pulverised coal injection into blast furnace - a …

PULVERISED COAL INJECTION INTO BLAST FURNACE -

A PRACTICAL STUDY OF AN INTEGRATED IRON AND

STEEL WORKS

S. W. DU

I

PULVERISED COAL INJECTION INTO BLAST FURNACE -

A PRACTICAL STUDY OF AN INTEGRATED IRON AND

STEEL WORKS

A Thesis Submitted in Fulfillment of

the Requirement for the Degree of

Doctor of Philosophy

by

SHAN-WEN DU

(MChE)

Department of Chemical Engineering

The University of Newcastle, Australia

December, 2015

II

I hereby certify that this thesis is submitted in the form of a series of

published papers of which I am a joint author. I have included as part of the

thesis a written statement from each co-author; and endorsed by the Faculty

Assistant Dean, attesting to my contribution to the joint publications.

Signed: _______________________

( Shan-Wen Du)

III

To my family

IV

ACKNOWLEDGEMENTS

First and foremost, I would like to express my sincere gratitude to Professor John Lucas

for his enthusiastic supervision and guidance throughout the period of research. Without

his support and assistance, this work would not have been possible accomplished.

Great thanks go to Professor Terry Wall and Dr. Harold Rogers for their experienced

advice in coal combustion modelling and experiment, to Professor Jian-Long Yu and

Dr. Chatphol Meesri for their encouragements, and to Professor Ai-Bing Yu (New

South Wales University) and Dr, Yan-Song Shen for sharing their research works in

CFD modelling.

I would like to extend my appreciation to Professor Wei-Hsin Chen (National Cheng

Kung University in Taiwan) for his valuable suggestions to coal and biofuel combustion

tests, to Professor Chien-Hsiung Tsai (National Pingtung University of Science and

Technology) for his selfless help in CFD modelling, to Dr. Cheng-Peng Yeh for his

assistance with the prediction of raceway shape, and to Professor Wei-Kao Lu

(McMaster University) for his insightful lectures on blast furnace theory at China Steel

Corporation (CSC).

At CSC, continuous encouragement from Mr. Sing-Tsu Tsai is deeply appreciated. I

would like to thank Mr. Chung-Ken Ho (R&D, BF), Dr. Yung-Chang Ko (R&D, BF),

Mr. Ming-Tsai Hung (R&D, Cokmaking), Dr. Li-Heng Hsieh (R&D Sintering), Mr.

Chi-Sheng Chou (BF) and Mr. Che-Hsiung, Tung (BF) for their challenging comments

and constructive suggestions to this work.

Last but not least, I would like to acknowledge my wife Dr. Shao-Wen Su and my

children Won-Yu, Jia-Yu and Jei-Ruei for their support. They are always my motivation

to finish this thesis.

V

ABSTRACT

The economic benefits of pulverised coal injection (PCI) into blast furnace include a

reduction in the cost of hot metal, resulting primarily from decreased coke consumption

and an increase in hot metal production. Since coal is consumed directly, without going

through the cokemaking plant, PCI is also thought to be environmentally friendly.

Therefore, PCI has become a standard practice in many blast furnaces worldwide. To

improve the performance of PCI operation, a comprehensive understanding of

pulverised coal combustion behaviours are required. The goal of this thesis was to study

the coal burning characteristics in the regions of blowpipe, tuyere and raceway through

both numerical and experimental methodology.

From the validation of model, the calculation region and the application of calculated

results in practice, the PCI combustion model was developed through 4 phases in this

work: (1) validation of the coal combustion model by comparing its predictions with

experimental data; (2) investigation into the influence of operation conditions to coal

burnout in the regions of blowpipe and tuyere; (3) performance evaluation of coal blend

injection in terms of pressure loss caused by combustion within a simplified raceway

space; and (4) examination of combustion characteristics of oxy-coal injection

technology in the regions of blowpipe, tuyere and raceway, which is a porous space

featured by Eulerian-Eulerian multi-fluid approach.

In the first phase, the performance of coal devolatilisation models and kinetic

parameters were validated by comparing predicted gas temperature profiles with the

experimental results of Burgess et al. (1983). It is found that the kinetic parameters

proposed by Ubhayakar et al (1976) for the two competing devolatilisation model

permit a reasonable simulation of the measured results for blast furnace conditions.

The coal combustion in the regions of blowpipe and tuyere was modelled under the

conditions of CSC’s No3 blast furnace in the second phase. The influence of operation

conditions to coal burnout was comprehensively studied. It is found early ignition can

be achieved with higher coal burnout when the double lance is employed instead of the

VI

single one. Accordingly, the injection lance used at CSC was changed from the single to

the double air-cooled coaxial lance arrangement in 2002.

In the third phase, the calculation was extended to the raceway with simplified

configuration. The performance of coal blend injection was examined. As indicated in

the calculation results, a decrease in coal burnout is found with decreasing the coal

volatile content, while the pressure loss within the raceway can be abated due to less

volatile released to gas and moderate gas expansion in the combustion region. With

improved permeability, more hot blast air can be introduced into the blast furnace for

higher productivity. Consequently, the high volatile coal injection was replaced by the

coal blend (mixtures of high and low volatile coals) injection at CSC in 2003.

In the last phase, the Eulerian-Eulerian multi-fluid model was employed for the

prediction of raceway configuration with consideration of coke combustion in all coke

operation. Validation work against measured raceway shape and gas composition

distribution by Nogami et al. (2005) indicates that the model is acceptable for the

simulation. The calculation results show the oxy-coal lance injection enables to fulfil

two contradictory conditions at the same time: (1) to retard the coal combustion for

moderating the pressure loss in the upstream of coal plume; and (2) to enhance coal

combustion and reduce unburnt char generation in the downstream of coal plume.

Taking these advantages from the oxy-coal lance injection, blast furnaces can be

operated with more blast for higher productivity, or with higher PCI rate for lower fuel

cost, thereby achieving the goal of hot metal production with energy saving.

In this work, a drop tube furnace has been established and used to provide fundamental

insights on PCI coal combustion behaviours. The experiments were carried out in three

stages. In the first stage, the volatile release and the generation of char particle and tiny

aerosols in the region of coal plume were studied. Only the tested low volatile coal

(HGI=85) with larger size (100-200 mesh) exhibits fragmentation during heating. This

may encourage the use of low volatile coals in granular coal injection. Significant char

agglomeration is found for both tested high and low volatile coals with smaller size

VII

(200-325 mesh). It implies that excessive grinding may be avoided in PCI operation.

Considering the generation of tiny aerosols composed of soot particles and tar droplets,

it is mainly determined by the content of volatile matter and elemental oxygen.

In the second stage, a technology has been developed and employed at CSC to evaluate

the combustion efficiency of PCI coals. It is found that the coal burnout increases with

decreasing the fuel ratio (FC/VM), except for certain coals departing from the general

trend. It can be explained by the effect of maceral content to coal combustion. When the

coal size is smaller than 200 mesh, the burnout can not be improved further, resulting

from the agglomeration of fine particles. In the PCI operation at CSC, the coal quantity

passing through 200 mesh has been reduced from 80 to 60%.

The experiments for the last stage aim to gain a fundamental insight into the combustion

characteristics of pulverised biofuels under conditions pertinent to the raceway of blast

furnace. From the van Krevelen diagram, it is found that the rate of hydrogen release

from biomass fuels is faster than that of oxygen during the pre-treatment. An increase in

pretreating temperature almost linearly decreases the burnout of biofuels. As revealed in

the experimental results, the fuel properties, such as fuel ratio, burnout, and ignition

temperature, of biomass torrefied at 300 °C or pyrolysed between 400 and 500 °C, are

between a high-volatile bituminous coal and a low-volatile one. Therefore, the

pretreated biomass can partially replace the coals consumed for PCI and blends with

coals to keep reasonable burnout in raceways.

It is emphasised that, due to the objectives of this thesis, some results or

countermeasures obtained from the comprehensive experimental and numerical studies

have been taken into PCI operation at CSC. This seems as a limitation of this study, but

it may have a wide range of applications for the improvement of PCI operation.

,

VIII

LIST OF PUBLICATIONS INCLUDED AS PART OF THE THESIS

(1) Du, S. W., and Chen, W. H. (2006), Numerical prediction and practical

improvement of pulverised coal combustion in blast furnace, International

Communications in Heat and Mass Transfer, 2006, vol. 33, p. 327-334.

(2) Du, S. W., Chen, W. H. and Lucas, J. A. (2007), Performances of pulverised coal

injection in blowpipe and tuyere at various operational conditions, Energy

Conversion and Management, vol. 48, p. 2069-78.

(3) Du, S. W., Yeh, C. M., Yang, M. K. and Ho, C. K. (2004), Practice of high

productivity at No.3 blast furnace of China Steel Corporation", Conference of

Association for Iron and Steel Technology Proceedings (USA), p. 195-204.

(4) Du, S. W., Yeh, C. P., Chen, W. H., Tsai, C. H. and Lucas, J. A. (2015), Burning

characteristics of pulverized coal within blast furnace raceway at various injection

operations and ways of oxygen enrichment, Fuel, vol. 143, p. 98-106.

(5) Chen, W. H., Du, S. W. and Yang, T. H. (2007), Volatile release and particle

formation characteristics of injected pulverised coal in blast furnace, Energy


(6) Du, S. W., Chen W. H. and Lucas J. A. (2010), Pulverised coal burnout in blast

furnace simulated by a drop tube furnace, Energy, vol. 35, p. 576-581.

(7) Du, S. W., Chen, W. H. and Lucas, A. J. (2014), Pretreatment of biomass by

torrefaction and carbonization for coal blend used in pulverized coal injection,

Bioresource Technology, vol. 161, p. 333-339.

XII

STATEMENT 4

Cheng-Peng Yeh, Wei-Hsin Chen, Chien-Hsiung Tsai and John Lucas attest that

Research Higher Degree candidate Shan-Wen Du contributed to the (1) construction of

calculation model; (2) case studies; and (3) writing to the paper entitled: Burning


operations and ways of oxygen enrichment, published in Fuel (2015).

Cheng-Peng Yeh

Date:

Wei-Hsin Chen

Date:

Chien-Hsiung Tsai

Date:

John Lucas

Date:

Shan-Wen Du

Date:

Suzanne Ryan (Assistant Dean Research Training)

Date:

XVI

LIST OF ADDITIONAL PUBLICATIONS

(1) Du, S. W., Ho, C. K., Tsai, S. T. and Yeh, C. M. (2001), Development of pulverised

coal injection lance at China Steel Corporation”, World Coal, vol. 10, p. 39-42.

(2) Chen, W. H., Du, S. W., Yang, H. H. and Wu, J. S. (2008), Formation

characteristics of aerosol particles from pulverised coal pyrolysis in

high-temperature environments, Journal of the Air & Waste Management

Association, vol. 58, p. 702-710.

(3) Chen, W. H., Du, S. W., Tsai, C. H. and Wang, Z. Y. (2011), Torrefaction of

biomasses in a drop tube furnace to evaluate their utility in blast furnaces”,


(4) Yeh, C. P. Du, S. W., Tsai, C. H. and Yang, R. J. (2012), Numerical analysis of

flow and combustion behavior in tuyere and raceway of blast furnace Fueled with

pulverized coal and recycled top gas”, Energy, vol. 42, p. 233-240.

(5) Du, S. W., Ho, C. K. and Tung, C. H. (2013), Numerical Investigations into burning

characteristics of pulverised coal within the BF raceway at various injection lances”,

Proceedings of the Fifth Baosteel Biennial Academic Conference. Shanghai, China,

p. A300-A304.

XVII

CONTENTS

ACKONWLEGEMENTS IV

ABSTRACT V

LIST OF PUBLICATIONS INCLUDED AS PART OF THE THESIS VIII

STATEMENT OF CONTRIBUTION OF OTHERS IX

LIST OF ADDITIONAL PUBLICATIONS XVI

LIST OF FIGURES XXV

LIST OF TABLES XXXII

CHAPTER 1 INTRODUCTION 1

1.1 Background 2

1.2 Auxiliary Fuel injection into blast furnace 5

1.3 Achievement of high coal injection rate 6

1.4 Utilisation of auxiliary fuels at CSC 8

1.5 CO2 emission at CSC 11

1.6 Objectives of the work 11

1.7 Thesis outline 12

CHAPTER 2 LITERATURE REVIEW 15

2.1 Coal combustion experiments under simulating PCI operation

conditions 16

2.1.1 Experiments using empty combustion rig 17

2.1.1.1 Effect of volatile matter content on combustion 17

2.1.1.2 Effect of coal size on combustion and granular coal

injection 19

2.1.1.3 Effect of hot blast conditions on combustion 19

2.1.1.4 Effect of injection rate on combustion 20

2.1.1.5 Effect of lance configurations on combustion 20

XVIII

2.1.1.6 Effect of co-injection on combustion 20

2.1.1.7 Effect of coal blend operation 21

2.1.1.8 Ignition and combustion of volatile matters 21

2.1.1.9 Ash fusion temperature 21

2.1.2 Experiments by coke-packed bed rigs and actual blast furnace 21

2.1.2.1 Effect of lance configuration on combustion 22

2.1.2.2 Effect of hot blast conditions on combustion 23

2.1.2.3 Movement of small coke within raceway 23

2.1.2.4 Effect of volatile content on combustion 24

2.1.2.5 Raceway control during operation at CSC 25

2.1.3 Coal combustion experiments by drop tube furnace 26

2.1.4 Summary of factors affecting coal combustion from

experiments 28

2.2 Modelling of pulverised coal combustion in blast furnace 29

2.2.1 Development of one-dimensional model 29

2.2.1.1 Model of Kuwabara et al. 30

2.2.1.2 Model of Burgess et al. 30

2.2.1.3 Model of He et al. 31

2.2.1.4 Model of Jamaluddin et al. 32

2.2.1.5 Model of Yamagata et al. 33

2.2.1.6 Model of Sato et al. 34

2.2.1.7 Summary of 1-D model 35

2.2.2 Development of two-dimensional model 44

2.2.2.1 Model of Aoki et al. 44

2.2.2.2 Model of Takeda and Lockwood 48

2.2.2.3 Model of Haywood et al. 51


2.2.3 Development of 3-D model 54

2.2.3.1 Model of Picard 54

XIX

2.2.3.2 Model of Guo et al. 56

2.2.3.3 models of Shen et al. 60

2.2.3.4 Model of by Gu et al. 68

2.2.3.5 Model of Nogami et al. 70


2.3 Sub-models for integrated calculation 73

2.3.1 Devolatilisation of coal 76

2.3.1.1 Single overall reaction model 77

2.3.1.2 Two competing reaction model 77

2.3.2 Char Oxidation 84

2.3.2.1 Field approach 84

2.3.2.2 Gibb Mode 87

2.3.3 Turbulence model 88

2.3.4 Gas combustion in turbulent flow field 89

2.3.4.1 Probability density function (PDF) of turbulence

chemistry 89

2.3.4.2 Eddy break up and eddy dissipation models 90

2.3.5 Lagrangian approach 91

2.3.6 Summary 93

2.4 Raceway shape 93

2.4.1 Observation of raceway 93

2.4.2 CSC’s raceway shape prediction model 96

2.4.2.1 Validation of raceway shape prediction model 99

2.4.2.2 Prediction of raceway shape in an operating blast

furnace 100

2.4.3 Summary of raceway shape prediction 103

2.5 Injection of biofuel into blast furnace 103

2.5.1 Combustion experiments and modelling 104

2.5.2 Summary of the biofuel injection 111

XX

2.6 Summary 111

2.7 Methodology 111

CHAPTER 3 NUMERICAL PREDICTION AND PRACTICAL

IMPROVEMENT OF PULVERIZED COAL

COMBUSTION IN BLAST FURNACE

113

Abstract 114

Nomenclature 115

Greek symbols 115

Subscripts 115

3.1 Introduction 116

3.2 Mathematical Formulation 117

3.2.1 Burning process of pulverized coal 117

3.2.2 Momentum and energy balance of a coal particle 117

3.2.3 Model of devolatilisation of coal particle 118

3.2.4 Turbulent combustion model 119

3.3 Results and discussion 120

3.3.1 Numerical validation and parameter selection 120

3.3.2 Impact of injection pattern 123

3.3.3 Practical improvement of blast furnace 125

3.4 Conclusions 126

CHAPTER 4 PERFORMANCES OF PULVERIZED COAL

INJECTION IN BLOWPIPE AND TUYERE AT

VARIOUS OPERATIONAL

127

Abstract 128

Nomenclature 129

Greek symbols 129

Subscripts 130


XXI

4.2 Methodology 132

4.2.1 Gas-phase continuity and momentum equations 133

4.2.2 Coal particle momentum and energy equations 135


4.3.1 Trajectories and residence times of coal particles 140

4.3.2 Injection pattern 140

4.3.3 Oxygen concentration and hot blast temperature 143

4.3.4 Hot blast temperature 144

4.3.5 Mass flow rate of carrier gas 144

4.3.6 Installation of ceramic sleeve 145

4.4 Conclusions 146

CHAPTER 5 PRACTICE OF HIGH PRODUCTIVITY AT NO

3 BLAST

FURNACE OF CHINA STEEL CORPORATION 148

Abstract 149


5.2 Development of low flux sinter 150

5.3 Establishment of burden terrace 152

5.4 Development of one bit drilling method 154

5.5 Coal blend injection 155

5.5.1 Analysis of permeability of the furnace 155

5.5.2 Coal Combustion model within tuyere-raceway area 157

5.5.3 Calculation results and discussion 161

5.5.4 Plant trial of coal blend injection 162

5.6 Increase of hot metal production in No3 blast furnace 163

5.7 Conclusions 164

CHAPTER 6 BURNING CHARACTERISTICS OF PULVERIZED

COAL WITHIN BLAST FURNACE RACEWAY AT

VARIOUS INJECTION OPERATIONS AND WAYS OF

OXYGEN ENRICHMENT

165

XXII

Abstract 166

Nomenclature 167

Greek symbols 167

Subscripts 168


6.2 Methodology 171

6.2.1 Gas-particle flow 171

6.2.1.1 Gas phase 171

6.2.1.2 Single particle in dispersed phase 172

6.2.1.3 Turbulence model 172

6.2.2 Turbulent combustion 173

6.2.3 Devolatilization of coal 174

6.2.4 Physical geometry and operating conditions 175


6.3.1 Trajectories of coal particles 177

6.3.2 Oxygen consumption within the combustion region 179

6.3.3 Ignition and temperature distribution 181

6.3.4 Combustion efficiency of coal particles 184

6.3.5 Pressure loss 185

6.4 Conclusions 186

CHAPTER 7 VOLATILE RELEASE AND PARTICLE FORMATION

CHARACTERISTICS OF INJECTED PULVERIZED

COAL IN BLAST FURNACES

188

Abstract 189


7.2 Experiment 192


7.3.1 Devolatilisation extent 195

XXIII

7.3.2 Particle formation of the low-volatile bituminous coal 196

7.3.3 Particle formation of the high-volatile bituminous coal 199

7.3.4 Aerosol formation and reactivity 204

7.3.5 Reactivity of char and soot 204

7.4 Conclusions 207

CHAPTER 8 PULVERIZED COAL BURNOUT IN BLAST FURNACE

SIMULATED BY A DROP TUBE FURNACE 208

Abstract 209


8.2 Experiments 212

8.2.1 Reaction system 212

8.2.2 Experimental procedure and conditions 214


8.3.1 Combustion efficiency of individual coals 216

8.3.2 Influences of reaction temperature and particle size 219

8.3.3 Burnout of blended coals 221

8.4 Conclusions 223

CHAPTER 9 PRETREATMENT OF BIOMASS BY

TORREFACTION AND CARBONIZATION FOR COAL

BLEND USED IN PULVERIZED COAL INJECTION

224

Abstract 225


9.2 Experimental 228

9.2.1 Materials and preparation 228

9.2.2 Burnout and ignition tests 230


9.3.1 Proximate analysis and van Krevelen diagram 231

XXIV

9.3.2 Solid yield and energy yield 234

9.3.3 Ignition and burnout 238

9.4 Conclusions 241

CHAPTER 10 CONCLUSIONS AND RECOMMENDATIONS 243


10.2 Achievements and conclusions 244

10.2.1 Modelling 244

10.2.2 Coal combustion experiments 247

10.3 Recommendations 249

10.3.1 Raceway control 249

10.3.2 Improvement of permeability by charging pattern 252

REFERENCES 254

XXV

LIST OF FIGURES

Figure 1.1 Schematic diagram of internal structure with five district zones

in a blast furnace. 2

Figure 1.2 Reduction reactions in blast furnace (Biswas, 1981). 3

Figure 1.3 Change in blast furnace operating conditions in Japan

(Ariyama et al, 2007). 6

Figure 1.4 Changes and problems of blast operation with high injection

rate. 9

Figure 2.1 Schematic of combustion rig attached with an empty

combustion chamber. 17

Figure 2.2 Influence of blast temperature on Q-factor. 18

Figure 2.3 Degree of combustion of the coals as a function of injection

rate. 20

Figure 2.4 Schematic of combustion rig attached with a coke bed. 22

Figure 2.5 Effect of lance arrangement on pulverised coal flow and

combustion efficiency by hot model (η (%): combustion at 300 and

600 mm from lance tip). 23

Figure 2.6 Change of coal burnout at the height of 700mm above tuyere

level with change of volatile content in coals. 24

Figure 2.7 Image of coal flow patterns at CSC. (a) single lance; (b) double

lance injection. 25

Figure 2.8 Image of blockage of tuyere by un-melted scab. 26

Figure 2.8 Schematic representation of a DTF. 27

Figure 2.9 Mechanism of pulverised coal combustion. 27

Figure 2.10 Critical operational factors on coal burnout. 28

Figure 2.11 Schematic view of combustion in front of a tuyere. 30

Figure 2.12 Schematic view of combustion with the injection of

pulverised coal. 31

Figure 2.13 Gas flow, entrainment and disentrainment patterns inside the

raceway. 32

Figure 2.14 Relationship between coal burnout and injection rate. 33

Figure 2.15 Modelling concept of pulverised coal flows in blow pipe. 34

Figure 2.16 Influence of injection lance on coal burnout along tuyere axis. 34

XXVI

Figure 2.17 Calculation domain in the region of blowpipe-tuyere. 45

Figure 2.18 Effect of coal volatile content on gas distribution in the

blowpipe. 46

Figure 2.19 Calculated particle trajectories of low volatile coal injection in

the blowpipe. 46

Figure 2.20 Calculated particle trajectories of high volatile coal injection

in the blowpipe. 47

Figure 2.21 Distributions of gas composition and gas temperature along

the centre line of tuyere. 47

Figure 2.22 Schematic representation of raceway structure used in the

simulation. 48

Figure 2.23 Contours of gas temperature and oxygen concentration in the

regions of blowpipe and raceway. 49

Figure 2.24 Lance design for coal injection. 50

Figure 2.25 Coal burnout comparisons for the various modifications of

lance. 51

Figure 2.26 Schematic of the experimental combustion rig. 52

Figure 2.27 Oxygen mass fraction (top) and gas temperature contours

(bottom) in the near injector region. low volatile coal (a) low volatile

coal; (b) high volatile coal injection. 52

Figure 2.28 3D meshed calculation domain. 55

Figure 2.29 Trajectories of coal particles within the tuyere and raceway. 55

Figure 2.30 Evolution of coal residence time within the raceway versus

particle diameter. 55

Figure 2.31 Main dimensions (in mm) of the coal combustion model (plan

view). 57

Figure 2.32 Gas velocity vectors in Y-Z plane (a) vector length to scale,

and (b) vector normalised showing recirculation zone. 57

Figure 2.33 Typical particle trajectories with colour scaled to particle size. 58

Figure 2.34 Gas species fraction isopleths in Y–Z plane. (a) Oxygen; (b)

volatile matter. 58

Figure 2.35 Calculated mass fraction of volatiles (a) and calculated mean

coal burnout (b) as a function of distance from the lance tip along the

centreline averaged for different particle sizes (in μm). 59

Figure 2.36 Comparison of burnout evolutions predicted by the previous

model (Case 1), and present model with char gasification reactions

(Case 2) and without char gasification reactions (Case 3). 62

XXVII

Figure 2.37 Burnout for Cases I, II and III: (a) along the centreline and (b)

at the exit. 62

Figure 2.38 Effect of blast temperature on burnouts at the distances of 300

mm and 925 mm from the lance tip, respectively. 63

Figure 2.39 Effect of oxygen enrichment on coal burnout 63

Figure 2.40 Effect of cooling gas type: (a) final burnout and (b) O2

distributions at the cross-plane of 550 mm from the lance tip. 64

Figure 2.41 Geometry of the model: (a), the whole model; (b), porosity

distribution (Zone 0: 1, Zone 1: 0.25, Zone 2: 0.5, Zone 3: 0.4); (c),

blowpipe and raceway; and (d), lance tip. The detailed dimensions are,

(1) for blowpipe, radius: 90 mm, and length: 800 mm; (2) for tuyere,

radius: 75/90 mm, and length: 135 mm; (3) for raceway, depth: 1600

mm, height: 1000 mm (925 + 75), and width: 710 mm; and (4) for

coke bed, depth: 3700 mm, height: 4500 mm, and width: 1000 mm. 66

Figure 2.42 Flow pattern of gas-particle flow: (a), vectors of gas phase in

the raceway; (b), streamlines of gas flow; (c), particle trajectories

coloured by particle mean size; and (d), particle trajectories coloured

by particle travelling time. 67

Figure 2.43 Combustion characteristics of coal along particle trajectories

in the coke bed. 67

Figure 2.44 Schematic of computational domain: (a) side view; (b) top

view. 69

Figure 2.45 Distributions of (a) gas velocity vectors and (b) gas

temperature (K) in the computational domain. 69

Figure 2.46 Coal burnout at the exit of the computational zone. 70

Figure 2.47 Schematic figure of hot model. 71

Figure 2.48 Comparison of calculated raceway shape with observation of

test. (a) All coke operation. (b) PCI operation. 71

Figure 2.49 Characteristics of raceway: (a) calculated raceway shape, and

(b) calculated gas velocity vectors. 72

Figure 2.50 Illustration of combustion phenomena of pulverised coal

(Ishii, 2000). 74

Figure 2.51 Framework of the CFD code and computational procedure of

the gas phase and solid (coal particle) phase (Du, et al., 2007). 75

Figure 2.52 Size effects on physical and temperature profile (Smoot and

Smith, 1985) 76

Figure 2.53 Schematic of the Blowpipe/Tuyere (combustion test section)

assembly of the pilot scale raceway hot model. 81

XXVIII

Figure 2.54 Gas temperature distributions for pulverised coal burning in a

reactor from experimental measurement and numerical predictions

using different devolatilisation models. 81

Figure 2.55 Comparison of the relationships between Y1 and Y2 in the

literature (Shen et al., 2008). 82

Figure 2.56 Rate-controlling regimes for char reactions (Smoot and

Smith, 1985). 85

Figure 2.57 Schematic illustration of raceway structure. 94

Figure 2.58 Representation of the movement of coke through the raceway. 95

Figure 2.59 Comparison of numerical and experimental results (Nogami

et al., 2004) of: (a) raceway shape; and (b) gas composition

distribution along central axial of tuyere. 100

Figure 2.60 Profile of CSC’s No3 blast furnace: (a) main dimensions

(unit: m); (b) calculation domain of a single tuyere. 101

Figure 2.61 Void fraction contours in combustion zone of 3D coke

packed furnace model: (a) top view; (b) side view. 102

Figure 2.62 The simplified calculation domain. Note that αg is the volume

fraction of gas inside the raceway. 103

Figure 2.63 Burnouts as a function of volatile matter of the injectants with

an air cooled lance and O/C = 2.0. Comparison is made with previous

results for PCI coals. 106

Figure 2.64 Differential pressure across the tuyere as a function of the

volatile matter of the injectants. 106

Figure 2.65 Schematic of the reaction system (1) cylinder; (2) carrier gas;

(3) secondary gas; (4) rotameter; (5) hopper; (6) preheater; (7) lance;

(8) DTF; (9) thermocouple; (10) ceramic tube; (11) heater; (12)

sampling probe; (13) cooling water; (14)cyclone; (15) residual solid

particles; (16) induced suction fan; (17) exhausted gas. 108

Figure 2.66 Distributions of burnout versus fuel ratio of raw and torrefied

biomasses as well as a HV coal. 109

Figure 2.67 Geometry and computational domain used in numerical

simulation. 110

Figure 2.68 Temperature profiles at an injection rate of 36 (kg solid fuel) /

(1000 Nm3 feed gas). 110

Figure 3.1 A schematic diagram of internal structure in a blast furnace. 118

Figure 3.2 A comparison of gas temperature distribution among

experimental measurement and two devolatilization models. 122

Figure 3.3 A schematic diagram of blowpipe and tuyere as well as their

sizes. 124

XXIX

Figure 3.4 Isothermal contours in blowpipe and tuyere under the

operations of (a) single lance and (b) double-lance injections. 125

Figure 4.1 A schematic diagram of internal structure in blast furnace. 133

Figure 4.2 Framework of the CFD code and computational procedure of

the gas phase and solid (coal particle) phase. 137

Figure 4.3 Gas temperature distributions for pulverize coal burning in a

reactor from experimental measurement and numerical predictions. 138

Figure 4.4 Trajectories and residence times of coal particles under the

operation of the base case. 140

Figure 4.5 Burning ratios of PC at various injection patterns. 142

Figure 4.6 Combustion situations of pulverized coal in (a) case 1 and (b)

case 3. 142

Figure 4.7 Burning ratios of pulverized coal at various oxygen

concentrations. 143

Figure 4.8 Burning ratios of pulverized coal at various hot blast

temperatures. 144

Figure 4.9 Burning ratios of pulverized coal at various mass flow rates of

carrier gas. 145

Figure 5.1 AE sensor system for measuring burden falling point. 153

Figure 5.2 Burden profile before (a)/ after (b) changing charging pattern. 154

Figure 5.3 Typical pressure distribution of No 3 blast furnace. 156

Figure 5.4 Physical geometry of combustion region. 157

Figure 5.5 Trajectories and residence time of coal particles in the

combustion region. 161

Figure 5.6 Oxygen concentration contour at cross section along

combustion region. 162

Figure 5.7 Pressure distribution along combustion region from lance exit. 162

Figure 6.1 Schematics of (a) physical sizes of computational domain as

well as the arrangements of (b) CSC’s double air-cooled lance and (c)

single and oxy-coal lance (α is the porosity within the raceway). 177

Figure 6.2 Distributions of coal particle trajectory and residence time

under (a) single lance, (b) double air-cooled lance, and (c) oxy-coal

lance injections. 178

Figure 6.3 Distributions of oxygen mole fraction under (a) single lance,

(b) double air-cooled lance, and (c) oxy-coal lance injections 180

Figure 6.4 Distributions of hydrogen mole fraction under (a) single lance,

(b) double air-cooled lance, and (c) oxy-coal lance injections. 181

XXX

Figure 6.5 Distributions of isothermal contours under (a) single lance, (b)

double air-cooled lance, and (c) oxy-coal lance injections. 183

Figure 6.6 Distributions of gas temperature along the centreline of tuyere

under single lance injection and oxy-coal lance injections at different

proportions of enriched oxygen. 183

Figure 6.7 A comparison of coal combustion efficiency among single

lance injection, double air-cooled lance injection, and oxy-coal lance

injections with different proportions of enriched oxygen. 185

Figure 6.8 A comparison of pressure loss among single lance injection,

double air-cooled lance injection, and oxy-coal lance injections with

different proportions of enriched oxygen. 186

Figure 7.1 Schematic diagram of pulverized coal injection and internal

structure of blast furnace around raceway. 192

Figure 7.2 Profiles of R-factor of two different coals at various reaction

temperatures. 196

Figure 7.3 Particle size distributions of coal F before and after

experiencing reactions with (a) larger feed particles and (b) smaller

feed particles 197

Figure 7.4 Peak locations of particle size distributions for coal F before

and after experiencing reactions. 198

Figure 7.5 SEM images of unburned chars of coal F at larger feed

particles (a-c) and smaller feed particles (d-f). 198

Figure 7.6 Particle size distributions of coal L before and after reactions

with (a) larger feed particles and (b) smaller feed particles. 200

Figure 7.7 Peak locations of particle size distributions for coal L before

and after experiencing reactions. 201

Figure 7.8 SEM images of unburned chars of coal L at larger feed

particles (a-c) and smaller feed particles (d-f). 201

Figure 7.9 SEM images of feed coal and unburned char particles shown in

cross-section. 203

Figure 7.10 Profits of soot and tar formations with respect to reaction

temperature. 205

Figure 7.11 Thermogravimetric analyses of the produced unburned chars

and soots at 1400oC. 205

Figure 8.1 Schematic of the reaction system. 213

Figure 8.2 Tests of experimental stability of four different coals under the

base experimental conditions. 215

Figure 8.3 Correlation between burnout and fuel ratio under the standard

combustion conditions. 218

XXXI

Figure 8.4 Distributions of burnout of Coal C, Coal D and Coal I at

various reaction temperatures. 219

Figure 8.5 Distributions of burnout of Coal B, Coal E and Coal I at

various particle sizes. 221

Figure 8.6 Distributions of burnout with respect to blending ratio for Coal

K individually blended with Coal A and Coal C. 222

Figure 9.1 (a) Volatile matter, (b) fixed carbon, and (c) fuel ratio values of

raw and pretreated biomass materials. 233

Figure 9.2 Atomic H/C versus O/C ratio (van Krevelen diagram) of raw

and pretreated biomass materials. 234

Figure 9.3 (a) HHVs and replacement factors of biomass materials based

on (b) Coal A and (c) Coal B. 237

Figure 9.4 (a) Solid yield and (b) energy yield of pretreated biomasses

materials. 238

Figure 9.5 Ignition temperatures of raw and pretreated biomass. 240

Figure 9.6 Burnout versus (a) pretreated temperature and (b) fuel ratio. 241

Figure 10.1 Zonal structures in a drill core. 250

Figure 10.2 Formation of cavity and bird’s nest in a pack bed. 251

Figure 10.3 The measured burden profiles and calculated descending rate

at No3 blast furnace. 253

XXXII

LIST OF TABLES

Table 1.1 Number of blast furnaces worldwide equipped with substitute

fuel injection (Schott, 2013). 6

Table 1.2 Average fuel rates of the blast furnaces in the EU 15 (Peters and

Bodo, 2009). 8

Table 1.3 General features of CSC’s blast furnaces. 10

Table 2.1 PCI combustion models and sub models used. 36

Table 2.2 Operational conditions selected by Burgess et al. (1983). 80

Table 2.3 Three sets of parameters used for predicting PC devolatilisation. 80

Table 2.4 Kinetics of single overall reaction and two competing reaction

models used in the PCI calculation models. 83

Table 2.5 Kinetics of char oxidation employed in the PCI combustion

models. 86

Table 2.6 Comparison of raceway observations in cold and hot models. 95

Table 2.7 Chemical reactions in coke–packed furnace model. 99

Table 2.8 Operating conditions and properties of coke–packed bed. 100

Table 2.9 Typical operating parameters of No3 BF for all coke operation 101

Table 2.10 Key properties of the bulk coal and charcoal samples. 105

Table 2.11 Key properties of the bulk coal and charcoal samples. 105

Table 2.12 Enhancement factor of higher heating value. 108

Table 2.13 Computational conditions for biofuel injection. 110

Table 3.1 Operational conditions selected by Burgess et al. (1983) 121

Table 3.2 Two sets of parameters used for predicting PC devolatilization. 122

Table 3.3 Operating conditions of PCI at CSC. 124

Table 4.1 Operating conditions (base case) of PCI at CSC. 139

Table 5.1 Main features of CSC’s No3 blast furnace. 150

Table 5.2 Typical energy consumption in the late period of the first

campaign. 152

Table 5.3 Reduction of SiO2 in sinter and slag volume. 152

Table 5.4 Variation of wall heat loss. 154

XXXIII

Table 5.5 Comparison between soaking bar tapping and one bit drilling. 155

Table 5.6 Parameters of devolatilisation kinetics.

160

Table 5.7 PCI Operation condition used in the calculation. 160

Table 5.8 Changes of pressure drop and permeability for coal blend

injection. 163

Table 5.9 Hot metal Production in CSC’s No 3 blast furnace. 163

Table 6.1 A list of fuel properties and operating conditions. 176

Table 7.1 Proximate and ultimate analyses of the investigated coals 195

Table 7.2 Summary of reaction physics of the two coals 206

Table 8.1 Proximate analyses (dry basis), fuel ratios and higher heating

values (HHV, dry basis) of the investigated coals. 216

Table 8.2 Maceral analyses of the investigated coals. 218

Table 9.1 Proximate, elemental, fiber, and calorific analyses of two coals

and raw biomass materials. 229

Table 9.2 Ash contents in pretreated biomass materials. 236

1

CHAPTER 1

INTRODUCTION

Pulverised coal injection technology is considered as an effective countermeasure to

reduce operation cost and CO2 emission of blast furnace. The injection rates achieved

by the blast furnaces worldwide are reviewed in this chapter. The development of

auxiliary fuel injection at China Steel Corporation is reported. The objectives and

outline of this work are briefly presented.

2

1.1 Background

Blast furnaces (BF) are currently a crucial and the most commonly employed facility in

ironmaking processes. It is also predicted that the blast furnace will remain the

successful process for hot metal production for the coming 30 years due to its

advantages in high productivity and heat utilisation (Geerds et al., 2011).

From the thermal point of view, the blast furnace works as a counter-current reactor in

which solids descend and gases ascend. Practically, iron oxides (i.e. sinter, pellet and

lump ore) and coke (reductant) are charged through the rotating chute with discrete

dumps to form alternating layers in the furnace. When descending, the charged iron

oxides are heated and reduced by ascending gases from tuyeres installed in the low zone

of blast furnace. Softening and melting of reduced iron and gangue materials begin

when the temperatures are high enough. Notably, the coke remains in solid state and

descents to the furnace hearth constructing a porous coke bed named deadman. The

generated liquid iron (hot metal) and slag trickle through the deadman to the hearth

bottom of the blast furnace. The hot metal and slag are cast into a main through when

the taphole is opened. For the generation of thermal energy and reducing gases required

for hot metal production, blast air heated to the temperature of 1100-1250°C is

introduced at a velocity around 180 m/s into the furnace through the tuyeres in the lower

zone of blast furnace. Consequently, a cavity call a raceway is formed in front of the

tuyere exit. It is found that the raceways of adjoining tuyeres are not connected with

each other (Nakamura et al., 1978).

According to the reaction characteristics, the entire blast furnace can be generally

divided into five individual zones (Figure 1.1) from the top downwards:

(1) lumpy zone for pre-reduction and reduction of iron oxides;

(2) cohesive zone for reduced iron and gangue materials fusing and melting;

(3) dripping zone for gas and liquid refining;

(4) raceway for carbon gasification and heat generation; and

(5) deadman zone for irrigated bed slag/metal refining.

3

Clearly, the five reaction zones are effectively integrated into a single shell in the blast

furnace ironmaking process. As a result, the blast furnace is still competitive for hot

metal production even though many new ironmaking technologies have been studied

and developed.

Figure 1.1 Schematic diagram of internal structure with five district zones in a

blast furnace.

Oxides: sinter, pellet and lump ore

Reductant: coke

Top gas: CO, CO2,

H2, N2, H2O

Blowpipe and

tuyere

Hot blast air

and PCI

Taphole

Lumpy zone with coke

and ferrous layers

Cohesive zone

Raceway

Dripping zone

Deadman

Hot metal and slag

4

An idealised blast furnace reduction process can be found in Figure 1.2 (Biswas, 1981).

Generally, the reduction of iron oxides can be catalogued into indirect reduction and

direct reduction, which occur at the upper and lower zones of a BF, respectively. In

indirect reduction process, iron oxides, including Fe2O3, Fe3O4 and FeO, are

exothermically reduced by reducing gases (CO and H2) into other iron oxides (Fe3O4

and FeO), resulting in CO2 and H2O as products. These normally occur at temperatures

below 850-900oC. In direct reduction at the low zone of blast furnace, iron and CO are

produced by carbon (from coke) reacting endothermically with iron oxides above

900oC. It is noted that hydrogen regeneration reaction (C + H2O = CO + H2) is less

endothermic and proceeds faster than the carbon monoxide regeneration loss reaction (C

+ CO2 = 2CO). For a blast furnace with all coke operation, partial replacement of coke

by hydrogen containing fuels may be an effective way for the improvement of blast

furnace performance. .

Figure 1.2 Reduction reactions in blast furnace (Biswas, 1981).

5

1.2 Auxiliary fuel injection into blast furnace

In the blast furnace ironmaking process, coke plays a particularly important role for

stable operation, because it provides mechanical support for the descending materials

and ensures permeability for the ascending gas within the cohesive zone, and for the hot

metal within the deadman zone. Besides, it reacts with hot blast air in the raceway to

generate energy and reducing gases for the hot metal production. Because the

metallurgical coals for cokemaking are expensive, the reduction of coke consumption is

always desirable for the blast furnace operation. With the developments in high

temperature stoves, big oxygen plants (providing large amount inexpensive oxygen),

measurements and operation models, the injection of auxiliary fuels, such as natural gas,

oil and pulverised coal, into the blast furnace as a substitute of coke in the raceway has

become an industrial reality. Apart from the economic benefit, the advantages below are

generally expected from the auxiliary fuel injection:

(1) to increase the productivity of the blast furnace, resulting from more oxygen being

added in the blast for maintaining the flame temperature of raceway;

(2) to smooth down the burden movement to assist in maintaining furnace stability; and

(3) to prolong the life span of coke battery since less coke has to be produced.

Due to ease of operation, natural gas followed by oil were popular injectants. Although

the adoption of the pulverised coal injection (PCI) technology was discouraged initially

primarily because there was no strong economic incentive before second oil crisis in

1978, some significant achievements on PCI operation have been reached. For instance,

Armco installed a commercialised PCI system at the Ashland, Kentucky plant in 1963

(Nolde et al., 1996), and the Shoudu Iron and Steel Corporation in China first achieved

a successful PCI practice in 1964 (Wei and Qi, 1983). After the crises, many companies

stopped injecting oil into blast furnaces and turned to PCI operation, because the coal

preparation and pneumatic transportation have become proven technologies by that

time. As shown in Table 1.1 (Schott, 2012), nearly half of all blast furnaces in the world

6

(47.7%) use PCI; while only 11.9% inject gas or oil, and 35.9 % remain all coke

operation because the blast furnaces are rather small or old, making cost effectiveness of

an application of substitute fuel injection questionable.

Table 1.1 Number of blast furnaces worldwide equipped with substitute fuel injection

(Schott, 2012).

PC

Oil Gas

All Coke Plastics

Oil with gas gas with PC

Africa 2 1 4 2

America 36 3 7 11 17 26

Asia 278 4 1 220

Australia 2 1 2

Europe 74 15 4 61 3 47 2

Total, % 47.7 4.1 11.9 35.9 0.2

1.3 Achievement of high coal injection rate

In the beginning of the adoption of PCI operation, it was predicted that the PCI rate

might be limited to 60 kilograms per tonne hot metal (Gudenaua and Kiesler, 1991) or

to about 15% of the fuel rate of blast furnace (Poos and Ponghis, 1990). Efforts towards

enhancing the PCI operation have been made by the blast furnaces worldwide for higher

coal injection rates.

As can be seen in Figure 1.3 (Ariyama et al., 2007), pulverised coal injection into blast

furnaces in Japan began in 1983, and the average pulverized coal injection rate achieved

by the Japanese blast furnaces has been gradually increased to 140 kg/tHM in 1999.

Since then, the coal injection rates were remained in a range of 120 to 135 kg/tHM.

Notably, a remarkable injection rate of 230 kg/tHM and monthly average of 218

kg/tHM have been attained by NKK’s Fukuyama No4 BF (Maki et al., 1996). In Korea,

POSCO’s Gwangyang No4 BF (inner volume: 5500 m

3) has reached a PCI rate of 200

7

kg/tHM with a low coke rate of 290 kg/tHM. This achievement has been set as the

operation target for the big blast furnaces (inner volume >4000 m3) in China (Zhang et

al., 2013). Besides, it should be noted that Baosteel No4 BF only spent one month after

blow in operation to reach a high PCI level of over 200 kg/tHM (Li et al., 2007).

In the 1990s, there were still 45 integrated works operated in whole Europe. Since then

blast furnace works have been shut down and partly been replaced by electric arc

furnaces (Peters and Lüngen, 2009). During this period, three European blast furnaces,

Thyssen Schwelgern No1 BF, Sidmar B BF and Hoogovens N

o7 BF, recorded high PCI

rates of 178, 197 and 199 kg/tHM respectively with low coke rate (Buss et al., 2000;

Peters et al., 1991). In 2008, the average PCI rate of the blast furnaces in the EU 15 is

123.9 kg/tHM, while extraordinary operation modes with PCI rate, as well as low coke

consumption, have been reached at some blast furnaces as indicated in Table 1.2 (Peters

and Lüngen, 2009). It is found that highest coal rate was realised at Tata Corus

(formerly Hoogoven) No6 with 235.1 kg/tHM as yearly average. The lowest coke rate

with 289.9 kg/tHM with a resulting total fuel rate of 516.9 kg/tHM was achieved at this

furnace. As shown in the Table, the lowest total reductant rate was achieved at Ruukki

blast furnace No1 (458.5 kg/t HM) in oil injecting operation mode.

Figure 1.3 Change in blast furnace operating conditions in Japan (Ariyama et al,

2007).

8

1.4 Utilisation of auxiliary fuels at CSC

China Steel Corporation (CSC) group is the only integrated and the largest steel

producer in Taiwan. It has 4 blast furnaces (No1 to 4) on its major production side in

Kaohsiung, and 2 blast furnaces (No5 and 6) in Taichung, with a designed annual hot

metal output of approximately 15 million metric tonnes. The general features of CSC’s

blast furnaces are listed in Table 1.3. In general, the hot metal production consumes the

most energy of an integrated steelwork (Babich et al., 2002). At CSC, more than 55% of

the entire energy is consumed by the blast furnace ironmaking process. Therefore many

technologies, such as auxiliary fuel injection and top gas recovery turbine (TRT), have

been applied for the reduction of energy consumption at the blast furnace process of

CSC.

Before the second oil crisis, oil was injected through tuyeres into CSC’s blast furnaces

as the substitute of coke. The injection rates were in a range of 50 to 60 kg/tHM.

Table 1.2 Average fuel rates of the blast furnaces in the EU 15 (Peters and Lüngen,

2009).

9

Concerned by the uncertain oil supply and substantial increase in prices after the crisis,

CSC introduced PCI technology to its blast furnaces in 1987 (Du et al., 2001).

To make the injected coals ignited earlier in the tuyere for higher combustibility, high

volatile coals (VM＞35%) were solely injected through single lances into the blast

furnaces since CSC commenced its PCI operation. For the reduction of the unburnt char

generated in the raceway, double air-cooled coaxial lance was developed and applied at

CSC in 2001 (Du et al., 2001; Yeh et al., 2002; Du and Chen, 2006). To improve the

permeability in the lower zone of blast furnace, as well as to diversify the coal types for

the blast furnace injection, low volatile coals were blended into high ones in the PCI

operation of CSC in 2003 (Du et al., 2004). At present, the proximate volatile matter of

coal blend for injection operation at CSC’s blast furnaces is generally kept in the range

of 19 to 21%.

Apart from its economic benefit from the replacement of expensive coking coals by

cheaper thermal coals, PCI operation is also thought to be environmental friendly at

CSC since the injected coals are consumed directly, without going through the

coke-making plant. Therefore, high PCI rate is one of operation targets of CSC’s blast

furnaces. It should be emphasised that the blast furnace ironmaking is a complicated

process, so the replacement of coke with pulverised coal is not as simple as just

increasing the injection rates or improving the combustibility of coal injected, especially

when one considers the impact of coal combustion to the stability of raceway, fuel rates

and permeability of furnace. Figure 1.4 shows the impact of high coal injection rate on

blast furnace operation (Ishii, 2000). It suggests in high PCI rate, knowledge of the

details in the combustion region becomes more critical for high PCI rate. A brief

understanding on the coal combustion behaviours in the regions of blowpipe, tuyere and

raceway can lead to a more effective and safe operation with high PCI rate.

10

Table 1.3 General features of CSC’s blast furnaces.

BF No1 N

o2 N

o3 N

o4 N

o5 N

o6

Inner volume, m3 2624 3274 3606 3422 3274 3274

Hearth diameter, m 10.2 12 12.5 12.5 12 12

Cooling Stave Stave Stave Plate Stave Stave

Number of tapholes 2 2 4 4 3 3

Number of tuyeres 30 30 32 32 32 32

Blow-in/ campaign 2010

(4th

)

2006

(3rd

)

2009

(2nd

)

2014

(2nd

)

2010

(1st)

2013

(1st)

First campaign 1977 1982 1987 1996 2010 2013

Figure 1.4 Changes and problems of blast operation with high injection rate.

11

1.5 CO2 emission at CSC

The greenhouse gas of most relevance to the world steel industry is carbon dioxide. As

a matter of fact, CO2 liberated from blast furnaces approximately account for 3.5–5% of

total anthropogenic CO2 emissions (Wang et al., 2009). On average, 2.1 tonnes of CO2

are emitted for every tonne of crude steel produced at CSC. The utilisation of

alternatives to coal in PCI can abate the consumption of fossil fuels and, in the case of

biofuel or charcoal, mitigate CO2 emissions. It was reported by Babich et al. (2010) that

the injection of charcoal fines has been successfully practiced in some small charcoal

blast furnaces in Brazil with injection rates of 100 to 150 kg/tHM.

For the reduction of CO2 emission, a pilot plant for producing biofuel from local

agricultural wastes has been established in Malaysia by CSC. To successfully use

biofuel as partial replacement for pulverised coals through injection, it is required to

examine the fuel properties of biomass pretreated by torrefaction and carbonisation and

compare to those of high-volatile and low-volatile coals.

1.6 Objectives of the work

To provide brief insights into the coal burning characteristics in the lower zone of blast

furnace, one of objectives of this study is focused on the development of a 3-D CFD

based model to predict the coal combustion phenomena occurring in the regions of

blowpipe, tuyere and raceway, identification of the major operation parameters affecting

the coal combustion and the permeability of raceway, and application of the calculation

results to the PCI operation of CSC’s blast furnaces.

Although the combustion conditions within a drop tube furnace are less intense than

those in the region of blowpipe-tuyere-raceway, the drop tube furnace may provide a

better view at the combustion properties of coal without influences from gas flow

pattern, coal injection rate and char sampling. A drop tube furnace has been established

in this research to evaluate the combustion performance of PCI coals, coal blend and

12

biofuels in an environment with high heating rate (>104 K/s). Besides, the coal

combustion characteristics, such as volatile release, swelling and agglomeration, are

also examined using the furnace.

1.7 Thesis outline

The thesis begins with brief introductions to blast furnace ironmaking process, PCI rates

achieved in blast furnaces worldwide and the general features of PCI operation in the

blast furnaces of CSC. A review of the relevant literature is made in Chapter 2. The

development and application of the 3D CFD based coal combustion model are

presented in Chapters 3 to 6. In this work, the development of the model can be divided

into 4 phases: (1) validation of the coal combustion model by comparing its predictions

with experimental data; (2) investigation into the influence of operation conditions to

the coal burnout in the region of blowpipe and tuyere; (3) performance evaluation of

coal blend injection in terms of pressure loss due to coal combustion within a simplified

raceway space; and (4) examination of burning characteristics of pulverised coal with

different ways oxygen enrichment in the regions of blowpipe, tuyere and raceway,

which is a porous space featured by Eulerian-Eulerian multi-fluid approach.

The publications presented in this thesis are listed below:

Chapter 3

Du, S. W. and Chen, W. H. (2006), Numerical prediction and practical improvement of

pulverized coal combustion in blast furnace, International Communications in Heat and

Mass Transfer, vol. 33, p. 327-334.

Chapter 4

Du, S. W., Chen, W. H. and Lucas, A. J. (2007), Performances of pulverized coal

injection in blowpipe and tuyere at various operational conditions, Energy Conversion

and Management, vol. 48, p. 2969-78.

13

Chapter 5

Du, S. W., Yeh, C. M., Yang, M. K. and Ho C. K. (2004), Practice of high productivity

at No.3 blast furnace of China Steel Corporation, Proceedings of AISTech Conference,

Tennessee, USA, p. 195-204.

Chapter 6

Du, S. W., Yeh, C. P., Chen, W. H., Tsai, C. H. and Lucas, J. A. (2015), Burning



By using the drop tube furnace established at CSC, the research is designed to (1) study

the volatile release and particle formation within the coal plume, in which the oxygen is

insufficient; (2) establish a method to evaluate the coal combustion efficiency for the

selection of PCI coal at CSC; and (3) examine the combustion of bio fuels pretreated by

torrefaction and carbonisation for their utilisation in PCI operation. The experimental

results and findings are presented in Chapters 7 to 9 by the publications below:

Chapter 7

Chen, W. H., Du, S. W. and Yang, T. H. (2007), Volatile release and particle formation

characteristics of injected pulverized coal in blast furnace, 2007, Energy Conversion


Chapter 8

Du, S. W., Chen, W. H. and Lucas, A. J. (2010), Pulverized coal burnout in blast


Chapter 9

Du, S. W., Chen, W. H. and Lucas, A. J. (2014), Pretreatment of biomass by



14

Finally, Chapter 10 summaries the main conclusions from this work, followed by

recommendations for future work on the subject.

15

CHAPTER 2

LITERATURE REVIEW

In this chapter, a comprehensive review of the previous experimental and numerical

studies on the pulverised coal combustion in the blast furnace is presented. An overview

of factors affecting coal combustion is made. It also discusses the application of sub

models in the integrated coal combustion models. The CSC’s internal research works,

including the validation of devolatilisation models and the prediction of raceway shape

are briefly described in this chapter. The application of biofuel in the PCI operation is

outlined. Finally, the methodology of the work is presented.

16

2.1 Coal combustion experiments under simulating PCI operation conditions

Combustion process of injected coal in the blowpipe-tuyere-raceway region can be

characterised by (Steiler et al., 1996):

(1) high heating rate (>104 K/s) of injected coal due to very high blast temperature

(1050-1250oC);

(2) short residence time for injected coal in the combustion area (<20 ms) caused by

high blast velocity (>160m/s); and

(3) insufficient mixing of injected coal particles with hot blast gas (Du et al., 2007;

Goto et al., 2002).

It is predicted that as PCI rate is increased, more amount of unburnt char is generated

and accumulated in the blast furnace due to incomplete conversion of injected coal,

leading to dirtying of the deadman and finally decrease in the furnace productivity and

increase in the fuel rate. Therefore, considerable attention has been focussed on the

improvement of coal combustion efficiency for the PCI operation.

For studying the factors influencing coal combustion during PCI operation, many coal

combustion experiments have been carried out by using combustion rigs with empty

combustion chambers (Guo et al., 2011; Mathieson et al., 2005; Du et al., 2001; Picard

et al., 2000; Kim et al., 1996; Babich et al., 1996; Steeghs et al., 1996; Ueno et al.,

1993; Suzuki et al., 1984; Burgess et al., 1983; Bortz and Flament, 1983; Narita et al.,

1982;), or, in some cases, with coke bed reactors (Nogami et al., 2004; Ariyama et al.,

1994; Sato et al., 1994; Yamagata et al., 1992; Yamaguchi et al., 1992). Practically, the

operation conditions of the rigs, such as blast air (temperature, oxygen level and

velocity), lance configuration and coal flow rate, can be set as closely as possible to

those of actual blast furnaces. Notably, some operation difficulties, including pulverised

coal preparation, char sampling, temperature measurement and keeping a constant coal

injection rate, may be encountered with the combustion rigs (Du et al., 2010). Despite

the inherent difference between the realistic combustion environment of the raceway

17

Figure 2.1 Schematic of combustion rig attached with an empty combustion chamber.

and the drop tube furnace, the drop tube furnace is considered as an effective device

when one attempts to evaluate the combustion performance of PCI coals in an

environment with high heating rate (Du et al., 2010). The sections below will

summarise the findings of the coal combustion experiments.

2.1.1 Experiments using empty combustion rig

Figure 2.1 shows a combustion rig connected with an empty combustion chamber

(simulated as the raceway) developed by BlueScope Steel Research (Mathieson et al.,

2005). Due to operational ease in comparison with the coke bed reactors, empty

combustion rigs have been preferred for practising pulverised coal combustion

experiments. A summary of the effects of operation parameters and coal properties to

the coal combustion obtained from those experimental works is presented as followed:

2.1.1.1 Effect of volatile matter content on combustion

(1) The burnout of injected coal was strongly dependent on the coal rank. High volatile

coals were burnt preferentially to low volatile coals (Vamvuka et al., 1996; Steeghs

et al., 1996; Ueno et al., 1993; Malgarini, 1991; Suzuki et al., 1984; Wakimoto et

al., 1983; Burgess et al., 1983; Bortz and Flament, 1983). However, high volatile

https://getinfo.de/app/subject-search?action=search&author=%22Malgarini%2c+G.%22&form=advanced

18

matter content may not be sufficient to characterise coal combustion efficiency in

the blast furnace operation, because the formation of tar and soot generated from the

released volatile (Chen et al., 2008; Bortz and Flament, 1983) and the pressure loss

due to the combustion within the raceway (Du et al., 2004; Du et al., 2015) should

be considered as the operation challenges with high coal injection or high blast air

flow rate.

(2) Some low volatile coals showed higher combustion efficiency than that expected

from their volatile levels (Mathieson et al., 2005). It was explained by the formation

of fragments from chars, leading to the increase of surface area for combustion (Li

et al., 2014).

(3) Under nitrogen atmosphere, the final volatile yield of injected coal was higher than

the proximate analysis (Q factor >1), and the yield increased with increasing the

blast temperature as indicated in Figure2.2. Moreover, the coal particle started to

swell significantly when the weight loss of coal was higher than 20%. With 51%

weight loss, the diameter of coal could be 1.3 times higher than the initial size

(Ueno et al., 1992). The Q factor is the ratio of the measured volatile content in high

heating rates to the proximate volatile matter.

Q factor: ratio of

final volatile yield

to that by

proximate analysis

Figure 2.2 Influence of blast temperature on Q-factor.

19

2.1.1.2 Effect of coal size on combustion and granular coal injection

(1) The burnout of injected coal increased with decreasing the coal particle size

(Wakimoto et al., 1983; Bortz et al., 1983; Narita et al., 1982). Alternatively, it was

found in the research work by Du et al. (2010) that the coal combustion could not

be improved further as a result of the agglomeration effect when the feed particle

size was smaller than 75 μm.

(2) To get higher combustion efficiency, many blast furnaces inject pulverised coal

(80% < 75 m) (Hutny et al., 1991). On the other hand, British Steel injects

granular coal (100% < 5mm, and 95% < 2mm) into the blast furnace (Gathergood

and Jukes, 1996). A sampling probe has been developed by Guo et al. (2012) to

take granular coal particles from the raceway of an operating blast furnace (Lai-Wu

Steel in China). Owing to rapid release of moisture and volatile matters, significant

cracking of injected granular coal was observed through microscope structure

analysis. Consequently, coal combustion was boosted with the increase of reaction

surface area by coal cracking. Practically, No1 blast furnace of Lai-Wu Steel

adopted granular coal injection (GCI) in practice in 2005, and an average injection

rate of 168 kg/tHM was achieved in 2012 (Zhu and Xu, 2014). However, injecting

granular coal into blast furnace has not yet been widely practised by the

international steel producers.

2.1.1.3 Effect of hot blast conditions on combustion

(1) It was reported (Guo et al., 2011; Du et al., 2001; Ueno et al., 1992) the increase in

oxygen content and temperature of blast were the effective countermeasures for

improving the coal combustion efficiency.

(2) Increasing blast pressure up to 3 kg/cm2 promoted combustion efficiency markedly

(Wakimoto et al., 1983).

20

2.1.1.4 Effect of injection rate on combustion

As indicated in Figure2.3 (Vamvuka et al., 1996), the coal burnout was decreased with

the increase of coal injection rate (Vamvuka et al., 1996; McCarthy et al., 1986; Suzuki

et al., 1984). This implies the increase of unburnt char generated in the raceway is

unavoidable when one attempts to promote the PCI rate to a certain amount.

2.1.1.5 Effect of lance configurations on combustion

The double lance was found to have the highest combustion efficiency followed by

oxy-coal lance and single one (Du et al., 2001).

2.1.1.6 Effect of co-injection on combustion

The coal combustion was improved when MgO was added in amounts of 3-4% to the

PCI coal (Guo et al., 2011). On the other hand, it was reported by Vamvuka et al. (1996)

that an increase of additive amount of shredded light fractions (from scrap automobiles)

resulted in a decrease in the coal burnout. Besides, by injecting coke oven gas to an

outer peripheral area of the pulverised coal plume, the combustion efficiency of coal

was improved significantly (Suzuki et al., 1990).

Figure 2.3 Degree of combustion of the coals as a function of injection rate.

21

2.1.1.7 Effect of coal blend operation

Adding a high volatile coal to a low volatile coal, a linear relationship of combustion

efficiency was reported by Yu (1999). In some experiments of POSCO, the combustion

efficiency of coal blend is even higher than that of high volatile coal (Kim et al., 1996).

Despite the increase in the generation of unburnt char, the pressure loss caused by the

coal combustion in the lower zone of blast furnace could be abated when a low volatile

coal was blended into a high one (Du et al., 2004).

2.1.1.8 Ignition and combustion of volatile matters

(1) The ignition of injected coal within a tuyere was observed by using an endoscope

(Picard et al., 2000). It was found inflammation never occurred when an 11% VM

content coal was injected; at 23% VM, a flame was not present at every trial; at 30%

VM, a flame attached to the lance was always observed.

(2) During the coal combustion, there were two zones observed in terms of temperature

in the flame (Wakimoto et al., 1983). The first zone (higher heating rate) close to the

tuyere was for volatile combustion, and second zone was for char combustion. It is

similar with the calculated temperature profile in a combustion chamber (Du and

Chen, 2006).

2.1.1.9 Ash fusion temperature

From the amount of slag/ash deposits in blowpipe, it suggested that high ash fusibility

(low ash fusion temperature) was a desirable characteristic of coals for PCI (Scaife,

1983; Ostrowski, 1983).

2.1.2 Experiments by coke-packed bed rigs and actual blast furnace

For simulating the raceway of blast furnace, the coke-packed rigs have been operated

under blast furnace operation conditions. Figure 2.4 shows a coke-packed combustion

rig developed by NKK (Ariyama et al., 1994).

22

The main findings obtained from the rigs or actual furnaces are summarised below.

2.1.2.1 Effect of lance configuration on combustion

(1) From the images taken by a high speed camera in the blowpipe, the pulverised coal

combustion was not uniform across the blowpipe with operating a single lance. The

injected coal was likely to move as a group of particles, and formed the luminous

flame around it. This suggested that the hydrocarbons from coal pyrolysis were

decomposed to generate soot (Sato et al., 1994).

(2) Compared with single lance operation, double lance operation revealed a relatively

smooth change of brightness measured by a high speed camera, and the average

brightness level was higher than that given by the single lance operation.

Consequently, higher combustion efficiency could be reached by using the eccentric

double lance (Sato et al., 1998) as shown in Figure 2.5. In practice, NKK has been

using the eccentric double-lance arrangement at its Fukuyama No4 BF, and an

injection rate of 218 kg/tHM was announced by the furnace (Maki et al., 1996).

Figure 2.4 Schematic of combustion rig attached with a coke bed.

23

(3) An oxy-coal lance, with a coal flow in the inner pipe and an oxygen flow in the

annulus of the coaxial lance, was tested in a single tuyere at an actual blast furnace

(Wikström et al., 1996). The extension of the mixing and the combustion zones

were measured by a thermovision camera. The combustion of coal was found to be

improved markedly by using the coaxial lance as compared with single lance

operation.

2.1.2.2 Effect of hot blast conditions on combustion

The increase of blast temperature and oxygen content in hot blast gas were effective

countermeasures for improving the coal burnout in the raceway, and an optimum

oxygen enrichment of 4% for hot blast gas was suggested by Yamaguchi et al. (1992).

2.1.2.3 Movement of small coke within raceway

Upon inspection of the raceway coke shell (bird nests) taken from the rig, it was

concluded by Nogami et al. (2004) that the movement of small coke particles contracted

by gasification almost determined the formation of raceway shell. Small particles

delivered by gas flow to the raceway boundary did not return inside raceway space and

Figure 2.5 Effect of lance arrangement on pulverised coal flow and combustion

efficiency by hot model (η (%): combustion at 300 and 600 mm from lance tip).

24

enter into the packed beds. Then, most of them lost the momentum by collision with

coke in packed beds and accumulated on the surface domain of the packed beds with

decreasing their mass by solution-loss reaction.

2.1.2.4 Effect of volatile content on combustion

Measurements by a sideways tuyere probe in the raceway of an actual furnace show the

combustibility of injected coal was improved with the increase in volatile matter

content (Takeda et al., 1990). This is similar to the experimental results given by the

empty combustion rigs.

From the combustion experiments with use of a coke-packed hot model (Yamagata et

al., 1992), the combustion efficiency of coal decreased with the decrease of volatile

matter in tuyere level. On the other hand, under 200kg/tHM injection rate, the coal

burnouts measured at the height of 700mm above tuyere were at high level of over 95%

for all tested coals with different volatile contents (18.9-44.9%) as indicated in Figure

2.6. The reaction of char with CO2 (carbon solution loss reaction: C + CO2 = 2CO) was

thought to be predominant after the char particles leave the raceway.

Figure 2.6 Change of coal burnout at the height of 700mm above tuyere level with

change of volatile content in coals.

25

2.1.2.5 Raceway control during operation at CSC

As indicated in Figure 2.7a taken by the tuyere monitoring system, the dispersion of

coal plume to hot blast gas within the raceway was poor with the single lance (Du et al.,

2001). Obviously, the contact area between coal particles and hot blast gas could be

increased when the double lance is employed (Figure 2.7b). To prevent the coal plume

from being forced backwards into the bustle pipe, leading to possibility of explosion,

coal injection should be stopped at the tuyere blocked by un-melted scab as shown in

Figure 2.8 (Ho and Du, 2008).

Figure 2.7 Image of coal flow patterns at CSC. (a) single lance; (b) double lance

injection.

(a)

(b)

26

2.1.3 Coal combustion experiments by drop tube furnace

As a matter of fact, the PCI coal combustion experiments by the combustion rig

with/without coke bed are manpower and cost consuming process. Alternatively, a drop

tube furnace (Figure 2.8) can be used to simulate the situation of fuel particles suddenly

exposed to a high-temperature environment with heating rates ranged from 104 to10

5

K/sec.

The experimental results given by the furnace are indicated below:

(1) Through optical pyrometer observation (Figure 2.9), two-step combustion of single

coal particle was found; a first and very fast step of volatile combustion followed by

a much slower step of char combustion (de Lassat de Pressigny et al., 1990).

Devolatilisation experiments performed with coals with different volatile levels at

1200oC in a nitrogen atmosphere showed the ultimate volatile released was close to

1.7 times higher than that given by ASTM standard method (de Lassat de Pressigny

et al., 1990).

(2) The performance of a combustion rig and a drop tube furnace was compared by Li et

al (2014). It was found the measured burnouts from both the drop tube furnace and

the rig (Mathieson et al., 2005) produced approximately linear trends as a function

Figure 2.8 Image of blockage of tuyere by un-melted scab.

Lance

Un-melted

scab

Lance Coal

plumes

27

of coal volatile matter. In addition, the coal burnout in drop tube furnace was more

sensitive to the coal volatile matter than that in the rig, resulting from differences in

operation, temperature, residence time, and heating rate between the drop tube

furnace and the PCI rig.

Figure 2.8 Schematic representation of a DTF.

Figure 2.9 Mechanism of pulverised coal combustion.

28

2.1.4 Summary of factors affecting coal combustion from experiments

Based on the experimental results, some key findings for the coal combustion in blast

furnace can be drawn:

(1) The factors, which inference the coal burnout significantly, can be summarised into

three categories as shown in Figure 2.10:

(2) From the temperature profile, the combustion of injected coal in the raceway can be

characterised into two stages: (1) a rapid rise in gas temperature caused by the

combustion of volatile matters emitted from the injected coal, and (2) a slow process

of char reactions.

(3) Due to very short residence time for coal particles in the raceway, complete

combustion of the injected coal in the raceway is quite unlikely in the blast furnace

operation. Notably, because char gasification proceeds much faster than that of coke

(Iwanaga, 1991; de Lassat de Pressigny et al., 1990), the permeability of gas and

liquid in the furnace may not be significantly affected by the generation of unburnt

Blast conditions

1. Temperature

2. Oxygen enrichment

3. Pressure

Coal burnout

Coal properties

1. Volatile content

2. Particle size

3. Coal blend

4. Cracking

Injecting facilities

1. Configuration of

lance (mixing of coal

with blast or oxygen)

2. Secondary fuel

injection

3. Injecting additives

Figure 2.10 Critical operational factors on coal burnout.

29

char as long as the consumption rate of unburnt char exceeds its accumulation rate in

the furnace.

(4) With rapid heating environment at blast furnace tuyere, the reactions in the raceway

include fast volatile release and combustion followed by slow char reactions.

Moreover, the final (or ultimate) volatile yield of injected coal in the raceway is

higher than that by proximate analysis. It suggests the conversion of injected coal in

the raceway may come largely from the coal devolatilisation. Furthermore, the

raceway combustion behaviours are more likely to be dominated by the release and

combustion of volatile matter rather than by the gasification of char and coke.

2.2 Modelling of pulverised coal combustion in blast furnace

As indicated above, experimental combustion rigs and drop tube furnaces have been

established and applied for the investigation of pulverised coal combustion behaviour in

the blast furnace. It should be noted the experimental studies mainly focused on the

evaluation of operation parameters to coal burnout. Alternatively, numerical models,

including Computational Fluid Dynamics (CFD), are able to analyse the physical fluid

system of raceway more cost effectively and rapidly than the experimental procedures.

From the perspective of mathematical modelling, the pulverised coal combustion

process, involving turbulent multiphase flow coupled with momentum, mass and heat

transfer, and various homogenous and heterogeneous chemical reactions, can be studied

comprehensively.

2.2.1 Development of one-dimensional model

In the 1980s, one-dimensional (1-D) mathematical models were developed to study the

combustion behaviours in raceway for all coke operation, or for PCI operation.

Common features of the 1-D models include: (1) the injected coal particles were

dispersed uniformly with the hot blast gas; and (2) the raceway was assumed as

30

cylindrical and non-spreading jet, and (3) gas leaved from the main flow zone through

the roof of the cylindrical path. A brief discussion on the key models is presented

below.

2.2.1.1 Model of Kuwabara et al.

Based on the equations of motion and continuity of gas and coke, a 1-D kinetic model

was developed to describe the behaviour of coke combustion in front of a tuyere area

without injecting auxiliary fuel (Kuwabara et al., 1981). According to the observation

using an endoscope by Greuel et al (1974), the raceway was assumed to comprise a

cylindrical jetting space which had the same diameter as the tuyere exit as shown in

Figure 2.11. With the increase of oxygen level from 21% to 23%, early ignition of coke

in the raceway was found, resulting in shifting the peak levels of gas temperature and

concentration of CO2 towards the tuyere exit.

2.2.1.2 Model of Burgess et al.

A plug flow model was developed to simulate the fuel-lean combustion of pulverised

coal in the blowpipe (Burgess et al, 1983). The coal particle velocity was calculated by

Tuyere

Figure 2.11 Schematic view of combustion in front of a tuyere.

31

the equation of motion of the particle. From the sensitivity tests of the model, blast

temperature and the volatile content of injected coal, as compared with blast velocity,

coal size, and oxygen content in blast, showed a more effect on coal burnout. Due to the

assumption of instantaneous mixing of coal particles with hot blast gas, the calculated

gas temperatures in the blowpipe were significantly higher than the measured ones.

2.2.1.3 Model of He et al.

Competitive combustion of coke and pulverised coal in the raceway was considered in

the 1-D model developed by He et al. (1986). As indicated in Figure 2.12, the raceway

was modelled as a cylindrical jet surrounded by coke bed. Heterogeneous reactions of

coke and pulverised coal with O2, H2O and CO2 were taken into account. The model

results showed that (1) an increase in blast temperature resulted in increases in the

amount of released volatile and gas temperature; (2) higher burnout could be given by

the coal with higher volatile content; (3) when pulverised coal was injected, the peak

temperature of the raceway was slightly lower than that of all coke operation; and (4)

the peak temperature of gas in the raceway moved closer to the tuyere exit when

operating higher PCI rate.

Figure 2.12 Schematic view of combustion with the injection of pulverised coal.

32

2.2.1.4 Model of Jamaluddin et al.

In the model developed by Jamaluddin et al. (1986), the flow in the region of

blowpipe-tuyere was divided into three regions, the potential core (coal flow), the main

flow (blast), and mixing layer in which the injected coal particles and the blast were

mixed uniformly and the heating and pyrolysis of the injected coal particles took place.

The area of mixing layer grew in cross-section. The particle velocity was calculated

from the equation of motion of the particle. The char was treated as pure carbon and its

combustion proceeded in two steps, the first one char was oxidised by O2, CO2, H2O, O

and OH to produce CO and H2, and second constituting the oxidation of the primary

combustion products to CO2 and H2O. The raceway combustion was treated as a 1-D

process. The model assumed that falling cokes with recirculated combustion gas entered

the raceway from the roof of raceway as shown in Figure 2.13. Notably, the model is

the first published model to consider raceway with recirculated gas.

The calculated results showed the predicted gas temperature and concentration were

sensitive to (1) combustion of gases recirculated from the upper part of the raceway

Figure 2.13 Gas flow, entrainment and disentrainment patterns inside the raceway.

33

cavity, (2) combustion or gasification of the coke particles falling into the raceway, and

(3) reactions at the coke bed forming the raceway boundary.

2.2.1.5 Model of Yamagata et al.

A calculation model was developed by Yamagata et al. (1994) to examine the effects of

the rate of PCI up to 300 kg/tHM, coal volatile content and particle size on combustion

in the region of blowpipe-raceway. Both gas and solid phase were modelled as one

dimensional flow with complete mixing in the blowpipe and raceway cavity. This model

consisted of the equations of continuity and motion for gas and coal particles. Collision

of fine particles with packed bed was also considered. The heterogeneous reactions of O2,

CO2 and H2O with char and coke particles were controlled by mass transfer in the gas

film and chemical reactions. As can be seen in Figure 2.14, the coal burnout was

decreased with an increase in injection rate, leading to more unburnt char generated in

the raceway.

Figure 2.14 Relationship between coal burnout and injection rate.

34

2.2.1.6 Model of Sato et al.

Sato et al. (1996) introduced a dispersion angle of coal plume in the blowpipe as

indicated in Figure 2.15. In the model, the dispersion angle increased with an increase in

lance diameter. Furthermore, the dispersion of multiple lances in the blowpipe was

simulated by that of single one with larger diameter. In Figure 2.16, the performance of

double lance and single lance injection in terms of coal burnout were compared.

Obviously, the double lance operation achieved higher coal burnout than that given by

the single one due to better dispersion of coal particles with hot blast gas.

Figure 2.15 Modelling concept of pulverised coal flows in blow pipe.

Figure 2.16 Influence of injection lance on coal burnout along tuyere axis.

35

2.2.1.7 Summary of 1-D model

Table 2.1 shows a summary of key 1-D models and the sub-models applied. A brief

summary of the calculation results obtained from the 1-D model is shown below.

(1) The calculated peak levels of gas temperature and concentration of CO2 are shifted

towards the tuyere exit with an increase of PCI rate. Practically, for the prevention of

tuyere from burn out, the lance tip should be moved towards the tuyere exit when the

coal injection rate is significantly increased.

(2) Double lance is more effective than single one in mixing injected coal particles with

hot blast gas. Consequently, the coal burnout can be improved.

It is noted that many important operational features, especially relating to the flow

spatial properties of the process, are not considered in the 1-D models. For example, in

the 1-D model, the injected coal particles were assumed to be dispersed uniformly, or

dispersed with a growth angle in the hot blast gas. In fact, the assumptions are quite

different from the observations given by the tuyere monitoring system (as shown in

Figure 2.7). Practically, accurate information about the dispersion of injected coal is

required for the PCI operation, because it is critical not only to control the

devolatilisation of injected coal particles in the raceway, but also to protect tuyeres from

failure due to PCI abrasion. Therefore the practical application of the 1-D models in PCI

operation may be limited.

36

Table 2.1 PCI combustion models and sub models used.

Author(s) Kuwabara et al. (1981) Burgess et al. (1983)

Flows patterns 1-D flow 1-D flow

Coal particles are dispersed

uniformly with the blast gas

Calculation

domain

Blowpipe, tuyere and raceway

(cylindrical jetting space)

Cylindrical combustion

chamber

Devolatilisation - Two competing reaction model

Volatile Reaction - Instantaneous combustion

Char Reaction - C + 1/2 O2 = CO (combustion

heat to solid phase)

CO + 1/2 O2 =CO2

(combustion heat to gas phase)

Coke Reaction C+ O2 CO2

C+CO2 2CO

C +H2OH2 + CO

Coke temperature is assumed

to be 0.8 time of gas

temperature

-

Raceway shape As a cylindrical jetting space

with the same diameter of

tuyere and 1.5meters in length

Voidage: 0.6

Diameter of coke: 20 mm

-

Sensitivity

Analysis

Blast conditions (O2, moisture,

temperature)

Blast conditions, coal type,

coal size, injection rate.

Validation Measured data by Inatani et

al. (1973)

Measured temperatures from

combustion rig

Remarks Without injecting auxiliary

fuel

Raceway not included

Author(s) He et al. (1986) Jamaluddin et al. (1986)

37

Table 2.1 contd.

Flows patterns 1-D flow

Injected coal dispersed

uniformly with the blast air

Quasi-two-dimensional flow

Modelled as a combination of

plug flow regions, viz., coal

flow, blast, and mixing layer

which is quantified through a

jet-spread angle

Calculation

domain



Coke and injected coal

dispersed uniformly with the

blast air

Modelled as a one dimensional

jet with recirculating gas and

coke particles entrained at the

regions of 0.2 – 0.73 depth of

raceway roof

Devolatilisation First order reaction

Two-competing reaction model

Volatile reaction Instantaneous combustion Instantaneous combustion

Char reaction C + 1/2 O2 CO

C + CO2 2CO

C + H2O H2 + CO

C + 1/2 O2 CO

C + CO2 2CO

C + H2O CO + H2

C + O CO

C + OH CO + 1/2 H2

Coke reaction in

raceway

C + 1/2 O2 CO

C + CO2 2CO

C + H2O H2 + CO

C + 1/2 O2 CO

C + CO2 2CO

C + H2O CO + H2

Raceway shape Cylindrical jet with spreading

gas through coke wall

Coke diameter: 20mm

Proposed as a pipe with

non-smooth wall (coke bed)

Coke diameter: 30 mm

Sensitivity

Analysis

Blast conditions, injection rate,

coal properties,

Blast conditions, injection rate,

soot conc., coal properties

Validation Measured data by Ariyama et

al. (1994)

Measured data by Inatani et al.

(1973)

Remarks The model comprises warm-up

zone and burning zone

The first model considering

recirculating gas entering the

raceway

38

Table 2.1 contd.

Author(s) Yamagata et al. ( 1992 ) Sato et al. (1996)

Flow patterns 1-D flow

Injected coal dispersed

uniformly with the blast air

Quasi-two-dimensional flow

Coal plume is dispersed along

an angle determined by the

lance diameter

Calculation

domain

Blowpipe tuyere and raceway


Coke and injected coal

dispersed uniformly with the

blast air

Blowpipe and coke bed

(cylindrical space)

Injected coal dispersed into a

packed bed (the void fraction of

the raceway is linearly

predictable)

Devolatilisation Overall reaction model:

DV/dt = k(Vo-V)0.5

K=4.0 105 exp(-17500/RT)

Single overall reaction model

Volatile reaction - n≦2m+2

CmHn + m/2O2 = mCO + n/2H2

n>2m+2

CmH2m+2 +m/2O2 = mCO +

(m+1)H2


C + CO2 2CO

C + H2O CO + H2

C + 1/2 O2 CO

C + CO2 2CO

C + H2O CO + H2

Coke reaction in

raceway

C + 1/2 O2 CO

C + CO2 2CO

C + H2O CO + H2

C + 1/2 O2 CO

C + CO2 2CO

C + H2O CO + H2

Raceway shape 480mm in depth Cylindrical jet

Sensitivity

Analysis

Injection rate, volatile content Multiple lance and oxy-coal

lance injection

Validation Measured burnout and

gas composition from a coke

bed

Measured dada by Fukuyama

No5 blast furnace

Remarks The first 1D model available

for the evaluation of double

lance

The first 1D model available

for the evaluation of multiple

lance and oxy-coal lance

39

Table 2.1 contd.

Author(s) Aoki et al. (1993) Takeda and Lockwood (1997)

Flow patterns Axi-symmetric 2-D flow

Standard k- turbulent model

for gas phase

Lagrangian approach for the

trajectory of coal particles

Eulerian approach for coke

Axi-symmetric 2-D flow

Modified k- turbulent model

for gas phase

Lagrangian calculation for the

trajectory of coal particles

Calculation

domain

Blowpipe tuyere and raceway

determined by the force

balance between gas and solid

Blowpipe, tuyere and raceway,

including a jet zone and

transition zone

Devolatilisation Two competing reaction model Single overall reaction model

Q factor: 1.5

Volatile reaction Eddy dissipation model Eddy dissipation model

Char reaction C + O2 CO2

C + 1/2 O2 CO

C + CO2 2CO

C + H2O H2 + CO

C + 1/2 O2 CO

C + CO2 2CO

Coke reaction in

raceway

C + O2 CO2

C + 1/2 O2 CO

C + CO2 2CO

C + H2O H2 + CO

C + O2 CO2

C + CO2 2CO

C + H2O H2 + CO

Raceway shape Determined by force balance

among gravity force of coke,

gas-solid interaction force and

fractional force of coke at the

boundary of raceway

Including jet zone and

transition zone

Sensitivity

Analysis

Influence of coal size and

volatile content on the

dispersion of coal plume

Ways of oxygen enrichment,

lance configuration, coal size

Validation Raceway shape by cold model Raceway gas compositions

measured form an operating

furnace

Remarks Residence time of PC in the

combustion region: 12 ms

A high turbulent lance applied

at Kawasaki Steel

40

Table 2.1 contd.

Author(s) Haywood et al. (1994) Picard (2001)

Flow patterns Axi-symmetric 2-D flow


for gas phase

Lagrangian approach for

calculating trajectory of coal

particles

3-D CFD based model


for gas phase



particles

Calculation

domain

A cylindrical combustion

chamber

Tuyere, empty raceway and

porosity media surrounding the

raceway

Devolatilisation Two competing reaction model Two competing reaction model

Volatile reaction Mass fraction/ PDF model Chemical reaction rate

(Arrhenius)


C + CO2 2CO

C + H2O H2 + CO

C + 1/2 O2 CO

Limited either by chemical

kinetic or diffusion of oxygen

Coke reaction in

raceway

Raceway shape Only blowpipe Measurements results from

blast furnaces

Sensitivity

Analysis

Injection rate and coal volatile

content

Particle size, coal VM, oxygen

enrichment, injection rate

Validation Gas composition measured

from a combustion rig

Remarks One of pioneers in PCI

combustion model by CFD

(1)Raceway shape featured by

tuyere coke sampling

(2)One of pioneers in 3D PCI

combustion model

41

Table 2.1 contd.

Author(s) Guo et al. (2005) Shen et al. (2008, 2009a,

2009b)

Flow patterns 3-D CFD based model

RNG k–e turbulence model for

gas phase



particles

3-D CFD based model


for gas phase



particles

Calculation

domain

blowpipe and combustion

chamber

blowpipe and combustion

chamber


Stoichiometric parameters

given by the tests of Australian

coals with VM ranged from

12-40%

Volatile reaction eddy break up model eddy dissipation model

Char reaction ΦC + O2 → 2(Φ-1) CO + (2-Φ)

CO2

Φ is function of particle

temperature (Gibb, 1985)

ΦC + O2 → 2(Φ-1) CO + (2-Φ)

CO2

CO2 2CO

C + H2O H2 + CO

Coke reaction in

raceway

Raceway shape Combustion chamber Only blowpipe

Sensitivity

Analysis

Coal properties (size, VM and

flow rate), lance configuration,

cooling gas types for co-axial

lance


flow rate), oxygen level in blast

gas, lance configuration,


lance, coal blend

Validation Burnout given by the

combustion tests (Mathieson, et

al, 2005)

Measured burnout given by the


al, 2005)

Remarks (1) evaluation of oxy-coal

lance

(2) fine particles dispersed

more widely

42

Table 2.1 contd.

Author(s) Shen et al. (2011) Gu et al. (2008)


Standard k–εturbulence model

for gas phase



particles

3-D CFD based model

k--kp two-phase turbulence

model for gas and particle

phases

Calculation

domain

(1) Blowpipe, tuyere and

raceway; (2)deadman;

(3)dripping zone; and (4)

cohesive zone

Blowpipe and tuyere


Volatile reaction Eddy dissipation model Eddy break up model

Char reaction ΦC + O2 → 2(Φ-1) CO + (2-Φ)

CO2

C + CO2 2CO

C + H2O H2 + CO

C + O2 CO2

2C + O2 2CO

C + CO2 2CO

C + H2O H2 + CO

Coke reaction in

raceway

C + CO2 2CO

Raceway shape Designed in a shape of balloon Determined by CFD

calculation (Selvarasu et al.,

2006)

Sensitivity

Analysis


flow rate), oxygen level in blast

gas, lance configuration,


lance

Tuyere diameter

Validation Measured burnout given by the


al, 2005)

Remarks Recirculation of fine particles

(<70um) within raceway

Fine particles dispersed more

widely

43

Table 2.1 contd.

Author(s) Nogami et al. (2004)


Standard k–εturbulence model

for gas and solid phases

DEM for coke movement

Calculation

domain


Devolatilisation Two competing reaction model

Volatile reaction Eddy dissipation model

Char reaction C + O2 CO2

Coke reaction in

raceway

C + O2 CO2

C + CO2 2CO

C + H2O H2 + CO

Raceway shape Determined by DEM model

Sensitivity

Analysis

Oxygen enrichment

With/without PCI

Top gas recycling operation

Validation (1) Measured gas composition

along tuyere axis in the hot

model experiments

(2) Comparison of calculated

raceway shapes with the

observations in the hot model

experiments

Remarks (1) 2-D raceway shape observed

through the hot experimental

model

(2) No recirculating flow found

in the raceway

44

2.2.2 Development of two-dimensional model

By the 1990s, few two-dimensional (2-D) models came into place to assist the search

for optimising blast conditions and design improvements in coal injection and its

operation.

2.2.2.1 Model of Aoki et al.

To analyse the flow patterns, heat transfer and reactions of injected coal in the blowpipe,

tuyere and raceway, a 2-D mathematical model of pulverised coal combustion was

developed by Aoki et al (1993). Calculations of coal combustion in the blowpipe-tuyere

region were carried out in a cylindrical space described by a 2D axi-symmetric system

(Figure 2.17). In the coke bed region, the space was considered as a 2D x-y plane set on

the tuyere axis. The calculated results in the outlet of blowpipe-tuyere region were used

as the inlet conditions of coke bed region. The two-equation turbulence model, k-ε

model, was employed to calculate the turbulent dispersion of coal flow, heat transfer

and reactions. In both domains, the gas phase was treated as continuous phase, and the

motion of coal particles was solved by Lagrangian approach (Kuo, 1986). Coke

particles in the raceway were assumed as a quasi-fluid, and the two-phase flow of coke

and gas were calculated in the model. The effect of coal volatile content on the

combustion was analysed. The effect of volatile matter content of coal on the gas

temperature distribution in the blowpipe is shown Figure 2.18. Clearly, an increase in

coal volatile content caused the rise in gas temperature within the region of

blowpipe-tuyere. The particle trajectories of low volatile coal (VM: 19.5%) are shown

in Figure 2.19. It was found particles under 30 μm could not disperse towards the radial

direction, while some particles over 65 μm dispersed in close to the inner wall of the

tuyere. This was caused by the small radial diffusion of fine particles with a diameter

less than 30 μm in the blowpipe where the gas velocity was extremely high (Ishii,

2000).

45

Figure 2.20 shows the coal particles reached the inner wall of tuyere when the particle

size was bigger than 65 μm for the high volatile coal (VM: 44.6%) injection. It was

explained that when high volatile coals were injected, more volatile was released,

followed by early ignition and combustion of volatile in the region of blowpipe-tuyere.

As a result, the dispersion of coal particles in the region could be significantly enhanced

by the intensified turbulence. Notably, from the calculated particle trajectories, it was

found that the mean residence time of pulverised coal in the region of blowpipe-tuyere

was 5 ms, and it was 7 ms in raceway. Totally, the travelling time for coal particles

from the injection point of lance to the raceway boundary was around 12 ms.

Computational results of the gas composition and temperature distributions in the

raceway along the centre line of tuyere are presented in Figure 2.21 for both with and

without PCI operation. Peak positions of CO2 and gas temperature were moved towards

the tuyere nose with PCI operation. After reaching the peak level in the raceway, the gas

temperature decreased, resulting from the endothermic reaction of CO2 with carbon

from unburnt char and coke.

Figure 2.17 Calculation domain in the region of blowpipe-tuyere.

46

Figure 2.19 Calculated particle trajectories of low volatile coal injection in the

blowpipe.

Figure 2.18 Effect of coal volatile content on gas distribution in the blowpipe.

47

Figure 2.21 Distributions of gas composition and gas temperature along the

centre line of tuyere.

Figure 2.20 Calculated particle trajectories of high volatile coal injection in the

blowpipe.

48

2.2.2.2 Model of Takeda and Lockwood

Takeda and Lockwood (1997) have developed an integrated axi-symmetric 2D

mathematical model of pulverised coal combustion in the regions of blowpipe and

raceway. In this model, a cylinder was used to represent the region of blowpipe and

tuyere, and the raceway was modelled into two regions: (1) a jet core region with an

empty space, and (2) a transition region with the linear decrease in voidage to the

raceway boundary. The assumed configuration of calculation domains is shown in

Figure 2.22. A modified turbulence model (k-lm model) was adopted for predicting the

gas and particle flows, and the Lagrangian approach for particle phase with stochastic

fluctuation was used to calculate turbulent dispersion of coal plume. The volatile

combustion in gas phase was simulated by the eddy dissipation model (Magnussen and

Hjertager, 1976).

Figure 2.22 Schematic representation of raceway structure used in the simulation.

49

The predicted gas temperature and oxygen concentration for the injection of a high

volatile coal (38.7%) at a rate of 50 kg/tHM are presented in Figure 2.23. The gas

temperature increased slightly in the blowpipe region where the injected coal particles

were in the early stages of heating and devolatilisation. On the other hand, rapid

changes in gas temperature and oxygen concentration were found in the raceway. This

indicated that the coal combustion was intensified by the volatile combustion.

Figure 2.23 Contours of gas temperature and oxygen concentration in the regions of

blowpipe and raceway.

50

In Takeda’s research (1994), four types of injection lances shown in Figure 2.24 were

also evaluated. Swirl motion was imposed to the carrier gas and coal flow in Case D.

An outer diameter of the injection lance was enlarged from 0.02 to 0.04m in order to

enhance gas turbulence in the vicinity of the injection lance in Case E. Radial velocities

were added to initial gas and particle velocity at the exit of the injection lance in Case F.

Figure 2.25 compares the performance of the lances in terms of coal burnout. As

indicated in the comparison, remarkable improvement in coal burnout could be achieved

by Case E because an increase in the outer diameter of the injection lance gave better

mixing of the coal plume and hot blast gas through the intense turbulence downstream

of the injection lance.

Figure 2.24 Lance design for coal injection.

51

2.2.2.3 Model of Haywood et al.

In an attempt to promote the utilisation of Australian coals in PCI operation, a 2-D CFD

(Computational Fluid Dynamics) based model has been developed by Haywood et al.

(1994). The coal burning characteristics in the experimental combustion rig (Figure 2.26)

was investigated. In the model, the geometry of the rig was considered to be an

axi-symmetric system. The model consisted of several sub-models for describing

turbulent mixing, particle tracking, pyrolysis of coal, volatile combustion and char

combustion.

Figure 2.27a and 2.27b shows the oxygen mass fraction (top) and gas temperature

contours (bottom) in the near injector region with a low volatile coal (VM: 9.67%) and

a high volatile coal (VM: 39.0%) injection respectively. A comparison of the gas

temperature contours indicated that early ignition could be provided by injecting high

volatile coal, resulting in rapid decrease of oxygen at the rig axis. This implied the

Figure 2.25 Coal burnout comparisons for the various modifications of lance.

52

oxygen consumption rate in the centre region was faster than that of oxygen diffusion

from bulk gas. Therefore it was concluded that mixing will be a key factor for high PCI

rates.

(a) low volatile coal injection

(b) high volatile coal injection

Figure 2.27 Oxygen mass fraction (top) and gas temperature contours (bottom) in

the near injector region. low volatile coal (a) low volatile coal; (b) high volatile

coal injection.

Figure 2.26 Schematic of the experimental combustion rig.

53


In the 2-D models, the gas phase is described by the transport equations of the

continuum phase, and the particle behaviours are calculated by Lagrangian approach, in

which the effects of the fluctuation of gas are considered on the particle dispersion. This

advantage allows one to precisely calculate the heating and devolatilisation of coal

particles, followed by volatile combustion. Table 2.1 shows a summary of key 2-D

models and the sub-models applied. The main findings obtained form the models are

indicated below.

(1) The dispersion of coal particle in the region of blowpipe-tuyere can be enhanced by

volatile release and combustion. Therefore, the extent of particle dispersion in radial

direction for high volatile coals is wider than that of low ones.

(2) The estimated resident time of coal particles in the regions of blowpipe, tuyere and

raceway is around 13 ms based on the 2-D calculation frame.

(3) As high volatile coals are injected, the gas temperature increases quickly in the

region of blowpipe-tuyere, while the oxygen concentration drastically decreases

within the coal plume.

(4) The burnout of injected coal in the region of blowpipe-tuyere-raceway is limited by

(a) short resident time of coal particles in the region, (b) mixing of coal particles

with hot blast gas, and (c) oxygen diffusion from bulk gas to the interior of coal

plume, and (d) coal properties, mainly volatile content and size.

Basically, the axi-symmetric 2D modes developed for the pulverised coal combustion in

the region of blowpipe-tuyere-raceway can generate results that are qualitatively useful

in practice (Chattopadhyay et al, 2010). However, with the axi-symmetric assumption,

some operation features can not be precisely analysed. For an example, all computations

are carried out with a lance located at the tuyere centre, resulting in underestimating the

influence of lance designs (such as inlet angle and configuration etc.) on the gas flow

pattern, trajectory of coal particle and coal combustion.

54

2.2.3 Development of 3-D model

To generate results which can be used for realistic operations, 3-D models by CFD

codes have been developed to simulate the PCI process under actual furnace conditions.

The models developed by different teams are described below.

2.2.3.1 Model of Picard

A 3D numerical model based on a CFD code was reported by Picard (2001).The

combustion domain of the model included tuyere, empty raceway (without coke) and a

porous media surrounding the raceway, as indicated in Figure 2.28. The raceway

dimensions were determined by the practical conditions of No5 blast furnace of

Dillingen works, and the porosity and deadman coke size were coming from probing

measurement performed in the Fos works. The standard k-ε model was applied for gas

turbulence, while the Lagrangian approach was used for simulating the dispersion of

coal particles in the gas. The release of volatile matters was modelled by two competing

devolatilisation model. In the model, the volatile matters were considered to be a

mixture of C3H8 and C6H6. In the model, the author considered radiative heat transfer

can be ignored in comparison with the convection heat transfer.

To understand the behaviours of coal particles in the region of tuyere-raceway, the

trajectories of coal particles were calculated in the model, as shown in Figure 2.29. It is

found that most of coal particles bumped into the raceway border and escaped there due

to high inertia. The particles average residence time in the raceway was 29 ms. Some of

particles which were smaller than 48 μm in diameter follow the recirculating gases in

the backward, resulting in longer residence time, as revealed in Figure 2.30. The effects

of injection rate and oxygen enrichment on coal burnout were examined in the model.

With 23% oxygen level in the hot blast gas, the coal burnout was decreased from 86.3

to 80.4% when the injection rate was increased from 170 to 230 kg/tHM. Besides, the

coal burnout could be promoted with higher oxygen level in the hot blast gas.

55

Figure 2.28 3D meshed calculation domain.

Figure 2.29 Trajectories of coal particles within the tuyere and raceway.

Figure 2.30 Evolution of coal residence time within the raceway versus particle

diameter.

56

2.2.3.2 Model of Guo et al.

In the research of Guo et al., (2005), the pulverised coal combustion behaviours in the

combustion rig (Figure 2.31) were simulated using CFD code. The empty combustion

chamber was adopted to simulate the raceway of blast furnace. In the model,

two-competing reaction model was chosen to simulate the de-volatilisation of coal, and

eddy breakup model was employed for the volatile combustion in gas phase. The model

assumed that the time scale of reaction was much shorter than that of turbulence and the

gas–gas reaction was controlled by turbulent diffusion.

The calculation results showed recirculation gas flow in the combustion chamber as can

be seen in Figure 2.32. The recirculation was resulted from the expansion design

between the tuyere and the combustion chamber. Figure 2.33 indicates typical coal

particle trajectories coloured by their sizes. The extent of radial dispersion of particles

exiting from a simple straight lance was very limited in the tuyere. On the other hand, a

particle segregation phenomenon could be found in the combustion chamber. Upon

exiting the injection lance, large particles with larger momentum maintained their initial

direction (the lance axial direction), while fine particle dispersed more widely. This is

quite different from the calculated particle trajectories given by Aoki et al. (1993) as

indicated in Figures 2.19 and 2.20. Figure 2.34 shows oxygen and volatile matter

distributions in the gas phase. The oxygen concentration in the coal plume region was

always low as the combustion of the volatile and char oxidation consume oxygen. Thus

the transport of oxygen primarily controlled the combustion rate.

57

Figure 2.31 Main dimensions (in mm) of the coal combustion model (plan

view).

Figure 2.32 Gas velocity vectors in Y-Z plane (a) vector length to scale, and (b)

vector normalised showing recirculation zone.

58

Figure 2.33 Typical particle trajectories with colour scaled to particle size.

Figure 2.34 Gas species fraction isopleths in Y–Z plane. (a) Oxygen; (b)

volatile matter.

59

Figure 2.35 shows the particle characteristics (devolatilisation and burnout) as a

function of axial distance from the lance tip along the centreline. At 0.2m, particles

below 30μm have nearly completed devolatilisation, whereas those above 80μm have

hardly started (Figure 2.35a). Up to 0.5m from the lance tip, the burnout levels were

significantly different for particles of different sizes (Figure 2.35b); however, beyond

0.5 m, there was little variation—at this distance, slow char oxidation was prevalent. In

other words, the coal burnout is mainly contributed by the release of volatile matter

from coal particles, and the char combustion is limited due to insufficient oxygen

around the particles.

(a) (b)

Figure 2.35 Calculated mass fraction of volatiles (a) and calculated mean coal

burnout (b) as a function of distance from the lance tip along the centreline

averaged for different particle sizes (in μm).

60

2.2.3.3 Models of Shen et al.

A series of PCI calculation models have been established by Shen et al. (2008, 2009a,

2009b) to analyse the combustion behaviours of injected coal in the empty tuyere

(Figure 2.31). Based on the measured Q factors of coals contenting 12-39 % volatile

matter (Ueno et al., 1993; Niksa et al., 1984; Maloney and Jenkins, 1984), a correlation

of stoichiometric parameters for two competing devolatilisation model was made (Shen

et al., 2008). In addition, char gasification reactions (C + CO2 = 2CO and C + H2O =

CO + H2) were considered in the model. Figure 2.36 compares the calculated burnouts

(with/ without considering char gasification) with those generated by Guo et al. (2005).

Case 1 and Case 2 were compared for downstream and upstream, respectively. In the

upstream, late ignition was found with the model. As for the downstream, the two

models predicted a similar burnout evolution beyond 0.6 m for the base condition. It is

found from Cases 2 and 3 that burnout curves were similar upstream, quite different

downstream. With the char gasification reactions, the downstream burnout level was

increased from 65 to 75%.

To evaluate the combustion performance of coal blend in the combustion rig, three

cases were simulated and compared by Shen et al. (2009a):

1. Case I: the blend of PC_A and PC_B, with the blend fraction 50% + 50%;

2. Case II: single coal case of PC_A (a larger but higher volatile(HV) coal); and

3. Case III: single coal case of PC_B (a fine but lower volatile (LV) coal)

Figure 2.37 compares the burnout evolutions of Cases I, II and III: (a) along the

centreline; and (b) at the exit. Along the centreline (Figure 2.37a), the final burnout of

Case I was closer to Case II, i.e., higher than the average value of the final burnouts of

two coals. Conversely, when the final burnout was calculated at the exit area (Figure

2.37b), the burnout of Case I was even slightly higher than both Cases II and III. That is,

the overall burnout of coal blend showed a non-additivity, i.e., synergistic effect. It was

concluded that the chemical interactions between two components in terms of particle

61

temperature and volatile content were responsible for the synergistic effect: the HV coal

releases more VM, helping form a higher gas temperature field, which then heated up

the LV coal and promoted its devolatilisation and combustion.

The effects of blast conditions on pulverised coal combustion were examined by Shen et

al. (2009b). It is found that the combustion efficiency of the injected coal could be

improved by increasing blast air temperature. However, as shown in Figure 2.38, the

burnout of the coal could not be increased further when the blast air temperature was

higher than 1200oC. Figure 2.39 shows the influence of oxygen enrichment on final

burnout. Clearly, the final burnout could be improved as more oxygen was added into

the blast.

The performance of coaxial lance with cooling gas of methane, oxygen and air flowing

through the annulus of the lance was compared by Shen et al. (2009c). Figure 2.40

shows the effect of cooling gas type on coal burnout. Note that the oxygen enrichment

in blast was kept with the same for all the three cases. It is shown that for the three types

of cooling gas, oxygen gave the highest burnout of 67%. In addition, a linear

relationship was found between the burnout and atomic O/C ratio in the gases delivered

to the tuyere. This was because, comparing the O2 distributions on a cross-section at the

distance of 550 mm from the lance tip (Figure 2.40b), by using methane as the cooling

gas, a larger amount of O2 was consumed when the methane was burning together with

the VM competitively. As a result, the VM combustion was slowed down, which

decreased the final burnout significantly. In addition, the amount of O2 available to the

subsequent char reactions was also reduced.

62

Figure 2.36 Comparison of burnout evolutions predicted by the previous

model (Case 1), and present model with char gasification reactions (Case 2)

and without char gasification reactions (Case 3).

Figure 2.37 Burnout for Cases I, II and III: (a) along the centreline and (b) at the

exit.

63

Figure 2.38 Effect of blast temperature on burnouts at the distances of 300 mm and

925 mm from the lance tip, respectively.

Figure 2.39 Effect of oxygen enrichment on coal burnout.

64

Figure 2.40 Effect of cooling gas type: (a) final burnout and (b) O2

distributions at the cross-plane of 550 mm from the lance tip.

65

For further study of the PCI operation in the regions of lance blowpipe tuyere raceway

and coke bed, a 3D mathematical model of the combustion of pulverised coal and coke

was developed by Shen et al. (2011) to simulate in-furnace phenomena of pulverised

coal injection in an ironmaking blast furnace. The model integrated not only pulverised

coal combustion model in the blowpipe-tuyere-raceway-coke bed (deadman zone) but

also coke combustion model in the coke bed. Figure 2.41 reveals the calculation domain

in the model.

Figure 2.42a shows the gas velocity vectors in the raceway cavity. It was found a

high-speed jet (220m/s) formed along the tuyere axis, which was similar to the cold

model observations by Inatani et al. (1976). After reaching the raceway boundary, the

gas flow started a large-scale recirculation above the main gas flow jet in the raceway. It

was probably because the porosity was set at a level of 0.25 for the deadman. In the

coke bed, as shown in Figure 2.42b, gas velocities decreased rapidly to <5 m/s within a

very short distance. It is shown in Figure 2.39c that, corresponding to the gas flow, the

coal particle trajectories inside the raceway had two different flow patterns: (i) an

inclined main coal plume located along the lower part of the raceway, where fine

particles were observed at the upper part of the plume initially and then left the main

coal plume before reaching the end of the raceway; and (ii) a large-scale recirculation of

the fine particles of up to 70 μm around the raceway centre. The coke bed also showed

two flow patterns of coal particles accordingly (Figure 2.42c): (i) the main coal plume

(relatively large particles of around 100 μm) penetrated into the deadman zone; (ii) the

recirculating fine particles exited mainly from the top of the raceway and then moved

upward into the dripping zone. Figure 2.42d shows the residence time of coal particles

along the main coal plume was around 10–50 ms before reaching the end of the raceway,

while the recirculating coal particles might be up to 0.9 s in the raceway. On the other

hand, compared with the raceway in the coke bed, the travelling time of the particles

penetrating the coke bed was quite long, around 1.0 s. The travelling time was even

longer in the deadman compared to the dripping zone. Moreover, Figure 2.43 shows the

66

coal burnout distribution along the particle trajectories in the coke bed. It is shown in

the simulation that the burnouts varied greatly in different zones of the coke bed, nearly

100% in the dripping zone, and around 75% in the deadman zone, in which the char

burnout could not be promoted further due to lack of O2 and CO2. Notably, the char in

the deadman could be consumed by the direct reduction of FeO in slag (Iwanaga, 1991).

Figure 2.41 Geometry of the model: (a), the whole model; (b), porosity distribution

(Zone 0: 1, Zone 1: 0.25, Zone 2: 0.5, Zone 3: 0.4); (c), blowpipe and raceway; and

(d), lance tip. The detailed dimensions are, (1) for blowpipe, radius: 90 mm, and

length: 800 mm; (2) for tuyere, radius: 75/90 mm, and length: 135 mm; (3) for

raceway, depth: 1600 mm, height: 1000 mm (925 + 75), and width: 710 mm; and

(4) for coke bed, depth: 3700 mm, height: 4500 mm, and width: 1000 mm.

67

Figure 2.42 Flow pattern of gas-particle flow: (a), vectors of gas phase in the

raceway; (b), streamlines of gas flow; (c), particle trajectories coloured by

particle mean size; and (d), particle trajectories coloured by particle travelling

time.

Figure 2.43 Combustion characteristics of coal along particle trajectories in the coke

bed.

68

2.2.3.4 Model of Gu et al.

Gu et al. (2008) has reported a three-dimensional multi-phase model using Eulerian

approach has been developed to simulate the coal heat transfer, devolatilisation and

combustion process inside a blast furnace tuyere. The calculation was based on an

Eulerian coordinate, while the particle phase was treated as a pseudo-fluid interacting

with the gas phase. The k–ε–kp two-phase turbulence model was used to model the gas

and particle turbulence, and the eddy dispersion model was adopted to quantify the

effect of turbulence on the combustion rates of volatiles, carbon monoxide and

hydrogen. The heterogeneous reactions considered in the model included char oxidation

with oxygen, carbon dioxide and moisture. The computational domain is shown in

Figure 2.44. Obviously, the raceway shape looks like a balloon.

Figure 2.45 shows the distributions of gas velocity vectors and gas temperature in the

computational domain. As indicated in Figure 2.45a, a high speed gas exiting the tuyere

was found to form a jet space inside the raceway, while a large scale recirculation zone

was generated between the jet and coke bed. Owing to well mix of oxygen (31.5% in

the hot blast gas) and fuel, the location of highest temperature (3115K) coincided with

the location of gas recirculation, as shown in Figure 2.45b. The effect of tuyere diameter

to the coal combustion was also examined. Figure 2.46 shows the coal burnout at the

exit of the computational domain. It can be seen that coal burnout was promoted from

60.6% to 85.3% when the tuyere diameter was enlarged from 0.156 m to 0.165m. It

could be explained by an increase of residence time of coal particles in the raceway

when the tuyere diameter was enlarged. It should be noted that the blast furnaces of

CSC have reduced their tuyere diameters for obtaining deeper penetration of hot blast

gas into the deadman, since it is thought of as an effective countermeasure to keep a

stable thermal level within the hearth.

69

Figure 2.45 Distributions of (a) gas velocity vectors and (b) gas temperature (K)

in the computational domain.

(a) (b)

Figure 2.44 Schematic of computational domain: (a) side view; (b) top view.

(a)

(b)

70

2.2.3.5 Model of Nogami et al.

Nogami et al. (2004) has reported a 3D transient analysis model for raceway and the

surrounding coke bed. In the model Finite Differencing Method (FDM) and Discrete

Element Method (DEM) were used for featuring the coke movement. As a result, the

trajectories of coke particles could be traced within and around the raceway.

Gasification rate of coke particles was calculated by shrinking core model, in which

chemical reaction on the surface of particle, mass transfer in boundary layer, and gas

diffusion inside particle were considered. The model was applied to the combustion

tests in an experimental coke bed as shown in Figure 2.47. The calculated raceway

shapes were validated by that obtained from the observations of the tests, as revealed in

Figure 2.48. It can be found that the raceway for PCI operation was slightly lager than

that for all coke operation. With the design of the combustion rig, the development of

raceway was constrained in both side walls of the rig, therefore the raceway could be

considered as 2-dimensional. Figure 2.49 shows (a) the calculated raceway shape, and

(b) the calculated gas velocity vectors in the lower zone of blast furnace. As indicated in

Figure 2.49b, the blast proceeded as a high speed jet along tuyere axis due to its inertia,

Figure 2.46 Coal burnout at the exit of the computational zone.

71

and there was no recirculation flow in the roof of the raceway. It is different from the

gas flow patterns in the raceway given by Shen et al. (2011) and Gu et al. (2008).

Figure 2.47 Schematic figure of hot model.

Figure 2.48 Comparison of calculated raceway shape with observation of

test. (a) All coke operation. (b) PCI operation.

72


As reported above, 3-D studies in PCI operation have been developed from blowpipe to

raceway, even to the lower zone of blast furnace. A summary of key 3-D models and the

sub-models applied can be found in Table 2.1. Discussion for the 3-D models is made

below:

(1) The segregation of coal particles can be found after exiting injection lance. Large

particles with larger momentum maintain their initial direction (the lance axial

direction), while fine particle disperse more widely.

(a)

(b)

Figure 2.49 Characteristics of raceway: (a) calculated raceway shape, and (b)

calculated gas velocity vectors.

73

(2) Owing to lack of O2 and CO2, the char burnout can not be promoted further when

the char particles enter the deadman. It should be noted that the char in the deadman

zone can be consumed by the direct reduction of FeO in slag.

(3) After leaving tuyere nose, blast proceeds as a high speed jet along tuyere axis due to

its inertia. In the region above the jet, the recirculation of gas may occur as a result

of low porosity around the raceway.

(4) The coal burnout in the raceway can be significantly promoted when the tuyere

diameter is enlarged. However, to keep a stable thermal level in the deadman, the

tuyere diameter has been reduced at CSC for obtaining deeper penetration of hot

blast gas into the deadman.

As revealed above, the 3-D models have been successfully developed for featuring the

coal combustion characteristics in the region of blowpipe-tuyere-raceway. The

parameters which are related to coal burnout are especially examined in the models. In

fact, higher pressure resistance (or poor permeability) has been encountered in the lower

zone of CSC’s blast furnaces when attempting to increase the PCI rates. For a stable

operation of blast furnace with high PCI rates, countermeasures for the reduction of

pressure loss caused by coal combustion in the raceway is needed at CSC.

2.3 Sub-models for integrated calculation

The combustion of pulverised coal in the region of blowpipe-tuyere-raceway includes

many physical and chemical processes as shown in Figure 2.50. After exiting the

injection lance tip, the coal particles are dispersed to the hot blast gas, while the coal

particles are heated by hot blast gas, and the moisture evaporates. Coal devolatilisation

then occurs after further heating. The released products, mainly hydrocarbon, are ignited,

which causes an increase in the temperatures. As a result, the devolatilisation of coal

can be enhanced. Finally, the residual char particles can be consumed by oxidants (O2,

CO2 and H2O). All of these processes take place sequentially with some overlap. Since

74

the residence time of coal particles in the combustion region is very short, generally less

than 20 ms, the characteristics of these physical and chemical processes are very

important for the effectiveness of a PCI system.

As mentioned above, the combustion of pulverised coal consists of several processes,

which can be described by some mathematical and empirical equations or simple model

component. These equations should be solved simultaneously to express entire coal

burning characteristics in the region of blowpipe-tuyere-raceway. In the CFD based

modelling for coal burning in the raceway, all the steps of geometry and grid

generations, boundary condition implementation and sub-model simulations are

integrated into a framework, as shown in Figure 2.51. Note that the use of appropriate

sub-models allows for a more accurate description of the coal burning in the combustion

region.

Figure 2.50 Illustration of combustion phenomena of pulverised coal (Ishii, 2000).

75

According to the experiments by Seeker et al. (1981), the combustion of a high volatile

coal in a high temperature environment (1750K) can be divided into three stages in

terms of residence time (Smoot and Smith, 1985): heat-up and devolatilisation (few ms

to 75 ms), followed by char oxidation, as shown in Figure 2.52. It implies that the coal

burnout in the raceway may be primarily contributed by the coal devolatilisation, rather

than by char oxidation and gasification. Therefore the sub-model review starts with the

coal devolatilisation, followed by other physical and chemical processes involved.

Solution

Gas phase

˙ Mass conservation

˙ Momentum conservation

˙ Turbulence model

˙ Energy conservation

˙ Turbulence combustion

model

Solid Phase

˙ Devolatilisation model

˙ Char combustion

˙ Reaction/ diffusion

controlling regime

˙ Momentum

conservation (particle

trajectory)


Initialisation

˙ Geometry (meshing)

˙ Boundary conditions

˙ Sub-models



model

Radiative Transport

Particle

properties

Radiative

energy

exchange

Figure 2.51 Framework of the CFD code and computational procedure of the

gas phase and solid (coal particle) phase (Du et al., 2007).

76

2.3.1 Devolatilisation of coal

Coal devolatilisation is a complex chemical and physical process. Various phenomena

are involved during coal devolatilisation, namely, heat transfer to and within the coal

particles, various bond-breaking and cross-linking reactions, and the transport of

volatile products. It should be noted that enhanced yields for rapid heating have been

correlated with “Q-factors”, which are the ratios of the weight loss after rapid heating to

the proximate volatile matter content of the coal, and values for various coals have been

reported (Anthony and Howard, 1976; Niksa and Lau, 1993; Yan et al., 2014).

As a matter of fact, a complete description of the chemical reactions that occur during

devolatilisation is not available. Practically, based on experimental results without

considering the chemical structure of coal and complicated physics process, simplified

models, such as the single overall reaction model (Badzioch and Hawksley, 1970) and

two-competing reaction model (Kobayashi et al., 1977), have been widely applied in the

40 μm

80 μm

Figure 2.52 Effects of coal size and residence time on physical and temperature

profile (Smoot and Smith, 1985)

77

CFD based simulations for pulverised coal combustion.

2.3.1.1 Single overall reaction model

The single overall reaction model proposed by Badzioch and Hawksley (1970) is the

simplest devolatilisation model, in which the kinetics of devolatilisation has been

simplified by assuming a first-order decomposition occurring uniformly throughout the

particle.

CharVMCoal k (2.1)

The coal conversion rate can be expressed as the lumped production rate of all volatile

species being proportional to the volatile matter yet to be released:

)( VVkdt

dV (2.2)

Where V∞ is the ultimate yield of volatiles at t = , i.e. the total volatile content of coal,

and k is the rate constant, usually expressed as an Arrhenius relationship.

pRTEAk /exp (2.3)

Where Tp is particle temperature, and A (pre-exponential factor) and E (activation

energy) are constants, determined experimentally for the coal. As coal devolatilisation is

not a single reaction but a wide range of overlapping decompositions, the use of a single

set of parameters to describe reactions occurring over a wide range of conditions may be

inadequate.

2.3.1.2 Two competing reaction model

Based on the assumption that the coal may decomposed via one of several possible

reaction paths depending upon the time-temperature history, two competing reaction

model has been developed by Kobayashi et al. (1977) to explain devolatilisation yields

http://en.wikipedia.org/wiki/Pre-exponential_factor

78

as a function of time, temperature and heating rate. The two parallel and competing

reactions are given as follows:

1111 )1( 1 VYSYCoal k (low temperature) (2.4)

2222 )1( 2 VYSYCoal k (high temperature) (2.5)

where Y, S, and V denote stoichiometric coefficient, char, and volatile, respectively. The

relative importance of the two equations is mainly determined by temperature.

Specifically, when the temperature is low, the devolatilisation reaction is dominated by

route one (Equation 2.4). Alternatively, it is governed by route two (Equation 2.5) once

the temperature is relatively high. Accordingly, the devolatilisation reaction kinetics are

written as:

CYkYkdt

dV 2211 (2.6)

pRTEAk /exp 111 (2.7)


To get calculation results that can sufficiently reflect the pulverised coal burning

characteristics, in a research work at CSC (Du, 2001), the kinetic parameters for two

competing reaction model and those for single overall reaction model were validated

using the measured gas temperatures from an experimental combustion rig, as show in

Figure 2.53 (Burgess et al., 1983). The investigated operation conditions are

summarised in Table 2.2. Details of the kinetic parameters are given in Table 2.3. In the

calculation, the volatile combustion was dealt by the probability density function (PDF)

approach. Figure 2.54 shows the comparison of predicted temperature distributions with

the gas temperatures measured along the blowpipe. It depicts that the temperature was

spatially uniform when using the parameters of Kobayashi et al. (1977), indicating that

the rate constants gave unrealistic devolatilisation rates for blast furnace operations.

Upon inspection of the prediction given by the single reaction model with the

79

parameters of Takeda (1994), the gas temperature increased at about 0.25 m from the

lance tip, resulting from the release and ignition of volatile matters. As can be seen in

Figure 2.54, the calculated temperature kept rising with moderate rates in the following

process. Obviously, the rates in region of 0.4-0.5 m from the tip were behind the

measurements. Instead, when the parameters proposed by Ubhayakar et al. (1976) were

employed, the calculated temperature profile exhibited a good agreement with the

measurements. As revealed, the calculated temperature distribution could be partitioned

into two stages: it composed of rapid rise in the upstream region and progressive

increase in the downstream one. The behaviour in the first region arose from the PC

devolatilisation or pyrolysis reaction followed by the combustion of the emitted volatile

matters with oxygen. Obviously, in this stage, the devolatilisation reaction was

accelerated by the route two (Equation 2.5) when the coal particles were heated up by

volatile combustion. In regard to the second region, the slow increase in temperature

was attributed to the reaction between char and oxygen. In a word, after the PC is blown

into the blowpipe, the chemical reaction is primarily achieved by the gas-phase

combustion (i.e., homogeneous reaction) and then implemented by the solid-phase

oxidation (i.e., heterogeneous reaction).

Apart from the importance of kinetic parameters, the stoichiometric coefficients Y1 and

Y2 play essential roles for the determination of the total amount of volatile being

released during the combustion. Since the route one (Equation 2.4) represents

low-temperature devolatilisation, Y1 is set to the fraction of volatiles given by proximate

analysis. The high temperature yield Y2 (in Equation 2.5) is often estimated as being

related to Y1. A summary of the relation between Y2 and Y1 used in the PCI combustion

models (Suzuki et al, 1986; Guo et al., 2003; Aoki et al., 1993; Du and Chen, 2006;

Guo et al., 2005; Haywood et al., 1994) has been reported by Shen et al. (2008), as

shown in Figure 2.55. Clearly, the ratio of 1.5 (red line) employed in this research is

acceptable for the modelling in comparison with the experimental results (Ueno et al.,

1993; Niksa et al., 1984; Maloney and Jenkins, 1984). Table 2.4 shows the kinetic

80

parameters and the stoichiometric coefficients employed in the PCI combustion models.

Table 2.2 Operational conditions selected by Burgess et al. (1983).

Rig diameter, mm 50

Hot blast temperature, K 1243

Hot blast velocity, m/s 68

Coal particle diameter, μm 40

Volatile matter of PC (db), % 35.9

PC injection rate, kg/h 5.3

Table 2.3 Three sets of parameters used for predicting PC devolatilisation.

Kobayashi et al.

(1977)*

Ubhayakar et al.

(1976)*

Takeda (1994)**

A1, 1/s 200000 3.7×105 8.36×10

4

A2, 1/s 1.3×107 1.46×10

13 N/A

1E , kJ/mol 1.046×102 74 65

2E , kJ/mol 1.674×102 251 N/A

* Two competing reaction model

** Single overall reaction model

81

800

1000

1200

1400

1600

1800

2000

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

Distance from lance tip, m

Ga

s te

mp

era

ture, K

Injection rate: 5.3 kg/h

Blast temperature: 1243K

O Burgess et al. (1983)

Ubhayakar et al. (1977)

Takeda (1994)

Kobayachi (1976)

Figure 2.53 Schematic of the Blowpipe/Tuyere (combustion test section)

assembly of the pilot scale raceway hot model.

Figure 2.54 Gas temperature distributions for pulverised coal burning in a reactor

from experimental measurement and numerical predictions using different

devolatilisation models (Du, 2001).

First stage Second stage

82

Y2/Y

1

Figure 2.55 Comparison of the relationships between Y1 and Y2 in the literature (Shen

et al., 2008).

83

Table 2.4 Kinetics of single overall reaction and two competing reaction models used in

the PCI calculation models.

Authors

Kinetic parameters Stoichiometric

coefficients

A (1/s) E (kJ/mole)

Y1

Y2 A1 A2 E1 E2

Takeda (1994) 8.36×104 65 1.5 VM

He et al. (1986) 5105.1 73.13 1.3VM

Sato et al. (1996)

-1.11×103Cb

2+1.96×

105Cb－8.5510

6

Cb: carbon in coal,

wt%

-2.26×102 Cb

2+

4.48×104Cb－

2.04×106

(0.96×10-3

Tb+4.6

×10-2

Cb－3.9)VM

Tb: B. T., oC

Ubhayakar et al. (1976) 5107.3 131046.1 74 251 VM 2Y1

Kobayashi et al. (1977) 5102 71037.1 104.6 167.4 0.3 1

Burgess et al. (1983) 5107.3 131046.1 74 251 VM 2 Y1

Jamaluddin et al.

(1986) 5107.3 131046.1 74 251 VM 2 Y1

Aoki et al. (1993) 5107.3 131046.1 149.6 251 VM 2 Y1

Picard (2001) 1.34×105 131046.1 74 251 VM 1.7Y1

Guo, et al. (2005) 5107.3 131046.1 150 251 VM 1.4+

1.6VM

Du and Chen (2006) 5107.3 131046.1 74 251 VM 1.5Y1

Gu et al. (2008) 5107.3 131046.1 74 251 VM 2 Y1

Shen et al. (2008) 5107.3 131046.1 150 251 VM

1.25Y12

+

0.92Y1

84

2.3.2 Char Oxidation

2.3.2.1 Field approach (Field, 1969)

Once the volatile component of coal is completely evolved, heterogeneous surface

reactions begin to consume the char fraction of the coal. The reactions occurring in the

raceway can be listed below:

C + 1/2 O2 CO (2.8)

C + CO2 2CO (2.9)

C + H2O CO + H2 (2.10)

The char oxidation (Equation 2.8) is exothermic, while char gasification with CO2 and

H2O are endothermic. The time required for consumption of char particles in a turbulent

environment can range from 30 ms to over hour. It should be noted that the char

reactions with CO2 and H2O are about 5-6 orders of magnitude slower than that with O2

(Smoot and Smith, 1985). Due to very short residence time for injected coal particles

within the raceway, the char burnout contributed by the gasification reactions may be

negligible.

The particle temperature affects the apparent reaction rate by shifting the reaction

controlling mechanisms. Figure 2.56 represents the effect of temperature on the reaction

rate of a char particle. In low temperature range (regime I), the chemical reaction may

determine the overall rate. The reaction rate in middle temperature range is controlled

by both chemical reaction and diffusion (regime II). In regime III (high temperature),

diffusion of reactants and products in the boundary layer limits the reaction rate

(diffusion control), while the temperature has little effect. In the actual PCI combustion

operation, reactions of char particles occur under various temperatures and gas

atmospheres. Therefore, the reaction and diffusion should be considered for the

estimation of the global reaction of char particles (Ishii, 2000; Picard, 2001).

85

The corresponding equations for global reaction of char particles proposed by Field

(1969) are given below:

RD

RD

kk

kk

oxp

ppA

dt

dm 2.11

where 2

pp πDA is the surface area of the coal, and oxp is the partial pressure of

oxidant species in the gas surrounding the combusting particle. kD and kR are the

diffusion rate and the kinetic rate, respectively, expressed by

0.75

p

1

2D

C

gp

D

TTk 2.12

pR

N

gR TETk R expAR 2.13

In this research, the mass diffusion–limited rate constant is 12

1 105C , the kinetics

limited rate pre–exponential factor (AR) and activation energy (ER) are 3050

kg/(m2-s-atm) and 179.4 kJ/mol (Smith, 1982), respectively. Besides, the temperature

exponent N is taken as zero (Smoot and Pratt, 1979). Table 2.5 summarises some

kinetics of char oxidation by PCI combustion models.

Figure 2.56 Rate-controlling regimes for char reactions (Smoot and Smith, 1985).

86

Table 2.5 Kinetics of char oxidation employed in the PCI combustion models.

Authors Chemical reaction rate, kR

Jamaluddin et al, 1987 )R

exp( R

p

RRT

EAk

AR=309 kg/m2-s-atm 1200<Tp<1800 K

ER=156.5 kJ/ mol

AR=3.15 kg/m2-s-atm 1800<Tp<2100 K

ER=87 kJ/ mol

Nogami, et al. (1992) )R

exp( R

p

RRT

EAk

AR=87100 kg/m2-s-atm Tp≦1273 K

ER=149.7 kJ/ mol

AR=－4.9+3.85×10-3

Tp Tp>1273 K

ER=0

Takeda and Lockwood

et al., 1997

)R

exp( R

p

RRT

EAk

AR=87140 kg/m2-s-atm

ER=102 kJ/ mol

Haywood et al. (1994)

Du and Chen (2006)

)R

exp( R

p

RRT

EAk

AR=3050 kg/m2-s-atm

ER=179.4 kJ/ mol

Picard (2001) )R

exp( R

p

RRT

EAk

AR=0.007 kg/m2-s-atm

ER=90 kJ/ mol

Gu et al. (2008) )R

exp( R

p

RRT

EAk

AR=1.813×103 m/s

ER=108.9 kJ/ mol

Aoki et al (1993)

He et al. (1986)

Sato et al., 1996

Nogami, et al. (2004)

g

p

R RTT

k )80001

exp(7260

87

2.3.2.2 Gibb Model (1985)

In Gibb approach (Gibb, 1985; Guo et al., 2005), the oxidation mechanism of carbon

can be characterised by the parameter Φ so that oxides are produced according to the

equation:

ΦC + O2 → 2(Φ-1) CO + (2-Φ) CO2 (2.14)

The value of Φ is assumed to depend on the particle temperature TP:

)exp(2

)1(2

p

SS

T

TA

(2.15)

where the constants are given by Gibb as AS=2500 and TS=6240 K (Gibb, 1985). An

analytical solution of the oxygen diffusion equation leads to the following equation for

the rate of change in char mass (mc):

11

32

1

1 ))((1

3

kkkM

M

edt

dm

COxy

CSC

(2.16)

The far field oxygen concentration ρ∞ is taken to be the time-averaged value obtained

from the gas phase calculation, and ρc is the density of the char. Physically, k1 is the rate

of external diffusion, k2 is the surface reaction rate, and k3 represents the rate of internal

diffusion and surface reaction. These are defined as follows:

21

PR

Dk (2.17)

P

R

R

kek )1(2 (2.18)

)exp(P

CpRR

T

TTAk (2.19)

aTkk PR

2

3 /)1coth( (2.18)

88

5.0)(eaD

kR

P

C (2.19)

The parameters include void fraction (e), volume/internal surface area ratio, a, and

particle radius (Rp). D is the external diffusion coefficient of oxygen in the surrounding

gas, and kc is the carbon oxidation rate, defined by the modified Arrhenius equation.

Besides, Gibb recommends a value for Dp an order of magnitude less than D. As

indicated in Table 2.1, Gibb approach has been adopted for char oxidation by Guo et al.

(2005) and Shen et al. (2008, 2009a, 2009b, 2011).

2.3.3 Turbulence model

In the region of blowpipe-tuyere-raceway, the fluid motion is practically fast. The

standard k-ε model has been used in the PCI combustion models (Aoki et al., 1993;

Haywood et al., 1994; Picard, 2001; Nogami et al., 2004; Shen et al., 2008). Some

modified k-ε models were also considered in the calculations (Takeda and Lockwood,

1997; Guo et al., 2005; Gu et al., 2008). For a better treatment in the mixing and

dispersion of coal particles in comparison with the conventional k-ε model, the RNG k-ε

model (Biswas and Eswaran, 2002) is thus applied in this research to simulate the

turbulent combustion (Du and Chen, 2006; Du et al., 2007; Du et al., 2015). The

complete formulation of the RNG k–ε turbulence model is given as follows:

ρεGkμρUk kefft (2.20)

ρεCGCk

εεμρUε k1εefft

*2 (2.21)

where k is the turbulence kinetic energy, ε is the kinetic energy dissipation rate, and Gk

is the generation of turbulence kinetic energy due to the mean velocity gradients and

expressed by:

89

UUkUUG t

T

tk

I

3

2)( (2.22)

UUkUUG tT

tk

I

3

2)( (2.23)

The coefficient *2εC is given by

3

3

22012.01

)38.4/1(

CCC*

ε (2.24)

where η=Sk/ε, and S is the modulus of the mean rate of strain. The coefficients Cμ, C1ε,

C2ε, and αt are empirical constants, and their values derived empirically are 0.0845,

1.42, 1.68, and 1.393, respectively.

2.3.4 Gas combustion in turbulent flow field

2.3.4.1 Probability density function (PDF) of turbulence chemistry

In this approach, individual species transport equations are not solved. Instead, the

solution of a single conserved scalar transport equation, the mixture fraction, is solved.

The individual component concentrations are derived from the predicted mixture

fraction distribution (Eghlimi and Sahajwalla, 1997; Zhou, 1993). The effect of

turbulence on the gas combustion is modelled using the partial equilibrium chemistry

model. An assumption is made that the reaction is mixing limited (i.e. fast chemistry)

and the diffusion coefficients of all the species are equal. (Note that this assumption is

reasonable since the temperature in the combustion zone of blast furnaces is extremely

high and the effects of turbulent convection dominate those of molecular diffusion.)

Given this simplifying assumption, the species equations can be reduced to a conserved

scalar quantity known as the mixture fraction f . The mixture fraction can be written in

terms of the atomic mass fraction as

Oi,Fi,

Oi,i

XXXX

f

, where iX is the

elemental mass fraction for element i . The subscript O and F denote the value at the

oxidizer stream inlet and the fuel stream inlet, respectively. It is essentially a numerical

90

construct used to describe the degree of scalar mixing between the fuel and the oxidant.

In accordance with the mixture fraction concept, the transport equations are given by:

f

σ

μfρ

t

g t,

ggU (2.25)

Zk

ερCfμCZ

σ

μZρ

g

g

g2Z

2

g t,1Z

t

g t,

gg

U (2.26)

The last two terms of Equation 2.26 represent, respectively, the production of

concentration fluctuation due to non–uniformity of mixture fraction and destruction rate

of the fluctuations. Note that in the present simulations, the values of tσ , 1ZC and

2ZC are specified as 0.85, 2.86 and 2.0, respectively. Importantly, the shape of the

assumed probability distribution for a variation of the mixture fraction, fp , is

described by the β–function form that more closely represents experimentally observed

features. The shape produced by this function is given by the following functions of

mean mixture fraction f and concentration fluctuation Z :

dff1f

f1ffp

1B1-A

1B1-A

(2.27)

1

Z

f1f fA (2.28)

f

A f1B (2.29)

Thus, given prediction of f and Z at each point in the flow field (Equations 2.25 and

2.26), the assumed PDF shape can be computed and used as the weighting function to

determine averaged values of variables.

2.3.4.2 Eddy break up and eddy dissipation models

Gaseous combustion in turbulent flow field can be modelled by eddy dissipation

approach (Magnussen and Hjertager, 1976), in which modelling concept assumes that

91

chemical reactions are instantaneous when molecular–level mixing of reactants occurs.

In such cases, the overall rate of reactions is essentially controlled by turbulent mixing

rather than chemical kinetics. Note that this assumption is reasonably applicable in the

combustion zone of blast furnaces where temperature is extremely high and reaction

rates are fast compared to reactant mixing rates. In turbulent flows, the mixing time is

dominated by the eddy properties, and therefore the reaction rate is proportional to the

rate of turbulent mixing (g

g

kε

), i.e.

S

m,mmin

ερC ox

fu

g

gA

g

fuk

R (2.31)

where Rfu is reaction rate of volatile, AC is an empirical constant, mox and mfu are the

time-average mass fraction of oxygen and volatile in the gas phase and S is the

stoichiometric oxygen required for volatile combustion. The sensitivity analysis of CA,

ranged from 4.0 to 0.5, to the coal burnout was carried out by Takeda (1994). It is found

that CA did not have a strong influence on a burnout profile. This parameter was

recommended by Takeda (1994) to be 0.8 because it gave a marginally better fit to the

measured data reported by Suzuki et al. (1984).

When eddy break up model (Spalding, 1971) is adopted, the reaction rate of volatile can

be written as:

1/2

prod

g

gA mε

ρCg

fuk

R (2.32)

where mprod is variance of mass fraction of gas product.

In practice, both models have been widely used in PCI combustion models as shown in

Table 2.1.

2.3.5 Lagrangian approach

For a stable operation of blast furnace with high coal injection rates, accurate

92

information related to the dispersion of coal particles is required, because it is critical

not only to control the devolatilisation and combustion of injected coal particles in the

raceway, but also to protect tuyeres from failure caused by coal abrasion. As indicated

in the 2-D and 3-D PCI combustion models summarised in Table 2.1, the Eulerian

model was used for describing the behaviours of gas phase, while Lagrangian approach

(Kuo, 1986) was used in tracking behaviours of individual coal particle. With the

approach, the coal particles are treated as discrete objects, and their motion is calculated

as the burning coal particles move through the combustion flow field (Toporov, 2014).

In the Lagrangian approach, the trajectory of the particle is calculated by integrating the

force on the particle which can be written as:

where the subscript “ p ” denotes the particulate phase, m is coal particle mass, gμ is

the molecular viscosity, pD is the particle diameter, DC is the drag coefficient given

by Morsi and Alexander (1972), and pRe is the relative Reynolds number, which is

defined as g

prel.gp μ

DρRe

U . The gas velocity fluctuation ( '

gu ) are randomly

sampled by assuming that they obey a Gaussian probability density function, so that

2

gg uζu , 2

gg vζv , 2

gg wζw (2.34)

where ζ is a normally distributed random number, and the remainder of the right–hand

side is the local RMS value of the velocity fluctuations. These values of the RMS

fluctuating components are obtained from solving the turbulence kinetic energy

equation and defined (assuming isotropy) as 3

2kwvu

g2

g

2

g

2

g .

Integration in time of Equation 2.33 yields the velocity of the coal particle at each point

along the trajectory via:

pggpDpgD

ppuReCπDμ

8

1

dt

mdUU

U f (2.33)

93

pi U

dt

dx (2.35)

Equations 2.33 and 2.35 are solved in each coordinates (xi) to calculate the trajectories.

2.3.6 Summary

The sub models used in the CFD based models are comprehensively reviewed in section

2.3. Due to very short residence time for the injected coal particles in the regions of

blowpipe, tuyere and raceway, the coal burnout in the combustion region is mainly

contributed by the volatile evolution, followed by the char combustion. Therefore, in

this research, the performance of single overall reaction and two competing reaction

model are compared, while the kinetic parameters, as well as the stoichiometric

coefficients, are validated by the experimental data reported by Burgess et al. (1985).

The calculated results show that the kinetic parameters and stoichiometric coefficients

used in this research can significantly reflect the coal combustion behaviours in the high

temperature environment.

2.4 Raceway shape

2.4.1 Observation of raceway

A cold model experiment was performed by Inatani et al. (1976) with the aid of a

high-speed camera to investigate the dynamic behaviour of coke in the raceway. As can

be seen in Figure 2.57, the raceway was divided into five regions, from A to E

exhibiting different features of coke movements. Each of tick interval made on the coke

flowing trajectories represents 0.001 second. In the region of the raceway immediately

downstream of the tuyere nose (labelled as section A in Figure 2.57), the resident coke

particles were accelerated by the blast supplied from the tuyere, while new particles

were introduced from the area of the raceway just above the tuyere nose (exit). In

section B, the hot blast generated a recirculation structure which caused the particles

falling in front of the tuyere nose to be caught by the gas stream and carried backward

to impact on the raceway wall (section C). In section D, the coke particles simply

94

circulated around point P. Finally, in section E, the coke particles delivered by the gas

flow accumulated with a high packing density and form an almost impermeable zone

referred to as the “bird’s nest”.

Using an endoscope enclosed by a cooling jacket, studies of the raceway phenomena by

photographic observations parallel and vertical to the axis of tuyere have been carried

out by Greuel et al. (1974) in an operating blast furnace (hot model). The observations

revealed that a hollow (jetting) space was formed in front of the tuyere. As shown in

Figure 2.58, coke particles fell in the flow of the hot blast gas from the tuyere exit, and

were accelerated towards the centre of the furnace. Contrary to the observations in a

cold model reported by Inatani et al. (1976), the coke particles did not circulate in the

raceway. Notably, no coke recirculation within the raceway was also observed in the hot

model experiment of Nogami et al. (2004), which will be discussed in more detail latter.

Owing to short residence time of coke in the raceway, a complete combustion of coke in

the raceway could not be achieved. Instead, it was presumed that main reaction between

the blast oxygen and the coke carbon took place deeper in the furnace. Table 2.6 shows

the comparison of raceway observed by the coal model and the hot models.

Figure 2.57 Schematic illustration of raceway structure.

95

Table 2.6 Comparison of raceway observations in cold and hot models.

To analyse the coal combustion behaviours in the lower zone of blast furnace, a proper

raceway shape should be taken in the calculation model. The raceway shapes used in the

calculation models were determined by means of:

(1) observation (Kuwabara et al., 1981; Takeda and Lockwood, 1997);

(2) tuyere probing (Gu et al., 2008);

(3) DEM modelling (Nogami et al., 2004);

Characteristics

Cold model

by Inatani et al. (1976)

Hot models

by Greuel et al. (1974),

Nogami et al. (2004)

Shape Balloon-like with a jetting space Hollow space with the shape

of curved hose

Entrance of coke Above the tuyere exit Above the tuyere exit

Coke recirculating Yes No

Combustion space

Figure 2.58 Representation of the movement of coke through the raceway.

96

(4) Eulerian approach (Aoki et al., 1993); and

(5) unpublished research (Jamaluddin et al., 1986; Shen et al., 2011).

In early phase of model development in this study (Du et al., 2004), the raceway was

assumed to be a cylindrical jetting space according to the observation by an endoscope

(Greuel et al., 1974). To achieve a more accurate calculation, a mathematical model

based on Eulerian approach coupled with CFD has been developed to predict the

raceway configuration of blast furnace at CSC (Du, 2011).

2.4.2 CSC’s raceway shape prediction model

The raceway prediction model was developed using the Eulerian–Eulerian multi–fluid

model implemented in a commercial CFD code (Du, 2011). It is noted that this model is

based on the fundamental concept of interpenetrating continua for multiphase mixtures

(Gidaspow, 1994). In performing the simulations, the conservation equations (i.e. mass,

momentum, energy and species) are derived by averaging the local instantaneous

balances, and are solved for each individual phase. Moreover, the different phases

present within the same control volume at the same time are characterized by means of

phase volume fractions. In the model, the energy transfer between solid (coke particles)

and gas phases and the heterogeneous reactions within the furnace are taken into

consideration. Calculation approaches for gas and solid phases can be described as

follows:

Gas phase

i. The gas mixture is an incompressible ideal fluid.

ii. The gaseous turbulent combustion rate can be described by the eddy

dissipation model (Magnussen and Hjertager, 1976).

Solid phase (coke particles)

97

i. The coke has the form of granular particles with a smooth and spherical

morphology.

ii. The reaction rate of the coke combustion and gasification processes can be

calculated using the kinetic / diffusion–limited rate model.

iii. The diameter of respective coke particles is kept constant.

iv. The coke degradation mechanisms are ignored.

Furthermore, the considered chemical reactions are shown in Table 2.7. Heterogeneous

reactions of carbonaceous fuels are assumed to be of first–order irreversible with respect

to the oxygen and carbon dioxide. In the model, the conservation equations of mass,

momentum, energy and species are constructed for each phase. The equations for gas

and coke are described in Eulerian coordinate system. Moreover, the different phases

present within the same control volume at the same time are characterised by means of

phase volume fractions. The conservations equations for gas and solid phases can be

written as

he i,ggggg mραρα

tU

(2.36)

he i,sssss mραρα

tU

(2.37)

where the subscript g and s denotes the gas and solid phase; α , ρ , and U are the

volume fraction, density, and mean velocity, respectively. The terms on the right–hand

side of Equations 2.36 and 2.37 express the mass transfer between phases due to

heterogeneous reactions.

Momentum equation:

he i,ssggggg

ggggggg

mβραpα

ραραt

UUUgτ

UUU (2.38)

98

he i,ssgsssss

sssssss

mβραPpα

ραραt

UUUgτ

UUU

(2.39)

where p is the gas pressure, τ is the stress–strain tensor, g is the gravitational

acceleration, β is the interphase momentum transfer coefficient ,and sP is the solid

pressure, which arises due to the both of translation and collision for particles in the

solid phase. The last two terms on the right–hand side of Equations 2.38 and 2.39 are

the source terms due to interaction between phases.

The energy equations for both phases are:

o

ihe i,

o

iho i,gsgg

t

g t,

g p,g

ggggggg

ΔhmΔhωαQTPr

μCλ

hραhραt

U

(2.40)

gssssssssss QTλhραhραt

U

(2.41)

where h is the sensible enthalpy, λ is the thermal conductivity, T is the mean

temperature, pC is the specific heat, sggs QQ is the intensity of heat exchange

between the gas and solid phases, o

ih is the formation enthalpy per unit mass of

species i , and g t,μ is the gas phase turbulence viscosity and is described later. In the

simulations, the turbulent Prandtl number, tPr , is set to be 0.85. For Equation 2.40, the

last two terms on the right–hand side are the source terms due to chemical reactions.

The rate of energy transfer between phases sgQ is modelled according to the correlation

of Gunn (1978).

The species equation can be described below:

he i,ho i,gi g,

t

g t,

i g,ggi g,gggi g,gg mωαYSc

μDραYραYρα

t

U

(2.42)

99

where i g,Y and i g,D are the mass fraction and molecular diffusivity of species i in

the gas phase, respectively. tSc is the turbulent Schmidt number and is set to be 0.7.

The last two terms on the right–hand side of Equation 2.42 are net rates of production of

species i form homogeneous and heterogeneous reactions respectively. Since coke is the

only species of solid phase, the species equation of solid phase is reduced to equation

2.37.

Table 2.7 Chemical reactions in coke–packed furnace model.

Homogeneous reactions (gas–gas) Heterogeneous reactions (gas–solid)

CO ＋ 1/2 O2 → CO2 C ＋ 1/2 O2 → CO

C ＋ CO2 → 2 CO

2.4.2.1 Validation of raceway shape prediction model

To evaluate the performance of the model, the combustion experiment using the coke

bed rig developed by Nogami et al. (2004), as shown in Figure 2.49, is modelled. Table

2.8 indicates the hot blast gas conditions and properties of coke–packed bed in the

calculation. Figure 2.59 compares the numerical and measurement results given by

Nogami et al. (2004). As shown in Figure 2.59a, the predicted raceway shape, which is

featured by the voidage contour of 0.4 (Mondal et al., 2005), for all coke operation is

only slightly smaller than the observation, showing the model is performed acceptable

accuracy in the prediction of raceway shape. Figure 2.59b shows the variation of the gas

compositions along the central axis of the tuyere (i.e. A→B in Figure 2.59a). Note that

for each gas, the symbols and lines denote the experimental observations and the

numerical predictions, respectively. It is evident that the numerical results predicted by

the present model are concurred with the experimental results of Nogami et al. (2004).

100

Table 2.8 Operating conditions and properties of coke–packed bed.

Hot blast Coke

Temperature, K 1100 Diameter, mm 25－35

Flow rate, Nm3/h 710 Density, kg/m

3 1081

Oxygen content, % 22 Initial volume fraction, - 0.55－0.65

Shape factor, - 0.7

2.4.2.2 Prediction of raceway shape in an operating blast furnace

The model was applied to predict the raceway shape of CSC’s No3 BF, in which 32

tuyeres with inner diameter of 140 mm are introduced into the blast furnace at an

inclination angle of 5 degrees in the downward direction. Figure 2.60a shows the

geometry and dimensions of the furnace. In the calculation, only 1/32 sector of blast

Figure 2.59 Comparison of numerical and experimental results (Nogami et al., 2004)

of: (a) raceway shape; and (b) gas composition distribution along central axial of

tuyere (Du, 2011).

(a) (b)

101

furnace corresponding to a single tuyere of the furnace was considered (Figure 2.60b).

The typical operating parameters of the furnace for all coke operation were used as the

inlet boundary conditions (Table 2.9).

Table 2.9 Typical operating parameters of No3 BF for all coke operation.

Blast pressure, atm 4.5

Blast flow mass flow rate, kg/s 3.9

Blast temperature, K 1423

Blast oxygen content, % 21

Coke size, mm 30

Initial voidage 0.4

Figure 2.60 Profile of CSC’s No3 blast furnace: (a) main dimensions (unit: m); (b)

calculation domain of a single tuyere.

(a) (b)

102

Figure 2.61 shows the calculated voidage distribution under all coke operation

conditions of the furnace. In this research, the raceway shape is featured by the voidage

contour of 0.4. It is seen that the depth of the raceway is around 0.95 m. In addition, it is

observed that the raceway is composed of an inlet region with a relatively high voidage

and a transition zone characterized by a reduction in voidage toward the raceway

boundary. The movement of coke is much slower than that of coal particles and hot

blast gas within the raceway, therefore the cave could be thought of as a porous media.

As indicated in Figure 2.62, a simplified computational domain is proposed for the 3D

coal combustion in this work. Obviously the predicted raceway shape is similar with

that observed by Greuel et al. (1974).

a

b

Figure 2.61 Void fraction contours in combustion zone of 3D coke packed

furnace model: (a) top view; (b) side view.

103

2.4.3 Summary of raceway shape prediction

In the cold raceway experiment model of Inatani et al. (1976), the raceway was featured

as a balloon, and it was a hollow space with a shape of cured hose in the hot model

(Greuel et al., 1974; Nogami et al., 2004). Notably, recirculting coke particles are found

in the raceway of cold model. On the other hand, the coal particles in the raceway of the

hot model are blown towards the centre of the furnace.

In this research, the raceway shape and size is determined by Eulerian-Eulerian

multi–fluid model implemented in a commercial CFD code. The measured 2-D raceway

shape reported by Nogami et al. (2004) was applied for the validation of calculated

raceway shape. Based on the operation conditions of CSC’s No3 blast furnace, the

voidage distribution in the raceway are determined. The voidage contour of 0.4 is used

as the boundary of the raceway in this research.

2.5 Injection of biofuel into blast furnace

The utilisation of alternatives to coal in PCI can abate the consumption of fossil fuels

and, in the case of biomass, mitigate CO2 emissions. If biomass can be used in blast

Raceway depth

950 mm

Blowpipe--Tuyere

405 mm 795 mm

Figure 2.62 The simplified calculation domain. Note that αg is the volume fraction

of gas inside the raceway.

104

furnaces as an alternative fuel to coal, it is anticipated that CO2 emissions from the steel

industry and ironmaking can be lessened to a certain extent. However, application of

biomass as an industrial fuel is limited due to its high moisture, low bulk and energy

densities as well as hard grindability. The upgrade of biomass can be fulfilled via

torrefaction and carbonization or pyrolysis where biomass is thermally degraded in an

inert or oxygen-free environment. The torrefaction temperature is in the range of

200-300o C (Peng et al., 2013; Lu et al., 2012; Sabil et al., 2014), and carbonisation is

operated at temperatures of 300–500oC (Abdullah and Wu, 2009). As reported by

Babich et al. (2010), the injection of charcoal fines has been successfully practiced in

some small charcoal blast furnaces in Brazil with injection rates of 100 to 150 kg/tHM.

2.5.1 Combustion experiments and modelling

Combustion experiments of four charcoals and a high volatile coal using the

combustion rig shown in Figure 2.1 (Mathieson et al., 2005) have been carried out by

BSL (Mathieson et al, 2012). Table 2.10 and 2.11 show the key properties of the

samples and operating parameters of the tests respectively. Figure 2.63 provides a

summary of combustion performance of the coal (C-HVM) and the three hardwood

charcoals as a function of the VM of the samples with an air cooled coaxial lance and

interpolated to an O/C of 2.0. Also included are results for previously studied PCI coals

at the same conditions, and the trend of increasing burnout with VM that was

established. It is observed from Figure 2.63 that the burnout of the hardwood charcoal

samples was (a) a function of VM, (b) of similar slope to that for the PCI coals, and (c)

the trend line for the charcoals was located approximately 40% above the trend line for

the PCI coals. The higher performance in burnout given by the charcoals might be

resulted from higher specific areas by heat pretreatment and fragmentation in

combustion. This implies that superior combustion performance and therefore higher

injection rates can be expected for similar hardwood charcoals at relatively low VM

contents. As shown in Figure 2.64, the pressure drop measured across to the tuyere was

105

proportional to the VM of the injectant. This is expected to be related to the amount of

additional gas from the combustion of the volatile matter within the tuyere.

Table 2.10 Key properties of the bulk coal and charcoal samples.

Table 2.11 Key properties of the bulk coal and charcoal samples.

106

Torrefaction and burning characteristics of bamboo, oil palm, rice husk, bagasse, and

Madagascar almond were studied and compared with a high-volatile bituminous coal

using the drop tube furnace developed by Du et al. (2010) at CSC to evaluate the

Figure 2.63 Burnouts as a function of volatile matter of the injectants with an

air cooled lance and O/C = 2.0. Comparison is made with previous results for

PCI coals.

Figure 2.64 Differential pressure across the tuyere as a function of the volatile

matter of the injectants.

107

potential of biomass consumed in blast furnaces (Chen et al., 2012). The schematic of

the reaction system for testing combustibility of fuel is demonstrated in Figure 2.65.

The particle size of solid fuels was in the range of 74 to 149 μm, and the reaction

temperature was 1000oC. The higher heating value (HHV) ratios (enhancement factors)

for raw and torrefied biomasses and coal are listed in Table 2.12. From a calorific point

of view, Madagascar almond is the most sensitive biomass to torrefaction. Specifically,

with torrefaction temperatures of 250 and 300oC, its HHV was amplified by factors of

1.36 and 1.54, respectively. For the torrefaction temperature of 300oC, the calorific

values of the biomasses were close to that of coal, except for rice husk, which had

enhancement factors of around unity. This reveals that 300oC is a feasible operating

condition to transform the biomasses into solid fuels resembling high-volatile

bituminous coal. The burnout versus fuel ratio of the samples is demonstrated in Figure

2.66. The fuel ratio is defined as the content ratio of FC to VM. The fuel ratio was in the

range of 0.13–1.4 in the experiments. Once the biomasses underwent torrefaction, the

burnout tends to decay, whereas the fuel ratio shifted rightwards in the diagram. In

contrast to the experimental results (Figure 6.21) reported by Mathieson et al. (2012),

the burnout of the biofuels did not their exhibit superior performance on burnout as their

fuel ratios were higher than that of HV PCI coal. This might result from the differences

in gas flow patterns in both combustion systems.

108

Table 2.12 Enhancement factor of higher heating value.

Figure 2.65 Schematic of the reaction system (1) cylinder; (2) carrier gas; (3)

secondary gas; (4) rotameter; (5) hopper; (6) preheater; (7) lance; (8) DTF; (9)

thermocouple; (10) ceramic tube; (11) heater; (12) sampling probe; (13) cooling water;

(14)cyclone; (15) residual solid particles; (16) induced suction fan; (17) exhausted gas.

109

The combustion of Taiheyo coal (VM: 44.6%) and Oak char (VM: 27.1%) in a 2-D

furnace (Figure 2.67) was simulated and compared by Wijayanta et al. (2014). Table

2.13 shows the computational conditions. Figure 2.68 are the temperature distributions

for both fuels at 23 wt.% O2 (Figure 2.68a) and at 27 wt.% O2 (Figure 2.68b). The

calculated results showed that Taiheiyo coal achieved a higher temperature distribution

in the furnace than that achieved by Oak char. Obviously, the higher temperature caused

more volatile being released from Taiheiyo coal. In other words, the burnout of the fuels

is strongly related to the volatile contents. Besides, an increase in flame temperatures

for both fuels could be achieved when the oxygen level is increased from 23% to 27%.

Figure 2.66 Distributions of burnout versus fuel ratio of raw and torrefied

biomasses as well as a HV coal.

110

Table 2.13 Computational conditions for biofuel injection.

Figure 2.68 Temperature profiles at an injection rate of 36 (kg solid fuel) / (1000

Nm3 feed gas).

Figure 2.67 Geometry and computational domain used in numerical simulation.

111

2.5.2 Summary of the biofuel injection

(1) The burnout results provided by the turbulent combustion rig indicate the trend line

for the charcoals is located approximately 40% above the trend line for the PCI

coals. The superior performance in combustion may be resulted from higher specific

area of the biofuel by heat pre-treatment and fragmentation in combustion

(Mathieson et al., 2012). Upon inspection of the experimental results by the drop

tube furnace, this advantage is not reproduced by the biofuels tested in the drop tube

furnace (Chen et al., 2012). This may result from the differences in gas flow

patterns in both combustion systems.

(2) The pressure drop measured across to the tuyere is proportional to the VM of the

injectant. This is expected to be related to the amount of additional gas from the

combustion of the volatile matter within the tuyere.

(3) The promotion of torrefaction temperature results in an increase in the calorific

value of biofuel, whereas the burnout of biofuel is decreased.

2.6 Summary

The experimental and numerical studies on pulverised coal combustion in PCI operation

have been reviewed in this chapter. Under conditions simulating blast furnace

environments, the experiments were carried out using the combustion rigs with/without

coke bed to investigate the influence of operation factors, including coal properties, hot

blast gas conditions and injecting facilities, on the coal combustion behaviours.

Alternatively, the drop tube furnace provides a better view at the combustion properties

of coal when one attempts to evaluate the combustion efficiency of coal.

In the 1980s, 1-D mathematical models were established to study the coal combustion

behaviours in raceway. However, the most important phenomena, the dispersion of the

coal particles to hot blast gas were not efficiently simulated. By the 1990s, some studies

have been carried out to develop the 2-D coal combustion models, in which the coal

particle trajectory was determined by Lagrangian approach. In the models, the

112

calculation domain and lance arrangement were assumed to be axi-symmetric. Recently

3-D CFD based models have been intensively developed to deal with the complicated

configurations of lance arrangement and raceway shape. Through parametric studies,

many countermeasures for promoting the coal burnout and the performance of blast

furnace have been proposed.

Practically, improvement of permeability of the raceway is one of key factors to achieve

high productivity with high PCI rates. Notably, none of the previous models consider

the relation between the pressure loss (permeability resistance) and the coal combustion

within the raceway.

2.7 Methodology

To improve the coal combustion efficiency and stability of raceway operation, the

combustion characteristics in the regions of blowpipe, tuyere and raceway were

numerically and experimentally studied in this research. The development of the

calculation model begun with the model validation. Secondly, only blowpipe and tuyere

were considered as the calculation domain. In the third phase, a cylindrical space with

the same diameter of tuyere exit was used as the raceway in the calculstion. In the final

phase, the raceway shape was determined by the Eulerian-Eulerian multi-fluid model.

The factors which influence the pressure loss and coal burnout were especially studied.

In the coal combustion experiment, a drop tube furnace was established to evaluate the

combustion performance of PCI coal, coal blend and biofuel in an environment with

high heating rates (>104 K/s). The volatile release and particle formation characteristics

in the blast furnace were also studied. The experiment results of the drop tube furnace

have been applied by CSC’s blast furnace operators as guidance for the selection PCI

coals and operation conditions.

113

CHAPTER 3

NUMERICAL PREDICTION AND PRACTICAL IMPROVEMENT

OF PULVERIZED COAL COMBUSTION IN BLAST FUTRNACE

A CFD base coal combustion model in the regions of blowpipe and tuyere is developed

and validated in this chapter. The performance of single lance and double lance

injection are compared. Earlier ignition is found with the operation of double lance in

comparison with the single one. Accordingly, the double lance injection became the

standard operation of CSC’s blast furnace in 2002.


pulverized coal combustion in blast furnace, International Communications in Heat and

Mass Transfer, vol. 33, p. 327-334.

114

ABSTRACT

The burning characteristics of pulverized coal in blowpipe and tuyere at two different

injection patterns are simulated numerically, to aid improving the practical performance

of blast furnace. With the condition of the same fuel and oxidant mass flow rates, the

predictions indicate that the combustion efficiency of pulverized coal using

double-lance can be substantially enhanced compared with that using single lance.

Accordingly, the pulverized coal injection in a practical blast furnace was modified

from single lance to double-lance. As a result, the practical injection rate of the

pulverized coal in the blast furnace was increased from 110 kg/tHM to 153 kg/tHM,

revealing that a profound decrease in operating cost of the blast furnace has been

implemented.

Keywords: Blast furnace; Pulverized coal; Blowpipe and tuyere; Combustion; Injection.

115

Nomenclature

pA Coal particle surface

pC Specific heat of coal particle

E Activation energy

h Convective heat transfer coefficient

k Reaction rate constant or turbulent kinetic energy

pm Coal particle weight

t Time

pT Temperature of coal particle

T Gas temperature

xi Spatially coordinate

21 , YY Mass fractions of emitted volatile at low and high temperatures

Greek Symbols

ε Dissipation of turbulent kinetic energy

p Emissivity of coal particle

μ Viscosity

σ Stefan-Boltzmann constant (=5.67×10-8

W/m2.K

4)

Subscripts

f Fuel

o Oxidizer

p Coal particle

∞ Gas phase

116

3.1 Introduction

It is known that coal plays an important role in energy developments and industrial

applications. For example, since the industrial revolution occurred in eighteenth century

(Boubel, et al., 1994), a considerable amount of coal has been utilized for power

generation in steam engines. In the last decade of 19th

century, pulverized coal (PC)

began to be used in cement industry for heating drying kilns (Singer, 1984). Nowadays,

pulverized coal is widely applied in coal-fired power plants for producing electricity

(Sami et al., 2001; Hinrichs and Kleinbach, 2002); it is also extensively used in

metallurgical industry for refining metals. As far as the process of ironmaking is

concerned, conventionally, coke, the product of high-temperature pyrolysis of coal,

serves as an important reactant in reducing iron ores into steels in blast furnaces

(Vamvuka et al., 1996). However, because of higher price of coke compared with

pulverized coal, the technique of pulverized coal injection (PCI) from tuyere has been

developed for several years to partially replace the consumption of coke. In other words,

the operating cost of the blast furnace can be substantially reduced if the injection rate

of PC is promoted significantly.

The pulverized coal can be used as auxiliary fuel in a blast furnace and possesses the

merit of reducing operating cost. Nevertheless, it should be addressed that, if coal

particles in combustion zones undergo incomplete combustion, the unburned or residual

char will accumulate in the blast furnace in which the char is depleted by means of

reaction with slag and carbon dioxide (Takahashi et al., 2002; Iwanaga, 1991). If the

accumulation rate of the char in the furnace is larger than the depletion rate, the

movement of hot blast will be retarded. This results in a pressure fluctuation which

further suppresses the operation of the blast furnace. In consequence, enhancing the

burning rate of PC and reducing the accumulation ratio of unburned char is one of

available methods to stabilize the performance of the blast furnace.

Because the operation of PCI is highly relevant to the performance of the blast furnace,

the purpose of the present study is to predict the combustion characteristics of

117

pulverized coal in a blast furnace through numerical simulation. By varying the

injection pattern of PC, its impact on the burning behaviors of the PC in the blast

furnace will be evaluated. Furthermore, based on the obtained results, a practical

strategy in improving PC combustion will be adopted.

3.2 Mathematical Formulation

3.2.1 Burning process of pulverized coal

A schematic diagram of the internal structure of a blast furnace is demonstrated in

Figure 3.1. Attention of the present study is focused on the pulverized coal combustion

in the regions of blowpipe and tuyere in the blast furnace. As shown in the Figure, when

coal particles are injected into the blowpipe, they will immediately immerse in a

high-temperature environment filled with hot blast and thereby experience rapid

heating, devolatilization reaction of the coal, oxidization of the volatile matters with hot

blast, combustion of residual unburned char, and gasification of the char. Recognizing

the above characteristics, it is known that the devolatilization reaction initiates coal

combustion, implying that the selection of parameters to model the devolatilization

reaction is of the utmost importance in predicting the PC combustion. Therefore, in the

current study, the initial chemical reaction of coal particles will be tested and verified to

ensure the validation of the numerical method.

3.2.2 Momentum and energy balance of a coal particle

Considering a moving coal particle, when it is assumed to be spherical and the

Lagrangian framework is used, the trajectory of the particle can be obtained by solving

a single particle momentum equation. That is, the rate of change of momentum is equal

to external forces on the particle. On account of very small coal particles investigated, it

is proper to neglect body force and only drag force is considered during computation.

Consequently, the equation of motion of the particle is expressed as:

118

p

p

p Fdt

vdm (3.1)

In regard to the energy balance, with conceiving the coal particle as a lump system,

the heating of the particle is carried out by convection and radiation; thus the

temperature of the particle can be described by the energy equation as the

following:

)()(44

ppppp

p

pp TTATThAdt

dTCm (3.2)

3.2.3 Model of devolatilization of coal particle

When one is concerned with the devolatilization process of coal particles, it depends

strongly on the heating rate, reaction time, and coal grade, and so forth. In fact, as the

heating process is fast, the volatile matters emitted from the coal is larger than the

analyzed result of ASTM, rendering that Q factor is larger than one. To describe the

devolatilization process more realistically, two-competing devolatilization model

Figure 3.1 A schematic diagram of internal structure in a blast furnace.

119

(Smoot and Smith, 1985) is employed. The two parallel and competing reactions are

given as follows:

) ( )1( 1111 etemperaturlowVolatileYCharYcoal

k (3.3)

) ( )1( 2222 etemperaturhigh VolatileYCharYcoal

k (3.4)

Furthermore, the reaction kinetics is written by:

CoalYkYkdt

dV )( 2211 ; )/exp( 111 pRTEAk and )/exp( 222 pRTEAk (3.5)

where V and R are mass fraction of volatile matter and universal gas constant,

respectively. In examining the preceding model, it is apparent that the parameters Y1, k1,

Y2, k2, E1, and E2 have a vital influence in predicting the devolatilization process. The

appropriate values will be suggested later.

3.2.4 Turbulent combustion model

In the gas phase the fluid motion is fast, the k model is thus applied to simulate the

turbulent combustion. In the operation of PCI, following the release of volatile matters

from coal particles, oxygen will encompass the volatile, yielding the diffusion flame

combustion. In such a situation, mass fraction probability density function (PDF) model

(Kobayashi et al., 1977) is an appropriate method to approach the reaction phenomena.

The model is established based on the concepts of mixture fraction, mix-is-burnt

(Eghlimi and Sahajwalla, 1997; Zhou, 1993), and probability density function. For a

system just having two reactants, consisting of fuel and oxidant, the PC combustion can

be approximated by a single-step reaction as:

ProductkgiOxidantkgiFuelkg )1( 1 (3.6)

The coefficient i represents the stoichiometric balance between the fuel and oxidant.

When the turbulent transport coefficients of reactant and oxidant in the flow field are

summed to be equivalent, employing the Zeldovich transformation the combined mass

fraction X can be obtained as the following:

iMMX of / (3.7)

120

and the mixture fraction f is defined by:

0

0

XX

XXf

f

(3.8)

where Ofof XXMM and , , , stand for mass fractions of fuel and oxidant as well as

combined mass fractions on the fuel and oxidant sides, respectively. The mixture

fraction f is a conservative scalar, and its value at a control volume can be calculated via

the solution of its instantaneous conservation equation for f (time-averaged):

m

it

t

i

i

i

Sx

f

xfu

xf

t

)()()(

(3.9)

In the above equation, mit Sx and , , , designate density (g/cm3), dynamics viscosity

(N.s/m2), spatial coordinate, and source term stemming from the reaction of coal into

the gas phase, respectively. Meanwhile, t is a computational parameter whose value

is given by 0.9 (Zhou, 1993). In the framework of PDF, mean square value of

concentration fluctuation g can be calculated through the following equation

gk

Cx

gC

x

g

xgu

xg

tdtg

it

t

i

i

i

2

)()()( (3.10)

where gC and dC are the computational parameters and they are given by 2.8 and

2.0, respectively. According to mixture fraction f, molar fraction of each gas species,

density, and temperature in control volume can be calculated.

3.3 Results and discussion

3.3.1 Numerical validation and parameter selection

Previous to simulating the physical phenomena, accurate selection of the parameters in

the devolatilization model has to be carried out. To achieve this goal, the presently

predicted results are compared with the experimental data of Burgess et al. (1983) to

confirm the validation of the simulation. The investigated conditions are summarized in

Table 3.1. In the meantime, two sets of parameters, reported by Kobayashi et al. (1977)

and Ubhayakar et al. (1976), are tested for comparison each other. Details of the

121

parameters are given in Table 3.2 and the predicted temperature distributions in the

blowpipe and tuyere are displayed in Figure 3.2. It depicts that the temperature is

spatially uniform when using the parameters of Kobayashi et al. (1977), implying that

the devolatilization reaction in the reactor is not exhibited. Clearly, the preceding result

is inconsistent with the experimental measurement. Regarding the parameters of

Ubhayakar et al. (1976), as shown in Fugure 3.2, the predicted temperature distribution

is close to the experimental data. It follows that the proposed parameters of Ubhayakar

et al. (1976) is capable of providing a more realistic prediction. Because of this, their

parameters are employed in the current study.

In examining the calculated temperature distribution, it is noteworthy that the curve is

characterized by monotonic increase with increasing distance away from the lance exit.

The profile can be partitioned into two stages; it composes of rapid rise in the upstream

region and progressive increase in the downstream one. The behavior in the first region

arises from the PC devolatilization or pyrolysis reaction followed by the combustion of

the emitted volatile matters with oxygen. In regard to the second region, the slow

increase in temperature is attributed to the reaction between char and oxygen. In a word,

after the PC is blown into the blowpipe, the chemical reaction is primarily achieved by

the gas-phase combustion (i.e., homogeneous reaction) and then implemented by the

solid-phase oxidation (i.e., heterogeneous reaction).

Table 3.1 Operational conditions selected by Burgess et al. (1983)

Reactor diameter, mm 50

Hot blast temperature, K 1243

Hot blast velocity, m/s 68

Coal particle diameter, μm 40

Volatile matter of PC (db), % 35.9

PC injection rate, kg/h 5.3

122

Table 3.2 Two sets of parameters used for predicting PC devolatilization.

Kobayashi et al. (1977) Ubhayakar et al. (1976)

1Y 0.3 VM

2Y 1 1.5×Y1

A1, 1/s 200000 3.7×105

A2, 1/s 1.3×107 1.46×10

13

1E , kJ/mol 1.046×102 74

2E , kJ/mol 1.674×102 251

Distance (m)

Ga

ste

mp

era

ture

(K)

0 0.2 0.4 0.6 0.8 1800

1000

1200

1400

1600

1800

2000

Kobayashi et al.

Ubhayakar et al.

Burgess et al.

2nd stage1st stage

Figure 3.2 A comparison of gas temperature distribution among experimental

measurement and two devolatilization models.

123

3.3.2 Impact of injection pattern

As mentioned previously, the emphasis of the present work is upon the burning

characteristics of pulverized coal in blowpipe and tuyere. To evaluate the behaviors of

PC combustion with different injection patterns, a typical running condition of the blast

furnace at China Steel Corporation (CSC), as shown in Table 3.3, is simulated.

Meanwhile, the physical geometries of the blowpipe and tuyere are illustrated in

Figure3.3. When attention is placed on the influence of injection pattern upon the PC

combustion, two different cases, consisting of single lance and double-lance, are

calculated where the mass flow rates of PC and carrier gas are fixed. Accordingly, the

diameters of the lances in the cases of single lance and double-lance are 20 and 14 mm,

respectively. To provide a reference for indicating combustion efficiency, the burning

ratio of PC is defined as:

%100(%) i

e

M

Mratioburning (3.11)

where Me and Mi are the PC weight-loss at the exit of tuyere and the original PC weight

at the entrance of blowpipe, respectively. The calculations suggest that, once the single

lance is modified to the double-lance, the burning ratio is substantially promoted from

4.9% to 12.2%. To proceed farther into the recognition of the burning mechanisms,

Figure 3.4 displays the isothermal contours in the blowpipe and tuyere in accordance

with the performances of the single lance and the double-lance. In the both cases,

because the PC and carrier gas are at room temperature prior to entering the blowpipe,

the temperatures in the vicinity of the entrance are relatively lower, as observed. When

comparing the isothermal contours in the downstream, it can be found that the ignition

of the latter case occurs earlier than that of the former. This obviously reflects that the

operation of the double-lance can facilitate the mixing between the PC and hot blast,

whereby the production rate of unburned char adjacent to the exit of the tuyere is

reduced.

124

Table 3.3 Operating conditions of PCI at CSC.

Hot blast conditions

Temperature: 1423 K; Pressure: 4.5 atm;

Mass-flow-rate: 3.9 kg/s; Oxygen content: 21 %.

Properties of PC

FC: 55.09%; VM: 35.13%; Ash: 6.23%; Moisture:

3.55%.

Particle distribution of

PC

90μm: 5%; 63μm: 25%; 45μm: 55%; 20μm: 15%.

Others Lance angle: 15o; Lance internal diameter: 20mm;

Carrier gas mass-flow-rate: 0.026 kg/s; PC injection

rate: 0.4 kg/s; Heat loss of tuyere: 900,000 W/m2.

180 mm

450 mm

140 mm

100 mm

Lance exit Tuyere Blowpipe

Figure 3.3 A schematic diagram of blowpipe and tuyere as well as their sizes.

125

3.3.3 Practical improvement of blast furnace

The aforementioned results have provided a practical insight into the performance of the

blast furnace. Based on the simulations, the injection pattern in one of the blast furnaces

in CSC has been the redesigned through changing the single lance to the double-lance.

After that, the injection rate of PC has been promoted to a great extent, from 110

kg/tHM (ton of hot metal) to 153 kg/tHM. For this reason, the goal of reducing the

operating cost of the blast furnace has been accomplished sufficiently.

PC

Blowpipe Tuyere

(a)

(b)

Figure 3.4 Isothermal contours in blowpipe and tuyere under the operations of (a)

single lance and (b) double-lance injections.

126

3.4 Conclusions

By utilizing two different injection patterns of pulverized coal, the burning

characteristics of the pulverized coal in the blast furnace have been examined. The

numerical simulations elucidate that the performance of PCI by means of double-lance

is capable of providing a superior burning, in contrast to the original single lance

design. This is attributed to the fact that the double-lance injection is conducive to

mixing between pulverized coal and hot blast, resulting in earlier ignition of the fuel.

The practical injection pattern of the PC in the blast furnace was modified from the

double-lance to the single lance, in accordance with the foregoing numerical

predictions. As a result, the injection rate of the PC has been amplified by a factor of

40%, from 110 kg/tHM to 153 kg/tHM. In summary, the numerical study has provided a

useful insight into the practical improvement of the blast furnace performance.

127

CHAPTER 4

PERFORMANCES OF PULVERIZED COAL INJECTION IN

BLOWPIPE AND TUYERE AT VARIOUS OPERATIONAL

CONDITIONS

The factors affecting the coal combustion are numerically studied in this chapter. The

calculated results provide useful insights for the assessment of blast, tuyere and PCI

operation conditions assisting improvement of coal burnout.

Du, S. W., Chen, W. H.and Lucas, A. J. (2007), Performances of pulverized coal

injection in blowpipe and tuyere at various operational conditions, Energy Conversion


128

ABSTRACT

Combustion efficiencies of pulverized coal in blowpipes and tuyeres under various

operational conditions are numerically predicted to recognize the performance of

pulverized coal injection in a blast furnace. A variety of parameters including injection

pattern of pulverized coal, oxygen content in hot blast, inlet temperature of the hot blast,

and mass flow rate of coal carrier gas are taken into consideration. The effect of

installing a ceramic sleeve around the tuyere on the pulverized coal combustion is also

evaluated. The predictions indicate that pulverized coal combustion is highly related to

the injection pattern, hot blast temperature, mass flow rate of the carrier gas, and

installation of ceramic sleeve, whereas it is insensitive to the oxygen concentration. The

present study is carried out based on the practical operational conditions of the blast

furnace at the China Steel Corporation. Consequently, the obtained results have

provided a useful insight into the operation of pulverized coal injection for improving

the blast furnace performance in the future.

Keywords: Blowpipe and tuyere; Blast furnace; Pulverized coal; Injection; Combustion

129

Nomenclature

pA Coal particle surface

C Coal

pC Specific heat of coal particle

E Activation energy

f Mixture fraction

F Fuel

h Convective heat transfer coefficient

k Reaction rate constant or turbulent kinetic energy

pm Coal particle weight

M Mass fraction

O Oxidant

R Universal gas constant

S Source term

t Time

pT Temperature of coal particle

T Gas temperature

v Mass fraction of solid lost as volatiles

V Volatile

X Combined mass fraction


21 , YY Stoichiometric coefficients of emitted volatiles at low and high

temperatures

Greek Symbols

130

ε Dissipation of turbulent kinetic energy


μ Viscosity


W/m2.K

4)

ρ Density

Subscripts

f Fuel

o Oxidizer

p Coal particle

∞ Gas phase

131

4.1 Introduction

Pulverized coal (PC) has become an important auxiliary fuel in the iron and steel

industry since the technique of pulverized coal injection (PCI) was developed for

ironmaking (Babich et al., 1996). When pulverized coal is injected into blast furnaces

through blowpipes and tuyeres, because of the reactions of devolatilization, gasification,

and combustion as well as the formation of unburned char, the coal becomes sources of

heat and reductant in raceways (Ohno et al., 1994). For this reason, PC is extensively

employed in blast furnaces to partially replace the metallurgical coke at the present time

(Babich et al., 1996; Ohno et al., 1994; Chung and Hur, 1997). In fact, utilizing PCI

also possesses the advantages of reducing operation costs of blast furnaces and

decreasing emissions of carbon dioxide. This arises from the fact that the price of coal is

relatively lower than that of coke and, from the viewpoint of energy conversion and

management, the PCI is more efficient than metallurgical coke. Accordingly, under the

situation of stabilizing blast furnaces, how to extend the PC injection rate for increasing

yield in the ironmaking process has become a prime concern for blast furnace engineers

(Babich et al., 1996; Ariyama et al., 1994).

When one attempts to recognize the combustion situations of PC in a blast furnace in

order to evaluate the performances of the furnace at various operating conditions, it is

practically difficult by virtue of intrinsic high-temperature environment and close

system of the furnace. Also the practical analysis will expose one to a high risk

environment. Over the past several decades, on account of rapid progress in computer

simulating capability as well as advanced development in numerical algorithm,

computational fluid dynamics (CFD) has become a powerful tool to aid understanding a

variety of scientific and industrial phenomena. For instance, Chen et al. (2000a; 2000b)

developed a numerical code with multi solid progress variables to simulate coal

gasification in an air blown entrained flow gasifier. Their simulation illustrated that

carbon conversion was independent of devolatilization rate, but sensitive to the

chemical kinetics of heterogeneous reactions on char surfaces, and less sensitive to a

132

change in coal particle size. On the other hand, Choi et al. (2001) numerically predicted

coal gasification in an entrained flow gasifier with slurry feed. By varying O2/coal ratio,

their developed models in predicting carbon conversions and syngas concentrations

agreed with the measured results. Aside from the difficulty of practical analysis and the

safety issue mentioned above, other obvious merits of CFD include that the numerical

predictions can be achieved in a short time and their analyses are much more economic

than the practical measurements.

In order to increase the performance of blast furnaces, according to the realistic

operation of a blast furnace at China Steel Corporation (CSC) the preset study is

intended to investigate PC combustion in blowpipes and tuyeres by means of a

numerical simulation. A variety of operating parameters, composed of injection pattern

of PC, oxygen content and inlet temperature of hot blast, as well as mass flow rate of

coal carrier gas, will be taken into consideration to account for their impacts on the

burning behaviors of PC. The PC combustion in the blast furnace and the

presence/absence of installing ceramic sleeve is also evaluated. From these predictions,

the obtained results will be adopted as potential countermeasures to enhance the

productivity of the blast furnace in the future.

4.2 Methodology

Schematic diagram of the internal structure of the investigated blast furnace is

demonstrated in Figure 4.1. Attention of the present study will be focused on the

pulverized coal combustion in the regions of blowpipes and tuyeres. Physically, the

phenomena involve fluid dynamics and reactions of the gas phase and solid particles.

The relevant governing equations are stated below.

133

4.2.1 Gas-phase continuity and momentum equations

In the gas phase, with the assumptions of steady-state flow and Newtonian fluid, the

continuity and momentum equations are:

m

pi

i

Sux

(4.1)

and

u

p

j

i

ji

ji

i

Sx

u

xx

puu

x

(4.2)

where the subscript p shows sources to the corresponding property due to the presence

of particulate phase. For the accurate and efficient predictions of the turbulent mixing

and dispersion of injected pulverized coal into the hot blast (>150 m/s), the RNG

Figure 4.1 A schematic diagram of internal structure in blast furnace.

134

(Re-Normalization Group) k model is adopted to predict turbulent combustion

(Biswas and Eswaran, 2002). This is because that the RNG k model can provide a

better treatment in the mixing and dispersion of coal particles in comparison with the

conventional k model.

Considering the operation of PCI, when coal particles are heated, volatile matters will

be released to react with oxygen, resulting in diffusion flame combustion (Khalil, 1982).

To approach the gaseous combustion, it is proper to assume that the chemical reactions

are fast compared to fluid mixing rate (mixed-is-burned); hence the mixture fraction

probability density function (PDF) model (Sivathanu and Faeth, 1990) is employed. In

the mixture fracture PDF frame, individual species transport equations are not

considered. Instead, the mixture fraction transport equation is solved. The mixture

fraction, f, can be written in terms of the mass fraction as

iOiF

iOi

XX

XXf

(4.3)

where Xi represents the mass fraction for some element i and the subscripts F and O

respectively stand for the values at the fuel and oxidant sides. The mixture fraction f is a

conserved scalar and its value in a control volume can be calculated from the solution of

its time-averaged ( f ) instantaneous conservation equation:

p

it

t

i

i

i

Sx

f

xfu

xf

t

(4.4)

In the above equation, pt S and designate dynamics viscosity and source term,

stemming from the reactions of coal in the gas phase, respectively. Meanwhile, t is a

computational parameter whose value is given by 0.9 (Jones and Whitelaw, 1982). In

the framework of PDF, mean square value of concentration fluctuation g can be

calculated through the following equation:

gk

Cx

gC

x

g

xgu

xg

tdtg

it

t

i

i

i

2

(4.5)

where gC and dC are the computational parameters and they are given by 2.8 and

2.0, respectively (Jones and Whitelaw, 1982). According to the mixture fraction f, molar

135

fraction of each gas species, density, and temperature in every control volume can be

obtained.

4.2.2 Coal particle momentum and energy equations

When coal particles are injected into a blowpipe via injection lance, the pathway of

individual moving particle can be tracked by solving a single particle momentum

equation. It is known that a particle’s motion is subjected to relative velocity between

the gas phase and solid phase. If the coal particle is assumed to be spherical and its body

force is neglected, resulting from investigating very small particle, the equation of

motion of the particle can be expressed by means of Lagrangian framework as the

following:

ppDp

p

p uuuuCddt

dum

8

1 2 (4.6)

In regard to the energy balance, if we postulate that the coal particle is a lump system

(Myers, 1971), the heat conducted into the whole particle is equivalent to the convective

and radiative heat transfer onto the particle surface; thus the temperature of the particle

can be described by:

44

ppppp

p

pp TTATThAdt

dTCm (4.7)

In general, the residence time of coal particles in combustion zone is about 20 ms

(Steiler, et al., 1996) where the environmental temperature is very high. The result is

that the particles experience rapid heating. Specifically, the heating rate of coal particles

is commonly in the order of 105 K/s. Following the rapid heating, hot blast surrounding

the particles will trigger a sequence of physical and chemical reactions. The reactions

include devolatilization of the coal, combustion of volatiles and unburned char, as well

as gasification of the char (Steiler et al., 1996). It has been illustrated (Smoot and Smith,

1985) that significant devolatilization will be started when temperature is as high as

about 650K. Since the characteristic times of coal devolatilization and volatiles

combustion are much shorter than that of unburned char combustion and gasification

136

(Smoot and Smith, 1985), the devolatilization process becomes dominant mechanism in

the initial period of injection. To describe the devolatilization process, the following

parallel, first order, irreversible reactions (Kobayashi et al., 1977) are employed:

) ( )1( 11111 etemperaturlowVYSYC

k (4.8)

) ( )1( 22222 etemperaturhigh VYSYC

k (4.9)

where C, Y, S, and V denotes coal, stoichiometric coefficient, char, and volatile,

respectively. The relative importance of the two equations is mainly determined by

temperature. In other words, as long as the temperature is low, the reaction is dominated

by Equation 4.8. Alternatively, it is governed by Equation 4.9 once the temperature is

relatively high. Accordingly, the reaction kinetics is written to:

CYkYkdt

dv 2211 (4.10)



In the aforementioned equations, v is the mass fraction of volatiles, and the reactions are

characterized by 21 EE . The framework of the CFD code and overall computational

procedures of the gas phase and the solid phase are illustrated in Figure 4.2. In

examining the preceding model, it is apparent that the parameters Y1, k1, Y2, k2 , E1, and

E2 have a vital influence in predicting the devolatilization process (Du and Chen, 2006).

In the present study, two sets of parameters respectively recommended by Kobayashi et

al. (1977) (Model 1) and Ubhayakar et al. (1977) (Model 2) are applied to predict PC

combustion and they are compared with the experimental data of Matheson, et al.

(2005). The predicted results under the situations of low and high PC injection rates are

plotted in Figures 4.3a and 4.3b, respectively. The operational conditions and the

properties of the coal are also included in Figure 4.3a. Obviously, the model 2 is

capable of providing accurate predictions in the current phenomena when compared to

137

the model 1. Consequently, the model 2 is employed in the study. The profiles of the

model 2 shown in Figure 4.3 also suggest that the PC combustion can be simulated well

in the developed numerical method.

Gas phase

˙ Mass conservation


˙ Turbulence model



model

Solid Phase

˙ Devolatilization model

˙ Char combustion

˙ Reaction-rate controlling

regime


(particle trajectory)


Initialization

˙ Geometry (meshing)

˙ Boundary conditions

˙ Sub-models



model

Solution

Radiative Transport

Particle

properties

Radiative

energy

exchange

Figure 4.2 Framework of the CFD code and computational procedure of the gas

phase and solid (coal particle) phase.

138


In general, pulverized coal combustion in blowpipes and tuyeres is likely to be

influenced when certain important operating parameters are altered. The present work is

aiming to provide a reference for improving the performance of PCI. To recognize the

burning behaviors of coal particles under various operational parameters, a baseline

case, in accordance with the typical running conditions of a practical blast furnace in

Figure 4.3 Gas temperature distributions for pulverize coal burning in a reactor

from experimental measurement and numerical predictions.

139

CSC is chosen. The operation conditions (Table 4.1) of the case are used as the

boundary conditions in the baseline calculation. Detailed physical geometries of the

investigated blowpipe and tuyere are sketched in Figure 4.1 as well. In the following

discussion, moving dynamics for coal particles in blowpipes and tuyeres as well as the

influences of some possibly important operating parameters on the particles burning

will be examined.

Table 4.1 Operating conditions (base case) of PCI at CSC.

Hot blast conditions Temperature: 1423 K;

Pressure: 4.5 atm;

Mass flow rate: 3.9 kg/s;

Oxygen content: 21 %.

Proximate analysis of

pulverized coal

FC: 55.09%;

VM: 35.13%;

Ash: 6.23%;

Moisture: 3.55%.

Particle distribution of

pulverized coal

90μm: 5%;

63μm: 25%;

45μm: 55%;

20μm: 15%.

Other conditions Lance angle: 15o;

Lance internal diameter: 20mm;

Carrier gas mass flow rate: 0.026 kg/s;

PC injection rate: 0.4 kg/s;

Heat loss of tuyere: 900,000 W/m2.

140

Figure 4.4 Trajectories and residence times of coal particles under the operation of

the base case.

4.3.1 Trajectories and residence times of coal particles

Figure 4.4 demonstrates the trajectories and residence times of coal particles under the

operation of the base case. Upon inspection of the trajectories, it is evident that the

mixing between coal particles and hot blast is insufficient in the blowpipe and the

upstream of the tuyere. This will delay the ignition of the fuel and result in poorer

reactions. Figure 4.4 also reveals that the residence times of the particles in the

blowpipe and tuyere range from about 4 to 7 ms.

4.3.2 Injection pattern

The preceding observation suggests that if the mixing between the solid phase and the

gas phase can be intensified to a certain extent, it will be conducive to heating,

141

devolatilization, and combustion of the PC. At present, three different methods,

including injection through single lance (base case), single lance with larger diameter,

and double-lance are calculated and compared with each other. They are denoted by

cases 1, 2, and 3, sequentially. Since the mass flow rates of the PC and carrier gas are

controlled to be the same in all the cases, the diameters of the lances corresponding to

the three cases are 20, 25.4, and 14 mm, respectively. On the other hand, to figure out

the performance of the PCI, a parameter of burning ratio (BR) is evaluated which is

defined as the ratio between the PC weight-loss at the exit of tuyere to the original PC

weight at the entrance of blowpipe. As seen in Figure 4.5, once the diameter of the

single lance is enlarged from 20 mm (case 1) to 25.4 mm (case 2), the BR will be

increased from 4.9% to 8.4%. It is inferred that the enhancement of the PC

combustion is due to the decrease in the velocities of the particles, thereby elongating

their residence times. In addition to increasing residence time, physically, if the

originally single lance is divided into double-lance, it will facilitate particles dispersion

and thereby enlarge the mixing between fuel and oxidant. For this reason, as shown in

Figure 4.5, when the single lance injection is modified into the double-lance injection

(case 3), the BR is further promoted to 12.2%.This argument can be verified by

examining the practical combustion situations of PC in cases 1 and 3 which are

displayed in Figure 4.6. From the Figures, it is obvious that coal particles extending

outward away from the centerline of the tuyere in case 3 is much more pronounced than

that in case 1. As a whole, the relative values of the BR in the three cares are 1:1.7:2.5,

revealing that an appropriate design in lance arrangement is capable of sufficiently

increasing the performance of PCI.

142

Figure 4.5 Burning ratios of PC at various injection patterns.

(a) (b)

Figure 4.6 Combustion situations of pulverized coal in (a) case 1 and (b) case 3.

143

hot blast temperature (K)

bu

rnin

gra

tio

(%)

1350 1400 1450 1500 1550 16000

2

4

6

8

10

12

4.9%(base case)

6.2%

9.3%

mass flow rate of carrier gas (kg/s)

bu

rnin

gra

tio

(%)

0 0.01 0.02 0.03 0.04 0.05 0.060

2

4

6

8

10

7.2%

2.5%

4.9%(base case)

oxygen content (%)

bu

rnin

gra

tio

(%)

21 23 25

4.6

4.8

5

5.2

5.4

4.9%(base case)

5.1% 5.1%

4.3.3 Oxygen concentration and hot blast temperature

Intuitively, pulverized coal combustion in blowpipes and tuyeres will be intensified if

oxygen concentration in hot blast is increased. To understand the role played by oxygen

concentration upon the PC combustion, the burning ratios of PC at three different

oxygen concentrations, consisting of 21% (base case), 23%, and 25%, are simulated and

demonstrated in Figure 4.7. The Figure depicts that the BR is increased a bit as the O2

concentration is enriched from 21% to 23%. Specifically, only 0.2% of increment in the

BR is developed. Once the O2 concentration is further enlarged to 25%, it is noteworthy

that the BR ceases increasing. The reason causing this feature is that the characteristic

times for particles traveling in the tuyere are very short, as illustrated in Figure 4.4,

whereas time required for unburned char to react with oxygen is much longer.

Consequently, coal devolatilization becomes the dominant reaction in the tuyere and the

reaction is insensitive to the oxygen concentration. It follows that the enriched oxygen

almost plays no part in the enhancement of PC combustion.

Figure 4.7 Burning ratios of pulverized coal at various oxygen concentrations.

144

4.3.4 Hot blast temperature

Considering the effect of hot blast temperature, unlike the oxygen concentration, Figure

4.8 reveals that increasing temperature has a remarkable effect on the BR. In other

words, corresponding to the hot blast temperatures of 1423K (base case), 1473K, and

1523K, the values of BR are 4.9%, 6.2%, and 9.3%, respectively. This arises from the

fact that the utilized devolatilization model is principally determined by particle

temperature, which is closely related to the hot blast temperature. Therefore, the higher

the hot blast temperature, the more effective the coal reactions, as predicted.

4.3.5 Mass flow rate of carrier gas

For the mass flow rate of carrier gas, Figure 4.9 shows the burning ratios of the PC

at the mass flow rates of 0.015 kg/s, 0.026 kg/s (base case), and 0.05 kg/s,

respectively. In contrast to the base case, when the mass flow rate is decreased to

0.015 kg/s, the BR grows greatly, from 4.9% to 7.2%. Alternatively, when the

mass flow rate is increased to 0.05 kg/s, the BR declines to 2.5%. This elucidates,

in short, that the BR rises markedly as long as the mass flow rate of the carrier gas

is decreased. This is the result that the inlet carrier gas is in the state of room

mass flow rate of carrier gas (kg/s)

bu

rnin

gra

tio

(%)

0 0.01 0.02 0.03 0.04 0.05 0.060

2

4

6

8

10

7.2%

2.5%

4.9%(base case)

oxygen content (%)

bu

rnin

gra

tio

(%)

21 23 25

4.6

4.8

5

5.2

5.4

4.9%(base case)

5.1% 5.1%

hot blast temperature (K)

bu

rnin

gra

tio

(%)

1350 1400 1450 1500 1550 16000

2

4

6

8

10

12

4.9%(base case)

6.2%

9.3%

Figure 4.8 Burning ratios of pulverized coal at various hot blast temperatures.

145

temperature. As soon as the gas is blown into the blowpipe, contrary to the

function of the hot blast, the cooling effect stemming from the carrier gas will be

exhibited, rendering the coal reactions being suppressed. For PC in CSC, its

transport pertains to dilute phase transportation (DPT). In order to avoid PC choke

in pipes, the blown amount of coal particles is generally controlled below 18 kg per

kg of carrier gas. Therefore, despite the advantage of decreasing the mass flow rate

of the carrier gas observed above, it should be emphasized that we are still unable

to noticeably reduce the mass flow rate. Furthermore, according to the present

simulation and under the limitation mentioned above, how to promote the

efficiency of the pulverized coal injection via decreasing the mass flow rate of the

carrier gas will become an important and practical issue in the near future.

4.3.6 Installation of ceramic sleeve

Apart from the operations, redesigning equipment is another achievable method to

improve the performance of PCI. In the practical operation of the blast furnace, cooling

water is commonly used in the tuyere to lessen the damage caused by high-speed PC

erosion. Recently, a new method of installing ceramic sleeve around tuyere was

Figure 4.9 Burning ratios of pulverized coal at various mass flow rates of carrier gas.

146

developed in CSC to diminish the damage. Because the ceramic sleeve possesses an

adiabatic feature, one is able to assume that no heat is transferred through the wall of the

sleeve and this assumption is adopted in the simulation. As a result, the BR is promoted

to 7.7% when compared to the base case of 4.9% which is not adiabatic situation. In

consequence, the installation of the ceramic sleeve can simultaneously protect the tuyere

and promote PC burning while isolating the energy.

4.4 Conclusions

Proper operations of PCI are highly related to fuel consumption and energy

management in blast furnaces. By selecting various operating conditions in a blast

furnace, consisting of injection pattern, oxygen concentration and inlet temperature of

hot blast, mass flow rate of carrier gas, and construction of ceramic sleeve, the burning

characteristics of the pulverized coal in blowpipe and tuyere in the blast furnace has

been extensively examined. When tracking the trajectories of coal particles, it is

observed that the mixing of the coal particles and hot blast plays a crucial role in

heating, devolatilization, and combustion of the PC. Considering the injection form, the

performance of PCI by means of single lance with larger diameter or double-lance is

conducive to providing a superior burning, in contrast to the original single lance

design. This results from coal particles characterized by longer residence times and

better mixing between the fuel and oxidant. The improvement in PC burning is

especially significant for the double-lance injection. Besides, the simulations also

indicate that either increasing hot blast temperature or decreasing the mass flow rate of

carrier gas enable us to promote the burning ratio of the PC in the tuyere. However, in

view of the dominant reaction of the devolatilization in the tuyere, the combustion

efficiency of PC is hardly affected at all when the concentration of oxygen is enriched.

With regard to the design of device, the numerical prediction reveals that the installation

of adiabatic ceramic sleeve can reduce the heat loss of hot blast effectively, whereby the

147

PC burning ratio is enhanced greatly. The present study has provided a number of

practical insights into the improvement of blast furnace performance.

148

CHAPTER 5

PRACTICE OF HIGH PRODUCTIVITY AT NO3 BLAST FURNACE

OF CHINA STEEL CORPORATION

The coal blend (mixture of low and high volatile coals) combustion within a simplified

raceway is analysed through CFD in this chapter. The pressure loss due to the coal

combustion in the raceway can be abated when the coal blend is practiced. It was

confirmed by the plant trials at CSC’s No3 blast furnace. Consequently, the coal blend

injection has become standard practice at CSC’s blast furnaces since 2003.

Du, S. W., Yeh, C. M., Yang, M. K. and Ho, C. K. (2004), Practice of high productivity

at No.3 blast furnace of China Steel Corporation, Proceedings of AISTech Conference,


149

ABSTRACT

To meet ongoing rise in the market demand for steel products, China Steel has carried

out series of studies for promoting its productivity since 2000. No3 blast furnace was

exemplified in this paper. Some countermeasures to improve the stability of blast

furnace operation have been practiced as follows: (1) decreasing the slag volume for the

reduction of energy consumption, (2) taking proper charging patterns based on the

burden trajectory measurement to establish terrace in the vicinity of wall, and (3)

adopting one-bit drilling to stabilize the tapping process. As a result, the hot metal

production was gradually increased from 7469 (2000) to 8167 t/d (2002). To promote

the performance of the blast furnace further, a calculation model has been developed to

investigate the coal combustion behaviours in the raceway. The calculation results

indicated that the pressure resistance (pressure loss) in the raceway could be reduced

with the decrease of the volatile content of the PCI coal. With this advantage, the flow

rate of hot blast air may be increased for the increase of the hot metal production. This

was successfully confirmed by the plant trials of coal blend (adding low volatile coal

into the high one) injection in the furnace in 2002. As a result, the PCI operation at CSC

was shifted from high volatile coal injection to coal blend injection in 2003. The coal

blend injection was of key importance in upraising the hot metal production of the

furnace from 8167 t/d in 2002 to 8322 t/d in 2003, even reaching its record high 8536

t/d in September 2003.

Key words: Blast furnace; High productivity; Combustion model; Coal blend injection

150

5.1 Introduction

China Steel Corporation (CSC) is the only integrated steel producer in Taiwan. It

operates four blast furnaces, No1 to 4, having inner volumes of 2434, 2850, 3400 and

3400 m3 respectively. N

o3 blast furnace completed its first campaign on 18

th October

1999 (from November 1987) achieving an average productivity of 1.98 t/d-m3 (inner

volume base), and was again blow-in on 15th

January 2000. Table 5.1 shows the main

features of the furnace. The recovery of steel market has continued worldwide,

especially in Asia, since 2002. Therefore, to promote the productivity has become a

major challenge for No 3 blast furnace in its second campaign.

This paper describes the research works and countermeasures taken for raising the

productivity of No3 blast furnace.

Table 5.1 Main features of CSC’s No3 blast furnace.

5.2 Development of low flux sinter

The energy required for blast furnace process is mainly supplied from sensible heat of

the hot blast air and the combustion heat produced by reactions between the oxygen and

Inner volume, m3

3400

Working volume, m3 2850

Hearth diameter, m 12.5

Number of tapholes 4

Number of tuyeres 32

Top pressure (kPa) 2.3

Blast temperature, ℃ 1150-1180

Charging equipment Bell-less

Cooling system Stave cooler

151

fuels within the raceway. Therefore, the production rate of blast furnace is mainly

determined by blast volume supplied, and can be expressed as the quotient below:

Production Rate (t/h) = (Hot blast supplied through tuyere, m3/h)/(Specific hot blast

required for producing one tonne of hot metal, m3/tHM) (5.1)

Equation 5.1 suggests the production rate can be increased with the decrease of the

energy consumption for generating hot metal. Table 5.2 shows the typical energy

consumption of No3 blast furnace in 1999 (late period of the first campaign). It was

found the energy consumed by slag was significant in the process. Therefore, the

reduction of the slag volume might be one of effective countermeasures for energy

saving and for raising the productivity of the blast furnace. In CSC, the blast furnace

slag is mostly from sinter which accounts for more than 70% of the charged ferrous

burden and is fluxed by adding serpentine, silicon sand and limestone in the sinter plant.

Theoretically, the slag volume can be lowered as the fluxes addition is decreased.

Concerning negative effects on the properties of sinter due to decreasing the fluxes, an

investigation into low fluxes sintering has been carried out (Hsieh et al., 2002). The

experimental results given by sinter pot tests showed that under the conditions of

reducing SiO2 content (from 5.6% to 4.5%), increasing basicity (CaO/SiO2) slightly and

reducing MgO content slightly in sintering, the negative effect was limited and

negligible. According to the results, SiO2 content of sinter was reduced gradually, and

the slag volume was consequently lowered as shown in Table 5.3.

152

Table 5.2 Typical energy consumption in the late period of the first campaign.

Table 5.3 Reduction of SiO2 in sinter and slag volume.

1999 2000 2001 2002 2003

SiO2 content in sinter，% 5.60 5.04 4.84 4.74 4.72

Slag volume，kg/tHM 295 275 265 264 265

5.3 Establishment of burden terrace

The blast furnace has been operated with V shape burden profile since its first

campaign. Owing to unstable movement of the burden resulting in its collapse,

especially in the high productivity operation, the blast furnace has experienced sharply

increase in wall heat loss. To obtain stable movement of the descending burden, a new

charging pattern is required to establish a terrace of the burden profile in the vicinity of

the wall.

×103kcal/tHM %

Heat of Solution Loss 320.5 27.0

Reduction Heat by H2 12.0 1.0

Reduction Heat of Non-Ferrous Metal 12.5 1.1

Decomposition Heat of Moisture 106.8 9.0

Sensible Heat of Top Gas 73.5 6.3

Sensible Heat of Hot Metal 318.5 26.9

Sensible Heat of Slag 156.2 13.2

Heat Loss to Cooling Water 49.3 4.2

Heat Loss 134.3 11.3

Total 1186.6 100

153

Thus an acoustic sensor system (Figure 5.1) was developed to measure the falling points

of the discharged burdens from the rotating chute (Ho, 2000), and the trajectories of the

burdens were also modelled. From the measured falling points and the calculated

trajectories of the burdens, it was found the charged burdens hit the wall heavily and

rolled down towards the centre of the furnace; consequently the V shape burden profile

was formed. Obviously, this was caused by the charging pattern employed in the first

campaign, in which the rotating chute was located at an angle of 49o to the centre of the

furnace in the beginning of charging. A new charging pattern was developed and tested.

The main change of the new pattern was to reduce the angle from 49o to 46

o. As shown

in Figure 5.2, after taking that pattern into practice, the burden profile was shifted from

V shape to M shape, and a long terrace (1.2-1.5m) in the vicinity of wall was

successfully formed. Consequently, the wall heat loss was decreased (Table 5.4) due to

stable peripheral gas flow.

Above Burden Probe

Bar

Chute

AEo

Oscillator

PC

OreCoke

AEi

Pipe

N2

Li Lo

Above Burden Probe

Bar

Chute

AEo

Oscillator

PC

OreCoke

AEi

Pipe

N2

Li Lo

Figure 5.1 AE sensor system for measuring burden falling point.

154

Table 5.4 Variation of wall heat loss.

Wall heat loss, M cal/h

2000 8418

2001 8313

2002 7539

2003 7708

5.4 Development of one bit drilling method

At CSC, soaking bar tapping has been employed since 1984. In the soaking bar

operation, the percussion bar soaked was pulled out by the drilling machine when the

taphole was plugged by mud. However, it was found the oxygen lancing process was

always needed in the final stage of opening. Consequently, the soaking bar tapping

failed to match the gradually increased productivity of the furnace, and fluctuation of

blast pressure occurred due to high liquid level in the hearth. To solve these problems, a

(a 修改前)

前

焦炭層

爐壁

爐中心區域

燒結礦層

Coke TerraceCoke Terrace

(b after)

Coke Layer

Wall

Ferrous Layer

(a before)

(a) (b)

Figure 5.2 Burden profile before (a)/ after (b) changing charging pattern.

155

simplified tapping method called as one bit drilling was developed at CSC for replacing

the soaking bar tapping (Wu and Hsieh, 2003).

For drilling through the taphole in one bit, the diameter of the percussion bar was

increased from 40-45 mm to 42-48 mm, and the pressure of cooling nitrogen was

increased from 6 to 14 atm for better cooling. From the operation results as shown in

Table 5.5, it was found the new drilling practice has become more efficient resulting in

less consumption of the oxygen lance and percussion bar. Additionally, the tapping

frequency has been reduced due to longer life of the taphole.

Table 5.5 Comparison between soaking bar tapping and one bit drilling.

5. 5 Coal blend injection

5.5.1 Analysis of permeability of the furnace

To stably increase the productivity, it is essential to maintain a good permeability of the

furnace. At CSC, the permeability resistance is defined as:

1000B

ΔPk

v

t (5.2)

where

tΔP is the pressure drop between the tuyere exit and the top of the furnace, atm

Bv is the blast volume, Nm3/min

Soaking bar tapping One-bit tapping

Oxygen lance consumed, piece/month 800 207

Percussion bar consumed, piece/ month 498 197

Average number of tapping, -/day 9.3 8.8

156

Equation 5.2 indicates that the permeability can be improved by decreasing the pressure

drop across the furnace. Figure 5.3 shows the pressure distribution of the furnace across

the furnace. It is clear from the Figure, the maximum pressure gradient happens in the

lower zone of the blast furnace, especially in the area of the raceway, where pulverised

coal is injected into and combusted with oxygen. In other words, the pressure drop in

the lower zone of the furnace is influenced by the pulverised coal injection (PCI)

operation, and the pressure gradient in the lower zone might be abated by changing PCI

operation parameters.

At CSC, it was presumed the permeability of the furnace might be affected if the

accumulation rate of unburnt char exceeded its consumption rate in the furnace.

Therefore, high volatile coals have been injected solely into the furnace since it

commenced PCI operation. On the other hand, the advantages of injecting low volatile

coal has been reported below (Willmers, 1989):

(1) Coke replacement ratio increases due to higher calorific value of the low volatile

coal;

(2) The blast momentum decreases with decreasing volatile content of coal injected.

2.2

2.4

2.6

2.8

3

3.2

3.4

3.6

3.8

4

0 5 10 15 20 25 30 35 40 45

Vertical Distance from Tuyere Exit, m

Press

ure, a

tm

Raceway Area

Figure 5.3 Typical pressure distribution of No 3 blast furnace.

157

It suggested the performance of the furnace may be improved by blending low volatile

coal into high volatile coal. Therefore, a further study was needed to compare the

performance of high volatile coal injecting and blend coal injecting.

5.5.2 Coal combustion model within tuyere-raceway area

A calculation model based on a CFD code has been developed to analyse the coal

combustion behaviour within a tuyere-raceway area of blast furnace. The raceway was

assumed to comprise a cylindrical jetting space which had the same diameter (14cm) as

the tuyere exit. Figure 5.4 shows the physical geometry of the tuyere-raceway area. The

burning history of pulverised coal injected can be expressed as the following process

(Takeda, 1994):

(1) rapid heating;

(2) devolatilisation;

(3) oxidization of volatile matters released with hot blast;

(4) combustion of residual unburnt char; and

(5) char gasification.

Figure 5.4 Physical geometry of combustion region.

158

Having examined the aforementioned processes, it was illustrated that the residence

time of coal particle is about 20 ms (Steiler et al., 1996), while the second and third

reactions are implemented within 100 ms. As regards the char combustion, its

characteristic time is in the order of one second. The time required for completing the

last reaction (i.e., the char gasification) is even longer (Smoot and Smith, 1985).

Recognizing these characteristic times, it is known that the devolatilisation reaction

initiates the coal combustion, implying that the selection of parameters in modelling the

devolatilisation reaction is of the utmost importance. To describe the coal

devolatilisation process more realistically, two-competing devolatilisation model(8)

was

employed. The two parallel and competing reactions are given as follows:

Coal 1k

(1-Y1) Char1 + Y1Volatile (high temperature) (5.3)

Coal 2k

(1-Y2) Char2 + Y2Volatile (low temperature) (5.4)

Furthermore, the reaction kinetics can be written as:

Coal )YkY(kdt

dV2211 (5.5)

)/RTEexp( Ak p111 (5.6)

)/RTEexp( Ak p222 (5.7)

In examining the preceding model, it is apparent that the parameters of A1, A2, Y1, Y2,

E1, and E2 have a vital influence in predicting the devolatilisation process. According to

the validation of the model, the parameters suggested by Ubhayakar et al. (1976) were

adopted in the calculation (Table 5.6). In the gas phase, because the average blast

velocity faster than 160 m/s, k-εmodel was thus applied to simulate the turbulent

combustion. In the operation of PCI, following the release of volatile matters from the

coal particles, the oxygen will encompass the volatile, yielding the diffusion flame

combustion. In such a situation, mass fraction probability density function (PDF) model

(Zhou, 1993) is an appropriate method to approach the reaction phenomena. For a

159

system just with two reactants, consisting of fuel and oxidant, the combustion can be

approximated by a single-step reaction as:

Product kg i)(1Oxidant kg i Fuel kg 1 (5.8)

where the coefficient i represents the stoichiometric balance between the fuel and

oxidant. When the turbulent transport coefficients of reactant and oxidant in the flow

field are assumed to be equivalent, by employing the Zeldovich transformation, the

combined mass fraction can be obtained as the following:

iMMX of / (5.9)

where Mf and Mo respond to the mass fractions of fuel and oxidant, respectively.

The mixture fraction f is further defined by

f≡( X-Xo )/( Xf-Xo ) (5.10)

where Xo and Xf express values of X in the fuel side and oxygen side respectively.

Because f is a conservative scalar, its time-average, f , instantaneous conservation

equation in a control volume can be written as:

Sx

f

xfu

xf

t it

t

i

i

i

)()()(

(5.11)

where σt is a computational parameter and it is given as 0.9. On the other hand, in the

framework of PDF, the mean square value of concentration fluctuation, g, can be

calculated through the following equation

160

gk

Cx

gC

x

g

xgu

xg

td

itg

it

t

ii

i

2)()()()( (5.12)

where Cg and Cd are the computational parameters and they are given as 2.8 and 2.0,

respectively. The operation parameters used in the base case calculation is listed in

Table 5.7.

Table 5.6 Parameters of devolatilisation kinetics.

Table 5.7 PCI Operation condition used in the calculation.

Y1, - VM analyzed value

Y2,- 1.5×Y1

A1, 1/s 3.7×105

A2, 1/s 1.46×1015

E1, kJ/mol 74

E2, kJ/mol 251

Hot blast conditions Properties of PC Other operating conditions

Temperature: 1423K

Pressure: 4.5atm

Mass flow rate: 3.9kg/s

O2 content: 21%

FC：55.09%

VM：35.13%

Ash：6.23%

Moisture：3.55%

Particle size

distribution:

90μm：5%

63μm：25%

45μm：55%

20μm：15%

Lance angle: 15o

Lance inner diameter: 20mm

Carrier gas flow rate:0.026 kg/s

Injection rate of PC: 0.4kg/s

Heat loss of tuyere: 900000W/m2

161

5.5.3 Calculation results and discussion

In the base case calculation, the combustion efficiency of the high volatile coal reached

81.5%, and it was 63% when the volatile content of the coal was reduced to 25% (coal

blend case). Figure 5.5 shows the trajectories and residence times of coal particles under

the operation of the base case within the reaction area. It is evident that the mixing of

the coal particles and hot blast is poor, resulting in insufficient oxygen within the coal

plume as shown in Figure 5.6, where the released volatile may turn into soot, which is

not favourable in the blast furnace ironmaking process.

As can be seen in Figure 5.7, the smoother pressure distribution and less pressure loss

along the combustion region area is found with the coal blend operation in comparison

with the base case. This indicates that a decrease in volatile content of coal injected can

effectively abate the pressure loss due to less volatile released to the gas and moderate

gas volume expansion in the combustion area. Although the coal blend injection

generate more unburnt char than injecting high volatile coal, the performance of the

furnace can be improved as long as the consumption rate of the unburnt char exceeds its

accumulation in the furnace. The calculation results encouraged the use of coal blend

injection to improve the performance of the furnaces of CSC.

PC

C

鼓風嘴

Jet Zone

Coal Plum

Jet Zone

Tuyere

Residence Time, ms

Figure 5.5 Trajectories and residence time of coal particles in the combustion

region.

Coal plume

Jet zone

Tuyere

162

5.5.4 Plant trial of coal blend injection

A series of coal blend trials were carried out in the end of 2002, with the low volatile

coal in coal blend increasing gradually from 30% to 50%. Table 5.8 compares the

changes of pressure drop (resistance) in the lower zone of the furnace and permeability

before and after injecting coal blend. The operation results indicate the pressure loss in

the lower zone of the furnace was abated, and permeability of the furnace was improved

Oxygen Concentration, -

Low Oxygen Zone



Low Oxygen Zone

Figure 5.6 Oxygen concentration contour at cross section along combustion

region.

3.4

3.5

3.6

3.7

3.8

3.9

4

4.1

0 0.2 0.4 0.6 0.8 1 1.2 1.4

Horizotal Distance from Tuyere Exit, m

Pres

sure

, a

tm

VM=35%

VM=25%

Raceway

Tuyere Exit

Figure 5.7 Pressure distribution along combustion region from lance exit.

Distance from lance exit, m

163

when the low volatile coal was added. Having verified of the improvement of the

operation, the coal blend injection now becomes standard operation of CSC’s blast

furnaces. With the advantage in lower pressure resistance, the increase of PCI rate or

hot blast flow rate for higher productivity can be achieved.

Table 5.8 Changes of pressure drop and permeability for coal blend injection.

100% HV coal 30% C coal 40% C coal 30% F coal

Pressure drop between

blast and P1, atm 0.68 0.46 0.53 0.58

Permeability, - 0.286 0.267 0.259 0.257

* P1 is located 2.08 meters above tuyere level.

** C and F are low volatile coals with volatile content of 12 and 12.5% respectively.

5.6 Increase of hot metal production in No3 blast furnace

Table 5.9 shows the performance of the furnace on the hot metal production. Generally,

the increase of hot metal production from 2000 to 2002 was mainly contributed by the

countermeasures mentioned in sections of 5.2 to 5.4. On the other hand, the coal blend

injection was of key importance in upraising the productivity from 8167 t/d (2.4 t/m3-d)

in 2002 to 8322 t/d (2.45 t/m3-d) in 2003, even reaching its record high 8536 t/d (2.51

t/m3-d) in September 2003.

Table 5.9 Hot metal Production in CSC’s No 3 blast furnace.

2000 2001 2002 2003

Average blast air flow rate, Nm3/min 5335 5440 5493 5579

Average daily production rate, t/d 7469 7912 8167 8330

Productivity，t/d-m3

2.20 2.33 2.40 2.45

164

5.7 Conclusions

In the period of 2000 to 2003, the operation of CSC’s No3 blast furnace was improved

by means of (1) decreasing the slag volume for the reduction of energy consumption of

the furnace, (2) taking proper charging patterns based on the burden trajectory

measurement to establish terrace in the vicinity of wall, and (3) adopting one-bit drilling

to stabilize the tapping process. As a result, the hot metal production was gradually

increased from 7469 (2000) to 8167 t/d (2002). As revealed in the coal combustion

model in the raceway and plant trials, the pressure resistance in the lower zone of blast

furnace can be abated when low volatile coal is added to the high one in the PCI

operation. With this advantage, the flow rate of hot blast air can be increased for

upraising the hot metal production. In 2003, the PCI operation at CSC’s blast furnaces

was shifted from high volatile injection to coal blend injection. With the advantage of

the coal blend injection, the hot metal productivity was promoted from 8167 (2002) to

8330 t/d (2003), even reaching its record high 8536 t/d (2.51 t/m3-d) in September 2003.

165

CHAPTER 6

BURNING CHARACTERISTICS OF PULVERIZED COAL WITHIN

BLAST FURNACE RACEWAY AT VARIOUS INJECTION

OPERATIONS AND WAYS OF OXYGEN ENRICHMENT

This chapter evaluates the performance of injection lances associated with different

ways of oxygen enrichment in terms of pressure loss and coal burnout within the

regions of blow pipe, tuyere and raceway, which is featured by a voidage contour of 0.4.



operations and ways of oxygen enrichment, Fuel, vol. 143, p. 98-106

166

ABSTRACT

In this research, coal combustion behavior across the regions of blowpipe, tuyere, and

raceway of blast furnace are numerically examined. Three different lance

configurations, including a single lance, a double air-cooled coaxial lance, and an

oxy-coal lance with different oxygen enrichment patterns, are taken into consideration.

The coal combustion efficiency by the double lance injection is 5.1% higher than that

by single lance injection. From the calculated temperature by the oxy-coal lance, coal

ignition is retarded due to the cooling effect of enriched oxygen flowing through the

lance annulus, resulting in the moderation of pressure loss within the raceway. Most

importantly, the enriched oxygen becomes the combustion enhancer in the downstream

of coal plume after ignition is triggered. Consequently, the coal burnout under the

oxy-coal lance injection is comparative to that under the double air-cooled lance

injection. The performance of blast furnace may be improved with the advantages

provided by the oxy-coal lance injection. Compared with the single lance injection, coal

trajectories under the oxy-coal lance injection are closer to the tuyere exit due to the

higher inertia force of coal particles against hot blast. This should be taken into account

for the designs of the oxy-coal lance.

Keywords: Blast furnace; Pulverized coal injection; Oxy-coal lance injection;

Combustion efficiency; Pressure loss; Numerical simulation.

167

Nomenclature

pA Coal particle surface (m2)

C Coal

C0 Inertial loss coefficient (m-1

)

pC Specific heat of coal particle (J kg-1

K-1

)

Cμ, C1ε, C2ε Empirical constants for turbulence model

Dc Coke diameter in the coke bed (m)

Dp Coal diameter (m)

E Activation energy (kJ mol-1

)

f Mixture fraction

fD Drag force from a particle (N)

F Fuel

Gk Generation of turbulence kinetic energy (kg m-1

s-3

)

H Total enthalpy (J kg-1

)

h Convective heat transfer coefficient (J kg-1

)

k Turbulent kinetic energy (m2 s

-2)

k1, k2 Devolatilization rate constant (s-1

)

M Mass fraction

pm Coal particle weight (kg)

O Oxidant

Qreac Reaction heat (J kg-1

)

q Heat transfer from a particle (W)

Prt Turbulent Prandtl number

R Universal gas constant (kJ mole-1

K-1

)

S Source term (kg m-2

s-2

)

S1, S2 Char

t Time (s-1

)

pT Temperature of coal particle (K)

T Gas temperature (K)

U Mean velocity (m s-1

)

v Mass fraction of solid lost as volatiles

V1, V2 Volatile released at low and high temperatures

168

X Combined mass fraction


21 , YY Mass fractions of emitted volatile at low and high temperatures

Greek Symbols

α Porosity

αt Empirical constants for turbulence model

ε Dissipation of turbulent kinetic energy (m2 s

-2)


λ Thermal conductivity (W m-1

K-1

)

μ Viscosity (kg m-1

s-1

)

μeff Effect viscosity of gas (kg m-1

s-1

)

μt Turbulent viscosity (kg m-1

s-1

)


W m-2

K-4

)

σt Prandtl number of turbulence kinetic energy

ρ Density (kg m-3

)

Subscripts

f Fuel

o Oxidizer

p Coal particle

g Gas phase

169

6.1 Introduction

A number of new ironmaking processes have been developed over the last several

decades; however, blast furnaces are still the most important and commonly employed

facilities for hot metal production due to their superiority in productivity and heat

utilization (Geerds et al., 2011; Chen, et al., 2012; Suopajärvi et al., 2014). In order to

reduce iron ore into iron, metallurgical coke is fed from the top of the blast furnace.

Meanwhile, pulverized coal is injected and burned at the bottom of the furnace to

provide heat for the reduction reactions (Du, et al., 2010). On account of mass

consumption of coal for hot metal production, ironmaking is an energy-intensive

industry and a large amount of CO2 is emitted into the atmosphere (Chen et al., 2011;

Porzio et al., 2013; Hammond and Norman, 2014).

During the operation of blast furnace, blast air heated to temperatures of 1100–125oC is

blown into the furnace through tuyeres, and reacts with coke in raceways to generate

heat and reduction gases for iron ore reduction. To diminish the consumption of

expensive coke, some cheaper auxiliary fuels, such as oil, natural gas, and pulverized

coal, have been used as the substitutes of coke and injected through lances into

raceways. Due to the relatively low price and abundant reserve of coal in comparison

with other fossil fuels, nearly half of blast furnaces in the world (47.7%) use pulverized

coal injection (PCI), while only 4.1% use oil, 11.9% use gas, and 0.2% use plastic

injection (Schott, 2012). For a stable PCI operation, high coal burnout along with a low

pressure loss (high permeability) is always desirable in the regions of blowpipe, tuyere,

and raceway. However, it is difficult to simultaneously implement the two situations

because the enhancement of coal combustion intensity may raise the pressure loss of

blast flow within the raceway.

China Steel Corporation (CSC) is the only integrated steel producer in Taiwan, and has

four blast furnaces located in Kaohsiung and two in Taichung. For the cost reduction of

fuel and its stable supply, the auxiliary fuel injected at CSC has been changed from oil

to pulverized coal since 1988 (Du et al., 2001). In an attempt to increase PCI rate and

170

stabilize blast furnace operation, simulation models by computational fluid dynamics

(CFD) code have been developed at CSC to investigate the flow patterns of injected

coal and gas temperature distributions in combustion zone. Upon inspection of the

predicted trajectories of coal particles in the regions of blowpipe and tuyere (Du et al.,

2007), it was found that the dispersion of injected pulverized coal into hot blast was

poor when single lance injection was operated. This resulted in relatively low

combustion efficiency of coal and soot formation (Chen et al., 2007; Chen et al., 2008).

Besides, from the calculated temperature contours within the regions of blowpipe and

tuyere (Du and Chen, 2006), pulverized coal ignition could be triggered earlier under

the operation of double lance injection when compared to the single one. This

intensified the combustion efficiency of pulverized coal. For this reason, the pulverized

coal injection system in the blast furnaces of CSC was modified from the conventional

single lance to a double air-cooled coaxial lance system in 2001 (Yeh et al., 2002).

Instead of cooling air, the coal combustion temperature could be promoted with oxygen

flowing through the annulus of coaxial lance (Gudenau et al., 1994). This oxy-coal

injection technology has been adopted in many blast furnaces (Chung and Hur, 1997;

Peters et al., 2009; Austin et al., 2011; Hartig et al., 2011). The PCI operation of CSC’s

blast furnaces might be improved if the cooling air for the coaxial lance is replaced by

the enriched oxygen. However, the information regarding the application of oxygen

enrichment in pulverized coal injection remains insufficient. To recognize the

influences of lance configuration and oxygen enrichment pattern on the performance of

PCI operation, a CFD based simulation model has been established in this study. Three

different lance configurations, consisting of a single lance, a CSC’s double lance, and

an oxy-coal lance, are taken into account. The numerical predictions are able to provide

a useful insight into the coal combustion behavior in the regions of blowpipe, tuyere,

and raceway of a blast furnace. In addition, detailed discussion is made to reveal the

impact of enriched oxygen on coal ignition and pressure loss across the combustion

zone.

171

6.2 Methodology

The gas-particle flow and coal combustion in tuyere and raceway were calculated using

FLUENT V12 code. The flow field and temperature distribution were described using

3-D, steady-state Reynolds-averaged Navier–Stokes equations in association with the

RNG (Re-Normalization Group) k–e turbulence model. The pulverized coal particles

were treated as a dispersed phase using a Lagrangian method subject to the assumption

that each particle followed a discrete trajectory without interactions with any of the

other particles. The mathematical formulation is described below.

6.2.1. Gas-particle flow

6.2.1.1. Gas phase

Continuity:

mρU (6.1)

where ρ is the density, U is the mean velocity, and m is the mass transfer rate from

particulate to gas phase.

Momentum:

Deff

T

eff fρUUCUkpUUμρUU2

1

3

20 (6.2)

where μeff, p, and fD are the effective viscosity, pressure, and the drag forces of a

particle, respectively. Based on the Ergun equation (Ergun, 1952), the inertial loss

coefficient is expressed as 3

cαDα13.5C 0 where Dc represents the particle

diameter in coke bed and α is the voidage, which is defined as the total volume of the

voids divided by the total volume of the coke region.

Energy:

qHμ

C

λρUH

t

t

p Pr (63)

where H is the total enthalpy, λ is the thermal conductivity, Cp is the specific heat, μt is

the turbulent viscosity which is modeled by RNG k–ε turbulence model, and q is the

172

heat source from the particle phase. The turbulent Prandtl number, Prt, set in this study

is 0.85.

6.2.1.2 Single particle in dispersed phase

Mass:

mdt

dmp (6.4)

where the subscript p denotes the particulate phase, and m is coal particle mass.

Momentum:

ppDpD

ppUUUUρCπDf

dt

Umd 2

8

1 (6.5)

where Dp is the particle diameter, and CD is the drag coefficient given by Morsi and

Alexander (1972).

Energy:

reac

p

pgpppgp

pp

p,s Qdt

dmTTσεπD TTλNuπDq

dt

TmdC 442

(6)

where Cp,s , εp , σ , and Qreac. stand for the heat capacity of the particle, particle

emissivity, Stefan–Boltzmann constant, and heat released by the surface reaction,

respectively. The Nusselt number, Nu, is computed using the correlation of Ranz and

Marshall (1952a; 1952b).

6.2.1.3 Turbulence model

For the accurate and efficient predictions of the turbulent mixing and dispersion of

injected pulverised coal into hot blast (>160 m s-1

), the RNG k-ε model was adopted to

predict turbulent combustion (Biswas and Eswaran, 2002). This is because that the RNG

k-ε model can provide a better treatment in the mixing and dispersion of coal particles in

comparison with the conventional k-ε model. The complete formulation of the RNG k–ε

turbulence model is given as follows:

ρεGkμρUk kefft (6.7)

ρεCGCk

εεμρUε k1εefft

*2 (6.8)

173

where k is the turbulence kinetic energy, ε is the kinetic energy dissipation rate, and Gk

is the generation of turbulence kinetic energy due to the mean velocity gradients and

expressed by

UUkUUG tT

tk

I

3

2)( (6.9)

2kCteff (6.10)

The coefficient *2εC is given by

3

3

22012.01

)38.4/1(

CCC*

ε (6.11)

where η=Sk/ε, and S is the modulus of the mean rate of strain. The coefficients Cμ, C1ε,

C2ε, and αt are empirical constants, and their values derived empirically are 0.0845,

1.42, 1.68, and 1.393, respectively.

6.2.2 Turbulent combustion

When coal particles are heated, volatile matters will be released to react with oxygen,

resulting in diffusion flame combustion (Khalil, 1982). Owing to high blast temperature

(>1100 °C), it is reasonable to assume that the combustion reactions are fast compared

to fluid mixing rate. Therefore, the mixture fraction probability density function (PDF)

model (Sivathanu and Faeth , 1990) was employed. In the mixture fraction PDF frame,

individual species transport equations were not considered. Instead, the mixture fraction

transport equation was solved. The mixture fraction f in terms of the mass fraction is

written as

iOiF

iOi

XX

XXf

(6.12)

where Xi represents the mass fraction for some element i and the subscripts F and O

stand for the values at the fuel and oxidant sides, respectively. The mixture fraction f is a

conserved scalar and its value in a control volume can be calculated from the solution of

its time-averaged ( f ) instantaneous conservation equation:

p

t

t Sfσ

μfρU

(6.13)

174

In the above equation, t and pS designate dynamics viscosity and source term,

respectively, stemming from coal reactions in the gas phase, respectively. Meanwhile,

t is a computational parameter whose value is given by 0.85 (Jones and Whitelaw,

1982). In the PDF framework, the mean square value of concentration fluctuation g was

calculated through the following equation:

gk

ερCfμCg

σ

μρUg d

2

tg

t

t

(6.14)

where gC and dC are the computational parameters and they are 2.86 and 2.0,

respectively (Jones and Whitelaw, 1982). According to the mixture fraction f, molar

fraction of each gas species, density, and temperature in every control volume were

obtained.

6.2.3 Devolatilization of coal

The burning history of injected pulverized coal is featured by the following sequences:

(1) rapid heating; (2) devolatilization; (3) volatile oxidization; (4) residual char

combustion; and (5) char gasification. In examining the characteristic times of the

aforementioned reactions, the residence time of coal particle is around 20 ms (Steiler et

al., 1996), while the devolatilization reaction is implemented within 10-200 ms (Smoot

and Smith, 1985). With regard to char combustion, its characteristic time is in the order

of 1 s. The time required for completing char gasification is even longer (Smoot and

Smith, 1985). It is known that the devolatilization reaction initiates the coal combustion,

implying that the selection of parameters in modelling the devolatilization reaction is of

the utmost importance. To describe the coal devolatilization process more realistically, a

two-competing devolatilization model (Kobayashi, et al., 1977) was employed. The two

parallel and competing reactions are given as follows:

1111 )1( 1 VYSYC k (low temperature) (6.15)

2222 )1( 2 VYSYC k (high temperature) (6.16)

where C, Y, S, and V denote coal, stoichiometric coefficient, char, and volatile,

respectively. The relative importance of the two equations is mainly determined by

temperature. Specifically, when the temperature is low, the devolatilization reaction is

dominated by Equation 15. Alternatively, it is governed by Equation 16 once the

175

temperature is relatively high. Accordingly, the devolatilization reaction kinetics is

written as:

CYkYkdt

dv 2211 (6.17)



In a previous study (Du and Chen, 2006), it has been found that the kinetic parameters

suggested by Ubhayakar et al. (1977) could accurately predict the devolatilization

processes of coal particles and were thus adopted in this work. The Arrhenius rate

constant and activation energy in Equation 18 are set as 3.7 × 105 s

-1 and 74 kJ mol

-1,

respectively, while they were 1.46 × 1013

s-1

and 251 kJ mol-1

in Equation 19. The

stoichiometric parameter Y1 is taken as the volatile matter measured in proximate

analysis of coal, and Y2 is equal to 1.5Y1 (Du and Chen, 2006). The relationship of Y2

and Y1 fits the experimental data well with coals containing 10 to 40% volatile matter

(Shen et al., 2008). The developed CFD model has been validated by means of the

comparison of gas temperature between simulations and experimental measurements

(Du and Chen, 2006; Du et al., 2007).

6.2.4 Physical geometry and operating conditions

The operating conditions of CSC’ blast furnaces are given in Table 6.1 and used as the

boundary conditions in the numerical predictions. Detailed physical geometries and

sizes of the blowpipe, tuyere, and raceway are sketched in Figure 6.1. The coaxial lance

for double and oxy-coal system used a coal conveying pipe of half 1/2 inch stainless

steel tube (schedule 80), which were contained in a 1 inch tube (schedule 40).

Alternatively, a 3/4 inch straight pipe (schedule 80) was adopted as the single lance. The

placements for the all lance configurations studied in this work were the same at 9o with

respect to the centreline of the blow pipe. The tips of oxy-coal and single lance were

located at the centerline of the tuyere. The size and shape of the raceway were

determined by the porosity contour of 0.4, as shown in Figure 6.1a, and the raceway

was thought of as a porous medium (Du, 2011; Yeh et al., 2012). Because the specific

surface of injected pulverized coal was much higher than that of coke, only coal burning

176

was considered in the simulation. In the oxy-coal lance injection, 40–100% of enriched

oxygen was designed to go through the annulus of the coaxial lance, and the rest,

namely, from 60% to 0%, was added into the blast.

Table 6.1 A list of fuel properties and operating conditions.

Properties of pulverized coal

Proximate analysis (wt%, dry basis) Size distribution (wt%)

FC: 70.66%

VM: 20.02%

Ash: 8.32%

90μm: 5%

63μm: 25%

45μm: 55%

20μm: 15%

Hot blast conditions Other operating conditions

Temperature: 1423 K

Pressure: 4.5 atm

Flow rate of cold blast air: 3.753 kg s-1

Flow rate of enriched oxygen: 0.174 kg s-1

Single and double lance: 100% enriched

oxygen (0.174 kg s-1

) being added to

blast (O2 content: 25%)

Oxy-coal lance: 40-100% enriched oxygen

flowing through the lance, and the rest

proportion (60-0%) being added to

blast

Pulverized coal transportation

(1) Injection rate of pulverized coal:

0.51kg s-1

(2) Carrier gas (air) flow rate: 0.017 kg s-1

(3) Carrier gas and PC temperature: 298 K

Gas for the annulus of coaxial lance

(1) Cooling air flow rate for double lance:

0.0017 kg s-1

(2) Oxygen flow rate for oxy-coal lance:

40 to 100% of enriched oxygen (0.174

kg s-1

)

(3) Gas temperature: 298 K

177


6.3.1 Trajectories of coal particles

Figure 6.2 shows the trajectories of coal particles at the three different injection patterns

under the operating conditions given in Table 6.1. Meanwhile, the residence times of

coal particles are presented by different colors. As a whole, the residence time of coal

particles within the combustion region is less than 25 ms, regardless of which injection

pattern is operated. The dispersion of coal particles at the exits of the lances under the

CSC’s double lance injection (Figure 6.2b) is superior to those with single lance

injections (Figure 6.2a and c). Figure 6.2c displays the dispersion characteristics of

larger (e.g., 90 μlm) and smaller (e.g., 20 μm) particles under the operation of oxy-coal

lance injection to reveal the segregation of coal particles in terms of particle sizes. In the

region of lance exit, the larger particles (90 μm) tend to travel along the lance direction,

(a)

(b) (c)

Figure 6.1 Schematics of (a) physical sizes of computational domain as well as the

arrangements of (b) CSC’s double air-cooled lance and (c) single and oxy-coal lance

(α is the porosity within the raceway).

178

resulting from their higher inertia force against the hot blast. Under the influence of the

turbulent fluctuation of the blast, the smaller particles (20 μm) detach themselves from

the centerline of the lance quickly. This suggests the dispersion of injected coal particles

may be improved when the particle size is reduced. In examining the particle

trajectories under the operations of the single lance injection (Figure 6.2a) and oxy-coal

lance injection (Figure 6.2c), it is noted that the coal plume under the oxy-coal lance

injection is closer to the tuyere inner wall. Obviously, the coal particles with the

enriched oxygen flowing through the annulus of the oxy-coal lance could travel longer

along the direction of the injection lance.

(a) single lance

90 µm

20µm

(c) oxy-coal lance

(b) CSC’s double air-cooled lance

Residence time, s

Figure 6.2 Distributions of coal particle trajectory and residence time under (a)

single lance, (b) double air-cooled lance, and (c) oxy-coal lance injections.

179

6.3.2 Oxygen consumption within the combustion region

In a previous study (Du and Chen, 2006), it was found the coal burning was not

sensitive to the oxygen concentration in the blast within the region of blowpipe and

tuyere, in which the dominant mechanism of reactions was coal devolatilization rather

than char oxidation. When the raceway was taken into account, notable improvement in

coal combustion could be accomplished by oxygen enrichment (Shen et al., 2009b). It

follows that the operation of oxygen enrichment may play an important role in

intensifying coal combustion in the raceway. With conventional oxygen enrichment

(stove oxygen enrichment) in the single and the double lance injections, as can be seen

in Figure 6.3a and b, the oxygen in the coal plume is rapidly consumed by the volatiles

liberated from coal particles. Figure 6.3a and b depict that a certain amount of hydrogen

is remained within the coal plume. In other words, a volatile rich region is exhibited

within the coal plume due to insufficient oxygen. The unburnt volatiles in the coal

plume may undergo secondary pyrolysis reactions and be converted into tiny aerosols,

composed of soot particles and tar droplets (Chen et al., 2007; Chen et al., 2008). The

formation of solid carbon aerosols will raise energy loss, as a consequence of

incomplete combustion of fuel. Moreover, these tiny aerosols particles will cause

problems of their collection in the flue gas cleaning (Steeghs, 1992). When the oxy-coal

lance injection is operated, coal particles are surrounded by the enriched oxygen in the

upstream of the coal plume, as shown in Figure 6.3c. A comparison to Figure 6.4a and b,

the volatile rich region with lower levels is pushed downstream (Figure 6.4c). This

reflects that the oxy-coal injection consumes more oxygen for the volatile combustion

in the downstream after the chemical reactions are triggered. It is also implies that the

generation of tiny aerosols can be efficiently abated.

180

Mole fraction of oxygen, -

(a) Single lance


(c) Oxy-coal lance

Figure 6.3 Distributions of oxygen mole fraction under (a) single lance, (b) double

air-cooled lance, and (c) oxy-coal lance injections.

181

6.3.3 Ignition and temperature distribution

Figure 6.5 shows the isothermal contours in the regions of blowpipe, tuyere, and

raceway under the three lance injection patterns. Fuel ignition under the single lance

injection occurs in the vicinity of the tuyere exit (Figure 6.5a). With the operation of the

double lance injection, fuel ignition is triggered earlier (Figure 6.5b), stemming from

the better dispersion of coal particles and the increased contact surface between the

particles and hot blast. For the two injection patterns, low temperature zones within the

diffusion flames are observed, and they are consistent with the volatile rich zones within

the coal plumes (Figure 4a and b). The low temperature zones are attributed to the

endothermic reaction of char gasification (C + CO2 = 2CO). The developed diffusion

flame, which partition oxygen and volatiles into two different zones, is a result of the

Mole fraction of hydrogen, -

(a) Single lance


(c) Oxy-coal lance

Figure 6.4 Distributions of hydrogen mole fraction under (a) single lance, (b)

double air-cooled lance, and (c) oxy-coal lance injections.

182

reactions between oxygen and volatiles. The flame configuration for the oxy-coal

injection is fairly different from those found in the single and double lance injections.

Figure 6.5c depicts that fuel ignition and diffusion flame obviously shift toward the

downstream of the coal plume in the raceway. This is attributed to the cooling effect of

the enriched oxygen to the combustion. Moreover, the low temperature zone in the

flame under the oxy-coal injection is not as significant as those under the single and the

double air-cooled lance injections where diffusion flames accompanied by

low-temperature cores are exhibited. To farther into the recognition of the burning

characteristics of the oxy-coal lance injection, the gaseous temperature profiles along

the centerline of the tuyere at different proportions of enriched oxygen are plotted in

Figure 6.6. The profile under the single lance injection is also provided for comparison

because the lance is arranged at the same position as the oxy-coal lance. Under the

single lance injection, the rapid increase in gas temperature is located in the vicinity of

tuyere exit where the peak temperature is 2635 K. The temperature sharply drops to

1952 K, as a consequence of insufficient oxygen within the diffusion flame. The peak

temperatures for 40% (60% is added to the blast), 60%, 80%, and 100% of enriched

oxygen flowing through the annulus of coaxial lance are 2785 K, 2781 K, 2804 K, and

2785 K respectively. As expected, the peak temperatures for the oxy-coal injection are

located behind that for the single lance injection due to the cooling effect of the

enriched oxygen. Moreover, for a higher proportion of enriched oxygen, the peak

temperature is located at the deeper location of the raceway. The higher the proportion

of enriched oxygen injected through the annulus of oxy-coal lance, the lower the

magnitude of the gas temperature drop after reaching the peak level. This indicates that

coal combustion can be intensified at a higher oxygen level. Despite the late ignition

resulting from the cooling effect in the beginning of injection, the enriched oxygen

becomes a combustion enhancer to facilitate coal burning in the downstream of the coal

plume as long as the reactions are triggered.

183

6.3.4. Combustion efficiency of coal particles

Temperature, K

(a) Single lance


(c) Oxy-coal lance

Figure 6.5 Distributions of isothermal contours under (a) single lance, (b) double

air-cooled lance, and (c) oxy-coal lance injections.

Distance (m)

Ga

ste

mp

era

ture

(K)

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 2 2.2300

600

900

1200

1500

1800

2100

2400

2700

3000

Lance exit Tyuere exit

Single lance

40%60%80%100%

Figure 6.6 Distributions of gas temperature along the centreline of tuyere under

single lance injection and oxy-coal lance injections at different proportions of

enriched oxygen.

184

6.3.4 Combustion efficiency of coal particles

Figure 6.7 presents the profile of coal combustion efficiency (burnout) under the

oxy-coal lance injection at different proportions (40–100%) of enriched oxygen. The

combustion efficiencies under the operations of the single lance and the double lance

injections with 25% oxygen in the blast are also given in the figure for comparison. The

coal combustion efficiency is defined as the reduction percentage of combustible

portion in coal (Ishii, 2000). It is apparent that the combustion efficiency is improved

when the injection is changed from the conventional single lance injection to the double

or oxy-coal lance injection. Specifically, when the single lance injection is replaced by

the double lance injection, the combustion efficiency rises from 51.3% to 56.4%,

accounting for 5.1% of improvement. This can be explained by the better dispersion

(Figures 6.2 and 6.5) and earlier ignition of coal particles. The combustion efficiency of

the injected coal particles with oxy-coal lance injection goes up with increasing the

proportion of enriched oxygen, and it is in the range of 55.1–56.5%.; it is even higher

than that of the double lance injection when 100% proportion of enriched oxygen is

injected through the annulus.

185

Proportion of enriched oxygen through oxy-coal lance (%)

Co

alco

mb

ustio

ne

ffic

ien

cy

(%)

40 50 60 70 80 90 10051

51.5

52

52.5

53

53.5

54

54.5

55

55.5

56

56.5

57

Single lance

CSC's double air-cooled lance

Oxy-coal lance

6.3.5 Pressure loss

Practically, the pressure loss is an index of permeability resistance in the blast furnace.

To evaluate the performance of the lances in the regions of blowpipe, tuyere, and

raceway, it is essential to analyze the pressure loss due to the generation of gas and heat

from coal combustion. Figure 6.8 shows the profiles of pressure loss across the regions

of blowpipe, tuyere, and raceway at the three injection patterns. In comparison with the

single and double lance injections, relatively low resistance in the regions of blowpipe,

tuyere, and raceway is exhibited when the oxy-coal lance injection is operated. This is

attributed to the occurrence of ignition away from the tuyere exit (Figure 6.5c),

rendering that the effect of gas expansion is not as significant as the other two injections.

Obviously, the oxy-coal lance injection enables to fulfill two contradictory conditions at

the same time in the regions of blowpipe, tuyere and raceway: (1) to retard the coal

Figure 6.7 A comparison of coal combustion efficiency among single lance

injection, double air-cooled lance injection, and oxy-coal lance injections with

different proportions of enriched oxygen.

186

combustion for moderating the pressure loss in the upstream of coal plume; and (2) to

enhance coal combustion and reduce unburnt char generation in the downstream of coal

plume. Taking these advantages from the oxy-coal lance injection, blast furnaces can be

operated with more blast for higher productivity, or with higher PCI rate for lower fuel

cost, thereby achieving the goal of hot metal production with energy saving.

6.4 Conclusions

A three-dimensional CFD model for simulating coal combustion within the regions of

blowpipe, tuyere, and raceway under the operating conditions of CSC’s blast furnace

has been presented in this research. The performances of three different injection

patterns, in terms of coal burnout and pressure loss caused by the combustion, have

Proportion of enriched oxygen through oxy-coal lance (%)

Pre

ssu

relo

ss

(atm

)

40 50 60 70 80 90 1000.25

0.27

0.29

0.31

0.33

0.35

0.37

0.39

0.41

0.43

0.45

Single lance

CSC's double air-cooled lance

Oxy-coal lance

Figure 6.8 A comparison of pressure loss among single lance injection, double

air-cooled lance injection, and oxy-coal lance injections with different

proportions of enriched oxygen.

187

been examined as well. The predicted trajectories of coal particles show the poor

dispersion of injected coal particles. The dispersion can be improved when the coal size

is reduced. Owing to more contact surface between the injected coal particles and the

hot blast under the double air-cooled lance injection, early ignition is triggered; but this

causes higher pressure loss. With oxy-coal lance injection, the coal plume is surrounded

by enriched oxygen. Fuel ignition is delayed, as a consequence of the cooling effect of

the enriched oxygen. However, coal combustion is intensified at the downstream of the

coal plume, resulting from more oxygen consumed for the reaction. This can abate the

generation of tiny aerosols composed of soot particles and tar droplets. Compared to the

single and double air-cooled lance injections, less pressure loss across the combustion

region is exhibited from the oxy-coal lance injection in that the ignition position is away

from the tuyere exit. Increasing the proportion of enriched oxygen under the oxy-coal

lance injection facilitates the coal burnout, and the coal combustion efficiency is slightly

higher than that of the double lance injection when all enriched oxygen is used in the

lance. By virtue of higher combustion efficiency and lower pressure loss under the

oxy-coal lance injection, the performance of blast furnace may be improved. Compared

with the single lance injection, the coal particle trajectories under the oxy-coal lance

injection are closer to the tuyere exit due to the higher inertia force of the particles

against the hot blast. This should be taken into account for the designs of oxy-coal lance

injection. This model has provided useful information for understanding coal

combustion characteristics in the regions of blowpipe, tuyere, and raceway, and this is

conducive to the operation of pulverized coal injection in blast furnaces with

energy-saving.

188

CHAPTER 7

VOLATILE RELEASE AND PARTICLE FORMATION

CHARACTERISTICS OF INJECTED PULVERIZED COAL IN

BLAST FURNACES

This chapter experimentally investigates the volatile release and the generation of char

particle and tiny aerosols in the region of coal plume, where the oxygen is insufficient.

The experiment results provide directions for improving PCI operation.

Chen, W. H., Du, S. W. and Yang, T. H. (2007), Volatile release and particle formation

characteristics of injected pulverised coal in blast furnace, Energy Conversion and

Management, vol. 48, p. 2025-33

189

ABSTRACT

Volatiles release and particle formation for two kinds of pulverized coals (a

high-volatile bituminous coal and a low-volatile bituminous coal) in a drop tube furnace

are investigated to account for the reactions of pulverized coal injection in blast

furnaces. Two different sizes of feed particles are considered; one is 100-200 mesh and

the other 200-325 mesh. By evaluating the R-factor it is found the devolatilization

extent of larger feed particles is relatively poor. However, the swelling behavior of

individual or two agglomerated particles is pronounced which is conducive to the

gasification of chars in blast furnaces. In contrast, for the smaller feed particles,

volatiles liberated from the coal particles can be improved in a significant way as a

result of the amplified R-factor. This enhancement can facilitate the performance of

gas-phase combustion. Nevertheless, the residual char particles are characterized by

agglomeration, implying that the reaction time of the char particles will be elongated,

thereby increasing the possibility of furnace instability. Double-peak distributions in

char particle size are observed in some cases. This is possibly resulting from the

interaction of plastic state and blowing effect at particle surface. Considering the

generation of tiny aerosols composed of soot particles and tar droplets, the results

indicate that their production is highly sensitive to the volatile matter and elemental

oxygen contained in coal. Comparing the reactivity of the soot to that of the unburned

char, the former is always lower than the latter. Consequently, the lower the soot

formation, the better the blast furnace stability.

Keywords: Blast furnace; Pulverized coal; Drop tube furnace; Particle size; Soot.

190

7.1 Introduction

In pulverized coal injection (PCI) operation, the cheaper pulverized coal is injected with

hot air (1100-1200℃) into blast furnaces (BFs) as a substitute for coke. The economic

benefits of the PCI include a reduction in the cost of hot metal, resulting primarily from

decreased coke consumption and an increase in hot metal production. In addition, the

coal is consumed directly without going through the cokmaking plant, the PCI is also

thought to be environmentally friendly because it helps to reduce CO2 emissions (Du et

al., 2001). When coal particles are injected and proceeding from the blowpipe, tuyere,

and then to the raceway, as shown in Figure 7.1, they will experience rapid heating (the

heating rate ranges from 104 to 10

5 K/s), devolatilization, gas-phase combustion, and

char combustion and gasification (Ishii, 2000; Smoot and Smith, 1985). These processes

obviously include homogeneous and heterogeneous reactions. Regarding the particles

dynamics of the pulverized coal, while the coal particles are liberating volatiles they

may undergo swelling (Yu et al., 2003), fragmentation (Hurt and Davis, 1999), and

agglomeration (Shampine, et al., 1995), and eventually evolve into unburned char

particles, which can be consumed by CO2 in the furnace afterwards. Because the typical

residence time for coal particles reaching bird’s nest is around 20 ms (Steiler, et al.,

1996), the reactions between the gas phase and the solid phase are mainly governed by

devolatilization and pyrolysis followed by gas-phase combustion. In contrast, unburned

char combustion and gasification due to the interaction among carbon, oxygen, and

carbon dioxide are relatively unimportant in that the characteristic times of the

heterogeneous combustion and gasification are by far longer than the other reactions

mentioned above (Smoot and Smith, 1985).

In the past, in order to increase pulverized coal injection rate (PCR), a variety of

theoretical and experimental methods have been carried out. For example, Ohno et al.

(1994) investigated PC combustion in a raceway cavity by deriving theoretical

formulas; they also experimentally developed a coke-packed furnace to evaluate the

effect of mixing between the PC and oxygen on the combustion rate. Babich et al.

191

(1996) analyzed the effect of coal grinding on PC burning. Though an increase in coal

grinding level can raise coal combustion intensity, it also increases energy-waste of coal

grind. Hence an optimum coal grinding for PCI into BFs was suggested. The

configuration of injection lance was also studied to achieve higher combustion

efficiency. Chung and Hur (1997) experimentally conducted a coaxial lance with

enriched oxygen going through the annulus to improve coal combustion. They found

that the foregoing design is capable of increasing raceway depth and reducing bird’s

nest thickness in consequence of the enhancement of coal reactions. Ariyama et al.

(1994) experimentally studied pulverized coal combustion in tuyere zone by means of

single-lance and double-lance injections. It was illustrated that the performance of coal

combustion through the double lance injection is better than that by the other one. This

is a result of superior particle dispersion. Similarly, recent numerical predictions of Du

and Chen (2006) suggested that the double lance injection in a blast furnace can

facilitate the ignition of coal particles compared to that of single lance injection. They

also reported that the PCR in a practical BF has been promoted since the double-lance

injection was carried out, revealing that the double-lance injection does possess the

merit of increasing the practice of the PCI.

To pursue cost reduction, high PCR is always the operation target of a BF; however,

once the injection rate is promoted to a certain extent, it will seriously affect the furnace

stability. According to the PC trajectories calculated by Du et al. (2004), the dispersion

of coal plum to hot blast in the raceway is poor, resulting in low oxygen concentration

within coal plum. High PCR practice will not only produce more unburned char, but

also trigger secondary reactions and thereby generate a bit of soot (Fletcher et al.,

1997). The operation might go worse when the accumulation rate of the unburned char

and soot is faster than their consumption, mainly by CO2 inside of the BF. This results

in destroying the permeability of the furnace, excessive coke erosion, as well as

undesirable temperature distribution and cohesive zone shape (Kalkreuth, et al., 2005).

Accordingly, it is recognized that the formation of the unburned char and soot within

192

the coal plume and their reactivity in the BF have a vital effect on the stability of the

furnace operation. For these reasons, the main interests of the present study are on the

characteristics of volatile release from PC as well as related features of unburned char

and soot such as char particle size distribution, surface and internal structures of char,

and reactivity of char and soot. To provide a useful insight into the practice of PCI in

BFs, a drop tube furnace (DTF) will be developed to simulate coal particle reactions in

coal plum region under high-temperature and inert environments. The reactivity of the

unburned char and soot will also be evaluated through thermogravimetric analysis

(TGA). The obtained results enable us to provide a reference for choosing coals in the

performance of BFs.

7.2 Experiment

Two different coals, denoted by coal F and coal L, respectively, serve as the basis of the

present study. These coals are used in the form of PCI for the purpose of getting hot

Bird’s nest

Raceway

Injection

lance

Blowpipe

Tuyere

Coke

Pulverized coal Char particles

Figure 7.1 Schematic diagram of pulverized coal injection and internal structure of

blast furnace around raceway.

193

metal in BFs and their basic properties such as proximate analysis, elemental analysis,

and heating values are listed in Table 7.1. As seen in the table, the volatile matters

(VMs) of the coal F and coal L are 14.67% and 37.26%, respectively. According to the

coal classification of the ASTM, it is known that the former pertains to low-volatile

bituminous coal whereas the latter is classified into high-volatile bituminous coal. When

investigating the characteristics of coal reaction, conventionally, a variety of devices

such as thermogravimetry (TG) (Alonso et al., 2001), heated wire grid (HWG)

(Carpenter and Skorupska, 1995), and drop tube furnace (DTF) (Card and Jones, 1995;

Lee et al., 1996) can be employed. Corresponding to the TG, HWG, and DTF the

heating rates are approximately in the orders of 0.1-1, 103, and 10

4-10

5 K/sec (Anthony,

et al, 1976), respectively. As mentioned in the introduction, in general, the heating rate

for coal particles in raceway is around 104-10

5 K/sec. Consequently, it is proper to use

DTF rather than TG or HWG to simulate coal reactions in the raceways of BFs.

Because of this, a DTF and a standard testing procedure are conducted in the present

work. Alternatively, when the reactivity of char and soot is evaluated, the characteristic

time of heterogeneous reaction between these particles and carbon dioxide in BF is in

the order of second, thermogravimetric analysis is thus performed.

Regarding the developed DTF, as a whole, the entire system can be partitioned into

feeding, reactor, collection, and control subsystems. The feeding unit is constructed by a

310 stainless steel tube which is enveloped by a water jacket and cooled by water to

reduce the possible damage caused by the high temperature furnace. Pulverized coal is

stored in a hopper and fed into the furnace by means of a screw feeder. A hitter with

periodically striking the feed tube is mounted beside the tube to avoid bridging

phenomenon in the tube when charging coal particles. With regard to the reactor, its

diameter and length are 70 mm and 1000 mm, respectively, and the material of the

reactor is composed of crystal Al2O3 which can sustain up to 1600℃. The traveling time

of coal particles within the reactor is around 0.5 second. After the pulverized coal passes

through the reactor, the formed unburned char particles are directly collected from the

194

bottom of the reactor due to inertial impact. The residual particles will travel through a

cyclone in which the medium-size particles are caught by means of centrifugal force.

Thereafter, the tiny aerosols such as soot particles and tar droplets are captured on a

filter. For the heating element, the furnace is heated by a SiC rod which can promote the

reaction temperature up to 1650℃ and a R-type thermocouple is installed in the reactor

to provide a reference temperature of the controller. The reaction temperature is

controlled by a PID (proportional band integral derivative) temperature controller and a

SCR (silicon controlled rectifier) power controller.

As far as the experimental procedure is concerned, when PC is sent into the DTF, pure

nitrogen with volume flow rate of 1 L/min is blown as well to aid carrying the fuel into

the furnace. In addition, preheated gas (N2) with the temperature of 350℃ and volume

flow rate of 1 L/min is also sent into the furnace. The pulverized coal and the preheated

gas are in a co-current pattern. Two different particle size ranges, consisting of 100-200

mesh and 200-325 mesh, are considered to account for coal particles reactions in the

DTF. Moreover, three different reaction temperatures composed of 1000, 1200, and

1400℃ are included in that these temperatures cover the reaction environments of the

raceway in BFs. To ensure the measuring quality, previous to experiments air with fixed

flow rate is blown into the reaction system and then measured at the system exit. This

guarantees that no gas leakage occurs in the system. Following the collection of the

unburned char, particle size distributions are analyzed by laser particle size analyzer

(Coulter LS100) in which the particle sizes ranging from 0.4 to 1000 μm can be

measured. With regard to the reactivity of char particles and soots, they are analyzed in

a TG with the heating rate of 8℃/min. Before that, the soots are washed by acetone and

then dried to separate tar. Carbon dioxide with the volumetric flow rate of 2 L/min is

sent into the TG because the gasification of char and soot in BFs is simulated.

195

Table 7.1 Proximate and ultimate analyses of the investigated coals


7.3.1 Devolatilization extent

Volatile release (or devolatilization) extent from PC in a short time can be evaluated by

examining the value of R-factor (Gibbins et al., 1993). The parameter R-factor stands

for volatile release ratio under rapid heating condition to standardized volatile matter

(VM) test from the ASTM. Hence, the higher the R-factor, the better the devolatilization

extent for coal particles travelling in a reaction zone. That is to say, the coal featured

with high R-factor will facilitate volatile liberation from coal particles, thereby

intensifying gas-phase combustion. The R-factor, which is established based on the ash

tracer technique, is defined by

coal) daf (% coal ofcontent VM prox.

VR

where V, which essentially represents volatile yield, is given as

100chardry %char ofcontent ash

coaldry % coal ofcontent ash 1

V

Coal Coal F Coal L

Classification (ASTM) Low-volatile

bituminous

High-volatile

bituminous

Proximate Analysis

(wt %)

Volatile matter 14.67 37.26

Moisture 1.10 1.74

Fixed carbon 76.35 52.51

Ash 7.89 8.50

Ultimate analysis

(wt %, dry-ash-free)

C 91.46 83.19

H 4.15 5.60

N 1.78 1.75

S 0.57 0.53

O (diff.) 2.04 8.82

Heating value (kcal/kg) 7410 7060

196

The R-factor values of the two coals under the three reaction temperatures are displayed

in Figure 7.2. As a whole, the R-factor is obviously affected by both the reaction

temperature and feed particle size. For the case of larger feed particles (i.e., 100-200

mesh), as shown in Figure 7.2a, it is difficult to completely liberate volatiles from the

two coals in consequence of R-factor being smaller than unity. The value is larger than

one only when the reaction temperature is as high as 1400℃. Once the feed particle size

is reduced to 200 to 325 mesh, Figure 7.2b depicts that the R-factor is enlarged greatly,

with the exception of the coal L at 1000oC. It follows that the devolatilization process

can be improved in a significant way if the feed particle size is decreased to a certain

extent. The improvement in the devolatilization extent is particularly remarkable for the

low-volatile bituminous coal (coal F). Accordingly, it should be illustrated that, despite

higher VM contained in the coal L, its R-factor is still lower than that of the coal F. This

implies that the devolatilization extent is not determined by the VM in nature.

7.3.2 Particle formation of the low-volatile bituminous coal

Subsequently, particle size distributions of the coal F before and after reactions are

demonstrated in Figure 7.3. For the coal F with larger feed particles, once the coal

Temperature (0C)

R-f

acto

r

900 1000 1100 1200 1300 1400 15000

0.25

0.5

0.75

1

1.25

1.5

1.75

2

Coal F

Coal L

(a) Mesh 100-200

Temperature (0C)

R-f

acto

r

900 1000 1100 1200 1300 1400 15000

0.25

0.5

0.75

1

1.25

1.5

1.75

2

(b) Mesh 200-325

Figure 7.2 Profiles of R-factor of two different coals at various reaction

temperatures.

197

Particle diameter (m)

Vo

lum

e(%

)

10-1

100

101

102

1030

1

2

3

4

5

6

7

8

9

10

11

12

Raw coal

1000oC

1200oC

1400oC

(a)

Partialfragmentation

Swelling


Vo

lum

e(%

)

10-1

100

101

102

1030

1

2

3

4

5

6

7

8

9

10

11

12

Raw coal

1000oC

1200oC

1400oC

(b)

Swelling

Agglomeration

experiences reactions, it is evident that the size distributions of unburned char are not as

sharp as that of the feed coal no matter what the reaction temperature is. The curves also

depict that part of the char particles become smaller whereas some particles tend to

extend. It is thus recognized that the coal particles are simultaneously characterized by

fragmentation and swelling while they are evolving into unburned char. When the peaks

of the curves are examined in detail, prior to reaction the maximum value of the curve

occurs at d=130.7μm (the solid line shown in Figure 7.4), where d designates particle

diameter. In addition, corresponding to the three curves of the reaction temperatures

1000, 1200, and 1400oC, Figure 7.4 depicts that the maximum values or peaks take

place at d=130.7, 117.4, and 105.5μm, respectively. The maximum value decreases with

increasing reaction temperature, therefore it is inferred that the peaks are dominated by

little debris separated from their parent particles, as a consequence of thermal impact

upon the particle surface stemming from rapid heating in the DTF. To provide a more

clear observation on the unburned char formation, the scanning electron microscope

(SEM) images of the char particles are demonstrated in Figure 7.5. It is apparent that, in

the cases of larger feed particles (Figures 7.5a-5c), most char particles are characterized

by individual particles and accompanied by a few debris. The figures also indicate that

increasing reaction temperature makes the char particles rounder.

Figure 7.3 Particle size distributions of coal F before and after experiencing reactions

with (a) larger feed particles and (b) smaller feed particles.

198

Temperature (oC)

Pe

ak

loca

tio

n(

m)

800 1000 1200 1400 160050

100

150

200

250

300

Larger feed particlesSmaller feed particles

Second peak

First peakPeak location of feed coal

Temperature (oC)

Pe

ak

loca

tio

n(

m)

800 1000 1200 1400 160050

100

150

200

250

300

Larger feed particlesSmaller feed particles

Second peak

First peak

Peak location of feed coal

Figure 7.4 Peak locations of particle size distributions for coal F before and after

experiencing reactions.

Figure 7.5 SEM images of unburned chars of coal F at larger feed particles (a-c) and

smaller feed particles (d-f).

(a) 1000℃ (b) 1200℃ (c) 1400℃

(d) 1000℃ (e) 1200℃ (f) 1400℃

199

When the coal F with smaller feed particles is considered, unlike the preceding

behaviors, the three curves shown in Figure 7.3b shift rightward to a great extent; the

case of 1400oC even exhibits a double-peak distribution. In examining the maximum

values of the curves, it occurs at d=76.46 μm for the feed coal (the dashed line shown in

Figure 4) whereas they develop at 85.11 and 94.74 μm for the unburned chars under the

situations of 1000 and 1200oC, respectively. Physically, on account of smaller feed

particles, heat transferred from the gas phase to the particle phase is more effective,

rendering that more volatiles are released, as elucidated in Figure 2. This results in the

characteristic of particle fragmentation being intensified. At the same time, the plastic

state on the particles surfaces and the collisions among the particles make them

agglomerate each other. As a result, as presented in Figures 5d-5f, the generated char

particles are featured by agglomeration instead of the individual particles (Figures

5a-5c). In regard to the case of 1400oC, the first and the second peaks appear at 76.46

and 248.6μm, respectively. It is known that the Stefan flow (Chen and Jiang, 2000)

stemming from particle surface will be elicited as long as volatiles are liberated. If the

blowing effect of the Stefan flow is high to a certain extent, the tendency of particles

gather will be suppressed somewhere. This might be the mechanism which causes the

phenomenon of the double-peak distribution. In a word, there are two major forces at

play while coal particles are evolving into chars. The first one is the plastic state which

will trigger agglomeration; the second one is the blowing effect which will prohibit char

particles to combine together.

7.3.3 Particle formation of the high-volatile bituminous coal

When attention is focused on the high-volatile bituminous coal (viz., the coal L) with

larger feed particles, a comparison with Figure 7.3a reveals that the curves shown in

Figure 7.6a move rightward to a certain extent after undergoing reactions. This is

because the particles swelling and agglomeration are more significant, resulting from

200

higher VM owned by the coal L. The peak of the feed coal is located at d=105.7 μm (the

solid line shown in Figure 7.7). Once the coal experiences devolatilization,

corresponding to the reaction temperatures of 1000, 1200, and 1400oC the peaks move

to 200.6, 180.3, and 161.9 μm, respectively. In light of these char particles being larger

than the feed particles approximately by a factor of 2, it is realized that the chars

formations near the peaks are no longer dominated by individual particles but mainly

characterized by two particles agglomeration. The aforementioned mechanism can be

verified by observing the SEM images of the char particles (Figures 7.8a to 8c) where

the majority of char particles exhibit low level accumulation.


Vo

lum

e(%

)

10-1

100

101

102

1030

1

2

3

4

5

6

7

8

9

10

11

12

Raw coal

1000oC

1200oC

1400oC

(a)

Swelling

Agglomeration


Vo

lum

e(%

)

10-1

100

101

102

1030

1

2

3

4

5

6

7

8

9

10

11

12

Raw coal

1000oC

1200oC

1400oC

(b)

Agglomeration

Figure 7.6 Particle size distributions of coal L before and after reactions with (a)

larger feed particles and (b) smaller feed particles.

201

Temperature (oC)

Pe

ak

loca

tion

(m

)

800 1000 1200 1400 160050

100

150

200

250

300

Larger feed coalSmaller feed coal

Second peak

First peak

Peak location of feed coal

Temperature (oC)

Pe

ak

loca

tion

(m

)

800 1000 1200 1400 160050

100

150

200

250

300

Larger feed coalSmaller feed coal

Second peak

First peakPeak location of feed coal

Figure 7.7 Peak locations of particle size distributions for coal L before and after

experiencing reactions.

Figure 7.8 SEM images of unburned chars of coal L at larger feed particles (a-c) and

smaller feed particles (d-f).

202

When the situations of the smaller feed particles are taken into account, the curves

shown in Figure 7.6b indicate that the double-peak distributions develop in the cases of

1000 and 1200oC. From the profiles shown in Figure 7.7, the values of the second peak

are larger than that of the first peak approximately by a factor of four. It is inferred that

the second peaks yielded are mainly contributed by over two char particles together.

The SEM images shown in Figures 8d and 8e clearly suggest the situations of particles

accumulation. With regard to the result of 1400oC, both the Stefan flow and particles

fragmentation are further enhanced. In such a situation, the DTF tends to behave as a

mixed flow reactor and the char agglomeration is in a more uniform environment. As a

result, the characteristic of the double-peak distribution disappears and the observation

of particles accumulation is provided in Figure 7.8f.

To proceed farther into the recognition of particles formation under the condition of

larger feed particles, the SEM cross-sectional images of the feed coal F and coal L as

well as their char particles with the reaction temperatures of 1000 and 1400oC are

demonstrated in Figure 7.9 Basically, the internal structures of the feed particles of the

coal F and coal L are compact (Figures 7.9a and 9b). Following the reaction with the

temperature of 1000oC, the interior of the char of the coal F becomes porous

accompanied by some larger holes (Figure 7.9c). In view of the higher VM of the coal

L, volatiles emission from the coal is more pronounced so that the holes inside the char

particles are larger (Figure 7.9d). With promoting the reaction temperature to 1400℃,

the char interiors of both the coal F and coal L are further emptied (Figures 7.9e and 9f).

It is not surprising because more volatiles are transported into the gas phase compared

to the results of 1000oC. In the study of Wall et al. (2002), char internal structures were

categorized into three groups in accordance with geometric parameters and porosity.

Upon inspection of the SEM cross-sectional images shown in Figure 7.9, the three

groups are also obtained. Specifically, Figure 7.9c pertains to group III where the char

porosity is lower and the wall is very thick. Alternatively, the internal structures with

203

group I and group II can be observed in some char particles shown in Figures 9d to 9f

since the porosity is higher and the average wall thickness is thinner.

(a) Feed coal F (b) Feed coal L

(c) Char particles from coal F at 1000oC (d) Char particles from coal L at 1000

oC

(e) Char particles from coal F at 1400 oC (f) Char particles from coal L at 1400

oC

Group III

Group II

Group I

Group I

Group II

Group II

Figure 7.9 SEM images of feed coal and unburned char particles shown in

cross-section.

204

7.3.4 Aerosol formation and reactivity

It is known that soot and tar particles will be produced due to volatiles secondary

reactions if the fuel undergoes incomplete combustion. By virtue of higher crystal

structure of carbon in soot which causes lower reactivity, the accumulated soot in blast

furnaces will reduce gas permeability. The weight percentages of soot and tar from the

two coals reactions at the three reaction temperatures are plotted in Figure 7.10 where

the feed particles of 100-200 mesh are tested. For the coal F (or the low-volatile

bituminous coal), Figure 7.10a suggests that the generation of soot and tar increases

with increasing reaction temperature where the weight percentage ranges from 0.13 to

0.25. This is, of course, attributed to more volatiles release which results in more soot

and tar production. Alternatively, for the high-volatile bituminous coal, namely, the coal

L, one can find that the weight percentage is in the range from 1.8 and 2.2. These values

are by far higher than the coal F suggesting that soot and tar formation is highly

sensitive to the VM content. Contrary to the behavior shown in Figure 7.10a, it is of

interest that the weight percentage from the coal L decreases with temperature. It is

inferred that the preceding feature is a consequence of high elemental oxygen contained

in the coal, as shown in Table 7.1. At higher temperature, the reaction between volatiles

and oxygen is more complete. Hence, increasing temperature reducing the formation of

the aerosol particles, as observed.

7.3.5 Reactivity of char and soot

Under the condition of 1400℃ and 100-200 mesh feed particles, the thermogravimetric

analyses of the unburned char and soot from the coal F and coal L reacting with carbon

dioxide are shown in Figure 7.11. It is observed that the decaying trends of both the

unburned char and soot from the coal F are slower than that from the coal L. This

elucidates that the reactivity of the char and soot from the low-volatile bituminous coal

is worse than that from the high-volatile bituminous coal. Furthermore, from the

viewpoint of furnace stability, the performance instability of furnaces may be excited if

205

Temperature (0C)

We

igh

tp

erc

en

tag

e(%

)

We

igh

tp

erc

en

tag

e(%

)900 1000 1100 1200 1300 1400 15000

0.1

0.2

0.3

0.4

0.5

0.6

1.5

1.6

1.7

1.8

1.9

2

2.1

2.2

2.3

2.4

2.5

Coal F

Coal L

the low-volatile bituminous coal is used. In addition, for either the coal F or coal L, the

reactivity of soot is slower than the unburned char. It follows that the lower the soot

production ratio, the higher the gas permeability in blast furnaces. In light of the

obtained results, the reaction physics of the two coals are summarized in Table 7.2.

Temperature (0C)

We

igh

tp

erc

en

tag

e(%

)

600 700 800 900 1000 11000

20

40

60

80

100

Coal F unburned charCoal L unburned charCoal F sootCoal L soot

Figure 7.11 Thermogravimetric analyses of the produced unburned chars and soots

at 1400oC.

Figure 7.10 Profits of soot and tar formations with respect to reaction

temperature.

206

T

ab

le 7

.2

Sum

mary

of

reacti

on p

hysi

cs

of

the t

wo c

oals

207

7.4 Conclusions

Experimental studies on the characteristics of volatiles release, particles formation, and

reactivity of unburned char and soot of two different coals have been achieved in the

present study. The purpose was to recognize the fundamental physics of pulverized coal

reactions in blast furnaces. An examination of R-factor reveals that either decrease feed

coal particle size or increase reaction temperature can efficiently enhance volatile

liberation. As a whole, the value of the R-factor of the larger feed particles is relatively

small; however, the swelling behavior of the char particles is conducive to the

subsequent char combustion and gasification reactions in blast furnaces because of the

surface enlargement of individual or two agglomerated particles. On the contrary, for

the finer feed particles the devolatilization extent is more complete. This can facilitate

the gas-phase combustion. However, the char formation is dominated by over two

particles agglomeration which will disadvantage the furnace stability. The double-peak

distributions in particle size are exhibited in some cases possibly due to the interaction

between softening stage and blowing effect of the Stefan flow. Furthermore, in view of

volatiles release from particles interiors, three different internal structures in the char

particles are observed and they depend strongly on the VM and reaction temperature.

Considering the generation of tiny aerosols which consist of soot particles and tar

droplets, it is mainly determined by the content of VM and elemental oxygen. Finally,

from the examinations of the unburned char and soot in TG, we find that the reactivity

of the unburned char from the low-volatile bituminous coal is lower than that from the

high-volatile bituminous coal. Moreover, the reactivity of the soot is lower than the

char. The obtained results have provided a very useful insight into the choice of injected

pulverized coal from the viewpoints of gas-phase combustion as well as char and soot

gasification in blast furnaces.

208

CHAPTER 8

PULVERIZED COAL BURNOUT IN BLAST FURNACE

SIMULATED BY A DROP TUBE FURNACE

This chapter describes the method developed at China Steel Corporation to evaluate the

combustion efficiency of PCI coals using a drop tube furnace. The method is currently

employed at CSC for the selection of PCI coals.

Du, S. W., Chen, W. H. and Lucas, J. A. (2010), Pulverised coal burnout in blast


209

ABSTRACT

Reactions of pulverized coal injection (PCI) in a blast furnace were simulated using a

drop tube furnace (DTF) to investigate the burnout behavior of a number of coals and

coal blends. For the coals with the fuel ratio ranging from 1.36 to 6.22, the experimental

results indicated that the burnout increased with decreasing the fuel ratio, except for

certain coals departing from the general trend. One of the coals with the fuel ratio of

6.22 has shown its merit in combustion, implying that the blending ratio of the coal in

PCI operation can be raised for a higher coke replacement ratio. The experiments also

suggested that increasing blast temperature was an efficient countermeasure for

promoting the combustibility of the injected coals. Higher fuel burnout could be

achieved when the particle size of coal was reduced from 60-100 to 100-200 mesh.

However, once the size of the tested coals was in the range of 200 and 325 mesh, the

burnout could not be improved further, resulting from the agglomeration of fine

particles. Considering coal blend reactions, the blending ratio of coals in PCI may be

adjusted by the individual coal burnout rather than by the fuel ratio.

Keywords: Combustion; Burnout; Pulverized coal injection (PCI); Drop tube

furnace (DTF); Blast furnace; Blend; fuel ratio.

210

8.1 Introduction

Blast furnaces are a crucial and the most commonly employed facility in ironmaking

processes. The interior of a blast furnace is filled with alternating layers of coke and ore

burden charged from the top of the furnace to initiate the production of hot metal (Ishii,

2000). The fed coke descents to the hearth of the furnace and constructs a porous coke

bed named deadman. The formed hot metal and slag inside the blast furnace can

penetrate through the deadman and flow downward to the bottom of the blast furnace.

Meanwhile, blast air heated to the temperature of 1100-1250°C is blown into the furnace

through the tuyeres. As a result, a cavity called raceway is formed around the exit of a

tuyere. In the raceway, the heated air takes part in the combustion of deadman coke to

generate heat and reducing gases. The heat and reducing gases are used for the

production of hot metal. To reduce the consumption of coke, some cheaper auxiliary

fuels, such as oil, natural gas, pulverized coal (Perlov, 1987) and even plastic wastes

(Ziebik and Stanek, 2001), have been used as the substitute of coke and injected through

the injection lance into the raceway. It should be emphasized that the technique of

auxiliary fuel injection can not only promote the productivity of a blast furnace but also

increase the flexibility of practical operation. For example, the oxygen concentration in

hot blast can be enriched (Jianwei, et al., 2003), thereby intensifying the combustion of

the auxiliary fuel. Since the second energy crisis occurred in 1979, the technique of

pulverized coal injection (PCI) has become a vital method in ironmaking process due to

the relatively lower price and abundant reserve of coal compared to other fossil fuels

(Toyoda, 1983).

The operation of PCI was introduced into the blast furnaces of China Steel Corporation

(CSC) in 1987 (Lai et al., 1994). Since that, a high PCI rate is always a desirable

operation target of CSC to reduce the cost of fuel. To make the injected coals ignited

earlier in the tuyere for higher combustibility, the coals with low fuel ratio, which

generally had high ignitibility (Kurose, et al., 2004), were solely injected into the blast

furnaces since CSC commenced the PCI operation. The fuel ratio is defined as the

211

weight ratio of fixed carbon to volatile matter in a raw coal (Kurose, et al, 2004).

Practically, it is known that the operational cost for producing hot metal can be reduced

if the consumed fuel rate (coke rate plus PCI rate) based on per metric ton of hot metal

is decreased. To approach this goal, some coals with high fuel ratio, which have higher

calorific values, have been blended into low ones at CSC since 2002 (Du et al., 2004).

On the other hand, the past research has suggested that soot formation from coal

reaction with low fuel ratio will exceed that with high fuel ratio, and the reactivity of

soot is lower than that of unburned char (Smoot and Pratt, 1979; Chen et al, 2008).

Consequently, another advantage for coals with high fuel ratio injected into the blast

furnaces is that the formation of soot can be abated (Chen et al., 2007). For these

reasons, increasing the blending ratio of a coal with high fuel ratio in PCI coals has

become an important target in the operation of CSC’s blast furnaces. When coal

particles are injected into the blowpipe-tuyere-raceway area, by virtue of very short

residence time (Steiler et al., 1996) and poor dispersion of the particles into the blast

(Du and Chen, 2006; Du, et al., 2007), the generation of unburned char from the

operation of high PCI rate is unavoidable. This is especially significant when a coal

with high fuel ratio is injected. Although it was reported that the unburned char trapped

in the furnace can be consumed by CO2 and H2O and by the direct reduction of FeO

(Iwanaga, 1991), the permeability of gas and liquid in the furnace may be adversely

affected if the accumulation rate of unburned char exceeds its consumption rate.

Therefore, if one intends to achieve a smooth operation of blast furnace, it is essential to

evaluate the combustion efficiency of PCI coals, especially for the coals with high fuel

ratio.

To figure out the behavior of coal reaction in blast furnaces, many experiments of coal

combustion have been carried out using combustion rigs to simulate the situations

around blowpipes. The combustion rig was attached to an empty combustion chamber

(Mathieson, et al., 2005; Ueno et al., 1993) or to a coke bed in some cases (Yamagata,

et al., 1992; Ariyama et al., 1994). However, in a CSC’s internal research for the

212

combustion experiments using a combustion rig connected with an empty combustion

chamber (Du et al., 2001), many operation difficulties have been encountered; they

included pulverized coal preparation, char sampling, temperature measurement and

keeping a constant coal injection rate. As a matter of fact, it has also been found that the

experiments were a time, manpower and cost consuming process. Over the years a

number of studies concerning coal combustion using drop tube furnaces have been

reported. Despite the inherent difference between the realistic combustion environment

of the raceway and the drop tube furnace (DTF), the DTF is still an effective device

when one attempts to evaluate the combustion performance of PCI coals in an

environment with high heating rate (>104

K/s). In the present study, a DTF combustion

system has been developed and the effects of some parameters, composed of coal type

(or fuel ratio), reaction temperature, particle size and blending ratio of binary coal, on

the combustion efficiency have been taken into consideration. The obtained results are

able to provide useful information to the operation of PCI in blast furnaces.

8.2 Experiments

8.2.1 Reaction system

The experiments were performed by dropping pulverized coal particles into a drop tube

furnace reaction system, as shown in Figure 8.1. The DTF system consisted of a feeding

subsystem, a reactor and a particle collection subsystem. The feeding subsystem was

composed of a hopper, a screw feeder, a lance and an electric hammer. The hammer was

arranged to periodically knock the lance during the experiments to aid in dispersing coal

particles into the carrier gas. The reactor was heated by a SiC heating element

controlled by a PID (proportional band integral derivative) temperature controller and a

SCR (silicon controlled rectifier) power controller. An R-type thermocouple was

mounted in the reactor to detect the temperature in the reaction zone and the detected

temperature was used as the reference of the SCR. Two ends of the reactor were sealed

by high temperature o-rings. The char particles were collected by a cyclone through a

213

sampling probe. The sampling probe was enveloped by a water jacket and the water

jacket was filled with cooling water so as to protect the probe. A suction motor was

connected to the exit of the cyclone to keep the coal particles moving straightforward

into the probe. The position of the probe was adjustable for controlling different

residence times, and a cooling nitrogen stream was arranged at the entrance of the probe

to prevent further reaction of the collected char particles when they entered the probe.

1. Carrier gas 2. Secondary gas 3. Rotameter

4. Hopper 5. Preheater 6. Thermocouple

7. DTF 8. Heater 9. Sampling probe

10. Cooling water 11. Container 12. Cyclone

13. Unburned char 14. Pump 15. Exhausted gas

14

15

1

3

9

11

5

8

6 2

7

13

10

4

12

Figure 8.1 Schematic of the reaction system.

1. Carrier gas 2. Secondary gas 3. Rotameter

4. Hopper 5. Preheater 6. Thermocouple

7. DTF 8. Heater 9. Sampling probe

10. Cooling water 11. Container 12. Cyclone

13. Unburned char 14. Pump 15. Exhausted gas

14

15

1

3

9

11

5

8

6 2

7

13

10

4

12

Figure 8.1 Schematic of the reaction system.

214

8.2.2 Experimental procedure and conditions

The tested coals were dried in a nitrogen-purged oven at 100-105°C for 24 hours and

then pulverized and sieved to different size fractions for experiments. During the

experiments, the tested coal particles stored in the hopper were fed by a screw feeder.

The coal particles and carrier gas (2L min-1

) were introduced from the top of the reactor

and they travelled along the centerline of a ceramic tube with the internal diameter of

70mm. To lessen the significant agglomeration of the coal particles, the feeding rate of

the coal particles was controlled to be as lower as 5 g h-1

. Additional secondary gas (3L

min-1

) was preheated to 350°C followed by blowing into four metal tubes around the

lance. The compositions of the carrier gas and the secondary gas were the same as the

atmospheric air. The base experimental conditions used in this study included the coal

particle size of 100-200 mesh and the reaction temperature of 1200°C. The comparison

of the combustion efficiency among the tested coals was made in terms of the base

conditions. To investigate the effects of reaction temperature and coal particle size on

the coal combustion efficiency, the burning experiments were carried out for the

reaction temperature ranging from 1100 to 1400°C as well as the coal particle size in the

ranges of 60-100, 100-200 and 200-325 meshes. Combustion experiments of binary coal

blends were also performed in this study. As described in Introduction, to increase the

calorific values of coals used for PCI operation, a coal with low fuel ratio can be

blended with a coal with high fuel ratio. This blended fuel can reduce the fuel rate used

in PCI to a certain extent compared to the utilization of the coal with low fuel ratio

alone. Because of this, two coals with high fuel ratio (>5) were individually blended

with a coal with low fuel ratio (<2). The blending ratios for the coals with high fuel

ratio in the coal blends covered 25, 50 and 75 wt%. To ensure the measurement stability

and accuracy, the reaction system was leak tested before the experiments were

performed. Moreover, the R-type thermocouple was calibrated annually by a CSC’s

standard furnace to ensure the accuracy of the measured temperature in this study. In

comparison with the readings given by a standard wire, the errors of the thermocouple

215

were found to be within +/- 3°C in the temperatures ranged from 1000 to 1400°C. The

burnout experiments were carried out at least twice for some tested coals to guarantee

that the results could be reproduced. To provide an analysis of the reliability of the

experiments from the coals covering a wide range of fuel ratio, four different coals,

Coals A, B, D and K, with the fuel ratio in the range of 1.36-6.33 were thus tested. In

other words, the analyzed fuels included the coals with the lowest and the highest fuel

ratios. Figure 8.2 displays the values of burnout of the four coals. As seen in the figure,

the relative differences of the burnout tests were controlled below 5%, revealing that the

quality of the experiments was reliable.

Fuel ratio (-)

Bu

rno

ut

(%)

1 2 3 4 5 6 740

50

60

70

80

90

100

Test 1

Test 2

Coal D

Coal B

Coal A

Coal K

Figure 8.2 Tests of experimental stability of four different coals under the base

experimental conditions.

216


In the following discussion, eleven coals (Coal A to Coal K), covering a wide range of

fuel ratio from 1.36 to 6.22, were tested. The proximate analyses, fuel ratios and higher

heating values of the tested coals are summarized in Table 8.1. Meanwhile, this study

also considered the effects of reaction temperature, particle size and blending ratio of

two different coals on the burnout behavior. Some discussion from the viewpoint of

practical operation of PCI based on the obtained results will be addressed.

Table 8.1 Proximate analyses (dry basis), fuel ratios and higher heating values (HHV,

dry basis) of the investigated coals.

8.3.1 Combustion efficiency of individual coals

The combustion efficiency of a coal is often measured in terms of burnout. In this study,

the ash tracer method (Tate, 1993; Osório, et al., 2006] was adopted to determine the

burnout of the tested coals. The burnout can be expressed as

100

1001

(%)

rcuc

rcuc

AshAsh

AshAshBurnout (8.1)

Coal VM % (db) FC % (db) Ash % (db) Fuel ratio

(FC/VM)

HHV

kcal/kg (db)

Coal A 12.65 78.74 8.61 6.22 7,970

Coal B 12.69 78.93 8.38 6.22 7,890

Coal C 13.70 76.85 9.45 5.61 7,861

Coal D 15.74 74.61 9.65 4.74 7,716

Coal E 18.10 71.78 10.12 3.87 7,580

Coal F 23.71 67.39 8.9 2.84 7,638

Coal G 25.88 64.16 9.96 2.48 7,290

Coal H 32.82 58.46 8.72 2.37 7,180

Coal I 34.87 55.90 9.23 1.55 7,385

Coal J 35.79 55.51 8.7 1.54 7,531

Coal K 39.58 53.96 6.46 1.36 7,522

217

where ucAsh and rcAsh stand for ash contents in the unburned char and the raw coal,

respectively. Generally, the ash content analyses of the coal and unburned char were

carried out two times in this study, and the average value was taken if the difference

between two analyses was lower than 0.15%. Physically, by virtue of high reactivity of

volatile matter compared to that of fixed carbon, a coal with low fuel ratio represents

that the reactivity of the coal is high. Alternatively, the combustion efficiency or

burnout of a coal with high fuel ratio is usually lower than that with high fuel ratio. In

the present study, the burnout of the coals with the fuel ratio varying from 1.36 to 6.22

was tested individually under the base experimental conditions. A linear correlation

(R2=0.91) between the burnout and the fuel ratio was apparently exhibited, as shown in

Figure 8.3. This reflects that the burnout was increased with the decrease of the fuel

ratio. However, two of the coals with high fuel ratio, namely, Coal A and Coal C,

deviated from this rule. The fuel ratio (FR) of Coal A (FR=6.22) is much higher than

that of Coal D (FR=4.72), however, the burnout of the former (59%) was fairly close to

that of the latter (61.9%). Alternatively, the fuel ratio of Coal C is close to that of Coal

A, the burnout of the former (45.1%) was by far lower than that of the latter (59%). This

may be attributed to the effects of maceral and mineral contained in the coals rather than

the fuel ratio of the coals (Carpenter and Skorupska, 1993). For example, the maceral

analysis of Coal A, as shown in Table 8.2, indicated that the vitrinite, inertinite and

liptinite were 57.2%, 38.8% and 0.4%, respectively, whereas in Coal B they were

42.3%, 57.7% and 0%, respectively. Because the reactivity of vitrinite was larger than

the other two components and the vitrinite percentage in Coal C was especially low

compared to other coals, this resulted in the low burnout of Coal C, as observed. Seeing

that the combustion of some coals with exceptional performance was found from the

experiments, the evaluation of coal combustion from the DTF can provide useful

information on coal selection, especially for the coals with high fuel ratio.

218

Table 8.2 Maceral analyses of the investigated coals.

Coal Vitrinite, % Inertinite, % Liptinite, %

Coal A 57.2 38.8 4

Coal B 68.0 32.0 0

Coal C 42.3 57.7 0

Coal D 50.6 49.4 0

Coal E 49.0 51.0 0

Coal F 53.4 33.8 12.8

Coal G 66.6 30.8 2.6

Coal H 69.5 22.4 8.1

Coal I 68.0 23.0 9.0

Coal J 68.5 17.7 13.8

Coal K 63.2 23.5 13.3

Fuel ratio (-)

Bu

rno

ut

(%)

0 1 2 3 4 5 6 740

50

60

70

80

90

100

A

B

C

DE

F

G

H

IJ

K

Reaction temperature: 12000C

Particle size: 100-200 mesh

R2=0.91

Figure 8.3 Correlation between burnout and fuel ratio under the standard

combustion conditions.

219

8.3.2 Influences of reaction temperature and particle size

Subsequently, the influence of reaction temperature on the burnout of three selected

coals, Coal C, Coal D and Coal I with the particle size of 100-200 mesh, was examined

in Figure 8.4, where three reaction temperatures of 1000, 1100 and 1200°C were

considered. As seen in Figure 8.4, an increase in the reaction temperature facilitated the

burnout of the tested coals, and the promotion of the burnout on the coals with high fuel

ratio (i.e. Coal C and Coal D) was more pronounced than that on the coal with low fuel

ratio (i.e. Coal I). It follows that increasing blast temperature is an effective

countermeasure to reduce the production of unburned char in the raceway of blast

furnace, especially when the coals with high fuel ratio are blended.

Temperature (0C)

Bu

rno

ut

(%)

1050 1100 1150 1200 1250 1300 135020

30

40

50

60

70

80

90

100

Coal C (FR=5.61)

Coal D (FR=4.74)

Coal I (FR=1.55)


Figure 8.4 Distributions of burnout of Coal C, Coal D and Coal I at various

reaction temperatures.

220

For the purpose of increasing the reaction surface of injected pulverized coals for higher

burnout, in the PCI operation worldwide the raw coal is commonly ground to a size

specified by 70-80% of coal particles passing through a 200 mesh sieve (Jaffarullah and

Ghosh, 2005). On the other hand, the injection of granular coal (10-30% minus 200

mesh) is being applied at British Steel for the reduction of coal preparation energy

(Gathergood and Jukes, 1996). To investigate the effect of particle size on coal

combustion, Coal B, Coal E and Coal I were pulverized into three different size regions,

60-100 mesh, 100-200 mesh and 200-325 mesh, respectively. Figure 8.5 demonstrates

the burnout profiles of the three coals where the reaction temperature was 1200oC. The

results presents that the burnout of the tested coals rose when the size of coal particles

was decreased from 60-100 mesh to 100-200 mesh. However, the coal burnout was

slightly decreased when the finer coals with the size of 200-325 mesh were tested. In the

study of Chen et al. (2008) analyzing the particle sizes of unburned char after coal

devolatilization tests, it was reported that when the feed particle size was in the range of

100-200 mesh, the unburned char particles were characterized by swelling, partial

fragmentation and particle agglomeration. Alternatively, as the feed particle size was

200-325 mesh, particle agglomeration was the dominant mechanism of the unburned

char formation. In contrast to the present results, it was presumed that the particle

agglomeration became significant when the fine coals (200-325 mesh) were tested. As a

result, the coal combustion could not be improved further, as observed. This further

implies that the excessive grinding may be avoided in PCI operation, especially when a

high injection rate or dense phase transportation is practiced.

If one further examines the burnout percentages shown in Figures 8.4 and 8.5, it should

be pointed that the burnout difference between different coals tends to be enlarged when

the fuel ratio of coal becomes small, regardless of the variation of reaction temperature

or particle size. Specifically, in Figure 8.4 the burnout difference between Coal I and

Coal D is more pronounced than that between Coal C and Coal D, and in Figure 8.5 the

burnout difference between Coal I and Coal E is more obvious compared to that

221

between Coal E and Coal B. Physically, the lower the fuel ratio of a coal, the higher the

volatile matter contained in the coal. It is known that the reactivity of volatile matter is

generally much higher than that of fixed carbon. A coal with low fuel ratio can be

reacted and depleted easily compared to that with high fuel ratio. This is the reason that

the burnout difference between different coals is amplified as the fuel ratio of a coal

becomes small.

8.3.3 Burnout of blended coals

Upon inspection of Figure 8.3, significant differences in the combustion efficiency

among the coals have been found. It follows that the stability of the lower zone of a

blast furnace may be affected when coal blends are changed. Therefore, the experiments

of coal blend combustion were carried out to figure out the burning behavior of blended

coals. In this study, Coal K, a coal with low fuel ratio, was individually mixed with

Coal A and Coal C, characterized by high fuel ratios, at various blending ratios.

Coal particle size (mesh)

Bu

rno

ut

(%)

20

40

60

80

100

60-100 200-325100-200

Coal I (FR=1.55)

Coal B (FR=6.22)


Coal E (FR=3.87)

Figure 8.5 Distributions of burnout of Coal B, Coal E and Coal I at various

particle sizes.

222

Figure8.6 depicts that the burnout of the coal blends was linearly predictable from the

combustion performance of individual coals. In other words, no synergistic effect of

combustion (Chen and Wu, 2009) was obtained when two different coals were mixed

and burned. It is thus pointed out that the blending ratio of a PCI coal may be adjusted

from the burnout of individual coals given by the DTF rather than by the fuel ratio. On

the other hand, in examining the distributions shown in Figure 8.6, the coal blends

containing Coal A reached higher burnout than that containing Coal C. Taking the

advantage of the better combustion efficiency from the blends with Coal A, a higher

blending ratio of Coal A can be arranged for the increase of calorific value of a PCI

coal. This implies that the coke replacement ratio in an ironmaking process may be

promoted. Notably, this coincides with the operation experience at CSC’s blast

furnaces. It is thus concluded that the DTF system developed in this study is capable of

provide a practical evaluation for the selection of PCI coals employed in blast furnaces.

Blending ratio (%)

Bu

rno

ut

(%)

0 20 40 60 80 10020

40

60

80

100

Blended with Coal C

Blended with Coal A



Pure Coal K

Figure 8.6 Distributions of burnout with respect to blending ratio for Coal K

individually blended with Coal A and Coal C.

223

8.4 Conclusions

A variety of coals and coal blends have been tested using a developed drop tube furnace

to evaluate the combustion characteristics of the coals applied for PCI in blast furnaces.

The present study also considered the effects of reaction temperature, particle size and

blending ratio of two different coals on fuel burnout. In the individual coal tests, it was

found that a coal with higher fuel ratio had a lower value of burnout. However, one coal

with the fuel ratio of 6.22, denoted by Coal A, was characterized by its superiority in

combustion. On the contrary, the burnout of another coal with the fuel ratio of 5.61,

designated by Coal C, was significant lower than the general trend. This indicates that

the combustion behavior of coal depends not only on the fuel ratio but also on the coal

nature. The burnout of a coal could be enhanced by increasing the reaction temperature

or reducing the particle size of the coal. Nevertheless, when the particle size of the coal

was reduced from 100-200 mesh to 200-325 mesh, the combustion efficiency of the

coal could not be improved any more. It can be explained by particles agglomeration

happened when the particle size was small to a certain extent. The burnout of coal

blends was linearly predictable from the combustion efficiency of individual coals in

the blends. Therefore, the blending ratio of a PCI coal may be adjusted from the burnout

of individual coals given by the DTF rather than by the fuel ratio. From the burnout

tests, the DTF system developed in this study has been proved to be capable of

providing a useful evaluation for PCI coals utilized in blast furnaces.

224

CHAPTER 9

PRETREATMENT OF BIOMASS BY TORREFACTION AND

CARBONIZATION FOR COAL BLEND USED IN PULVERIZED

COAL INJECTION

This chapter presents a fundamental insight into the pre-treatment of biomass and the

combustion characteristics of pulverised biofuels under conditions pertinent to the

raceway of blast furnace.




225

ABSTRACT

To evaluate the utility potential of pretreated biomass in blast furnaces, the fuel

properties, including fuel ratio, ignition temperature, and burnout, of bamboo, oil palm,

rice husk, sugarcane bagasse, and Madagascar almond undergoing torrefaction and

carbonization in a rotary furnace are analyzed and compared to those of a high-volatile

coal and a low-volatile one used in pulverized coal injection (PCI). The energy densities

of bamboo and Madagascar almond are improved drastically from carbonization,

whereas the increase in the calorific value of rice husk from the pretreatment is not

obvious. Intensifying pretreatment extent significantly increases the fuel ratio and

ignition temperature of biomass, but decreases burnout. The fuel properties of pretreated

biomass materials are superior to those of the low-volatile coal. For biomass torrefied at

300 °C or carbonized at temperatures below 500 °C, the pretreated biomass can be

blended with coals for PCI.

Keywords: Torrefaction and carbonization; Biomass and biochar; Burnout and

ignition; Rotary furnace; Pulverized coal injection (PCI); Blast furnace.

226

9.1 Introduction

Biomass is able to fix atmospheric carbon while it grows; therefore, biomass is regarded

as a carbon-neutral fuel when it is burned. For this reason, using biomass as an

alternative fuel to fossil fuels is considered as an effective countermeasure to reduce

anthropogenic carbon dioxide emissions into the atmosphere and mitigate global

warming (Machado et al., 2010). For example, bioethanol and biodiesel have been

extensively employed for power generation in spark ignition engines and compression

ignition engines, respectively (Gustavo et al., 2013). In addition to the liquid biofuels,

biomass can also be combusted directly to get heat and power. Compared to coals, the

energy density of biomass is low and its moisture content is high (Rousset et al., 2011).

Moreover, more energy will be consumed to comminute biomass due to its

lignocellulosic nature (van der Stelt et al., 2011). These characteristics limit the

applications of biomass in industry.

As far as blast furnaces are concerned, coke, produced from metallurgical coal, is an

essential reducing agent and provides thermal energy for hot metal production (Du and

Chen, 2006). By means of the technique of pulverized coal injection (PCI), non-coking

or weakly coking coals are injected into the raceways of blast furnaces to partially

replace coke (Chen et al., 2007; Du et al., 2007). On account of mass consumption of

coals for cokemaking and PCI in blast furnaces, a large amount of CO2 is emitted from

the ironmaking processes (Wang et al., 2009). Solid biomass is a potential substitute to

coals and can be partially used for PCI without net carbon dioxide emissions into the

atmosphere (Chen and Wu, 2009). However, due to the disadvantages of raw biomass

described earlier, the upgrade of raw biomass is necessary for its application in blast

furnaces.

The upgrade of biomass can be fulfilled via torrefaction and carbonization or pyrolysis

where biomass is thermally degraded in an inert or oxygen-free environment. The

torrefaction temperature is in the range of 200-300 °C (Peng et al., 2013; Lu et al.,

2012; Sabil et al., 2014), whereas carbonization is operated at temperatures of 300-500

227

°C (Abdullah and Wu, 2009). The biomass materials pretreated from torrefaction and

carbonization are called torrefied biomass and biochar, respectively. By virtue of the

partial disruption of lignocellulosic structure in biomass from the two methods, biomass

grindability is improved greatly (Aris et al., 2008). Grinding coal for PCI is an

energy-intensive process. Therefore, the energy for grinding coal can be saved if

torrefied biomass and biochar are used for PCI (Abdullah and Wu, 2009; Phanphanich

and Mani, 2010). Torrefaction and carbonization lead to the release of volatile matter

from biomass and change the hygroscopic material to hydrophobic one. This

transformation improves the reactivity of solid biomass. Bridgeman et al. (2008) studied

raw and torrefied willow exposed to a methane-air flame, and found that the latter was

ignited more quickly than the former. Pimchuai et al. (2010) investigated rice husk

reaction in a spout-fluid bed combustor, and reported that torrefied rice husk ignited

faster and raised the bed temperature to a higher level when compared to raw rice husk.

These ignition observations were very likely due to the low moisture content in the

torrefied willow and rice husk.

When fuel particles are injected into blast furnaces, they proceed from blowpipes,

tuyeres, and then to raceways, and experience rapid heating, devolatilization, gas-phase

combustion, char combustion, and gasification (Hutny et al., 1991; Shen et al., 2009a;

Wijayanta et al., 2014). Devolatilization and gas-phase combustion correspond to the

mass transfer and reactions of volatile matter from fuel particles, while char combustion

and gasification account for the reactions of fixed carbon. Accordingly, particle

reactions are highly related to the volatile matter and fixed carbon contents in the fuels.

The ignition temperature of volatile is much lower than that of char. Therefore, the first

stage of fuel particle reactions is triggered by volatile ignition, while char combustibility

is subject to its residence time in the reactor and the surrounding temperature (Du et al.,

2010). However, after biomass is torrefied or pyrolyzed, part of the volatiles are

liberated from the material and relatively more fixed carbon is retained (Chen et al.,

2012). This may lower the ignition temperature of biomass in the gas phase.

228

Coals with high fuel ratios are frequently blended with low fuel-ratio coals to increase

the flexibility of PCI operation (Du et al., 2010). When biomass is used as an alternative

fuel to coals for PCI, its utility can be evaluated through a number of properties, such as

fuel ratio, ignition temperature, and burnout (Gao and Bian, 2013; Li, et al., 2014). To

the authors’ knowledge, the pretreatment of biomass simultaneously covers torrefaction

and carbonization has not been studied yet. The purposes of the present study are to

examine the fuel properties of biomass pretreated by torrefaction and carbonization and

compare to those of a high-volatile coal and a low-volatile coal. Particular emphasis will

be paid to the applications of upgraded biomass, from the viewpoint of coal blend used

in PCI.

9.2 Experimental

9.2.1 Materials and preparation

Five different biomass materials were studied in the present work; they are bamboo, oil

palm, rice husk, sugarcane bagasse (abbreviated by bagasse), and Madagascar almond.

The bamboo, rice husk, bagasse, and Madagascar almond were obtained in Taiwan. The

oil palm was the fiber fraction left after the nut was removed in a Malaysia oil

extraction mill. Oil palm is an important economic crop in some countries, especially in

Malaysia. Oil palm fibers are abundant wastes from palm oil fruit harvest and oil

extraction processing. The fibers are considered as a potential renewable energy source

due to its high calorific value and quantity (Shuit et al., 2009). Therefore, oil palm fiber

is adopted and studied in the present study.

Meanwhile, a high-volatile bituminous coal (Coal A) and a low-volatile coal (Coal B)

for PCI operation at China Steel Corporation (CSC) were tested for comparison. The

basic properties of the coals and biomass materials, such as proximate, elemental, fiber,

and calorific analyses, are given in Table 9.1. The proximate analysis was performed in

accordance with the standard procedure of the American Society for Testing and

Materials (ASTM E870-82). The elemental analysis was carried out by use of an

229

elemental analyzer (Vario EL III). The fiber contents (hemicellulose, cellulose, and

lignin) in biomass were analyzed through the measurements of neutral detergent fiber,

acid detergent fiber, and ash (Chen, et al., 2010b). The higher heating values (HHVs) of

samples were detected by a bomb calorimeter (IKA C2000 Basic). As shown in the

Table, the higher heating values (HHVs) and fixed carbon values of the five biomass

species are in the range of approximately 17-19 MJ kg-1

and 9-20 wt%, respectively,

which are much lower than those of Coal A (i.e. 23.99 MJ kg-1

and 40.47 wt%) and

Coal B (i.e. 31.01 MJ kg-1

and 76.35 wt%). Bamboo, oil palm, and Madagascar almond

were shredded by a cutting shredder to the particle sizes of 5-10 mesh (i.e. 2-4 mm),

whereas rice husk and bagasse were not treated. The samples were stored in a nitrogen

atmosphere at 75 ºC until biomass pretreatment was performed.

Table 9.1 Proximate, elemental, fiber, and calorific analyses of two coals and raw

biomass materials.

Coal A Coal B Bamboo Oil

palm

Rice

husk

Bagasse Madagascar

almond

Proximate analysis (wt%)

Moisture 13.92 1.09 5.76 7.20 8.00 7.03 10.17

VM 44.09 14.67 78.76 67.25 73.18 75.03 70.38

FC 40.47 76.35 14.40 19.03 9.27 13.61 18.62

Ash 1.52 7.89 1.08 6.52 9.55 4.33 0.83

Elemental analysis (wt%)

C 63.72 83.24 48.64 44.81 43.40 46.38 47.68

H 4.40 3.78 5.64 4.10 4.33 4.68 4.31

N 0.67 1,62 0.52 2.10 0.65 0.50 0.50

S 0.10 0.52 0.03 0.24 0 0 0

O * 29.59 1.86 44.09 42.23 42.07 44.11 46.68

Fiber analysis (wt%)

Hemicellulose 20.38 34.00 21.34 30.59 18.23

Cellulose 39.82 26.78 36.06 45.66 41.86

Lignin 12.16 16.08 21.16 19.38 16.17

Others 27.64 23.14 41.44 5.37 23.74

Higher heating value (MJ kg-1

)

23.99 31.01 18.95 17.12 17.46 18.31 17.32

* By difference

230

9.2.2 Burnout and ignition tests

In addition to the aforementioned basic analyses, the fuel properties of the pretreated

biomass samples were also analyzed through burnout and ignition tests. The burnout of

the samples was tested by a drop tube furnace. The details of the structure and

experimental procedure of the drop tube furnace can be found elsewhere (Chen et al.,

2012). In brief, the pretreated samples were ground into powders by a shedder followed

by sieving using vibrating screens to the particle sizes of 100-200 mesh (i.e. 74-149

μm). The particles were introduced into the drop tube furnace by nitrogen at a flow rate

of 2 L min-1 (25°C), while the secondary gas, namely, air, at a flow rate of 3 L min-1

was preheated to 350°C and blown into the furnace, acting as an oxidizing agent. The

feeding rate of the particles was approximately 5 g h-1 and the reaction temperature in

the furnace was 1000 °C. The combustibility or combustion efficiency of a solid fuel

can be measured in terms of burnout. The ash tracer method (Du et al., 2010; Chen et

al., 2012) was adopted to determine the burnout of a tested fuel. The burnout is

expressed as

010

100

Ash1Ash

AshAsh(%)Burnout

rawuc

rawuc

(9.1)

where Ashuc and Ashraw designate the ash contents in the unburned char and raw fuel,

respectively.

With regard to the ignition test, the biomass samples were mixed with sodium nitrite

(NaNO2) at a weight ratio of 1:075 and the total amount was 0.2 g in each test. The

sodium nitrite was used as an ignition booster for the ignition test. The mixture was

placed in an ignition analyzer which was heated at a heating rate of 95 °C min-1 under

the open atmosphere. The heating temperature was recorded by a computer, while the

ignition process was shot by a high-speed camera at a frequency of 30 Hz and the

images were stored in the computer. The brightness in the images rose rapidly when an

231

ignition occurred. From the recorded temperature and images, the ignition temperature

of a sample was identified.

To ascertain the analysis quality, the drop tube furnace was leak-tested using nitrogen,

and the axial temperature distribution in the furnace was measured. The measurements

indicated that no leakage occurred in the drop tube furnace, and the temperature

distribution was fairly uniform at the center of the furnace, revealing the spatially

homogeneous gas temperature in the reaction zone. The burnout experiments were

carried out at least twice for some tested samples to ensure the experimental stability

and accuracy. The relative error between each run was controlled within 5%. Moreover,

the ignition temperature of a sample was defined from the average temperature of nine

tests where the deviation of temperature was control within 3 °C.


9.3.1 Proximate analysis and van Krevelen diagram

The profiles of the volatile matter (VM), fixed carbon (FC), and fuel ratio (=FC/VM) of

the five biomass species before and after pretreatment are displayed in Figure 9.1. As a

whole, VM linearly decrease with increasing pretreating temperature (Figure 9.1a),

whereas FC linearly increases (Figure 9.1b). Similar behavior has been observed in the

studies of Couhert et al. (2009) and Lu et al. (2012). Because of this, increasing

pretreating temperature has a trend to raise the fuel ratio exponentially (Figure 9.1c).

When the biomass materials are pretreated at 300 °C or higher temperatures, their FC

values and fuel ratios are larger than those of Coal A (i.e. 40.47 wt% and 0.92).

Alternatively, the VM contents in the samples are lower than that in Coal A (=44.09

wt%) when the temperature is as high as 400 °C. In contrast, the VM contents of the

pretreated materials are always higher than that of Coal B (=14.67 wt%). The FC

contents and fuel ratios of the pretreated materials are lower than those of Coal B (i.e.

76.35 and 5.20), except for bamboo carbonized at 500 °C. From the viewpoint of

proximate analysis, the obtained biomass materials from the pretreatment at

232

temperatures of 300-500 °C are between a high-volatile bituminous coal and a

low-volatile bituminous coal. FC in a fuel gives a higher contribution on heat release

than VM when the fuel is burned (Parikh et el., 2005). This implies, in turn, that the

higher the fuel ratio in biomass, the higher the calorific value of the biomass.

Accordingly, the thermal pretreatment should be operated at temperatures equal to or

higher than 300 °C, from the viewpoint of coal replaced by biomass.

The profiles of atomic H/C ratio versus atomic O/C ratio, namely, the van Krevelen

diagram, of the investigated materials are plotted in Figure 9.2. The linear regression of

the data is also shown in the figure. The coefficient of determination (R2) of the linear

regression is 0.9312. This indicates that there exists a strong linear correlation between

the atomic H/C ratio and the atomic O/C ratio, whether the biomass is torrefied or

carbonized. The slope of the regression line is 1.681, revealing that the impact of the

thermal pretreatment on the H/C ratio is higher than on the O/C ratio. In other words,

more hydrogen is depleted from the thermal degradation when compared to oxygen.

This behavior is different from the observations of Rousset et al. (2011) and Chen et al.

(2012) where only torrefaction was practiced. It follows that the influence of higher

pretreating temperatures (i.e. carbonization) on hydrogen is more than on oxygen. In

light of the profiles shown in Figure 9.2, the van Krevelen diagram can be partitioned

into three different regions. For the raw biomass materials, the atomic O/C and H/C

ratios are larger than 0.53 and 1.17, respectively. After undergoing torrefaction, the O/C

and H/C ratios of the materials are in the ranges of approximately 0.16-0.53 and

0.65-1.17, respectively. When the samples are carbonized, the atomic O/C and H/C

ratios are smaller than 0.16 and 0.65, respectively. The atomic H/C and O/C ratios of

Coal A are 0.83 and 0.31, respectively, which are located in the torrefaction regime.

Alternatively, the H/C and O/C ratios of Coal B are 0.54 and 0.02, respectively,

situating in the carbonization regime.

233

Temperature (oC)

Fix

ed

ca

rbo

n(w

t%)

0

20

40

60

80

100

Raw 250 300 400 450 500

(b)

Temperature (oC)

Vo

latile

ma

tte

r(w

t%)

0

20

40

60

80

100

BambooOil palmRice huskBagasseMadagascar almond

Raw 250 300 400 450 500

(a)

Coal B

Coal A

Temperature (oC)

Fu

elra

tio

(=F

C/V

M)

0

1

2

3

4

5

6

Raw 250 300 400 450 500

(c)

Figure 9.1 (a) Volatile matter, (b) fixed carbon, and (c) fuel ratio values of raw and

pretreated biomass materials.

234

Atomic O/C ratio

Ato

mic

H/C

ratio

0 0.1 0.2 0.3 0.4 0.5 0.6 0.70

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

Bamboo

Oil palm

Rice husk

Bagasse

Madagascar almond

Coal A

Coal B

y = 1.681x + 0.305R

2= 0.9312

Raw

Carbonization

Torrefaction

9.3.2 Solid yield and energy yield

The profiles of the HHVs of the coals and biomass are displayed in Figure 9.3a. When

the biomass is torrefied, the HHV is intensified obviously. In contrast, the increment in

the HHV of the biomass under carbonization is not as significant as that under

torrefaction, especially for rice husk. The HHV of pretreated rice husk is always lower

than that of Coal A; hence, rice husk is not recommended to replace coals used in blast

furnaces. The ash contents of the pretreated biomass samples are given in Table 9.2.

When rice husk is thermally pretreated from 250 to 500 °C, the ash content goes up

from 13.54 to 26.60 wt%. In view of the high ash contents in the raw and pretreated rice

husks, their calorific values are always lower than that of Coal A. Alternatively, the ash

contents in bamboo and Madagascar almond merely increase from 1.08 to 3.71 wt% and

from 0.83 to 3.78 wt%, respectively. It has been addressed that the thermal degradation

of fibrous biomass (i.e. oil palm, rice husk, and bagasse) is inherently different from

Figure 9.2 Atomic H/C versus O/C ratio (van Krevelen diagram) of raw and

pretreated biomass materials.

235

that of ligneous biomass (i.e. bamboo and Madagascar almond) (Chen et al., 2013). On

the other hand, hemicellulose is the most active constituent in biomass (Chen and Kuo,

2010a) which will be thermally decomposed drastically from the torrefaction and

carbonization. Table 9.1 suggests that the hemicellulose contents in raw bamboo and

Madagascar almond are 20.38 and 18.22 wt%, which are relatively lower than the other

species. This is the reason that carbonization is still able to effectively intensify the

calorific values of bamboo and Madagascar almond. Meanwhile, it can be seen that the

HHVs of the two species pyrolyzed at 500 °C are almost equivalent to that of Coal B.

The intention of the present study is to evaluate the potential of torrefied or carbonized

biomass for the replacement of coals consumed in blast furnaces. A physical scale of the

replacement ratio is defined as the HHV ratio between pretreated biomass and a coal.

Consequently, the calorific value of a fuel is higher than that of a coal when the ratio is

larger than unity. Overall, the replacement ratios based on Coal A and Coal B are in the

ranges of 0.71-1.32 (Figure 9.3a) and 0.55-1.02 (Figure 9.3b), respectively. When the

five biomass materials are torrefied at 250 °C, their replacement ratios based on Coal A

are always lower than unity. The replacement ratios of bamboo and Madagascar almond

are larger than unity when they are torrefied at 300 °C, whereas the ratios of the other

three species are smaller than unity. After undergoing carbonization, the replacement

ratios of the samples are higher than unity, with rice husk giving an exception. From the

calorific view of point, the biomass materials except for rice husk can be employed to

replace high-volatile bituminous coals when they are carbonized. Alternatively, only

bamboo and Madagascar almond carbonized at 500 °C can be utilized to replace Coal B

or low-volatile bituminous coals.

The enhancement factors of the HHVs of the five biomass materials are tabulated in

Table 9.3. The factor stands for the HHV ratio between the pretreated biomass and its

parent biomass. Rice husk has the lowest enhancement factor among the five species, as

a result of high ash content (Table 9.2). Alternatively, the enhancement factors of

bamboo and Madagascar almond are relatively high, especially for the latter, which is

236

lifted up to 1.79 when the pretreating temperature is 500 °C. This can be explained by

the low ask contents in the two torrefied materials (Table 9.2). The profiles of the solid

yield and energy yield of the biomass at various pretreating temperatures are

demonstrated in Figure 9.4. The former and the latter are the weight ratio and the energy

ratio between the pretreated biomass and the raw one, and the energy yield is the

multiplication of the solid yield and the enhancement factor of HHV (Lu et al., 2012).

The solid yield is lower than 50% when the biomass is carbonized, regardless of which

sample is pretreated. This reflects that over 50 wt% of raw biomass is consumed from

biochar production at temperatures higher than 400 °C. Though the HHV of

Madagascar almond is promoted markedly from carbonization (Figure 9.3a), it is

noteworthy that its solid yield is lessened obviously, ranging from 22 to 25 wt%. This

behavior is similar to the pyrolysis of mallee wood (Abdullah and Wu, 2009). As a

consequence, only around 40% of energy is retained in the biochar carbonized from

Madagascar almond. The energy yield of rice husk is also lower than 50% from its

carbonization. In contrast, bagasse has the highest energy yield at temperatures of

300-450 °C. This can be explained by the highest cellulose content in bagasse (Table

9.1) among the five raw biomass species.

Table 9.2 Ash contents in pretreated biomass materials.

Temperature (°C) Bamboo Oil palm Rice husk Bagasse Madagascar

almond

250 1.48 10.35 13.54 5.56 1.21

300 2.02 13.12 23.68 6.33 1.79

400 2.69 14.36 24.68 8.61 3.35

450 3.11 15.18 25.41 9.19 3.54

500 3.71 17.85 26.60 13.25 3.78

237

Temperature (oC)

HH

V(

MJ

/kg

)

15

20

25

30

35

Bamboo

Oil palm

Rice husk

Bagasse

Madagascar almond

Raw 250 300 400 450 500

Coal B

(a)

Coal A

Temperature (oC)

Re

pla

ce

me

nt

ratio

0.4

0.6

0.8

1

1.2

1.4

1.6

(b)

Coal A

Raw 250 300 400 450 500

Temperature (oC)

Re

pla

ce

me

nt

ratio

0.4

0.6

0.8

1

1.2

1.4

1.6

(c)

Coal B

Raw 250 300 400 450 500

Figure 9.3 (a) HHVs and replacement factors of biomass materials based on (b)

Coal A and (c) Coal B.

238

9.3.3 Ignition and burnout

The profiles of the ignition temperatures of the coals and raw/pretreated biomass

samples are shown in Figure 9.5. VM is the main factor triggering biomass ignition due

to the gas-phase reaction in a high temperature environment. Without the consideration

of bagasse, the ignition temperatures of the raw biomass materials are in the range of

261-271 °C which are lower than that of Coal A (302 °C) to a certain extent and by far

lower than that Coal B (418 °C), attributing to the high VM contents in the raw biomass

Temperature (oC)

So

lid

yie

ld(%

)

200 250 300 350 400 450 500 5500

20

40

60

80

100

Bamboo

Oil palm

Rice husk

Bagasse

Madagascar almond

(a)

Temperature (oC)

En

erg

yyie

ld(%

)

200 250 300 350 400 450 500 5500

20

40

60

80

100

(b)

Figure 9.4 (a) Solid yield and (b) energy yield of pretreated biomasses materials.

239

samples (Table 9.1). The higher ignition temperature of raw bagasse (315 °C) might be

due to the high cellulose content in the biomass. An increase in pretreating temperature

almost linearly increases the ignition temperature, with the exception of bagasse. As

long as the biomass materials are carbonized, their ignition temperatures are higher than

that of Coal A. This implies that more traveling time for carbonized biomass particles in

raceways is required to activate combustion reaction when compared to Coal A.

Nevertheless, the pretreated biomass can be ignited more easily than that of Coal B.

The profiles of the burnout of the biomass materials versus pretreating temperature are

presented in Figure 9.6a. The burnout has a trend to decrease linearly with increasing

pretreatment temperature. For biomass carbonized under the same temperature, their

burnout values are close to each other, and the values are lower than that of Coal A but

substantially higher than that of Coal B. To keep a constant fuel rate for hot metal

production and reasonable combustion efficiency in the raceways of blast furnaces, the

carbonized biomass examined in this study can be used as the injection fuel, except that

from rice husk. The profiles of burnout versus fuel ratio are shown in Figure 9.6b. In a

previous study (Chen, et al, 2012), a linear correlation between the burnout and the fuel

ratio was exhibited when biomass was torrefied. This implies, in turn, that the

combusting behavior of biomass could be predicted from its fuel ratio, whether the

biomass is torrefied or not. However, when torrefaction and carbonization are

simultaneously considered, as seen in Figure 9.1c, the fuel ratio grows rapidly when the

pretreating temperature increases. The reactivity of VM in an oxidizing environment is

much higher than that of FC, stemming from the gas-phase reaction. Therefore, the

higher the fuel ratio in a fuel, the lower the reactivity or burnout of the fuel is. By virtue

of the significant growth in the fuel ratio with increasing pretreatment severity, the

burnout substantially decays and no linear correlation is observed.

The study of Basu et al. (2013) suggested that mass and energy yields from biomass

torrefaction increased with increasing particle length but decreased with particle

diameter. The analysis of Peng et al. (2012) indicated that the torrefaction rate was

240

affected by the particle size, especially at high temperatures. Accordingly, the particle

size distribution has a significant effect on the torrefaction and pyrolysis processes, even

the ignition and burnout of pretreated biomass. This arises from the fact that heat

transfer and decomposition behavior are related to the particle surface and the total

particle surface depends strongly on the particle size. In the present study, bamboo, oil

palm, and Madagascar almond were shredded by a cutting shredder to the particle sizes

of 5-10 mesh. Though rice husk and bagasse were not treated, their particle sizes, in

fact, are close to those of bamboo, oil palm, and Madagascar almond in the study.

Accordingly, it should be illustrated that the effect of particle size on in this study not

considered. Alternatively, in reviewing recent studies concerning economic analysis, the

overall cost, including the production, transportation and logistics costs, for torrefied

pellets has been evaluated and compared to regular pellets (Bergman, 2005; Peng et al.,

2010). It was reported that torrefied pellets were cheaper than regular pellets. However,

the economic analysis of biomass torrefaction alone is absent so far. Therefore, the

comparison in cost between torrefaction and pyrolysis remains unknown. The topics of

the effect of particle size on torrefaction and carbonization and economic analysis

deserve further investigation in the future.

Temperature (oC)

Ign

itio

nte

mp

era

ture

(oC

)

250

300

350

400

450

BambooOil palmRice huskBagasseMaragascar almond

Coal A

Raw 250 300 400 450 500

Coal B

Figure 9.5 Ignition temperatures of raw and pretreated biomass materials.

241

9.4 Conclusions

The utility potential of torrefied and carbonized biomass materials for PCI in blast

furnaces has been evaluated. A strong linear correlation between the atomic H/C ratio

and O/C ratio in biomass over the operations of torrefaction and carbonization is

exhibited. The fuel properties, such as fuel ratio, burnout, and ignition temperature, of

biomass torrefied at 300 °C or pyrolyzed between 400 and 500 °C, are between a

Temperature (oC)

Bu

rno

ut

(%)

0

20

40

60

80

100

Bamboo

Oil palm

Rice husk

Bagasse

Madagascar almond

Raw 250 300 400 450 500

Coal B

(a)

Coal A

Fuel ratio

Bu

rno

ut

(%)

0 1 2 3 4 5 620

40

60

80

100

Increasingpretretmentseverity

(b)

Coal A

Coal B

Figure 9.6 Burnout versus (a) pretreated temperature and (b) fuel ratio.

242

high-volatile bituminous coal and a low-volatile one. Therefore, the pretreated biomass

can partially replace the coals consumed for PCI and blends with coals to keep

reasonable burnout in raceways.

243

CHAPTER 10

CONCLUSIONS AND RECOMMENDATIONS

This chapter summarises the main achievements of the work and their applications in

the PCI operation at CSC. Besides, the further works for the improvement of PCI are

also recommended.

244

10.1 Introduction

The goal of this thesis is to advance knowledge on the coal combustion in the regions of

blowpipe, tuyere and tuyere of blast furnace through modelling and experimental

studies. The next section summarises the main achievements and conclusions of the

studies. In addition, the applications of the studies in the blast furnaces of China Steel

Corporation are also briefly talked. Further works for improving PCI operation are

recommended in the subsequent section.

10.2 Achievements and conclusions

10.2.1 Modelling

The 3-D CFD coal combustion model has been established in this work to find the

factors influencing the coal combustion for PCI operation and to develop methods to

stabilise the operation of blast furnace raceway. The development of the model

comprised 4 phases:

(1) validation of kinetic parameters of the coal devolatilisation model (Section 2.3.1.2

and Chapters 3 and 4) ;

(2) modelling of coal burning characteristics in the regions of blowpipe and tuyere at

different operation conditions and injection patterns (Chapters 3 and 4);

(3) examination of coal blend combustion in the regions of blowpipe, tuyere and

simplified raceway (Chapter 5); and

(4) development of a comprehensive coal combustion in the blowpipe, tuyere and

raceway featured by a porosity contour of 0.4 in a packed bed (Section 2.4 and

Chapter 6).

To validate the kinetic parameters used in the single and two competing devolatilisation

models, the coal combustion behaviours within the combustion rig (Burgess et al.,

1985) s simulated in the first phase. The RNG k-ε model (gas phase turbulent),

Lagrangian approach (particle trajectory) and the mixture fraction probability density

function model (turbulent combustion) were incorporated into the model. By comparing

245

the predicted temperature distribution with the experimental data of Burgess et al.,

(1985), it is found that the kinetic parameters proposed by Ubhayakar et al., (1976) for

the two competing devolatilisation model is able to sufficiently reflect the pulverised

coal burning characteristics. According to the calculated gas temperature, the PCI coal

combustion can be partitioned into two stages: a very fast step of volatile combustion

followed by a much slower step of char combustion. It suggests that the coal burnout

within the region of blowpipe-tuyere-raceway is largely contributed by the volatile

release rather than by char combustion.

In the second phase, the sensitivity analysis of the operation conditions on the coal

burning in the regions of blowpipe and tuyere was carried out. The simulation suggests

that either increasing the hot blast gas temperature or decreasing the mass flow rate of

carrier gas is able to promote the coal burning within the tuyere. Besides, the

application of the tuyere embedded with ceramic sleeve can be a countermeasure for

enhancing coal burnout. When tracking the trajectories of coal particles, it is observed

that the mixing of the coal particles and hot blast gas played a crucial role in heating,

devolatilisation and combustion of the injected coal particles. In other words, the coal

burnout can be increased when the mixing of coal particles with the hot blast is

improved. From the simulation, earlier ignition was achieved with the operation of

double lance in comparison with the single one. Accordingly, the double air-cooled

coaxial lance system was developed at No3 blast furnace of CSC in 2001. The average

PCI rate of No3 blast furnace was promoted from 110 to 153 kg/tHM within 6 months in

2001. Following successful plant trials, the single lance injection was replaced by the

double air-cooled coaxial lance system in 2002 at the blast furnaces of CSC.

The region of calculation was extended to the jet zone of raceway in the third phase. In

the calculation, the influence of volatile matter content in PCI coal on combustion was

especially examined. Practically, the reduction of volatile matter content in PCI coal can

be made by blending of low volatile coals into high ones. It was found that when the

coal volatile matter reduced from 35% to 25%, the coal burnout decreased from 81.5%

246

to 63%, leading to more unburnt char generated within the combustion region. The

increase of unburnt char might cause drop in gas permeability in the lower zone of the

blast furnace. On the other hand, the calculation also indicated that the pressure loss

across the combustion region could be effectively abated when the coal bend injection

was employed. With the coal blend injection, the performance of the blast furnace may

be improved as long as the consumption rate of unburnt char exceeds its accumulation

in the lower zone of blast furnace. As confirmed during the plant trials in CSC’s No3

blast furnace, the permeability resistance in the lower zone of the furnace was decreased

with the increase of low volatile coals in the coal blend from 30 to 50%. With this

advantage, the flow rate of hot blast could be increased for promoting its hot metal

production. Accordingly, the PCI operation of CSC’s blast furnaces has been shifted

from high volatile coal injection into coal blend since 2003. At present, the volatile

content of coal blend for PCI operation in CSC’s blast furnaces is kept in the range of

19 to 21%. Further improvement in PCI operation may be achieved by late ignition of

coal within the raceway while maintaining a high coal burnout.

In the last phase, the multi Eulerian-Eulerian multi–fluid model was used to determine

the raceway shape, which was featured by the voidage contour of 0.4. The calculated

raceway shape and gas composition distribution within the raceway agreed well with

the measurements reported by Nogami et al. (2004). Taking the operation conditions of

CSC’s No3 blast furnace as the boundary conditions, the burning characteristics of

pulverised coal within the regions of blowpipe, tuyere and raceway were simulated. It

was found that the coal plume was surrounded by enriched oxygen when oxy-coal lance

was used. Fuel ignition was delayed, as a consequence of the cooling effect of the

enriched oxygen. Therefore, less pressure loss across the combustion region is exhibited

in comparison with single and double lance injections. Most importantly, the coal

combustion was intensified by the enriched oxygen at the downstream of the coal

plume. As a result, the coal burnout was as high as that given by double lance injection.

247

It is concluded the oxy-coal injection could be an effective countermeasure at CSC to

optimise its blast furnaces operating.

10.2.2 Coal combustion experiments

In this research, a drop tube furnace was established to evaluate the combustion

performance of PCI coals and pretreated biomass in an environment with high heating

rate (>104

K/s). The drop tube furnace was employed to study:

(1) volatile release and particle formation characteristics of injected coal under

conditions simulating the PCI operation environments within the raceway (Chapter

7);

(2) the combustion behaviours of the PCI coals used in CSC’s blast furnaces,

considering the effects of fuel ratio (the weight ratio of fixed carbon to volatile

matter), reaction temperature, particle size and blending ratio of two different coals

on fuel burnout (Chapter 8); and

(3) the utility potential of pretreated biomass in the PCI operation in the blast furnaces

of CSC (Chapter 9).

In the first research with the drop tube furnace, the experiments were performed by

dropping coal particles into the drop tube furnace by nitrogen at temperatures between

1100 to 1400oC. An examination of the R-factor (volatile release ratio under rapid

heating condition to ASTM test) reveals that either decreasing the feed coal particle size

or increasing the reaction temperature can efficiently enhance volatiles liberation. With

rapid heating, significant agglomeration of the residual char particles is observed when

the coal particle size was reduced from 100-200 mesh to 200-325 mesh for both high

and low volatile coals tested in the experiments, implying that the reaction time of the

char particles will be elongated. Particle swelling during the coal devolatilisation is

found in all testes, regardless of coal rank and size. Notably, only the residual char

particles given by the coarser low volatile coal are characterised by fragmentation. The

formation of tiny soot particles and tar droplets converted from the release volatile

248

matters is found for the two tested coals. The experimental results indicate that their

production is highly sensitive to the volatile matter in the coal. Comparing the reactivity

of the soot to that of the unburnt char, the former is always lower than the latter. It

implies the utilisation of the injected fuels can be improved when the oxygen level

within the coal plume can be enriched.

In second step, a variety of coals with the fuel ratio (the weight ratio of fixed carbon to

volatile matter in coal ranging from 1.36 to 6.22) are tested by using the drop tube

furnace to evaluate the combustion characteristics of the coals, considering the effects

of fuel ratio, reaction temperature, particle size and blending ratio of two different coals

on fuel burnout. It is found the burnout of tested coals is increased with the decrease of

the fuel ratio. However, a coal with high fuel ratio deviates from this rule, showing

higher burnout than the trend. It is concluded to be contributed by its higher vitrinite

content in the maceral. Accordingly, this coal has become the primary low volatile PCI

coal of the blast furnaces of CSC. The experimental results show higher fuel burnout

can be achieved when the particle size of coal is reduced from 60-100 to 100-200 mesh.

However, once the size of the tested coals is in the range of 200 and 325 mesh, the

burnout can not be improved further, resulting from the agglomeration of fine particles.

According to this result, the coal quantity passing through 200 mesh has been reduced

from 80 to 60% in the CSC’s PCI operation for less energy consumed in coal grinding.

Notably, this test has become a standard procedure at CSC to evaluate the coal

combustion using the drop tube furnace before the coal is injected into the blast

furnaces.

In the third step, the fuel properties, including fuel ratio, ignition temperature, and

burnout, of five torrefied and carbonised biomasses were analyzed and compared to

those of a high volatile coal and a low volatile one used in pulverised coal injection at

CSC. A strong linear correlation between the atomic H/C ratio and O/C ratio in biomass

over the operations of torrefaction and carbonization is exhibited. The fuel ratio,

burnout, and ignition temperature, of biomass torrefied at 300 °C or pyrolysed between

249

400 and 500 °C, are close to those of the high volatile coal. Therefore, the pretreated

biomass can partially replace the high volatile coals in the coal blend for PCI operation.

Although it has been emphasised in many studies that high burnout of pulverised coal in

the regions of blowpipe, tuyere and raceway is required, as revealed in this work,

moderate pressure loss in the raceway is also essential for achieving high injection rate

or high productivity in a blast furnace.

In this work, some results obtained from the comprehensive experimental and numerical

studies have been taken into PCI operation at CSC. This seems as a limitation of the

current study, but it may have a wide range of applications for the improvement of PCI

operation.

10.3 Recommendations

10.3.1 Raceway control

To improve the performance of PCI operation by controlling the raceway, further

research works are recommended below:

(1) Investigation into raceway phenomena by considering the formation of bird’s nest.

In this study, the multi Eulerian-Eulerian multi–fluid model has been used in this

research to determine the raceway shape. As matter of fact, the structure of raceway

is more complicated than that used in the calculation. To investigate the raceway

structure and coke degradation, the sampling of coke at tuyere level has been

performed at CSC’s blast furnaces during scheduled time of shutdowns. Figure 10.1

indicates the coke particles taken from CSC’s No3 in 2013 during scheduled

shutdown. The sampled cokes could be characterised into three distinctive zones in

terms of coke size; bosh coke (0 to 55 cm in depth from wall), raceway coke (105 to

135 cm) and bird’s nest (135 to 160 cm), which makes the raceway boundary tight

and impervious. The penetration of hot blast into the deadman zone of blast furnace

could be adversely affected with an accentuated growth of bird’s nest. As a result,

250

the deadman will become inactive. For a stable operation of blast furnace, research

works related to the structure of raceway are required.

Experimental study on the raceway phenomena is undergoing at CSC using a cold

2-D model. To investigate the formation of raceway cavity and bird’s nest, the

model was filled by mung beans (breakable materials) rather than by rigid particles

in the first stage of experiments (Du and Tsai, 2014). Figure 10.2 shows the

relationship between the raceway shape and the thickness of bird’s nest. As observed

in the experiments, the bean particles circulate along the raceway boundary at a high

velocity, resulting in the generation of particle fines. In Figure 10.2c (elapsed time

of 15 minutes), a dense shell layer is found due to the accumulation of bean debris in

160 cm 135 cm 105 cm 80 cm

80 cm 55 cm 30 cm 0 (Wall)

Figure 10.1 Zonal structures in a drill core.

251

the boundary of the raceway, while the cavity volume increases. At elapsed time of

35 minutes (Figure 10.2d), the raceway is bended upwards, indicating the gas flow

within the raceway is driven towards the wall. It will increase the gas flow in the

periphery, resulted in higher heat load. Obviously, the shell formation and its

thickness play important roles for determining the permeability and gas distribution

in the lower zone of blast furnace. To preciously predict the pressure loss and the

raceway shape, the coke breakage caused by circulation and the combustion of coke

fines trapped in the bird’s nest should be taken into account in the calculation model.

This may be achieved by coupling DEM (discrete element method) with CFD code.

In DEM, the coke particle can be considered as an agglomerate of 2-3 sub particles,

and the coke breakage occurs when the impact force is higher than the bonding

strength among the sub particles.

(2) Modelling of co-injection of oxygen and pulverised coal in double lance operation.

(a) t = 0 (b) t = 2 min

(c) t = 15 min (d) t = 35 min

Injection nozzle

bird nest

Circulating fines

Figure 10.2 Formation of cavity and bird’s nest in a pack bed.

252

For the reduction the thickness of bird’s nest, individually injecting pure oxygen

into the raceway may be an effective way to improve the permeability of the

raceway.

(3) Further case studies on torrefied or carbonised bio-fuel injection in the raceway.

10.3.2 Improvement of permeability by charging pattern

The ore/coke (O/C) of burden in blast furnace increases with the increase of pulverised

coal injection rate due to the relative increase of the ferrous than the coke (Ishii, 2000).

With high O/C, the gas permeability resistance of blast furnace is increased, and the

stability of blast furnace may turn worse. Therefore, proper charging patterns are

required to improve the ascending gas distribution within the furnace. Practically, the

burden profile is considered as essential information for the evaluation of the charging

patterns and the status of the furnace. To measure the burden profile during the

operation, an online monitoring system has been successfully developed and applied at

CSC’s No3 blast furnace (Lu and Du, 2010). Accordingly, the measured profiles have

provided useful basis for determining the charging patterns of the furnace. As indicated

in Figure 10.3, the burden profile with a terrace in the vicinity of wall was found after

changing the first charging angle from 45 to 42.6o (Du et al., 2012). In general, the

burden profile with side terrace is preferred at CSC for better gas distribution and stable

burden movement in the furnace. As a result, the PCI rates of the furnaces were

gradually promoted from 140 to 170 kg/tHM, hitting the monthly highest injection

record (180.6kg/tHM) of CSC in August 2010. To increase PCI rate and fuel efficiency

of blast furnace, technologies on the on-line measurement and control of burden profile

are recommended.

253

Figure 10.3 The measured burden profiles and calculated descending rate at No3

blast furnace.

furnace

254

References

Abdullah, H. and Wu, H. W. (2009), Biochar as a fuel: properties and grindability of

biochars produced from the pyrolysis of mallee wood under slow-heating

conditions, Energy Fuel, vol. 23, p. 4174-81.

Alonso, M. J., Borrego, A .G., Alvarez, D., Kalkreuth, W. and Menendez, R. (2001),

Physicochemical transformations of coal particles during pyrolysis and combustion

Fuel, vol. 80, p. 1857-70.

Anthony, D. B. and Howard, J. B. (1976), Coal devolatilization and hydrogasification.

AIChE Journal, vol. 22, p. 625-656.

Aoki, H., Nogami, H., Suge, T., Miura, T. and Furukawa, T. (1993), Simulation of

transport phenomena around the raceway zone in the blast furnace with and

without pulverized coal injection, ISIJ International, vol. 33, p. 646-654.

Arias, B., Pevida, C., Fermoso, J., Plaza, M. G., Rubiera, F. and Pis, J. J. (2008),

Influence of torrefaction on the grindability and reactivity of woody biomass, Fuel

Processing Technology, vol. 89, p. 169-175.

Ariyama, T, Sato, M., Yamakawa, Y., Yamada, Y. and Suzuki. M. (1994), Combustion

behaviors of pulverized coal in tuyere zone of blast furnace and influence of

injection lance arrangement on combustibility, ISIJ International, vol. 34, p.

476-483.

Ariyama, T., Ueda, S., Natsui, S., Inoue, R. and Sato, M., (2007), Current technology

and future aspect on CO2 mitigation in Japanese steel industry, Proceedings of the

5th International Congress on the Science and Technology of Ironmaking,

Shanghai, China, p. 57-64.

Austin, P. R., Chew, S. J., Maldonado, D., Mathieson, J. G., Pinson, D. J. Rogers, H.

and Walsh, M. (2011), PC Injection studies at BlueScope Steel, Proceedings of the

METEC InSteel Conference, Düsseldorf, Germany, Section 27, p. 1-10.

Babich, A., Senk, D. and Fernandez, M. (2010), Charcoal behaviour by its injection into

the modern blast furnace, ISIJ International, vol. 50, p. 81-88.

Babich, A., Yaroshevskii, A. Formoso, A. Isidro, A., Ferreira, S. Cores, A. and Garcia,

L. (1996), Increase of pulverized coal use efficiency in blast furnace, ISIJ

International, vol. 36, p. 1250-58.

255

Babich, A., Gudenau, H. W., Mavrommatis, K. T., Froechling, C., Formoso, A., Cores,

A. and Garcia, L, (2002), Choice of technological regimes of a blast furnace

operation with injection of hot reducing gases, Revista de Metalurgia, vol. 33, p.

288-305

Badzioch, S. and Hawksley, P. G. W. (1970), Kinetics of thermal decomposition of

pulverized coal particles, Industrial and Engineering Chemistry Process Design

and Development, vol. 9, p. 520-530.

Basu, P., Roa, S. and Dhungana, A. (2010), An investigation into the effect of biomass

particle size on its torrefaction, The Canadian Journal of Chemical Engineering,

vol. 91, p. 466-474.

Bergman, P.C.A. (2005), Combined torrefaction and pelletization the TOP process,

Report ECN-C-05-073, The Netherlands: ECN.

Biswas, A. K. (1981), Principles of blast furnace ironmaking, Cootha Publishing House,

Brisbane, Australia.

Biswas G and Eswaran V. (2002), Turbulent flows: fundamentals, experiments and

modelling, Alpha Science.

Bortz, S. and Flament, G. (1983), Experiments on pulverized-coal combustion under

conditions simulating blast-furnace environments, Ironmaking and Steelmaking,

vol. 10, p. 222-229.

Boubel, R. W., Fox, D. L., Turner, D. B. and Stern, A. C. (1994), Fundamentals of air

pollution, Academic Press, San Diego.

Bridgeman, T.G., Jones, J.M., Shield, I. and Williams, P.T. (2008), Torrefaction of reed

canary grass, wheat straw and willow to enhance solid fuel qualities and

combustion properties, Fuel, vol. 87, p. 844-856.

Burgess, J. M., Jamaluddin, A. S., McCarthy, M. J., Mathieson, J. G., Nomura, M.

Truelove, J. S. and Wall, T. F. (1983), Pulverised coal ignition and combustion in

the blast furnace tuyere zone, Proceedings of Joint Symposium of ISIJ and AIMM,

Tokyo, Japan, p. 129–141

Buss, W. E., Toll, H. and Worberg, R. (2000), Developments in Cokemaking in Europe,

AISE Steel Technology, vol. 77, p. 47-52.

Card, J. B. A., and Jones, A. R. (1995), A drop tube furnace study of coal combustion

and unburned carbon content using optical techniques, Combustion and Flame, vol.

101, p. 539-547.

http://www.google.com.tw/url?url=http://pubs.acs.org/loi/iepdaw&rct=j&frm=1&q=&esrc=s&sa=U&ei=WU8uVeCVLcGm8AWD-IGYAg&ved=0CBMQFjAA&sig2=v6zIUdQOV8Nuuv-PEJekdw&usg=AFQjCNFrrKe7GLGYI75A9KwmbySxLkIBWg

http://www.google.com.tw/url?url=http://pubs.acs.org/loi/iepdaw&rct=j&frm=1&q=&esrc=s&sa=U&ei=WU8uVeCVLcGm8AWD-IGYAg&ved=0CBMQFjAA&sig2=v6zIUdQOV8Nuuv-PEJekdw&usg=AFQjCNFrrKe7GLGYI75A9KwmbySxLkIBWg

http://dx.doi.org/10.1016/0010-2180(94)00237-M

http://dx.doi.org/10.1016/0010-2180(94)00237-M

256

Carpenter, A. M. and Skorupska, N. M. (1993), Coal combustion – analysis and testing,

IEA Coal Research, London.

Chattopadhyay, K., Isac, M. and Guthrie, R. I. L. (2010), Review: Applications of

Computational Fluid Dynamics (CFD) in iron- and steelmaking: Part 1,

Ironmaking and Steelmaking, vol. 37, p. 554-561.

Chen, C., Horio, M., Kojima, T. (2000a), Numerical simulation of entrained flow coal

gasifiers. Part I: modeling of coal gasification in an entrained flow gasifier,

Chemical Engineering Science, vol. 55, p. 3861-74.

Chen, C., Horio, M. and Kojima, T. (2000b), Numerical simulation of entrained flow

coal gasifiers. Part II: effects of operating conditions on gasifier performance,

Chemical Engineering Science, vol. 55, p. 3875-83.

Chen W. H., Du, S. W., Yang, H. H. and Wu, J. S. (2008), Formation characteristics of

aerosol particles from pulverized coal pyrolysis in high-temperature environments,

Journal of Air and Waste Management Association, vol. 58, p. 702-710.

Chen, W. H., Du, S. W., Tsai, C. H., and Wang, Z. Y. (2012), Torrefied biomasses in a

drop tube furnace to evaluate their utility in blast furnaces, Bioresource

Technology, vol. 111, p. 433-438.

Chen W. H., Du, S. W. and Yang, T. H. (2007), Volatile release and particle formation

characteristics of injected pulverized coal in blast furnace, Energy Conversion and

Management, vol. 48, p. 2025-33.

Chen, W. H., Du, S. W., Yang, H. H., Wu, J. S. (2008), Formation characteristics of

aerosol particles from pulverized coal pyrolysis in high-temperature environments,

Journal of the Air & Waste Management Association, vol. 58, p. 702-710.

Chen, W. H. and Jiang T. L. (2000), Double, triple, and tetra-brachial flame structures

around a pair of droplets in tandem, Combustion Science and Technology, vol.

151, p. 105-132.

Chen, W. H. and Kuo, P. C. (2010), A study on torrefaction of various biomass

materials and its impact on lignocellulosic structure simulated by a

thermogravimetry, Energy, vol. 35, p. 2580-2586.

Chen. W. H., Lin, M. R., Leu, T. S. and Du, S. W. (2011), An evaluation of hydrogen

production from the perspective of using blast furnace gas and coke oven gas as

feedstocks, International Journal of Hydrogen Energy, vol. 36, p. 11727-37.

257

Chen, W. H., Lu, K. M., Lee, W. J., Liu, S. H. and Lin, T.C. (2013), Biomass

torrefaction characteristics in inert and oxidative atmospheres at various superficial

velocities. Bioresource Technology, vol. 146, p. 152-160.

Chen, W. H., Tu, Y.J. and Sheng, H. K. (2010), Impact of dilute acid pretreatment on

the structure of bagasse for producing bioethanol, International Journal of Energy

Research, vol. 34, p. 265-274.

Chen, W. H., and Wu, J. S. (2009), An evaluation on rice husk and pulverized coal

blends using a drop tube furnace and a thermogravimetric analyzer for application

to a blast furnace, Energy, vol. 34, p. 1458-1466.

Choi, YC., Li, X.Y., Park, T.J., Kim, J.H. and Lee, J.G. (2001), Numerical study on the

coal gasification characteristics in an entrained flow coal gasifier, Fuel, vol. 80, p.

2193-2201.

Chung, J. K. and Hur, N. S. (1997), Tuyere level coke characteristics in blast furnace

with pulverized coal injection. ISIJ International, vol. 37, p. 119-125.

Couhert, C., Salvador, S. and Commandre, J. M. (2009), Impact of torrefaction on

syngas production from wood, Fuel, vol. 88, p. 2286-2290.

de Lassat de Pressigny, Y., Picard, M., Prado, G., Aleboyeh, H. and Simonin, O. (1990),

Study of coal combustion with respect to blast furnace injection, Proceedings of

Ironmaking Conference, Iron and Steel Society, Detroit, USA, p. 473-480.

Du, S. W. (2001), Development of 3D coal combustion model for PCI operation, CSC’s

internal research report, PJ-90043.

Du, S. W. (2011), Modelling of raceway shape prediction and oxy-coal lance

evaluation, CSC’s internal research report, PJ-100050.


pulverized coal combustion in blast furnace, International Communications in Heat

and Mass Transfer, vol. 33, p. 327-334.

Du, S. W., Chen, W. H. and Lucas, A. J. (2007), Performances of pulverized coal

injection in blowpipe and tuyere at various operational conditions, Energy


Du, S. W., Chen, W. H. and Lucas, A. J. (2010), Pulverized coal burnout in blast


258



Bioresouce Technology, vol. 161, p. 333-339.

Du, S. W., Ho, C. K., Tsai, S. T. and Yeh, C. M. (2001), Development of pulverised

coal injection lance at China Steel Corporation, World Coal, vol. 10, p. 39-42.

Du, S. W., Lu, C. Y., C. Y. and Chou, C. S. (2012), Development and application of

blast furnace burden profile measurement system at CSC, Proceedings of 4th

Scanmet: International Conference on Process Development in Iron and

Steelmaking, Luleå, Sweden, p. 320-329

Du, S. W. and Tsai, C. H. (2014), Investigation into the formation of the raceway of

blast furnace, CSC’s internal research report, RE-102018.

Du, S. W., Yeh, C. M., Ho, C. K. and Yang, M. K. (2004), Practice of high productivity

at No.3 blast furnace of China Steel Corporation, Proceedings of AISTech

Conference, Tennessee, USA, p. 195-204.




Eghlimi, A. and Sahajwalla, V. (1997), Pulverized coal combustion in a 2D

furnace-generalised finite rate chemistry versus mixture fraction/PDF model,

Proceedings of International Conference Fluid and Thermal Energy Conversion,

Indonesia, p. 25-30.

Ergun, S. (1952), Fluid flow through packed columns, Chemical Engineering and

Processing, vol. 48, p. 89-94.

Field, M. A. (1969), Rate of combustion of size–graded fractions of char from a low

rank coal between 1200 K – 2000 K, Combustion and Flame, vol. 13, p. 237-252.

Fletcher, T. H., Ma, J., Rigby, J. R., Brown, A. L. and Webb, B. W. (1997), Soot in coal

combustion system, Progress in Energy and Combustion Science, vol. 23, p.

283-301.

Gao, Y. H. and Bian, L. T. (2013), A evaluation on pulverized coal combustion

properties using a thermogravimetric analyze method, Applied Mechanics and

Materials, p. 423-426, p. 609-613.

259

Gathergood, D. S. and Jukes, M. H. (1996), Blast furnace injection of granular coal,

Proceedings of Injection Technology in Ironmaking and Steelmaking, Brussels,

Belgium, p. 89-107.

Geerds, M., van Laarm R. and Vaynshteyn, R. (2011), Low-cost hot metal: the future of

blast furnace of blast furnace ironmaking, Iron and Steel Technology, vol. 8, p.

51-57.

Gibb, J. (1985), Combustion of residual char remaining after devolatilization. Lecture at

Course of Pulverised Coal Combustion, Imperial College, London.

Gibbins, J. R., Man, C. K. and Pendlebury, K. J. (1993), Determination of rapid heating

volatile matter contents as a routine test, Combustion Science and Technology, vol.

93, p. 349-361.

Gidaspow, D. (1994), Multiphase flow and fluidization: continuum and kinetic theory

descriptions, Academic Press, New York.

Goto, K., Murai, R. Murai, A., Murao, A. Sato, M. Asanuma M. and Ariyama, T.

(2002), Massive combustion technology of solid fuels injected into blast furnace,

Proceedings of International BF Lower Zone Symposium, Wollongong, Australia,

p. 1-10.

Greuel, M., Hillnhütter, F. W., Kister, H. and Kruger, B. (1974), Investigation on

movement mechanisms in front of the tuyere of an industrial blast furnace using an

endoscope, Stahl u. Eisen, vol. 94, p. 533-538.

Gu M. Y., Zhang M. H., Selvarasu, N. K. C., Zhao, Y. F. and Zhou, C. Q. (2008),

Numerical analysis of pulverized coal combustion inside tuyere and raceway, Steel

Research International, vol. 79, p. 17-24.

Gudenau, H. W. and Kiesler, R. (1991), Iron and steel making, especially pulverized

coal injection into the blast furnace, Proceedings of the Symposium on Raceway

Control for Optimum Blast Furnace Performance, McMaster University, Canada,

p. 1-23

Gudenau, H. W., Peters, M. and Joksch, M. (1994), Meβtechnische untersuchungen zur

kohleeinblasung am hochofen, Stahl and Eisen, vol. 114, p.81-86.

Gunn, D. J. (1978), Transfer of heat or mass to particles in fixed and fluidized beds,

International Journal of Heat and Mass Transfer, vol. 21, p. 467–476.

260

Guo B. Y., Zulli P., Rogers, H., Mathieson J. and Yu A. B. (2005), Three-dimensional

simulation of flow and combustion for pulverised coal injection, ISIJ International,

vol. 45, p. 1272-1281.

Guo, H. W., Su, B. X., Zhang, J. L., Shao, J. G., Zuo, H. B. and Ren, S. (2012), Energy

conservation for granular coal injection into a blast furnace, JOM, vol. 64, p.

1002-10.

Guo, X. Z., Wei, G, Zhou, Y. P., Ma, Z. F., Ding, Z. M. and Shen, F. M. (2011), Effect

of MgO addition on burnout rate of pulverized coal for blast furnace injection, Iron

and Steel, vol. 46, p. 7-10.

Guo, Y. C., Chan, C. K. and Lau, K. S. (2003), Numerical studies of pulverized coal

combustion in a tubular coal combustor with slanted oxygen jet, Fuel, vol. 82, p.

893-907.

Gustavo B. Leite, G. B., Abdelaziz, A. E. M., Patrick C. and Hallenbeck, P.C. (2013),

Algal biofuels: challenges and opportunities, Bioresource Technology, vol. 145, p.

134-141.

Hammond, G. P. and Norman, J. B. (2014), Heat recovery opportunities in UK industry,

Applied Energy, vol. 116, p. 387–397.

Hartig, W., Zewe, H., Leyser, P., Cano, S. and Cortina, C., Mahowald, P. and Muller, B.

(2011), Increase of pulverized coal injection at ROGESA, Proceedings of AISTech,

Warrendale, USA, p. 475-486.

Haywood, R. J., Truelove, J. S. and McCarthy, M. J. (1994) Modelling of pulverized

coal injection and combustion in blast furnaces, Proceedings of Ironmaking

Conference, Iron and Steel Society, Chicago, USA, p. 437-442.

He, J. C., Kuwabara, M. and Muchi, I. (1986), Analysis of combustion zone in raceway

under operation of pulverized coal injection, Tetsu to Hagane, vol. 72, p. 1847-54.

Hinrichs, R.A. and Kleinbach, M. (2002), Energy: Its use and the environment (3rd

Ed.),

Harcourt, Inc., Philadelphia.

Ho, C. K. (2000), Development of AE sensor system for measuring burden falling point

in blast furnace, CSC’s internal research report, PJ-89020.

Ho, C. Y. and Du, S. W. (2008), Development of tuyere monitoring system, CSC’s

internal research report, PJ-97075.

https://getinfo.de/app/subject-search?action=search&author=%22Mahowald%2c+P.%22&form=advanced

https://getinfo.de/app/subject-search?action=search&author=%22Muller%2c+B.%22&form=advanced

261

Hsieh, L. H., Kao, T. S., Liu, K. C. and Wang, G. C. (2002), Investigation of low fluxes

in iron ore sintering and its application, Proceedings of Ironmaking Conference,


Hurt, R. and Davis, K. (1999), Percolative fragmentation and spontaneous

agglomeration, Combustion and Flame, vol. 116, p. 662-670.

Hutny, W. P., Lee, G. K. LEE and Price, J. T. (1991), Fundamentals of coal combustion

during injection into a blast furnace, Progress in Energy and Combustion Science,

vol. 17, p. 373-395

Inatani, T., Fukutake, T. and Okabe, K. (1973), Investigations on the coke combustion

and coke size in the raceway, Der Hochofenprozess Hersg., Verein Deutscher

Eisenhüttenletute, Düsserdorf, p. 114-121.

Inatani, T., Okabe, K., Nishiyama, T., Serizawa, Y., Takahashi, H. and Saino M. (1976),

Heavy oil combustion in blowpipe and tuyere, Tetsu to Hagane, vol. 62, p.

514-524.

Ishii, K. (2000), Advanced pulverised coal injection technology and blast furnace

operation, Oxford, UK, Elsevier.

Iwanaga, Y. (1991), Gasification rate analysis of unburnt pulverized coal in blast

furnace, ISIJ International, vol. 31, p. 500-504.

Jaffarullah, R. and Ghosh, B. (2005), Alternate fuel in blast furnace to reduce coke

consumption, I.E (I) Journal-MM, vol. 86, p. 16-23.

Jamaluddin, A. S., Wall T. F., and Truelove, J. S. (1986), Mathematical modeling of

combustion in blast furnace raceways, including injection of pulverized coal,

ironmaking and steelmaking, vol. 13, p. 91-99.

Jamaluddin, A.S., Wall T.F., Truelove, J.S. (1987), Modelling of high intensity

combustion of pulverized coal in a tubular combustor, Combustion Science and

Technology, vol. 55, p. 89-113.

Jianwei, Y., Guolong, S., Cunjiang, K., Tianjun, Y. (2003), Oxygen blast furnace and

combined cycle (OBF-CC) - an efficient iron-making and power generation

process, Energy, vol. 28, p. 825-835.

Jones, W.P. and Whitelaw, J.H. (1982), Calculation methods for reacting turbulent

flows: A review. Combustion and Flame, vol. 48, p. 1-26.

http://www.sciencedirect.com/science/article/pii/S0010218098000613

http://www.sciencedirect.com/science/article/pii/S0010218098000613

262

Kalkreuth, W., Borrego, A. G., Alvarez, D., Menendez, R., Osorio, E., Ribas, M.,

Vilela, A., and Cardozo Alves, T. (2005), Exploring the possibilities of using

Brazilian subbituminous coals for blast furnace pulverized fuel injection, Fuel, vol.

84, p. 763-772

Khalil, E.E. (1982), Modelling of Furnace and Combustor, Abacus Press.

Kim, S. M., Chung, J. K. and Cho, C. M. (1996), Improvement coal combustion

efficiency for high injection rate and its effect on raceway behavior, POSCO

Technical Report, vol. 1, p. 110-114.

Kobayashi, H., Howard, J. B. and Sarofim, A. F. (1977), Coal devolatilization at high

temperatures, Proceeding of 16th Symposium (International) on Combustion, The

Combustion Institute, Pittsburgh, PA, p. 411-425.

Kuo, K. K. (1986), Principles of combustion, John Wiley and Sons, New York.

Kurose, R., Ikeda, M., Makino, H., Kimoto, M. and Miyazaki, T. (2004), Pulverized

coal combustion characteristics of high-fuel-ratio coals, Fuel, vol. 83, p. 1777-85.

Kuwabara, M., Hsieh, Y. S., Isobe, K. and Muchi, I. (1981), Mathematical modelling of

the tuyere combustion zone of the blast furnace, Proceedings of International Blast

Furnace Hearth and Raceway Symposium, Newcastle, Australia, P. 7.1-7.7.

Lai, F. C., Chang, D. Y. and Chen, E. K. (1994), Promotion of PCI rate in China Steel

Corporation’s No. 1 blast furnace, Proceedings of the First International Congress

Science and Technology of Ironmaking, p. 541-546.

Lee, J. G., Kim, J. H., Lee, H. J., Park, T. J. and Kim, S. D., 1996; Characteristics of

entrained flow coal gasification in a drop tube reactor, Fuel, vol. 75, p. 1035-42.

Li, H. Y., Elliott, L. Rogers, H. and Wall, T. (2014), Comparative study on the

combustion performance of coals on a pilot-scale test rig simulating blast furnace

pulverized coal injection and a lab-scale drop-tube furnace, Energy and Fuels, vol.

28, p. 363-368.

Li, R. R., Zhu, J. M. and Li, Y. Q. (2007), Practice of raising Baosteel No. 4 BF’s coal

injection rate rapidly, Baosteel Technology, No. 1, p. 11-14.

http://www.sciencedirect.com/science/article/pii/0016236196000841

http://www.sciencedirect.com/science/article/pii/0016236196000841

263

Lu, K. M., Lee, W. J., Chen, W. H. and Lin, T. C. (2013), Thermogravimetric analysis

and kinetics of co-pyrolysis of torrefied wood and coal blends. Applied Energy,

vol. 105, p. 57–65.

Lu, C. Y. and Du, S. W. (2010), Development of an online blast furnace burden profile

measuring system, China Steel Technical Report, no. 23, p. 25-30.

Lu, K. M., Lee, W. J., Chen, W. H., Liu, S. H. and Lin, T. C. (2012), Torrefaction and

low temperature carbonization of oil palm fiber and eucalyptus under nitrogen and

air atmospheres, Bioresource Technology, vol. 123, p. 98-105.

Machado, J. G. M. S., Osorio, E., Vilela, A. C. F., Babich, A., Senk, D. and Gudenau,

H. W. (2010), Reactivity and conversion behaviour of Brazilian and imported

coals, charcoal and blends in view of their injection into blast furnaces, Steel

Research International, vol. 81, p. 9-16.

Magnussen, B. F. and Hjertager, B. W. (1976), On mathematical modelling of turbulent

combustion with special emphasis on soot formation and combustion, Proceedings

of 16th

Symposium (International) on Combustion. Pittsburgh, The combustion

institute.

Malgarini, M. G. (1991), Direct use of coal in ironmaking, Proceedings of World Coal

Institute Conference and Exhibition: Coal in the Environment, London, p. 367-368.

Maloney D. J. and Jekins, R. G. (1984), Coupled heat and mass transport and chemical

kinetic rate limitations during coal rapid pyrolysis, Proceedings of 20th

Symposium

(International) on Combustion, The Combustion Institute, Pittsburgh, PA, p.

1435-43.

Maki, A., Sakai, A., Takagaki, N., Mori, K., Ariyama, T., Sato, M. and Murai, R.

(1996), High rate coal injection of 218 kg/t at Fukuyama No. 4 blast furnace, ISIJ


Mathieson, J. G., Rogers, H. Somerville, M. A. and Jahanshahi, S. (2012), Reducing net

CO2 emissions using charcoal as a blast furnace tuyere injectant, ISIJ international,

vol. 52, p. 1489-96.

Mathieson, J. G., Truelove, J. and Rogers, H. (2005) Towards an understanding of coal

combustion in blast furnace tuyere injection, Fuel, vol. p. 1229–37.

McCarthy, M. J. (1986), Replacement of oil by coal injection at the blast furnaces, BHP

Research Report, NERDDP/EG/86/602.

264

Mondal, S. S., Som, S. K. and Dash, S. K. (2005), Numerical predictions on the

influences of the air blast velocity, initial bed porosity and bed height on the shape

and size of raceway zone in a blast furnace, Journal of Physics D–Applied Physics,

vol. 38, p.1301–07.

Morsi, S. A. and Alexander A. J. (1972), An investigation of particle trajectories in

two–phase flow system, Journal of Fluid Mechanics, vol. 55, p. 193–208.

Myers, G.E. (1971), Analytical Methods in Conduction Heat Transfer, McGraw-Hill,

New York.

Nakamura, N., Togino, Y. and Tateoka, M. (1978), Behavior of coke in large blast

furnaces, Ironmaking and Steelmaking, vol. 5, p. 1–17.

Narita, K. Maekawa, M. Kanayama, H. Seki, Y. And Saito, T. (1982), Combustion of

pulverised coal in experimental furnace, Tetsu-to-Hagane, vol. 68, p. 2385-92.

Niksa, S. Heyd, L. E., Russel, W. G. and Saville, D. (1984), On the role of heating rate

in rapid coal devolatilization, Proceedings of 20th

Symposium (International) on

Combustion. The Combustion Institute, Pittsburgh, PA, p. 1445-53.

Niksa, S. and Lau, C. W. (1993), Global rates of devolatilization for various coal types,

Combustion and Flame, vol. 94, p. 293-307.

Nogami, H., Miura, T. and Furukawa, T. (1992), Simulation of transport phenomena

around raceway zone in the lower part of blast furnace, Tetsu-to-Hagane, vol. 78,

p. 1222-29.

Nogami, H., Yamaoka G. and Takatani, K. (2004), Raceway design for the innovative

blast furnace, ISIJ International, vol. 44, p. 2150-58

Nolde, H. D. Peters, C. and Wagner, E. D. (1996), Blast furnace coal injection,

Proceedings Blast Furnace Injection Symposium, Ohio, USA, p. 33-46

Ohno, Y., Furukawa, T. and Matsu-Ura, M. (1994), Combustion behaviour of

pulverized coal in a raceway cavity of blast furnace and its application to a large

amount injection. ISIJ International, vol. 34, p. 641-648.

Osório, E., Gomes, M. D. I., Vilela, A. C. F., Kalkreuth, W., de Almeida, M. A. A.,

Borrego, A. G., Alvarez, D. (2006), Evaluation of petrology and reactivity of coal

blends for use in pulverized coal injection (PCI), International Journal of Coal

Geology, vol. 68, p. 14–29.

265

Ostrowski, E. J. (1983), Factors influencing optimization of blast-furnace coal injection,

Ironmaking and Steelmaking, vol. 10, p. 215-221.

Parikh, J., Channiwala, S. A. and Ghosal, G. H. (2005), A correlation for calculating

HHV from proximate analysis of solid fuels. Fuel, vol. 84, p. 487–494.

Peng, J. H., Bi, X. T., Lim, C. J. and Sokhansanj, S. (2013). Study on density, hardness,

and moisture uptake of torrefied wood pellets, Energy Fuel, vol. 27, p. 967–974.

Peng, J. H., Bi, X. T., Sokhansanj, S. and Lim, C. J. (2012), A study of particle size

effect on biomass torrefaction and densification, Energy Fuel, vol. 26, p.

3826-3839.

Peng, J. H., Bi, X. T., Sokhansanj, S., Lim, C. J. and Melin S. (2010), An economical

and market analysis of Canadian wood pellets, International Journal of Green

Energy, vol. 7, p. 128-142.

Perlov, N. I. (1987), Technological approaches to energy saving in blast-furnace

operations in the iron and steel industry of the U.S.S.R, Energy, vol. 12, p.

1177-81.

Peters, M. and Lüngen H.B. (2009), Iron Making in Western Europe, Proceedings of the

5th International Congress on the Science and Technology of Ironmaking,

Shanghai, China, p. 22-28.

Peters, K. H., Beppler, E., Korthas, B. and Peters, M. (1991), Effect of high levels of

coal injection on blast furnace operation, Proceedings of 2nd European Ironmaking

Congress, Glasgow, U.K, p. 247-262.

Phanphanich, M.and Mani, S. (2010), Impact of torrefaction on the grindability and fuel

characteristics of forest biomass, Bioresource Technology, vol. 102, p. 1246-1253.

Picard, M. (2001), Pulverized coal combustion in the raceway, using a 3D numerical

simulation, Proceedings of Ironmaking Conference, Maryland, US, p. 229-239.

Picard, M., Bolsigner, P. J., Succurro, A., Eymond, J. L. and Valdan, G. (2000),

Development of injection lances in order to improve coal ignition inside the tuyere,

La Revue de Metallurgie-CIT, 36, p. 21-28.

Pimchuai, A., Dutta, A. and Basu, P. (2010), Torrefaction of agriculture residue to

enhance combustible properties, Energy Fuels, vol. 24, p. 4638-4645.

Poos, A. and Ponghis, N. (1990), Potentials and problems of high coal injection rates,

Proceedings of Ironmaking Conference, Michigan, USA, p. 443-453.

266

Porzio, G. F., Fornai, B., Amato, A., Matarese, N., Vannucci, M., Chiappelli, L., and

Colla, V. (2013), Reducing the energy consumption and CO2 emissions of energy

intensive industries through decision support systems – an example of application

to the steel industry, Applied Energy, vol. 112, p. 818–33.

Ranz, W. E. and Marshall, W. R. (1952a), Evaporation from drops, part I, Chemical

Engineering Process, vol. 48, p. 141–146.

Ranz, W. E. and Marshall, W. R. (1952b), Evaporation from drops, part II, Chemical

Engineering Process, vol. 48, p. 173-180.

Rousset, P., Aguiar, C., Labbe, N. and Commandre, J. M. (2011), Enhancing the

combustible properties of bamboo by torrefaction, Bioresource Technology, vol.

102, p. 8225-8231.

Sabil, K.M., Aziz, M.A., Lal, B. and Uemura, Y. (2014), Synthetic indicator on the

severity of torrefaction of oil palm biomass residues through mass loss

measurement, Applied Energy, vol. 114, p. 104-113.

Sami, M., Annamalai, M., and Wooldridge, M. (2001), Co-firing of coal and biomass

blends, Progress in Energy and Combustion Science, vol. 27, p. 171-214.

Sato, M., Ariyama, T., Yamakawa, Y. and Suzuki, M. (1994), Combustion mechanism

of pulverized coal injected into the blast furnace and its mathematical simulation,

Proceedings of ICSTI ’94, Iron and Steel Institute of Japan, Sendai, p. 511-516.

Sato, M., Murai, R. and Ariyama, T. (1996), Development of one-dimensional

mathematical model for pulverized coal combustion considering particle

dispersion, Tetsu to Hagane, vol. 82, p. 13-18.

Sato, M., Murai, R, Ariyama, T., Maki, A., Shimomura, A., and Mori, K. (1998),

Development of injection lance with high combustibility for high rate coal

injection, Tetsu to Hagane, vol. 84, p. 37-42.

Scaife, P. H., Mathieson, J. G., Rogers, H. and Nomura, S. (1983), Replacement of oil

by coal injection at the blast furnace, BHP Research Report, NERDDP/EG/84/268.

Scaife, P. H., Mathieson, J. G., Rogers, H. and Nomura, S. (1983), Replacement of oil

by coal injection at the blast furnace, BHP Research Report, NERDDP/EG/84/268

267

Schott, R. (2012), State-of-the-art PCI technology for blast furnace ensured by

continuous technological and economical improvement, Proceedings of the Iron

and Steel Technology Conference, Georgia, USA, P. 589-604.

Seeker, W. R., Samuelsen, G. S., Heap M. P. and Trolinger, J. D. (1981), The thermal

decomposition of pulverized coal particles, Proceedings of 18th

Symposium

(International) on Combustion, The Combustion Institute, Pittsburgh, PA, p.

1213-26.

Selvarasu, N. K., Huang, D., Chen, Z. M., Gu, M. Y., Chaubal, P. and Zhou, C. Q,

(2006), Prediction of raceway in a blast furnace, Proceedings of ASME 2006

International Mechanical Engineering Congress and Exposition, Illinois, USA, p.

297-303.

Shampine, R. W., Cohen, R. D., Bayazitoglu, Y. and Anderson, C. F. (1995), Eeffect of

agglomeration on pulverized-coal combustion, Combustion and Flame, vol. 101, p.

185-191.

Shen, Y. S., Guo, B. Y., Yu, A. B., Austin, P. R. and Zulli, P. (2011), Three

dimensional modelling of in-furnace coal/coke combustion in a blast furnace, Fuel,

vol. 90, p. 728-738.

Shen, Y. S., Guo, B. Y., Yu, A. B., Maldonado, D., Austin, P. and Zulli, P. (2008),

Three-dimensional modelling of coal combustion in blast furnace, ISIJ

International, vol. 48, p.777-786.

Shen, Y. S., Guo, B. Y., Yu, A. B. and Zulli, P. (2009a), A three-dimensional numerical

study of the combustion of coal blends in blast, Fuel, vol. 88, p. 255-263.

Shen, Y. S., Guo, B. Y., Yu, A. B. and Zulli, P. (2009b), Model study of the effects of

coal properties and blast conditions on pulverized coal combustion, ISIJ


Shen, Y. S., Maldonado, D., Guo, B. Y., Yu, A. B., Austin, P. and Zulli, P (2009c),

Computational fluid dynamics study of pulverized coal combustion in blast furnace

raceway, Industrial and Engineering Chemistry Research, vol. 48, p. 10314–23.

Shuit, S. H., Tan, K. T., Lee, K. T. and Kamaruddin, A.H. (2009), Oil palm biomass as

a sustainable energy source: A Malaysian case study, Energy, vol. 34, p.

1225-1235.

Singer, S. (1984), Pulverized coal combustion: recent developments, Noyes

Publications, New Jersey.

268

Sivathanu, Y. R. and Faeth, G. M. (1990), Generalized state relationships for scalar

properties in non-premixed hydrocarbon/air flames, Combustion and Flame, vol.

82, p. 211-230.

Smith, I. W. (1982), The combustion rates of coal char: A review, Proceedings of

Nineteenth Symposium (International) on combustion, Pittsburgh, The combustion

institute, p. 1045-65.

Smoot, L. D. and Pratt, D. T. (1979), Pulverized coal combustion and gasification, New

York, London, Plenum.

Smoot, L. D. and Smith P. J. (1985) Coal combustion and gasification, New York and

London, Plenum Press.

Spalding, D. B. (1971), Mixing and chemical reaction in steady confined turbulent

flames, Proceeding of 13th Symposium (International) on Combustion, Pittsburgh,

The Combustion Institute, p. 649-657.

Steeghs, A. G. S. (1992), Determination of coal combustion under simulated blast

furnace raceway conditions, Technical Steel Research Report of European

Commission.

Steeghs A. G. S., Schoone E. E. and Toxopeus H. L. (1996), High injection rate of coal

into the blast furnaces of Hoogonvens IJmuiden, Proceedings of Injection

Technology in Ironmaking and Steelmaking, Brussels, Belgium, p. 75-87.

Steiler J. M., Lao D. and Lebonvallet J. L. (1996), Development of coal injection in the

blast furnace at Usinor Sacilor, Proceedings of Injection Technology in Ironmaking

and Steelmaking, Brussels, Belgium, p.15-32.

Suopajärvi, H., Pongrácz, E. and Fabritius, T. (2014), Bioreducer use in Finnish blast

furnace ironmaking – Analysis of CO2 emission reduction potential and mitigation

cost, Applied Energy, vol. 124, p. 82-93.

Suzuki, T., Hirose, R. Morimoto, K. and Abe, T. (1984), High intensity combustion of

coal for application to a blast furnace, Proceedings of 20th

Symposium

(International) on Combustion, University of Michigan, Michigan, p. 1419-25.

Suzuki, T., Smoot, L. D., Fletcher, T. H. and Smith, P. J. (1986), Prediction of

high-Intensity pulverized coal combustion, Combustion Science and Technology,

vol. 45, p. 167-183.

269

Suzuki, T, Uehara, T. and Akedo, H. (1990), Combustion characteristics of pulverized

coal for blast furnace coal injection, Proceedings of Ironmaking Conference, Iron

and Steel Society, Warrendale, PA, USA, p. 465-471.

Tate, A. (1993), The Variation of Char Reactivity during the Combustion of Pulverized

Coal, Ph.D. Thesis, The University of Newcastle.

Takeda, K. (1994), Mathematical modelling of pulverised coal combustion in a blast

furnace, Ph.D thesis, Department of Mechanical Engineering, The University of

London, England.

Takahashi, H., Kawai, H. and Suzuki, Y. (2002), Analysis of stress and buoyancy for

solids flow in the low part of a blast furnace, Chemical Engineering Science, vol.

57, p. 215-226.

Takeda, K. and Lockwood F. C. (1997), Integrated mathematical model of pulverised

coal combustion in a blast furnace, ISIJ International, 1997, vol. 37, p. 432-440.

Takeda, K., Miyagawa, S. and Taguchi, S. (1990), Effect of coal properties on

combustibility of coal injected to blast furnace, Proceedings of Ironmaking

Conference, Iron and Steel Society, Detroit, USA, p. 455-465.

Toporov, D. (2014), Combustion of pulverised coal in a mixture of oxygen and recycled

flue gas, Esvier Science, London.

Toyoda, S. (1983), Changes of the use of energy in Japanese steel industry,

Transactions of ISIJ, vol. 23, p. 1-13

Ubhayakar, K., Stickler, D. B, Rosenberg, C. W. V. and Ganon, R. (1976), Rapid

devolatilization of pulverized coal in hot combustion gases, Proceedings of 18th

Symposium (International) on Combustion, The Combustion Institute, Pittsburgh,

p. 427–436.

Ueno, H., Yamaguchi, K and Tamura, K. (1993), Coal combustion in the raceway and

tuyere of a blast furnace, ISIJ international, vol. 33, p. 640-645.

Vamvuka, D., Schwanekamp, G. and Gudenau, W. (1996), Combustion of pulverized

coal with additives under conditions simulating blast furnace injection, Fuel, vol.

75, p. 1145-50.

van der Stelt, M. J. C., Gerhauser, H., Kiel, J. H. A. and Ptasinski, K. J. (2011), Biomass

upgrading torrefaction for the production of biofuels: a review. Biomass and

Bioenergy, vol. 35, p. 3748-62.

270

Wall, T. F., Liu, G. S., Wu, H. W., Roberts, D. G., Benfell, K. E., Gupta, S, Lucas, J. A.

and Harris, D. J. (2002), The effects of pressure on coal reactions during pulverised

coal combustion and gasification, 'Progress in Energy and Combustion Science,

vol. 28, p. 405-433.

Wakimoto, H. Sato, K., Fujiura, M. and Hara, H. (1983), Combustion characteristics of

pulverized coal in pressurized state, Tetsu-to-Hagane, vol. 69, p. S105.

Wang, C., Ryman, C. and Dahl, J. (2009), Potential CO2 emission reduction for

BF-BOF steelmaking based on optimised use of ferrous burden materials,

International Journal of Greenhouse Gas Control, vol. 3, p. 29-38.

Wei, S. M. and Qi, Z. D. (1983), Coal-powder injection on blast furnace of Shoudu Iron

and Steel Co., Ironmaking and Steelmaking, vol. 10, p. 109-113.

Wijayanta, A. T., Alam, M. S., Nakaso, K., Fukai, J., Kunitomo, K. and Shimizu, M.

(2014), Combustibility of biochar injected into the raceway of a blast furnace, Fuel

Processing Technology, vol. 117, p. 53-59.

Wikström, J. O., Sköld, B. E. and Karsrud, K. (1996), SSAB/MEFOS oxy-coal system –

3 Years of industrial experience, Proceedings of Ironmaking, Iron and Steel

Society, Pittsburgh, p. 24-27.

Willmer, R. R. (1989), The effect of blast furnace coal injection upon bosh coke

properties, coke combustion and furnace permeability, Proceedings of Ironmaking

Conference, Chicago, USA, p. 359-401.

Wu, J. B. and Hsieh, J. S., (2003), The development and result of one bit opening

method at CSC blast furnace”, Mining and Metallurgy, vol. 47, No. 3, p.62-67.

Yamagata, C, Suyama S., Horisaka, S., Takatani, K., Kajiwara, Y., Komatsu, S.,

Shibuta, H. and Aminaga, Y. (1992), Fundamental study on pulverized coal

injected into coke bed at high rate, ISIJ International, vol. 32, p. 725-732

Yamaguchi, K., Ueno, H. and Tamura K. (1992), Maximum injection rate of pulverized

coal into blast furnace through tuyeres with consideration of unburnt char, ISIJ


Yan, B. H., Cao, C. X., Cheng, Y., Jin, Y and Cheng, Y. (2014), Experimental

investigation on coal devolatilization at high temperatures with different heating

rates, Fuel, vol. 117, p. 1215-22.

271

Yeh, C. M., Ho, C. K., Du, S. W., Chou, C. S. and Hsu, C. H. (2002), Development of

pulverized coal injection lance with air cooling in No. 3 blast furnace of CSC,

Proceedings of 61st Ironmaking Conference, Tennessee, USA, p. 267-276.

Yeh, C. P. Du, S. W., Tsai, C. H. and Yang, R. J. (2012), Numerical analysis of flow

and combustion behavior in tuyere and raceway of blast furnace Fueled with

pulverized coal and recycled top gas”, Energy, vol. 42, p. 233-240.

Yu, J. L., Lucas, J., Strezov, V. and Wall, T. (2003), Swelling and char structures from

density fractions of pulverized coal, Fuel, vol. 82, p. 1977-87.

Yu, Q. Y. (1999), Experimental research on the combustion efficiency of blended coal

injection into BF, Baosteel Technical Report, vol. 1, p. 33-36.

Zhang, W. D., Chen, H., Ma, Z. J. Song, J. L. and Wu, J. L. (2013), Technological

development and competitiveness analysis of large capacity blast furnace,

Proceedings of the Fifth Baosteel Biannual Academic Conference, Shanghai,

China, p. 235-240.

Zhou, L. X. (1993), Theory and numerical modelling of turbulent gas-particle flows and

combustion, CRC Press, Florida, USA.

Zhu, M. and Xu H. Y. (2014), Production practice of keeping No.1 1880m3 BF long

term stable regular working in Laiwu Steel, Shandong Metallurgy, vol. 36, p. 4-6.

Ziebik, A. and Stanek, W. (2001), Forecasting of the energy effects of injecting plastic

wastes into the blast furnace in comparison with other auxiliary fuels, Energy, vol.

26, p. 1159-73.

http://pubs.acs.org/doi/abs/10.1021/ef030022h

http://pubs.acs.org/doi/abs/10.1021/ef030022h

pulverised coal injection into blast furnace - a …

Documents