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4–5 February 2013

5th Annual High Temperature Processing Symposium 2013

Book of Abstracts

High Temperature Processing Symposium 2013 Swinburne University of Technology 1

HIGH TEMPERATURE PROCESSING SYMPOSIUM 2013 Swinburne University of Technology 4 – 5 February 2013, Melbourne, Australia Editors

M. Akbar Rhamdhani Geoffrey Brooks

Abdul Khaliq

Organising Committee

M. Akbar Rhamdhani Geoffrey Brooks John Grandfield Abdul Khaliq Md Saiful Islam Ben Ekman Sazzad Ahmad Mohammad Mehedi Shabnam Sabah Reiza Mukhlis Md Abdus Sattar

Published in Australia by: Faculty of Engineering and Industrial Sciences, Swinburne University of Technology ISBN 978-0-9871772-6-1 © 2013 Swinburne University of Technology Apart from fair dealing for the purpose of private study, research, criticism or review as permitted under the Copyright Act, no part may be reproduced by any process without the written permission of the publisher. Responsibility for the contents of the articles rests upon the authors and not the publisher. Data presented and conclusions drawn by the authors are for information only and not for use without independent substantiating investigations on the part of the potential user.


HIGH TEMPERATURE PROCESSING SYMPOSIUM 2013 Swinburne University of Technology 4 – 5 February 2013, Melbourne, Australia We wish to thank the main sponsors for their contribution to the success of this symposium


5th High Temperature Processing Symposium 2013 Swinburne University of Technology

ATC 101, Hawthorn Campus

Sponsored by CSIRO, Furnace Engineering, OneSteel, Outotec

Symposium Program

Day 1 (4 February 2013) in ATC101 8.30 to 9.00 Registration in Foyer Advanced Technologies Centre (ATC) 9.00 to 9.10 Welcome by Prof George Collins (Deputy Vice-Chancellor

Research and Development, Swinburne University of Technology) Session 1 Chaired by: Dr M Akbar Rhamdhani (Swinburne) 9.10 to 9.40 01 – Keynote: Prof Aldo Steinfeld (ETH Zurich/Paul Scherrer

Institute) – Solar Thermochemical Processing of Fuels and Materials

9.40 to 10.00 02 – Dr Rene Olivares (CSIRO) - Lithium-Sodium-Potassium Nitrate Salt for Thermal Energy Storage Thermo-Chemical Evaluation in Different Atmospheres

10.00 to 10.20 03 – Mr Abdul Khaliq (Swinburne) – Mechanism of VB2 Formation in Molten Aluminium

10.20 to 10.40 04 – Dr Stephanie Vervynckt (Umicore) – Focus on Recycling of Critical Raw Materials: An Industry’s Perspective

10.40 to 10.55 Coffee/Tea in ATC Foyer/ATC105 Session 2 Chaired by: Prof Geoffrey Brooks (Swinburne) 10.55 to 11.25 05 – Keynote: Prof Veena Sahajwalla (UNSW) – Recycling End-

of-Life Waste Materials as Resources in EAF Steelmaking – Fundamentals of High Temperature Reactions and Industrial Implementations

11.25 to 11.45 06 – Ms Elien Haccuria (University of Queensland) – Recycling Lithium Ion Batteries

11.45 to 12.05 07 – Mr Saiful Islam (Swinburne) – Kinetics of Si Refining using Slag Treatment

12.05 to 12.25 08 – Prof Doug Swinbourne (RMIT) - Minor Element Distributions during Copper Flash Converter

12.25 to 1.15 Lunch in ATC Foyer/ATC105


Session 3 Chaired by: Assoc Prof Brian Monaghan (Uni of Wollongong) 1.15 to 1.45 09 – Keynote: Prof Evgueni Jak (University of Queensland) –

Integrated Experimental and Modelling Research Methodology for Phase Equilibria, Thermodynamics and Viscosities of Metallurgical Slags

1.45 to 2.05 10 – Dr Nazmul Huda (Swinburne) – Aluminium Production Route through Carbosulfidation of Alumina utilising H2S

2.05 to 2.25 11 – Mr Ata Fallah Mehrjardi (University of Queensland) – Investigation of Freeze-Linings in Copper-Containing Slag Systems

2.25 to 2.45 12 – Mr Ross Baldock (Outotec) – Circosmelt: Outotec's Alternative Ironmaking Process

2.45 to 3.00 Coffee/Tea in ATC Foyer/ATC105 Session 4 Chaired by: Mr Warren Bruckard (CSIRO) 3.00 to 3.30 13 – Keynote: Dr Mark Pownceby (CSIRO) – Insights into the

Formation of Iron Ore Sinter Bonding Phases 3.30 to 3.50 14 – Dr Joe Herbertson (The Crucible) – The Crucible Process 3.50 to 4.10 15 – Mr Hamed Abdeyazdan (University of Wollongong) – The

Effect of Slag Basicity on Spinel Inclusion Wettability 4.10 to 5.00 Panel Discussion – “What is the future of energy supply for

metallurgical industries” – led by Adjunct Prof John Grandfield

Close of Day 1

Day 2 (5 February 2013) in ATC101 8.30 to 9.00 Registration in Foyer Advanced Technologies Centre (ATC) Session 5 Chaired by: Mr Richard Simpson (Furnace Engineering) 9.00 to 9.30 16 – Keynote: Prof George Kaptay (University of Miskolc) -

Current Issues of High Temperature Thermodynamics 9.30 to 9.50 17 – Dr Nawshad Haque (CSIRO) – Life Cycle Based Greenhouse

Gas Emission Assessment from Ferroalloy Production 9.50 to 10.10 18 – Dr Yvonne Durandet (Swinburne) – Special Alloy Strip

Production in a Micro-Mill Environment 10.10 to 10.30 19 – Dr Sri Harjanto (University of Indonesia) - Carbothermic

Reaction of High-Combined-Water Iron Ore and Coal Composite Pellet

10.30 to 10.45 Coffee/Tea in ATC Foyer/ATC105


Session 6 Chaired by: Mr Leo Frawley (OneSteel) 10.45 to 11.05 20 – Prof Geoffrey Brooks (Swinburne) – Development of Dynamic

Process Models for Oxygen Steelmaking 11.05 to 11.25 21 – Assist/Prof Youn-Bae Kang (Pohang University of Science and

Technology) – Coupled Experimental and Thermodynamic Modelling for Metallurgical Slags and Inclusions Containing Sulphur

11.25 to 11.45 22 – Mrs Shabnam Sabah (Swinburne) - Analysis of Waves in a Cavity and Their Significance to Splashing in Steelmaking

11.45 to 12.05 23 – Mr Michael Sommerville (CSIRO) – Injection of Charcoal to Slag Bath

12.05 to 12.25 24 – Dr Sun-Joong Kim (Tohoku University) – Innovative Process of Manganese Recovery from Steelmaking Slag by Sulphurisation

12.25 to 1.15 Lunch in ATC Foyer/ATC105 Session 7 Chaired by: Dr Guangqing Zhang (University of Wollongong) 1.15 to 1.35 25 – Dr Zulfiadi Zulhan (Insitute of Technology Bandung) - Vacuum

Treatment of Molten Steel in RH (Ruerhstal Heraeus) and VTD (Vacuum Tank Degasser): A Comparative Study

1.35 to 1.55 26 – Dr Luckman Muhmood (CSIRO) - Experimental Investigations on the Dynamics of Interfacial Phenomena in Synthetic Blast Furnace Slags

1.55 to 2.15 27 – Mr Xiang Li (University of Wollongong) - Synthesis of High Purity Silicon Carbide for Solar Silicon Production

2.15 to 2.35 28 – Mr Taufiq Hidayat (University of Queensland) - Thermodynamic Optimization of Iron-Silicate Slag for Simulation of Copper Smelting Processes

2.35 to 2.50 Coffee/Tea in ATC Foyer/ATC105 Session 8 Chaired by: Mr Jacob Wood (Outotec) 2.50 to 3.10 29 – Dr Ryan Cottam (Swinburne) - Thermodynamic Assessment

of In-Situ Formation of Hard Phase Materials 3.10 to 3.30 30 – Mr Zhe Wang (University of Wollongong) – Behaviour of New

Zealand Ironsand during Iron Ore Sintering 3.30 to 3.50 31 – Mr Imam Santoso (University of Queensland) – Phase

Equilibria Studies of Cu-S and Cu-Fe-S systems 3.50 to 4.10 32 – Mr Abdus Sattar (Swinburne) - A Comprehensive Approach

for CFD Modelling of Slag Foaming with Population Balance Modelling

4.10 to 4.30 33 – Mr Tijl Crivits (University of Queensland) - Investigation of Phase Equilibria in the Cu-Fe-Si-Mg-O System at Low MgO Concentrations in Equilibrium with Copper Metal

4.30 to 4.35 CLOSING

Close of Symposium


Day 3 (6 February 2013) in ATC205 Post-Symposium Workshop 8.45 to 9.00 Registration in ATC205, Advanced Technologies Centre (ATC) 9.00 to 13.00 Application of Thermodynamics to Industrial Processes – by

Dr M Akbar Rhamdhani, Swinburne This workshop is specifically designed for Engineers and Applied

Scientists who would like to refresh their understanding of thermodynamics and/or learn its applications to industrial processes (evaluation and optimisation processes).

13.00 Close of Workshop

Campus Map – Swinburne @ Hawthorn, Melbourne


KEYNOTE: PRESENTATION - 1

Solar Thermochemical Processing of Fuels and Materials

Aldo Steinfeld Department of Mechanical and Process Engineering, ETH Zurich, Switzerland

and Solar Technology Laboratory, Paul Scherrer Institute, Switzerland

www.pre.ethz.ch

Keywords: solar, thermochemical, fuel, syngas, redox, metal, gasification, carbothermal.

Solar thermochemical processes for the production of synthetic fuels and materials make use of concentrated solar radiation as the energy source of high-temperature process heat. Considered are H2O/CO2-splitting thermochemical cycles via metal oxide redox reactions, gasification processes for the thermal conversion of biomass and other carbonaceous feedstock, and carbothermic reduction processes in the extractive metallurgy. These processes inherently operate at high temperatures and utilize the entire solar spectrum, and as such provide thermodynamic favorable paths to efficient and clean production [1].

Solar Fuels — Considered is the ceria-based redox cycle for splitting H2O and CO2 [2]. A 3-kW solar cavity-receiver containing a reticulated porous ceramic (RPC) foam made of pure CeO2 has been experimentally investigated [3]. The RPC was directly exposed to concentrated thermal radiation at mean solar flux concentration ratios exceeding 3,000 suns. During the endothermic reduction step, solar radiative power inputs in the range 2.8-3.8 kW and nominal reactor temperatures from 1400 to 1600°C yielded CeO2-δ with oxygen deficiency δ up to 0.042. In the subsequent exothermic oxidation step at below about 1000°C, CeO2-δ was stoichiometrically re-oxidized with CO2 to generate CO. The solar-to-fuel energy conversion efficiency, defined as the ratio of the calorific value of the fuel produced to the solar radiative energy input through the reactor’s aperture, was 1.73% average and 3.53% peak. These are the highest solar-to-fuel energy conversion efficiency values reported to date for a solar-driven device converting CO2 to CO. We also demonstrated the simultaneous splitting of H2O and CO2 for the co-production of H2 and CO (syngas), whose molar ratio H2:CO was controlled by adjusting the H2O:CO2 molar ratio.

Solar Gasification — The advantages of the solar-driven gasification vis-à-vis the conventional autothermal gasification are [4]: a) It delivers higher syngas output per unit of feedstock, as no portion of the feedstock is combusted for process heat; b) It avoids contamination of syngas with combustion by-products, and consequently reduces costly downstream gas cleaning and separation requirements; c) It produces syngas with higher calorific value and lower CO2 intensity, as the energy content of the feedstock is upgraded by up to 33% through the solar energy input; d) It allows for higher gasification temperatures


(>1200°C), resulting in faster reaction kinetics and higher quality of the syngas produced with low – or without – tar content, and further enabling the processing of virtually any type of carbonaceous feedstock, resulting in higher exploitation of the available resources; e) It eliminates the need for an upstream air separation unit, as steam is the only gasifying agent, which further facilitates economic competitiveness. Within the framework of a joint ETH-PSI-Holcim R&D project, a 250 kW solar industrial pilot plant was experimentally demonstrated at the solar tower concentrating facility of the Plataforma Solar de Almería (Spain) under realistic operating conditions relevant to industrial solar concentrating systems. Residual biomass and other carbonaceous wastes such as tire chips, plastics, and industrial and sewage sludges of heterogeneous characteristics (in terms of chemical composition, particle size, morphology, moisture, volatile, ash, and fixed carbon contents) were thermochemically converted to high-quality syngas with a calorific value upgraded over that of the input feedstock.

Solar Metals — The extractive metallurgical industry is characterized by its energy-intensive processes and their concomitant CO2 emissions, derived mainly from the combustion of fossil fuels for heat and electricity generation. These emissions can be significantly mitigated by applying concentrated solar energy as the source of high-temperature process heat. Considered are the solar-driven productions of Fe, Zn, Mg, Al, and Si by carbothermal reduction of their oxides. When the reducing agent is derived from a biomass source, the solar-driven carbothermal processes are CO2 neutral.

The EU-project SOLZINC demonstrated the solar pilot production of Zn by carbothermal reduction of ZnO. The key component is a 300-kW solar chemical reactor. It consists of two cavities in series. The upper cavity functions as the solar absorber and contains a windowed aperture to let in concentrated solar radiation. The lower cavity functions as the reaction chamber and contains a packed-bed of the reacting ZnO and bio-charcoal. Testing at a solar tower in the 1300–1500 K range yielded up to 50 kg/h of 95%-purity Zn with energy conversion efficiency (ratio of the reaction enthalpy change to the solar power input) of about 30%.

Using concentrated solar process heat, the carbothermal reductions of Al2O3 to Al and SiO2 to Si were examined thermodynamically and demonstrated experimentally at vacuum pressures [5]. Reducing the system pressure favors Al(g) and Si(g) formation, enabling their vacuum distillation and avoiding contamination by carbides and/or oxycarbides. Exploratory

quartzwindow

upper cavity(absorber)

lower cavity(reaction chamber)

ZnO/Cpacked-bed

carrier gasinlet

300 kWconcentrated solar

power

gaseousproducts


experimentation in a solar reactor was performed with mixtures of charcoal with alumina and silica in the ranges of 1300-2000 K and 1997–2263 K, respectively, at ~3×10-3 bar by direct exposure to concentrated thermal radiation. Distilled samples contained up to 19 wt.% of Al in Al-Al2O3 mixtures and 79 wt% of Si in Si-SiO2 mixtures.

For all these solar thermochemical processes, R&D work encompasses fundamental studies on thermodynamics, reaction kinetics, heat/mass transfer, and chemical reactor engineering. Solar reactor prototypes – at the 10 kW power level – are designed, fabricated, modeled, and tested in a high-flux solar furnace, further optimized for maximum solar-to-chemical energy conversion efficiency, and finally scaled-up for industrial applications – at the MW power level – using concentrating solar tower technology.

References

1. Romero M., Steinfeld A., “Concentrating Solar Thermal Power and Thermochemical Fuels”, Energy & Environmental Science, Vol. 5, pp. 9234–9245, 2012.

2. Chueh W.C., Falter C., Abbott M,, Scipio D., Furler P., Haile S.M., Steinfeld A., “High-Flux Solar-Driven Thermochemical Dissociation of CO2 and H2O using Nonstoichiometric Ceria”, Science, Vol. 330, pp. 1797-1801, 2010.

3. Furler P., Scheffe J., Gorbar M., Moes L., Vogt U., Steinfeld A., “Solar thermochemical CO2 splitting utilizing a reticulated porous ceria redox system”, Energy & Fuels, in press.

4. Piatkowski N., Wieckert C., Weimer A.W., Steinfeld A., “Solar-driven gasification of carbonaceous feedstock – A review”, Energy & Environmental Science, Vol. 4, pp. 73-82, 2011.

5. Kruesi M., Galvez M.E., Halmann M., Steinfeld A., “Solar Aluminum Production by Vacuum Carbothermal Reduction of Alumina – Thermodynamic and Experimental Analyses”, Metallurgical and Materials Transactions B, Vol. 42B, pp. 254-260, 2011.


PRESENTATION - 2

Lithium-Sodium-Potassium Nitrate Salt for Thermal Energy Storage Thermo-Chemical Evaluation in Different Atmospheres

Rene I. Olivares

CSIRO Energy Centre PO Box 330, Newcastle NSW 2300, Australia

Keywords: thermal energy storage, concentrating solar power, high temperature stability Introduction With the development of tower technologies for concentrating solar power (CSP), the possibility of reaching very high temperatures has become a reality. The recent commercially demonstrated system, a 20MWe power tower Torresol Gemasolar in Spain [1], has a molten salt thermal storage capacity of 15 hours and uses a binary solar salt of 60wt% NaNO3 - 40wt% KNO3 that has a relatively elevated melting point of 222oC and is available up to temperature of 565oC. It would be desirable to identify lower melting point salt that is stable at higher temperature than 565oC; this would increase the heat-to-electricity conversion efficiency. Using nitrates-based salts at higher temperature may be possible by exercising rigorous atmosphere control to delay thermal decomposition [2]. The exploration of this possibility on the low melting point ternary eutectic LiNO3-NaNO3-KNO3 was studied in this work by simultaneous differential scanning calorimetry, thermogravimetry and mass spectrometry (DSC/TG-MS). Stability and thermal decomposition of nitrates The principal mode of thermal decomposition in nitrate salts has been agreed amongst researchers [3-5] to be NO3

- ↔ NO2- + ½O2. Further decomposition also takes place with the

evolution of the oxides of nitrogen, particularly at higher temperatures [2,6]. In some studies [3,7,8] the decomposition temperature has been reported as the temperature at which oxygen, nitrogen or nitrous oxide is detected in the gas phase. Gordon and Campbell [7] reported that the alkali metal nitrates were observed to bubble undergoing a thermal reaction at temperatures as low as 100oC to 300oC above their melting points and the evolution of nitrous fumes observed to occur at temperatures ranging from about 200oC to 350oC above the initial bubbling reaction. The later stage of the decomposition of these nitrates was still occurring at temperatures as high as 900oC which was the upper limit of the apparatus employed [7]. Experiments at 300oC with a pure equimolar sodium potassium nitrate melt in an oxygen rich atmosphere [9] reported an oxide ion concentration (O2

-) in the molten salt much higher than that which would correspond to the dissociation constant for nitrite ion (NO3

-), based on the equilibrium NO3- NO2

- + O2-, demonstrating that nitrate ion (NO3

-) could decompose to a measurable extent at temperatures as low as 295oC. The thermal stability of binary mixes of alkali metal nitrates as investigated by means of differential thermal analysis (DTA) and evolved gas analysis (EGA) [10] concluded that the nitrate solutions are thermally unstable at temperatures higher than 500oC. Evolved gas analysis measurements [10] indicated that the decomposition of nitrates took place independently and followed the ranking order of thermodynamic stability as determined by the reactions sequence MNO3 MNO2 + ½O2 followed by 2MNO2 M2O + NO + NO2 where M can


represent Li, Na or K respectively. On this basis the decomposition of a mixture of LiNO3-NaNO3-KNO3 would be determined by the least thermodynamically stable species, in this case LiNO3 followed by NaNO3 and then KNO3, a possibility if the solution is close to ideal. In the present work the thermo-chemical behaviour and thermal stability of the LiNO3-NaNO3-KNO3 eutectic was studied from at least three different points of view: (i) the temperature at which the rapid evolution of gases, NO and/or NO2 and O2, are detected by means of mass spectrometry (MS), (ii) the temperature of the melt at which an irreversible endothermic peak of decomposition is resolved by differential scanning calorimetry (DSC), and (iii) the temperature at which rapid weight loss is observed in a thermo-gravimetric curve (TG). A combination of these three criteria is used for elucidation of the limiting temperature for operation of the salt as would apply in a TES installation. The chemistry and equilibrium reactions by which molten nitrites/nitrates may interact with the atmosphere can be found in [2,11-13]. Experimental results It was found that the stability of the LiNO3-NaNO3-KNO3 eutectic, as measured by the gases evolving from the melt, was influenced by the atmosphere. Evolution of the main gaseous species NO was detected at 325oC in an atmosphere of argon, at 425oC in an atmosphere of nitrogen, at 475oC in an atmosphere of air and at 540oC in an atmosphere of oxygen. Examples of the mode of decomposition and thermal stability of the ternary eutectic are shown in Figure 1 for an argon cover gas and in Figure 2 for an air cover gas.

Figure 1: DSC/TG-MS analysis of thermal decomposition of lithium-sodium-potassium nitrate eutectic under a cover gas of argon

The delay in the thermal decomposition as measured by the evolution of gaseous NO is clearly demonstrated in Figure 2 for the salt under a cover gas of air.

Sample Temperature (°C)1000900800700600500400300200100

TG |s

c (m

g)

25

20

15

10

5

Heat

Flow

|sc

(mW

)

0

-20

-40

-60

-80

MS

Ion

Curre

nt In

tens

ity (A

)

1E-11

1E-10

1E-9

1E-8

1E-7

Rapid wt loss begin, T: 602.81 (°C)

Exo

Heat : 4.2 (J/g) Peak Maximum : 94 (°C) Solid-solid transf : 88 (°C)

Heat : 163.4 (J/g) Peak Maximum : 129 (°C) Melting T : 121 (°C)

Onset, bulk of decomposition : 671 (°C)

(N) T: 584.12 (°C)

(O) T: 588.42 (°C) (N2) T: 693.19 (°C)

(NO) T: 325.09 (°C)

(O2) T: 418.29 (°C)

(NO2) T: 659.54 (°C) NO2

N

NO

O2N2

N2O

OT : 895.11 (°C)


Figure 2: DSC/TG-MS analysis of thermal decomposition of lithium-sodium-potassium nitrate eutectic under a cover gas of air

Cycling the melt several times between 50oC and 400oC allowed the determination of the melting and solidification points respectively. Figure 3 shows that prior to melting; the eutectic underwent endothermic (α/β) solid-solid type transformation. From an average of four cycles; the (α/β) type transformation occurred at 87oC, the melting point was 121oC, and the solidification point 98oC. Under-cooling of the salt coincided with the onset of the (α/β) solid-solid transformation upon heating. The measurement of temperature is accurate to

2.7oC.

Figure 3: Melting point and solidification point determination of lithium-sodium-potassium

nitrate eutectic


TG |s

c (m

g)

24

20

16

12

8

4

Heat

Flow

|sc (

mW

)

20

0

-20

-40

-60

MS

Ion

Curre

nt In

tens

ity (A

)

1E-12

1E-11

1E-10

1E-9

1E-8

Rapid wt loss begin, T: 601.93 (°C)

Exo

Heat : 135.5 (J/g) Peak Maximum : 128 (°C) Melting T : 122 (°C)

Onset, bulk of decomposition : 625 (°C)

Heat : 4.1 (J/g) Peak Maximum : 95 (°C) Solid-solid transf : 88 (°C)

(NO) T: 475.72 (°C)

(NO2) T: 550.41 (°C)


Hea

tFlo

w |s

c (h

eatin

g) (m

W) 0

-10

-20

-30

-40

-50

-60

Hea

tFlo

w |c

[6] (

Coo

ling)

(mW

)40

30

20

10

0

Exo

Exo

Heat : 6.9 (J/g) Peak Maximum : 94 (°C) Onset (solid-solid transf): 86 (°C) Heat : 140.5 (J/g)

Peak Maximum : 128 (°C) Onset (melting temperature) : 122 (°C)

Heat : -97.4 (J/g) Peak Maximum : 93 (°C) Onset (solidification temperature) : 98 (°C)


Long term stability At a temperature of 500oC in air, TG analysis indicated that the long term stability of the salt was limited and this was confirmed by DSC. Figure 4 shows that for a salt sample after four hours at 500oC, although the melting and solidification points were unaffected, the DSC signature of the salt was significantly different.

Figure 4: Long term stability of salt at 400oC and 500oC respectively and DSC analysis after

4 hours

Although the salt can be stabilised in an oxygen rich atmosphere to reach 540oC [2], the long-term stability of the ternary eutectic at greater than 500oC is considerably impacted, this being evident by the increased rate of salt vaporization as measured by TG analysis. The use of the salt in a closed or slightly pressurised containment arrangement is expected to minimise this problem.

References

1. R.I. Dunn, P.J. Hearps, M.N. Wright, Molten-Salt Power Towers: newly Commercial Concentrating Solar Storage, Proceedings of the IEEE 100(2) (2012) 504-515.

2. R. Olivares, The Thermal Stability of Molten Nitrite/Nitrates Salt for Solar Thermal Energy Storage in Different Atmospheres, Solar Energy 86 (2012) 2576-2583.

y = -2E-05x + 1R² = 0.999

y = 2E-07x2 - 0.0002x + 0.9997R² = 0.9999

0.90

0.92

0.94

0.96

0.98

1.00

0 60 120 180 240

Frac

tiona

l wei

ght l

oss

Time at temperature, min

400oC

500oC

Sample Temperature (°C)400350300250200150100H

ea

tFlo

w |c |sc (

he

atin

g)

(mW

)

20

10

0

-10

-20

-30

He

atF

low

|c |sc (

co

olin

g)

(mW

)

20

10

0

-10

-20Exo

Exo

Peak Maximum : 90 (°C) Onset : 82 (°C)

Peak Maximum : 127 (°C) Onset : 120 (°C)

Peak Maximum : 92 (°C) Onset (solidification temperature) : 97 (°C)


He

atF

low

|c |s

c (h

ea

ting

) (m

W)

30

25

20

15

10

He

atF

low

|c |s

c (c

oo

ling

) (m

W)

5

0

-5

-10

-15

-20Exo

Exo

Peak Maximum : 122 (°C) Onset (melting temperature) : 118 (°C)

Peak Maximum : 219 (°C) Onset (exothermic): 213 (°C)

Peak Maximum : 222 (°C) Onset (endothermic): 224 (°C) Peak Maximum : 130 (°C)

Onset (endothermic): 135 (°C)

Peak Maximum : 90 (°C) Onset (solidification temperature): 93 (°C)


3. E.S. Freeman, The kinetics of the thermal decomposition of sodium nitrate and of the reaction between sodium nitrite and oxygen, Journal of Physical Chemistry 60(11) (1956) 1471-1600.

4. G.D. Sirotkin, Equilibrium in melts of the nitrates and nitrites of sodium potassium, Russian Journal of Inorganic Chemistry 4(11) (1959) 1180-1184.

5. A. Buchler, J. Stauffer, Gaseous Alkali-Nitrogen-Oxygen and Alkali-Phosphorous-Oxygen Compounds, The Journal of Physical Chemistry 70 (1966) 4092-4095.

6. R.F. Bartholomew, A study of the equilibrium KNO3(l) = KNO3(l) + 1/202(g) over the temperature range 550-750oC. The Journal of Physical Chemistry 70(11) (1966) 3442-3446.

7. S. Gordon, C. Campbell, Differential thermal analysis of inorganic compounds nitrates and perchlorates of the alkali and alkaline earth groups and their subgroups, Analytical Chemistry 27(7) (1955) 1102-1109.

8. B. Bond, P. Jacobs, The thermal decomposition of sodium nitrate, Journal of the Chemical Society A (1966) 1265-1269.

9. R.N. Kust, J.D. Burke, Thermal decomposition in alkali metal nitrate melts, Inorganic Nuclear Chemistry Letters 6 (1970) 333-335.

10. O. Abe, T. Utsunomiya, T. Hoshino, The thermal stability of binary alkali metal nitrates, Thermochimica Acta 78 (1984) 251-260.

11. D.A. Nissen, D.E. Meeker, Nitrate/nitrite chemistry in sodium nitrate-potassium nitrate melts, Inorganic Chemistry 22(5) (1983) 716-721.

12. R.W. Bradshaw, N.P. Siegel, Molten nitrate salt development for thermal energy storage in parabolic trough solar power systems, Energy Sustainability ES (2008) Jacksonville, Florida 4p.

13. C. M. Kramer, DSC of Sodium and Potassium Nitrates and Nitrites, Thermochimica Acta 55 (1982) 11-17.


PRESENTATION - 3

Mechanism of VB2 Formation in Molten Aluminum

A. Khaliq1, 2, M.A. Rhamdhani1, 2, G.A. Brooks1, 2, J. Grandfield1, 2, 3 1Swinburne University of Technology,

Faculty of Engineering and Industrial Sciences, Melbourne, Australia 2CAST Cooperative Research Centre (CAST CRC), Australia

3Grandfield Technology Pty, Ltd, Victoria, Australia Keywords: Transition metal borides, kinetics, mechanism, molten Al, boron treatment Smelter grade aluminium is used for electrical grade conductor applications. Impurities present in solution, especially transition metals such as V, Ti, Zr and Cr reduce the electrical conductivity of smelter grade aluminium [1-2]. In the cast houses, these impurities such as V, Ti, Zr and Cr are removed by the addition of Al-B (AlB12/AlB2) master alloys, called boron treatment [1-7].The thermodynamic analysis of transition metal impurities in molten aluminium [8], it was predicted that the order of impurities removal will be from Zr and Ti to V in the temperature range of 650oC to 900oC. It was further predicted that diborides of transition metals are more stable as compared with their other possible borides in the temperature range investigated [8]. Although the process for the removal of transition metal impurities is explained in literature, limited information has been reported related to the kinetics of process and the mechanism of borides formation. For the better understanding of boron treatment of molten aluminium, kinetics experiments were performed. In this paper, selected results and the mechanism of VB2 formation in molten aluminium are presented. In the kinetic experiments, removal of vanadium was investigated by the addition of boron (stoichiometric amount to form VB2) addition in the form of Al-B (AlB12) master alloy, assuming reactions given in Eq. (1) and (2). An alloy of Al-1wt%V was prepared in the resistance pot furnace and weighed ingots of Al-B master alloy were added assuming 100% recovery of boron. Samples were taken at 0, 2, 4, 6, 8, 10, 15, 30, 45 and 60 minutes interval after the addition of boron. The experiments were performed at 700oC, 750oC and 800oC and melt was held in the clay bonded graphite crucible. Samples were analysed for vanadium in solution with aluminium using ICP-AES technique. The possible reactions during the formation of VB2 are shown in Eq. (1) & (2). The thermodynamic feasibility of transition metal borides formation has already been explained elsewhere [8].

(1)

(2) Where “[ ]” indicates that elements are dissolved in solution with molten aluminium and “(s)” represents that compounds are present in solid state. The change in concentration of [V] at 750oC and SEM image of early stage reaction are shown in Figure 1. Figure 1(a) shows that the concentration of V in solution with aluminium much less than its solubility (0.61wt %) at 750oC. Vanadium out of solution was most likely present in the form of Al10V/Al7V intermetallics. These particles were settled during the melt holding period and were observed under SEM in the borides sludge collected from the bottom


of crucible. The general observation of curve shown in Figure 1(a) also revealed that the rate of [V] removal was rapid in the beginning of reaction and slower down after 10 minutes of reaction time. This observation along with microstructural study suggested that the kinetics of VB2 formation may be controlled by two kinetic mechanisms. In the beginning, VB2 formation may be controlled by chemical reaction or liquid metal phase mass transfer and at the later stage 2 (e.g beyond 10 minutes) is most likely controlled by diffusion through the product boundary layer (VB2). The formation of product (VB2) layer has already been reported in our previous work [7] and is also shown in Figure 1(b). Similar trends of [V] removal were observed in experiments carried out at 700oC and 800oC.

(a)

(b)

Figure 5. (a) The change in [V] with time of Al-V-B alloy and (b) SEM image of Al-1wt%V alloy after the addition of Al-10wt%B (AlB12) master alloy at 750oC One of the important aspects of heterogeneous (solid-liquid) reaction is the interfacial area. Careful estimation of interfacial area will predict the kinetic of reaction close to actual situation which will be helpful in designing the chemical processes. In this particular case, AlB12 particles surface area was estimated using image analysis technique. Assuming the kinetic was controlled by the chemical reactions and follow the first order with respect to boron, the reaction rate constant was calculated for the first 10 minutes of process. The values of rate constant/mass transfer at 700oC, 750oC and 800oC were calculated to be 1.61x10-

3m/sec, 2.095x10-3m/sec and 2.685x10-3m/sec respectively. The activation energy was calculated to be 23.73kJ/mol. The curve fitting with Jander model as shown in Figure 2 (a) and other diffusion control models such as Ginstling-Brounshtein after 10 minutes of reaction gave a linear fit with R2 values more than 99%. This suggested that the later stage of reaction during the formation of VB2 is control of diffusion through VB2 shell. It could be observed from the preliminary kinetic analysis of [V] removal and the formation of VB2 in molten aluminium that the reaction mechanism is complex which show mixed control. The kinetics of reaction was most likely controlled by chemical reactions or liquid metal phase in the first 10 minutes of reaction and by the diffusion through product layer in the later stage during the boron treatment of molten aluminium. The detail investigation of the kinetic analysis is under progress that will conclude the rate limiting steps during the formation of VB2 during boron treatment of molten aluminium.


(a)

(b)

Figure 2. (a) Experimentally determined [V] conversion curves at 700oC, 750oC and 800oC and (b) The kinetics data after 10 minutes reaction were plotted using Jander model.

References 1. Gauthier, G. G. 1936. The conductivity of super-purity aluminium: The influence of

small metallic additions. J. Inst. Met., 59, 129-150. 2. Dean, W. A. 1967. Effects of Alloying Elements and Impurities on Properties.

Aluminum, 1, 174 3. Cooper, P. S. and Kearns, M. A. 1996. Removal of transition metal impurities in

aluminium melts by boron additives. Aluminium Alloys: Their Physical and Mechanical Properties, Pts 1-3, 217, 141-146.

4. Dube, G. 1983. Removal of Impurities from molten aluminium. 833068034. 5. Stiller, W. and Ingenlath, T. 1984. Industrial Boron Treatment of Aluminium Conductor

Alloys and Its Influence on Grain Refinement and Electrical Conductivity. Aluminium (English Edition), 60

6. Karabay, S. and Uzman, I. 2005. Inoculation of transition elements by addition of AlB2 and AlB12 to decrease detrimental effect on the conductivity of 99.6% aluminium in CCL for manufacturing of conductor. Journal of Materials Processing Technology, 160, 174-182.

7. Khaliq, A., Rhmadhani, M. A., Brooks, G. A., Grandfield, J., Mitchell, J. and Davidson, C., 2011. Analysis of Transition Metal (V, Zr) Borides Formation in Aluminium Melt. In: EMC (European Metallurgical Conference) June 26-29 Dusseldorf, Germany. 825-838.

8. Khaliq, A., Rhamdhani, M. A., Brooks, G. A. and Grandfield, J. 2011. Thermodynamic analysis of Ti, Zr, V and Cr impurities in aluminum melt. In: TMS 2011, February 27 - March 3, San Diego, CA. 751-756.


PRESENTATION - 4

Focus on Recycling of Critical Raw Materials: An Industry’s Perspective

Stephanie Vervynckt, Mieke Campforts Umicore Group Research & Development, Kasteelstraat 7, 2250 Olen, Belgium

Today’s transition to a low carbon economy boosts the demand of technology metals. As these metals have multiple competing applications, their supply chain is pressurized. Take for example precious metals which have show their use in catalysts for the conversion of harmful components in car exhaust gases and in industrial catalysts but are also needed in electronics such as cell phone and laptops. Because of the development of catalysts and electronics the precious metals demand has risen resulting in an increase in mining. To give some numbers, 9 times as much Ru, Pd and Ru are mined between 1980 and 2010 compared to the 80 years before. Also recycling of these products has increased as the metal prices can pay off for the processing even though in some products such as electronics only small volumes are present. This increase in metal prices is due to the potential metal scarcity that is general encountered for technology metals. Primarily the main driver is of a temporary base, namely that there is a mismatch between demand and supply due to new developments, speculation, trade barriers because of securing access to raw materials, time lag and investment risk for new mines and smelters. But in case of the technology metals, also structural scarcity needs to be taken into account. These ‘minor’ metals are only accessible as by-product of major metals due to coupled production. As a result it is not surprising that worldwide exercises are done to map the criticalness of materials and programs are setup in order to assure access to raw materials. The challenge to tackle resources scarcity is further complicated by several issues. Here four additional challenges are defined. First of all everything is connected. The high degree of linkages among resources means strong demand for one can spread to others. The production of technology metals will use energy, land, water and produce carbon dioxide. Secondly products are complex (containing multiple metals) and getting more complex with time. Moreover, element combinations that are not encountered in nature put new challenges to recycling industry as most extraction processes are developed to treat ore. This complexity is mirrored in the supply chain and the recycling process flow sheets. A third challenge is to keep materials in the cycle with the main challenge being end of life consumer goods. A last challenge is the knowledge, awareness and engagement of people on material life cycle in order to support and drive optimal closing of the loop. This presentation describes how Umicore contributes to achieve a secured supply chain through recycling and discusses which challenges are still ahead.



Recycling End-of-Life Waste Materials as Resources in EAF Steelmaking – Fundamentals of High Temperature Reactions and Industrial

Implementations

M. Zaharia(*, M. Rahman(*, N. F.Yunos, R. Khanna, N. Saha-Chaudhury, P. O’Kane(*, A. Fontana(**, J. Dicker(*, C. Skidmore(* and V. Sahajwalla

SMaRT Centre, University of New South Wales, NSW, 2052, Australia

(*OneSteel, Sydney Steel Mill, Rooty Hill, NSW (**OneSteel, Laverton Steel Mill, VIC

Keywords: waste, steelmaking, slag foaming, reduction The steel industry consumes a large amount of energy GHG emissions [1]. Rubber tires and agricultural wastes have the potential to be used in industries seeking alternative fuel and sustainable raw materials sources. Previous studies focused on recycling these materials as fuel resources, i.e. rubber in cement industry [2-3] and agricultural materials for power production [4]. The present paper focuses on investigations of carbon /slag reactions, namely slag foaming using rubber and palm shell wastes as sustainable carbon sources through quantitative estimation of the slag volume. An improved volume ratio for the rubber blend compared to coke was seen. Foaming was also improved when palm shell char was used as carbon material. Industrial implementations at OneSteel showed reductions in electrical energy and carbon consumption. These results indicate that partial replacement of coke with rubber and palm shell is efficient due to improved interactions with EAF slag. Iron and steelmaking processes use carbon as one of the main input materials. Anthracite and metallurgical coke are the conventional materials employed for reduction of iron oxides, slag foaming in EAF steelmaking. In order to address the issues of cost, availability and restrictions on greenhouse gas emissions, alternative carbon sources are required to replace, at least partially, these conventional materials. Postconsumer plastics such as HDPE (High-density polyethylene), end-of-life rubber tyres as well as palm shells contain both carbon and hydrogen. Meanwhile, expansion in the use of polymeric materials over the last 3 decades has also been accompanied by increasing problems over their disposal at the end of their life cycle. The utility and relatively short lifespan of plastics and rubbers have produced a massive waste problem. Conventional recycling technologies are unable to deal with the high volume of wastes produced. Agricultural waste materials derived from palm shells are among the main renewable waste sources in Malaysia and their quantities were seen to increase steadily in the last few years [5]. Approximately 4.7 million tonnes of palm shells were generated in 2006 alone, and a steady 5% increase has been seen in the last few years [6]. Palm shell waste is a true renewable material, so it does not contribute to the increase in the global CO2 concentration. The high temperature environments offer sustainable pathways for utilising chemical reactions to re-purpose waste materials as resources; such as reducing iron oxide to iron. The


current study describes the broad opportunities available to iron and steelmakers to utilise waste materials – ranging from polymeric materials to agricultural wastes -- as raw materials. At University of New South Wales (UNSW) a technology was pioneered that allowed polymer injection into the EAF steelmaking [7]. The focus was set on using the polymer material as a carbon replacement for coke to achieve slag foaming. Based on this in 2006 OneSteel adopted the new technique and started replacing part of the metallurgical coke with HDPE plastic. In 2007, rubber tyres were also considered and are now standard practice for the Sydney and Melbourne EAF plants [8-9]. The trials indicated that a reduction in the overall power consumption can be achieved (in the order of millions of kWh/year), therefore enhancing productivity with an associated reduction in energy [10]. This represents significant cost savings for an average EAF plant. The experimental procedure for carbon slag reactions involved: (1) reactions in a custom-made horizontal furnace having the capability to reach 1550°C, (2) visual observation using a charge-coupled device (CCD) camera, (3) off-gas analysis. Details can be found elsewhere [8]. As the current study introduces more complex materials, polymers and palm shells, the reactions occurring at the slag/carbon interface are expected to be affected by the presence of an increased level of hydrocarbons that could further decompose into carbon and hydrogen bearing products. When put in contact with an iron oxide rich EAF slag, the presence of iron oxide leads to a reduction reaction depending on the reducing agents including C, CO and H2 released at high temperature [11]. The produced gases (CO, CO2, CH4, H2O, H2) will allow the slag to foam. The rate of gas generation following the interaction of the HDPE-coke blend was seen to be the fastest, followed by the rate of gas generation from rubber-coke blends, while the lowest rate was seen when coke was the carbon material [12]. The high amount of CO and CO2 generated when HDPE and rubber replaced part of the coke could be attributed to a certain extent to the volatiles in the carbonaceous mixture. These volatiles are predominantly CH4 gas, which, at the temperature of the tests, transforms into CO and H2. Gas entrapment in the slag phase was quantified through slag volume measurements. In order to qualitatively demonstrate the slag foaming behaviour, a few representative dynamic images of the slag droplet in contact with metallurgical coke (MC) and its blends with rubber and HDPE and palm char are shown in Figure 1. The size of the slag droplet in contact with MC decreased with reaction time. HDPE blend showed significantly higher levels of gas entrapment and also rubber showed increased volume with time compared to coke In order to quantify the changes in slag volumes, attributed to formation, entrapment and release of gases, Vt/V0 was calculated and plotted as a function of time (Figure 2). Rubber/ HDPE blends and palm shell showed an increased slag volume compared to coke. HDPE blend revealed significantly higher volume ratios as a result of increased gas generation and entrapment. The volume ratio, Vt/Vo, of the EAF slag reacting with metallurgical coke records an initial value of 1.0; with time it decreased without any wide fluctuations to 0.5. This indicates a lower extent of gas entrapment by the slag. For the rubber blend, the slag volume ratios showed significantly different trends with much higher levels of droplet volumes (Figure 2a). The palm char showed fluctuations, with the drop volume


continuously growing and decreasing, however maintaining a higher volume compared to coke. Kinetics of reduction when HDPE partially replaced coke usage is reported elsewhere [11].

Figure 1 High temperature images of slag droplets in contact with 100% MC, HDPE Blend, Rubber Blend, and Palm Char at 1550ºC as a function of time [11-12]

Figure 2:Volume ratio of a) 100% MC - Rubber Blend, b) 100% MC -HDPE Blend, c) 100% MC - 100% Palm Shell interacting with EAF iron oxide slag as a function of time [15-16]

0 sec 60 sec 480 sec



HDPE Blend

Rubber Blend

100% Coke

0

1

2

3

4

0 200 400 600 800 1000Time, sec

Volu

me

Ratio

, Vt/

V 0

100%MC

MC-HDPE Blend

MC_HDPE Blend

100% MC

0

1

2

3

4

0 200 400 600 800 1000

Time, sec

Volu

me

Ratio

, Vt/

Vo

100%MC

Palm Shell

100% MC

Palm Shell

0

1

2

3

4

0 200 400 600 800 1000Time, sec

Volu

me

Ratio

, Vt/

V 0

100%MC

MC-Rubber Blend 1

MC-Rubber Blend 2

100% MC

MC_Rubber Blend 1

a)

MC_Rubber Blend 2

b)

c)


The results were subsequently confirmed, as of 2006 in industrial implementatios at OneSteel EAFs at Sydney Steel Mill, New South Wales, and Laverton Steel Mill, in Melbourne, Victoria. OneSteel has been using this technology as standard practice for over four years at SSM and LSM and 40,000 heats, consuming over one million recycled tyres [13-14]. This technology, referred to as, Polymer Injection Technology (PIT), was consequently patented, and introduced to third parties and is currently in use at UMC Metal, Thailand, where it was commissioned in May 2011 [14]. In all three cases, benefits derived from the implementation of this technology translated into cost and productivity benefits, including the reduction of foaming agent injectant, the reduction of power on time. A summary of the data for the trial heats clearly shows that the HDPE/rubber blend performs better than coke.

Table I. Summary of benefits at SSM and LSM EAF [15-16]

Injected materials Specific EE (KWh/t) Carbon (Kg/heat) FeO (%) Coke 424 462 27.6 Recycled tyres 412.4 406 26.2 HDPE 406 379 26.1

This study has shown that blends of rubber/HDPE with metallurgical coke and palm char could be used to partially replace some of the conventional metallurgical coke used in EAF steelmaking for its injecting carbon requirements. The laboratory work was reflected in the industrial implementations at OneSteel where improvements in the slag foaming behaviour and furnace efficiency were seen. References: 1. World Resources Institute, July 2009 2. E Mokrzycki, A. U.- Bocheńczyk, Applied Energy 2003, Vol 74, Issues 1–2, pp. 95-100 3. M. Bluementhal, World Cement, 1992, Vol. 23, n12, pp.14-20 4. J. L. Easterly, M. Burnham, Biomass and Bioenergy , 1996, Vol 10, Issues 2–3, Pages 79–92 5. L. Chor, Y. Laveraging on Sustainability; Malaysian Palm Oil Institute: 2010. 6. K. Mae, I. Hasegawa, N. Sakai, K. Miura, Energy & Fuels 14, 1212 (2000). 7. V. Sahajwalla, L. Hong, and N. Saha-Chaudhury: Iron and Steel Technology: AIST magazine

2006, pp. 99–96. 8. M. Zaharia, V. Sahajwalla, R Khanna, P. Koshy and P. O’Kane, ” Carbon/slag interactions

between Coke/Rubber Blends and EAF Slag at 1550°C”, ISIJ International 2009, 49(10) 9. V. Sahajwalla, M. Rahman, R. Khanna, N. Saha-Chaudhury, P. O’Kane, C. Skidmore and D.

Knights: Steel Research Int. (2009), Vol. 80(8), pp. 531–539. 10. V. Sahajwalla, R. Khanna, M. Zaharia, S. Kongkarat, M. Rahman, B. C. Kim, N. Saha-Chaudhury,

P. O’Kane, J. Dicker, C. Skidmore and David Knights: Iron and Steel Tech magazine, April 2009: 43(6)

11. J. R. Dankwah, P. Koshy, P. O’Kane and V. Sahajwalla, C. Skidmore, D. Knights, D. ISIJ International 2011, 51(3), 498–507

12. V. Sahajwalla, M. Zaharia, S. Kongkarat, M. Rahman, B. C. Kim, N. Saha-Chaudhury, P. O’Kane, J. Dicker, C. Skidmore and David Knights: Energy and Fuels, 2012, 26 (1), pp 58–66, DOI: 10.1021/ef201175t

13. A. Fontana, P. O’Kane, V. Sahajwalla, M.Zaharia, Steel research International, pp. 17-20, 2012 14. A. Fontana, P. O’Kane, D. O’Connell, & M. Schroer (2012), in Proceedings 2012 SEAISI

Conference and Exhibition, Bali, Indonesia.


PRESENTATION - 6

Recycling Lithium Ion Batteries

Elien Haccuria, P. Hayes and E. Jak Pyrosearch, The University Of Queensland, Brisbane, QLD 4072, Australia.

Keywords: lithium ion batteries, manganese, phase diagram, Al2O3-“MnO”-SiO2 slag system Introduction The lithium ion battery (LIB) market has grown considerably during the recent years and is expected to keep growing due to the application in the automotive sector. Safe recycling and extraction of valuable metals present in the batteries, including nickel, cobalt, copper and manganese, are important. Pyro-metallurgical processes involving a slag (molten oxide) phase, are especially being developed for that purpose [1-4]. Hence, the effect of manganese on the liquidus temperatures and thermodynamic properties of the possible Al2O3-CaO-“MnO”-SiO2 slag system are important. Phase diagram Al2O3 - “MnO”- SiO2 Revision of phase equilibria in the key Al2O3-“MnO”-SiO2 pseudo-ternary slag sub-system in equilibrium with metallic alloy is the focus of the present study. More specific, the discrepancies between previous investigators Roghani et al. [5], Jung et al. [6] and Snow [7] are being resolved. Experiments and preliminary results The applied experimental technique is based on the equilibration/quenching/electron probe X-ray microanalysis (EMPA) approach. Slag samples are prepared from pure oxide powers with addition of excess Mn or a Mn-Si alloy and equilibrated on a silica crucible in an argon atmosphere. The equilibration temperatures vary between 1150 and 1250 °C and the equilibration time ranges from 0 h to 24 h. Particular attention in this initial stage of the study was paid to the investigation of the system behaviour, improvements in accuracy and above all, to the confirmation of the achievement of equilibria, this included the study of i) the effect of equilibration time, ii) the homogeneity of phases, iii) the attainment of equilibria from different directions, and iv) the analysis of the

reactions taking place in samples during experiments. It was found that a Mn5Si3 intermetallic compound is formed in equilibrium with the slag at investigated conditions after more than 6 hours equilibration. If pure metallic Mn was added to the slag, the following changes in bulk composition were identified (Figure 1). The slag phase initially enriched with MnO due to the reactions between alloy and slag systems. Subsequently, the SiO2 concentration in the slag increases due to the dissolution of the substrate. Although initial work on the improvement of

Figure 1: Changes in bulk composition if pure metallic Mn is added to the slag


accuracy is still in progress, preliminary results of the liquidus in the tridymite and rhodonite primary phase fields can be analysed. Figure 2 presents some preliminary results obtained from the samples equilibrated for 0.5, 1, 2 and 24 hours. A range of local areas with different liquid phase compositions were found in the samples equilibrated for 0.5 and 1 hour. If local equilibrium is achieved between liquid, solid (Tephroite Mn2SiO4 and Rhodonite MnSiO3), and alloy system, then the points may represent equilibrium. This assumption is yet to be confirmed, however consistency between results from 2, 6 and 24 hours samples are an initial indication for that. These measurements are in agreement with Roghani [5] (Figure 2). Note that these results are preliminary and cannot be taken as final equilibrium results. Further work is going on to verify these preliminary findings.

Figure 2: Preliminary liquid phase composition measurements in local areas of incompletely equilibrated samples at 1200 °C in comparison with the 1200 °C liquidus isotherms by Roghani et al. [5] and Jung et al. [6]. Phases present: R) rhodonite, Tr) tridymite, Te) tephroite, A) alloy Several other issues are also under investigation. For example, a tridymite ring is formed around the alloy particles in samples equilibrated for two hours or more, resulting in the seclusion of the alloy from the liquid (Figure 3). The mechanism for this ring formation and measures to avoid it, are being investigated. The research into such reactions is essential to ensure accuracy and reliability of the final results.

Figure 3: Alloy enclosed by a tridymite ring after longer equilibration times Phases present: R) rhodonite, Tr) tridymite, A) alloy

L + R+ Te

L + R+ Tr

L + R

L + Tr

1 h 6 h 24 h

R

R

R

A

Te

Tr

Tr

A A

A

R

R

Tr Tr

1200 °C


Future plans Further study is necessary to further improve the accuracy of the results. Additional research will be focused on resolving the discrepancies in the Al2O3-“MnO”-SiO2 pseudo-ternary slag sub-system. The work is planned to be expended to the quaternary system containing CaO. References 1. Maschler, T., et al., Development of a recycling process for Li-ion batteries. Journal

of Power Sources, 2012. 207: p. 173-182. 2. Dewulf, J., et al., Recycling rechargeable lithium ion batteries: Critical analysis of

natural resource savings. Resources, Conservation and Recycling, 2010. 54: p. 229-234.

3. Xu, J., et al., A review of processes and technologies for the recycling of lithium-ion secondary batteries. Journal of Power Sources, 2008. 177: p. 512-527.

4. Umicore. Umicore Battery Recycling. 2012; Available from: http://www.batteryrecycling.umicore.com/UBR/process/.

5. Roghani, G., E. Jak, and P. Hayes, Phase equilibrium studies in the "MnO"-Al2O3-SiO2 system. Metallurgical and materials transactions B, 2002. 33B: p. 827-38.

6. Jung, I., et al., Thermodynamic evaluatioin and optimization of the MnO-Al2O3 and MnO-Al2O3-SiO2 systms and applicatioins to inclusion engineering. Metallurgical and materials transactions B, 2004. 35B: p. 259-268.

7. Snow, R.B., Equilibrium relationschips on the liquidus surface in part of the MnO-Al2O3-SiO2 system. Journal of The American Ceramic Society, 1943. 26(1): p. 11-20.


PRESENTATION - 7

Kinetics of Silicon Refining using Slag Treatment

Md Saiful Islam, M. Akbar Rhamdhani, Geoffrey A. Brooks Faculty of Engineering and Industrial Sciences, Swinburne University of Technology,

Victoria, Australia [email protected]

Keywords: silicon; solar grade silicon; metallurgical grade silicon; kinetics

Silicon, an important semiconducting material and alloying element in metallurgy and chemistry, is one the most abundant elements in the earth as oxides and silicates. The rapid growth of solar cell demand is creating a shortage of solar-grade silicon feedstock. Expensive high-purity scrap silicon (99.9999999% Si) is mainly used as the raw material to produce solar-grade silicon (SOG-Si) (99.9999% Si). Many researchers have reported that relatively inexpensive metallurgical grade silicon (MG-Si) (98-99% Si) can be used as an alternative raw material. Of all the impurities present in MG-Si, boron and phosphorus are usually the most difficult to remove. Slag refining is one of the few metallurgical methods for the efficient removal of boron from silicon. In order to produce silicon for photovoltaic applications, the relationship between the slag composition and the mass transfer rate of boron from silicon to slag is of great importance. The slag treatment on MG-Si for SOG-Si production is based on the principle of liquid-liquid extraction similar to those applied in the steel industry. Impurities with a higher oxygen affinity, in comparison with silicon, oxidise and pass into the slag. The thermodynamics of the oxidation removal of boron by the slag treatment have been of interest to researchers, and numerous studies have been carried out on slag refining [1-2]. In the case of boron oxidation, the equilibrium reaction is based on Equation 1. From Equation 1, in order to remove boron from molten silicon by the slag treatment, it needs to maintain high oxygen potential, which is restricted by Si/SiO2 equilibrium at the interface between molten silicon and slag, and simultaneously decreases the activity of borate in the slag. Because borate is an acidic oxide, its basicity should be kept high with basic oxides (CaO or Na2O). In this method, liquid silicon is treated with CaO-SiO2, CaO-SiO2-CaF2, CaO-SiO2-Al2O3, CaO-SiO2-Al2O3-MgO and other molten slags [3].

(1)

During refining, the rate of boron removal into the slag where the refining takes place is of great importance. It is therefore useful to assess the removal rate of boron with the equilibrium boron distribution. A list of selected studies on kinetics of boron removal from silicon and mass transfer coefficient values are shown in Table 1. The relationship between slag composition and optimal refining efficiency is still not well known, so a better understanding about the relationship between slag composition and mass transfer can improve the basic understanding of the refining process.


Table 1: Selected kinetics experiments for boron removal

Investigators Initial Slag & Alloy Composition Experimental Conditions Mass Transfer

Coefficient (m/s) Rate controlling

step

(Krystad, Tang et al. 2012) [4]

SiO2-CaO CaO/SiO2=1 Si with 250 ppm B

Electric resistance furnace Temperature=1873K;atm=Ar Equilibrium time=3hr Slag/Silicon mass ratio=1 & 2

2.1 x 10-6 (S/M=1) --

1.2 x 10-6 (S/M=2)

SiO2-CaO-20%MgO CaO/SiO2=1 Si with 250 ppm B

Electric resistance furnace Temperature=1873K;atm=Ar Equilibrium time=3hr Slag/Silicon mass ratio=2

3.2 x 10-6 --

(Nishimoto and Morita 2011) [5]

SiO2-CaO CaO/SiO2=1.22 Si with 300 ppm B

Electric resistance furnace Temperature=1873K;atm=Ar Equilibrium time=3hr Slag/Silicon mass ratio=2.5

1.4 x 10-6 Mass transport in slag

The kinetics of boron removal from liquid silicon during slag refining have been investigated in this research by means of several small-scale experimental studies at temperatures of 14500C to 16500C. Slag and silicon, in batch weights of 7 g, were heated together in an Alumina crucible placed in a resistance-heated tube furnace. The slags were produced from powdered SiO2, CaO, and Al2O3. Experiments were carried out at slag-to-silicon ratios of 1.5, 2 and 2.5, where the silicon initially contained approximately 350 ppm boron. From the results of kinetics of boron removal from molten silicon through slag (CaO-SiO2, CaO-SiO2-10%Al2O3 & CaO-SiO2-15%Al2O3) was investigated.

Figure 1: Changes in boron contents in silicon during the reaction with CaO-SiO2 slag

Figure 1 shows the change in boron contents during the experiments for CaO-SiO2 slags. The experimental results show that the rate of boron removal was very high at the beginning of the reaction; this rate dropped to zero after 120 min of the reaction, depending on the ratio of the amount of silicon to slag. In this study, the rates of boron removal were found to increase with a decrease in the initial weights of the silicon for both types of slags. The data shown in Figure 1 can be used to analyze the reaction kinetics by considering several kinetics models. The data plotted using models assuming slag and silicon mass transfer controlled. The results in Figure 2 show that all the data can be fitted into a straight line for silicon phase mass


transfer controlled. Therefore, it can be argued that the overall kinetics was controlled by mass transfer of B in silicon phase. Currently, further investigations are being carried out to confirm and evaluate the detailed mechanism.

Figure 2: Integrated rate plots obtained using the kinetic equation of (a) silicon mass-transfer control and (b) slag mass-transfer control for CaO-SiO2 slag References 1. Suzuki, K., T. Sugiyama, et al. (1990). "Thermodynamics for removal of boron from

metallurgical silicon by flux treatment." Journal of the Japan Institute of Metals 54(2): 168-172.

2. Johnston, M. D. and M. Barati (2010). "Distribution of impurity elements in slag-silicon equilibria for oxidative refining of metallurgical silicon for solar cell applications." Solar Energy Materials and Solar Cells 94(12): 2085-2090.

3. Teixeira, L. and K. Morita (2009). "Removal of Boron from Molten Silicon Using CaO–SiO2 Based Slags." ISIJ international 49(6): 783-787.

4. Krystad, E., K. Tang, et al. (2012). "The Kinetics of Boron Transfer in Slag Refining of Silicon." JOM Journal of the Minerals, Metals and Materials Society: 1-5.

5. Nishimoto, H. and K. Morita (2011). "The Rate of Boron Elimination from Molten Silicon by Slag and Cl2 gas Treatment." Supplemental Proceedings: Materials Processing and Energy Materials, Volume 1: 701-708.


PRESENTATION - 8

Minor Element Distributions during Copper Flash Converting

Douglas R. Swinbourne RMIT University, Latrobe Street, Melbourne, VIC, 3122, Australia

Keywords: Modelling, Copper, Flash converting, Minor elements, Distributions.

Peirce-Smith converting of copper matte is simple and reliable, but has many well-documented disadvantages[1]. They have largely been overcome by the advent of continuous converting technologies such as the Mitsubishi bath converting process and the Kennecott flash converting process[2]. Both processes use lime (CaO) rather than silica as a flux and so produce calcium ferrite slag. Flash converting was pioneered at the KUCC smelter in Utah, USA, and has been described by Newman et al.[3] The control of minor elements is an important issue for all copper smelters. Kaur et al.[4] reported on minor element deportments (lead, arsenic, bismuth, cadmium and molybdenum ) at KUCC. HSC Chemistry for Windows v.7.1 was used to predict the distributions of these minor elements and the results were compared to published data. The first requirement is a reliable set of operating data. The most complete data set is for 2001 operations[2] The masses of dusts consumed/produced are given but not their compositions. The copper mass balance is near to closing but the iron balance is not. Kaur et al.[4] provided the most recent set of data but contains no information on dust nor the quantity of lime (CaO) flux. Again the copper and iron balances do not close. Neither data set gave minor element concentrations in matte so the data of Asteljoki and Kyto[5] was assumed, although the actual amounts are not essential when calculating fractional distributions. The input masses based on the operating data sets and the assumed activity coefficients, all taken from the literature, are given in the table below. No data for CdO was found so it was assumed to be the as that of ZnO. Where activity coefficients are not given, they are unity.

Table 1. Species, the masses [ ] and activity coefficients ( ) used in the HSC model. gas matte slag copper solid O2(g) CuS0.5 [855 kg] CaO (1) [23 kg] Cu (1) FeO1.33 N2(g) FeS [122 kg] FeO1.33 (1.7) CuS0.5 (26) S2(g) PbS [14 kg] FeO (35) CuO0.5 (20) SO2(g) AsS1.5 [6 kg] CuO0.5 (3) Fe (10) SO3(g) BiS1.5 [2 kg] PbO (2) Pb (5) Pb(g) CdS [1 kg] AsO1.5 (0.2) As (0.005) PbO(g) BiO1.5 (0.8) Bi (2.4) PbS(g) CdO (2) Cd (0.73) AsO(g) Bi(g) Cd(g)


The mass of oxygen required to yield blister copper containing 0.2 wt% S was determined as 258kg/tonne matte, then all other process parameters were calculated at that level of oxygen addition. The copper content of the slag as a function of the sulphur content of the blister copper is shown below and the operating point for the KUCC converter is very close to the calculated equilibrium value. The figure shows why aiming for low sulphur contents in blister copper,

which would lessen the load on the anode refining furnaces, carries a significant penalty. The distribution of lead by species is given to the left. More lead would report to the slag and less to the blister copper if the oxygen addition was increased past the target level of 258 kg, but the change in the proportion reporting to the waste gas would not change much. It can be seen that PbO(g) and Pb(g) are present in similar proportions at the target oxygen addition.

The distribution of arsenic by species is given below. It matches the industrial data well, although the amount of arsenic being volatilised is underestimated. This could be because in practice the flash furnace is an open system i.e. gas is passing through the furnace and this would increase the amount of volatile species leaving the furnace. The very small amount of volatile arsenic as AsO(g) is striking because arsenic is widely regarded as a volatile element.

The reasons for this are that AsO1.5 has a low activity coefficient in calcium ferrite slag due to the basic nature of this slag, and because arsenic has a very low activity coefficient in copper. About one half of all arsenic entering the converter reports to the blister copper. Poor arsenic elimination is a feature of continuous converting processes, as opposed to the Peirce-Smith converter where significant arsenic elimination occurs during the first stage slagging blow before any copper forms.


The fractional distributions of minor elements between offtake gas, slag and blister copper at an oxygen addition of 258 kg are given in the Table below, together with the data from Kaur et al.[4]. The agreement is generally excellent.

Table 2. Comparison of fractional distributions with published data for the flash converter

Element Phase Mass fraction (%) Published This work

lead blister 33 33 slag 57 57 gas 10 10

arsenic blister 52 60 slag 42 37 gas 6 3

bismuth blister 73 68 slag 14 25 gas 13 7

cadmium blister - 4 slag 8 9 gas 92 87

The success of this thermodynamic model indicates that the Kennecott flash converter can be regarded as approximating an equilibrium reactor and that minor element distributions appear to be controlled mostly by thermodynamic factors.

References 1. K.J. Richards, D.G. George & L.K. Bailey: Advances in Sulfide Smelting, Proc. Fall

Meeting of TMS-AIME, San Francisco, Nov. 6-9, 1983, 489-498. 2. W.G. Davenport, M. King, M. Schlesinger & A.K. Biswas: “Extractive Metallurgy of

Copper”, Pergamon, Great Britain, 4th edition, 2002. 3. C.J. Newman, D.N. Collins & A.J. Weddick: Copper 99-Cobre 99 International

Conference, Phoenix, Arizona, 1999, vol. 5, The Met. Soc. of CIM, 29-45. 4. R. Kaur, C. Nexhip, M. Wilson & D. George-Kennedy: Proceedings of Copper 2010,

Hamburg, GDMB, 2010, vol.6, 2415-2432. 5. J.A. Asteljoki & S.M.I Kytö: TMS Paper A86-57, The Minerals, Metals & Materials

Society, Warrendale, PA, 1986.



Integrated Experimental and Modelling Research Methodology for Phase Equilibria, Thermodynamics and Viscosities of Metallurgical Slags

Evgueni Jak

PYROSEARCH, The University of Queensland, Brisbane, Queensland, 4072, Australia. [email protected]

Keywords: metallurgical slags, phase equilibria, thermodynamic modelling, liquidus. Coupled experimental and modelling studies are combined into an integrated research program on phase equilibria, thermodynamics and viscosities of the metallurgical slag systems. Key issues derived from experiences in continuing development and application of both experimental and thermodynamic modelling research are outlined. Particular emphasis is given to the details of the research methodologies, analysis of reasons for uncertainties and the ways to continuously improve the accuracy of both studies. The ways how the advanced research tools can be implemented into industrial operations are presented. Experimental part of the phase equilibria study involves high temperature equilibration in controlled gas atmospheres, rapid quenching and direct measurement of equilibrium phases with electron probe X-ray microanalysis (EPMA). Thermodynamic modelling undertaken using computer package FactSage with the quasi-chemical model for the liquid slag phase is closely integrated with the parallel experimental research. Experiments are planned to provide specific data for thermodynamic model development as well as for pseudo-ternary liquidus diagrams which can be used directly by process operators. Thermodynamic assessments are used to identify priorities for experiments. Experimental and modelling studies are combined into an integrated research program contributing to and enhancing outcomes of each other and of the overall program. The continuous development of experimental methodologies has brought significant advances. Importantly, these novel approaches enable measurements to be made in systems that could not previously be characterised, for example, due to uncontrollable reactions with container materials or changes in bulk composition due to vapour phase reactions. The approach, however, requires particular attention to ensure accurate information is obtained. An ongoing dedicated program of improving accuracy of all possible elements of the research revealed a number of possible sources of uncertainties and the ways developed to mitigate those shortcomings are systematically summarised in this paper. The thermodynamic modelling has progressed significantly, and achieved a level of prediction of phase equilibria and thermodynamics of complex multi-component multi-phase systems with improved accuracy. The adequate description of the systems however requires a combination of various types of data and still demands continuous further development. The outcomes of both experimental and modelling studies are applied to assist in improvements of the industrial metallurgical operations. High certainty of the predictions of the behaviour of complex industrial processes provides a strong basis for optimisation of operations. The stage of implementation of the outcomes of the laboratory experimental and theoretical modelling, however, is frequently overlooked, but requires high level of research expertise to establish the actual conditions in the real industrial process and relate them to the advanced laboratory and theoretical research tools.


PRESENTATION - 10

Aluminium Production Route through Carbosulfidation of Alumina utilising H2S

Nazmul Huda1, M. Akbar Rhamdhani1, G.A. Brooks1, B. J. Monaghan2, L. Prentice3

1HTP Research Group, Swinburne University of Technology, VIC 3122, Australia 2PYRO Research Group, University of Wollongong, NSW 2522, Australia

3CSIRO Process Science and Engineering, VIC 3169, Australia

Keywords: Aluminum, carbosulfidation, H2S Indirect carbothermal reduction of alumina for the production of aluminum utilizes different reducing agents to convert alumina into intermediate aluminum compounds. In the present study, the carbosulfidation route for aluminum production utilizing H2S(g) as the reductant and sulfur source has been investigated, in particular the formation of Al2S3 in the first step of the process. The results of the thermodynamic analysis predicted that conversion of Al2O3(s) to Al2S3(l) significantly increases above 1400°C at 1 atmosphere pressure. Experimental investigations were carried out at temperatures of 1100 to 1500°C using dilute H2S(g) gas in argon. The reaction products were analyzed using scanning electron microscopy (SEM), energy dispersive X-ray spectroscopy (EDS), X-ray diffraction (XRD), inductively-coupled plasma absorption emission spectroscopy (ICP-AES) and chemical filtration. The X-ray diffraction results confirmed the presence of Al2S3(s). Percentage of conversion from Al2O3 to Al2S3 was found to be over 80% at 1500°C.

Equilibrium Calculations of Al2O3-C-H2S Reaction Systems The equilibrium calculations were carried out using FactSage 6.1 thermodynamic package. The equilibrium calculations for Al2O3-C-H2S system were carried at temperatures 1000°C to 2000°C at different pressures. For all equilibrium calculations, 3 moles of C and 3 moles of H2S were considered for 1 mole of Al2O3. Figure 1 shows equilibrium calculation of Al2O3+3C+3H2S for temperature range of 1000 to 2000°C at 1 atm pressure. Figure 1(b) show that significant amounts of gases are produced with majority of H2(g) and CO(g) at higher temperatures. Al2S3 is predicted to be the main intermediate aluminum compound when H2S is reacted with Al2O3 and C at 1000 to 2000°C at 1 atmospheric pressure. Formation of Al2S3 is predicted to be very low at 1100 to 1300°C at 1 atm pressure (0.1012 mol Al2S3/ mol Al2O3) and predicted to increase with increasing temperature to 1800°C. Formation of CO is predicted to be lower at 1100°C (0.035 mol/mol Al2O3) and significantly increases with increasing temperature (2.6 mol/mol Al2O3 at 1800°C). Along with CO and other gases significant amount of H2(g) gas is also predicted to form at 1100°C (1.37 mol/ mol Al2O3). This content of H2(g) was predicted to increase to 2.62 mol/mol Al2O3 when temperature is at 1800°C.


a) Predicted condensed phases b) Predicted gaseous phases

Figure 1: Predicted equilibrium phases in the Al2O3+3C+3H2S system at T = 1000°C to 2000°C, at 1 atm pressure: a) condensed phases, b) gaseous phases

Experimental results

Experimental investigation on carbosulfidation of Al2O3(s) by using C(s) and dilute H2S(g) (5% H2S – 95% Ar) at different temperatures (1100 to 1600 °C) and reaction duration were carried out using a horizontal tube resistance-furnace (Nabertherm RHTV 200-600). A schematic diagram of the experimental setup is shown in Figure 2.

Figure 2: A schematic diagram of the experimental set up using a horizontal tube furnace Figure 3 shows the comparison of XRD pattern of the samples after experiments at 1400°C for three different times (3, 6 and 9 hours). Al2O3 and Al2S3 peaks are marked by “1” and “2”, respectively. As shown in Figure 3, significant aluminum sulfide (Al2S3) was detected after 6 and 9 hours of reaction. This is indicated by the higher and sharper Al2S3 peaks at 6 and 9 hours compared to those from at 3 hours. Al2O3 peaks are still present, indicated that some Al2O3 remains and unreacted in the samples. However, it can also be seen clearly that there is a gradual decrease of the intensity with increasing reaction time.


Figure 3: X-ray diffraction pattern of the samples after 3, 6, and 9 hours experiments at 1400 °C. (1 = corundum (Al2O3), 2 = aluminum sulfide (Al2S3) and 3 = Graphite (C)) The percentage of conversion from Al2O3 to Al2S3 was determined by chemical dissolution and filtration. As pure Al2S3 completely dissolves in hydrochloric acid (HCl), a portion of the experimental samples were dissolved in HCl (36% w/w aqueous solution) and the solution was then filtered out. The amount of mass that dissolves in HCl represents the formed Al2S3 while the residues are the unreacted Al2O3 and C. From the filtration results, the percent of conversion ( ) of Al2O3 to Al2S3 was calculated using following equation:

The details of calculated conversion from selected experiments are shown in Table I. The highest conversion was found for experiment at 1500°C and 9 hours duration. The conversion showed an increasing trend with respect to time and temperature.

Table I: The conversion of Al2O3 to Al2S3 from selected samples at 1400°C and 1500°C

Temperature (°C)

Duration (hours)

Weight of Sample (g)

% of Conversion

( )

1400 6 0.2012 75.4 9 0.2051 77

1500 6 0.2186 78.9 9 0.2060 81.6

In summary, the results, from XRD, SEM, EDS, ICP and conversion calculation, indicate that it is possible to form high amount of Al2S3 from Al2O3 using C and H2S gas in the range of conditions studied. The results also suggest that the conversion to Al2S3 increases with increasing temperature and duration of experiments.


PRESENTATION - 11

Investigation of Freeze-linings in Copper-Containing Slag Systems

Ata Fallah Mehrjardi, Peter C. Hayes, Evgueni Jak PYROSEARCH, The University of Queensland, Brisbane, Australia

Keywords: Freeze-lining, slag, copper production, deposit

Ways of increasing productivity have been always a challenging issue for pyrometallurgical processes. Increasing the throughput of the reactor through higher temperature and vigorous agitation in bath are possible options for enhancing the kinetics of reactions in the multi-phase system processes. On the other hand, these measures also lead to rapid degradation of the refractory and premature shutdown of the reactor for relining, imposing additional costs on processes in the form of planned and unplanned maintenance. An alternative solution to this problem is the formation of a slag freeze lining rather than direct contact of refractory layers with the hot bath. Slag freeze-linings are increasingly used in industrial pyrometallurgical processes to ensure furnace integrity is maintained in aggressive high temperature environments 1-4). Most previous studies of freeze-linings have analysed the formation of slag deposits based solely on heat transfer models 5, 6). The focus of the present research is to determine the impact of slag chemistry and local process conditions on the microstructures, thickness, stability and heat transfer characteristics of the frozen deposit.

To gain a full understanding of freeze lining formation and the effect of experimental variables on the stability, thickness and heat transfer characteristics of the freeze lining two particular approaches have been adopted. First, a 1-D heat transfer model in cylindrical coordinates was developed to approximately evaluate the deposit thickness and interpolate temperature distribution in the freeze lining as a function of key process variables such as coolant flow rate, thermal conductivity, bath convection and superheat. Second, experimental studies of the freeze lining formation and kinetics at steady state condition were undertaken including cold finger and supporting experiments. The cold finger was immersed into a synthetic slag bath heated by an induction furnace. The temperature profile across the deposit and bath was measured.

A Cu-Fe-Si-Al-O slag was selected for study; the liquidus temperature for the slag in equilibrium with metallic copper has been experimentally determined to be approximately 1140◦C and the primary phase under these conditions has been found to be delafossite. The phase assemblages and microstructures of the deposits formed in the cold finger experiments differ significantly from those expected from equilibrium considerations. The freeze-lining deposits have been found in general to consist of several different layers. Starting from the cold wall these layers consist of glass; glass with microcrystalline precipitates; multiphase sub-liquidus material containing delafossite and cuprite crystal phase assemblages and high-silica metastable liquid that was separated from the bulk liquid (closed crystalline layers); phase assemblages containing delafossite and cuprite crystals and a high-silica liquid phase that is connected to the bulk liquid (open crystalline layers), and the outer layer containing a complex mixture of liquid and solid phases.


It has been previously widely assumed that the interface between the stationary frozen layer and the molten bath at steady-state consists of the primary phase in contact with the bulk liquid at the liquidus temperature, Tliquidus. It has been shown in the present laboratory-based studies through the use a of cold finger technique that, at steady-state and in selected ranges of process conditions and bath compositions, the phase assemblage present at the deposit/liquid interface is not that of the primary phase alone. The microstructural observations clearly demonstrate that the temperature of the deposit/liquid bath interface, Tf, can be lower than the liquidus temperature of the bulk liquid, Tliquidus. These observations point to a significant change in proposed mechanism and behaviour of the systems.

Acknowledgements The authors would like to thank the Australian Research Council Linkage program, Rio Tinto Kennecott Utah Copper Corp., Xstrata Technology, Xstrata Copper, BHP Billiton Olympic Dam Operation and Outotec Finland Oy for their financial support. Appreciations are extended to all PYROSEARCH and CMM staff at the University of Queensland.

References 1. M. Campforts, B. Blanpain, and P. Wollants, "The importance of slag engineering in

freeze-fining applications". Metall. Mater. Trans. B., 2009. 40B: p. 643-655. 2. M. Campforts, et al., "Freeze-lining formation of a synthetic lead slag: Part

I.microstructure formation". Metall. Mater. Trans. B., 2009. 40B: p. 619-631. 3. M. Campforts, et al., "Freeze-lining formation of a synthetic lead slag: Part II. thermal

history". Metall. Mater. Trans. B., 2009. 40B: p. 632-642. 4. M. Campforts, et al., "On the microstructure of a freeze lining of an industrial

nonferrous slag". Metall. Mater. Trans. B., 2007. 38B: p. 841-851. 5. K. Verscheure, et al., "Continuous fuming of zinc-bearing residues: Part II. The

submerged-plasma zinc-fuming process". Metall. Mater. Trans. B., 2007. 38B: p. 21-33.

6. F. Guevara and G. Irons, "Simulation of Slag Freeze Layer Formation: Part II: Numerical Model ". Metall. Mater. Trans. B., 2011. 42: p. 664-676.


PRESENTATION - 12

CIRCOSMELT: Outotec’s Alternative Ironmaking Process

Ross Baldock

Outotec Pty Ltd Melbourne Australia

Current high iron ore prices have resulted in increased interest in alternative iron sources, which cannot be processed through conventional routes such as blast furnace or shaft furnace respectively rotary kiln based direct reduction.

Outotec has been developing circulating fluidized bed (CFB) processes for more than 50 years. One such CFB application is the coal based direct reduction process - Circofer®.

Since Ausmelt joined the Outotec group in 2010, further development of the Circosmelt process, combining the Circofer prereduction with the coal based AusIron® smelting reduction has taken place.

The Circosmelt process uses fines, thus avoiding agglomeration, such as sintering or pelletizing, and can utilize a wide range of bituminous and sub bituminous coals.

Test work in Outotec’s pilot reduction and smelting plants with different raw materials and a variety of coals has been performed.

This presentation will give an overview of the Circosmelt process principles. Furthermore results achieved during pilot plant campaigns with different raw materials will be discussed.



Insights into the Formation of Iron Ore Sinter Bonding Phases

Mark I. Pownceby1 and Nathan A.S. Webster1,2 1CSIRO Process Science and Engineering, Box 312, Clayton South, VIC, 3169, Australia

2Australian Nuclear Science and Technology Organisation, Locked Bag 2001, Kirrawee DC, NSW, 2232, Australia

Keywords: Iron ore sintering, ‘SFCA’, phase equilibria, in situ X-ray diffraction. During the iron ore sintering process, iron ore fines (<6.3mm) are mixed with limestone flux and coke breeze and heated to ~1300ºC. This results in partial melting of the mixture, converting the loose raw materials into a porous but physically strong composite material in which the iron-bearing minerals are bonded together by a range of complex ferrite-like phases known collectively as ‘SFCA’ (Silico-Ferrite of Calcium and Aluminium). These ‘SFCA’ phases can be divided on the basis of composition and morphology into two main types: a low-Fe form that is simply referred to as SFCA [1], and a second high-Fe, low-Si form called SFCA-I [2]. SFCA and SFCA-I are believed to be the most desirable bonding phases in iron ore sinter because of their high reducibility [3], high mechanical strength and low reduction degradation [4,5], all of which are significant factors in determining the efficiency of the blast furnace.

Figure 1. Schematic showing; a) a typical mixture of iron ore fines, and b) a typical iron ore sinter product composed of hematite nuclei (ore) and a porous ‘SFCA’ matrix. Also shown are photomicrographs showing typical SFCA-I (c) and SCFA (d) matrix microtextures. Mt = magnetite (white), G = glass (dark).

Despite their importance in controlling the quality of iron ore sinter, the stability range and mechanisms of ‘SFCA’ formation from precursor phases are not well understood. CSIRO has attempted to improve understanding of phase relations within iron ore sinter by; a) conducting experimental phase equilibria studies within the Fe2O3-CaO-SiO2 (FCS) and Fe2O3-Al2O3-CaO-SiO2 (FACS) model sinter systems to establish the key thermal and compositional parameters that influence the bonding phase chemistry and stability [6,7,8,9], and, b) conducting in situ X-ray diffraction experiments to determine the formation mechanisms of SFCA and SFCA-I under simulated sintering conditions [10]. The combination of techniques will provide insights into sinter formation mechanisms which are

SFCA – I (Good matrix)

50 m

SFCA (Poor matrix)

100 m Mt

G

CaCO3 flux

Ore

Coke

Ultra-fine coating

SFCA phases, other Ca-rich ferrites, calcium silicates, glass, Fe3O4/Fe2O3

Ore

Pore

a) b)

c) d)


important in designing strategies to counter some of the key problems facing the Australian iron ore industry. FCS System in Air and at pO2 = 5x10-3 atm. Results for the FCS system in air show that above 1270°C a continuous Liquid-only phase field extends from Fe2O3-rich to SiO2-rich compositions. Below 1270°C, the Liquid-only phase field segregates into two melt regions; one at high Fe2O3 contents and basicities (CaO/SiO2) greater than B=2.0 and a second melt field at low basicities (B<2.0) and high SiO2 contents. The melt field at high basicities results in Ca ferrite(s) crystallising upon cooling while the low basicity melt field crystallises Ca silicate(s) and hematite. At T≤1240°C (Fig. 2a) a new phase, SFC (alumina free variety of SFCA) forms part of the crystallising assemblage. In equivalent experiments at low pO2, the liquidus is depressed with the single Liquid-only melt field remaining until below 1255-1250°C. SFC could not be produced as a single crystalline phase under reduced oxygen partial pressures (Fig. 2b).

Figure 2. Summary of phase relations determined for the Fe2O3-CaO-SiO2 ternary system at 1240ºC [6,8]. Figure 2a is from experiments in air, whilst Fig. 2b is from experiments at 5x10-3 atm. H = hematite; Mt = Magnetite; Lα = High Fe-liquid; Lβ = High Si-liquid; C2F = 2CaO.Fe2O3; C2S = 2CaO.SiO2; CS = CaO.SiO2; SFCss = silico ferrite of calcium solid solution (hatched line in Fig. 2a). FACS System in Air At high temperatures, the initial effect of adding small amounts of Al2O3 (<1%) to the FCS system is to lower the liquidus temperature. As more alumina is added, Al2O3 begins to substitute into SFC forming SFCA. Alumina-free SFC has a melting point of ~1250-1255°C [7] however incorporation of Al2O3 via solid solution to form SFCA increases the melting point and stabilises SFCA to higher temperatures [9]. Mechanism of Formation of SFCA and SFCA-I Figure 3 shows synchrotron-based XRD data collected for a sinter mixture (4/5) with composition 77.36 wt% Fe2O3, 14.08 wt% CaO (added as CaCO3), 3.56 wt% SiO2, and 5.00 wt % Al2O3 (added as Al(OH)3). The sample was heated to 1350ºC under a pO2=5 10-3 atm. and then cooled back down to room temperature to observe phase formation during cooling. During heating, SFCA-I formation was associated with reaction between Fe2O3, 2CaO.Fe2O3 (C2F, i.e. C=CaO and F=Fe2O3) and SiO2. SFCA formation was associated with reaction of CF, SiO2, and a new phase designated CFA with average composition 49.60, 9.09, 0.14, 7.93 and 32.15 wt% Fe, Ca, Si, Al and O, respectively. Increasing Al2O3 concentration in the starting sinter mixture increased the temperature range over which SFCA-I was stable before

SiO2

Fe2O360 70 80 90CaO

30

20

10C S + L2 L

C S + C F + L2 2

Mt + C S + L2

C S + L2 L

Mt + CS + LL

Mt + LL

MtHem

C F + L2 Mt + L (Fe O )3 4

( SFC )ss

L

LL

SiO2

H + CS + L 30

20

10

H + C S + SFC2 ss

H + SFCss

Fe2O360 70 80 90C F + L2 H + SFC + Lss

CaO

SFC + Lss L

C S + L2 L

H + C S + L2

H + L

C S + SFC + L2 ss

C S + C F + L2 2

a) b)


the formation of SFCA, and stabilised SFCA to higher temperature before it melted to form a magnetite Fe3O4+melt phase assemblage (1213-1308ºC). During cooling, the first phase to crystallise from the melt was a phase, designated Fe-rich SFCA, with average composition 58.88, 6.89, 0.82, 3.00 and 31.68wt% Fe, Ca, Si, Al and O, respectively. At lower temperatures, SFCA crystallised. Increasing Al2O3 increased the temperature of Fe-rich SFCA crystallisation, and increased the temperature of SFCA crystallisation, during cooling. Figure 4 summarises the proposed SFCA and SFCA-I formation mechanisms.

Figure 3. In situ XRD data for sinter mixture SM4/5. Annotated on the plot are: the low-temperature (<650ºC) phase transformation (α βSiO2) and decomposition (e.g. CaCO3 CaO) events; the formation events of C2F, CF, SFCA-I, SFCA and the Fe3O4+melt phase assemblage during heating; and the Fe-rich SFCA, SFCA and Fe2O3 formation events during cooling [10].

Figure 4. Summary of the reaction sequences in the formation of SFCA phases during heating (zone 1) and cooling (zone 3) in the range 25-1350ºC and at pO2=5 10-3 atm. Zone 2 is the region where Fe3O4 and melt coexist [10]. Acknowledgements ANSTO is acknowledged for partial funding of this research. The in situ experiments were undertaken on the powder diffraction beamline (10BM1) at the Australian Synchrotron under beamtime awards AS093/PD1639 and AS113/PD4160.


References 1. J.D.G. Hamilton, B.F. Hoskins, W.G. Mumme, W.E. Borbidge and M.A. Montague,

“The crystal structure and crystal chemistry of Ca2.3Mg0.8Al1.5Si1.1Fe8.3O20 (SFCA): solid solution limits and selected phase relationships of SFCA in the SiO2-Fe2O3-CaO(-Al2O3) system,” Neues Jahrbuch für Mineralogie, Abhandlungen, Vol.161, 1989, pp. 1-26.

2. W.G. Mumme, J.M.F. Clout and R.W. Gable, “The crystal structure of SFCA-I, Ca3.18Fe3+

14.66Al1.34Fe2+0.82O28, a homologue of the aenigmatite structure type, and new

crystal structure refinements of -CFF, Ca2.99Fe3+14.30Fe2+

0.55O25 and Mg-free SFCA, Ca2.45Fe3+

9.04Al1.74Fe2+0.16Si0.6O20,” Neues Jahrbuch fur Mineralogie, Abhandlungen,

Vol. 173, 1998, pp. 93-117. 3. N.J. Bristow and A.G. Waters, Transactions of the Institution of Mining and Metallurgy

(Section C: Mineral Processing and Extractive Metallurgy), Vol. 100, 1991, C1-C10. 4. I. Shigaki, M. Sawada and N. Gennai, “Increase in low-temperature reduction

degradation of iron ore sinter due to solid solution and columnar calcium ferrite,” Trans. ISIJ, Vol. 26, 1986, pp. 503-511.

5. C.E. Loo, K.T. Wan and V.R. Howes, “Mechanical properties of natural and synthetic mineral phases in sinters having varying reduction degradation indices,” Ironmaking Steelmaking, Vol. 15, 1988, pp. 279-285.

6. M.I. Pownceby, J.M.F. Clout, and M.J. Fisher-White, “Phase equilibria for the Fe2O3-rich part of the system Fe2O3-CaO-SiO2 in air at 1240°-1300°C,” Transactions of the Institution of Mining and Metallurgy (Section C: Mineral Processing and Extractive Metallurgy), Vol. 107, 1998, pp. C1-C10.

7. M.I. Pownceby, and T.R.C. Patrick, “Stability of SFC: solid solution limits, thermal stability and selected phase relationships within the Fe2O3-CaO-SiO2 (FCS) system,” European Journal of Mineralogy, Vol. 12, 2000, pp. 455-468.

8. M.I. Pownceby, and J.M.F. Clout, “Phase relations in the Fe-rich part of the system Fe2O3(-Fe3O4)-CaO-SiO2 at 1240-1300oC and pO2=5x10-3 atm: Implications for iron ore sinter,” Transactions of the Institution of Mining and Metallurgy (Section C: Mineral Processing and Extractive Metallurgy), Vol. 109, 2000, pp. C36-C48.

9. T.R.C. Patrick, and M.I. Pownceby, “Stability of SFCA (silico-ferrite of calcium and aluminium) in air: solid solution limits between 1240oC and 1390oC and phase relationships within the Fe2O3-CaO-Al2O3-SiO2 (FCAS) system,” Metallurgical Transactions B, Vol. 33B, 2002, pp. 79-89.

10. N.A.S. Webster, M.I. Pownceby, I.C. Madsen and J.A. Kimpton, “Silico-Ferrite of Calcium and Aluminum (SFCA) iron ore sinter bonding phases: new insights into their formation during heating and cooling,” Metallurgical and Materials Transactions B, Vol. 43, 2012, pp. 1344-1357.


PRESENTATION - 14

The Crucible Process

Dr Joe Herbertson The Crucible Group Pty. Ltd

Keywords: Bioenergy, Pyrolysis

The Crucible Process is a proprietary technology for decomposition of abundant and undervalued biomass resources, such as agricultural and forest/timber industry residues and urban industrial and municipal waste streams into energy and water products for industrial applications.

The technology is a fundamental innovation, which combines the functions of dewatering, char making (pyrolysis), tar cracking and gas scrubbing within one reactor system.

The unique thermo-chemical profile leads fundamentally to key competitive advantages:

• Licensing of proprietary technology • Process efficiency • Feedstock flexibility • Carbon abatement • No need for pre-drying of feed materials • Clean gas production without gas treatment plant • Enhanced capacity to process contaminated feedstocks • Low capital costs

The process outputs all have economic value, and there are no process wastes.

The gas fraction (containing hydrogen, CO and methane) has energy and chemical value (eg. for heat, electricity generation, and synthesis of liquid fuels)

The char fraction (carbon rich) has energy and chemical value in industry (eg in metallurgical processes and power generation), or as a soil conditioner (improved fertility; carbon capture and storage)

The water fraction has value in horticulture and beneficial use in industrial operations (eg. waste water treatment).

The technology has been under development for some five years and has now been taken to the commercial stage. The science and business foundations for proliferation across multiple sectors are being established to create carbon neutral heat and electricity to manufacturing plants and co-firing at existing power stations. The Crucible Process will contribute to the transition to a clean technology economy by making new fuels, feedstocks and co-products available to industry on a commercially attractive basis.


PRESENTATION - 15

The Effect of Slag Basicity on Spinel Inclusion Wettability

Hamed Abdeyazdan1, Brian J. Monaghan1, Neslihan Dogan1 and M. Akbar Rhamdhani2

1 University of Wollongong, Wollongong, NSW 2522, Australia

2 Swinburne University of Technology, Hawthorn, VIC 3122 Australia

Keywords: spinel inclusion, wettability, inclusion removal Steel cleanness is an important and growing research area driven by the demands to produce high quality steel. Inclusion content in steel is an important criterion to assess clean steel. MgO.Al2O3 spinel inclusions cause problems in steel processing and are generally deleterious to steel products due to their high melting point and high hardness. Inclusions are generally removed by reacting with slag. This is primarily achieved by optimizing the process conditions to promote contact and reaction between the inclusion and slag 1. For efficient removal from the steel, the inclusions must attach to and dissolve in the slag phase. If this attachment is weak, then local fluid conditions are likely to result in the shearing of this attachment and the inclusions re-entrapment in the steel. The strength of attachment (reactivity) between the inclusion and the slag phase can be characterized by the wettability of the slag on inclusions 2-3. Research on inclusion removal in steel refining is principally divided into categories of flotation of inclusion to the steel/slag interface 4-5, modification to improve reactivity/separation with the slag phase 6 and dissolution in the slag phase. 2, 7-12 Much research 2, 7-12 has been carried out on the dissolution of inclusions in slag. Monaghan and Chen 11 and Valdez et al. 12 used a laser scanning confocal microscope (LSCM) to investigate the effect of slag basicity on spinel dissolution they both found that the rate of dissolution of the spinel particles increases with increasing basicity of the slag. Wettability of slag on a substrate of an oxide representing the inclusion phase can be considered an indirect measurement of the inclusions reactivity with slag. However, there are only limited data on slag on typical inclusion phases. Recently Choi and Lee 3 investigated the wettability of alumina on slag and concluded that for a slag with a given CaO/SiO2 ratio, an increase in Al2O3 results in a decrease in wettability. This may in part be explained by the change in thermodynamic driving force of the reaction and/or a change in the physical characteristics of the slag with increasing alumina. To fully understand the wetting behavior of the inclusion phase with liquid slag a similar technique as used by Choi and Lee 3 has been developed to study alumina, spinel, calcium aluminates and iron sulfides in CaO-SiO2-Al2O3 slags. The dynamic contact angle for these systems will be measured using a modified sessile drop apparatus, though only measurements using on a spinel type inclusion are reported here. The experiment involves allowing a liquid drop to spread over a substrate of solid material in an inert atmosphere and the measurement of the change in contact angle ( ) with time 13. Key details of the experiment are given below.


Slags used were based on ladle type slags. They have a CaO/Al2O3 ratio of 1.02, 1.24 and 1.52 and contain 9% SiO2 and 6% MgO. The phase stability of the spinel-slag was calculated using MTDATA thermodynamic software and is given in Fig 1.

Substrates were prepared using laboratory grade reagents. A high density substrate was achieved by using -38μm particle size of pre-fused spinel which was pressed and sintered at 1600 C for 24 hours. Phases were confirmed by XRD analysis.

In the sessile drop experiments the slag and substrate where not in contact when heated to 1500 C the experimental temperature in an argon atmosphere. Once stabilized at temperature the slag was contacted to the substrate and its change in wetting on the substrate measured. A schematic of the experiment apparatus used is given in Fig 2. The experiment was carried out based on heating the slag and substrate to temperature but not in contact. The change in wetting was captured by high definition video and still frames were analysed using the ImageJ 14 software package.

alumina support tube slag drop substrate alumina block

Pt wire alumina twin bore alumina tray Ar inlet Ar outlet monitor camera, lens and filter

Fig 1: phase stability diagram of spinel-slag Fig 2: A schematic of the sessile drop technique

Calculation of the contact angle was based on the geometry of a spherical cap, given in Fig 3 (a) and calculated via equation (1) where R is the radius dissected by a base plane. Also shown in Fig 3 are wetting and non-wetting conditions that have <90° and >90° respectively.

(1)

Fig 3: (a) geometry of a sessile drop approximated by a spherical cap defined by radius of the base and height above basal plane (b) contact angle for wetting liquid, (c) contact angle for non-wetting liquid,

The liquid-solid interface was examined using SEM and EDS analysis.

Frames from a sessile drop experiment using spinel substrate reacting with a slag of C/A=1.24 is shown in Fig 4 from before time zero to 30 sec after time zero. The time that the slag left the Pt wire completely defined time zero. The preliminary results of contact angle measurements were compared with those by Choi and Lee 4, who looked at a similar type slag (C/A=1.26 containing 13% SiO2) on alumina at 1600 °C. The results and comparison are


given in Figure 5. From this Figure it can be seen that the change in wetting in the two different studies have similar characteristic i.e. they start off high and decay quickly to something like a plateau value. Based on the current study it was found that the slag has lower with spinel than Choi and Lee 3 reported for alumina. It should be noted that the slag composition and temperature are not equivalent in these two studies. Further work will be undertaken to study wettability characteristics of slags on various (inclusion phase) substrates and to investigate the inclusion dissolution kinetics using laser scanning confocal microscopy.

t<0

t<0

t<0

t=0

t=1

t=2

t=3 sec

t=4 sec

t=5 sec

t=6 sec

t=7 sec

t=8 sec

t=9 sec

t=10 sec

t=15 sec

t=20 sec

t=25 sec

t=30 sec

Fig 4: frames of the sessile drop experiment

Fig 5: contact angle measurements of current study compared with Choi and Lee’s investigation

Acknowledgments The authors would like to acknowledge BlueScope Steel for their support of this research.


References 1. B. Deo and R. Boom: Fundamentals of steelmaking metallurgy, 84-96, 1993, New York,

Prentice Hall International 2. B. J. Monaghan, L. Chen and J. Sorbe, “Comparative Study of Oxide Inclusion

Dissolution in CaO-SiO2-Al2O3 Slag”, Ironmaking Steelmaking, 2005, 32, 258-264 3. J. Y. Choi and H. G. Lee, “Wetting of Solid Al2O3 with Molten CaO-Al2O3-SiO2, ISIJ

Int., 43(2003), 1348 4. L. Zhang, S. Taniguchi and K. Matsumoto, “Water Model Study on Inclusion Removal

from Liquid Steel by Bubble Flotation under Turbulent Conditions”, Ironmaking and Steelmaking, 29(2002), 326

5. L. Jonsson and P. Jönsson, “Modeling of Fluid Flow Conditions Around the Slag/Metal Interface in a Gas-Stirred Ladle”, ISIJ Int., 36(1996)

6. Y. Miki, H. Kitaoka, T. Sakuraya and T. Fujii, “Mechanism for Separation Inclusions from Molten Steel Stirred with a Rotating Electro-Magnetic Field”, ISIJ Int., 32(1992)

7. K. H. Sandhage and G. J. Yurek, “Indirect Dissolution of (Al,Cr)2O3 in CaO-MgO-Al2O3-SiO2 (CMAS) Melts”, Journal of the American Ceramic Society, 74(1991), 1941

8. S. Sridhar and A. W. Cramb, “Kinetics of Al2O3 dissolution in CaO-MgO-SiO2 -Al2O3 slags: in situ observations and analysis”, Metallurgical and Materials Transaction B, 31(2000), 406

9. X. Yu, R. J. Pomfret and K. S. Coley, “Dissolution of Alumina in Mold Fluxes”, Metallurgical and Materials Transaction B, 28(1997)

10. M. Valdez, K. Prapakorn, A. W. Cramb and S. Seetharaman, “A study of the dissolution of Al2O3, MgO and MgAl2O4 particles in a CaO-Al2O3-SiO2 slag”, Steel Research, 72(2001), 291

11. B. J. Monaghan and L. Chen, “Effect of changing slag composition on spinel inclusion dissolution”, Ironmaking and Steelmaking, 2006, Vol. 33, No. 4, 323-330

12. M. Valdez, K. Prapakorn, A. W. Cramb and S. Sridhar, “Dissolution of Alumina Particles in CaO-Al2O3-SiO2-MgO Slags”, Ironmaking Steelmaking, 29 (2002), 47

13. N. Eustathopoulos, M. G. Nicholas, and B. Drevet, Wettability at High Temperatures, Elsevier Science Ltd, (1999)

14. ImageJ, Version 1.38. 2008, National Institutes of Health, Bethesda, Maryland, USA, http://rsb.info.nih.gov/ij/, 1997-2008



Current Issues of High-Temperature Thermodynamics

George Kaptay1,2, Csaba Mekler1, Adam X. Vegh1,2

1Bay Zoltan Applied Research Public Nonprofit Ltd, BAY-LOGI, Dep. Nano-materials, 2 Igloi, Miskolc, Hungary 3519

2University of Miskolc, Dept. Nanotechnology, Egyetemvaros, Miskolc, Hungary 3515 [email protected]

Keywords: 4th law of thermodynamics; excess Gibbs energy of solutions; surface tension, interfacial energies; surface phase transition; nano-phase diagrams; Gibbs; Kelvin; Ostwald; Butler, SI-system of units. High-temperature processing technologies are based on many disciplines: chemical thermodynamics, kinetics, transport, interfacial phenomena, etc. The relative importance of thermodynamics compared to kinetics and transport in high-temperature processing is due to the fact that kinetic and transport limitations are much less expressed compared to room-temperature systems. On the other hand, the relative importance of interfacial phenomena in high-temperature processing compared to room-temperature processing is due to much higher values of interfacial energies of high-temperature systems. Therefore, the focus of the present talk will be on high-temperature bulk and interfacial thermodynamics. Thermodynamics is based on two absolute laws and a 3rd law, predicting the behaviour of materials when extrapolated to zero Kelvin of absolute temperature. For high-temperature thermodynamics a logical question arises: can we declare anything general for the case when temperature is extrapolated to infinity? The answer is the 4th law of materials thermodynamics [1], claiming that real solutions at fixed composition and pressure tend towards ideal solutions when temperature approaches infinity, supposing the standard states of all the components and that of the solution are selected identically [1]. Mathematically this claim is formulated as:

0lim 0imm E

TG (1)

where EGG (J/mol) is the excess Gibbs energy of the solution. One of the equations satisfying the 4th law for binary solutions is the extended Redlich-Kister equation [2], describing the temperature and concentration dependence of the excess Gibbs energy of high-temperature solid or liquid solutions at any pressure below 100 bar:

0exp)21()1(

j j

jj

E TxhxxGj

(2)

where x (dimensionless) is the mole fraction of one of the components, T (K) is the absolute temperature, jh (J/mol) is the enthalpy-part of the jth interaction energy, jj is the temperature at which the jth interaction energy would change its sign if described by the oversimplified linear equation, often leading to a high-temperature calculated Calphad-artefact. One of the most recent applications of Eq.(2) is the Calphad-artefact-free description of the ternary Al-Mg-Si system [3]. To control the phenomena in multi-component and multi-phase high-temperature systems the knowledge of all interfacial forces [4-5] and interfacial energies [6] are of high importance.


Although the surface tension of liquid alloys is routinely described by the Butler equation, it has been only recently extended to describe the concentration dependence of the solid/solid, liquid/liquid [7] and solid/liquid interfaces [8]. In all these methods the great insight of Butler is applied, according to which the “integral” interfacial energy of the solution ( ) equals the partial interfacial energies of its components ( ii ):

ii (3) This equation is similarly fundamental to the much better known condition of Gibbs for the equilibrium of heterogeneous substances:

)(i)(i GG )) G (4) where )(iG ) and )(iG ) are partial Gibbs energies of component i in contacting phases and

. Eq.(3) also serves as a basis to calculate surface phase transition (SPT) in multicomponent systems [9-10]. When the SPT line is crossed, the A-rich liquid becomes covered by a nano-meter thin B-rich layer (B is a surface active component, soluble in A). This nano-layer changes all kinetic parameters through the interface of the A-rich liquid, and also changes the sign of the temperature coefficient of surface tension. The SPT line in the Fe-O system can be used to explain the inversed Marangoni flow during A-TIG welding of steel [11]. Nano-Calphad, i.e. the calculation of phase diagrams of systems with at least on of its dimensions below 100 nm is a natural combination of bulk and interfacial thermodynamics [12]. One of the key equations is the extension of the bulk Gibbs energy to the Gibbs energy of phases with high specific surface (interface) area:

ss/s/

oo A

VVGG (5)

where oG (J/mol) is the Gibbs energy of phase if the role of interfaces is negligible, oVV (m3/mol) is the molar volume of phase , V (m3) is the volume of phase , s/A s/ (m2) and s/ s/ (J/m2) are the interfacial area and interfacial energy between phases and „s”, respectively. Let us mention, that the molar Gibbs energy of the phase is proportional to its specific surface (interface) area, in accordance with thermodynamics of Gibbs, and not to the curvature of the phase, as follows from the Kelvin equation. As a consequence of Eq.(5), the Kelvin equation (on size dependence of the vapour pressure above a small droplet) [13], the Gibbs-Thomson equation (on the melting point of a nano-crystal) [13] and the Ostwald-Freundlich equation (on the solubility of nano-crystals) [14] have been corrected. Finally, let us mention a more general topic of natural sciences and engineering. As known, the SI system of units is based on 7 base quantities and the corresponding base units: length (m), time (s), mass (kg), temperature (K), electric current (A), amount of matter (mole) and intensity of light (cd). However, as was recently shown [15], the smallest necessary number of base quantities and units to describe nature is only 5: length (m), time (s), mass (kg), temperature (K), electric charge (C). The amount of matter and the Avogadro number are based on an arbitrary definition (12 g of C-12 – why 12 g?), thus mole is not a base unit. To count atoms and molecules positive integer numbers discovered by ancient mathematicians is sufficient, i.e. mole is a useful, but not a necessary base unit. The intensity of light is a typical derived unit: it is defined as a given power (W) per given 3-D angle (sr) of light of given frequency (Hz). So, intensity of light should be treated as a derived unit and not as base unit. All electromagnetic quantities are suggested to be described though charge as base unit instead of electric current, as the latter is the velocity of charges, but length is not defined through velocity, either. The author believes that education of about 100 million children per


year worldwide will become easier if the current SI system of units is simplified as suggested here [15]. Acknowledgement This work was financed by project the European Union through the TÁMOP-4.2.2.A-11/1/KONV project and by the Hungarian Academy of Sciences through the K101781 OTKA project. References 1. G. Kaptay, "On the tendency of solutions to tend toward ideal solutions at high

temperatures”, Metall Mater Trans A, vol.43, 2012, pp. 531-543. 2. G. Kaptay, “A new equation for temperature dependence of the excess Gibbs energy of

solution phases”, Calphad, vol.28, 2004, pp.115-124. 3. Y. Tang, Y. Du, L. Zhang, X. Yuan and G. Kaptay, “Thermodynamic description of the

Al–Mg–Si system using a new formulation for the excess Gibbs energy”, Thermochimica Acta, vol.527, 2012, pp.131-142.

4. G. Kaptay, „Classification and general derivation of interfacial forces, acting on phases, situated in the bulk, or at the interface of other phases”, J Mater Sci, vol.40, 2005, pp.2125-2131.

5. G. Kaptay, “Interfacial Forces in Dispersion Science and Technology - Journal of Dispersion Science and Technology”, J Dispersion Sci Technol, vol.33, 2012, pp.130-140.

6. G. Kaptay, “Modeling Interfacial Energies in Metallic Systems”, Mater Sci Forum, vols. 473-474, 2005, pp. 1-10.

7. G. Kaptay, “On the interfacial energy of coherent interfaces”, Acta Mater, vol.60, 2012, pp. 6804-6813.

8. Z. Weltsch, A. Lovas, J. Takács, Á. Cziráki, A. Tóth, G. Kaptay, “Measurement and Modelling of the Wettability of Graphite by a Silver-Tin (Ag-Sn) Liquid Alloy”, Applied Surface Science, 2013, doi: 10.1016/j.apsusc.2012.11.150.

9. G. Kaptay, “A method to calculate equilibrium surface phase transition lines in monotectic systems”, Calphad, vol.29, 2005, pp. 56-67 (Erratum, same volume, p.262)

10. C.Mekler, G.Kaptay, “Calculation of surface tension and surface phase transition line in binary Ga-Tl system”, Mater Sci Eng A, vol.495, 2008, pp.65-69.

11. T. Sándor, C. Mekler, J. Dobránszky, G. Kaptay, „An improved theoretical model for A-TIG welding based on surface phase transition and reversed Marangoni flow”, Metall Mater Trans A, 2013, doi: 10.1007/s11661-012-1367-2.

12. G. Kaptay, “Nano-Calphad: extension of the Calphad method to systems with nano-phases and complexions” J Mater Sci, vol.47, 2012, pp. 8320-8335.

13. G. Kaptay, “The Gibbs equation versus the Kelvin and the Gibbs-Thomson equations to describe nucleation and equilibrium of nano-materials”, J Nanoscie Nanotechnol, vol.12, 2012, No.3, pp. 2625-2633.

14. G. Kaptay, “On the size and shape dependence of the solubility of nano-particles in solutions” Int J Pharmaceutics, vol.430, 2012, pp.253-257.

15. G. Kaptay, “On the five base quantities of nature and SI (The International system of Units)”. JMM B, vol.47, 2011, pp.241-246.


PRESENTATION – 17

Life Cycle Based Greenhouse Gas Emission Assessment from Ferroalloy Production

Nawshad Haque, Terry Norgate

CSIRO Minerals Down Under Flagship Bag 312, Clayton South, VIC 3168

Keywords: ferroalloys, LCA, metal production, energy intensity, carbon footprint Ferroalloys are defined as iron-bearing alloys with a high proportion of one or more other elements typically manganese, chromium, silicon, molybdenum and nickel. Ferroalloys are mainly used by the iron and steel industry and ferroalloy production is closely related to steel production. The leading ferroalloy-producing countries in 2008 were, in decreasing order of production, China, South Africa, Russia, Kazakhstan, and Ukraine. These countries accounted for 77% of world ferroalloy production. The major ferroalloys are ferrochromium (FeCr), (ferro)-silicomanganese (FeSiMn or often referred to as SiMn), ferrosilicon (FeSi), ferromanganese (FeMn), ferronickel (FeNi), ferromolybdenum, ferrotitanium, ferrotungsten and ferrovanadium. The increased emphasis on sustainability in recent years has seen the value chains for the production of materials including metals, come under close scrutiny. Life cycle assessment (LCA) methodology has been developed to assist in this task, particularly in regard to assessing environmental impacts of these value chains. Despite the significance of the ferroalloy industry, there have been very few LCAs of ferroalloy production reported in the literature [1]. The study described in this paper uses LCA methodology to estimate the greenhouse gas (GHG) footprint of ferroalloy production, in particular, FeMn, SiMn and FeSi, and to update the GHG footprints of FeCr and FeNi previously estimated [2]. This paper has been prepared assuming ferroalloy production is based in Tasmania with some broader Australian comparisons. Comparisons with other studies have also been presented [3]. LCA methodology was used to estimate the greenhouse gas (GHG) footprint of a number of the selected ferroalloy production processes using SimaPro LCA software. The results of the study showed that the GHG footprints of ferroalloy production in Australia were 1.8 t CO2-e/t FeMn, 2.8 t CO2-e/t FeSiMn and 3.4 t CO2-e/t FeSi alloy metal. GHG footprints were for ferronickel and ferrochromium of 13.9 t CO2-e/t and 3.0 t CO2-e/t, respectively. These GHG footprint estimates were calculated using the hydroelectricity dominated Tasmanian electricity greenhouse gas emission factor. The major difference in greenhouse gas emissions between the various ferroalloys is largely due to their respective amounts of electricity use and coke/coal consumption. The GHG emission footprints would increase if the ferroalloy production plants used electricity generated elsewhere in Australia (Figure 1). In this study the Australian average electricity, which is more fossil fuel (ie. coal) based, was 1 kg CO2-e/kWh while the Tasmanian electricity which is more hydroelectricity-based was 0.1 kg CO2-e/kWh. These results are similar to the results from a limited number of studies reported in the literature when compared on a similar electricity source basis.


Figure 1: Carbon footprints of various ferroalloy metals using electricity from two grids (Haque and Norgate, 2013). The LCA results also showed that coke and coal usage contributed close to 60% or more of the total GHG emissions from the various ferroalloy production processes. In light of this finding, there is an opportunity to reduce GHG emissions from ferroalloy production by direct replacement of fossil fuel-based coal with biomass-based renewable carbon. A first estimate of the potential reduction in greenhouse gas emissions from the use of biochar in ferroalloy production ranged from 0.7 t CO2e/t FeMn to 5.2 t CO2e/t FeNi for a 50% replacement of coal and/or coke in the respective processes. As Tasmania is rich in waste biomass such as forestry waste, it is likely that biochar application to ferroalloy making in Tasmania, if possible, would yield ferroalloy products with some of the lowest GHG intensities in the world. These results suggest environmental benefit from substituting biochar for coal/coke and a technical assessment to confirm the suitability of biochar for such an application is recommended. References 1. T. Lindstad, “CO2-emissions and the ferroalloys industry”. Proceedings of the 8th

International Ferroalloys Congress (INFACON8), 1998, Peking, China, pp. 87-92. 2. T. Norgate, S. Jahanshahi, and W. J. Rankin, “Alternative routes to stainless steel – a life

cycle approach”. Proceedings of Tenth International Ferroalloys Congress, February 2004, Cape Town, South Africa, pp. 693-704.

3. N. Haque, and T. Norgate, "Estimation of greenhouse gas emission from ferroalloy production using life cycle assessment with particular reference to Australia”. Journal of Cleaner Production, Vol.39, 2013, pp. 220:230.

Corresponding author’s email: [email protected] Tel: 03-9545 8931. Web: www.csiro.au/people/Nawshad.Haque.html

1.82.8 3.4 3.0

13.9

3.6

6.1

10.2

7.2

24.1

0

5

10

15

20

25

30

FeMn SiMn FeSi 75 FeCr FeNi

t CO

2-e/

t allo

y m

etal

Tasmania Australia


PRESENTATION - 18

Special Alloy Strip Production in a Micro-Mill Environment

Yvonne Durandet1, Kannappar Mukunthan2,3 and Joe Herbertson2 1 Swinburne University of Technology, Hawthorn, VIC 3122 Australia

2 The Crucible Group Pty Ltd, Newcastle, NSW, Australia 3 Deakin University, Geelong, VIC 3220, Australia

Keywords: net shape manufacturing, solidification simulation

In the current context of relentless global competition and shortages of energy and other resources, new business models and advanced manufacturing technologies are required to realise more value from less. Net and near net shape conversion processes, such as additive manufacturing and direct strip casting, may be part of the solution when designed for scaled economy and efficiency. A study conducted with the support of CAST CRC indicated that The Crucible Group’s micro mill concept for compact alloy production of low volume, high value added flat products (referred to as “MicroCAP”) was feasible for iron-based specialty alloys. Specialty Fe-based alloys that were investigated include a ferritic grade FeCrAl alloy [1] and an austenitic grade FeNi alloy [2]. FeCrAl alloys such as Kanthal are used in heating elements because of their relatively high electrical resistance and excellent corrosion resistance at high service temperatures of 1100-1300ºC. FeNi alloys such as Invar are commonly used in glass-to-metal seals in electron tubes, transistors, headlights, thermostats, and other similar applications because of their extremely small thermal expansion. Current processing of these alloys is complex, with many conventional steps consisting of ingot or slab casting, solution treating, hot forging and/or hot rolling, cold rolling, and possibly surface grinding before final cold rolling to remove surface cracks and oxide layers. Simulating the casting of these alloys directly into thin strip showed the potential for significant reductions in yield losses from liquid metal to final products and hence substantial improvements in productivity. Both FeCrAl [1] and FeNi [2] alloys were found to be castable with as cast properties suitable for direct finishing operations. The cast intermediate material responded well to conventional rolling and annealing conditions, giving comparable properties to commercially available products. The laboratory simulation was based on an experimental approach which approximates the solidification conditions encountered during rapid solidification, such as strip casting [3]. It involves the rapid immersion under controlled conditions of thermocouple instrumented copper substrates into a bath of molten alloy (Fig. 1) to determine the interfacial heat transfer with millisecond resolution (Fig 2). Parametric studies of the influence of melt temperature and chemistry, gas atmosphere, immersion speed and substrate surface conditions on heat transfer can thus be integrated with observations of the behaviour of inclusions formed at the interface during initial contact, cast defect formation and solidification microstructures, and analyses of microstructure/property evolution during subsequent thermal-mechanical processing [4, 5]. The experimental parameters can be adjusted quickly and many immersion solidification experiments can be completed in a single day.


Figure 1: Schematics of immersion apparatus (left) and instrumented mould substrate (right) for solidification simulation experiments [6].

Figure 2: Example of heat flux histories influenced by increasing level of oxide build-up from six consecutive immersion sequences (left) and SEM of initial surface oxides after the first immersion [7]. The MicroCAP concept is underpinned by (1) streamlined processing from liquid metal to final products, (2) avoiding complex and costly ingot processing stage, and (3) rapid solidification to enable direct processing of cast intermediates into final products. Thus, product development steps tend to be combined in concurrent process engineering. In this context, the flexible experimental approach to initial solidification, based on immersion simulation, can accelerate the research of special alloys and exploit the benefits of rapid solidification and compact alloy processing. Understanding interfacial heat transfer is of vital importance to understanding the rapid solidification step at the heart of the MicroCAP concept. For highly alloyed steels and other specialty ferrous alloys, the solidification simulation approach showed a myriad of ways of altering the interfacial heat transfer, and the subsequent evolution of microstructures and final properties [5]. For non-ferrous metals such as aluminium alloys, horizontal strip cast manufacturing is more common, but the full range of possibilities for controlling heat transfer during initial solidification has not been investigated. Certainly for magnesium very little interfacial heat


transfer research has been carried out. Experimental research using the immersion solidification technique in a low risk, low cost environment can unlock new opportunities in near net shape casting of non-ferrous alloys, to assess the range of conditions needed to understand the major factors controlling interfacial heat transfer and subsequent evolution of microstructures and final product properties, without being constrained or limited by a specific casting technology or engineering design [8, 9]. The presentation will summarise the results to date with special Fe-based alloys and discuss how the experimental methods specifically and the MicroCAP concept generally can be applied to the development of improved non-ferrous alloy production or products. References 1. Mukunthan, K., et al., Castability and microstructural development of iron-based alloys

under conditions pertinent to strip casting. Part 1 – Specialty Fe-Cr-Al alloys. ISIJ International, Submitted 15 January 2013.

2. Mukunthan, K., et al., Castability and microstructural development of iron-based alloys under conditions pertinent to strip casting. Part 3 – Specialty Fe-Ni alloys. ISIJ International, Submitted 15 January 2013.

3. Strezov, Les and Joe Herbertson, Experimental studies of interfacial heat transfer and initial solidification pertinent to strip casting. ISIJ International, 1998. 38(9): p. 959-966.

4. Mukunthan, K., et al., Evolution of microstructures during the solidification of 304 austenitic stainless steel. Metallurgia Italiana, 2002. 94(7-8): p. 39-46.

5. Mukunthan, K., et al., Evolution of microstructures and product opportunities in low carbon steel strip casting. Canadian Metallurgical Quarterly, 2001. 40(4): p. 523-532.

6. Strezov, L., Interfacial heat transfer mechanisms associated with the direct contacting of copper substrates by liquid stainless steel, in PhD Thesis (Mechanical Engineering)1994, The University of Newcastle: Newcastle, Australia. p. 248.

7. Strezov, Les, Joe Herbertson, and Geoffrey R. Belton, Mechanisms of initial melt/substrate heat transfer pertinent to strip casting. Metallurgical and Materials Transactions B: Process Metallurgy and Materials Processing Science, 2000. 31(5): p. 1023-1030.

8. Abbott, Trevor and Yvonne Durandet, Strip casting of light alloys. CAST Technical File Note 248/02, 2002.

9. Ferry, Michael, Direct strip casting of metals and alloys : processing, microstructure and properties Woodhead Pub. and Maney Pub., 2006.


PRESENTATION - 19

Carbothermic Reaction of High-Combined-Water-Iron Ore and Coal Composite Pellets

Sri Harjanto1, Adji Kawigraha1,2, Johny W. Soedarsono1, Pramusanto3

1Department of Metallurgy and Materials Engineering, Universitas Indonesia 2Center of Mineral Ressources Technology, Agency for Assessment and Application of

Technology (BPPT), Indonesia 3 R & D Centre for Mineral and Coal Technology, Ministry of Energy and Mineral

Resources, Indonesia

Keywords: lateritic iron ore, geothite, low-grade coal, reduction, pellet, ironmaking The availability of high grade iron ore and carbon materials for ironmaking industry decreases gradually. On the contrary, low-grade iron ore and coals are extensively utilized not only as suplement materials but also as alternative raw materials. In Indonesia context, the pull of government policy to increase added value of local natural resources makes a bigger opportunity to process low grade iron ore and carbon materials for ironmaking industry, such as lateritic iron ore and coal. The iron ores usually compose of a major component minerals with high combined water, namely geothite or FeOOH and with high amount of gangue. Generally, it is still required beneficiation process to obtain concentrated feed for ironmaking. The possibility of low grade iron ore utilization was investigated by using nugget iron ore [1]. The result showed that the experiment which simulated direct reduction process gave iron recovery about 96-97%. Recent works on low grade iron ore which contain high combined water and coal composite reports gasification reaction during the process which may decrease the reduction reaction temperature [2]. The result implies the advantage of utilization low grade iron ore with high combined water. Decomposition of combined water occurs at temperature about 300oC during calcination according to the reaction below:

2 FeOOH(s) Fe2O3(s) + H2O(g) (1) At higher temperature more than 450oC, generation of CO and CO2, probable gasification and other reactions may take place as follows [2].

C(s) + CO2(g) = 2CO(g) (2) C(s) + H2O(g) = CO(g) + H2(g) (3) CH4(g) + H2O(g) = CO(g) + 3H2(g) (4) C(s) + 3Fe2O3(s) = CO(g) + 2Fe3O4(s) (5)

Considering the above condition and recent results of low grade iron ore and coal composite researches, this study investigates carbothermic reaction behavior of Indonesia lateritic iron ore and coal composite pellets. It emphasizes to understand the reduction behavior of low grade feed in the form of composite pellet for direct reduction ironmaking process, in terms of their reduction degree, metallization, density and microstructures. Indonesia lateritic iron ore which contains geothite, FeOOH was employed as sample materials in the experiment. The ore was classified as two different samples, i.e. original ore (A) and concentrate (C). (A) samples were ground ore samples (less than 105 m) without any other treatment; (C) samples were ground ore samples (less than 105 m) with additional beneficiation process such as gravity concentration and magnetic separation. The ores samples were mixed with coal based on the Fe:C molar ratio of 1:3. Binder was added to each


mixtures at 1 %wt. The mixture was pelletized to diameter of 1.2 – 1.5 cm. They were then heated treated in tube furnace to the desired temperature at the range of 1100 – 1350oC with heating rate of 10oC/minutes in air atmosphere. Heat treated samples were characterized their substances by using x ray diffractometer. Reduction degree of was determined as a difference ratio of oxygen content in product and initial samples. Metallization of the samples were measured by the ratio of metallic iron in reduced pellets and total iron in the initial samples. Density of the samples were also measured based on Archimedes Law. Scanning electron microscope was utilized to characterized the microstructures of the samples. Fig. 1 shows the different weight loss characteristic of ore (A), concentrate (C) samples and coal. Ore beneficiation process reduced the combined water substance content in (C) samples to about 4% from 8% in those of (A) samples. Thermal decomposition of crystalline water of the ores ((A) and (C)) occurs at temperature range of 267 – 332oC. Decomposition of moisture in coal sample was about 8%, but more accurate analysis of coal gave higher moisture content of 14%. The coal contains 42% fixed carbon, 38% volatile matter, 6% ash and 5610 cal/g calorific value.

Fig. 1 Weight loss of ore, concentrate and coal. Fig. 2 Metallization of ore (A) and concentrate (C). Fig. 2 indicates the metallization of the fresh ore (A) and concentrate (C) samples in the reduction temperature range of 1200 – 1350oC. Removal of partial high combined water substance by concentration and magnetic separation may increase the metallization of the samples. It indicates that in this experiment, reduction degree and metallization of concentrate, which less geothite content, are still higher in the same temperature at the range of 1200-1350oC compared with those of original ores which contain high combined water substance. References

1. B. Anameric, S.K. Kawatra, ”Properties and Features of Direct Reduced Iron,” Mineral Processing and Extractive Metallurgy Review., 28, 2007, pp.59-116.

2. T. Murakami, T. Nishimura and E. Kasai,”Lowering Reduction Temperature of Iron Ore and Carbon Composite by Using Ores with High Combined Water Content, “," ISIJ International, Vol. 49, No. 11, 2009, pp. 1686-1693.


PRESENTATION - 20

Development of Dynamic Process Models for Oxygen Steelmaking

Geoffrey Brooks

FEIS, Swinburne University of Technology, Hawthorn, Victoria 3122, Australia

Oxygen steelmaking is a mature technology with over fifty years of continuous process development. Over the last twenty years, there have been significant advances in lance technology (e.g. coherent jets), refractories (e.g. C-MgO linings) and process control (e.g. use of cameras to monitor slopping). However, two inter-related challenges have only been partially addressed, namely:

a) how to minimise flux additions whilst still providing adequate protection to the refractory lining and meeting the required grade of steel (particularly, in regard to phosphorus).

b) how to lower blowing times without excessive slopping whilst still meeting the required grade of steel and ensuring process stability.

These problems are not easily addressed through the traditional static models because by they reflect interrelated issues of a dynamic nature, for example, lowering the lance close to the bath may accelerate decarburisation, lowering the potential blowing time, but this may result in slopping under certain circumstances. Similarly, changing the sequence of flux additions to a furnace may reduce the risk of slopping but maybe also expose the refractory to attack early during the blow. “Black box” modelling techniques (i.e. neural networks, fuzzy logic and multi-variate statistics) have been used to address the issues such as these and there is some evidence that these techniques can help in optimising the operation.1,2 & 3 These types of approaches tend to be quite specific to the operation being modelled and also rarely provide insight into the underlying physics and chemistry of the process.

Several groups around the world have recently attempted to address this shortfall in understanding of Oxygen Steelmaking through the development of “Grey box” process models i.e. models based on underlying scientific principles but with semi-empirical corrections. At Swinburne University of Technology, we have previously developed a two zone model of Oxygen Steelmaking that is focused on understanding how decarburisation is affected by lance height, blowing rate, flux dissolution and scrap melting. 4,5 & 6 A schematic diagram outlining the basic components of the model are provided in Figure 1. The model has been successfully validated against the industrial results of Cicutti et al.7 but is limited in its wider application because it doesn’t include slag generation, slopping behaviour and de-phosphorisation. More recent work at Swinburne has made significant progress in developing models for predicting slag foaming behaviour and droplet generation, with the ultimate goal that these models are incorporated into a general process model of Oxygen Steelmaking.

Guo et al. from Arrcelor Mitall have recently published a paper in which they outline some of the features of a model that calculates slag chemistry with time, as a function of lance height and flux additions.8 This model has been used to form strategies to limit slopping and minimise flux additions. Guo et al. claims to limit lime addition through accurate determination of lime saturation and phosphorus content from equilibrium relationships. They also provide evidence that slopping incidents have been greatly reduced since applying these strategies and linking their models to practice..8


Figure 1: Schematic overview of the model developed by Dogan, Brooks and Rhamdhani for describing the kinetics of decarburisation in top blown Oxygen steelmaking.4,5.6

Recent work at JFE Steel has focused on optimising phosphorus removal through dynamic modelling of the slag chemistry.9 In particular, they focused on predicting the FeO content (critical to understanding the phosphorus behaviour) of the slag by means of a dynamic Oxygen balance and relatively simple semi-empirical kinetic expressions to predict chemistry changes. It is not clear how they incorporate lance position into their model but they provide some evidence that their FeO estimations correlate reasonably well to plant data. This work and other related projects around the world point toward the eventual development of full dynamic kinetic models of Oxygen steelmaking that should allow optimisation of the process through a fuller understanding of the dynamics of the process.

References 1. J. Falkus and P. Pietrzkiewicz. Proc. Conf. High Technologies in Advanced Metal

Science and Engineering. 2001. St. Petersburg, Russia 2. C. Kubat, H. Taskin, R. Artir, and A. Yilmaz, Robotics and Autonomous Systems 49 (3-

4), 193-205 Vol.49, No. 3-4, 2004, 93-205 3. C. Stroomer-Kattenbelt, Ph.D Thesis, Twente University, The Netherlands, 2008. 4. N. Dogan, G.A. Brooks, and M.A. Rhamdhani., ISIJ, Vol 51, No. 7, 1086-1092 5. N. Dogan, G.A. Brooks, and M.A. RhamdhaniISIJ, Vol 51, No.7, 1093-1101 6. N. Dogan, G.A. Brooks, and M.A. Rhamdhani.. ISIJ, Vol 51, No.7, 1102-1109 7. C. Cicutti, M. Valdez, T. Perez, R. Donayo, and J. Petroni, Latin American Applied

Research, Vol.32, No.3, 2002, 237-240. 8. D. Guo, D. duBios, D. Swickard, M. Alavanja, R. Kostyo and J. Bradley, Iron and Steel

Technology, Vol. 9, No. 4, 45-51. 9. Y. Ogasawara, N. Kikuchi, A. Matsui and Y. Kishimoto, AIST Transactions – Iron and

Steel Technology, Vol. 9, No. 1, 220-226.


PRESENTATION - 21

Coupled Experimental and Thermodynamic Modelling for Metallurgical Slags and Inclusions Containing Sulphur

Youn-Bae Kang

Graduate Institute of Ferrous Technology, Pohang University of Science and Technology, Rep. of Korea

email: [email protected] Keywords: Oxysulphide, Slag, Sulphide capacity, Modified Quasichemical Model, Thermodynamic modelling Dissolution behaviour of sulphur in molten oxides is of importance in practical point of view. One of key processing steps for a clean steel production is desulphurisation of molten steel, which is controlled by sulphide capacity of flux used for the desulphurisation.1) Control of inclusion evolution in high-S free-cutting steel requires accurate knowledge on sulphide solubility limit in the liquid inclusion of oxysulphide.2) Recently proposed low-P FeMn alloy production via liquid sulphide – liquid oxide reaction also requires phase equilibria and their thermodynamics of the liquid sulphide/oxide phase.3) Those metallurgical processes may be well designed with the aid of computational thermodynamic calculations along with a predictive thermodynamic model and an accurate thermodynamic database for oxysulphide melts.

Figure 1: Calculated sulfide capacity in MnO-SiO2 and iso-log(CS) lines in CaO-SiO2-Al2O3 melts compared with experimental data. In order to have a tool for the accurate prediction of sulphur dissolution in molten oxide melts (slag and inclusion of oxysulphide), a thermodynamic model has been developed in the framework of the Modified Quasichemical Model in the quadruplet approximation.4) This permits the calculation of solubilities of sulphide in molten oxide slags and vice versa. The model calculates the solubilities from a knowledge of the thermodynamic activities of the component oxides and the Gibbs energies of the pure components (oxides and sulphides). In particular, solubilities of sulphur as sulphide in the Al2O3-CaO-FeO-Fe2O3-MgO-MnO-Na2O-SiO2-TiO2-Ti2O3 multi-component slag, predicted from the present model with very few adjustable model parameters, were found to be in good agreement with all available experimental data. The model applies at all compositions from pure oxides to pure sulphides,


and from basic slags to acidic slags. A few of example calculations for the sulphide capacities in slags are shown in Figure 1. In order to confirm validity of the model prediction or to refine the model calculations, several experimental investigations were carried out in the author’s laboratory. Phase equilibria of several oxysulphide systems were measured employing high temperature chemical equilibration-quenching-EPMA technique. Figure 2 shows a micrograph of a quenched sample composed of SiO2, CaS, and glass oxysulphide composed of CaO-SiO2-CaS(-SiS2), and a comparisons between the experimental phase diagram data and the model calculation.5) Shown in Figure 3 is a comparison between the experimental phase diagram data and the model calculation for the MnO-SiO2-Al2O3-MnS(-SiS2-Al2S3) system.6)

Figure 2: Microstructures of CaS and SiO2 crystals, all in equilibrium with liquid oxysulfides (left) . Isothermal section of the CaO–SiO2–CaS system at 1550°C (right). Symbols are experimental data obtained in the present study as well as those from the literature. Lines are calculated from the thermodynamic model employed in the present study.

Figure 3: MnS saturation boundaries in MnO-SiO2-Al2O3-MnS liquid oxysulfide at 1200°C with various wt pct Al2O3. Lines are calculated from the thermodynamic model.


The developed thermodynamic database has been integrated into FactSage thermochemical computing system.7) By coupling the developed database with other evaluated databases for steel, oxides and gaseous phases, practically important slag/steel/inclusion/gas equilibria can be computed such as the S-distribution ratio in molten steel, sulphide inclusion precipitates during solidification and heat treatment of steel, etc. Acknowledgement This study was partially supported by a grant from the Fundamental R&D Program for Core Technology of Materials funded by the Ministry of Commerce, Industry and Energy, Republic of Korea, and POSCO Ltd. through the Steel Innovation Program to Graduate Institute of Ferrous Technology, Pohang University of Science and Technology, Republic of Korea. The author would like to thank Prof. Arthur D. Pelton at Ecole Polytechnique de Montreal for the model development, Prof. Hae-Geon Lee at Pohang University of Science and Technology and Prof. Joo-Hyun Park at the University of Ulsan for constructive discussions, and Dr. Dae-Hee Woo, Ms. Ye-Jin Kim, Mr. Young-Gyu Jo, and Mr. Rongxun Piao at Pohang University of Science and Technology for their contributions in the experiments. References 1. J. Chipman, “Chemical Behavior of Sulphur in Iron and Steelmaking”, Metal Progress,

Vol. 62, No.6, 1952, pp. 97-107. 2. D.-H. Woo, Y.-B. Kang, H. Gaye, and H.-G. Lee, “Experimental Investigations of Phase

Equilibria of MnS Containing Sub-systems in the MnO-SiO2-Al2O3-MnS System”, ISIJ International, Vol.49, No.10, 2009, pp. 1490-1497; D.-H. Woo and H.-G. Lee, “Phase Equilibria of the MnO–SiO2–Al2O3–MnS System”, J. Am. Ceram. Soc., Vol. 93, No. 7, 2010, pp. 2098-2106.

3. S.-J. Kim, H. Shibata, N. Maruoka, S. Kitamura, and K. Tamaguchi, “Novel Recycling Process of Mn by Sulfurization of Molten Slag from a By-Product of Steelmaking Process”, High Temp. Mater. Proc., Vol. 30, No. 4, 2011, pp. 425-434; S.-J. Kim, H. Shibata, N. Maruoka, S. Kitamura, K. Tamaguchi, Y.-B. Kang, “Influence of Partial Pressure of Sulfur and Oxygen on Distribution of Fe and Mn between Liquid Fe-Mn Oxysulfide and Molten Slag”, Metall. Mater. Trans. B., Vol. 43B, No. 5, 2012, pp. 1069-1077.

4. Y.-B. Kang and A.D. Pelton, “Thermodynamic Model and Database for Sulfides Dissolved in Molten Oxide Slags”, Metall. Mater. Trans. B., Vol. 40B, No. 6, 2009, pp. 979-994.

5. R. Piao, H.-G. Lee, and Y.-B. Kang, “Experimental investigation of phase equilibria and thermodynamic modeling of the CaO-Al2O3-CaS and the CaO-SiO2-CaS oxysulfide systems”, Acta Mater., Vol. 61, No. 2, 2013, pp. 683-696.

6. Y.J. Kim, D.-H. Woo, H. Gaye, H.-G.Lee, and Y.-B. Kang, “Thermodynamics of MnO-SiO2-Al2O3-MnS Liquid Oxysulfide: Experimental and Thermodynamic Modeling”, Metall. Mater. Trans. B., Vol. 42B, No. 3, 2011, pp. 535-545.

7. C. W. Bale, E. Be´ lisle, P. Chartrand, S. A. Decterov, G. Eriksson, K. Hack, I.-H. Jung, Y.-B. Kang, J. Melanc¸on, A. D. Pelton, C. Robelin, and S. Petersen, “FactSage Thermochemical Software and Databases. Recent Developments,” Calphad, Vol. 33, No. 2, 2009, pp. 295-311.


PRESENTATION - 22

Analysis of Waves in a Cavity and their Significance to Splashing in Steelmaking

Shabnam Sabah, Geoffrey Brooks, Jamal Naser

Swinburne University of Technology, Hawthorn, Victoria 3122, Australia Email: [email protected]

In steelmaking, metal droplets are generated due to the impact of supersonic jets onto the liquid metal bath. This phenomenon is generally referred to “splashing” [1]. Splashing is an important phenomenon as it provides large interfacial area between reacting phases and promotes high overall reaction rates. Dogan et al. [2] estimated that 60% of decarburization in steelmaking process takes place in the gas-metal-slag emulsion phase during the main blow. Therefore, it is important to understand droplet generation if we are to optimize steelmaking processes. In literature only a few works [3,4] discuss the formation of waves in the cavity. But there is lack of understanding how waves inside the cavity affect splashing. In the present work, a 1/10th cold model of oxygen steelmaking was used to study wave phenomenon in the cavity and how its characteristics vary with the changing momentum of impact jet and thereby affecting splashing. Compressed air was passed through a top straight nozzle. Water was used to simulate liquid steel. As the first phase of the investigation, the slag phase was not included in this study. A transparent cylindrical rig, made of perspex sheet was used as the vessel. A high speed camera (MotionPro Y, model Y4L (1024x1024 – 4000 fps)) was used to take photos of the cavity and the splashing at a rate of 30 frames/ second for a time duration of 4 minutes. In addition, a video of 2 second was taken at 2000 frames/ second rate in order to observe the droplet generation process.

(a) (b)

Figure 1: (a) Peregrine sheet [5]; (b) sheet structure at flow rate 50 L/min, lance height 150 mm

As air jet hit the water surface, it created waves inside the cavity. These waves were pushed towards the edge of the cavity by deflected gas flow. At the edge of the cavity, these waves grew to certain critical amplitudes. These waves were like sheets structure at the edge of cavity. They were similar to “peregrine sheet” described in the general splashing literature [5]. Figure 1(a) and (b) shows an irregular peregrine sheet from an earlier study and a typical sheet structure observed in the cold modeling experiment respectively.


From high speed video, droplet formation process was observed and three stages could be identified- amplitude of the waves (which will be called “sheets” from now on) enlarged at the edge of the cavity, instability grew formed at the rim of the sheets and fingers were formed and finally, fingers broke up into one or number of droplets. Depending upon lance height and air flow rate, any one or two or all three stages were found. For example, at gas flow rate 40 L/min and lance height 150 mm, only sheet structures were observed. As gas flow rate increased from 40 L/min to 50 L/min, fingers as well as droplet generation began. It was observed that frequency and amplitude of the waves in the cavity, the height of the sheets increased as gas momentum on bath surface increased which ultimately, increased droplet generation. Also, the shape of the droplets changed from circular to irregular shapes as momentum of impact jet grew. It was found that the characteristics of the waves have an effect on the splashing droplet amount and shape. A variety of mathematical techniques (including Fast Fourier Analysis) were used to analyze the wave behavior and the results from this analysis will be discussed at the symposium. References 1. Tago Y. and Y. Higuchi, Fluid flow analysis of jets from nozzles in top blown process.

ISIJ International, 2003. 43(2): p. 209-215. 2. Dogan, N., G.A. Brooks, and M.A. Rhamdhani, Comprehensive Model of Oxygen

Steelmaking Part 3: Decarburization in Impact Zone. ISIJ International, 2011. 51(7): p. 1102-1109

3. Peaslee, K.D. and D.G.C. Robertson. Model studies of splash, waves, and recirculating flows within steelmaking furnaces. in Steelmaking Conference Proceedings. 1994. Warrendale, PA: Iron & Steel Society 77: p. 713-722.

4. Lee, M., V. Whitney, and N. Molloy, Jet-liquid interaction in a steelmaking electric arc furnace. Scandinavian Journal of Metallurgy, 2008. 30(5): p. 330–336.

5. Deegan, R., P. Brunet, and J. Eggers, Complexities of splashing. Nonlinearity, 2008. 21: p. C1.


PRESENTATION - 23

Injection of Fine Charcoal into a Slag Bath

Michael Somerville1, Justen Bremmell1, Jason Donnelly1, David Langberg2 1CSIRO Process Science and Engineering, Box 312, Clayton South, VIC, 3169, Australia

2Formerly CSIRO Process Science and Engineering

Keywords: charcoal, slag bath, injection, reduction, carbon utilisation, iron, Sirosmelt The use of charcoal as a metallurgical reductant has a number of advantages compared to coal and coke including: low ash and sulphur levels, higher reactivity and lower net CO2 emissions when substituted for coal and coke. One of the disadvantages of charcoal is low density which could affect its specific transport and handling costs. It is possible to utilize fine charcoal as a reductant through injection into a slag bath. However the low density of charcoal particles could affect the capture and utilisation of the contained carbon if the injection velocity is not sufficiently high. Injected particles with low momentum can be carried away in the gas plume without reacting with the slag bath. Alternatively unreactive carbon particles may reach the slag bath but react slowly. This work presents a technique which can be used to measure the reduction efficiency of injected carbon containing particles. The reduction efficiency is a combination of capture of the particles by the slag bath and reaction to remove oxygen. This work is relevant to slag bath reduction processes such as zinc fuming and melter/reduction iron making processes. A slag bath was prepared by melting either 150 or 200 kg of an ironmaking slag in the Sirosmelt furnace at Clayton, Victoria. The iron content of the slag was increased by the co-addition of ferric iron pellets. The slag had a total iron content of 14 wt % and a CaO/SiO2 ratio of 1.3. Heat used to melt the slag and maintain a constant temperature was generated by the combustion of natural gas either in the bath during heat up or above the bath during times of particle injection. The change in composition of the slag bath was determined through the analysis of samples collected during the heat. Slag samples were collected at regular intervals by pushing a steel rod into the slag. A thin layer of slag which adhered to the steel dip rod was quenched in water before being pulverized and analysed through XRF (element totals expressed as oxides) and wet chemical methods (ferrous iron). Fine charcoal or petroleum coke (pet coke) was injected into the slag bath through a side tuyere. This tuyere consisted of a stainless steel tube (5 mm ID) which penetrated the furnace wall at an angle of 45° to the vertical. The end of the tuyere was about 10 cm above the base of the furnace and was between 50 and 60 cm under the surface of the slag. Nitrogen gas was used as the pneumatic transport media. The charcoal and pet coke were stored in a hopper under pressure and were transported first by a screw feeder then pneumatically through a 1 inch hose to the tuyere and hence into the furnace. A load cell under the hopper was used to measure and control the fine solids injection rate. Table 1: Proximate and ultimate analysis of the injected charcoal and pet coke, % dry basis,

(ar: as received, diff: by difference) Moisture(ar) Ash VM FC C H N S O(diff) Charcoal 7.4 1.2 12.7 86.1 88.1 1.8 0.29 0.03 9.78 Pet coke 4.86 0.16 13.0 86.4 87.4 3.5 1.3 7.2 0.6


The proximate and ultimate analysis of the injected charcoal and pet coke is shown in Table 1. The size range of the charcoal and pet coke was between 155 and 500 μm. In one test charcoal with a maximum size of 155 μm was used. The test program consisted of 4 Sirosmelt heats where the variables of slag height (or weight), injectant density (charcoal or pet coke) and injectant paticle size were varied. Table 2 shows the variables used for each test. The injection feed rate was 3 kg/h, transport nitrogen was 10 Nm3/h and the bath temperature and injection velocity were 1450 °C and 142 m/s respectively.

Table 2: Variables used in the testwork

Heat No. Injectant Size range (μm) Slag weight (kg)

1 Charcoal 155 - 500 200 2 Pet coke 155-500 150 3 Charcoal 155-500 150 4 Charcoal -155 150

Carbon injected into a slag bath can remove oxygen contained in the slag components according to equation 1 and/or 2.

O(slag) + C(char) = CO [1] 2O(slag) + C(char) = CO2 [2]

Equation 2 represents a more efficient reduction process because two moles of oxygen is removed for every mole of carbon injected. However in a high temperature process equation 1 is thought to be a more likely reduction mechanism. The carbon utilisation efficiency of the injected charcoal or pet coke can be calculated from the change in the number of moles of oxygen in the slag, assumed to be the sum of FeO and FeO1.5 divided by the net moles of carbon added. This formula is shown in equation 3 where Ceff is the carbon utilisation efficiency.

[3] The net carbon addition can be calculated from the fixed carbon content of the charcoal or pet coke but corrected for the moisture and oxygen content. Oxygen contained in the charcoal or pet coke is assumed to consume carbon which is then unavailable to react with the slag. The calculated net carbon is given in equation 4 where the subscript (db) means on a dry basis.

[4] Figure 1 shows a graph of moles of oxygen removed plotted against net moles of carbon injected to the slag bath for the four injection heats. The data points represent the cumulative carbon added and the cumulative oxygen removed during the heat. All the plots shown in Figure 1 are linear and the slopes of these lines represent the overall carbon efficiency of the different heats. Heat 2 had the lowest carbon efficiency of the four heats at 43 %. Pet coke is much denser than charcoal1) and hence injected pet coke particles would be more likely to reach the slag bath. However it may be possible that the lower reactivity of the pet coke compared to charcoal2) minimized the slag reduction compared to the heats where charcoal was injected. The higher sulphur content of the pet coke may also inhibit the reduction reactions. Sulphur


is known to have a negative effect on the reduction of iron3) and has also been shown to reduce the rate of carbon dissolution into liquid iron4). Heat 3 can be considered as a base case for the three heats involving charcoal. The measured carbon efficiency was 78 %. Increasing the slag depth from 60 to 80 cm (heat 1) would be expected to capture more of the injected charcoal and allow time for the injected particles to more fully react and reduce the slag. This could explain the slightly increased carbon efficiency calculated for heat 1 at 86 %. In heat 4, fine charcoal was injected into the slag. The fine charcoal would be expected to react with the slag at a faster rate than for heat 3 and hence result in increased carbon efficiency. The calculated carbon efficiency for heat 4 was 108 % which is 38 % higher than for heat 3. The carbon efficiency of greater that 100 % indicates that equation 2, i.e. oxygen removal through CO2, may be contributing to the slag reduction. Alternatively the high apparent carbon efficiency may reflect the errors inherent in the calculations and analysis.

Figure 1: Graph of moles of oxygen removed plotted against moles of net carbon added for the four injection heats.

In summary the utilisation efficiency of carbon containing particles injected into a slag bath was found to increase with slag depth and finer charcoal particles. Injected charcoal particles had a higher carbon utilisation efficiency than pet coke particles of the same size. The exact cause of the decreased carbon efficiency of the pet coke was beyond the scope of this work but may be due to lower reactivity or higher sulphur content of the pet coke particles. References 1. CRC Handbook of Chemistry and Physics, 64th Edition, 1983-1984, CRC Press, Boca Raton

Florida, p F-1. 2. D J Harris and I W Smith, Intrinsic reactivity of petroleum coke and brown coal char to carbon

dioxide, steam and oxygen, Twenty Third Symposium (International) on combustion/The Combustion Institute, 1990, pp 1185-1190.

3. G R Belton, How fast can we go? The status of our knowledge on the rates of gas-liquid metal reactions, Proceedings of 10th PTD conference, 1992, pp 3-18

4. D E Langberg, M A Somerville, D E Freeman and B M Washington, The use of Mallee charcoal in metallurgical reactors, Proceedings of Green Processing 2006, Newcastle, AusIMM, pp 69-75.

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Heat 4


PRESENTATION - 24

Innovative Process of Manganese Recovery from Steelmaking Slag by Sulfurization

Sun-Joong Kim1, Hiroyuki Shibata2, Shin-ya Kitamura2, and Katsunori Yamaguchi3

1JSPS Postdoctoral Fellow, Institute of Multidisciplinary Research for Advanced Materials, Tohoku University, 2-1-1 Katahira, Aoba-ku, Sendai 980-8577, Japan

2Institute of Multidisciplinary Research for Advanced Materials, Tohoku University, 2-1-1 Katahira, Aoba-ku, Sendai 980-8577, Japan

3Iwate University, 4-3-5 Ueda, Morioka 020-8511, Japan Keywords: recycling, sulfurization, steelmaking slag, matte, manganese, phosphorus, basicity

Manganese is a key alloying element in various advanced steel productions for improving mechanical properties such as strength and formability. The production of advanced steel has led to an increase in Mn consumption, especially in high-purity grade of Fe-Mn alloys. Despite the important role of Mn in steel products, a recycling strategy to stabilize Mn resources has not yet been investigated. In Japan, the amount of Mn consumed as an alloying element in the steelmaking process is similar to the amount of Mn emitted as a steelmaking slag. However, the separation of P from steelmaking slag is necessary to obtain valuable resources of Mn because the steelmaking slag generally contains about 1.7 mass% P in the form of P2O5

1). The authors have proposed an innovative process for recycling Mn from steelmaking slag via sulfurization of slag because the phosphorus sulfide is unstable at high temperature 2-5). Figure 1 shows the concept of the proposed process.

Figure 1. Concept of innovative recycling process of Mn from steelmaking slag.

The first step is to sulfurize the steelmaking slag containing Mn and P and the liquid oxysulphide phase (matte) containing Mn without P. The next step is to oxidize the obtained matte and form the oxide phase. This oxide phase is named “oxidized slag” to distinguish it from the steelmaking slag. In the second step, the ratio of Mn/Fe in the oxidized slag increases when compared to the matte for the production of Fe–Mn alloys with high Mn content. Furthermore,the sulfur in the remaining matte can be reused in the sulfurization step. In our previous work2-6), we investigated the influences of the atmospheric conditions and slag basicity on the distribution of Mn, Fe, and P between the matte and the slag at 1673K. Since the P content in matte was below ~0.1 mass%, the separation of P could be achieved by sulfurizing the slag. In addition, the distribution of Mn and Fe between the matte/slag increased as the partial pressure (P) of sulfur and the slag basicity increased.The distribution


of the M element between the matte and the slag (LM(M/S)) was determined by the following

equation:

slaginMmassmatteinMmassLM )%(

}%{(M/S) { (1)

When the value of log PS2 was larger than -2 and the slag basicity was larger than 2.8, the value of LMn

(M/S) was larger than LFe(M/S).

Based on the equilibrium distribution of Mn and Fe between the matte and the slag [2-6], we evaluated the Mn yield at the sulfurization stage of steelmaking slag and the oxidation stage of matte by using the mass balance. Mass balances of M elements between the matte and the slag can be derived as follows:

MatteOutputSlag

InputSlag ... MWMWMWM WWWM (2)

Where W.MSlag

Input and W.MSlagOutput are the weights of M elements in the input slag (before

sulfurization) and the output slag (after sulfurization), respectively; ΔM is the weight of M that is transferred from the slag to the matte, i.e., ΔM is equal to the weight of M in the matte. When the matte is assumed to be an FeS–MnS binary system, the mass balance equations can be obtained from ( Eq. 1) as

)().(

)( MnOFeOIntput

Slag

IntputSlag(M/S)

FeMnSFeS CMnCFeW

FeFeWL

CMnCFeFe

CMCF(F

WL

CMCFF (3)

)().(

)( MnOFeOIntput

Slag

IntputSlag(M/S)

MnMnSFeS CMnCFeW

MnMnWL

CMnCFeMn

CMCF(M

WL

CMCFM (4)

where WSlag

Input is the weight of the input slag; CFeO and CMnO are the stoichiometric constants to convert the weight percents of FeO and MnO to that of Fe and Mn, respectively; CFeS and CMnS are the stoichiometric constants to convert the weight percents of Fe and Mn to that of FeS and MnS, respectively. Table 1 shows the composition of input steelmaking slag. When the weight of input slag was100 kg, W.FeSlag

Input and W.MnSlagInputwere19.44 and 7.75 kg, respectively. The respective

values of CFeO, CMnO, CFeS, CMnSwere1.286, 1.291, 1.573, and 1.584. When the slag basicity, log PO2, and log PS2 were 2.8, -11, and -1.9, respectively, LMn

(M/S) and LFe(M/S) were 22 and 17,

based on our previous results 6).

Table 1. Typical composition of input slag.

Figure 2 shows the Mn yield and the Mn ratio in the matte as a function of LMn

(M/S) at 1673K. The Mn yield and Mn ratio were determined by following equations:

100..(%)yieldMn Intput

Slag

Matte 1MnWMnW (5)

100)..(.(%)ratioMn

MatteMatte

Matte 1FeWMnWMnW (6)

As shown in Fig. 2, the Mn yield of the sulfurization stage was 92.7% when log PO2 and log PS2were-11 and -1.9, respectively. However, despite sulfurization of the slag under optimum conditions, the Mn content in the matte was not very high. Therefore, it is important to form Mn-rich oxide phase by oxidation of the matte for the production of Fe–Mn alloys with high

CaO SiO2 FeO MnO P2O5 MgO47 14 25 10 5 7


Mn content. In order to increase the Mn content in the oxidized slag, LMn(M/S)should be

smaller than LFe(M/S). The method for the mass balance calculation was the same as shown in

(Eq. 1–4). However, in the oxidation stage, an assumed value of the sum of CaO and SiO2 percent in the "oxidized slag" was used to calculate the MnO and FeO content. In this research, this value was assumed to be 60%. When the slag basicity, log PO2, and log PS2 were 0.4, -9.8, and -1.4, respectively, LMn

(M/S) and LFe(M/S) were 1.2 and 13.8 based on the

previous experiments. In this condition, , as shown in Fig. 3, LMn(M/S), and LFe

(M/S), the Mn ratio and Mn yield in the "oxidized slag" were determined to be79.2%and 45%, respectively. The large value of Mn ratio in the oxidized slag was achieved. And it is necessary to improve the Mn yield with keeping the large Mn ratio.

Figure 2. Mn yield and Mn ratio in the matte as a function of LMn

(M/S). LFe(M/S) has a constant value of 17.

Figure 3. Mn yield and Mn ratio in the "oxidized slag" as a function of LFe

(M/S) at 1673K. LMn(M/S)has

a constant value of 1.2. References 1. http://slg.jp/slag/slag-seisitsu.htm, Nippon Slag Association: Chemical Composition of

Iron and Steel Slag. 2. S.-J. Kim, T. Hotta, H. Shibata, S. Kitamura, and K. Yamaguchi: Proceeding of the 6th

European slag, Spain, Madrid, 2010, EUROSLAG Publication No. 5, p. 183. 3. S.-J. Kim, H. Shibata, N. Maruoka, S. Kitamura, and K. Yamaguchi: High Temp. Mater.

Proc., 2011, Vol. 30, No. 4/5, pp. 425-434. 4. S.-J. Kim, S. Kitamura, T. Hotta, H. Shibata, and K. Yamaguchi: Proceeding of Fray

International Symposium, Cancun, Mexico, 2011, Vol. 2, p. 183. 5. S.-J. Kim, H. Shibata, J. Takekawa, S. Kitamura, K. Yamaguchi, and Y.-B. Kang,:

Metallurgical and Materials Transactions B, Vol. 43, (2012) pp.1069-1077. 6. S.-J.Kim, H.Shibata,S.Kitamura, and K.Yamaguchi, CAMP-ISIJ, Vol. 25 (2012) p. 242

(M//////SS


PRESENTATION - 25

Vacuum Treatment of Molten Steel in RH (Ruerhstal Heraeus) and VTD (Vacuum Tank Degasser): A Comparative Study

Zulfiadi Zulhan1, Christian Schrade2

1. Metallurgical Engineering Department, Institute of Technology Bandung, Indonesia 2. Technometal GmbH, Duisburg, Germany

Keywords: RH, VTD, vacuum treatment, decarburization, desulphurization. On the beginning of their development, RH and VTD were applied to produce quality steel in term of low hydrogen content for reducing of “hair crack” formation. In the mean time, both of these technologies are installed in steelworks in order to obtain high quality of steel products containing low hydrogen content, low nitrogen content, ultra low carbon content, low total oxygen content as well as ultra low sulphur content. The criteria of technology selection of these vacuum treatment techniques are strictly dictated by the steel grade to be produced. Intensive slag-metal interaction was observed during molten steel treatment on VTD which promotes a good condition for sulphur removal. On the other hand, less slag-metal interaction is taken place during RH treatment. Since its development in Germany in 1950s, a lot of process improvements were performed on RH plant including the installation of oxygen lance in Hattingeni), the enlargement of snorkel and vessel diametersii), as well as the application of powder injection for desulphurizationiii). Comprehensive model for decarburization on RH plant was introduced by Kuwabaraiv) considering the vacuum pressure, liftgas flowrate, vessel as well snorkel diameters. Result on carbon content calculation based on this model compared to plant data is shown on Figure 1. It was reported that the time required to achieve carbon content of less than 20 ppm can be completed in less than 15 minutes on RH plantsii). Typical treatment diagram for ULC steel production on RH plant is shown on Figure 2.

Figure1 - Carbon content during RH-vacuum treatment.

0

50

100

150

200

250

300

350

0 2 4 6 8 10 12 14 16

Time [minutes]

Car

bon

cont

ent [

ppm

]

Data 1Data 2Data 3Model


Figure 2 - Typical treatment – time diagram for ULC steel production on RH plant.

On the other hand, decarburization rate in vacuum tank degasser is less compared to RH plant. Vacuum pressure and argon stirring gas shall be well controlled to avoid molten steel boiling during vacuum treatment. Furthermore, enough ladle freeboard (ca. 1000-1200 mm) shall be provided. Typical carbon content during treatment on VTD is depicted on Figure 3. Intensive slag metal agitation due to stronger gas expansion during vacuum treatment enhances desulphurization rate. The negative effect of slag metal interaction on VTD is the reduction of SiO2 which increase silicon content in the steel (silicon pick-up). Nitrogen can be removed well as oxygen and sulfur contents in molten steel are low. Nitrogen removal usually takes place inline with desulphurization. For production of ULC steel, a model to predict carbon and oxygen contents as well as temperature shall be established for VTD since the temperature and oxygen activity measurements under vacuum are sophisticated compared to RH plant. Typical treatment time for ULC production in VTD is shown in Figure 4. The comparison of RH and VTD are summarized in Table 1.

Figure 3 - Carbon content during VTD-treatment.

0 5 10 15 20 25 30 35 40

Ladle Transfer to Treatment Position (1)Increase ladle position (1)

S1, T1, a[O]1 (2)Immersed snorkel (1)

Vacuum Treatment (26)Decarburization (15)

Oxygen Blowing (0 Nm3) (0)Alloying weighing + transfer to vacuum hopper (2)

T2, a[O]2 (1,5)Al Deoxidation + Alloying (3)

Alloying addition (3)Mixing + Homogenization (5)

Flooding (1)Lowering Ladle (1)

S3, T3, a[O]3 (0,5)Transfer ladle to wire feeding (0,5)

Wire feeding (0)Transfer ladle to covering powder (0,5)

Covering Powder addition (0,5)Transfer ladle to lifting position (0,5)

Time [Min]

1500

1520

1540

1560

1580

1600

1620

1640

1660

Tem

pera

ture

[°C

]

Steel Temperature (calculated)

Steel Temperature Measured

0

20

40

60

80

100

120

140

160

180

200

0 5 10 15 20 25Time [minutes]

Car

bon

cont

ent [

ppm

] Plant Data

Calculated


Figure 4 - Typical treatment – time diagram for ULC steel production on VTD plant.

Table 1. Comparison of RH and VTD

VTD RH Ladle freeboard 600 – 1200 mm 200 – 300 mm Sampling, temperature and oxygen activity measurements

Taking sample and temperature measurement under vacuum is complicated

Taking sample and temperature measurement under vacuum is simple

Observation of molten steel height / surface during vacuum treatment

Depends on the opening diameter of ladle cover heat shield.

Easy by installation a video camera on hot-off take

Boiling Vacuum pressure and argon stirring shall be well controlled to avoid excessive boiling.

Easy to control due to high feeboard of RH vessel.

Decarburization < 30 ppm in 25 minutes < 20 ppm in 15 minutes Denitrogenization < 40 ppm < 40 ppm Dehydrogenation < 1.5 ppm < 1.5 ppm Desulphurization Intensive slag metal reaction Possible through powder

injection Silicon pick-up? Yes, SiO2 content in slag shall

be controlled to produce free silicon steel.

No, less slag metal interaction.

References 1. H.P. Haastert, “Entwicklungsrichtungen der Sekundärmetallurgie, im besonderen das RH-

Verfahren zur Vakuumbehandlung”, Stahl und Eisen, Vol. 111, No. 3, 1991, pp. 103-109. 2. C. Schrade, M. Huellen and Z. Zulhan “New Concepts for High Productivity RH Plants”, Stahl

und Eisen, Vol. 126, No. 11, 2006, pp. S109-S120. 3. N. Van Poucke “Use of the MESID Lance in the RH-Degasser at the Sidmar Steel Plant”, 4th

European Oxygen Steelmkaing Conference, Graz, Austria, 12th – 15th May 2003. 4. T. Kuwabara, K. Umezawa, K. Mori, H. Watanabe “Investigation of decarburization behaviour

in RH-reactor and its operation improvement”, Trans. ISIJ, Vol. 28, 1988, pp. 305-314.

0 5 10 15 20 25 30 35 40 45 50 55 60 65

Ladle Seated into VD Tank (2)Argon connection and started (1)

S1, T1, a[O]1 (2)Cover closed (2)

Vacuum Treatment (37)Decarburization (25)

Al (CaO/Flux) weigh. + trans. to lower vac hopper (4)Oxygen Blowing for de-C (0 Nm3) (0)

Alloying weigh. + trans. to upper vac hopper (1)Al(CaO/Flux) Addition for Deox + Alloy + Heating (3)

Alloying Charging (3)Mixing + Homogenization (5)

Flooding (1)Cover opened (2)

S2, T2, a[O]2 (2)Trimming (6)

Wire feeding (0)S3, T3, a[O]3 (2)Ladle lifting (2)

()

Time [Min]

1500

1520

1540

1560

1580

1600

1620

1640

1660

1680

1700

Tem

pera

ture

[°C

]

Steel Temperature Predicted

Steel Temperature Measured


PRESENTATION - 26

Experimental Investigations on the Dynamics of Interfacial Phenomena in Synthetic Blast Furnace Slags

Luckman Muhmood and Mirco Wegener

CSIRO Process Science and Engineering, Clayton, VIC 3168 Australia Keywords: dynamic surface tension, droplet formation, blast furnace slag, interfacial phenomena The term "interfacial phenomena" implicates a near endless scope of applications amongst various scientific fields. All multiphase systems do exhibit at least one if not more interfaces. In the pyrometallurgical area, slags (multicomponent molten oxides) are used in many sub-processes for different purposes such as protection of the metal phase against atmosphere (sealing the metal from hydrogen, nitrogen etc.), removal of undesired impurities (e.g. sulphur, phosphorus, oxygen), gangue and coke ash, thermal insulation, prevention of "skull formation" (solid pieces of steel left in refractory-lined vessels after use), optimum slag composition for low refractory interaction/dissolution (e.g. protection of lining from arc in Electric Arc Furnaces).1-3 A change in interfacial tension can be created by differences in temperature, composition or electric charge caused by chemical reactions, fluid dynamics, mass transfer, preferential adsorption of surfactants, etc. The local difference in surface energy causes an additional tangential shear stress which results in the movement of surfactant species along the interface from regions of lower to regions of higher interfacial tension - the Marangoni convection. The Marangoni convection is a chaotic interfacial phenomenon which is inherently three-dimensional in nature and so far analytically unpredictable. It in turn, significantly affects the fluid dynamics and mass transfer, which are always coupled and therefore creates highly complex interactions in the affected systems, as for example shown in liquid/liquid extraction systems.4,5 Industrial multicomponent blast furnace slag always contains surface active components at different concentration levels which may significantly affect the surface tension and the dynamics of surface related phenomena, e.g. the droplet formation, coalescence, droplet oscillating behaviour, and many others. Sub-processes like droplet breakup or coalescence may happen within fractions of milliseconds, and thus equilibrium values of surface tension are not always applicable. For example, the time scale of droplet or bubble ejection6-8 or droplet disintegration from ligaments in the dry slag granulation technology9,10 (see Fig. 1) is in the order of magnitude of milliseconds. This means that static values for the interfacial or surface tension may not be applicable if the time scale of the equilibration transport processes is significantly longer. The use of static values in these dynamic cases would

Fig. 1: Slag ligament formation with consecutive droplet breakup at a spinning disc. CSIRO’s Dry Slag Granulation Process.10


consequently lead to erroneous calculations for heat and mass transfer correlations or reaction kinetics. CSIRO’s dry slag granulation technology incorporates the pouring of molten slag on a disc rotating at a predetermined rotation speed. The molten slag spreads after impinging on the disc on its surface and emerges at its periphery as multiple curved jets which consecutively disintegrate into droplets within a specific size range. The success of the technology for effective cooling and heat recovery of molten slag depends on the size distribution of the droplets formed - too small drop size range causes the formation of accretions on the walls of the rig, thus reducing the heat recovery, while a larger drop size range results in ineffective cooling, agglomeration and compulsory grinding process before transport of the final product. The objective of the present project is to investigate the dynamic surface tension, droplet formation (including necking) and formation criteria of satellite drops in synthetic blast furnace slags (calcia/silica/alumina/magnesia based) in the presence of surfactant oxides (K2O, Na2O) in the temperature range 1450-1600°C. In literature, to the authors’ knowledge, no attempts have been reported to address these dynamic interfacial phenomena in molten oxides at small time scales (fractions of seconds and lower). For such time scales, the oscillating jet method, which relates the instability growth rate of a vertical laminar liquid jet emerging from a nozzle to the surface tension, was found to be suitable to extract the relevant data in aqueous systems and low viscosity melts at relatively low temperatures. The method will be adapted to blast furnace slags with comparatively high viscosity and high interfacial tension. In case of primary and satellite droplet formation, droplet detachment and dynamic surface tension in relatively larger time scales (from fractions of a second to seconds), the pendant drop method is used. This incorporates the careful generation and growth (until detachment) of droplets of molten slag under different conditions and variables like composition, temperature, flow rate and nozzle/drop diameter. In order to carry out these investigations, two critical elements need to be focussed on. Firstly, designing a set-up which has the flexibility to control the flow of slag (addressing both the jet and drop regime) and at the same time having the provision to observe the slag flow contours with a high-speed camera via a quartz window; and secondly, selecting a slag composition with an appropriate surfactant and its equilibrium thermophysical properties like surface tension, density and viscosity. Addressing the primary requirement of a suitable furnace, a 3-zone split furnace with provision to load an alumina cross tube with optical access was constructed. The equilibrium surface tension of slag with various concentrations of surfactant is required to obtain an idea of the sensitivity of the surface tension with respect to the surfactant concentration. A larger change in surface tension with the addition of small quantities of surfactant ensures that the value measured is a reflection of the effect of surfactant and not due to random errors. In case of slags, for initial trials 50% CaO - 50% Al2O3 slag was used and silica was added in various concentrations with a limit of 10% by weight. A graphite crucible and capillary arrangement is used for this slag. Fig. 2 shows the results on the recent equilibrium surface tension measurements for a CaO/Al2O3/SiO2 slag with varying silica content at constant CaO/Al2O3 ratio (= 1). The sessile drop method was used and the drop was kept on a graphite substrate in argon atmosphere under a partial oxygen pressure of 10-10 atm.


The current paper focuses on the challenges of the methodology adapted to measure dynamic surface tension, necking and satellite drop formation in slags. Details on the speciality furnace, slag selection and equilibrium surface tension measurements will also be discussed.

Fig. 2: Equilibrium surface tension of a CaO/Al2O3/SiO2 slag with varying silica content and fixed CaO/Al2O3 ratio (equals 1) at 1773 K. The error in measurement was ± 4 mN/m. References 1. L. Holappa, S. Nurmi, S. Louhenkilpi, Role of slags in steel refining: is it really

understood and fully exploited?, Revue de Metallurgie Vol. 106, No. 1, 2009, pp. 9-20. 2. K.C. Mills, S. Sridhar, Viscosities of ironmaking and steelmaking slags, Ironmaking &

Steelmaking, Vol. 26, No. 4, 1999, pp. 262-268. 3. K.C. Mills, The estimation of slag properties, in: Proceedings of South African

Pyrometallurgy 2011, South Africa. 4. M. Wegener, J. Grünig, J. Stüber, A.R. Paschedag, M. Kraume, Transient rise velocity

and mass transfer of a single drop with interfacial instabilities-experimental investigations, Chemical Engineering Science, Vol. 62, No. 11, 2007, pp. 2967-2978.

5. M. Wegener, A.R. Paschedag, The effect of soluble anionic surfactants on rise velocity and mass transfer at single droplets in systems with Marangoni instabilities, International Journal of Heat and Mass Transfer, Vol. 55, No. 5-6, 2012, pp. 1561-1573.

6. I. Hahn, D. Neuschütz, Ejection of steel and slag droplets from gas stirred steel melts, Ironmaking & Steelmaking, Vol. 29, No. 3, 2002, pp. 219-223

7. Z. Han, L. Holappa, Bubble bursting phenomenon in Gas/Metal/Slag systems, Metallurgical and Materials Transactions B, Vol. 34, No. 5, 2003, pp. 525-532.

8. F.D. Richardson, Drops and bubbles in extractive metallurgy, Metallurgical and Materials Transactions B 2, Vol. 10, 1971, pp. 2747-2756.

9. S. Jahanshahi, D. Xie, Y.H. Pan, P. Ridgeway, J. Mathieson, Dry slag granulation with integrated heat recovery, in: Proceedings of 1st International Conference on Energy Efficiency and CO2 Reduction in the Steel Industry, 27 June - 1 July 2011, Düsseldorf, Germany.

10. D. Xie, S. Jahanshahi, T. Norgate, Dry granulation to provide a sustainable option for slag treatment, in: Proceedings of Sustainable Mining Conference, 17-19 August 2010, Kalgoorlie, Western Australia.

= -4.9311wt% SiO2 + 649.3R 2 = 0.9759

550

600

650

700

0 3 6 9 12SiO2 (wt %)

surf

ace

tens

ion

(mN

/m)


PRESENTATION - 27

Synthesis of High Purity Silicon Carbide for Solar Silicon Production

Xiang Li1, Guangqing Zhang1, Ragnar Tronstad2, Oleg Ostrovski3 1School of Mechanical, Materials and Mechatronic Engineering, University of Wollongong,

Wollongong, NSW 2522, Australia 2Elkem AS, Hoffsveien 65B, 0377 Oslo, Norway

3School of Materials Science and Engineering, University of New South Wales, Sydney, NSW 2052, Australia

Keywords: silicon carbide, carbothermal reduction, trace elements Solar energy alone has the potential capacity to meet the planet’s entire energy needs in the future [1]. The majority of solar cells are made of high purity silicon [2], in which the maximum impurity level is below ~1 ppm, that is, six nines (99.9999%) purity. The greatest limitation for silicon based photovoltaic systems is the cost competitive high purity silicon, having especially low levels of boron and phosphorus. Therefore, the ultimate aim of this research project is to develop an innovative technology to produce low-cost high purity solar grade silicon, which will be achieved by two steps: 1) Conversion of silica to silicon carbide, and removal of impurities in the course of SiC formation; 2) Silicon carbide further conversion to high purity silicon. This work is focused on step 1). Synthesis of SiC was studied in a fixed bed reactor in the H2-Ar gas mixture. The quartz lumps supplied by Elkem were crushed into <70 μm using zirconia grinding set. The quartz powder was wet mixed uniformly with graphite powder (<45μm) with a desired C/O molar ratio and then dried. The mixture powder was compressed into pellets (8 mm diameter and 10 mm length, 1.0 g mass) at 10 MPa. A pellet was loaded into an alumina reactor tube in which quartz was carbothermally reduced in a vertical furnace. After reduction the raw materials and reduced samples were analysed by XRD, SEM/EDS and LA-ICP-MS. The experimental conditions are listed in Table 1.

Table 1. Experimental conditions of carbothermal reduction

Parameter examined

H2 concentration (%)

C/Si mole ratio

Temperature ( )

Reaction time (h)

Gas flow rate (L/min)

H2 concentration 0~100 3 1500 4 1

C/Si mole ratio 100 2~6 1500 4 1 Temperature 100 3 1100~1550 4 1

Reaction time 100 3 1500 0.5~6 1 Effect of H2 concentration Fig.1 shows XRD patterns of samples reduced in gas atmosphere with different H2 concentration (1L/min). The sample reduced under 100% Ar atmosphere consisted of β-SiC, unreacted C and residual quartz. Addition of 20% of H2 resulted in disappearance of quartz peaks, and strengthening SiC peaks. Graphite peaks indicate incomplete reduction of silica. With the increase of H2 concentration to 100%, the graphite peak also disappeared.


Carbothermal reduction of silica by Reaction (1) can be enhanced by Reaction (2) which assists transfer of carbon:

SiO2 + 3C = SiC + 2CO (1) C + 2H2 = CH4 (2)

Figure 1. XRD patterns of samples reduced at 1500oC in the Ar-H2 gas mixture with different H2 concentration

Trace impurity elements in samples have been determined using laser ablation inductively coupled plasma mass spectrometry (LA-ICP-MS). Generally, carbothermal reduction of quartz in hydrogen resulted in higher removal rate of impurities. Aluminium content in some samples increased after reaction, which can be attributed to reduction of alumina tube generating Al and Al2O vapour which deposited on the sample pellets [3]. Effect of C/Si mole ratio, temperature and reaction time C/Si ratio varied from 2 to 6. The XRD patterns shown in Fig.2 (a) indicate that SiC peaks intensity got stronger with the increasing of the C/Si ratio from 2:1 to 3:1. Reaction temperature was varied from 1100 °C to 1550 °C. The XRD patterns of reduced samples at different temperatures are shown in Fig.2 (b). Increase in temperature above 1200 °C resulted in formation of SiC. When temperature was increased to 1400 °C, all quartz was reduced. Reaction time was varied from 0.5 h to 6 h. The XRD patterns of reduced samples are shown in Fig.2 (c). Reaction time had a significant effect on the SiC peak strength only within the first hour. Residual graphite peak decreased with the extension of the reaction time due to further reaction with residual quartz and H2.


Figure 2. XRD patterns of samples: a) reduced with different C/Si ratio; b) reduced at different temperature; c) reduced for different reaction time.

It can be concluded that carbothermal reduction in hydrogen facilitates the removal of impurities. The effect of C/Si mole ratio, temperature and reaction time on behaviors of impurities during carbothermal reduction requires further examination, which is currently in progress. References 1. N. Lewis, G. Crabtree, A. Nozik, M. Wasielewski and P. Alivisatos, Basic research

needs for solar energy utilization, U.S. Department of Energy, 2005. 2. S. Ranjan, S. Balaji, Rocco A. Panella and B. Erik Ydstie, Silicon solar cell production,

Comput. Chem. Eng., Vol.35, No. 8, 2011, pp.1439-1453. 3. J. Li, G. Zhang, D. Liu and O. Ostrovski, Low-temperature synthesis of aluminium

carbide. ISIJ International, Vol. 51, No. 6, 2011, pp. 870-877.


PRESENTATION - 28

Thermodynamic Optimization of Iron-Silicate Slag for Simulation of Copper Smelting Processes

Taufiq Hidayat 1, Denis Shishin 2, Sergei Decterov 2, and Evgueni Jak 1

1 PYROSEARCH, the University of Queensland, Brisbane, Queensland, Australia 2 Centre de Recherce en Calcul Thermochimique (CRCT), École Polytechnique, Montréal,

Québec, Canada

Keywords: phase equilibria, copper, iron silicate slag, thermodynamic database

Iron-silicate slag is of primary importance to various metallurgical processes, in particular to copper smelting. The ability to predict phase equilibria and thermodynamic properties of iron-silicate slag in the complex chemical system pertaining to copper smelting, converting and slag cleaning is essential for further improvements in industrial operations. The current thermodynamic database for this system [1] still requires further improvement for more accurate predictions. Thermodynamic optimization of the “Cu2O”-FeO-Fe2O3-SiO2-S system is the focus of the present stage of the work. During the course of the optimization, it was realized that re-evaluation of basic Cu-free system of FeO-Fe2O3-SiO2 is also required; the recently improved phase diagram of FeO-Fe2O3-SiO2 system is given in Figure 1. The present thermodynamic optimization is carried out as part of the overall integrated research program on complex copper-containing slag-matte-metal-solids equilibria for copper production, which combines experimental and thermodynamic modelling studies in the multi-component Al2O3-CaO-“Cu2O”-FeO-Fe2O3-MgO-SiO2-S system.

Figure 1 - Calculated liquidus surface and liquidus oxygen isobars (atm) in the FeO-Fe2O3-

SiO2 system. A set of self-consistent parameters of thermodynamic models for the Gibbs energies of all stoichiometric and solution phases has been optimized using the available experimental phase


diagram and thermodynamic data. The spinel (magnetite) and monoxide (wustite) phases which were modelled previously [2] using the sublattice (based on the compound-energy formalism) and polynomial models, respectively, have been further optimized to take into account the presence of calcium and copper in the solid solutions. The Gibbs energy of the molten slag phase, which exhibits a tendency towards short-range ordering, has been described by the Modified Quasichemical model available in the FactSage computer package [3]. The slag model has gone through several improvements including: complete flexibility to choose “Kohler” or “Kohler-Toop” interpolation of binary parameters into any ternary subsystem based on basic, acidic, or amphoteric behaviour of different oxides; the flexibility to set the composition of the maximum short-range ordering in a particular binary system regardless of all other systems; the possibility to combine the Bragg-Williams model with the Modified Quasichemical model to treat liquid solutions with positive deviations from ideality; and the introduction of sulphur in addition to oxygen on the anion sublattice using the quadruplet approximation. The models permit extrapolation into regions of temperature and composition where experimental data are not available. Details of the recent thermodynamic optimization of the investigated system are presented. Examples of good agreement between experimental data and the current model are shown in Figure 2.

Figure 2 – Solubility of Cu in tridymite-saturated slag as a function of Log10[P(O2), atm]:

(A) at equilibrium with pure copper; and (B) at equilibrium with copper-gold alloy (aCu(l)=0.73).

0.0

5.0

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-12.0 -11.0 -10.0 -9.0 -8.0 -7.0 -6.0 -5.0

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ct C

u in

slag

Log10[P(O2), atm]

[1966Ruddle et al] - 1300 ⁰C[1978Elliot et al] - 1300 ⁰C[1983Oishi et al] - 1300 ⁰C[2008Kim & Sohn] - 1250 ⁰C[2012Henao et al] - 1250 ⁰C[2012Henao et al] - 1300 ⁰C[2012Hidayat et al] - 1250 ⁰C[2012Hidayat et al] - 1300 ⁰CThis study - 1250 ⁰CThis study - 1300 ⁰C

0.0

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Log10[P(O2), atm]

[1972Altman & Kellogg] - 1220 ⁰C[1972Altman & Kellogg] - 1260 ⁰C[1972Altman & Kellogg] - 1285 ⁰CThis study - 1220 ⁰CThis study - 1260 ⁰CThis study - 1285 ⁰C

(A)

(B)


The final optimized database can be used as a basis for simulation of copper smelting processes and other high temperature metallurgical processes. The ability of FactSage (with the optimized thermodynamic database) to calculate and predict the phase equilibria, partitioning of major and minor elements between phases, activities, vapour pressures, and other thermodynamic properties over a wide range of process conditions in copper production can enable informed decision to be made in relation to the optimum slag compositions, fluxing strategies, and operating conditions in copper smelting, converting, and slag cleaning processes for given feed and product requirements. References 1. Decterov, S.A. and Pelton, A.D.: Metall. Mater. Trans. B, 1999, vol. 30B (4), pp. 661-9. 2. Decterov, S.A., Jak, E., Hayes, P.C., and Pelton, A.D.: Metall. Mater. Trans. B, 2001,

vol. 32 B (4), pp. 643-57. 3. FactSage. 2012: Ecole Polytechnique, Montréal, http://www.factsage.com/.


PRESENTATION - 29 Thermodynamic Assessment of In-Situ Formation of Hard Phase Materials

Ryan Cottam1, Samuel Ng’anjo1, M. Akbar Rhamdhani1 1Industrial Laser Applications Laboratory, IRIS, Faculty of Engineering and Industrial Sciences, Swinburne University of Technology, Hawthorn, Victoria, 3122, Australia

Keywords: Thermodynamics; Laser Cladding; Hard Phase Materials The hardfacing of metallic components for applications in the mining as well as the oil and gas sectors is common. Therefore method to enhance the wear resistance of the applied coatings is attractive. One of the ways to increase the wear resistance of the coatings is to introduce a hard phase into the coating, the most common of which is tungsten carbide. Typically the carbide compound is mixed with the hard facing materials such as Stellite and laser clad on the surface to be coated. Recently there has been increasing interest in forming the hard phases in-situ [1, 2]. In this process the hard phase material is formed in the melt pool formed during laser processing and the elements that constitute the hard phase combine in the melt pool prior to solidification of the melt pool. Using thermodynamic analysis potential new hard phase materials could be determined. In this study, a systematic thermodynamic evaluation on the formation of hard phase materials during laser cladding of Stellite was carried out. Hard phases are typically carbides, nitrides, borides or silicates to a lesser extent. Therefore a range of metals Ca, Sc, V, Mn, Co, Sr, Zr and Nb where combined with one of the non-metals of B, N, C and Si and the phases that formed when combined with the Stellite composition was simulated using FactSage. The composition of Stellite is shown in Table 1. It is a cobalt based material with major alloying additions of Cr and W. Therefore to form a hard phase in-situ these elements need to be considered.

Table 1- Composition of Stellite 6

Element Co Cr Si W C Mn Ni,Fe,Mo wt% 64.3 28 1.1 4.5 1.2 1 <2

Figure 1 – Thermodynamic assessment of Hard phase materials; (a) – Ellingham diagram for Zr hard phase materials showing Gibbs free energy formation of ZrN, ZrC, ZrC4 and ZrB2; (b) predicted equilibrium composition of hard phases formed for Stellite 6 mixed with Zr and N2 at temperatures between 800 to 3000oC


The Gibbs free energy formation of various hard phase materials was evaluated and followed by equilibrium calculations for the candidate elements. It was found that the following combinations of elements could potentially form the in-situ hard phase; ZrN, ZrC, and ZrB, for example in the case of laser cladding in the presence of pure N2. Figure 1(a) shows the Gibbs free energy formation ( Gf

o) of ZrN, ZrC, ZrC4 and ZrB2. The curves are relatively flat for ZrN, ZrC and ZrB2 indicating the stability of the compound over the temperature range studied. Figure 1(b) shows the predicted equilibrium composition of the condensed phases at temperature range of 800 to 3000oC. It can be seen that ZrN is quite stable at the temperature range studied. Due to the ease with which the additions of Zr and N could be added to the Stellite this hard phase reaction was experimentally trialled. Stellite 6 was mixed with Zr with an amount that made 15wt%Zr. The powder mixture was then laser clad in a nitrogen atmosphere allowing the in-situ hard phase reaction to occur. It is clear from the microstructural analysis that the phase has formed and the matrix material has not undergone any dramatic changes, as shown in Figures 2 and 3. The XRD spectra of the Stellite 6 and Stellite 6 + ZrN, as shown in Figure 4, reveals the formation of new hard phase indicated by new peaks. Further studies including characterisations of the samples to identify optimum process parameters are being carried out.

a b

Figure 2 - Etched optical micrograph of; a – Stellite 6; b – stellte 6 +ZrN

Figure 3 – un-etched optical micrograph of laser clad Stellite 6 + ZrN


Figure 4 – XRD scan confirming the exisitance of a new pahse in the Stellite6 +ZrN materials.

References 1. Xu, J. and W. Liu, Wear characteristic of in situ synthetic TiB2 particulate-reinforced

Almatrix composite formed by laser cladding. Wear. 2006, 260, 486-492. 2. Du, B., Z. Zou, X. Wang, and S. Qu, Laser cladding of in situ TiB2/Fe composite

coating on steel. Applied Surface Science. 2008, 254, 6489-6494.


PRESENTATION - 30

Behaviour of New Zealand Ironsand during Iron Ore Sintering

Zhe Wang1, Harold Rogers2, David Pinson2, Brian J. Monaghan1, Sheng Chew2, Paul Zulli2, Sharon Nightingale1, and Guangqing Zhang1

1University of Wollongong, Wollongong, NSW2522 Australia 2BlueScope Steel Research, P.O. Box 202, Port Kembla, NSW 2505 Australia

Keywords: New Zealand ironsand; titanomagnetite; sintering; iron ore Titanium-bearing burdens are commonly introduced into blast furnaces to protect the hearth because the so-called “titanium bear” which is a precipitate of carbide, nitride and carbonitride of titanium may form in the blast furnace hearth if TiO2 is present in the feed [1, 2]. New Zealand ironsand is a titanomagnetite, containing around 60 wt.% iron, 8 wt.% titanium and other substances such as silica, phosphorus and lime [3, 4]. Since it is competitive in price, introduction of the ironsand into the ferrous feed can reduce the production cost and potentially increase blast furnace campaign life. An appropriate method of introduction of ironsand is as a component of the sinter as the small size of ironsand precludes direct charging into the blast furnace. Although the effect of introducing titanomagnetite into iron ore blends has been investigated [1, 5, 6], little is known about the detailed sintering mechanism. The present study is aimed at identifying the sintering behaviour of New Zealand ironsand as well as the interaction between New Zealand ironsand and CaO to gain better understanding of sintering mechanism of titanomagnetite. Experimental Procedure The raw materials for this study were iron ore blend, limestone, silica sand, manganese ore and Cold Return Fines (CRF). Each was crushed and screened to a particle size smaller than 200 μm before use. These materials were proportioned according to BlueScope Steel’s Sinter Plant practice. To investigate the effect of ironsand concentration, 0 wt%, 3 wt% and 5 wt% ironsand were introduced to the mixture of which 1.0 g was then pressed into cylindrical tablets (8 mm diameter, ~5 mm height. The samples were sintered in a vertical tube furnace at three different temperatures (1100°C, 1200°C, and 1300°C) in a gas mixture of CO = 1%, CO2 = 24%, and Ar = 75% (previous work by Hsieh and Whiteman [7] had shown the gas mixture simulated industrial sintering). After sintering for the desired time, the samples were directly removed to the cool top end of the reaction tube. Then the samples were mounted in epoxy resin and cut parallel to the original top surface of the sample, and polished in preparation for optical and scanning electron microscopy (SEM). To study the microstructural development greater detail, interaction couples were prepared for ironsand and CaO powder and heat treated under Argon at 1100, 1200, and 1300°C as before. After sintering, the pellets were sectioned perpendicular to the interface and polished. The compositions of the phases in the cross sections were analysed by SEM/EDS.


Results

Assimilation of New Zealand ironsand into melts during sintering The extent of coalescence of particles in raw mixture increased gradually with an increase in sintering temperature and time. At 1100°C, the assimilation effect was not remarkable. The amount of silicate bonding phase became pronounced at 1200°C. At 1300°C, the agglomeration of oxide materials was near complete and domains were dominated by polycrystalline masses of magnetite distributed in continuous silicate matrix. Titanomagnetite grains remained identifiable in the samples sintered at 1100°C and 1200°C and the hematite exsolution lamellae coarsened marginally with time. In a sample sintered at 1300°C for 12minutes, very limited evidence of relict titanomagnetite grains suggests a large proportion of the ironsand had assimilated into the sintered mass. EDS results showed the reactions started with lime in contact with acidic oxide compounds in the mixture. Therefore, calcium silicate phase containing iron oxides and alumina formed. As temperature and time increased, the new phase (with melting point of 1200°C) gradually melted and seeped between the large grains of the mixture including ironsand grains, disolving these grains progressively. As the surface of titanomagnetite dissolved into the melt, a reaction zone around a ironsand grain was establshed, gradually increasing with time. In this zone, Ca2+ from silicate melt diffused into the Fe3O4 lattice, with Fe2+ diffusing in the opposite direction towards the melt. On one hand, the diffusion of Ca2+ to Fe3O4 decreased the melting point of Fe3O4 and accelerated the melting rate of titanomagnetite grain. On the other hand, CaO·TiO2 was generated as the result of the reaction between TiO2 and CaO.

Figure 1. EDS line analysis across the reaction zone of a New Zealand sand and CaO couple heated at 1200°C for 20 minutes

Figure 2. EDS line analysis across the reaction zone of a New Zealand sand and CaO couple heated at 1300°C for 12 minutes


Interaction between New Zealand ironsand and CaO A reaction zone between the ironsand and CaO layers was formed in the sample heated at 1200°C that became wider with increasing heating time and temperature. Figure 1 shows the elemental distribution across the reaction zone of a sample heated at 1200°C for 20 minutes. Ca2+ from CaO diffused into the lattice of Fe3O4, and Fe2+ diffused in the opposite direction into CaO. As a result, the reaction zone contained a high and uniform level of Ca and a relatively low level of Fe. The reaction zone consisted of two subzones, one being next to the unreacted ironsand layer with a relatively high TiO2 concentration (notably higher than the average for an ironland particle), while the other close to the CaO layer having a high level of iron oxides. In the sample heated at 1300°C for 12 minutes(Figure 2), the subzone containing more iron oxides melted and penetrated into the CaO layer due to formation of a large amount of calcium ferrite (with low melting point), while the TiO2-rich subzone became wider and did not melt since TiO2 increases the liqudus temperature of CaO-TiO2-FeOx system. References 1. N.J. Bristow and C.E. Loo, "Sintering Properties of Iron Ore Mixes Containing

Titanium", ISIJ International, Vol. 32, No. 7, 1992, pp. 819-828. 2. Y. Li, Y.Q. Li, and R.J. Fruehan, "Formation of Titanium Carbonitride from Hot Metal",

ISIJ International, Vol. 41, No. 12, 2001, pp. 1417-1422. 3. J.B. Wright, "Iron-Titanium Oxides in Some New Zealand Ironsands", New Zealand

Journal of Geology and Geophysics, Vol. 7, 1964, pp. 424-444. 4. H.A. Cocker. et al., "Where is the Titanium in the Ironsands?-Ti Partitioning in the

Magnetic Fraction", AusIMM New Zealand Branch Annual Conference 2010, 2010, pp.165-174.

5. T. Paananen and K. Kinnunen, "The Effect of Titanium on Reduction Degradation of Iron Ore Agglomerate", Iron Ore Conference 2007, 2007, pp. 361-367.

6. Z.K. Yin, J.S. Li and S.F. Yang, "Sintering Pot Test on Improving TiO2-Containing Ore's Allocated Proportion", Advanced Materials Research, Vols. 311-313, 2011, pp. 850-853.

7. L.H. Hsieh and and J.A. Whiteman, "Effect of Oxygen on Mineral Formation in Lime-Fluxed Iron Ore Sinter", ISIJ International, Vol. 29, No. 8, 1989, pp. 625-634.


PRESENTATION - 31

Phase Equilibria Studies of Cu-S and Cu-Fe-S systems

Imam Santoso, Hector Henao, Evgueni Jak and Peter Hayes PYROSEARCH, University of Queensland, St Lucia, QLD 4072

Australia Keywords: copper smelting, matte, EPMA, phase equilibria The study of phase equilibria in the Cu-S and Cu-Fe-S system at high temperature is important for improving the operation of copper smelting and converting processes. A review of previous studies has shown however, that there are significant differences and omissions in the available data sets. To investigate this system, the experimental procedures developed at PYROSEARCH at the University of Queensland have been used, which involve equilibration of mixtures at high temperatures, rapid quenching, and measurement of phase compositions using electron probe X-ray microanalyses (EPMA). The samples were equilibrated in Ar atmosphere using different substrates i.e Mo foil for Cu-S equilibration and SiO2 for Cu-Fe-S matte equilibration. The phase equilibria of Cu-S and Cu-Fe-S systems have been characterized between 1150oC and 1250oC. The experimental results show that the technique can be applied successfully for accurate determination of phases in the Cu-S system and the Cu-Fe-S system at copper liquid saturation. The matte phase can be quenched to an amorphous or microcrystalline state. Figure 1 shows the microstructure of a sample containing Cu-S matte and liquid copper. The dispersion of copper droplet ensures the matte throughout the sample is at a metal saturation. The cracks in matte phase are formed on cooling of the solid phase

Figure 1. Backscattered image of microstructure of Cu-S system at copper saturation at 1250OC equilibrated using a Mo substrate for 8 hrs in high purity Ar gas


Figure 2 shows that there is a close agreement between FactSage prediction and the present study for Cu-S system. For the matte phase, all the experimental data from previous studies indicate the liquid matte (L2) is lower in sulfur content than that predicted by FactSage. Figure 3 shows new experimental data in the Cu-Fe-S matte in equilibrium with liquid copper. The only other data are available in the Cu-Fe-S matte at iron saturation. The data obtained in the present study at iron saturation will be compared with FactSage prediction. References 1. Bale, C.W., and Toguri, J.M., Thermodynamics of the Cu-S, Fe-S and Cu-Fe-S systems,

1976, Canadian Metallurgical Quarterly, 15: p. 305-318. 2. Johannsen, V.F.and Vollmer, H., Untersuchungen im System Kupfer-Kupfersulfid, 1960,

Erzmetall, 13: p. 313-322. 3. Judin, V.P. and Eerola, M., Thermodynamics of metallic impurities in copper-saturated

copper sulphide melts, 1979, Scandinavian Journal of Metallurgy, 8: p. 128-132. 4. Schmiedl, J., Repcak, V. and Cempa, S., Equilibrium studies in the system Cu-S-O,

1977, Trans.-Inst. Min. Metall., Sect. C, 86: p. C88-C91. 5. Sudo, K., Fundamental researches on smelting of sulphide ores. VII. On the equilibrium

in the reduction of copper sulphide in molten copper by hydrogen gas, 1950, Sci. Rep. Res. Inst. Tohoku Univ. A, 2: p. 519-530.

6. Schuhmann, Jr., R. and Moles, O.W., Sulphur activities in liquid copper sulphides, 1951, Journal of metals, 191: p. 235-241


PRESENTATION – 32

A Comprehensive Approach for CFD Modeling of Slag Foaming with Population Balance Modeling

Md.Abdus Sattar, Jamal Naser, Geoffrey Brooks

Faculty of Engineering and Industrial Science, Swinburne University of Technology

Hawthorn VIC-3122, Australia. Email:[email protected]

Keywords: Bubble break up, Coalescence, Film rupture, Population balance, Slag foaming A comprehensive approach for simulating the formation of slag foam is proposed. Foam is considered as a separate phase which is comprised of a mixture of air and liquid. A computational fluid dynamic (CFD) model has been developed for the simulation of slag foaming on bath smelting slag (CaO-SiO-Al2O3-FeO). The model accounts for the formation of foam due to transformation of both gas and liquid into foam and it’s destruction due to liquid drainage and bursting of bubble. The bubble break up model of Luo and Svendsen 1) and the coalescence model of Prince and Blanch 2) are used in the present study. The bubble coalescence model based on film rupture by Tong et al. 3) has been used in the foam layer of the present study. Population balance modeling was used to track the number density of different bubble class. The advent of high speed computing machine facilitates the use of computational fluid dynamic model in many engineering fields. Numerous CFD model of multiphase flow have been developed, and numerical data has been validated through comparison with experimental data. It has been successfully used in metal processing involving gas-liquid flow 4, 5, 6, 7, 8). The CFD solver employed uses the finite volume discretization method which rests on integral conservation statements applied to a general control volume. The Eulerian multiphase flow model solves the conservation of mass, momentum, and energy of each phase in each cell. The exchange of mass, momentum, and energy are dealt with the interphase exchange term in the conservation equation. The governing equations in this approach can be derived by ensemble averaging the fundamental conservation equations for each phase to describe the motion of liquid and gas in a bubble column. Both the continuous and the dispersed phases are modelled in the Eulerian frame of reference as interpenetrating continua. The momentum equations of the phases interact with each other through inter-phase momentum exchange terms. The numerical simulation was performed by a commercial CFD solver in AVL FIRE 2009.2 9), but the full foaming model comprising of equation of bubble break up, coalescence and foam formation and destruction were incorporated into AVL FIRE 2009.2 as subroutine written by the author in FORTRAN. The foaming index as a function of percentage of FeO in slag is presented in Figure 1. The predicted results from the present study are validated against the experimental data 10, 11) and was found to be in reasonable agreement with the available experimental data. Dimensionless analysis was performed based on the model available in the literature to correlate the foaming index with the physical properties of the slag and the coefficient of the foaming index was found to be 109 and 127 when the dimensionless group of Jiang and Fruehan 10) and Ito and Fruehan 11) was used respectively. The predicted results from the present study are in reasonable agreement with the experimental data.


The CFD model predicted the foaming height and the number density of different bubble class. The influences of different parameters on the foaming index and bubble number density are discussed in the presentation.

Figure 1 Foaming index of slag with different composition at 1773K.

References

1. H.,Luo, H.F., Svendsen, “Theoretical model for drop and bubble breakup in turbulent dispersions”. A.I.Ch.E. J. 42, 1996, p.1225–1233.

2. M.J., Prince, H.W., Blanch, “Bubble coalescence and break-up in air-sparged bubble columns” A.I.Ch.E. J. 36, 1990 p.1485–1499.

3. M. K. Tong, K., Cole, S.J. Neethling, “Drainage and stability of 2D foams: Foam behaviour in vertical Hele-Shaw cells” Colloids and Surfaces A: Physicochem. Eng. Aspects 382, 2011 p.42–49.

4. M., Alam, J., Naser, G., Brooks, “Computational Fluid Dynamics Modeling of Supersonic Coherent Jets for Electric Arc Furnace Steelmaking Process” Metallurgical and Materials Transactions B, 41b, 2010a p.1354 – 1367.

5. M.,Alam, J., Naser, G., Brooks, “Computational Fluid Dynamics Simulation of Supersonic Oxygen Jet Behavior at Steelmaking Temperature” Metallurgical and Materials Transactions B, 41b, 2010b p. 636 – 645.

6. M., Alam, G., Irons, G., Brooks, A., Fontana, J., Naser, “Inclined Jetting and Splashing in Electric Arc Furnace Steelmaking” ISIJ International, 51 (9), 2011 p.1439–1447.

7. M.,Alam, J., Naser, G., Brooks, A., Fontana, “Computational Fluid Dynamics Model of Shrouded Supersonic Jet Impingement on a Water Surface” ISIJ Int., 52( 6), 2012.

8. N., Huda, J., Naser, G., Brooks, M., Reuter, R., Matusewicz, “Computational Fluid Dynamic Modeling of Zinc Slag Fuming Process in Top-Submerged Lance Smelting Furnace” Metall. Mater. Trans. B, DOI: 10.1007/s11663-011-9558-6, 2011.

9. AVL FIRE v2009.2. AVL LIST GmbH, Graz, Austria. 10. K. Ito and R.J. Fruehan: Metall. Trans. B, 1989, vol. 20B, pp. 509-14. 11. R. Jiang and R.J. Fruehan, “Slag Foaming in Bath Smelting” Metall. Trans. B, 1991, vol.

22B, pp. 481-48.

0

1

2

3

0 5 10 15 20

Foam

ing

Inde

x ∑

(Sec

)

Percentage of FeO in Slag

Jiang and Fruehan,1991, CaO/SiO2=1.25" Jiang and Fruehan, 1991, CaO/SiO2=1.0"

Jiang and Fruehan, 1991,Large scale" Ito and Fruehan,1989

Present CFD


PRESENTATION – 33

Investigation of Phase Equilibria in the Cu-Fe-Si-Mg-O System at Low MgO Concentrations in Equilibrium with Copper Metal

Tijl Crivits, Evgueni Jak

PYROSEARCH, The University of Queensland, St Lucia, QLD 4072 Australia Keywords: phase equilibria, copper, silicate slag, freeze lining Introduction In pyrometallurgical processes where high temperatures and/or corrosive slag systems are used, excessive deterioration of the refractory lining is often a problem. An increasingly popular alternative is freeze lining. Previously, the interface between the bath and freeze lining was mostly assumed to be the primary phase at the liquidus temperature [1, 2]. Recent research, though, has demonstrated that this is not always the case [3]. The present research focuses on the determination of the effect of several slag and process parameters on this bath-freeze lining interface. Further study will be undertaken with a Cu-Fe-Si-O slag in an MgO crucible. This slag system is important in copper smelting, particularly in the “direct to blister” process. The MgO crucible was chosen to minimise the uncertainties introduced by the high solubility in slag of Al2O3 crucibles used in previous studies [3]. The liquidus in the multi-component Cu-Fe-Si-Mg-O system has not yet been investigated. The focus of the present stage of the study is to characterise the liquidus in this system at low MgO concentrations in equilibrium with copper at temperatures between 1100 and 1300°C. Procedure A high-temperature equilibration/quenching/electron-probe X-ray microanalysis (EPMA) technique was used to determine the composition of the phases in the system. Particular attention was paid to confirmation of the achievement of equilibrium and to the improvement of accuracy of results, and many steps in that direction were taken. For example, the homogeneity of the samples was investigated. First of all, several measurements were performed in different parts of the sample – near the gas-sample interface, in the middle of the sample and near the substrate-sample interface. Secondly, the composition was determined starting from the tip of a tridymite crystal, moving into the liquid. In both cases, no substantial differences in composition were observed Improvement of quenching Significant scatter of the liquid compositions was identified as a common issue in preliminary experiments. This scatter was attributed to an imperfect quench. Figures 1a and 1b illustrate examples: When the cooling speed of the liquid is too low, either small – submicron – copper particles (Figure 6a), or dendritic crystals with a high copper content (Figure 6b) are formed. To improve the quenching speed, several adjustments were made to the quartz (SiO2) substrate used to contain the slag during equilibration. For example, holes were drilled in the bottom to improve contact between the coolant (iced water) and the sample. The bottom of the crucible was also partially grinded off to reduce the amount of material being quenched. A well quenched area can often be found near the surface (Figure 6a: circle). Further active research is still under way to achieve adequate quenching in this system more frequently.


Figure 6 - Small, bright copper precipitates (a) or dendritic crystals (b) are observed in the liquid phase after quenching. The phases present are pyroxene (P), Tridymite (T), Liquid (L) and Copper (C). The well quenched area is marked by a circle. Results The boundary surface between tridymite (SiO2) and pyroxene (MgSiO3) at 1200 °C is projected onto the ‘Cu2O’-‘Fe2O3’-SiO2 plane (Figure 7) and the corresponding pseudo-ternary section with preliminary results is given in Figure 8. Present results for the 1200 °C pyroxene-tridymite boundary isotherm (Figure 8) are in agreement with the previously reported data [4]. It was found that addition of iron oxide into the system increases MgO solubility as well as the SiO2/‘Cu2O’ ratio in the liquid in equilibrium with tridymite, pyroxene and metallic copper.

Figure 7 - Sketch of the projection method used in Figure 8.

T

C L

C T

P

L

a b


Figure 8 - projection of the results on the 1200 °C isotherm on the boundary surface between tridymite and pyroxene (full line) and estimation of the 1200 °C isotherm on the pyroxene-olivine boundary surface (dashed line) A high iron oxide solubility in the pyroxene was observed. The stabilizing effect of MgO on the high silica liquid phase relative to the tridymite solid reported by Hidayat et al. [4] was also confirmed. Conclusions The preliminary results show the approximate position of the tridymite-pyroxene boundary isotherm at 1200°C. Further research is under way to characterise the liquidus in the tridymite and spinel primary phase field. Acknowledgements The authors would like to thank Umicore for the financial support for this research. References

1. M. Campforts, The formation of freeze linings: a microstructural perspective, in: KU Leuven, 2009.

2. K. Verscheure, High-Temperature Zinc Fumin in Water-Cooled Reactors - Thermodynamic Modeling and Experimental Inverstigation, in: KU Leuven, 2007.

3. A.F. Mehrjardi, P.C. Hayes, E. Jak, Met. Trans. B (2012). 4. T. Hidayat, H.M. Henao, P.C. Hayes, E. Jak, Metallurgical and Materials

Transactions B 43B (2012) 1290-1299

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