to determine an overall conveying rate (i.e. for a proposed installation) the average conveying rate avg. m s must be modified to allow for cycle overheads (e.g. filling time, valve switching time). However, this is not required for the purpose of the present investigations.

The type of conveying produced by the 0.113 m3 blow tank (i.e. for Test Rig F, see Figures 2.12 and 2.13) leaves a significant portion of material in the pipeline especially when the latter had been emptied or purged prior to an experiment (e g' after a planned shutdown). This is caused by the relatively low values of conveying velocity used for operation (typically, avg. vs = 1 to 3 m s"

1, which is based simply on the time taken for the material to reach the end of the pipeline and also the assumption that the product commences to move into the pipeline shortly after the air supply is turned on). However, there exists a minimum flow rate condition, below which an excessive amount of material will remain in the pipeline and gradually build up until either unstable, strong plugging or a blockage will occur. For the S G A tested in these investigations, mf - 0.07 kg s'

1 represented this minimum air flow condition. Hence, the above operating conditions are considered quite safe and reliable, although for an actual system, mf > 0.085 kg s"1 may have to considered. As both SGA and UGA were required to be conveyed pneumatically (and also with low levels of degradation), it was considered necessary to undertake additional experiments on U G A for the purpose of determining reliable operating conditions, and

establishing whether an increased level of fines would have any effect on the conveying performance of U G A (e.g. refer to particle size distributions presented in Figure 4.9).

Subjecting UGA to the operating conditions, which were found reliable for SGA (Le. orifice-air only, mf * 0.08 kg s-1), a stable plug formed during transportation and this produced substantial pipe vibrations and much higher pipeline air pressures (e.g. Pbt > 300 kPag). Although U G A was conveyed successfully under these conditions, the method of transport was considered too damaging to the particles and it was decided to pursue the requirements of a more gentle duning mode of conveying. Note it is believed that the majority of such degradation occurs during the final stages of the cycle, where the stable plug of product is released suddenly from the end of the pipeline (due to a high upstream air pressure, usually referred to as back-pressure).

EXPERIMENT NO. 1274 TEST DBTE I 20.5.1988

200.

cc cu

UJ

cc 103.

UJ

cc CL

20. 30. 40.

CYCLE TIHE (SECS)

^ i I 1 1 i i i

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EXPERIMENT NO. 1274 TEST DATE « 20.5.1988

200.

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20. I i I i • I I I I l _

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EXPERIMENT NO. 1274 TEST DRTE l 28.5.1988 TOTAL HflSS OF RIR USED (KGS) - 2.719

.09

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ORIFICE PLATE NO. 5

Figure 4.8 Transient plots of major conveying parameters for SGA (Exp. No. 1274, Test Rig F2).

75

100

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Screened Product

Unscreened Product

J I T ^ i I

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Particle Diameter (mm)

10.0

Figure 4.9 Particle size distributions of S G A and UGA.

Selecting orifice- and ring-air (see Figure 2.12) and carrying out an additional series of tests, the minimum flow rate condition described previously for S G A was found to occur at mf« 0.076 kg s_1. At or above this value, the discharge of product from the end of the pipeline occurred more in the form of a dune (i.e. instead of a stable plug), which was seen to be more gentle and less damaging to the particles (note, by visual inspection only). Also, the resulting operating pressures were found to be considerably less (e.g. pbt = 200 to 250 kPag). Typical transient plots of major conveying parameters are presented in Figure 4.10. Note that this experiment was operated in the vicinity of the minimum flow rate condition. The following conveying parameters were found to be most reliable for the plug-phase conveying of UGA. Orifice- and ring-air (see Figure 2.12), mf « 0.085 kg S"1, Avg. ms = 2.40

kg s'1 = 8.61 fr1, Max. pbt« 250 kPag, Avg. vs « 1.7 m S"1 (where the time

taken for the material to reach the end of the pipeline was « 35 s).

EXPERIMENT NO. 13S6 TEST DATE i 12.8.1986

300.

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cc 3

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EXPERIMENT NO. 1356 TEST ORTE 1 12.8.1988

200.

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0. 10. 20. j 1 1 I 1 1 1 J I i_i 1 1 L_i 1 1 1 I 1 1 1 L

30. 40. 50. 60. 70

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EXPERIMENT NO. 1356 TEST DRTE « 12.8.1986 TOTAL MASS OF RIR USED (KGS) - 2.610

.08 . 1 1 1 1 | 1 1 1 1 [ 1 1 1 1 | 1 1 1 1 | 1 1 1 1 | 1 1 1 1 | 1 1 1

1 1 1 1 I 1 1 1 1 I 20. 30. 40. 50.

CTCLE TIHE (SECS)

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60. 70.

ORIFICE PLATE NO. 4

Figure 4.10 Transient plots of major conveying parameters for U G A (Exp. No. 1356, Test Rig F2).

77

The following observations are made as a result of these experiments.

(a) SGA was found to be unsuitable for conventional dense-phase conveying and had to be transported under dilute-phase conditions. This produced high velocities and an excessive as well as unacceptable level of product degradation. Plug-phase conveying (using orifice-air only) was found to be more suitable and efficient than the dilute-phase mode (i.e. in terms of air flow and degradation requirements). However, it should be noted that as only a limited amount of product is transferred per cycle, relatively low conveying rates must be tolerated (e.g. 2.5 kg s-1 instead of 6.8 kg s-1 for dilute-phase) and may require the selection of larger sizes of pipe (i.e. when scaling-up to required system capacities).

(b) UGA, which possesses a wider particle size distribution (and also a larger amount of fines) than SGA, displayed a greater propensity to form stable plugs in the pipeline. For example, when U G A was subjected to the same operating conditions (i.e. which were found reliable for SG A ) , a strong plug formed during the cycle and resulted in substantial pipe vibrations, higher back-pressures and excessive degradation.

(c) The selection of orifice- and ring-air prevented the formation of such plugs and produced a more gentle duning type of flow (and hence, less degradation). Note that for this comparison, the air supply valves were adjusted to maintain similar values of the supplied air mass flow rate mf (i.e. with respect to orifice-air only).

(d) The minimum flow rate condition for UGA (i.e. mf « 0.076 kg s-1) was found to be slightly greater than the more mono-sized S G A (i.e. mf= 0.07 kg s-1).

(e) An increased level of fines definitely has an impact on conveying performance. Also, these results demonstrate that the plug-phase mode of transport is quite sensitive to changes in material properties (i.e. particle size) and also the blow tank configuration and air injection. Test work is required to select the most efficient mode of conveying (and hence, the method of air injection) for a given material and its behavioural properties. For example, S G A may be suitable for low-velocity [2,3] conveying, although higher operating pressures are expected. Such matters of powder classification are considered further in Chapter 5.

The following case study on bone char presents in more detail the effect of air injection on conveying performance and especially product degradation. 4.3.2 Bone Char This material is required for the clarification of sugar liquor and it is desirable that it be conveyed into the cisterns with a minimum amount of degradation. Note the material is recycled pneumatically as often as possible and rejected once the level of fines below 425 u.m becomes too great. Similar to SGA, conventional dense- or even dilute-phase conveying was considered unsuitable. Hence, plug-phase conveying was pursued with the aim of establishing reliable operating conditions and an efficient method of air injection based on particle sizing information.

78

The following test program was carried out on Test Rig F1 (0.113 m3 plug-phase blow tank, L = 41 m & D = 105 m m ) .

(a) Test Group No. 1 Using a 100 kg batch size, a reliable air flow of mf » 0.09 kg s-1

(determined from preliminary test work) with a combination of orifice- and ring-air, 10 tests were carried out on essentially the same sample of bone char. Three grab samples were taken after the fifth and tenth experiments (i.e. for later size analysis and averaging). Note that to compensate for product being left in the pipeline (as described in Section 4.3.1), additional fresh material had to be added to the blow tank during the first two transient experiments (i.e. to maintain a batch size of« 100 kg). The following summary of results is based on the transient plots of the major conveying parameters (which were obtained from each experiment) and the sieve analyses performed on each grab sample.

Conveying cycle time varied between 28 and 30 s. Based on the latter, an average conveying rate of ~ 3.33 kg s_1 or 121 h-1

is calculated. The air flow was constant at mf« 0.09 kg s*1. As more experiments were carried out, the operating pressure increased gradually from « 370 to 470 kPag (based on Exp. Nos. 1227 and 1235, which were the second and last experiments of Test Group No. 1). Refer to the pressure plots presented in Figure 4.11. Table 4.4 summarises the averaged particle sizing data, which were obtained from each of the three grab samples (i.e. collected after the fifth and tenth experiments). Data relevant to the original or as received sample is included.

Sample

Original After 5 exps. After 10 exps.

Sieve Size (nm)

425

17.0 26.0 32.0

300

3.5 11.0 17.0

212

0.5 5.0 9.0

Table 4.4 Cumulative % mass passing through sieve size (for orifice- and ring-air).

(b) Test Group No. 2 Using a fresh batch of bone char, an additional 10 experiments were carried out using a combination of orifice-, ring- and supplementary-air for the blow tank. All other test procedures and set-up conditions were similar to Test Group No. 1 (e.g. grab samples, batch size « 100 kg, mf« 0.09 kg s'1). The following major results were obtained.

EXPERIMENT NO. 1227 TEST DRTE i 19.3.1986 79

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Figure 4.11 Transient plots of blow tank and pipeline air pressure for bone char (Exp. Nos. 1227 & 1235, Test Rig F2).

80

Conveying cycle time varied between 30 and 31 s and hence, the average conveying rate was « 3.23 kg s_1 or 11.61 Iv1. The air flow was constant at mf « 0.088 kg s-1 (slightly lower than the 0.09 kg s_1 recorded during Test Group No. 1). The operating pressure increased gradually from = 255 to 285 kPag (based on Exp. Nos. 1237 and 1245, which were the second and last experiments of Test Group No. 2). Refer to the pressure plots presented in Figure 4.12. Table 4.5 summarises the average results obtained from the particle size analyses.

Sample

Original After 5 exps. After 10 exps.

Sieve Size (\xm)

425

18.0 20.0 26.0

300

4.1 6.5 12.0

212

0.6 2.2 6.0

Table 4.5 Cumulative % mass passing through sieve size (for orifice-, ring- and supplementary-air).

From the above results, it can be seen that (with respect to orifice- and ring-air) the orifice-, ring- and supplementary-air combination produces

a mode of conveying that is more gentle to the material, lower operating pressures (e.g. for the tenth experiment, pbt « 285 kPag instead of« 470 kPag), lower levels of degradation (e.g. a difference of« 6 % in the cumulative % mass passing through a 425 u.m sieve size) which would allow the material to be recycled a greater number of times (i.e. for the sugar liquor clarification process), hence, generally a more reliable and efficient method of transportation.

4.3.3 Crushed Bath Using Test Rig F3 (0.113 m3 plug-phase blow tank, L = 161 m & D = 105 mm), numerous experiments were carried out recently to determine reliable operating conditions and also examine the effect of blow tank air injection on the conveying performance of this material. Note that crushed bath possesses a very wide particle size distribution (e.g. « 30 % by weight of product < 0.5 m m , dso = 3.9 mm, » 10 % by weight of product > 11.2 m m , top size « 16 to 18 m m ) and usually contains a certain amount of alumina. Preliminary experiments established a minimum flow rate condition at mf» 0.28 kg s_1. An air flow of mf» 0.31 kg s*1 and a batch size of «100 kg of fresh material were maintained for each experiment. Four tests were carried out initially with orifice-air only (to establish the consequences of conveying this material in the form a stable plug). These were followed by three further experiments using a combination of orifice-, ring- and supplementary-air. The results are summarised in Table 4.6.

EXPERIMENT NO. 1237 TEST DATE i 21.3.1986

300.

£200.

ui

cc <n ui £100.

0.

" i — 1 — 1 — 1 — | — 1 — 1 — 1 — 1 — j — 1 — 1 — 1 — 1 — | — i — » — 1 — 1 — | — 1 — i — 1 — r — j — 1 — r — i — r

CH. NOS « 1 — : B-.-. -2 : 3 4

10. 20. 30. 40. SB. 60

CYCLE TIME (SECS)

EXPERIMENT NO. 1245 TEST ORTE * 21.3.1986

300. r—1—1—1—1—j—1—1—j—r

cs ^200. *:

UJ

cc CO CO Ui

cc °"100.

0. 60

CYCLE TIME tSECS)

Figure 4.12 Transient plots of blow tank and pipeline air pressure for bone char (Exp. Nos. 1237 & 1245, Test Rig F2).

82

Exp. NO.

108-10 -11 -12 -13

108-14 -15 -16

Mbt (kg)

100 100 100 100

100 100 106

Air Inj. O/R/S

0 O O 0

O+R+S O+R+S O+R+S

Max. mf (kg s-1)

.30

.31

.31

.31

.31

.31

.31

Mf (kg)

8.6 8.1 8.6 8.7

11.8 11.7 12.5

Ms

(kg)

95 104 99 104

94 93 101

Max. Pbt (kPag)

510 550 480 600

66 71 75

tc (S)

31 29 30 32

40 40 42

Avg. ms

(kg s-1)

3.1 3.6 3.3 3.3

2.4 2.3 2.4

Table 4.6 Summary of plug-phase conveying parameters for crushed bath (Test Rig F3, L = 160 m & D = 105 m m ) .

Transient plots of the major conveying parameters for Exp. Nos. 108-12 (orifice-air only) and 108-16 (orifice-, ring- and supplementary-air) are presented in Figures 4.13 and 4.14, respectively. The following observations are based on the above results. (a) The selection of only orifice-air caused the formation of a stable plug

during the conveying cycle. Although this produced severe pipe vibrations, very high pressures (e.g. 500 to 600 kPag) and would tend to suggest imminent blockages, Exp. Nos. 108-10 to -13 demonstrated good repeatability. However, such results are not considered practical and certainly would create serious hardware problems (i.e. for an actual installation).

(b) Orifice-, ring- and supplementary-air prevented the formation of the stable plug described in (a), and provided a surprising improvement in conveying performance (although the values of average m s were « 25 % lower than those determined from Exp. Nos. 108-10 to -13). The discharge of material from the end of the pipeline was very gentle (due to the duning mode of flow) and operating pressures did not exceed 100 kPag.

4.3.4 Summary The results presented for the previous case studies involving screened and unscreened granulated aluminate, bone char and crushed bath, demonstrate that

the plug-phase mode of conveying is able to handle efficiently a large range of conventionally difficult materials that otherwise could be considered only for dilute-phase transport (i.e. low-velocity or by-pass technology may not be possible),

orifice-air only tends to form a stable plug of material in the pipeline, usually resulting in higher operating pressures and levels of degradation,

500.

a.

400. -

300.

100.

i I I — i — | — i — i i i | - 1 i i i | i i i i | i i i i | i • i i i | i i i i

CH. NOS 5 28__

'~ I- *- •• ' ' L_i i i i_J I i L_L 10. 20. 30. 40. 50. 60. 70.

CTCLE TIME (SECS)

200.

o

a

IX. 3L

100.

" T - 1 — I | I I I I | — I I I—I | I I I I | — I I I I | I I I I | I I I I

CH. NOS

I i i i • I t t ' i I i i ' ' L_i I I i I i_ 01 8. 10. 20- ' 30- 4 0. 50. 60- 70.

CYCLE TIHE (SECS)

+1 i i—i—i [ i i i i [—i i i—i | i—n—i | n i i |-"i i i i—p

u UJ in

cr cc 3 O

in

•n cr z cc a

I t i i i I • • • i I • • r i I i i T i-l i i B. 10. 20. 38- 40. 58.

CYCLE TIHE (SECS)

60. 70.

Figure 4.13 Transient plots of major conveying parameters for crushed bath (Exp. No. 108-12, orifice-air only, Test Rig F3).

80. * ' ' i • ' i ' | ' ' ' ' | i • i i | i i i i | i i i i | i i i i | i i 1 1 | i i i i | i i i i

CH. NOS

S A 28_.

I i i_T I - L J .I I I I I I I I I I I I I I I

0. 10. 20. 30. 40> 50. 60. 70. 60. 90. 100.

CYCLE TIME (SECS)

200.

5 100.

i i i i I i i i ' I i i i i i i i i i I i i i i I i i i i I i i i i I i i i i I i i i i 1 i i i i i i i | i i i i | i i i i I I i i r | I i I i | i I i I | i I i i |

CH. NOS 1

n l i i - f l i i i i I i i i i I i i i i I i i i i I i i i i I i t i i I i i i i I i i i i I i i i

-J. 10. 20. 30. 40. 50. 60. 70. 60. 30. IBB

CYCLE TIME (SECS)

u UJ in u •XL

a cc •x a

in in

a x CC cc

4 i i i i i | i i i i | i i i i | i i i i | i i i i | i i i i | i i i i | i i i i | i i i i | i i i

• l . i i i I • i i i i i i t t i i i • l , ._• • l . ^ i i I - » • • • • • • < •

0 0 . 10. 20. 30. 40. 50. 60. 70. 60. 90. 120.

CYCLE TIME (SECS)

Figure 4.14 Transient plots of major conveying parameters for crushed bath (Exp. No. 108-16, orifice-, ring- and supplementary-air, Test Rig F3).

orifice-, ring- and/or supplementary-air prevents the formation of such plugs and produces a more gentle duning mode of flow (with lower levels of product degradation),

this method of transport is relatively sensitive to changes in the physical properties of a material (e.g. particle size), although this may be compensated to some extent by selecting a different method of air injection (usually orifice-, ring- and supplementary-air), generally, the method of blow tank air injection has a significant impact on the overall performance of a plug-phase pneumatic conveying system and should be given careful and sufficient consideration in the design, selection and operation of such systems.

86

CHAPTER 5

87

5. P O W D E R CHARACTERISATION

5.1 Introduction

The characterisation of bulk solids is becoming an increasingly important design requirement to

assess the suitability of conveying a material in the dense-phase mode, establish a more efficient (economic) means of transportation, determine the feasibility of long-distance pneumatic conveying (viz. > 800 m) for a particular material, and generally, select the most suitable mode of transport (e.g. dilute-, dense-, pulse-, plug-phase or low-velocity conveying).

To obtain an appreciation of the parameters which are relevant, the physical properties, fluidisation performance and pneumatic pipeline conveying characteristics were obtained for the seven fly ash samples listed in Table 3.2, as well as the Tallawarra pulverised coal sample that was considered previously in Section 3.1. As introduced in Section 3.3.1, the results obtained from this section of work are intended to provide a data base, from which the relative assessment of a given fly ash could be undertaken with only a small amount of sample (e.g. < 3 kg). Also, results from recent investigations into predicting the performance of some coarser products (e.g. PVC powder, screened coke, coarse ash) are presented and/or referenced to provide additional information and comparisons for the development of a general procedure to classify bulk solids (i.e. to meet the above design requirements). 5.2 Physical Properties The following bench-type experiments were conducted on the pulverised coal and each fly ash sample.

(a) Particle size analysis. (b) Solids or particle density measurement (viz. ps) using a Beckman Model

930 air-comparison pycnometer. (c) Loose-poured bulk density, pbi-

The above analyses also were carried out on PVC powder [14,16], screened coke [14,16] and coarse ash [27]. However, it should be noted that three different devices were employed for the measurement of particle size (viz. a Coulter Counter Model ZM for the pulverised coal and fly ash samples, sieving tests for PVC powder and screened coke and a Malvern Model 2600C laser diffraction analyser for the coarse ash). Before analysing and presenting the results obtained from this section of work, it is necessary to become familiar with the various definitions of particle size. 5.2.1 Definitions of Particle Size

For several decades, it has been accepted that the size and density of particles have a significant influence on the behaviour of a fluidised bed [28] and the performance of a pneumatic conveying system. Particle size distribution also has been recognised as an influential factor [29,30].

88

Possibly, the most difficult aspect of determining particle size is selecting initially the correct or relevant definition and then calculating a mean diameter to represent the complete bulk solid. To some extent, this will depend on

the measuring apparatus and its principle of operation, the final application or requirements (e.g. prediction of free settling velocity Voo, minimum fluidisation velocity V m f or pipeline air pressure drop Apt), and the basis of definition used in a theoretical or empirical relationship (e q sieve or volume measurement).

In some cases (especially for very fine powders), researchers have looked at other properties to explain/classify product behaviour. For example, Geldart et al [251 have found that the ratio of tapped to aerated bulk density provides a good indication of the likely fluidisation characteristics of fine and cohesive powders However, in this thesis, it was decided to pursue particle size measurement and evaluate its importance (as well as density) for powder characterisation. As a result of using three different devices to analyse particle size distribution the following different definitions and related properties [31,32] were required to calculate the mean particle diameter.

d pm

(5.1)

-pwm

Arithmetic mean of adjacent sieve sizes.

Mean particle size from a standard sieve analysis,

Z(AM)

where AM is the mass percent of product between adjacent sieves.

Weighted mean diameter [32] based on a sieve analysis,

£ (AM dp) E(AM) (5.2)

Jsv

Jsvm -

Diameter of a sphere with the same surface area to volume ratio as the particle.

Mean surface volume diameter,

2 (AM)

d vm

Diameter of a sphere with the same volume as the particle.

Mean (equivalent) volume diameter,

SCAM)

(5.3)

(5.4)

89

dvwm = Volume weighted mean diameter [32],

X (AM dv) Z(AM)

d5o = Median particle diameter [32].

dV50 - <%() f°r a volume diameter distribution.

°p50 = ^50 for a sieve size distribution.

\\r = Particle sphericity [31],

dSv dv

Although the Coulter Counter Model ZM has been found to have certain disadvantages [33], it still does make use of a convenient principle [32], where within a given size range, particles are counted according to their displaced volume (viz. dv). The resulting number count distribution then is transformed to a weight % frequency or cumulative distribution. This measurement technique was applied to the pulverised coal and fiy ash samples, and after obtaining the resulting particle size distributions, Equations (5.4) and (5.5) then were used to calculate dVm and dvwm 0-e- based on the various measured values of A M and dv). For PVC powder and screened coke, Equations (5.1) and (5.2) were applied to the sieving results (viz. A M and dp) to calculate the mean and weighted mean diameters d p m and dp w m . Assuming that the measured particle diameters for the coarse fly ash (i.e.using the Malvern analyser) were « dv, allowed d Vm and d V wm to be determined from Equations (5.4) and (5.5). Table 5.1 summarises the mean and weighted mean diameters obtained from the above section of work, as well as the various median particle diameters which were read directly from the actual size distributions. Note that the actual values of A M and dv (or dp), which were obtained from the cumulative size distribution of each material, are summarised in tabular form in Appendix A. Using the following results of Geldart and Abrahamsen [31] and the assumed values of sphericity, the surface volume diameter dSvm was calculated for each material, and values of dV50 and dVm also were estimated for the P V C powder and screened coke. 1.1 <-7^-< 1.2, Average ~» 1.127.

dp dp

For pulv. coal, \|/« 0.7 (angular, tetrahedron), fly ash, y « 0.9 (spheroids), P V C powder, y « 0.85 {rough spheroids), screen, coke, Y « 0.80 {coarse spheroids), coarse fly ash, \\r« 0.85 (spheroids and some unburnt coal particles).

(5.5)

(5.6)

90

NO.

1

2

3

4

5

6

7

8

g

10

11

Product

Tallawarra Pulv. Fuel

Tallawarra Fly Ash

Eraring Fly Ash

Munmorah Fly Ash

Vales Point Fly Ash

Gladstone Fly Ash

Wallerawang Fly Ash

Liddell Fly Ash

PVC Powder

Screened Coke

Coarse Fly Ash

Ps (kg nr3)

1600

2350

2160

2100

2130

2250

2195

2415

1400

1940

1860

Pbl (kg nv3)

760

500

880

650

700

1030

455

640

575

985

787

dv50 (u.m)

30.0

19.6

27.4

25.4

18.8

17.6

11.5

13.3

152.1

529.7

99.0

^vwm (u.m)

37.4

27.1

49.4

38.1

28.7

27.6

17.5

29.3

-

122.8

dvm (Jim)

14.1

12.0

14.0

13.1

14.1

10.1

8.0

8.6

148.8

378.7

59.1

Cp50 (lim)

-

135.0

470.0

-

dpwm

(nm)

-

150.0

538.7

-

dpm

(nm)

-

132.0

336.0

-

dsvm (nm)

9.9

10.8

12.6

11.8

12.7

9.1

7.2

7.7

126.4

302.9

50.2

Table 5.1 List of samples and physical properties.

5.3 Fluidisation Analysis

5.3.1 Experimental Apparatus

The fluidisation test facility, which was employed during the initial investigations on pulverised coal and fly ash, comprised the following major components.

A 102 mm internal diameter pyrex vertical column (750 mm long) with an attached graduation scale for bed height measurement (viz. hb). A 6 m m thick x 35 u.m Porex™ permeable plastic gas distributor covered with an epitropic Goretex™ filter fabric to prevent particulate penetration of the plastic and also allow the discharge of excessive electrostatic charge.

91

A plenum chamber base with a retaining ring assembly to house the 6 m m thick Porex™ gas distributor. A Goretex™ filter fabric covering the top of the vertical column to prevent atmospheric contamination and any loss of product. Two Rosemount Model 1151 D P differential pressure transmitters (0 to 152 m m H 2 O and 0 to 762 m m H2O) for direct measurement of the air pressure drop across the material via Goretex™ protected pressure tappings (of which the design is similar to that used on the pneumatic conveying test rig, as shown in Figure 2.4). Four in-line rotameters connected to the air supply line to monitor the amount of air passing through the bed of material.

A schematic layout of the fluidisation test rig is presented in Figure 5.1.

5.3.2 Results

With approximately 2.5 to 3.0 kg of product and using the air flow reduction technique [34], the following parameters were recorded for Sample Nos. 1 to 8.

Height of material above the Porex™ gas distributor, hb (cm). Air pressure drop across the bed of material, Apb (mm H2O). Volumetric flow rate of fluid (air) passing through the rotameter(s), Qf

(cm3 s-1). Operating conditions of the rotameter(s) (i.e. air pressure and temperature). Atmospheric conditions (i.e. air pressure, temperature and relative humidity).

For each value of Qf, the corresponding value of mf was calculated using the operating conditions of the rotameter in question. Assuming the pressure drop across the final Goretex™ filter fabric to be negligible, the superficial velocity of air, Vf, leaving the bed of material also was determined. The variation of the average air pressure gradient, Apb Iv1 (mm H 2 0 cm"

1), with respect to Vf was plotted for each coal and fly ash sample and the resulting fluidisation curves have been reproduced and superimposed onto the one plot shown in Figure 5.2. The following general observations also were noted during the experimental program.

Samples 1, 3 and 4 fluidised well and retained aeration for a considerable length of time (i.e. after the air supply valve was turned off). Fluidisation was accompanied by extensive bed expansion. Bubbling occurred at air flows greater than that required for fluidisation. For Samples 2, 5, 6 and 7, bad channelling (or rat-holing) occurred throughout the range of air flow rates considered. This produced large regions of dead material (preventing uniform and complete fluidisation of the sample). Although some channelling and poor mixing was observed for Sample 8, the material still displayed a semi-fluidised condition with some bed expansion (e.g. refer to the fluidisation curve shown in Figure 5.2).

Using the above procedure, measurements also were taken for PVC powder and screened coke and the resulting fluidisation curves are presented in Figure 5.3.

•i- O-

oo

92

77. fO

S-CD

-a

cu +-> ro

t/J

O S-

o

s-cu

X cu a> S-

o C£J

CU +J

a: cn

- cn c X E v-CU T - ro 5- o : 4->

o i cu a. o a:

>

s v

>

1

\ N N '

1

.1 /. .. 1 t ' V f 1

k :

s i

;

CO

cu

>{XJ-

l-txj-

T-txj-

>txH

i-

cu cu +J I—

cu o E cn (O c +J o o o oc

t/1

cu

. CU

• 13

X CU CU o. cn O c s- ro CU i— D_ U.

. CU Q CL • "^ i—i Q .

E n -E CU OJ CU o +> •—I 00

J S- $-3 CU i/> +-> 00 cu S-

o 03 •+— •*—»

CO CD +-'

c o to co |g |g a> 13

o _ca

o

ca

E cu x: o W

3 D)

(/) c ro s-

93

6.0

4.0

Apb/hb

(mmHgOcrrf1)

2.0

0

Sample No. (Table 5.1)

2.0 4.0 6.0

(cm s"1)

Figure 5.2 Comparison of fluidisation curves for pulverised coal (Sample 1) and fly ash (Samples 2 to 8).

94

10

8

6

A?b V1

(mm H-0 cm-1) 4

2

0 0 2 4 6 8 10

Vf Ccm s"1)

Figure 5.3 Fluidisation curves of PVC powder (Sample 9) and screened coke (Sample 10).

T r

/

Screened / Coke /

/ /

/ /

/ /

/

/

PVC Powder

J L

95

Note that insufficient material prevented a fluidisation analysis being performed on the coarse ash (Sample 11). Additional observations, which were made during each experiment on P V C powder and screened coke, are summarised below.

PVC powder (Sample 9) displayed good fluidisation characteristics but deaerated quickly when the air supply to the plenum chamber was turned off. Bubbling occurred at approximately the same air flows which were required for fluidisation and was accompanied by only a small amount of bed expansion.

For the range of air flow rates considered (viz. 0 < Vf < 8 cm s'1), the screened coke (Sample 10) did not display a fully fluidised condition. However, at the higher air flows (e.g. Vf > 5 c m s-1), a top section of the bed was seen to bubble strongly and hence, exhibit some form of fluidisation. This depth of active material decreased as the air flow was reduced, and for Vf < 2.0 cm s-1, almost all the activity had disappeared. On closer inspection of the test chamber (i.e. after the experiment), the bed of material was seen to be made up of stratified layers, each having a different particle size range (viz. coarse at the bottom and fine powder at the top).Jn fact, the lower section of material resembled a granular bed. This is believed to be caused by the relatively wide size distribution of the product (e.g. 60 (im < dp < 1500 urn) and the gradual segregation/settling of particles as the air flow was reduced during the experiment. The screened coke also deaerated very quickly when the air supply to the fluidisation rig was turned off.

5.4 Pipeline Conveying Characteristics The conveying characteristics for each fly sample have been presented previously in Section 3.3.3 (viz. Figures 3.10 to 3.16). The test rig that was employed for this work has been described in Section 3.3.2. For ease of comparison, the m s contour lines of 1, 2 and 3 kg S"1 have been reproduced in Figure 5.4.

5.5 Powder Classification Techniques

The Geldart [24] fluidisation and Dixon [23] slugging classifications have been found useful in explaining

some of the feeding problems that can occur in blow tanks [35] (e.g. refer to Section 4.2.1), and

the differences that can occur in flow performance and minimum transport behaviour [16] (e.g. refer to Section 3.4.3).

Similar findings have been obtained from the current work on pulverised coal and fly ash, and are discussed further in the following sections. Modifications to the Geldart [24] fluidisation diagram have been proposed by Molerus [36] and Zenz [37], but will not be considered here as they require some knowledge or measurement of particle adhesion forces and bulk surface tension, respectively. That is, detailed investigations into evaluating and/or developing such fluidisation diagrams were considered beyond the present scope of work and the overall objectives of this thesis. A more recent classification technique that makes use of two different bench-type experiments (viz. permeability and deaeration) has been

(a) ms = 3.0 kgs"1

180

Apt

(kPa)

160 -

140

Fly Ash Sample No. " (Table 5.1)

0 .02 .04 .06

mf (kgs-1)

(b) iru = 2.0 kgs-1

140 -Apt

(kPa)

120 -

100 -

.02 .04

mf (kgs-1)

.06

(c) ms = 1.0 kgs-1

100 -Apt

(kPa)

80 -

60 -

0 .02 .06 .04

mf (kgs-1)

Figure 5.4 Comparison of pipeline conveying characteristics for fly ash (Samples 2 to 8, Test Rig B1).

97

presented recently by Mainwaring and Reed [38]. As deaeration experiments were not carried out for these investigations and as most of the fly ash samples were not able to be fluidised readily (i.e. for the range of air flow rates considered) this technique also will not be considered in detail for the present study. However, where relevant some comments and references will be made to the results [38].

5.5.1 Fluidisation

Using fluidisation data obtained from several researchers, Geldart [24] characterised powders into four groups (viz. A, B, C and D) according to their fluidisation behaviour and developed a classification diagram, as shown in Figure 5.5. The reader is directed to the Geldart [24] paper for detailed descriptions of the various groups (including a numerical technique to distinguish between each one), and the Geldart et al. [25] paper for recent investigations into the fluidisation of cohesive powders. Note that the mean diameter used by Geldart [24] is actually a surface volume mean diameter, based on Equation (5.3). Hence, using the values ofdsvm listed in Table 5.1, it was found that according to the Geldart classification diagram [24] Samples 1 to 8 are Group C powders (i.e. difficult to fluidise due to

cohesive properties or large interparticle forces), Samples 9 and 11 are Group A powders (i.e. easy to fluidise, retain aeration, bubbling occurs some time after fluidisation and considerable bed expansion), Sample 10 is a Group B material (i.e. easy to fluidise, deaerate quickly, bubbling occurs at or just after fluidisation, small bed expansion).

10'

cn

<+-Q. I Q.

CU U c cu S-<D M -<4-

1— C/l

C cu o

10;

10<

10: io2 io3

Mean Particle Diameter (nm)

10l

Figure 5.5 The Geldart [24] fluidisation diagram.

98

However, based on the actual results presented in Figures 5.2 and 5.3 as well as the observations listed in Section 5.3.2, it was found that

Samples 1, 3 and 4 exhibited typical Group A behaviour, Samples 2, 5, 6 and 7 displayed poor fluidisation performance similar to that described by Geldart [24] for Group C powders (e.g. poor particle mixing due to cohesive properties, channelling), Sample 8 did display a semi-fluidised condition, but this was accompanied by poor mixing and channelling (similar to Group C materials), Samples 9 and 10 produced characteristics similar to that described for Group B, although the test rig was found to have insufficient capacity to fluidise completely the relatively coarse and heavy particles of screened coke (Sample 10).

On closer inspection of the particle size definitions and equations presented in Section 5.2.1 and the actual values of diameter listed in Table 5.1, it was found that (a) the reciprocal form of definition (e.g. Equations (5.1), (5.3) and (5.4))

tends to over-emphasise the influence of the finer particles, (b) the weighted or product type of definition (e.g. Equations (5.2) and (5.5))

tends to under-emphasise the influence of the finer particles (or overemphasise the coarse end of the size distribution),

(c) the relatively wide particle size distributions of pulverised coal and fly ash (e.g. 1 to 200 urn) seem to be the major contributing factor to the effects described in (a) and (b) (e.g. refer to the actual size range of products considered by Geldart [24]),

(d) the large differences described in (a) and (b) are not so apparent for PVC powder, which has a fairly narrow size distribution,

(e) for the materials considered in these investigations (especially Samples 1 to 8 and even Sample 9), the median particle diameter dV50 seems to provide a better indication of fluidisation performance (i.e. as described by Geldart [24]).

To demonstrate this further, the locations of Samples 1 to 11 have been included on the fluidisation diagram [24], which is repeated in Figure 5.6. Note as a result of (e) above, the dvso size was taken to represent the average or mean particle diameter of each material. According to this classification: Samples 1, 3 and 4 are Group A powders (i.e. easily fiuidised and retain aeration); Samples 2, 5, 6, 7 and 8 are Group C powders (i.e. cohesive, difficult to fluidise, channel or rat-hole, poor mixing); Sample 10 belongs to Group B (deaerates quickly); Samples 9 and 11 are Group A but lie very close to the A B boundary (this indicates that such border-line materials may exhibit characteristics from either one of the adjoining categories). Generally, these classifications confirm the experimental observations reported in Section 5.3.2 and explain the large differences that occurred in the fluidisation performance of Samples 1 to 10.

99

10'

cn

Q. 1 (/) Q.

CU U

c cu S-

cu <+-

in

c cu

10:

10' IO1 10< 10; 10'

Mean Particle Diameter , dm (nm)

Figure 5.6 The Geldart [24] fluidisation diagram showing the location of Samples 1 to 11.

Therefore, based on the results of this investigation, the Geldart [24] classification diagram does seem to provide a reliable technique for predicting fluidisation behaviour (i.e. using dvso instead of dSvm. especially for products having a wide particle size distribution). However, it should be noted that when Geldart proposed his classification diagram [24], he suggested a shaded boundary region between Groups A and C. This indicates that some degree of overlap may exist between the two categories (i.e. some typical Group C powders could display Group A performance, or vice-versa). For example, refer to the results obtained from Sample 8 (which is classified Geldart [24] Group C but almost displayed Group A performance). Similar results were obtained recently from a classified Group C fly ash (i.e. dV5o = 14 urn and p s = 2155 kg nr

3), which was tested for the Electricity Commission of N.S.W. and displayed good fluidisation characteristics (e.g. similar to Samples 3 and 4). However, for the following reasons, it is believed that ultimately the actual product(s) should be tested in a large-scale fluidisation test rig (e.g. similar to the one shown in Figure 5.1), so that actual characteristics and performance may be established.

100

Perhaps the greatest difficulty in predicting fluidisation performance (viz. via the Geldart [24] classification) is deciding on a particular diameter to represent the complete material (especially if the product possesses a wide particle size distribution). This is supported to some extent by the new bulk density approach proposed by Geldart et al. [25]. There is some doubt over the location of the boundary separating Groups A and C. A particle size versus density relationship (i.e. as proposed by Geldart [24]) may not be sufficient to define these regions (e.g. refer to [25]). Products lying close to a particular boundary may exhibit fluidisation behaviour from either one of the adjacent categories (e.g. a product which is in close proximity of the A B boundary may exhibit characteristics from either Group A or B). It is difficult to estimate the error associated with each boundary (e.g. between Groups A and B or B and D).

Appreciating the possible problems and inadequacies of the Geldart [24] classification diagram, this technique still provides a good initial indication of what to expect when a given product is fluidised or mixed with air. Even though fluidisation may be confirmed by experiment and used subsequently in the design of feeders (e.g. to establish possible rat-holing problems inside a blow tank, as demonstrated in Section 4.2.1), the application of such information to predicting pneumatic conveying performance is a different matter and in fact, has been found recently [14,22] to be inadequate. This is discussed further in the following section. 5.5.2 Slugging In an attempt to describe the natural behaviour of different solids in dense-phase, Dixon [39] developed theoretical slugging diagrams for different pipe diameter systems on the basis of the Geldart [15] fluidisation classification diagram (i.e. Groups A, B, C and D). Although Dixon's diagrams were based on slugging criteria [40] for vertical transport, he indicated that they also generally support the observed behaviour of materials in horizontal pipes. However, Dixon [23] stated that there is some doubt over the location of the boundary between Groups A and C, and he subsequently reproduced this Geldart boundary directly onto the slugging diagrams. Refer to Figures 5.7 and 5.8 for examples of slugging diagrams for 50 and 100 m m N B pipe diameter systems. For detailed descriptions of the various slugging classifications and the mathematical formulae to distinguish between each classification group, the reader is referred to the Dixon papers [23] and [39]. 5.5.2.1 Slugging Diagram Modifications After attempting to reproduce the slugging diagrams on the University's mainframe computer (using Fortran 77 and the Plot Package), the terminal velocity equations which were used by Dixon [39] to generate the Group A B boundary were found to be limited to a maximum particle Reynolds number of 1000. Alternative empirical expressions were investigated and the standard drag curve correlations recommended by Clift et al. [41] were incorporated finally. Examples of the modified slugging diagram, which were obtained from the mainframe computer for 50, 80, 100, 150 and 200 m m N B Schedule 40 pipe diameter systems, are presented in Appendix B. Note that the Group A-C boundary has been shown as a dashed line on each diagram due to the uncertainty over its precise location, as indicated by both Geldart [24] and Dixon [23].

101

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CU

o e cu %-cu «+-4 -•r— C/l

CD O

6.0

2.0

1.0

D

STRONG

AXISYMMETRIC

SLUGS


Figure 5.7 The Dixon [23] slugging diagram for a 50 m m pipe diameter system.

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AXISYMMETRIC SLUGS

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100


Figure 5.8 The Dixon [23] slugging diagram for a 100 m m pipe diameter system.

102

5.5.2.2 Results

The classification of each sample listed in Table 5.1 has been represented on the modified slugging diagram as shown in Figure 5.9. Note that a 50 m m N B Schedule 40 pipe diameter system was selected to represent the test rig employed in these investigations (viz. Test Rig B1) and dvso was used to represent the mean particle diameter (as was found necessary in Section 5.5.1). The following classifications are based on Figure 5.9 and the suggestions of dense-phase suitability are obtained from Dixon [23,39] (i.e. based on the concept of a moving fluidised bed). Samples 1, 3 and 4 are Group A (good candidates for dense-phase

conveying; high values of m*). Samples 2, 5, 6, 7 and 8 are Group C (usually considered to be too cohesive for the dense-phase mode and hence are poor candidates; but may demonstrate good properties attributable to Group A powders). Samples 9 and 11 are Group B (possible candidates for dense-phase but at higher velocities where dune flow and asymetric slugging are prevalent; could produce severe pipe vibrations if the dunes are allowed to fill the pipe; generally, low values of m*). For high operating pressures (e.g. > 4 atmospheres), Figure 5.9 suggests that Sample 11 becomes a Group A material. For low pressures, Sample 10 is Group D (can be conveyed in dense-phase over a wide velocity range; moderate values of m* less than those obtained for Group A but greater than Group B materials). However, for moderate to high operating pressures (e.g. > 2 atmospheres), this material may behave like a Group B powder (i.e. due to a change in air density).

The change in slugging characteristics of a given material due to increasing or decreasing operating pressure (i.e. as suggested above for Samples 10 and 11), is difficult to confirm. At this stage, it is suggested that such materials (i.e. those that lie very close to or on a classification boundary) may exhibit slugging behaviour from either one of the adjacent categories. Note in Section 5.5.1, a similar suggestion was proposed for the Geldart [24] fluidisation diagram. Intuitively, it would be expected that the pulverised coal and fly ash samples with the good fluidisation characteristics (viz. Samples 1, 3 and 4) would yield similar dense-phase pneumatic conveying results, with the performance of the Geldart [24] or Dixon [23] Group C powders (viz. Samples 2, 5, 6, 7 and 8) being quite different. In fact, Dixon [23] suggested that dense-phase conveying can be regarded as a moving fluidised bed, and that the Group C powders are considered usually to be too cohesive for this mode of transport. However, all the samples conveyed reasonably well with the minor exception of the Group C powders (viz. Samples 2, 5, 6, 7 and 8), which produced generally higher pressure drops (refer to Figure 5.2) and slightly greater irregular flow characteristics. An example of the latter is shown in Figure 5.10 (b), where the steady-state pipeline air pressure (at location G1, see Figure 3.8) is seen to fluctuate in the range 110 to 120 kPag. It is interesting to note that, although considerably higher pressures were obtained when the Group C samples were conveyed in the dense-phase mode (compare the conveying characteristics with the two-phase flow diagram [7]), such differences were reduced significantly in the dilute-phase regime (typically, mf > 0.04 kg s*1).

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Generally, the pulverised coal and fly ash samples investigated in this project were all found to be good dense-phase materials displaying a wide rangeability of conveying parameters (e.g. a max. m* « 500 was achieved for the coal sample over a distance of 25 m, and a value of 150 was obtained easily for the fly ash samples over a distance of 71 m). Hence, even though slightly higher pressure drops were obtained with the Group C materials, the precautions expressed by Dixon [23] in relation to this category, seem exaggerated. However, it is known from experience that difficulties still may occur for other products that belong to this category (especially those that are fine, heavy and cohesive). For example, manganese oxide (dV50 < 10 u.m, p s « 5000 kg nr

3, pbi « 1030 kg nr3), which was tested

recently for industry, not only displayed extensive feeding problems similar to that shown in Figure 4.4, but also required high transport velocities to prevent solids deposition and subsequent blockage (i.e. dense-phase transport was not possible). The PVC powder also would be expected to display good dense-phase performance similar to Samples 1 to 8 (i.e. based on the good fluidisation results, free flowing properties, the suggestion of Dixon [23] that dense-phase conveying can be regarded as a moving fluidised bed and excellent air-gravity conveying characteristics [26]). However, of all the materials tested in this investigation, the P V C powder possibly displayed the worst conveying performance due to its inability to be conveyed in the conventional dense-phase [7] or non-suspension mode (as provided by Test Rig B1). That is, as presented previously in Section 3.4.3 (e.g. refer to Figures 3.24 and 3.25), the P V C powder exhibited unstable plugging or blockage conditions in the vicinity of saltation or minimum pressure [7] (i.e. prior to entering the dense-phase regime). Hence, this material was able to be transported only in the dilute-phase mode, which resulted in fairly low values of mass flow ratio (e.g. max. m* ~ 20 kg kg-1). In contrast, for fly ash conveyed on the same Test Rig B1, values of m* « 150 kg kg-1 were obtained quite easily (e.g. refer to Figures 3.10 to 3.16). Although the P V C powder lies very close to the Geldart [24] Group A B boundary (suggesting possible behaviour from either category), these results demonstrate the danger of predicting the suitability of dense-phase based on only fluidisation characteristics. Similar limitations have been observed by Lohrmann and Marcus [22] with three Geldart [24] Group A materials. In contrast, the Group B suggestions of Dixon [23] not only confirm the observed minimum transport behaviour (e.g. pipe vibrations, require high velocities), but also seem to explain the flow behaviour (e.g. dunes grow to fill the pipe causing the high velocity slugs of air to force their way through the material). This is supported further by the results obtained on coarse ash (Sample 11), which was found to display similar problems in dense-phase (e.g. strong plugging, blockages, pipe vibrations). Unfortunately, fluidisation experiments on this material were not able to be carried out (as mentioned previously in Section 5.3.2). The screened coke also displayed similar plugging tendencies (although somewhat stronger than the P V C powder), pipe vibrations and relatively low values of mass flow ratio (e.g. max. m* « 35, for Test Rig A1). Pipeline conveying characteristics have been presented elsewhere [14,16,26] and are reproduced in Figure 5.11, for ease of comparison. Also, transient plots of major conveying parameters have been presented by Wypych and Arnold [16] to demonstrate the plugging nature of this material. From Figure 5.11, it can be seen that for the range of air flows considered, the screened coke did display s o m e dense-phase or non-suspension characteristics (i.e. refer to the two-phase flow diagram [7] shown in Figure 3.3).

106

80

60

Apt

(kPa)

40

20

, Blockage Conditions

.02

Blockage Boundary

Air Only 0.2^" (kgs-1)

.03 .04 .05 .-l" mf (kg s )

Figure 5.11 Pipeline conveying characteristics of screened coke [14,16,26] for L = 25 m & D = 52 m m (Test Rig A1).

107

This could be explained by the Group D behaviour suggested by Dixon [39] (e.g. dense-phase is possible but at relatively low values of m*). However, the natural formation of slugs described by Dixon [23] did not occur and in fact, the plugging which was observed for this material was quite strong (e.g. in relation to the P V C powder) and considered unreliable. Perhaps the relatively wide particle size distribution (and hence, possible Group B behaviour) could explain this apparent discrepancy. Also, the coarse shape of the coke particles may have contributed to this effect. Note that Dixon [39] based most of his suggestions and observations on plastic powders and granules (i.e. smoother particles and relatively narrow size ranges). Nevertheless, the screened coke generally was found to be a better dense-phase material than P V C powder (a typical Dixon [23] Group B material). Another interesting comparison was made recently when both Samples 9 and 10 were conveyed on Test Rig F3 (viz. 0.113 m 3 plug-phase blow tank, L = 161 m & D = 105 m m ) . Sample 9 (PVC powder) displayed both unreliable and unrepeatable characteristics (e.g. the max. operating pressure varied between 150 and 560 kPag). However, in contrast, Sample 10 (screened coke) produced good, reliable operating conditions. For example, maximum pbt ~ 250 kPag, maximum mf« 0.085 kg s"1, average m* « 48 kg kg-1 (based on 2.3 kg of air required to convey 110 kg of product) and an average plug velocity of « 3.1 m s_1. In fact, based on the success of these results, an existing vacuum pneumatic conveying system, which was being subjected to excessive rates of erosion, was replaced by a parallel plug-phase conveying system [26]. Hence, of the two materials, the screened coke was found to be better suited to the plug-phase mode of conveying. This may be explained by the Dixon [39] suggestion that Group D products produce axisymmetric slugging (viz. stable full diameter slugs) and Group B asymmetric slugging (viz. weak slugs and duning). Also, it should be noted that these screened coke results indicate that although Group D materials may not perform as suggested by Dixon [39], it is possible that such materials still could be conveyed successfully and efficiently in the plug-phase mode. Examples of materials that fit into this category include crushed coal (based on several tests undertaken for industry), crushed bath (e.g. refer to Section 4.3.3), sub 20 m m blue metal and diamond ore aggregate [42]. Otherwise, it is possible that such materials could be conveyed only in dilute-phase, which would result in high energy consumption and excessive system erosion (e.g. pipelines, bends). That is, these materials even may prove to be too coarse for specialised pipeline techniques involving say, by-pass technology [2]. On the other hand, typical Group B materials like P V C powder and alumina (e.g. dvso ~ 80 urn, ps = 4000 kg m"3) may be well suited to such by-pass conveying systems or other low-velocity techniques [2] (i.e. to avoid the high velocity gas slugs breaking or forcing through the powder, as suggested by Dixon [23], and also prevent the formation of long plugs of material and hence very high pressure drops and possible blockage).

The apparent anomaly between poor dense-phase performance and good fluidisation characteristics (as suggested by the P V C powder results) seems to be explained by the property of deaeration. That is, although this material displayed good fluidisation characteristics and is classified as Geldart [24] Group A, it was found to deaerate quickly. For example, on several occasions the expanded bubbling bed was seen to lose its height in less than one second (i.e. after the air supply valve was turned off). A similar property was observed for the coarse ash, although a proper fluidisation test was not able to be carried out due to insufficient sample. On the other hand, the pulverised coal and fly ash samples, which fluidised well (viz. Samples 1, 3 and 4), all were found to retain their aeration for considerable lengths of time (e.g. 10 to 30 minutes).

108

The empirical classification diagram proposed by Mainwaring and Reed [38] for the purpose of classifying bulk solids, emphasises the importance of permeability (obtained from a fluidisation test) and deaeration. Jones et al. [43] came to a similar conclusion and suggested further that the ratio of tapped to poured bulk density provides a good indication of the air retentive properties of a given material. Also, Geldart [25] proposed a similar ratio to distinguish between Group A and Group C powders. Hence, there seems to be sufficient evidence to suggest that powder classification (viz. to select ultimately the most suitable mode of conveying for a given product and its behavioural properties) depends on the following properties. (a) Particle size and density. (b) Particle size distribution. (c) Particle shape or sphericity (as indicated by the definitions of diameter). (d) Deaeration and permeability. (e) The ratio of tapped to poured (or perhaps fluidised) bulk density. Also, it seems that most of these properties are interdependent. For example, deaeration and permeability [38] (and perhaps the bulk density ratio [43]) seem to provide an adequate mechanism to detect changes in material performance due to different (a), (b) and/or (c). However, possibly the greatest disadvantage or limitation of the classification techniques proposed by Mainwaring and Reed [38] and Jones et al. [43], is the need to standardise the experimental apparatus and/or techniques. For example, the measured values of deaeration rate [38] depend on the size of the plenum chamber and to some extent the type of gas distributor. Also, different devices and techniques are available to determine vibrated or tapped bulk density. Standardisation is necessary so that the results will be applicable on an international level and can be used/compared by other researchers. In conclusion, as long as one is aware of the limitations of the Dixon [23,39] slugging diagram (similar to those found for the Geldart [24] fluidisation diagram), this technique still is able to provide a good initial indication of what to expect when a given material is conveyed in the dense-phase mode. In fact, it is believed that possibly the one main factor that will upset the suggestions of Dixon [23,39], is particle size distribution. This has been supported by results and experience (e.g. refer to the screened coke, crushed coal and crushed bath considered previously in this section) and the dilute-phase results observed by Mainwaring and Reed [38] for slate dust and pulverised fuel ash grits. Eventually, a standardised technique (similar to [38] and [43]) based on easily measured and relevant properties (i.e. instead of particle size, shape and distribution) is required for an accurate classification of bulk solids. Also, such techniques should be modified or extended to consider all possible modes of conveying (e.g. dilute-, dense-, plug-phase, low-velocity, by-pass conveying, and so on). However, it is believed that for many years to come, the final decision in relation to a given material, specification and method of transport will rely on conclusive and comprehensive experimentation (i.e. using a large-scale test rig), as well as accurate predictions of operating conditions.

109

CHAPTER 6

110

6. SCALE-UP CONVEYING CHARACTERISTICS

6.1 Introduction

The following four main areas of investigation usually require a knowledge of pneumatic conveying characteristics for a given material and pipeline configuration.

General System Design

When it is necessary to design or evaluate a proposed pneumatic conveying system, it is recommended strongly that the designer obtain as much information on the material as possible. With a knowledge of the product's conveying characteristics for various configurations of pipeline, it is quite a simple task to determine the blower or compressor rating (including operating pressure), the optimal pipe diameter and any other components that are dependent upon the operating conditions (e.g. filter size). System Optimisation

Conveying characteristics also may be used to investigate operational problems that a particular plant may be experiencing. For example, a reduced solids throughput simply may be the result of inefficient operating conditions. Also, problems of product degradation and pipeline wear may be reduced easily by determining an optimum value of mf for a given ms. In other words, product conveying characteristics will determine whether an existing plant is operating at an optimum condition. If not, they will indicate what modifications are necessary to achieve the desired result. Upgrading of an Existing Plant

In the event of an existing plant being upgraded to, say, a higher solids throughput, it is necessary to determine whether the system and the material will be able to cope with the increased pressure and/or air requirements (viz. whether the combination of pipe size and compressor rating is sufficient). The pipeline conveying characteristics for the material in question will provide useful information for this purpose. Feasibility Study

Before selecting pneumatic conveying as the method of transport for a particular application and material, the design engineer usually is required to undertake a feasibility study. Not only will conveying characteristics fulfil this purpose, but also they will indicate optimal operating conditions and disclose any unforeseen operational problems due to the material (e.g. PVC powder, which was found to be unreliable in dense-phase, as shown in Sections 3.4.3 and 5.5.2).

For reasons of confidence and reliability, the following procedure is recommended to meet the requirements of the four applications mentioned above.

(a) Conduct a series of pneumatic conveying experiments on a representative sample of the material in a test rig of appropriate size.

111

(b) Determine the steady-state pipeline conveying characteristics [11] (i.e. based on the test rig configuration).

(c) Scale up the test rig data to meet the requirements of a given specification.

(d) Modify the recommended steady-state operating conditions to allow for any relevant transients (e.g. blow tank filling, initial pressurisation, pipeline purging) and the mode of operation (e.g. single batch or tandem blow tanks).

As indicated in Chapter 3, there are two distinct advantages derived from obtaining test rig pneumatic conveying characteristics of a given product.

1. The information is accurate and applicable directly to the test rig employed.

2. By adopting the standardised-test procedure [21], minimum transport boundaries and any unforeseen operational problems also may be identified. In certain cases (e.g. P V C powder [21]), such effects if not allowed for in the initial design calculations would be detrimental to the operation of a plant.

The scaling-up of test rig data is considered to be the most important stage of the design process, in that it provides the necessary link between laboratory-scale apparatus and full-scale industrial installations. Hence, accuracy and reliability of scale-up predictions are essential. This section of work investigates the macro or systems approach to scale-up, where the conveying characteristics for the total pipeline are scaled to an existing or proposed pipeline configuration and then modified to allow for any differences in vertical lift, number and/or type of bends. Investigations into an alternative technique, which is considered more as a micro approach and designed to analyse a pipeline section by section (viz. using a solids friction factor correlation), are presented in Section 8.3. Note that Chapter 8 presents a description of the various mathematical models (including the solids friction factor correlations) which were reviewed and developed in thesis and also includes results and worked examples, where relevant. The scale-up procedures (i.e. scale-up of ms with respect to pipeline length and diameter), which were used initially in this section of work, have been based on the following criteria proposed by Mills et al. [15] and Mason et al. [28].

for a constant mf and A p s (6.1)

for a constant mf D 2 and A p s (6.2)

where subscript 1 refers to the test rig and 2 to the actual or proposed pipeline.

Length :

ms2 = rns1 -j— L2

Diameter :

(D2f ms2 = msi TT

112

Note that A p s remains unchanged during scale-up (i.e. Apsi = Ap s 2) but the corresponding value of Apt2 must be modified to allow for a change in the air-only

pipeline pressure drop, Apf. The results presented by Mills et al. [15] seem to indicate the following approximate relationships.

Length :

Apf2 =

Diameter :

Apf2 =

APfi -1

for a constant mf

A °1 for a constant mf D -2

(6.3)

(6.4)

Combining Equations (6.1) to (6.4) together with their stated conditions, the following generalised scale-up equations have been derived and presented recently [33].

mt2 =

ms2 = msl I-

L2

(D9\2

K^j

(6.5)

(6.6)

Apt 2 = Apt1 - Apf1 + Apf2 (6.7)

where

Apf2 =

and

A M APfiT-

L2

' D ^

Apf1 = xm,^ for mf1 < m f m 1

(6.8)

(6.9)

Note that Equation (6.9) is an empirical expression relevant to the test rig used. Equations (6.5) to (6.8), which will be referred to as the original scale-up equations, have been used quite extensively in several initial investigations (e.g. feasibility studies, troubleshooting, system uprating and general design). However, the equations and their predictions frequently have been found to contain the following inaccuracies and limitations.

(a) The values of ms2, as predicted by Equation (6.6) and when compared with available experimental data, are found quite often to be extremely conservative (i.e. too low) for a given value of mf2 and Apt2.

(b) The values of Apf2, as predicted by Equation (6.8), are found occasionally to be inaccurate.

113

(c) N o allowance during scale-up is m a d e for the relatively longer lengths of vertical pipe that are used usually in industry.

(d) The current method [15] used to scale up minimum transport boundaries [21] seems to be inadequate.

(e) The scale-up equations do not allow for different numbers and/or types of bend (i.e. between the test rig and actual plant).

The present work examines the above five main limitations of scale-up and presents results obtained from suggested areas of improvement (especially in relation to (a), (b) and (c)). It does not seek to provide a complete and proven theory on scale-up, rather it encourages other researchers to pursue similar lines of investigation for their own purpose and verification, so that perhaps in the near future a reliable and unified scale-up model may be available to engineers involved in the selection and design of pneumatic conveying systems. 6.2 Scaling Relationships

The scale-up of experimental data inevitably is required for the design-of pneumatic conveying systems and should involve the consideration of several aspects which could affect overall accuracy (e.g. vertical lift, number and type of bends, pipeline length and diameter, air-only pressure drop).

However, the present study is concerned mainly with the investigation of scale-up with respect to both length and diameter. Generally, previous results have indicated that whenever the diameter effect is involved (i.e. D 2 > Di), the prediction of m s 2 via Equation (6.6) is extremely conservative (i.e. m s for a given rrif2 and Apt2 is found frequently to be much lower than the actual value). Although conservatism may be a desirable feature for design, it also may preclude for a particular application the ultimate selection of pneumatic conveying as the most appropriate means of transportation (especially when long distances and/or large throughputs are involved).

The first aspect considered is the origin of the scale-up criterion [15], where ms a D2

for constant mf D*2 and Aps. This is examined by modifying two popular forms of definition for Aps. Note that Equations (6.5) and (6.7) are a direct result of the conditions specified in Equations (6.1) and (6.2). Hence, modifications to Equation (6.6) are sought to accommodate these conditions/requirements (i.e. generally, Aps

and mf D~2 are constant during scale-up with respect to both diameter and length).

6.2.1 Definitions for Aps

Two popular expressions for Aps are used in the literature. The first results from an analogy to the Darcy-Weisbach friction factor equation, where the pressure drop due to solids may be expressed as

Aps = *• yjL (6.10)

s 2D

114

Using a simulated continuity equation

vs

4ms

w Pb D 2

Equation (6.10) may be rewritten as

Ap s

8 ̂ L m s2

K2 pb D5

(6.11)

(6.12)

In order to use Equation (6.12), an expression for pb is required. From the definition for the volumetric concentration of solids (per unit time)

c< = (1+^F)1 (6-13)

and assuming that this may be approximated by (pf ps-1 m*), Equation (6.14) is

obtained.

Pf ms

pb - ^r <6-14) mf

The percentage error difference between values of cv calculated by Equation (6.13) and the approximation is less than 5 % for pfm* < 100. Hence, in most cases, Equation (6.14) is considered as a reasonable approximation. The substitution of Equation (6.14) into Equation (6.12), results in the following expression for Aps.

8 X,. L mf 11% A P s = -^| L-i (6.15)

TC pf Db

Applying the scale-up condition Apsi = Aps2 and Equation (6.5) to Equation (6.15) yields

m s 2 = m *»1 Pf2 L, (£2 Y

S1 ^s2 Pf1 L2

(6.16)

Another popular expression used for A p s is based on the following definition of Barth [44].

(6.17) Ap s =

Applying the

vf =

A.sm*pfVf2L

2D

continuity equation

4 mf

TC pf D2

(6.18)

115

and the scale-up conditions for Ap s and mf D~2 to Equation (6.17), results in

.3

m S2 = ms1 hi fi2 i± f£i>

^s2 Pf 1 L2 V D1 >

(6.19)

which is identical to Equation (6.16).

The ratio (pf2 pfr1) will be affected by scale-up with respect to both length and

diameter. However, the friction factor ratio (Xsi Xs2-1) depends primarily on the pipe

diameter D. By using appropriate values of Apt2 obtained from, say, Equation (6.7), corresponding values of (pf2 pf-f

1) may be determined for use in Equation (6.19).

However, a reliable expression for Xs is still required before any calculations and/or future evaluations of m s 2 may be carried out. The following section presents results obtained by comparing these equations with other published empirical expressions for Aps.

6.2.2 Empirical Relationships

Empirical expressions for Xs and/or Aps were sought in the literature, with the intention of producing an expression for m s 2 of the general form

m s2 = m r\-<\ \^ZJ

C n

(6.20)

so that direct comparisons with Equation (6.6) may be made. As Xs largely is diameter dependent, length will be taken constant (i.e. L| = L2) for the remainder of this section. The first expression considered is obtained from Weber [45], where

= 2.1(m*)-°-3Fr-2Frfi0-5[£

-0.1 (6.21)

where Fr = Vf (g D) - 0 - 5 and Frs = v«, (g D)'0-5. Note the different form of definition for

Froude number (i.e. to the one used by Weber [45]).

Using the continuity Equation (6.18) and applying the scale-up condition of Apsl =

Aps2. results in the following relationship.

^s1

^s2

ms2

msi

,0.3/ \-2.0

i £li I VPf1>

1.45

(6.22)

After substituting Equation (6.22) into Equation (6.19), Equation (6.23) is obtained.

-1.429/p >2.214

m s 2 = m - (SJ v D i y

(6.23)

116

The second expression is based on the work of W e n and Simons [46], who conducted a series of experiments on glass beads and various coal powders (having particle diameters from 71 to 754 u.m) using 12.7, 19.05 and 25.4 m m internal diameter pipelines. As each length of test section was fixed at 3.048 m, a relationship in terms of diameter will be investigated. Conversion to SI units of the empirical pressure gradient expression presented by W e n and Simons [46], yields

%• - 41.846X10- pbv -(If , i^ (6.24)

where pb = bulk density or dispersed density of the material (kg nr3) vs = solids velocity (m s-1).

Assuming Apt» Aps (due to the relatively high values of mass flow ratio obtained by Wen and Simons [46], i.e. m* = 80 to 750) and after substituting Equations (6.11) and (6.14) into Equation (6.24) and letting L| =L2 )the following equation is obtained.

m s 2 = ms1

„ .-0.55 ,_ ,2.25

Pf2^ (6.25)

The final expression is determined by considering a slightly modified version of the pressure drop equation presented by Ostrovskii et al. [47], where

Aps = km*VfLpfa5D-°-22 (6.26)

where k is a constant which depends exclusively on the properties of the material being conveyed. Applying the previous analyses to Equation (6.26), results in

m s 2 = ms1 | — | | — (8.27)

A comparison of Equations (6.23), (6.25) and (6.27) reveals fairly consistent values for the power index of D 2 D-r

1 (i.e. TJ = 2.214, 2.25, 2.22) but relatively inconsistent

values for pf2 pfi"1 (i-©- £ = -1-429, -0.55, 0.5). These values of TJ support the experimental trends described previously in Section 6.1 and to some extent indicate the inadequacy of Equation (6.6), where the power index of D 2 Di'

1 is 2. However, they are still somewhat lower than the value of 3 suggested by Equation (6.19). The following section presents results obtained from investigations into determining accurate experimental data for the verification of scale-up with respect to diameter, as well as pipeline length. Note that due to the inconsistent values of the power index for pf2 pfi'1 and also as the present study is concerned mainly with the length and diameter effect on scale-up, the index e will be taken as zero for the remainder of these investigations.

117

6.3 Experimental Investigations

Results obtained from three materials (fly ash/cement mix, screened coke and PVC powder) are presented to investigate scale-up with respect to diameter only, length only, and both length and diameter, respectively. Physical properties of the three products are given in Table 6.1. Note that the fly ash/cement mix comprised 89 wt.% fly ash and 11 wt.% cement. Also, note that the P V C powder and screened coke are the same materials listed in Table 5.1 (viz. Samples 9 and 10).

Material

Fly Ash 89 wt.% fly ash Cement Mix 11 wt.% cement

Screened Coke

PVC Powder

Median Particle Diameter

dso (nm)

19 20

470

135

Solids Density

ps (kg nr3)

2130 3100

1940

1400

Loose-Poured Bulk Density Pbl (kg nr3)

700 950

985

575

Table 6.1 Physical properties of test materials.

6.3.1 Fly Ash/Cement Mix

Steady-state pipeline conveying characteristics of the fly ash/cement mix were obtained from Test Rigs C1 and C 2 (refer to Section 2.3) and are reproduced in Figures 6.1 and 6.2, respectively. Applying the original scale-up Equations (6.5) and (6.6) to the data presented in Figure 6.1 (i.e. using L| = 162 m & Di = 0.060 m and L2 = 162 m & D 2 = 0.105 m) results in the 'predicted' conveying characteristics shown in Figure 6.3. Note to minimise errors due to the scaling up of Apfi, experimental values of Apf2 were taken from Figure 6.2 and then simply added to Apsi ( = Aps2) to obtain Apt2-

It is evident from Figures 6.2 and 6.3 that Equation (6.6) significantly underpredicts the values of ms2. In fact, for given values of air flow (> 0.3 kg s

-1) and Apt, it is found that the values of m S 2 are underpredicted by « 33 %, as indicated in Table 6.2.

It is interesting to note that the general shape of the ms curves, when predicted from the Test Rig C1 results (i.e. shown in Figure 6.1), tends to be retained during scale-up, although some change is noticed due to the Apf effect. Hence, not only does this tend to cast some doubt over the validity of scaling up typical dense-phase results to an essentially dilute-phase regime, but also seems to suggest that there is a significant change in flow characteristics. That is, for a given ms, the variation of Apt

with respect to mf in the larger diameter system generally tends to be more linear and indicative of dilute-phase transport.

118

400

300 -

(kPa)

200

100 -

Figure 6.1 PipeJine conveying characteristics of fly/ash cement mix for Li = 162 m & Di = 0.060 m (Test Rig C1).

300

200 Apti (kPa)

100

d: I A.ir Onl^

L .2 .3

mfl (kgs-1)

.5

Figure 6.2 Pipeline conveying characteristics of fly/ash cement mix for Li = 162 m & Dt = 0.105 m (Test Rig C3).

119

300

200

A p t2 (kPa)

100

.1 .2 .3

mf2 (kgs-1)

Figure 6.3 Scale-up conveying characteristics of fly/ash cement mix for L-2 = 162 m & D2 = 0.105 m based Figure 6.1 and Equation (6.6).

Actual ms-| (kg s"1)

10.0 15.0 20.0

m S 2 , Eqn. (6.6) (kg s-1)

6.5 10.0 13.5

Error (%)

-35.0 -33.3 -32.5

Table 6.2 Comparison of predicted and actual values of m s for mf > 0.3 kg s-1 (i.e. based on Figures 6.2 and 6.3).

One possible reason for this may be due to the dispersion of particles (via velocity fluctuation) being more frequent in the smaller diameter pipeline resulting in a more efficient mixing of the two-phase flow. Hence, it is possible that the typical trends displayed by Rizk [7] for dense-phase transport (i.e. m s lines having negative slope on an mf and Apt plot) may be less pronounced for larger diameter pipelines (i.e. where flow instabilities seem to be more dominant). A comparison of the transient plots of some of the major conveying parameters obtained from Test Rigs C1 and C3 indicate support for this hypothesis. However, it should be mentioned that the main objective of this experimental program is to investigate the validity of the scale-up Equation (6.6) and not to examine the effect of minimum transport behaviour (i.e. further detailed studies still are required before any conclusive findings may be presented). Hence, for the present work, comparisons of typically

120

dilute-phase trends only will be considered (e.g. refer to Table 6.2 where results for air flows > 0.3 kg S"1 only were presented). By using a trial and error procedure, the determination of a suitable value for the power index TT in the equation

m s2 ~ m S1 D, (6.28)

resulted in the selection of T\ = 2.8. Application of Equations (6.5) and (6.28) to the test rig data presented in Figure 6.1, produces the scale-up conveying characteristics shown in Figure 6.4. From this plot it can be seen that there is good agreement with the experimental m s l curves shown in Figure 6.2.

300

200

Apt2 (kPa)

100

0 .1 -t— JL_ .2 .3

mf2 (kgs-1)

Air Only

.4 .5

Figure 6.4 Scale-up conveying characteristics of fly/ash cement mix for L2 = 162 m & D 2 = 0.105 m based on Figure 6.1 and Equation (6.28) with T\ = 2.8.

6.3.2 Screened Coke

Steady-state conveying characteristics of the screened coke were presented previously in Section 5.5.2.2 (viz. Figure 5.11, Test Rig A1, Li = 25 m and Dn = 0.052 m). Four additional tests were carried out on Test Rig A 2 (viz. Li = 71 m and Di = 0.052 m ) and the corresponding values of mf, Apt and m s are summarised below in Table 6.3. These results are employed in this section to investigate the accuracy of scale-up with respect to pipeline length.

121

mf

(kg s"1)

0.0405 0.0400 0.0390 0.0385

Apt (kPa)

15.0 42.0 53.0 62.0

ms

(kg s-1)

0.12 0.29 0.33 0.39

Table 6.3 Summary of screened coke results for Test Rig A2 (l_i = 71 m & Di = 0.052 m).

Applying the original scale-up Equations (6.5), (6.6) and (6.7) with Di = D 2 = 0.052 m, to the data presented in Figure 5.11 (using L2 = 71 m), results in the conveying characteristics depicted in Figure 6.5. Note that the minimum transport (or blockage) boundary was scaled up on the basis of Mills et al. [15], where the superficial minimum conveying velocity Vfimjn is assumed to be a function of m*. The experimental data presented in Table 6.3 have been superimposed on this plot for ease of comparison. As can be seen, the scale-up predictions do seem reasonable, although a comparison of experimental or actual m s curves would have been more conclusive. Also, it should be noted that during the scale up of Figure 5.11, no allowance was made for the different number of bends used on Test Rigs A1 and A2 (viz. 5 and 13, respectively). The experimental technique described and used by Mills etal. [48] and Mills and Mason [49] could be employed to obtain an empirical, equivalent length of bend, so that adjusted values of l_i and L2 could be employed in Equation (6.6). 80

60

Ap t 2

(kPa)

40

Predicted Blockage Boundary

20 -

0.3 "

O-^Ckgs-1)

.02 .03 .04 .05 .-i< m f 2 (kgs" )

Figure 6.5 Scale-up conveying characteristics of screened coke for L2 = 71 m & D 2 = 0.052 m (based on Figure 5.11 and Equations (6.5) to (6.7)) with four experimental data points from Test Rig A2 (L| = 71 m & Di = 0.052 m).

122

However, this additional work was not considered paramount to the purpose of the present investigation. Furthermore, l_i and L2 represent effective transport distances [9] and hence, do allow for bends to some extent. Nevertheless, considering a hypothetical equivalent bend length of 4 m, the following evaluation is made.

mS2

whereas

m s2

m s1

m s1

25 m 71 m m S1 0.352 ms1

25 m + 5 x 4 m 71 m + 13x4 m

ms1 = 0.366 m s 1

which would produce only a 4 % increase to the values of m s 2 presented in Figure 6.5. In either case, scale-up in terms of pipeline length is considered reasonable.

6.3.3 P V C Powder

Both Test Rigs A 2 (L-| =71 m & D-| = 0.052 m) and C 3 (L| = 162 m & D 2 = 0.105 m) were used to determine the steady-state pipeline conveying characteristics of the PVC powder. The former characteristics have been presented previously in Figure 3.23 (viz. Section 3.4.3) and the latter are shown in Figure 6.6.

300

200 -

A?ti (kPa)

100

Blockage Conditions

Blockage Boundary

mf (kg s a)

Figure 6.6 Pipeline conveying characteristics of P V C powder for Li = 162 m & Di = 0.105 m (Test Rig C3).

123

Applying the suggested scale-up equation

v2.8

mS2 = msi T~

to the data presented in Figure 3.23, results characteristics shown in Figure 6.7. Note that Apf2

experiment and not predicted using Equation (6.8) boundary was scaled-up on the basis of Mills et between Figures 6.6 and 6.7, it can be seen that the by » 33 %. An obvious reason for this discrepancy bends used on Test Rig A2 (i.e. 13 as opposed to 5 done previously for screened coke, a hypothetical 4 used to determine the extent of variation to ms2. That

(6.29)

in the predicted conveying was calculated once more by . Also, the minimum transport al. [15]. From a comparison m s 2 values are underpredicted would be the larger number of used on Test Rig C3). As was m equivalent length of bend is is,

Li' 71 m + (13x4m)

162 m + (5 x 4 m)

and

m s2 m s1 123 182

ro.iosY*-8

10.052 J

= 123 m

182 m

4.835 ms1

which is « 54 % greater than the value stated in Equation (6.29).

(6.30)

300

200

Apt2 (kPa)

100 -


.1 .2 .3 .4

mf2 (kgs-1)

.5

Figure 6.7 Scale-up conveying characteristics of P V C powder for L2 = 162 m & D 2 = 0.105 m, based on Figure 3.23 and Equation (6.29).

124

Application of Equation (6.30) to the test rig msi values of Figure 3.23, produces the adjusted scale-up conveying characteristics presented in Figure 6.8. Subsequent comparisons reveal that Equation (6.30) still overpredicts the values of ms2', but only by an amount of = 3.0 %. However, despite this apparent improvement there is still one area of concern that warrants further discussion (viz. the scale-up of minimum transport boundaries). It can be seen from Figures 6.6 and 6.8 that the blockage boundary, which was scaled up from the data obtained on Test Rig A2, displays values of Vfimjn lower than were obtained actually from Test Rig C3. Hence, the dependency of Vf>mjn on m*. as suggested by Mills et al. [15], seems to be inadequate and that a diameter effect may be involved. For example, Zenz [50] carried out numerous experiments on a variety of products and found that the single particle saltation velocity VfS0 in a horizontal pipe increases with respect to pipeline diameter according to the following relationship.

V fso (6.31)

where 0.4 < 8 < 0.6. lt is expected that a similar effect also would occur for the saltation velocity under load conditions, VfS (i.e. for a given value of m s ) . This would help explain the differences between the minimum transport boundaries depicted in Figures 6.6 and 6.8. Note that these blockage boundaries for P V C powder seemed to occur in the vicinity of saltation (see Section 5.5.2.2). An alternative approach based on minimum Froude number (viz. Frmjn = Vf>mjn (g D)-°-

5) is considered later in Section 8.3. However, further detailed investigations into the minimum transport behaviour of bulk solids in different pipe diameter systems still are required to test the applicability and accuracy of such relationships (especially for fine powders such as P V C powder, fly ash, cement, and so on).

300

200

APt2 (kPa)

100

ms2 ( k g S -l-


0 .1 .2 .3 .4

m f 2 (kg s )

.5

Figure 6.8 Scale-up conveying characteristics of P V C powder for L2 = 162 m & D 2 = 0.105 m, based on Figure 3.23 and Equation (6.30).

125

6.3.4 Effect of Vertical Lift on Scale-Up

As was mentioned previously in Section 6.1, one of the limitations of scale-up is that it does not allow for the relatively longer lengths of vertical pipe that are used usually in industry. This section briefly investigates and suggests a technique that could be employed to allow (to some extent) for this influence of vertical lift.

The scale-up condition Apsi = Aps2 is assumed valid for an installation containing a total length of vertical lift, which is proportional to that used on the test rig, so that

-v2 = v2

where Lv2' = Lvi (L2/L1). Similarly,

Lh2 = Ln2'

where Lh2' = Lhi (L2/L|).

The pipeline air pressure drop component due to solids may be considered in the general form

Ap s = ^ ( L K + K L , ) (6.32)

where K is the ratio of vertical to horizontal air pressure gradient due to solids. That is,

K fdPs") fdp s

IdL 1 VdL

,-1

(6.33)

Note that initially K will be assumed constant with respect to operating conditions (i.e. mf, Apt and m s ) . Applying the condition Apsi = ApS2 to Equation (6.32) and assuming that the pipeline air pressure gradient (due to solids) does not vary significantly between pipeline configurations having the same L2 but different total lengths of vertical pipe (i.e. LV2 > LV2'), then it may be shown that

Aps 2* = Aps 1 — L2

fUo + KoL,^ -h2 2 Lv2 < Lhi + Ki Ly-i

(6.34)

where A p s 2 * is the value of A p s 2 (which equals Apsi) modified to allow for a relatively longer length of vertical pipe.

Despite the fairly short length of vertical pipe used in Test Rig A2 (i.e. Lv = 3.6 m), measurements of pipeline pressure gradient were taken for fly ash [33] and the P V C powder [51], and K was found initially to be contained in the range 1.7 < K < 2.2. However, these results were obtained for a limited range of operating conditions and on further inspection of the fly ash results, it was found that K approached unity

128

for dilute-phase (e.g. mf > 0.05 kg s*1) and values of 3 to 4 for dense-phase conveying (e.g. mf < 0.02 kg s-1). Hence, it seems that it would be possible for the test material to determine an empirical relationship between K and pertinent operating conditions (e.g. mf, m s and Apt). This then could be used in the scale-up procedures. However, the necessary measurements of (dps/dL)n and (dps/dL)v

must be accurate and hence, should be carried out on a vertical pipe which is significantly longer than those being used currently on the test rigs described in this section of work. This additional effort was not considered paramount to the present study and it is suggested further that in the absence of experimental data, the approximation K1 « K2 « 2 (6.35)

could be used in Equation (6.34) for an initial estimation of Aps. This agrees well with the findings of Reed [52], who also has suggested an average value of K « 2.0 for a variety of materials.

6.4 Scale-Up of Apt

The scale-up predictions according to Equation (6.8) are found occasionally to be inaccurate, especially when relatively large values of L 2 are being considered. Furthermore, as a large proportion of the total pressure drop, which occurs during either dilute-phase or long-distance transportation, is caused by the air-only component, then it becomes necessary to predict values of Apf with reasonable accuracy. Numerous techniques were investigated, and the following equation based on [53] is suggested.

0.0016 Qf185L

Apf = , kPa (6.36) Pfi D 5

where Qf = volumetric flow rate of free air (N m3 s-1) L = length of the pipeline (m) D = actual internal pipe diameter (m) Pfi = initial pipeline air pressure (kPa abs).

For the case of pneumatic conveying and single-diameter pipelines, where the outlet or final pipeline air pressure (viz. Pf2) usually is atmospheric, Equation (6.36) may be rewritten in the form

Apf = 0.5[(1012 + 0.004567mf

1-85LD-5)°"5-101] , kPa (6.37)

for Pf2 = 101 kPa abs and T = 293.15 K. Note the following units required for Equation (6.37): mf (kg s"1), L (m) and D (m). To evaluate the accuracy of Equation (6.37), constant values of Apf were calculated for corresponding values of (L D'5 x

10"7) and mf as shown in Figure 6.9. Four to five experimental values of Apf also were determined from each of the six pipeline configurations outlined in Table 6.4 and superimposed onto Figure 6.9.

127

3.

1.

.3 -

m. f (kgs"1)

.03 -

.01 10. 30. 100.

m D5 IO7 I in5

Figure 6.9 Variation of Apf according to Equation (6.37) with experimental data points obtained from six different pipeline configurations.

128

Test Rig No.

A1 A2 A3 C1 C3 C4

L (m)

25 71 96 162 162 553

D (m)

0.052 0.052 0.052 0.060 0.105 0.068

L D'5x 10"7

(m nr5)

6.58 18.67 25.25 20.83 1.27

38.03

Coeff. x Eqn. (6.9)

293.34 803.55 1507.86 738.70 56.23

1783.73

Coeff. y Eqn. (6.9)

1.513 1.525 1.623 1.555 1.233 1.610

mfm (kg s"1)

0.10 0.09 0.10 0.17 0.45 0.15

No. of Bends

Nb

5 13 l 17 5 5 17

Table 6.4 Empirical expressions for Apf.

Despite the significantly different number of bends used on the test rigs, there is still good correlation between the predicted and experimental values of Apt.

Figure 6.9 also could be used to predict Apf for each section of a stepped diameter pipeline, which quite frequently is used for long-distance transportation. However, the final values would tend to be conservative (i.e. too high for a given mf) because Equation (6.37) would assume that the final air pressure of each pipeline section is atmospheric. Hence, in this case, the following slightly modified version of Equation (6.36) is suggested.

Apf = 11418.3 m f

1 8 5 L

Pf1 D5x107

, kPa (6.38)

Again note the following units: mf (kg s"1), L (m), D (m) and Pfi (kPa abs). Using Equation (6.38), the analysis procedure consists of the following steps.

(a) Assume Pfi at the start of the stepped pipeline. (b) Estimate Apf using Equation (6.38) for the first or initial diameter section. (c) Calculate Pfi for the second diameter section. (d) Repeat until the last pipeline section is reached and calculate the exit or

final air pressure Pf2 of the stepped-diameter pipeline. (e) Repeat steps (a) to (d) until atmospheric pressure is reached in step (d)

(e.g. Pf2 = 101 kPaabs).

To investigate the accuracy of Equation (6.38) for this application, the above analysis procedure was applied to the 940 m long pipeline of Test Rig D2, which comprises four different pipe diameter sections, as summarised in Table 6.5.

The variation of Apf with respect to mf was determined experimentally, where

Apf = 1819.0 mf 1.55 for m f < 0.15 kgs

-1 (6.39)

Values of mf = 0.10 and 0.15 kg s'1 were selected for this investigation and the

corresponding Apf values are summarised in Table 6.6.

129

L (m)

146 390 261 143

D (m)

0.060 0.069 0.081 0.105

No. of Bends Nb

3 13 8 5

Table 6.5 Long-distance pneumatic conveying pipeline (Test Rig D2).

mf (kg s-1)

0.10 0.15

Experimental Apt, Eqn. (6.39)

(kPa)

51.3 96.1

Theoretical Apt, Eqn. (6.38)

(kPa)

61 107

Error

(%)

+19 +11

Table 6.6 Comparison of experimental and theoretical values of Apf for the long-distance pneumatic conveying stepped diameter pipeline (Test Rig D2).

Although the maximum error shown in Table 6.6 is almost 20 %, the predicted values of Apf are only ~ 10 kPa greater than the experimental values. For a stepped-diameter pipeline almost 1 km in length, this result is considered quite reasonable.

Note that the analysis procedure described previously (viz. to predict Apf using Equation (6.38)) may be simplified significantly by replacing the atmospheric pressure (viz. 101 kPa abs) in Equation (6.37) by the final air pressure (viz. Pf2), as shown in Equation (6.40).

Apf = 0.5 [(Pf22 + 0.004567 m,1 '85 L D^5)05 - Pf2] , kPa (6.40)

The analysis procedure commences with the final section of pipeline (i.e. Pfi = Pf2 + Apf = 101 + Apf kPa abs) and then proceeds upstream until the initial or starting section of pipe is reached. In this way, iterative calculations are not required.

6.5 Summary

The experimental results obtained and presented on the fly ash/cement mix and PVC powder indicate that the original scale-up Equation (6.6) is inadequate when data are scaled up with respect to pipeline diameter. In conjunction with results obtained from the screened coke, the scale-up equation

-*' = m«u{%Y (641)

130

is suggested as an improvement, where Li' and L2' represent adjusted values of l_i and L2 to allow for any differences between the number (and type) of bends used on the test rig and actual plant.

Equation (6.34) also should be used to modify the values of Aps2 if Lv2 > (L2 Lr1)

Un (i.e. if the total length of vertical lift for the actual or proposed plant proportionally is greater than that used on the test rig). In the absence of experimental data, the pressure gradient ratio K-i « K 2 - 2 could be used as an initial estimation for use in Equation (6.34).

If prediction of Apf is required (e.g. to determine Apt = Apf + Aps), then it is suggested that Figure 6.9 or Equation (6.37) be used for conventional single-diameter pipelines, and Equation (6.40) for stepped-diameter pipelines.

6.6 Generalised Pipeline Conveying Characteristics

For the purposes of general design and feasibility studies, the construction of multiple scale-up conveying characteristics (i.e. for different values of L and D) is time-consuming and tiresome. Hence, it would be advantageous to either computerise scale-up procedures or develop generalised conveying characteristics (for a particular material), which would be applicable to any system of different L and/or D (provided the extrapolation is within acceptable limits).

Interpretation of the recommended scale-up Equations (6.5), (6.41) and the condition Apsi = Aps 2 > indicates that the coordinate system mf D"

2 (abscissa) and Ap s (ordinate) could be used to represent the parameter (ms L' D"

2-8) for the purpose of generalising pipeline conveying characteristics (obtained on a particular test rig).

The subsequent conversion of the test rig data displayed in Figures 6.1 and 6,2 for the fly ash/cement mix, results in the generalised conveying characteristics displayed in Figures 6.10 and 6.11, respectively. Note that both L|' and L2' have been assumed equal to 162 m for these calculations. Either graph can now be used to predict values of A p s (for a given ms, L' and D) or estimate suitable values of D (for a given m s a n d L').

For example, consider the requirement to estimate the smallest value of D that would transport the fly ash/cement mix at a rate of 151 rr1 (i.e. 4.167 kg s'1) over a distance of 620 m for an operating pressure not exceeding 200 kPag. From Figure 6.11, an initial value of (ms L' D-2-8 x 10'

5) = 20 is selected, from which D = 0.093 m is calculated. Note that for the purpose of this example, no allowance has been made for bends (i.e. L' has been assumed equal to 620 m).

131

250

200 -

150 _ Ap

rs (kPa) 100 -

mfD -2 (kg s-V"2)

Figure 6.10 Generalised pipeline conveying characteristics of fly ash/cement mix based on Test Rig C1 results (L|' = 162 m & Di = 0.060 m).

Ap rs

(kPa)

250

200 -

150 -

100 -

50 -

10

-

-

.

1

1

1

m L s D2'8 105

•

1

/kg s-1m\

\ -2'a i 1

1

_20

_ — 15

10

5

/

»

-

-

m D"

20 30

(kg S^m"2)

40 50

Figure 6.11 Generalised pipeline conveying characteristics of fly ash/cement mix based on Test Rig C3 results (LT = 162 m & Di = 0.105 m).

132

Selecting a 100 m m N B Schedule 40 (i.e. D = 0.102 m ) mild steel pipeline, the following corrected value of Aps is obtained.

msL' — — = 15.4, Aps « 145 kPa (Figure 6.14) D 2 8 x 1 0 5

for a minimum safe value of mf D"2 = 20 kg s-"1 nr2. Therefore, mf = (20) (.102)2 = 0.208 kg s_1. The corresponding value of Apt may then be determined by predicting Apt from either Figure 6.9 or Equation (6.37). That is, for (L D~5) = 5.616 x 107 and mf = 0.208 kg S'1, Apf is estimated at 27 kPa, so that Apt = 172 kPa.

Not only does the above example demonstrate the usefulness of generalising pipeline characteristics, but it also indicates that this procedure may be used effectively in future investigations on other materials to further evaluate and improve the accuracy of the scale-up equations presented in this thesis.

133

CHAPTER 7

134

7. THEORETICAL ANALYSES

7.1 Introduction

The problem of mathematically modelling the pneumatic transport of particles in the dilute-phase mode for the purpose of predicting pipeline air pressure drop (for a given set of operating conditions), has been the subject of numerous investigations over the past two decades [54-58]. However, empiricism has been used widely and a unified theory applicable to all materials (especially fine powders) and the numerous configurations of pipeline has yet to be formulated. Theoretical predictions for dense-phase pneumatic transportation have been few in number and the majority have been concerned only with straight and relatively short horizontal pipes [17,59,60]. Generally, the applicability of these models to industry is very limited and is reduced further for materials possessing small particle size (typically less than 40 u.m), relatively wide particle size distributions (e.g. 10 to 300 urn) and/or complex physical properties (e.g. binary products). This section of work is aimed at reviewing and developing mathematical models to predict system design parameters and verifying these predictions with experimental data. Particular objectives include presenting the analytical and numerical results which were obtained from

investigations into modelling the discharge characteristics of blow tanks, reviewing various existing pipeline theories to predict pressure drop for

the dense-phase mode of transport, developing new design strategies to predict more accurately the

operating conditions for single- and stepped-diameter pipelines (based on the correlation analysis of experimental data) and optimising the configuration of long-distance pipelines.

7.2 Blow Tank Discharge Characteristics The mathematical models developed by McLean [61] to predict the discharge rate of powders from mass flow hoppers are extended to simulate the flow from a pressurised blow tank. In formulating the governing equations, McLean used the concept of a simplified bulk solid flowing through an Enstad [62] element of a converging channel, as shown below in Figure 7.1. The resulting differential equations (applicable to an axisymmetrical flow channel) consist of: 1. Continuity Equation of the Bulk Solid dvs 2 vs vs dpb

_ • + — l + -±-p. = 0 (7.1) dr r pb dr

2. Interstitial Air Continuity Equation

135

Figure 7.1 The Enstad [62] element of a converging flow channel.

from which the interstitial air pressure gradient equation

4+dp(! + ldoV_^(YYi) . „ d r2 dr vr c dry p s c dr v

was obtained assuming Darcy's Law

dp U = ~ d r

Ph

and using r = 1 - — . Ps

3. Equation of Motion for the Bulk Solid

PbVs dv£ 1 -sin8 dr

da Xi a dp 1 ~z + Pb 9 Yi + rr z =-? dr r ™ w i dr 1 - sin8

(7.3)

(7.4)

(7.5)

136

7.2.1 Approximate Analytical Solution

Assuming a static consolidation stress distribution, the above differential equations were applied to the case of a bottom discharge blow tank (pressurised from the top) and the following equations for vs, vso and mSo were obtained.

'so

m so

where

Ai

T4

\N,

2-b

i1-b

" v$o w,

2A, A3 A!

0.5

2Ai

K _ 2 4 D o Pbovso

(7.6)

(7.7)

(7.8)

rw Pbo

T 3

C

- Gi r0 Tlln f, rT\ ^(Zi-D

T1T4(Z1-1)(1 -sinS)

PboSroYi (1-b) ZiT^pj-Po) + XÎ-W-1 T4(Z1-1)(1-sin5)

a + b-1 (1-b)

3b-4 1 -b

T2-X

Ti-X

2-b (1 - b) (1 - sinS)

'o

VrT>

Pbo = Pbc flo

C, PUo,

(7.9)

(7.10)

137

CT1o

Gi

C

Xi

Yi

YY!

=

=

=

=

=

=

o0 (1 + sinS)

Pbob(YY1)

(1 - b) ps c0

1 1 - sinS

2 sinS 1 - sin5

sin(a+2p)~ sina

2 [1 - cos(a+(3)] sina + sinp* sin2(a+p)

(1 - sinS) sin3(a+p)

Y! (1 - s 2[1-

inS) sin2(a+p)

cos(a+p)]

By integrating Equation (7.3) with respect to r, the following general expressions for the air pressure and the air pressure gradient inside the blow tank were derived (i.e. for the case of a static consolidation stress distribution).

d v 1 vso 'o rVi

T^-DUo ^W-M-M PT -Pc <-(*f

Zi-1 (7.11)

dp dr

Gi v 1 vso r V1"1

- T

VZt-DU PT-PO

Ziln — +ln — + (Zi-D

(Zi-DrJUT rfi (7.12)

Note that the above equations are similar to the ones derived by McLean [61], for the case of a bin discharging under gravity, except for the additional pressure drop terms appearing in Equations (7.11), (7.12) and in the expression for the coefficient A3 given in Equation (7.7).

7.2.1.1 Results

For a given material, blow tank geometry and transition pressure, values of vso and mso were able to be calculated by adopting the following procedure.

1. Calculate the material parameters Xi, Yi, YY1 and C.

2. Assume vso = 0.25 m s-1 and pt>o = Pbc-

138

Calculate the corresponding value of o 0 using Equation (7.13).

Pbo C vso ,_, „ . o"Q = 2 (7-13)

3. Calculate Ti, T2, T3, T4, W1 and Z1.

4. Set up an iteration loop by determining pbo and c0 from Equations (7.9) and (7.10).

5. Calculate Gi, Ai, A2, A3 and the new value of vso according to Equation (7.7).

6. Recalculate o0 with this new value of vso and using Equation (7.13).

7. Compare the old and new values of o0. If the difference is greater than 0.1%, go to Step 4 to continue the iteration procedure. Otherwise, calculate m So using both Equation (7.8) and the final value of vso obtained from Step 5.

The above procedure was incorporated onto the University's mainframe computer (using Fortran 77) and enabled the variation of ms 0, with respect to D 0 and ApDt = P T - po» to be predicted for a given a, as shown in Figure 7.2. Note that the parameters a = 20° and D j = 0.917 m were selected to simulate the Sturtevant™ blow tank geometry shown in Figure 2.1. Also, the material properties, which were used to generate this plot, were obtained from the Tallawarra pulverised coal sample (refer to Table 5.1) and consisted of

Pbc

cic b c

0"1p a

8

0 Ps

=

=

=

=

=

=

=

=

=

832.4 kg m *

6020.57 Pa 0.0351 11.9 x 10"9 m4N-l s-

1

6020.57 Pa 0.2036

40°

25°

1600 kg nr3.

7.2.1.2 Discussion

Several alternatives to the static consolidation stress assumption, which was used to formulate the equations (7.6), (7.7), (7.11) and (7.12), also were investigated (e.g. a linear decrease in p from the transition to the outlet; constant pb and c with respect to r). The subsequent detailed formulations were postponed until an accurate set of experimental data could be obtained for the blow tank. However, due to the following reasons, this did not eventuate during the course of this thesis project.

1. The necessary blow tank parameters such as p and vs0 were extremely difficult to measure. For example, the variation of p with respect to r would have had to been recorded for relatively low values of pressure drop over

139

a short distance of approximately 1.3 metres. Results from the fly ash test program seemed to indicate that Apbt occurs in the range 1 to 10 kPa, which would be approximately 1 to 5 % of a 200 kPag typical operating pressure (blow tank top-air). A direct measurement of Apbt was attempted without success, mainly due to the inappropriateness of the pressure tapping design (similar to that shown in Figure 2.4) to monitor accurately small differences in air pressure. Other designs were investigated (e.g. an open pressure hole with an in-built purge system) but also were found to be unsuitable.

2. The demands of the fly ash test program, which commenced effectively in October 1983, generally reduced the emphasis on the blow tank investigations.

3. The pneumatic conveying test results obtained during the fly ash program also revealed that the blow tank top-air only conveying mode (which is required for the application of the mathematical model) provided insufficient air for the transport of this material over the test rig distance of 71 m. Hence, not only did this prevent any 'parallel' investigations to be carried out, it also initially threw some doubt over the practical usefulness of such a model. However, it was later realised that continued effort in this area still could ultimately lead to a unified blow tank theory, which also would allow for the various types of configuration [63].

4. Due to the requirements of the fly ash test program and industry in general, it was decided to concentrate the theoretical investigations on the development of a reliable scale-up model to predict plant operating conditions based on test rig data (e.g. refer to Chapter 6) and correlation analysis to design and optimise long-distance stepped-diameter pipelines (considered later in Section 7.3).

Although a thorough assessment of the mathematical model was not able to be completed due to the lack of experimental data, approximate values of Apbt were calculated for a number of experiments, which were carried out during the fly ash program (i.e. using the extrapolation technique depicted in Figure 3.20). After comparing the corresponding pipeline values of m s with the m s o predictions shown in Figure 7.2, the results seem to indicate that the model underestimates the m s o

values. 7.2.2 Numerical Analyses

The numerical solution of the three governing differential equations (viz. (7.2), (7.3) and (7.5)) was investigated using the Runge-Kutta technique of numerical integration. A number of software programs were written on the Univac computer to predict the parameters vs, p, dp/dr and a at various levels inside the blow tank (based on a given set of initial boundary conditions at either the outlet or the transition, depending on the direction of integration). However, occasional convergence problems, which occurred during the integration procedure, prompted a more detailed investigation of the actual equations and the initial assumptions which were used. Several modifications and improvements to the numerical model were introduced and a summary of the latest version is presented below.

140

20.0

10.0

2.0 -•

1.0

0.2

0.1

Figure 7.2 Example of blow tank model results (approximate analytical solution).

The solids continuity equation remained unchanged (i.e. as presented in Equation (7.1)).

dvs _ 2v s vs dpb

dr r pb dr

The following modified air continuity equation was formulated

(7.14)

du dr

by expanding

— (pf vs r As

2u r

u + vs

P r dp

dr dPb dr

the differential equation

+ Pf uAs) = 0

vs

Pb (7.15)

(7.16)

where As is the surface area of the Enstad element (m2)

P = p+Patm = absolute air pressure (Pa abs).

The variation of p with respect to r was assumed to be governed by the following Ergun [64] equation.

$• = Evu + Eku2 (7.17)

dr v K

E 150uf(1 - D 2

V ~ H 2T-3 QV50 ~

1.75 P(1 -D k " RTd^T3

Note that the median particle size dvso (based on a volume diameter distribution, see Section 5.2.1) was used to represent the material in these equations.

The original compressibility equation for pb (viz. Equation (7.9)) also was modified to allow for zero values of o (i.e. at the top surface of the material in the blow tank).

Pb = PblO+<T)b (7'18)

where pbi is the loose-poured bulk density of the material (kg m-3).

Note that, due to the use of the Ergun [64] equation (instead of Darcy's Law), the permeability coefficient c of the bulk solid is not required for this analysis.

142

5. The equation of motion for the bulk solid essentially remains unchanged (as presented in Equation (7.5)) except for the differences introduced by the modified compressibility equation.

da dr

r -pbgYi-

1

1 dp [ 2 p b v s2

1 - sinS dr ' r (1 - sinS)

Pbbvs2 (7.19)

(1 + a) (1 - sinS)

6. The variation of pb with respect to r, which was also required for the application of the Runge-Kutta technique, was obtained by simply differentiating Equation (7.18).

^ = 7^£ (7.20) dr 1 + G dr

7.2.2.1 Results

Due to similar reasons stated in Section 7.2.1.2, only preliminary results were able to be obtained from the numerical model. For example, refer to Figure 7.3, which presents typical variations of vs, p, a and pb (with respect to r/rj) obtained for Tallawarra pulverised coal where

pb = 760(1+o)-0138kgrrr3

dV5o = 30 u.m ps = 1 6 0 0 kg nr3

8 = 40° $ = 25° a = 20° D 0 = 0.052 m Sturtevant™ blow tank (Figure 2.1) Dy = 0.917 m Patm = 1010 hPa T = 293.15 K,

and for the following initial conditions selected at the transition VST = 0.002 m s-

1

(dp/dr)T = 0 P T = 150kPag or = 1.0 kPa.

The corresponding value of ms, which is constant with respect to r, is calculated to be 1.07 kg s"1. Note the slight convergence problem which still prevails towards the outlet of the blow tank (i.e. referring to Figure 7.3(c), the mean consolidation stress, o, is seen to increase as r/rj approaches r0/rj). Various values of initial condition were tested and the trends seem to indicate that the Ergun [64] equation provides a more realistic prediction for dp/dr, u and, hence, m s. However, extensive experimentation is still required before any conclusive assessment can be made.

143

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7.3 Dense-Phase Pipeline Conveying Characteristics

Several existing theories [17, 59, 60, 65, 66] were reviewed for the purpose of predicting dense-phase pneumatic conveying parameters in horizontal and vertical pipes and around different types of bends. Although all these theories were found to be limited to horizontal pipes, they still were considered to be worthy of investigation for the following reasons.

1. A large number of pipeline configurations that are used in industry contain at least 90 % horizontal pipe.

2. For the applications where the vertical pipe cannot be neglected, appropriate modifications to the horizontal data (similar to that proposed in Equation (6.34)) can be incorporated into the model.

3. To allow for the different types and numbers of bends, the experimental techniques described by Mills et al. [48] and Mills and Mason [49] could be employed to obtain an equivalent length of pipe for the bend(s), so that the total length of pipeline could be increased accordingly.

However, of the five theories that were investigated, the more simplified approach of Muschelknautz and Krambrock [59] was found to be most applicable to the work on pulverised coal and fly ash (viz. dense-phase conveying of fine powders) and is reviewed in the following section. The other four theories were found to be limited to particular applications. For example,

1. Konrad et al. [17] mainly considers the specialised case of transporting discrete plugs of cohesionless material (viz. in the pulse-phase mode, which is more applicable to granular products).

2. Chari [60] presents empirical correlations for solids friction factor based on materials with mean particle diameters in the range 140|im to 2629 urn.

3. The Fortier [65] theory requires a knowledge of material parameters, such as momentum loss coefficient (for solid particles impacting against a wall), which are extremely difficult to measure for fine powders. Also Fortier only reports on experimental results obtained for 3 m m particles.

4. When Wilson [66] applied his sliding bed model [67] to the case of pneumatic transportation, he still incorporated the empirical correlations (e.g. turbulent threshold velocity), which were based on hydraulic data of relatively coarse particles. Furthermore, some of the recent fly ash results, which were obtained from the test program described in Section 3.3, revealed that the pneumatic transportation of sliding beds was not possible for such fine materials and that a form of duning and/or moving bed flow [38] seemed to predominate.

7.3.1 Pressure Loss Predictions by Muschelknautz and Krambrock [59]

The mathematical model presented by Muschelknautz and Krambrock [59] is well documented and only a brief description is presented.

145

7.3.1.1 Theory

Considering the full-bore plug transport system displayed in Figure 7.4 and assuming that the pressure drop acting across the plug of length L is proportional to the wall friction caused by the body forces, it has been shown that

3f2 = exp

fym'gLV^ RTV„ (7.21)

where Pfi and Pf2 are the initial and final pipeline air pressures (Pa abs)

Y L = R = T = Vf = V p =

and m* =

material coefficient » 0.6 (average), total length of plug (m), gas constant for air = 287.1 (N m kg-1 K"1), absolute temperature of air (K), superficial air velocity (m s'1), velocity of plug (m s-1), mass flow ratio = m s mf

1 (kg kg-1).

,— D

Figure 7.4 Full-bore plug transport system.

Operational data from actual dense-phase systems have shown that there is a relationship between the ratio (Vp Vf

1) and the parameters m*. p* and £, where

Pf1 + Pf2

2 Pw (7.22)

146

Pfi Pf2

Pbl

Initial value of air density (kg nr3)

Final value of air density (kg nr3)

Loose-poured bulk density of the material (kg nr3)

1.6x10"^%* Frc

(7.23)

Fr V,

V?TJ (7.24)

Fre go

(7.25)

and v0 free settling velocity of the material based on pbi (m s-1).

Note that pbi is used instead of ps- Figure 7.5 presents the relationship that was obtained [59] between (Vp Vf

1) and (m* p*) for % = 0.01, 0.1,1.0 and 10.0.

7.3.1.2 Calculation Procedure

For ease of calculation, the following step-wise procedure is proposed for the application of the Muschelknautz and Krambrock [59] model.

1. Determine Voo using either Figure 7.6 (for air at 20° C) or the drag correlations recommended by Clift et al. [41]. The latter method requires use of

3Cn

0.5

(7.26)

and

Re s = pf Voo d v 5 0 (7.27)

where C o = drag coefficient, Re s = Reynolds number at free settling velocity (m s-1),

and (if = absolute or dynamic viscosity of air (Pa s).

2. Calculate Fr and Frs using Equations (7.24) and (7.25).

3. Using Equations (7.22) and (7.23), obtain expressions for (m* p*) and % in

terms of (Pn Pf2'1).

4. Assume a value for (P« Pf2"1) and calculate (m* p*) and %.

147

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149

5. Determine (Vp Vf1) from Figure (7.5).

6. Calculate (Pfi Pf2-1) using Equation (7.21).

7. Compare calculated and assumed values of (Pfi Pf2_1) and repeat Steps

4 to 6 until the difference is less than 1 %.

8. Determine the total pipeline air pressure drop using

fPfi ^ Apt f2 P ~\

VM2 J

(7.28)

assuming Pf2 to be atmospheric pressure.

7.3.1.3 Worked Example

The following steady-state operating conditions were obtained from a typical pneumatic conveying experiment, which was conducted on Tallawarra pulverised coal, using the original Test Rig A1 (L = 25 m and D = .052 m) described in Section 2.1).

mf =

ms =

m* =

Apt =

0.00935 kg S'1

2.93 kg s-1

313.4

75 kPa. and

Also note that the air temperature was 20° C and Patm = 101000 Pa abs. These data now are used as part of a worked example to demonstrate the Muschelknautz and Krambrock [59] model and the calculation procedure described previously.

1. From Figure 7.6, v„o = 0.02 m s_1, for dV50 = 30 u.m and pbi = 760 kg nr3

(obtained from Table 5.1). This value of v^ agrees well with the one calculated using the Clift era/. [41] drag correlations, where

Voo

Res

cD

=

=

=

0.0204 m s'1

0.0406 596

for w = 1.81 x Pa s and pf = 1.2 kg nr3. Hence, use v^, = 0.02 m s_1.

2. Calculate Vf.

V, = -JS-_ . <-00935> W = 3.67ms"1

Pf [j D2J 1 -2 it C052)2

150

Therefore, from Equation (7.24)

Fr V, 3.67

Vg~D

and Equation (7.25)

Frc =

V9.81 x .052 = 5.14

.020

Vg~D V9.81 x .052

Rewriting Equation (7.22)

m*Pf2(Pfi ."i 313.4 x 1.2

,028

m" p" = 2 Pbi VPf2 e + 1 2x760

rPJl *<Pf2

+ 1

.2474 (P fi \

+ 1 VH2 J

Also, from Equation (7.23)

\ = 1.6x10"^!^* (1.6 x 10"3) (5.14) 5 r

Fre .028 1.2

2x760

fPJl VPf2

+ 1

= .1614 f ^ + 1

4. to 7. The following table summarises the results obtained from applying Steps 4 to 7.

Assumed

(Pfi Pf2-1)

2.00 1.57 1.65 1.63

m* p*

.74

.64

.66

.65

\

.48

.42

.43

.43

(Vp Vf1) (Figure 7.5)

1.21 1.10 1.13 1.12

Calculated (Pfi Pf2"1) (Equation (7.21))

1.57 1.65 1.63 1.63

Table 7.1 Summary of results obtained from Steps 4, 5, 6 and 7.

Note that the calculated values of (Pfi Pf2'1) shown in Table 7.1 were

based on L = 25 m. That is, no allowance was made for the 3.6 m vertical lift or the five 1 metre radius 90° bends.

151

8. Substituting the final calculated value of (Pfi Pf2"1) = 1.63 into Equation

(7.28) yields

Apt = rdir--A V H 2 J

= (101000) (1.63-1)

= 63630 Pa

- 64 kPa.

This prediction is 15 % lower than the experimentally obtained value of 75 kPa. However, as no allowances for vertical pipe or bends were made during the calculations, then this result is considered to be reasonable. That is, if an equivalent pipe length of say 30 m was used for L in Equation (7.21), then a much closer result would have been obtained.

7.4 Correlation Analysis and Stepped-Diameter Pipelines

The case studies and proposed applications presented recently by Wypych and Arnold [42] indicate the need to optimise the design of the conveying pipeline in terms of minimising installation cost, air flow, pressure drop (hence, running cost), and transport velocity (hence, pipe/bend erosion). For this purpose, stepped-diameter pipelines should be considered, especially when large-throughput or long-distance conveying is required. Also, with improved design procedures, it is believed that this type of pipeline will become more popular for short-distance applications (e.g. L = 50 to 100 m) involving coarse and/or abrasive products. For example, refer to the Australian Industrial Refractories case study [42], which describes the use of three 80/100 m m N.B. stepped-diameter pipelines (viz. where L = 54, 56 and 62 m) to transport a wide range of heavy, coarse and abrasive materials at rates between 25 and 40 t lr1 with 70 < m* < 200. In relation to single- and stepped-diameter pipelines, the main aims of this section are to review generalised solids friction factor correlations for the prediction of

total pipeline air pressure drop Apt (viz. for a wide range of materials and single-diameter pipelines) and identify possible areas of improvement,

examine the general principles of designing stepped-diameter pipelines (including minimum transport behaviour),

suggest a combined test-design procedure to determine an optimal configuration for the pipeline,

present results from a recent investigation into predicting operating conditions for long-distance conveying (by using an improved correlation analysis).

152

7.4.1 Generalised Correlation for Solids Friction Factor

Quite often, correlation analysis is used to determine an empirical expression for the pipe friction coefficient due to solids Xs in the equation

pf Vf2 AL Ap = Apf + Ap s = (̂ + m* Xs)

KT 2Tp (7.29)

which is based on the definition presented by Barth [44].

Unfortunately, most of these correlations usually are limited to one or more of the following factors.

Only two or three different products (generally coarse). Low values of m* (typically, less than 30). One or two different pipe sizes (usually for similar values of L). Small values of D (typically, less than 50 m m ) .

Consequently, the practical use in industry of such correlations is quite limited and must only be exercised with extreme caution. Furthermore, complete sets of experimental data (viz. pipeline conveying characteristics which display the variation of mf, m s and Apt) generally are not published by the researcher or are located in articles which are extremely difficult to obtain.

To date, the most significant and applicable work seems to be by Stegmaier [68], who summarised the data from various products by the correlation, which is shown in Figure 7.7. Modifying this correlation to allow for the different definitions of Froude number (i.e. in this thesis, Fr = Vf (g D)-°-5 & Frs = v*, (g D)-°-

5, whereas Stegmaier [68] actually used Fr = Vf2 (g D)"1 & Frs = v*,

2 (g D) _ 1 , as shown in Figure 7.7) results in the following expression for Xs (presented previously in Equation (6.21).

Xs = 2.1(m*)-°-3Fr-2Frs

0-5(^)°'1 (7.30)

A number of researchers [45,69,70] also have applied Stegmaier's correlation [68] to their own work on predicting values of Aps. For example, Chambers and Marcus [70] recently have compared the experimental data of several researchers with the theoretical predictions from a mathematical model (based on Stegmaier [68] and Weber [45,69]) and found the difference to be less than a factor of 2 (for 2 < m* < 530).

Unfortunately, the data that were used by Stegmaier [68] to generate the line-of-best-fit given in Equation (7.30), were based mainly on small pipe diameters (e.g. 8 and 40 m m ) . Also in Figure 7.7, it can be seen that the proposed correlation does not represent accurately the data points that are plotted for Fr > 30. It was decided therefore, to investigate this apparent inadequacy in an attempt to develop an improved correlation for fly ash (and other fine powders, such as pulverised coal), which would be applicable to large values of L, D and m*.

153

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112

»6

y.

War5!

1500

26 50

2200

2J60

1500

2650

2200

26«0

2»80

MOO

MOO

D

M B

8

8

to

»0

to

to

TO

to

to

to

Quelle

ref.

Bohnet

KBUer

Llppert

•

V 8 \

\g

fi

U A o

' 0.<b

S a t *

•VA ITFT \A —=*

6 8 10 20 40 60 80

Figure 7.7 Correlation of pipe friction coefficient due to solids according to Stegmaier [68].

The experimental data, that are used for this purpose (i.e. to develop an improved correlation for solids friction factor), are obtained from

the dense-phase operating conditions of pulverised coal (Test Rig A1, Section 3.1 and [33]) and additional tests carried out on Test Rig A3, the pipeline conveying characteristics of seven fly ash samples (Test Rig B1, Section 3.3.3 or Figure 5.4) and additional tests carried out on Gladstone fly ash (Test Rig C3 and [71]), the pipeline conveying characteristics of fly ash/cement mix (Test Rigs C1 and C3, Figures 6.1 and 6.2), Muschelknautz and Krambrock [59], who presented data on fly ash conveyed over a distance of 1200 m.

154

A summary of the data and associated pipeline configurations are presented in Table 7.2. Appendix C contains a complete list of all steady-state operating conditions (viz. mf, m s and Apt) for each material and pipeline configuration. Note that the values of m* applicable to the data presented in Table 7.2, vary in the ranqe 7<m*<413.

No.

1

2

3

4

5

6

7

8

12

13

Material

Tallawarra pulv. fuel Tallawarra fly ash Eraring fly ash Munmorah fly ash Vales Point fly ash Gladstone fly ash Wallerawang fly ash Liddell fly ash Fly ash -cement mix Fly ash f591

Code Name

TWPF

TWASH

ERASH

MNASH

VPASH

GLASH

WGASH

LDASH

PFAC

PFA

Test Rig

A1 A3 B1

B1

B1

B1

B1 C3 B1

B1

C1 C3 -

L (m)

25 96 71

71

71

71

71 162 71

71

162 162 1200

D (m)

.052

.052

.052

.052

.052

.052

.052

.105

.052

.052

.060

.105

.200

Total No. Data Pts.

5 2 11

13

12

12

12 8 12

12

17 12 1

Values of m s

(kg s-1)

2.62, 2.93, 3.99, 4.47, 4.50 1.41, 1.47 1.0, 2.0,3.0

1.0,2.0,3.0

1.0,2.0,3.0

0.5, 1.5,2.5

1.0,2.0,3.0 j 5.0, 10.0, 15.0 0.5, 1.5,2.5

1.0,2.0,2.5

2.0,4.0,6.0,8.0, 10.0 5.0, 10.0, 15.0, 20.0 13.98

Table 7.2 Summary of products and experimental data for correlation analyses.

Also, note that

the physical properties for Samples 1 to 8 were presented previously in Table 5.1 (Section 5.2), to avoid confusion with the other materials listed in Table 5.1, the fly ash/cement mix and fly ash [59] listed above were numbered Samples 12 and 13, respectively, the pressure drops Apt listed in Appendix C are for the total pipeline and hence include any effects due to bends and vertical lifts, the values of dV50 and ps listed in Table 5.1 for Samples 1 to 8 were used in the following correlation analyses, the physical properties for Samples 12 and 13 have been presented elsewhere [71] and are repeated here for ease of comparison

Fly ash/cement mix : dV50 = 19 |im, ps = 2230 kg nr3, pbi = 720 kg nr3

Fly ash [59] : dV50 = 15 nm, ps = 2350 kg nr3, pDi = 1200 kg nr

3.

155

Using a Fortran 77 program on the University's mainframe computer, values of Fr, m*. Xs and Frs were calculated initially from the data sets listed in Appendix C (and the physical properties of each material). This information then was grouped according to Equations (7.31) and (7.32) and plotted on the same axes which were used by Stegmaier [68]. The resulting plot is shown below in Figure 7.8, which includes the correlation given by Equation (7.33).

X = Fr

Y = ^(m^Fr;0- 5^) 0' 1

= 2.1 XT'

(7.31)

(7.32)

(7.33)

T — i — i — T r

ie

in •

i

« L U_

ca co

<*

X

o

•

o A

+ X

X K S X I

1 i i i i i r

MATERIAL (L.D)

THPF(2S..052) TWPFC96..0S2) THASH171.-B52I ERASH171..052) MNASHH1..052) VPASHC71..052) GLASH(71..052J " CLASH 1162.. 1051 HGASH(71.-052) LDASH171..052) PFACtl62..060) PFAC 1162..105) PFA 11200. -200)

II*

i » I _ L 100.

X = Fr

Figure 7.8 Comparison between experimental data and the Stegmaier [68] correlation.

156

Note that Fr was assumed to represent the velocity Vf at the end of the pipeline (i.e. at essentially atmospheric conditions) and similarly pf was assumed to be constant at 1.2 kg nr3. As various groups of data (for a given product) were plotted by Stegmaier [68] at a constant value of Fr (e.g. refer to quartz powder in Figure 7.7), these assumptions do seem to be valid. Also, note that dV5o was employed to represent each product, whereas Stegmaier [68] may have used a mean value (no definition w a s presented). However, the subsequent effect on the location of the data points shown in Figure 7.8 is expected to be minor.

One further aspect that should be considered and discussed in more detail is the calculation of Xs (i.e. refer to Equation (7.29)). As indicated previously, the friction

factor is based on the total pipeline air pressure drop Apt and hence, includes any effect due the bends and vertical lift. Unfortunately, at the time of undertaking the various test programs, the pipelines had not been instrumented sufficiently to extract accurate straight pipeline air pressure gradients from the recorded data. Also, it is believed that the sections of pipe used on most of the test rigs (viz. Test Rigs A1, A 3 and B1) were not long enough to ensure fully established flow (i.e. due to reacceleration of the product after each bend). As far as vertical lift on Test Rigs A1, A3, B1, C1 and C 3 is concerned, the ratio (Lv L

_1) is seen to vary between 0.027 and 0.144. Also, most of the data presented in Figure 7.8 were obtained on Test Rigs B1, C1 and C3, where 0.027 < (Lv L'

1) < 0.051. To evaluate the possible effect on X-s, the ratio of vertical to horizontal pressure gradient K is assumed = 2 (i.e. refer to Equation 6.35 in Section 6.3.4). This allows an equivalent length of 2 x Lv to be calculated for the vertical lift. The subsequent percentage increase in total pipeline length (i.e. L + Lv) for Test Rigs B1, C1 and C 3 is seen to vary between 2.7 and 5.1 %, which only would have a minor effect on the data points plotted in Figure 7.8. In fact, as the values of Xs and hence Y would reduce slightly as a result of this allowance, any correlations representing such data would be considered slightly conservative. Also, it is believed that the trends displayed in Figure 7.8 would not be affected by this modification. In contrast, the effect due to bends may have a significant influence on Apt (e.g. refer to [48,49,72]). However, the results obtained from more recent investigations on pulverised brown coal (presented in the following section) indicate that under certain circumstances (e.g. relatively few number of bends for a given length of pipeline, moderate conveying rates for a given D), the presence of bends may be ignored for the calculation of straight-pipeline air pressure gradient and hence, Xs. Note that the effective length [9] of a 1 m radius 90° bend (i.e. 2 metres) is included in the determination of L (i.e. the total effective length of a pipeline) and hence, to some extent would compensate for the pressure loss caused by the bend. However, for the data presented in Figure 7.8, it is difficult to estimate the degree of compensation. Nevertheless, for the main objective of identifying possible areas of improvement, the use of Apt to calculate Xs is considered sufficiently accurate. In fact, it is believed that the trends displayed in Figure 7.8 (and future analyses) are not affected severely by this approximation. For example, in both Figures 7.7 and 7.8, the data points for Fr > 20 are seen to diverge away from the Stegmaier correlation (i.e. for increasing Fr). During the preliminary stages of developing an improved correlation (i.e. based on the data presented in Table 7.2 and Appendix C), it was found that this effect was caused largely by not including the variation of average air density pf in the correlation

157

analysis. Also, the power index for m* used by Stegmaier [68] (i.e. 0.3) did not seem to represent adequately the changes in pipe diameter (for the same product).

The final set of axes that were selected as most appropriate (i.e. to represent the the experimental data), are summarised in Equations (7.34) and (7.35).

X =

Y =

Frmpf -0.5

m

,1 \0.8 , n N0.85

Mm*)°-8Fre°-6(^ D Jv50.

(7.34)

(7.35)

The resulting correlation (i.e. line-of-best-fit), which is presented in Figure 7.9, was found to be

Y = 9 x 1 0 6 X " 1 J 8 2 (7.36)

.18+808

co

CS CD

10+007 -

ca to

GO

il

ea 09

10*006 -

M

10+005 100.

X = Fr m tpfm> -.50

Figure 7.9 Improved correlation of pipe friction coefficient due to solids.

158

Note the values of pfm and Frm, which were used for this correlation analysis, were calculated from the following equations.

Pf2 + 0.5 Aps Pfm = RJ (7.37)

4 mf Frm = *n no.5n2.5 (7.38)

« Pfm 9 D

It is evident from Figure 7.9 that the scatter of experimental data has been reduced significantly and the divergence effect observed previously (i.e. in Figures 7.7 and 7.8) for the larger values of Fr has been eliminated. It is interesting to note that to improve the accuracy of the correlation (i.e. as shown in Figure 7.9), the pipeline length L had to be included in Equation (7.35). This was unexpected and possibly could be a result of either not selecting an optimal power index for pfm (i.e. -0.5) or using Apt in Equation (7.29). That is, in relation to the latter possibility, the inclusion of L may be compensating for the different number of bends, as well as the various values of (Lv L

-1). Nevertheless, even though there exists some doubt over the numerical accuracy of this correlation (i.e. mainly due to using Apt in Equation (7.29)), it still may be concluded from the trends observed in this investigation, that employing the mean values pfm and Frm significantly reduces the scatter

shown in Figure 7.8 (which is based on the exit values of pf and Fr), including air density pfm in the correlation analysis reduces the scatter of data points (i.e. for a given material and pipeline) and also improves the divergence problem for large Frm, the power index for m* provides a better representation of larger and different values of D (i.e. for a given material).

In applying the improved solids friction factor correlation (i.e. as presented in Equation (7.36)), to the prediction of pressure drop for various materials, pipeline configurations and operating conditions, the resulting iterative calculations were found to display poor convergence and in fact, on several occasions unable to provide a confident value for Ap s and hence, Apt (i.e. using the suggested Equation (6.40) to calculate Apt). However, due to the following reasons, it was decided to postpone any further investigations into this matter. Some uncertainty exists over the accuracy of the solids friction factor

correlation shown in Figure 7.9 (i.e. due to including the effect of bends and vertical lift in the pressure drop data). Good accurate data (e.g. obtained from long straight sections of pipe of different diameter) are essential for the efficient investigation of such problems. That is, numerous experiments on different products and pipelines would have to be carried out for this purpose. The general objectives of this section of work largely have been fulfilled (i.e. reviewing generalised correlations and identifying possible areas of improvement).

159

Instead, it was decided to concentrate on developing a test-design procedure for the purpose of determining an accurate correlation for a given material and different values of L and D (i.e. after isolating the relevant physical variables and dimensionless groupings). In this way, as several different materials are tested and correlated in the same manner, the generalisation of solids friction factor (and the troubleshooting of any divergence problems) can proceed with confidence. Investigations into determining an accurate correlation for pulverised brown coal (applicable to single- and stepped-diameter pipelines, over short and long distances) are considered in the following sections. However, the principles of designing a stepped pipeline are reviewed initially.

7.4.2 Design of Stepped-Diameter Pipelines

Over the past decade, only a number of investigations [5,6] have been carried out in this area of pneumatic conveying design and these have been limited to longdistance pneumatic transportation. That is, the application of stepped-diameter pipelines to large throughput conveying and/or over relatively short distances has not received much attention. The major difficulties that need to be overcome before a stepped-diameter pipeline can be designed and optimised for a particular application, include the determination of the following items. (i) Number of different pipe diameters (i.e. n) for a given L (ii) Size of each pipe section (i.e. Dj for i = 1 to n). (iii) Location of each transition (i.e. ALj for i = 1 to n). (iv) Minimum mass flow rate of air (i.e. min. mf) to suit (i) and (ii).

However, the first decision to make is the maximum operating pressure. This will depend on the type of feeder (e.g. 100 kPag for a rotary valve, 300 kPag for a single (i.e. batch) blow tank, 500 kPag for a tandem blow tank system) and the required overall reliability of the system. For example, the operating pressure may need to be reduced to increase the service life of the hardware and components (e.g. discharge and vent valves for blow tanks, rotary valve clearances). Once deciding on this parameter, it then is necessary to optimise Items (i), (ii) and (iii), which are dependent on the material (as well as its minimum transport properties), the operating conditions (e.g. conveying rate ms, solids to air mass flow rate ratio m* = m s m r 1 ) and the available sizes of pipe (e.g. 100, 125, 150 m m N.B.). It is suggested from experience that the increment in pipe diameter (i.e. in the direction of flow) be kept to a minimum (e.g. 100 to 125 m m N.B. and not 100 to 150 m m N.B.). Selecting the number and sizes of the different pipe sections (i.e. n and Dj for i = 1 to n) is a difficult task and usually relies on experience and/or trial and error. Note that it would be advantageous to minimise the value of D at the end of the pipeline for reasons of ease/cost of installation and avoiding the use of large bends and valves that may be required (e.g. diverters, unblocking systems). A practical upper limit seems to be 250 m m N.B., although 200 m m N.B. would be preferred, if possible. Note that the pipe sizes 50, 80, 100, 125, 150, 200 and 250 m m N.B. would cater for most applications with the latter five being used more frequently for long-distance and large-throughput transportation.

160

Once selecting n and the values of Dj (i.e. for i = 1 to n), the next stage of the design process involves determining the location of each pipe transition (i.e. the values of ALj for i = 1 to n). The ultimate aim of selecting the combinations of Dj and ALj is to minimise the total pipeline air pressure drop Apt, which then is compared with the selected value of maximum operating pressure. For the purpose of this approach, it is suggested initially to maximise Al_i for Di (where i = 1 represents the final or largest size of pipe) and then A L 2 for D2, and so on until i = n for the initial or smallest size of pipe. In this way, it may be determined whether

the number of different pipe sizes (i.e. n) is excessive, adequate or insufficient for the total length of pipeline L, new combinations of Dj and ALj may be necessary to either increase or decrease Apt (i.e. with respect to the maximum operating pressure).

Ideally, it is desired to achieve this result with the smallest possible value of Dn which (in conjunction with the other different pipe sections) produces a Apt just below the maximum operating pressure. In this way, the required values of air flow (i.e. mf) are minimised as well. For the purpose of maximising ALj for each Dj (i.e. starting from the end of the pipeline), the following information must be determined for the product(s) in question and also each section of pipe.

(a) Minimum transport behaviour and properties (e.g. minimum superficial air velocity Vffmin or Froude number Frmjn).

(b) Maximum transport velocity Vf)max or Froude number Frmax. (c) Pressure drop for the air and solids components (viz. Ap,- = Aptj + ApSi).

Note that Apt represents the total air pressure drop for the overall pipeline,

whereas Ap,- represents the total pressure drop for pipe section No. i, which may

include bends and/or vertical lifts. That is, Apt = Z (Apj) for i = 1 to n. Items (a), (b) and (c) constitute the mechanism that is necessary to optimise the combinations of Dj and AL,- and hence, the design of the overall pipeline. The following section considers both Items (a) and (b) in the light of existing criteria to step pipelines, and Item (c) is investigated in Section 7.4.3.

7.4.2.1 Stepping pipe criteria

In relation to Item (a) above, a minimum Froude number (Frmjn) approach often is used by researchers. For example, refer to Bohnet [73]. However, from experience it is believed that for a given material, Frmjn should be modified by air density pt, as suggested by Equation (7.39), where i and v are exponents and E is a constant.

Frmin(pf)T = EOrfT (7.39)

This is supported by the findings of Roski [5], who determined the following relationship for natural anhydrite

, x0.15

Frmin - ^ - = ( 1 1 4 . 3 6 +159.23 m * ) 0 5 (7.40) V pf.atm /

161

where pf.atm is the atmospheric air density. S o m e attempts have been made to generalise Frmjn for different products. For example, Weber [45] suggested the following relationships for Frmjn in terms of the single particle settling velocity v<„ and the particle to pipe diameter ratio (d D"1).

f ft N r H \®~^ Frmin = (̂— v^ + 7J (m*) 0 2 5 (̂ —J forvoo<3ms-1 (7.41)

x0.25 I d 0.1

Frmin = 15(m*r^^J for^^rns"1 (7.42)

r- M mjn pfFrmin = Pf-j== (7.43)

Vg D Also, Schade [74] predicted the critical or minimum conveying velocity Vf>min for 9 coarse products (d = 0.3 to 3.5 m m ) with a mean linear deviation of 19 % using the relationship / w x-0.025 / _, N-0.34

Fr m i „ = ( m T ' ( £ ) £ ) (7.44)

which was based on 151 experimental measurements and where ps is the particle density. Note that from Equations (7.40), (7.43) and (7.44), it can be seen that as pf increases Frmjn decreases. This effect due to air density indicates that Frmin and hence Vf>mjn (as well as mf) may be reduced for higher operating pressures (i.e. towards the upstream end of each pipeline section). Therefore, in order to allow for this effect (so that the stepped-pipeline configuration may be optimised), the experiments must be carried out at air densities ranging from pf« pf,atm to pt.max. which corresponds to the selected value of maximum operating pressure. Rizk [7] proposed an alternative approach by correlating the pressure minimum curve (defining the transition between the dilute- and dense-phase regimes) and suggesting the relationship m* = _J__ Frx = Ki Fr* (7.45)

10<°

where the exponents co and % are dependent on the equivalent particle diameter of the product and the pipe material (for bulk solids with an approximately constant value of ps). In fact, Rizk [7] demonstrated that for three different sizes of polystyrol (d = 1 to 2.5 m m ) conveyed through four different pipe diameters (D = 50 to 400 m m ) , the resulting pressure minimum curves could be correlated by the equation m* = Ki Fr4 to an accuracy of ± 15 %. This could be applied to the design of a stepped-diameter pipeline by assuming Frmjn «the Froude number given in Equation (7.45). Note this will give a conservative result, as for most materials the onset of solids deposition or flow instability occurs at a critical or minimum Froude number which is less than the pressure-minimum value (i.e. Fr in Equation (7.45)). Also, again it is believed that such a relationship should be modified to allow for air density. At this stage, therefore, it is suggested that in order to determine an accurate expression for Frmin (e.g. similar to Equation (7.39)), experiments should be carried out on the material(s) in question and for the expected range of air densities. Once obtaining a

162

complete set of data, it then is possible to evaluate the various approaches described above and determine the best relationship between Frmjn, m* and pf.

As far as Item (b) in Section 7.4.2 is concerned, Marcus and Bettman [6] recommended simply that an increase in pipe diameter should occur when Fr reaches a selected maximum (i.e. Frmax) and also suggested that Fr m a x is a function of the energy utilisation and the required levels of product degradation or damage. It is believed that erosion of the bends and pipeline also should be considered in the selection

of Frmax, increasing the length of a particular section of pipe (i.e. increasing ALj for a given Dj in the direction of flow) to ensure Fr = Fr m a x is not necessary and in fact, could hamper the ultimate objective of minimising pressure drop, F r m a x should be used mainly as a guide to ensure Frmjn < Fr < Fr m a x (i.e. for any section of pipeline).

7.4.3 Test-design procedure

Using generalised correlations for the determination of solids friction factor Xs in the equation [45]

ft} = (Xf + m*U^ (7.46)

where Xf is the air-only friction factor, certainly avoids the use of extensive test work. For example, refer to Weber [69]. However, with typical mean linear deviations ranging between 30 and 40 %, it is believed that the resulting degree of uncertainty would be too great for long-distance applications, especially when the operating pressures are in excess of 300 kPag. Also, the more recent predictions of Chambers and Marcus [70] to within a factor of 2 are considered too inaccurate for this application. Hence, as suggested in Section 7.4.2, sufficient experiments should be carried out to determine an accurate relationship between Fr, m* and Xs. For this purpose, the following test-design procedure is recommended. Note, to demonstrate this procedure, results from recent investigations into the long-distance pneumatic conveying of pulverised brown coal (median particle diameter dV50 = 21 [im, p s = 1500 kg rrr

3, loose-poured bulk density pbi = 515 kg nr3) are presented where appropriate. (i) For a wide range of mf, ms and pf, record the pressure drop Apj and upstream

static gauge pressure for a test section of pipeline (of diameter Dj), which should have a length of ALj « 10 to 20 m. Note the test section should be located along a straight section of pipe and sufficiently distant from any bend effects. To minimise the total number of experiments it may be convenient to use two test sections, one at the beginning of the pipeline and the other towards the end. Note, the latter could be installed in a larger size of pipe (i.e. using a stepped-diameter test rig pipeline) to establish/verify the variation of Xs with respect to D. Also, it is suggested to monitor and develop a correlation for the pressure drop caused by a bend. As demonstrated by Bradley [72], this

163

effect may contribute considerably to the overall pressure drop of a pipeline. Table 7.3 summarises the length of each diameter section of the 947 m long pipeline (i.e. Test Rig E1), which was employed for the recent investigations on pulverised brown coal. Note the numbering system for the different diameters commences at the end of the pipeline (to be compatible with later calculation procedures).

Pipe No. 1

4 3 2 1

D| (m)

.060

.069

.081

.105

ALi (m)

146.0 390.0 261.0 150.0

No. of Bends

Nb

3 13 8 5

Table 7.3 Pipeline configuration for Test Rig E1.

Note that at the time of undertaking the test program on the pulverised brown coal, the preferred method of pressure drop measurement described above had not been installed. Instead, three to five single pressure tappings points were used along selected sections of straight pipe for each of the different diameters listed in Table 7.3. To determine the required values of pressure gradient, lines of best fit were drawn through each set of data. For the nine experiments which were data-logged, the effect of the bends on pressure drop was found largely to be insignificant (possibly due to the relatively few number of bends with respect to each ALj and also the effective length [9] of each bend being included in ALj). Examples of the pressure gradients which were obtained for the Dj = 0.060 m pipe (two test sections) are presented in Figure 7.10. Therefore, it was decided for each experiment to calculate the pressure drop caused by each Dj and hence, obtain an average pressure gradient. Also, this technique was found to minimise some of the experimental scatter. The overall operating conditions obtained from each experiment as well as the various values of Apj (for i = 1 to 4) are summarised in Table 7.4.

164

Exp. NO.

0 5 6 7 8 10 11 12

1 3

mf

(kg S'1)

.120

.127

.120

.096

.087

.125

.180

.200

.200

ms (kg s"1)

2.50 2.94 1.80 2.08 2.14 1.67 1.43 1.79 2.94

m* (kg kg-1)

20.83 23.15 15.00 21.67 24.60 13.36 7.94 8.95 14.70

AP4 (kPa)

91 96 72 76 75 76 76 85 110

AP3 (kPa)

165 175 147 143 150 138 145 166 209

Ap2

(kPa)

81 91 73 72 72 69 72 97 108

AP1 (kPa)

29 32 25 27 28 23 27 32 38

Apt (kPa)

366 394 317 318 325 306 320 380 465

Table 7.4 Steady-state operating conditions for the 947 m Test Rig E1 pipeline.

500

60

U

CO CO CU

u P-, Vi

•H <J CU

C •H H CU O. •H PM

400 -

300 -

200 50 100

Distance from Blow Tank (m)

150

Figure 7.10 Examples of air pressure drop for the Dj = 0.060 m section of pipe, showing the vertex location of the three 1 m radius x 90° bends.

165

(ii) Using the data obtained from each experiment, the following calculations are performed. A summary of the results are presented in Appendix D.

Pfm (Pa abs) = Mean absolute air pressure of test section = f1 + f2

2

Pfm pfm (kg m

- 3) = Mean air density = •=•= H \

Pfi (Pa abs) = Absolute air pressure at upstream end of test section. Pf2 (Pa abs) = Absolute air pressure at downstream end of test section.

4 mf Vfm (m s 1) = Mean superficial air velocity =

Vfm Frm = Mean Froude number =

V g D i

71 pfm D 2

X\ = Overall friction factor for test section = 2DjAp

pfm V f m AL,

Apf (Pa) = Air only component of pressure drop [53] =

2 Dj Apf

1.85 1141.83 rry-ÂL,

Pfi Dj5

Xf = Air only friction factor = 2 Pfm Vfm ALj

X[ — Xf Xs = Solids friction factor = 7—

m

(iii) The values of Xs then are plotted against Frm, as shown in Figure 7.11. Note

the actual value of m* for each A,s and Frm has been included on this plot with the decimal point representing its actual location . It can be seen from Figure 7.11 that due to the large amount of scatter, it is difficult to establish a relationship between Xs and Frm (i.e. for any given value of m*). Although a correlation could be fitted statistically to these data, it is believed that the resulting errors would be too great for the design of long-distance pipelines. The results presented previously in Section 7.4.1 indicated that this type of scatter was caused largely by not including air density in the relationship between Xs and Frm.

025

020

14-6

21-7

166

.015 IS'O

.010

005

/3«4

20-6

24-6

2+* 24-6 23-2

21-7 2/«7

2I«7

10

Fr_ m

7*9

9-0

20-8 2og

23-2 23-2

(£"•0

20-tf 23-2

)3-+ (4-7

/4*7

15

9-o 7-9

9-0 7-9

/4»7

1

9-0

7»9

J 20

Figure 7.11 Relationship between Xs and Frm showing actual values of m*.

167

(iv) To generalise the relationship between Xs and Frm (i.e. in terms of pfm and m*),

values of Xs are plotted initially against X = Frm (pf m)e for different values of e

(e.g. refer to Figure 7.12), and then modified using Y = Xs (m*)f. After determining the best values for exponents e and f, X and Y are replotted on log-log scales, as shown in Figure 7.13, which includes the following line-of-best-fit for the pulverised brown coal.

Y = 1.8076 X-1-423 (747)

where X = Frm (pfm)0-2 and Y = Xs (m*)°-5. From Equation (7.47), the following

expression was obtained for Xs.

Xs = 1.8076 (m*)-0-5 (Frm)-1 -423 (pfm)'0-2846 (7.48)

(v) To evaluate the accuracy of Equation (7.48), predicted values of Xs are plotted against the experimental values, as shown in Figure 7.14. It can be seen that for the pulverised coal, Equation (7.48) predicts Xs well within ±10 %. Also, a good comparison can be seen in Figure 7.12 between the actual values of m* and the curves predicted from Equation (7.48). Noting that the experimental data were obtained from four different diameters of pipe section and also would be subjected to a certain degree of scatter (i.e. due to the tests being undertaken over a total distance of 947 m ) , this is considered to be a good result. Also, this indicates that the same correlation (i.e. Equation (7.47)) would have been obtained if only one pipe section had been monitored and analysed (i.e. 0.060, 0.069, 0.081 or 0.105 m ) . However, as operating pressures of the test section must approach the maximum value suggested for the actual plant, either the 0.060 m or 0.069 m sections of pipe would have been preferred.

A computer program also was written to predict values of Apt for the test rig pipeline (using the operating conditions listed in Table 7.5). Note that the calculation procedure commences at the end of the pipeline (where Pf2 = 101000 Pa abs and T = 293.15 K are assumed) and then involves iteration for each different diameter section of pipe. The corresponding predicted values of total pipeline air pressure drop have been included in Table 7.5.

(vi) Plotting the predicted operating conditions listed in Table 7.5 and superimposing the experimental data given in Table 7.4, results in the pipeline conveying characteristics for the pulverised coal, as shown in Figure 7.15. This graph demonstrates further the good accuracy of Equation (7.48) and the design/analysis technique. To determine the amount of relative error, values of Apt also were predicted for the experimental values of mf and m s listed in Table 7.4 and have been summarised in Table 7.6.

168

.020 -

015 -

.010 -

.005

Fr m 'fm

0-2

Figure 7.12 Relationship between Xs and X = Frm pfm0,2 showing experimental

values of m* and predicted curves, based on Equation (7.48).

169

200

.100

.080

.060

.040

020

010

~i 1 1 r 1 1 1 r

Y = 1.8076 X-1'1*23

- i 1 i i _ J ' ' '

10 20 30

Figure 7.13 Relationship between Y and X, where Y = XS (m*)0-5

and X = Frm ptm02-

170

0 .01 .02

Xg (Experiment)

Figure 7.14 Comparison between actual and predicted values of Xs, based on Equation (7.48).

m s (kg s*1)

3 3 3 2 2 2 1 1 1

mf (kg s-1)

.10

.15

.20

.10

.15

.20

.10

.15

.20

Predicted Apt (kPa)

388.1 425.4 464.4 319.7 357.1 397.1 234.7 I 273.2 i 314.8

Table 7.5 Predicted values of pressure drop for the test rig pipeline, based on Equation (7.48).

171

(kg s-i)

2

1

0

0 0.1 0.2 0.3

mf (kg s_I)

Figure 7.15 Pipeline conveying characteristics of pulverised coal for L = 947 m and D = .060/.069/.081/.105 m (Test Rig E1), showing experimental data

points and predicted curves, based on Equation (7.48).

Exp. No.

0 5 6 7 8 10 11 12 13

mt (kg s-1)

.120

.127

.120

.096

.087

.125

.180

.200

.200

ms

(kg s-1)

2.50 2.94 1,80 2.08 2.14 1.67 1.43 1.79 2.94

Exp. Apt (kPa)

366.0 394.0 317.0 318.0 325.0 306.0 320.0 380.0 465.0

Predicted Apt (kPa)

370.0 404.2 319.0 322.8 320.7 312.6 336.2 381.4 460.7

Error

(%)

+1.1 +2.6 +0.6 +1.5 | -1.3 +2.2 +5.1 +0.4 j -0.9 I

Table 7.6 Comparison between experiment and predicted values of Apt.

400 -

Apt

(kPa)

200 -

172

From these results, it can be seen that the largest difference in pressure drop is 16.2 kPa, which represents a percentage error of +5.1 %. This is considered very accurate for a long-distance stepped-diameter pipeline. However, it should be noted that the test rig bends are included in the correlation presented in Equation (7.48). Hence, it is believed that any predictions for pipelines containing a relatively smaller number of bends will be conservative. Further experiments and detailed investigations are required before an accurate correlation for bends can be developed and at this stage are considered beyond the present aims of this section.

(vii) To determine for a proposed length of pipeline, the optimal number and location of pipeline transitions (i.e. stepped diameters), the minimum Froude number (Frmin) approach, as defined by Equation (7.39), is suggested. Refer to Section 7.4.2.1. Unfortunately, insufficient data were generated during the initial test program on pulverised coal to determine accurate values for z and i). Nevertheless, to demonstrate the principles of this approach and also the advantages of stepping pipelines (i.e. in relation to using a single-diameter pipeline), Frmin » 6 is used to optimise the design of a hypothetical 1.8 km pipeline, which is required to transport the pulverised brown coal at a continuous rate of 24 t rr1. Due to this substantial conveying rate (and hence, relatively high operating pressure), a tandem blow tank system is suggested. Also, for long term reliability of the hardware and components, a 400 kPag maximum operating pressure is suggested (i.e. Apt < 400 kPa).

(viii) Using the same computer program referred to in (v) and adjusting the various lengths of selected diameter sections to maintain Frmin ~ 6.0, results in the prediction of operating conditions for five different pipeline configurations (D = .127/.154 m, .127/.154/.203 m, .154/.203 m, .154 m and .203 m). The final results are presented in Table 7.7.

(m)

.154

.127

.203

.154

.127

.203

.154

.154

.203

AL, (m)

1140 660

475 710 615

880 920

1800

1800

ms

(kg s*1)

6.7

6.7

6.7

6.7

6.7

mf

(kg s-1)

.61

.55

.78

.95

1.37

m* (kg kg-1)

11.0

12.2

8.6

7.1

4.9

Ap| (kPa)

273 232

69 169 211

137 240

479

319

Apt

(kPa)

505

449

377

479

319

Vfi (m s'1)

7.4 6.7

8.4 7.3 6.6

8.5 7.4

7.4

8.5

FM (-)

6.0 6.0

6.0 6.0 6.0

6.0 6.0

6.0

6.0

Vf2

(m s"1)

27.3 10.8

14.2 14.6 10.8

20.1 14.8

42.5

35.3

Fr2 (-)

22.2 9.7

10.0 11.9 9.7

14.2 12.0

34.6

25.0

Table 7.7 Suggested pipeline configurations and predicted operating conditions for pulverised brown coal conveyed at 241 fr1 over L = 1800 m.

173

In relation to minimising air pressure, velocity and the total amount of air required, the advantages of selecting a stepped pipeline (i.e. instead of a single-diameter) can be seen from the results presented in Table 7.7.

Note that the above predicted values of Ap, (and hence Apt) are considered conservative for the following two reasons.

(i) The effect of bends were included in the pressure drop data, which were used to generate the correlation represented by Equation (7.48).

(ii) It is believed that Frmin (and hence, Vfimin as well as mf) can be reduced for higher operating pressures' (i.e. towards the start of each pipeline section). Refer to Section 7.4.2.1.

Nevertheless, they provide a good relative indication of what to expect for the final installation. In fact, based on this information and also for reasons of providing long term reliability (e.g. minimising system erosion and hardware problems), the following configuration and operating conditions are suggested.

Di = .203 m, AL| = 880 m, D2 = .154 m, AL2 = 920 m, m s = 6.7 kg s-1 = 241 h-\ mf = 0.78 kg s-1 (39 Nm

3min-1), m* = 8.6, Apt - 377 kPa < 400 kPag.

The .203/. 154/. 127 m stepped pipeline also could be selected, but it is believed that hardware components (e.g. discharge and vent valves), which are in contact with the material and subjected to operating pressures greater than 400 kPag, will have a greater chance of fatiguing and malfunctioning over a period of time.

174

CHAPTER 8

175

8. CONCLUSIONS

While some progress has been made during the last decade in the understanding of pneumatic transportation, the technology that is required to design or select industrial pneumatic conveyors with complete confidence is somewhat limited. The primary objective of this thesis is to provide industry with some of this technology in relation to the pneumatic transportation of fine powders (e.g. pulverised coal, fly ash, P V C powder) and some coarser products (e.g. crushed bath, bone char, screened coke). The work undertaken considers the main areas of determining pipeline conveying characteristics, evaluating various types of conveying mode and blow tank configuration, developing powder characterisation techniques and formulating mathematical models to predict some of the more important design parameters. Pipeline conveying characteristics and the transient plots of major conveying parameters demonstrate the large differences that can occur in the flow performance and the minimum transport behaviour of bulk solids. Pilot-plant testing is essential for reliable design, where it is necessary to be aware of any operational problems resulting from unusual material properties (e.g. strong plugging). A standardised-test procedure comprising three different types of experiment is developed to generate data efficiently for the presentation of pipeline conveying characteristics and also delineate any unstable or unreliable transport phenomena. The application of this procedure to different materials on the one test rig provides an accurate means of comparing their relative performance in a pneumatic conveying system (e.g. dense-phase). Also, in conjuction with accurate scale-up equations or solids friction factor correlations, the technique provides a basis for the future design of pneumatic conveying systems. The pipeline conveying characteristics and the transient plots of the major conveying parameters demonstrate the wide range rangeability of conveying parameters and hence, good dense-phase performance of pulverised coal and fly ash. However, other results from slightly coarser materials (e.g. P V C powder, screened coke and coarse ash) emphasise the large differences that can occur in flow performance and minimum transport (dense-phase) behaviour (e.g. unstable plugging in the vicinty of saltation). The introduction of supplementary conveying-air at the blow tank outlet provides a smoothing effect upon conveying parameters. This result is particularly important if lower values of conveying pressure or air flow rate are required or if longer transport distances are desired. Additional results from the fly ash test program emphasise the importance of considering material properties when selecting an appropriate blow tank configuration and an efficient method of air injection (for the dual purpose of fluidisation and pressurisation). The transient plots of major conveying parameters demonstrate the importance of blow tank air injection on the overall performance of a plug-phase pneumatic conveying system. Although this method of transport is able to handle a wide range of conventionally difficult dense-phase materials (e.g. crushed bath, screened coke), it is relatively sensitive to changes in the physical properties of a material (e.g. particle size). However, to some extent, it is possible to compensate for such changes by selecting a different method of air injection.

176

Experimental results from fly ash/cement mix and P V C powder indicate the inadequacy of the original scale-up criteria, which were employed initially to design pneumatic conveying systems. Improved design equations are derived from these results and modified to allow for the relatively longer lengths of vertical that are used usually in industry. The scale-up of the air-only component of pressure drop is an equally important design consideration (especially for long-distance applications). Accurate equations are developed, and their predictions compare favourably with 27 data points from 7 different configurations of pipeline (including a 940 m long stepped-diameter pipeline). A technique to generalise pipeline conveying characteristics also is proposed and simplifies scale-up procedures for design. Despite certain limitations (e.g. particle size distribution), the Geldart [24] and Dixon [23] classifications generally provide a useful technique to indicate the fluidisation performance of bulk solids (important for efficient transportation and blow tank operation) and assessing a material's suitability to be conveyed in the dense-phase mode. However, although such information is important for design, it still only provides a qualitative evaluation of the material. The results from an approximate analytical model to represent the governing differential equations of a converging flow channel seem to underestimate the rate of solids discharging from the outlet of a blow tank (i.e. for a given pressure drop across the material). A thorough assessment of the mathematical mode! still is required and depends on the accurate measurement of the relevant variables, such as pressure and velocity inside the blow tank. After several changes and improvements, the predictions of solids discharge rate from the latest version of a numerical model to predict blow tank characteristics seem to be more realistic (i.e. when compared with the analytical model). Five pipeline theories also are evaluated with one being recommended for predicting pressure drop in the dense-phase pneumatic transportation of fine powders. A worked example is presented and the results compare favourably with data obtained from the initial test program on pulverised coal. Generalising solids friction factor correlations for the prediction of pressure drop avoids the need of extensive test work. However, the applicability of such correlations to industry is limited, especially when good accuracy is required (e.g. for long-distance and/or large throughput conveying). The results from preliminary investigations into identifying possible areas of improvement indicate that mean air density should be included in the correlation analysis. Stepped-diameter pipelines provide definite advantages for long-distance and large-throughput pneumatic conveying, but also could be applied to short-distance applications involving heavy, coarse and/or abrasive materials. Based on existing criteria and the results of this thesis, a combined test-design procedure is proposed and demonstrated for the ultimate objective of obtaining an optimal configuration of pipeline. Results from recent investigations into the long-distance pneumatic conveying of pulverised brown coal, demonstrate the applicability and good accuracy of this test-design procedure by predicting to within ± 5 % the operating conditions obtained from a 947 m x 60/69/81/105 m m I.D. stepped pipeline.

177

These results demonstrate further that the same correlation and predictions could have been obtained using only one size of pipe. This also indicates that experimental work in the future only may need to be carried out on a short-distance pipeline (i.e. to predict operating conditions for long-distance and/or large-throughput applications). However, operating pressures in the test rig still are required to approach those expected for the actual or proposed system. Perhaps this could be achieved by replacing the test rig receiving silo with a pressure vessel having good back-pressure and flow rate control. Various configurations of pipeline are optimised (based on a simplified minimum Froude number approach) for a hypothetical application requiring 24 t h"1 of pulverised coal over a total distance of 1800 m and demonstrate the advantages of selecting a stepped-diameter pipeline, in terms of minimising pressure, velocity and air flow requirements. 8.1 Further Work

Although the results obtained in this thesis have contributed significantly to predicting and understanding the pneumatic transportation of bulk solids, the following areas of investigation still require immediate attention.

(a) The equipment and test procedures that are used to measure the permeability (fluidisation) and deaeration characteristics of products need to be developed and standardised for the formulation of a unified classification technique (i.e. to evaluate the dense-phase suitability of bulk solids). Numerous experiments (as well as pneumatic conveying tests) should be carried out on a wide range of materials having narrow and wide particle size distributions. These investigations also should be extended to encompass the different possible modes of dense-phase conveying (e.g. by-pass pipelines, plug-phase, pulse-phase) and include an evaluation of

the relevance of mean (or median) particle diameter and size distribution to the prediction of fluidisation behaviour and slugging (dense-phase) performance,

other possible influential factors or variables (e.g. ratio of tapped to poured bulk density [43]) that may assist in the classification of fine or cohesive powders (e.g. [25]).

(b) Despite the developments made in improving the accuracy of the scale-up equations presented in Chapter 6, it is believed that the application of this macro or systems approach to design still is somewhat limited (e.g. approximations are used to allow for vertical lifts and bends). The correlation results presented in Section 7.4.3, indicate that a micro or components approach may be more useful and accurate in predicting pipeline operating conditions (even for large values of L and D). To evaluate the general applicability of this approach to design, the procedure proposed in Section 7.4.3 should be employed to examine a wide range of materials (e.g. fine powders and coarse products). These investigations also should include

178

a test program to identify the relevant variables that affect and define minimum transport behaviour (e.g. air density, minimum Froude number, mass flow ratio),

the introduction of additional instrumentation to monitor the pressure drop caused by bends,

an evaluation of undertaking the necessary experiments on only a small-scale test rig (i.e. replacing the receiving silo with a certified pressure vessel to generate high operating pressures and simulate long-distance conveying),

the development of accurate empirical correlations that will allow each component of a pipeline (e.g. bends, straight pipe sections, vertical lifts) to be analysed,

the development of computer software to optimise the correlations for solids friction factor and bend pressure loss.

These correlations may then be used in the development of an analysis package to predict with confidence the operating conditions for a given material and different configurations of pipeline (including long-distance, stepped-diameter pipelines).

(c) After a wide range of materials have been investigated in (b) above, it then will possible to conduct investigations into generalising correlations for solids friction factor and bend pressure loss. The importance of particle size distribution and a representative diameter also will need to be investigated for this purpose (e.g. refer to Werner [29]).

With the results presented in this thesis, together with those obtained from the above investigations, it then may be possible to predict with confidence and good reliability the most suitable mode of conveying and the pneumatic conveying performance for any material and configuration of pipeline (i.e. based on bench-type experiments and perhaps only a small-scale test rig).

179

CHAPTER 9

180

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56. Scott, A. M. The influence of particle properties on the pressure drop during the pneumatic transport of granular materials. Pneumotransport 4, Int. Conf. on the Pneumatic Transport of Solids in Pipes, Carmel-by-the-Sea, California, U.S.A., 26-28 June, 1978, Paper A3.

57. Molerus, O. Prediction of pressure drop with steady-state pneumatic conveying of solids in horizontal pipes. Chemical Engineering Science, Vol. 36, No. 12, 1981, pp. 1977-1984.

58. Marcus, R. D. Pneumatic conveying of bulk solids. Notes for Short Course on Pneumatic Conveying of Bulk Solids, T U N R A Bulk Solids Handling Research Associates, University of Newcastle, N.S.W., Australia, 1983.

59. Muschelknautz, F. and W. Krambrock. Simplified calculations on horizontal pneumatic feed pipes at high loading with finely divided granular products (In German). Chemie-lngenieur-Technik, Vol. 41, No. 21, 1969, pp. 1164-1172.

60. Chari, S. S. Pressure drop in horizontal dense-phase conveying of air-solid mixtures. AlChE Symposium Series-Fluidisation, Vol. 67, No. 116, 1970, pp. 77-84.

185

61. McLean, A.G. Flow Rate of Simple Bulk Solids from Mass Flow Bins. PhD Thesis, Dept. of Mech. Eng., The University of Wollongong, 1979.

62. Enstad, G. On the theory of arching in mass flow hoppers. Chemical Engineering Science, Vol. 30, 1975, pp. 1273-1283.

63. McLean, A. G. Blow tank design. Bulk Solids Handling, Vol. 5, No. 1, February, 1985, pp. 213-218.

64. Ergun, S. Fluid flow through packed columns. Chemical Engineering Progress, Vol. 48, No. 2, 1952, pp. 89-94.

65. Fortier, A. Two-phase turbulent steady flow of air and solid particles mixture with high mass concentration in a pipe. Pneumotransport 3, Int. Conf. on the Pneumatic Transport of Solids in Pipes, University of Bath, England, 7-9 April, 1976, Paper C2.

66. Wilson, K. C. Analytic modelling of energy consumption in solids pipelining. Pneumotransport 5, Int. Conf. on the Pneumatic Transport of Solids in Pipes, London, England, 16-18 April, 1980, Paper C3.

67. Wilson, K. C. A unified physically-based analysis of solid-liquid pipeline flow. Hydrotransport 4, Int. Conf. on the Hydraulic Transport of Solids in Pipes, Alberta, Canada, 18-21 May, 1976, Paper A1.

68. Stegmaier, W. Zur Berechnung der horizontalen pneumatischen Forderung feinkomiger Feststoffe. Fordern und Heben, Vol. 28, No. 5/6, 1978, pp. 363-366.

69. Weber, M. Correlation analysis in the design of pneumatic transport plant. Bulk Solids Handling, Vol. 2, No. 2, June, 1982, pp. 231-233.

70. Chambers, A. J. and R. D. Marcus. Pneumatic conveying calculations. 2nd Int. Conf. on Bulk Materials Storage, Handling and Transportation, Wollongong, 7-9 July, 1986. Organised by The Institution of Engineers, Australia. Preprints of Papers, pp. 49-52.

71. Wypych, P. W. and P. C. Arnold. Classification and prediction of fly ash handling characteristics for dense-phase and long-distance pneumatic transportation. TIZ-Fachberichte, Vol. 111, No. 11, 1987, pp. 753-761.

72. Bradley, M. S. A. and D. Mills. Approaches to dealing with the problem of energy loss due to bends. 13th Annual Powder & Bulk Solids Conf., Rosemont, Illinois, U.S.A., 9-12 May, 1988.

73. Bohnet, M. Advances in the Design of Pneumatic Conveyors. International Chemical Engineering, Vol. 25, No. 3, July, 1985, pp. 387-405.

74. Schade, B. Zum Ubergang Sprung-Strahnenforderung bei der Horizontalen Pneumatischen Feststofforderung. Dissertation, University of Karlsruhe, 1987.

186

APPENDIX A

Compilation of Particle Size Data (Samples 1 to 11, Table 5.1)

187

No. I

1 2 3 4 5 6 7 8 9 10 11

Particle Size Range Advi (nm)

0.0 - 4.7 4.7 - 6.6 6.6 - 9.4 9.4-13.0 13.0-19.0 19.0 - 27.0 27.0 - 38.0 38.0 - 53.0 53.0 - 75.0 75.0-106.0 106.0- 150.0

Average Size dVi (nm)

2.35 5.65 8.00 11.20 16.00 23.00 32.50 45.50 64.00 90.50 128.00

Mass % in Range AMj

6.0 5.1 6.1 7.2 9.7 12.4 14.3 15.7 12.3 7.4 3.8

Table A.1 Mass percentage frequency distribution for Tallawarra pulverised coal (Sample 1), using the Coulter Counter.

NO. i

1 2 3 4 5 6 7

Particle Size Range AdVi (nm)

2.0 - 4.0 4.0 - 7.0 7.0-10.0 10.0-20.0 20.0 - 40.0 40.0 - 70.0 70.0-100.0

Average Size dvi (nm)

3.0 5.5 8.5 15.0 30.0 55.0 85.0

Mass % in Range AM;

8.0 9.0 9.0 26.0 26.0 16.0 6.0

Table A.2 Mass percentage frequency distribution for Tallawarra fly ash (Sample 2), using the Coulter Counter.

No. i

1 2 3 4 5 6 7 8 9


2.0 - 4.0 4.0 - 7.0 7.0-10.0 10.0-20.0 20.0 - 40.0 40.0 - 70.0 70.0-100.0 100.0 - 200.0 200.0 - 300.0

Average Size dV| (nm)

3.0 5.5 8.5 15.0 30.0 55.0 85.0 150.0 250.0

Mass % In Range AM i

7.0 9.0 I 7.0 18.0 21.0 15.0 9.5 11.0 2.5

Table A.3 Mass percentage frequency distribution for Eraring fly ash (Sample 3), using the Coulter Counter.

188

No. i

1 2 3 4 5 6 7 '8


2.0 - 4.0 4.0 - 7.0 7.0-10.0

10.0-20.0 20.0 - 40.0 40.0 - 70.0 70.0-100.0

100.0 - 200.0

Average Size dv| (nm)

3.0 5.5 8.5 15.0 30.0 55.0 85.0 150.0

Mass % in Range AMf

8.5 8.0 6.5

20.5 21.5 21.0 8.0 6.0

Table A.4 Mass percentage frequency distribution for Munmorah fly ash (Sample 4), using the Coulter Counter.

No. I

1 2 3 4 5 6 7

Particle Size Range AdV| (nm)

4.0 - 7.0 7.0-10.0

10.0 - 20.0 20.0 - 40.0 40.0 - 70.0 70.0-100.0

100.0-200.0


5.5 8.5

15.0 30.0 55.0 85.0

150.0

Mass % In Range AM]

14.0 15.0 24.0 23.5 16.5 5.5 1.5

Table A.5 Mass percentage frequency distribution for Vales Point fly ash (Sample 5), using the Coulter Counter.

No. i

1 2 3 4 5 6 7 8

Particle Size Range AdV| (nm)

2.0 - 4.0 4.0 - 7.0 7.0 -10.0

10.0-20.0 20.0 - 40.0 40.0 - 70.0 70.0-100.0

100.0-200.0

Average Size dVi (nm)

3.0 5.5 8.5 15.0 30.0 55.0 85.0

150.0

Mass % In Range AMj

12.0 13.0 10.0 19.0 22.0 16.0 7.0 1.0

Table A.6 Mass percentage frequency distribution for Gladstone fly ash (Sample 6), using the Coulter Counter.

NO. i

1 2 3 4 5 6 7 8

Particle Size Range Advi Oim)

2.0 - 4.0 4.0 - 7.0 7.0-10.0

10.0-20.0 20.0 - 40.0 40.0 - 70.0 70.0-100.0

100.0-200.0

Average Size dV| (nm)

3.0 5.5 8.5 15.0 30.0 55.0 85.0 150.0

Mass % In Range AM]

16.0 17.0 13.0 25.0 19.0 10.0 0.0 0.0

Table A.7 Mass percentage frequency distribution for Wallerawang fly ash (Sample 7), using the Coulter Counter.

NO. i

1 2 3 4 5 6 7 8

Particle Size Range AdVi (nm) |

2.0 - 4.0 4.0 - 7.0 7.0-10.0

10.0-20.0 20.0 - 40.0 40.0 - 70.0 70.0-100.0 100.0-200.0


3.0 5.5 8.5 15.0 30.0 55.0 85.0

150.0

Mass % In Range AMj

17.0 16.0 10.0 15.5 15.5 14.0 7.5 4.0

Table A.8 Mass percentage frequency distribution for Liddell fly ash (Sample 8), using the Coulter Counter.

No. I

1 2 3 4 5 6 7

Particle Size Range Adpi (nm)

40.0 - 75.0 75.0-100.0

100.0-130.0 130.0- 150.0 150.0-200.0 200.0 - 300.0 300.0 - 420.0

Average Size dpi (nm)

57.5 87.5

115.5 140.0 175.0 250.0 360.0

Mass % in Range AMj

4.0 7.0

29.0 26.0 20.0 12.0 2.0

Table A.9 Mass percentage frequency distribution for P V C powder (Sample 9), using the sieve test.

190

No. i

1 2 3 4 5 6 7 8 9 10 11 12

Particle Size Range Adpi (nm)

60.0-100.0 100.0-200.0 200.0 - 300.0 300.0 - 400.0 400.0 - 500.0 500.0 - 600.0 600.0 - 700.0 700.0 - 800.0 800.0 - 900.0 900.0-1000.0 1000.0 - 1200.0 1200.0 - 1500.0

Average Size dpi (nm)

80.0 150.0 250.0 350.0 450.0 550.0 650.0 750.0 850.0 950.0 1100.0 1350.0

Mass % in Range AMj

3.2 12.1 13.4 13.2 11.3 6.8 11.1 6.6 6.4 5.1 8.1 2.7

Table A.10 Mass percentage frequency distribution for screened coke (Sample 10), using the sieve test.

No. i

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16

Particle Size Range AdVi (nm)

0.0-5.8 5.8 - 7.2 7.2-9.0 9.0-11.4

11.4- 14.5 14.5 - 18.5 18.5-23.7 23.7 - 30.3 30.3 - 39.0 39.0 - 50.2 50.2 - 64.6 64.6 - 84.3 84.3- 112.8 112.8- 160.4 160.4-261.7 261.7-564.0


2.90 6.50 8.10 10.20 12.95 16.50 21.10 27.00 34.65 44.60 57.40 74.45 98.55 136.10 211.05 412.85

Mass % in Range AMj

0.9 0.4 0.2 0.7 0.8 0.4 1.1 1.3 4.1 7.4 7.2 14.5 20.1 20.2 15.1 5.6

Table A.11 Mass percentage frequency distribution for coarse fly ash (Sample 11), using the Malvern analyser.

191

APPENDIX B

Modified Slugging Diagrams based on Dixon [23,39] and Clift et al. [41]

192

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197

APPENDIX C

Compilation of Operating Conditions for Correlation Analysis (Samples 1 to 11, 12 and 13, Table 7.2)

Test Rig

A1

A3

mf (kg s"1)

.0068

.0123

.0115

.0109

.0105

.0105

m s (kg s-1)

2.62 3.99 4.47 4.50

1.41 1.47

Apt (kPa) |

75 87 95 95

110 115

Table C.1 Steady-state operating conditions of pulverised coal (Sample 1) for Test Rigs A1 (L = 25 m & D = .052 m) and A3 (L = 96 m & D = .052 m).

Test Rig

B1

mt (kg s-1)

.018

.030

.040

.050

.016

.026

.036

.044

.014

.024

.038

I

m s (kg s'1)

1.0 1.0 1.0 1.0 2.0 2.0 2.0 2.0 3.0 3.0 3.0

Apt (kPa)

66 65 67 74 110 109 110 116 155 154 157

Table C.2 Steady-state operating conditions of Tallawarra fly ash (Sample 2) for Test Rig B1 (L = 71 m & D = .052 m).

Test Rig

B1

mf (kg s'1)

.010

.020

.036

.050

.016

.028

.042

.056

.022

.028

.036

.044

.050

m s (kg s*1)

1.0 1.0 1.0 1.0 2.0 2.0 2.0 2.0 3.0 3.0 3.0 3.0 3.0

Apt (kPa)

61 56 60 75 102 94 100 120 ; 140 135 134 141 149

Table C.3 Steady-state operating conditions of Eraring fly ash (Sample 3) for Test Rig B1 (L = 71 m & D = .052 m).

Test Rig

B1

mf (kg s_1)

.009

.020

.032

.042

.014

.022

.034

.046

.014

.030

.040

.049

m s (kg s"1)

1.0 1.0 1.0 1.0 2.0 2.0 2.0 2.0 3.0 3.0 3.0 3.0

Apt (kPa)

70 60 60 68 113 103 100 110 152 135 140 162

Table C.4 Steady-state operating conditions of Munmorah fly ash (Sample 4) for Test Rig B1 (L = 71 m & D = .052 m).

200

Test Rig

B1

mf (kg s"1)

.008

.020

.040

.068

.017

.040

.056

.068

.023

.036

.052

.068

m s (kg s*1)

0.5 0.5 0.5 0.5 1.5 1.5 1.5 1.5 2.5 2.5 2.5 2.5

Apt (kPa)

48 40 46 72 109 98 111 127 160 145 145 166

Table C.5 Steady-state operating conditions of Vales Point fly ash (Sample 5) for Test Rig B1 (L = 71 m & D = .052 m).

Test Rig

B1

C3

mf (kg s"1)

.020

.030 !

.045

.065

.023

.030

.045

.069

.025

.035

.055

.071

.200

.300

.400

.150

.300

.400

.300

.450

m s (kg s*1)

1.0 1.0 1.0 1.0 2.0 2.0 2.0 2.0 3.0 3.0 3.0 3.0

5.0 5.0 5.0

10.0 10.0 10.0 15.0 15.0

, -—.

Apt (kPa)

72 71 76 94

118 115 116 148 164 158 164 190

70 78 87

110 122 130 170 182

Table C.6 Steady-state operating conditions of Gladstone fly ash (Sample 6) for Test Rigs B1 (L = 71 m & D = .052 m) and C3 (L= 162 m & D = .105 m).

Test Rig

B1

mf (kg s'1)

.010

.030

.050

.070

.014

.030

.050

.070

.026

.040

.050

.062

m s (kg S"1)

0.5 0.5 0.5 0.5 1.5 1.5 1.5 1.5 2.5 2.5 2.5 2.5

Apt (kPa)

45 45 57 76 105 93 102 130 150 136 141 156

Table C.7 Steady-state operating conditions of Wallerawang fly ash (Sample 7) for Test Rig B1 (L = 71 m & D = .052 m).

Test Rig

B1

mf (kg s"1)

.020

.032

.048

.070

.024

.034

.048

.070

.026

.038

.056

.070

m s (kg s'1)

1.0 1.0 1.0 1.0 2.0 2.0 2.0 2.0 2.5 2.5 2.5 2.5

Apt (kPa)

79 78 83 111 126 124 129 162 153 150 163 190

Table C.8 Steady-state operating conditions of Liddell fly ash (Sample 8) for Test Rig B1 (L = 71 m & D = .052 m).

Test Rig

C1

C3

mf (kg s"1)

.035

.080

.120

.165

.045

.080

.120

.165

.055

.120

.165

.060

.090

.130

.065

.080

.100

.112

.300

.450

.134

.300

.450

.158

.300

.450

.184

.300

.450

m s (kg s'1)

2.0 2.0 2.0 2.0 4.0 4.0 4.0 4.0 6.0 6.0 6.0 8.0 8.0 8.0

10.0 10.0 10.0 5.0 5.0 5.0

10.0 10.0 10.0 15.0 15.0 15.0 20.0 20.0 20.0

Apt (kPa)

108 108 122 156 170 172 188 226 222 246 304 270 280 320 310 316 334 66 74 89 98

108 124 132 143 163 167 180 202

Table C.9 Steady-state operating conditions of fly ash/cement mix (Sample 12) for Test Rigs C1 (L = 162 m & D = .060 m) and C3 (l_= 162 m & D = .105 m).

Test Rig

-

mf (kg s_1)

1.389

m s (kg s_1)

13.89

Apt (kPa)

350

Table C.10 Steady-state operating conditions of fly ash [59] (Sample 13) for L = 1200 m & D = .200 m.

APPENDIX D

Summary of Solids Friction Factor Calculations for Pulverised Brown Coal (Test-Design Procedure, Section 7.4.3)

204

mf (kg s-1)

.120

.127

.120

.096

.087

.125

.180

.200

.200

m s

(kg s-1)

2.50 2.94 1.80 2.08 2.14 1.67 1.43 1.79 2.94

Ap-, (kPa)

29 32 25 27 28 23 27 32

! 38

Pfm (kPag)

14.5 16.0 12.5 13.5 14.0 11.5 13.5 16.0 19.0

Pfm (kg m-3)

1.372 1.390 1.349 1.360 1.366 1.337 1.360 1.390 1.426

Vfm (m s"1)

10.098 10.550 10.276 8.149 7.353 10.800 15.280 16.615 16.200

Frm (-)

9.950 10.395 10.125 8.030 7.245 10.641 15.055 16.371 15.962

Xi (-)

.2901

.2895

.2458

.4184

.5306

.2065

.1190

.1167

.1422

Apf (kPa)

2.043 2.218 2.108 1.373 1.136 2.310 4.393 5.138 4.916

Xf

(-)

.0204

.0201

.0207

.0213

.0215

.0207

.0194

.0187

.0184

Xs (-)

.0129

.0116

.0150

.0183

.0207

.0139

.0125

.0109

.0084

Table D.1 Solids friction factor calculations for pipe section No. 1 (Di = 0.105 m & AL-, = 150.0 m).

mf (kg s-1)

.120

.127

.120

.096

.087

.125

.180

.200

.200

m s

(kg s-1)

2.50 2.94 1.80 2.08 2.14 1.67 1.43 1.79 2.94

Ap2

(kPa)

81 91 73 72 72 69 72 97 108

Pfm (kPag)

69.5 77.5 61.5 63.0 64.0 57.5 63.0 80.5 92.0

Pfm (kg m-3)

2.026 2.121 1.931 1.949 1.960 1.883 1.949 2.157 2.293

Vfm (m s-1)

11.495 11.621 12.061 9.561 8.612 12.881 17.926 17.998 16.925

Frm (-)

12.896 13.036 13.531 10.725 9.661 14.450 20.110 20.190 18.987

x2 (-)

.1877

.1971

.1613

.2508

.3072

.1370

.0713

.0862

.1020

Apf

(kPa)

8.020 8.390 8.504 5.600 4.644 9.456 17.915 18.930 17.628

Xf (-)

.0186

.0182

.0188

.0195

.0198

.0188

.0178

.0168

.0166

A-s (-)

.0081

.0077

.0095

.0107

.0117

.0089

.0067

.0077

.0058

Table D.2 Solids friction factor calculations for pipe section No. 2 (D2 = 0.081 m & Al_2 = 261.0 m).

205

mf (kg s-1)

.120

.127

.120

.096

.087

.125

.180

.200

.200

m s

(kg s-1)

2.50 2.94 1.80 2.08 2.14 1.67 1.43 1.79 2.94

AP3 (kPa)

165 175 147 143 150 138 145 166 209

Pfm (kPag)

192.5 210.5 171.5 170.5 175.0 161.0 171.5 212.0 250.5

Pfm (kg m-3)

3.487 3.701 3.238 3.226 3.279 3.113 3.238 3.719 4.176

Vfm (m s-1)

9.203 9.177 9.912 7.959 7.095 10.739 14.868 14.382 12.807

Frm

(-)

11.185 11.154 12.047 9.673 8.624 13.052 18.071 17.481 15.566

x3 (-)

.1975

.1985

.1634

.2475

.3213

.1359

.0716

.0763

.1079

Apf (kPa)

15.000 15.697 16.299 10.881 8.863 18.374 34.609 36.641 31.820

Xf

(•)

.0180

.0178

.0181

.0188

.0190

.0181

.0171

.0168

.0164

Xs (")

.0086

.0078

.0097

.0106

.0123

.0088

.0069

.0066

.0062

Table D.3 Solids friction factor calculations for pipe section No. 3 (D3 = 0.069 m & Al_3 = 390.0 m).

mf

(kg s-1)

.120

.127

.120

.096

.087

.125

.180

.200

.200

m s

(kg s-1)

2.50 2.94 1.80 2.08 2.14 1.67 1.43 1.79 2.94

Ap4

(kPa)

91 96 72 76 75 76 76 85 110

Pfm (kPag)

320.5 346.0 281.0 280.0 287.5 268.0 282.0 337.5 410.0

Pfm (kg m-3)

5,008 5.311 4.539 4.527 4.616 4.384 4.551 5.210 6.072

Vfm (m s-1)

8.475 8.457 9.351 7.500 6.666 10.084 13.990 13.577 11.650

Frm

(-)

11.046 11.023 12.188 9.776 8.689 13.143 18.235 17.696 15.186

X4

(-)

.2078

.2076

.1490

.2451

.3003

.1400

.0701

.0727

.1096

Apf

(kPa)

9.092 9.526 10.158 6.706 5.498 11.251 21.353 22.712 19.301

Xf

(-)

,0208 .0206 .0210 .0216 .0220 .0207 .0197 .0194 .0192

Xs (-)

.0090

.0081

.0085

.0103

.0113

.0089

.0063

.0060

.0061

Table D.4 Solids friction factor calculations for pipe section No. 4 (D4 = 0.060 m & AL4 = 146.0 m).

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University of Wollongong Thesis Collection University of ... · quite extensively in the design of pneumatic conveying systems, is investigated and found to be inadequate in particular