advanced technology for smelting mcarthur river ore

Minerals Engineering, Vol. 2, No. l, pp. 3-32, 1989 0892-6875/89 $3.00 +0.00 Printed in Great Britain © 1989 Pergamon Press plc

ADVANCED TECHNOLOGY FOR SMELTING McARTHUR RIVER ORE

N.A. WARNER

Department of Chemical Engineering, University of Birmingham Birmingham B15 2TT, U.K. (Received 6 October 1988)

ABSTRACT

Major research on a new polymetallic smelting process for treating complex sulphides is currently being sponsored by the Commission of the European Communities and the Mineral Industry Research Organization (MIR0) in the Chemical Engineering Department of Birmingham University. In the same laboratory, the Science and Engineering Research Council (SERC) are supporting a project on energy recovery from molten slag. This research on advanced smelting technology could pave the way for exploitation of the vast zinc-lead-silver deposit at McArthur River in the Northern Territory of Australia. The proposed process smelts ore directly to produce metallic zinc, lead and copper in the one furnace. Zinc and lead recoveries are predicted to be in the nineties and it is shown that with the adoption of innovative energy recovery technology, the thermal requirements can be satisfied by the inherent energy content of the ore itself. The electrical energy consumption for the total process from ore to metal is estimated and it is apparent that the gross energy requirement for zinc production from McArthur River ore by oxygen smelting is very considerably lower than that for current technology.

Keywords Direct smelting; McArthur River; complex sulphides; oxygen smelting

INTRODUCTION

The research on advanced smelting in the Chemical Engineering Department at Birmingham University covers two principal areas of activity, separately funded but clearly very closely related. The experimental programme began in 1982 with the award of a research grant from the Science and Engineering Research Council (SERC) to study energy recovery from slag using the technique of self-impinging jets to dry granulate molten slag and effect transfer of heat to a carrier gas. A second phase of the work, aimed at quenching partly solidified slag droplets in a fluidised bed, began in 1985, again with SERC support, and is still in progress. SERC has appointed the Mineral Industry Research Organization (MIRO) to monitor progress via an industrial steering committee.

The other major activity relates to direct smelting of complex sulphides. This work formally began towards the end of 1984 in a MIRO co-ordinated project with funding shared between five industrial companies (MIRO members), the British Technology Group (BTG) who are holders of the patent rights and the Department of Trade and Industry (Support for Innovation Scheme).

In January 1988, Phase 2 of the MIRO direct smelting project commenced with 2.5 year funding from the Commission of the European Communities via the Materials Programme and MIRO member companies presently comprising Billiton Research BV, BP Minerals International Limited and RTZ Metals Limited. In total, the financial contribution supporting the programme on advanced smelting technology since 1982 is in excess of one million pounds sterling.

In relation to the exploitation of the huge Zn-Pb-Ag deposit at McArthur River, it has been stated in the literature [1] that among other prime objectives "recent Mount Isa Mines (MIM) research has been directed towards

4 N A WARNER

processes which utilize the natural beneficial attributes of the ore to minimise large scale metallurgical processing on-site". This is completely the opposite viewpoint to that being advocated in this paper. The products leaving the mine site should ideally be high added value metals produced directly from the ore. Concerning the natural beneficial attributes of the ore, apparently this means different things to different people. Royston et al.[1] list five of these, all of which relate to either mineral beneficiation by flotation or hydrometallurgical leaching.

The natural beneficial attributes identified in this paper are:

I. McArthur River ore is high grade in terms of non-ferrous metals content.

2. It occurs in association with dolomitic material.

3. The sulphur content is more than enough to sulphidize associated iron oxide minerals.

4. The fuel value of the contained sulphides is comparable to that required for total ore fusion.

To make effective use of these natural beneficial attributes neither hydrometallurgy nor minerals beneficiation has a major role to play. What is required is the development of radically innovative smelting technology. Only by these means is it conceivable that the developers of the McArthur River orebody will be able to secure the competitive advantage that will be needed to ensure that they become low-cost high-volume market leaders.

MIM's McARTHUR RIVER PROJECT

A comprehensive review of the project was presented in Darwin by Buchanan [2] at the 1984 Annual Conference of the Aus.I.M.M. Some years earlier in 1979 MIM had submitted a major feasibility and environmental report to the N.T. government. This report was studied by external consultants appointed by the government, who agreed with the conclusions that the McArthur River HYC deposit was uneconomic.

The principal minerals in the deposit, sphalerite and galena, are finely laminated, extremely fine grained and intimately intergrown with non-economic minerals such as dolomite and pyrite [3]. McArthur River probably ranks as one of the world's largest undeveloped, high grade near surface sulphide orebodies with reserves estimated at 227 million tonnes with an average grade of 9.2% zinc, 4.1% lead and 41 g/tonne silver.

The use of conventional grinding - flotation processes was studied in a 50 tonne per day pilot plant during 1977 and 1978. Results of the pilot plant operation indicated that an extremely fine grind would be necessary with a comparative power usage for the concentrator of 45 kWh per tonne for McArthur compared with 21 kWh per tonne for the Mount Isa concentrator. Low grade bulk concentrates were produced initially but effort was then switched to the production of separate zinc and lead concentrates. The ultra-fine grained nature of the ore required grinding to an average size of 95% passing 37 microns or 400 Tyler mesh. The final development of the pilot flowsheet produced lead concentrates of 54.6 percent grade at 23.4 percent recovery and zinc concentrates of 51.9 percent grade at 67.8 percent recovery [2].

The feasibility study already referred to accordingly concluded that the production of separate concentrates was not an economic reality under the then current conditions. It also recommended that future research should be redirected towards hydrometallurgy.

DIRECT ORE SMELTING

Pyrite and non-ferrous metal sulphides are normally separated from gangue minerals by fine milling followed by froth flotation. Further separation of the non-ferrous sulphides by differential flotation or the production of a bulk flotation concentrate is conventional practice. In some cases, however, there would appear to be a real incentive to avoid the energy-intensive process of fine milling and, instead, there may well be a case for feeding certain minerals directly at relative coarse size to a pyrometallurgical

Advanced technology for smelting McArthur River ore 5

process such as that now being proposed. This is particularly true for finely disseminated sulphides, where even the production of a bulk concentrate can be associated with relatively poor metals recovery.

Compelling reasons for considering the direct ore smelting option include the desire to:

a) increase the overall metals recovery beyond established practice b) secure benefits of polymetallic production c) fix deleterious elements in a non-reactive slag d) reduce the environmental impact of process residues. Disposal of a fused

ore slag is straightforward and may even generate a small profit e) overcome variations in input ore characteristics by incorporation of

downstream processing unaffected by the initial physical structure of the ore.

However, direct ore smelting without energy recovery from the fused ore slag demands excessive energy input. The requirement is decreased as process intermediates or pre-concentrates are considered but this results in a decrease of overall recovery.

Large amounts of energy are consumed in melting down gangue material but dry granulation of the molten ore slag can transfer the slag energy into a carrier gas thereby allowing sensible heat to be passed back to the front-end of the process in an ore preheater. Under these conditions only the thermal losses have to be added and the energy demand is then more modest. In fact, when one considers the inefficiency of thermal power generation, the energy intensity of fine comminution and beneficiation procedures, the energy requirements of direct ore smelting can be competitive with conventional mineral processing. This is particularly the case for ores containing sulphides.

The key requirement is the development of an effective means of recovering thermal energy from slag. The current status of the worldwide effort in this direction has been reviewed by Broadbent and Warner [4]. There are reasonable grounds for optimism that new technology will emerge in the not too distant future.

ENERGY RECOVERY FROM SLAG

Slag production associated with several pyrometallurgical operations is listed in Table I. The slag make in an iron blast furnace is very low in line with the requirement of high intensity ironmaking units requiring high quality charge. This does not mean that energy recovery from iron and steelmaking slags has a lesser role to play than in the other processes listed in Table I but, on the contrary, the very large quantities of iron produced more than justifies worldwide effort in this direction.

TABLE 1 Slag Production per Unit of Metal in Product

Slag/Product Metal Process Product Mass Ratio

Iron blast-furnace Zinc-lead blast-furnace Inco oxygen flash furnace Noranda reactor Outokumpu flash furnace (WMC) Outokumpu flash furnace (BCL) Laterite electric furnace (Exmibal, Guatemala) Laterite electric furnace (Falconbridge Dominicana)

Hot metal 0.3 Zinc and lead metal 0.7 47% Cu matte 1.4 Blister copper 3.5 45% Ni matte 8.5 32% (Ni + Cu) matte 15.5 25% Ni matte 43

32-40% Ni Ferronickel 56

A typical slag tapped at say 1400oc retains a heat content around 1.8GJ/t and although it is desirable to recover this thermal energy, it is also necessary

6 N . A . WARNER

to dispose of the slag at minimum cost. Traditionally, slag is discarded with little or no attempt to recover the thermal energy that is contained within it and water granulation is often the preferred means of disposal. Several dry granulation processes are reported to be in the pilot plant stage in Japan. These include atomization by means of rotating drums and high speed air jets. An alternative approach involving impinging jets of slag for disintegration prior to energy recovery has been studied by Broadbent and Warner [5, 6].

Work on the novel impinging jet technique is continuing in our laboratory and a further phase of the project is now attempting to study the partial quenching of slag droplets in a high temperature fluidized bed.

The test rig now being constructed embodies the following essential features:

a) an induction heated graphite crucible and stopper rod/impinging jet assembly inside a cylindrical stainless vessel to contain the nitrogen protective atmosphere.

b) a 2m diameter free fall chamber with an uninterrupted drop of about 4m beneath the point of impingement.

c) a fluidized bed of crushed slag granules, so positioned that the major portion of the disintegrated slag stream is captured after the initial free fall period.

The ancillary equipment required for the test rig includes a closed loop nitrogen circuit with a Roots-type positive displacement blower rated at 700m~/h, a graphite resistor element preheater capable of heating the fluidizing inlet gas to a temperature around 1200oc, a cyclone separator designed to permit bed recirculation at high space velocities and entrainment ratios and two convective gas coolers. The free-fall chamber is shown in Figure I. This unit is now installed above a 2m diameter by 3m deep sump which contains the fluid bed and gas preheater and the whole assembly is positioned immediately adjacent to the smelting furnace. The sump can be seen in the foreground in the view of the furnace in Figure 5.

Fig.1 Free-fall chamber for slag energy recovery


POLYMETALLIC SULPHIDE SMELTING

Polymetallic smelting is not a new concept. It is established practice in the copper-nickel industry for producing matte from bulk Cu-Ni-Co sulphide concentrates and it is now over thirty years since the first zinc-lead blast furnace was commissioned at Avonmouth in the United Kingdom. More recently the KIVCET process has become available for smelting complex Cu-Zn-Pb sulphides. The thrust in all these developments is the recognition that a smelter dedicated to a single metal is not necessarily the best approach to resource recovery.

The metallurgical features that distinguish the process being developed at Birmingham University from others are:

(1) The process is fully continuous and uses rapid reactions to simultaneously produce copper, zinc and lead in the one reactor.

(2) Zinc metal is produced directly from zinc sulphide without going through the oxide phase. This feature is unique among smelting processes and reduces substantially the amount of purchased energy required for zinc metal production and, particularly, eliminates the need for energy in the expensive forms (i.e.D.C. electrical power or high grade metallurgical coke) which existing processes require for oxide reduction.

(3) The need for submerged tuyeres or lances is eliminated. The oxygen requirements are provided by a number of top blow lances operating in the non-splash mode.

(4) Some in-situ slag cleaning would appear possible because it is present as a moving layer only a few millimetres thick, thus ensuring good contact between reductant and oxidised slag components. Subsequent retention of the slag in the main circuit with matte flowing underneath should facilitate matte slag separation and contribute further to in-situ slag cleaning.

(5) Transfer of heat from fuel combustion (e.g. natural gas) occurs across a relatively clean matte surface enabling efficient and highly intensive heat transfer to take place. In other matte- slag processes heat has to be transferred to matte through a layer of 10 cm or more of slag.

FORCED MATTE CIRCULATION

Basic Process Concept

Forcibly circulating a melt within a reactor overcomes limitations inherent in conventional pyrometallurgical furnaces such as back-mixlng and non- countercurrent contacting. Also, if closed loop circulation is promoted either within a single partitioned furnace or between slde-by- side furnaces, processing can be conducted in a sequential manner within the loop and the melt can traverse both oxidising and reducing conditions to effect the pre-requisite chemistry essential for multi- metal production with a single reactor system. Circulation also establishes a mechanism for direct recovery of heat from exothermic slagging and converting reactions as well as from fuel combustion and its transference as sensible heat to sites of endothermic chemical reaction, charge melting, volatilisation and slag cleaning.

In principle the precise means of inducing circulating is immaterial. In the context of matte smelting, however, an adaptation of technology already well established in steelmaking appears to offer many advantages. Of particular interest is the RH vacuum degassing system shown in Figure 2. An RH degasslng vessel has two legs or snorkels, the molten steel being carried up one leg with the assistance of argon injection to that leg and flowing out into the refractory lined vacuum vessel and then out down the other leg, back into a ladle positioned beneath the RH vessel. The circulation rate of steel from the ladle to the degassing vessel is normally around 45 to 60 tonne per minute.

For continuous matte smelting, the batch ladle of steelmaking would be replaced either by a metallurgical reactor with internal partitions to direct matte flow or more probably the matte smelting circuit would comprise two

8 N.A. WARNER

side-by-side reverberatory type furnaces linked in closed circuit with each other. A schematic plan view of such an arrangement is shown in Figure 3.

AUoy Additions

Inert

Vocuum

Vessel

To_Vcuum !Pumps

Dust Trap

Ladle

Fig.2 Essential features of RH vacuum degassing

MATTE CIRCULATION

r

°22 Fig.3 RH arrangement for matte circulation between two furnaces

Depending on the material being smelted, the effectiveness of the vacuum in removing zinc and lead, and the mass transfer intensity in the top blow region, steady state and self regulating conditions will tend to be established within the matte circuit.

The matte discharge from the RH reactor is distributed in the bath of the second reverberatory-type furnace and flows countercurrent to hot furnace gases resulting from sulphide oxidation, combustion of excess oxygen from the top-blowing region and additional combustion resulting from pulverised coal and oxygen enriched air admitted above the bath surface. High intensity heat transfer conditions are maintained between the clean turbulent matte surface and the furnace enclosure for at least half the length of the bath extending up to the top-blowing region, where oxygen is blown onto the top of the matte. The slag generated in this region is removed continuously as a thin surface layer and overflows along with matte out of the main furnace into a small forehearth in which matte separation and slag accumulation is allowed to occur. The matte overflows continuously from this forehearth back again to the first of the two principal furnaces.

Crushed ore, say -3 cm, or pelletized concentrates and dross reverts would be pre-heated by direct countercurrent contact with hot smelter gases in vertical furnaces which discharge preheated solids at around 1200K directly into the


circulating matte. The general features envisaged are shown in Figure 4 and a photograph of the experimental furnace taken during construction in the MIRO Phase I programme is given in Figure 5.

VaC:UUITI

Fig.4 Schematic representation of proposed plant for direct smelting of zinc-lead ore

DIRECT SULPHIDE SMELTING

There is an ambiguity in the expression 'direct smelting': direct smelting could imply metal production directly from ore without extensive mineral beneficiation, but its more important meaning is the production of metal directly from a sulphide without reliance on intermediate chemical reactions. The following chemical reactions are involved in direct smelting:

<ZnS> + 02 = <Zn> + SO 2

AH298 = -22.75 kcal

<PbS> + 02 = <Pb> + SO 2

AH298 = -48.15 kcal

<Cu2S> + 02 = 2<Cu> + S02

AH298 = -51.35 kcal

In all three instances, metal is produced directly from a sulphide, whether it is a sulphide concentrate or a sulphide ore mineral, without an intervening

10 N.A. WARNER

oxide phase. The ZnS, PbS and Cu2S reactions are strongly exothermic and the heat of reaction increases in the order shown. That the reactions are exothermic is obviously important from an energy viewpoint and it is also worth noting, the requirement for a carbonaceous reducing reagent is eliminated. Table 2 focuses attention on zinc sulphide and compares the chemistry of the established technology with that of the reactions that are proposed for direct oxygen smelting. The carbothermic reduction of the oxide produced in reaction 1(a) is the primary reaction of the zinc-lead blast furnace. The direct smelting route has certain thermodynamic limitations. If the temperature is sufficient, significant partial pressures of zinc gas and sulphur dioxide can coexist in stable equilibrium, but the problem is how to recover the zinc. Once the temperature is lowered in an effort to condense liquid or solid zinc, the reaction proceeds from right to left and zinc reverts to a mixture of ZnS and ZnO. That thermodynamic limitation can be overcome through two reactions in which copper and cuprous sulphide take part: reactions involved in 2(a) and 2(b) can be made to occur spontaneously in different regions of a reactor system, the net result being the direct smelting reaction.

Fig.5 General view of MIRO smelting furnace during construction

For complex sulphide feeds containing similar proportions of copper, zinc and lead, direct smelting in the oxidising region under the top blow air/oxygen lances produces blister copper which collects in the furnace hearth. The solid charge is assimilated into the circulating matte under neutral or reducing conditions and then zinc and lead are produced by reactions (I) and (2) as the copper-saturated matte is exposed to the vacuum inside the RH vessel.

ZnS + 2Cu = Cu2S + Zn(g) (I)

PbS + 2Cu = Cu2S + Pb(g) (2)

Elemental zinc and lead together with some lead sulphide are flashed off, but as the gases cool from around 1200oc in passing to a zinc condenser (probably


direct liquid zinc contact but possibly with molten lead), a lead sulphide mist or fog can be expected to form and equilibrium changes represented by reactions (3) and (4) can be anticipated, resulting in the formation of a sulphide/oxide dross in the condenser sump. This dross would be reverted but clearly it is imperative to limit the extent of SO 2 evolution within the vacuum vessel.

PbS(£,s ) + Zn(g) = ZnS(s ) + Pb(£) (3)

SO 2 + 3Zn(g) = 2ZnO + ZnS (4)

Both zinc and lead are preferably collected in molten zinc or Pb- Zn alloy, subsequent cooling of which yielding zinc 4 and zinc lead bullion. Alternatively, it may be desirable to produce directly a refined zinc product with a lead content meeting Grade I specification as outlined in Figure 6. This involves vacuum re-evaporation from the circulating liquid zinc quench alloy without further energy input, but some decadmiumising would still be required to meet high grade purity standards.

TABLE 2 Zlnc Smelting Reactions

I Established technology

(a) Roast reaction (gas-solid reactor; fluid-bed or sinter roast)

3 ZnS + - 0 2 = ZnO + SO 2

2

(b) Carbothermic reduction

ZnO + I/2 C = Zn + I/2 CO

(c) Overall reaction

3 ZnS + I/2 C + - 0 2 = Zn + S0 2 + I/2 CO

2

2 Proposed direct smelting route

(a) ZnS + 2Cu = Cu2S + Zn

(b) Cu2S + 0 2 = 2Cu + SO 2

ZnS + 0 2 = Zn + SO 2

To enable reactions (I) and (2) to proceed it may be necessary in some cases to add metallic copper to the vacuum vessel, but normally the matte will be circulating at a rate sufficient for the mutual solubility of copper and copper sulphide to provide the requirement with no external movement of copper whatsoever.

Effectively, copper matte is a solvent for the other sulphides being smelted. Adding these to matte under non-oxidising conditions implies raising the solids to bath temperature followed then either by solid/liquld dissolution or melting. With both processes being strongly endothermic it is essential that efficient heat transfer conditions are established and again forced matte circulation is the key requirement not only to supply the heat but also to ensure rapid heat and mass transport by forced convection.

Both charge assimilation and the reactions occurring in the vacuum vessel are intensively energy consuming. Exothermic reactions however, also form part of the direct smelting sequence: the reaction in which cuprous sulphide is converted to copper is exothermic and any iron sulphide in the feed to the smelter has to be oxidised to form a slag, slagging reactions being strongly exothermic. The energy released in these reactions is transferred to the

12 N . A . WARNER

sites of the endothermic processes by means of the sensible heat contained in the circulating matte stream, so the minimum circulation rate needed is dictated by the thermal requirements and the temperature range over which matte can reasonably be handled.

To Smelter =

Nitrogen --

To Vacuum I Zinc Pumping System <] J Condenser

Dross extraction I @' Zinc Splash ~ Zinc Condenser J-JRe-evapom~

I

J Separator Bath l

Lead (Zn]

.quid zinc :ircuit

_,.Refined Zinc

l iquid zinc ircuif

Fig.6 Flowsheet for refined zlnc production

VACUUM REMOVAL OF ZINC FROM MATTE

An indication of the maximum rate of zinc elimination from matte can be obtained from published data on carbon removal in the normal batch RH degassing of steel, assuming that this cannot exceed the liquid phase transport rate. For example, it has been reported [7] that with an iron circulation rate of 45t/min on a 120t heat of plain carbon steel it takes 9.3 minutes for the carbon level to drop from 0.09% to 0.01%. This includes a 3 minute pump-down time so the effective degassing time is 6.3 minutes.

For liquid phase diffusion control in batch degassing it can be shown that the relationship between the initial solute concentration C o and the concentration C at time t is given by:

In (Co/C) = (DA/6V)t (5)

where A = area of the gas-metal interface V = volume of the melt 6 = effective thickness of the diffusion boundary layer D = Diffusion coefficient

Now for continuous vacuum dezincing of matte under physical conditions similar to RH steel degassing, the value of (DA/6) evaluated from the above equation can be used as a first approximation for the maximum zinc distillation rate nzn as given by

nzn = (DA/~)(Czn - CZn*) (6)

The concentration value CZn in the above equation will be the bulk outlet value if complete backmixing is assumed to occur whilst Czn, refers to the surface concentration. At very large matte/zinc circulation ratios, the change in concentration of zinc in passing through the vacuum chamber is relatively small and for say 360 mol matte per one mol zinc it corresponds to a concentration change of only 0.14 percent.


The absolute value of the zinc concentration in the recirculatlng matte will find its own steady value. It may be, for example, that matte leaves the RH vacuum vessel containing 0.5 percent zinc as dictated by the overall zinc mass balance and the kinetic conditions existing elsewhere in the matte circuit.

Combining equations (5) and (6) gives

KV In(Co/C )

(Czn - Czn ,) nzn - t (7)

The term K is a factor taking into account that flowrates of matte are considerably greater than the 45t/min value for steel degassing in the example quoted. At best the same degree of stream break-up and general turbulent activity will occur for each volume element of melt exposed to the vacuum. Thus if the volumetric matte flow is equivalent to a steel flow of 200t/min, the factor K is simply (200/45) and then by substitution in the equations it can be shown that to distil 10t/h zinc, the concentration difference required is 0.12 weight percent, giving a surface concentration of (0.50-0.12) = 0.38 weight percent.

From published thermodynamic data [8,9] the calculated equilibrium constant for equation (I) at 1573K is 2.013, the activity coefficient of zinc sulphide in matte is 5.4 [12] so the zinc partial pressure in equilibrium with the surface where acu = 0.9 and acu2s = I can be shown to be about 62 mm Hg. The significance of the above partial pressure value is that vacuum dezincing of matte is kinetically much more favourable than vacuum dezincing of lead (VDZ). When continuous VDZ was operating as a commercial process, the interfacial partial pressure of the zinc could never have exceeded about 5 mm Hg and indeed, kinetic analysis of the VDZ process [10,11] indicates that the actual zinc partial pressure in the distillation gap between evaporating surface and condenser was more in the region of I or 2 mm Hg.

It can be argued, of course, the vacuum dezincing of matte in an RH type reactor utilizing an inert gas as the lift gas will not produce the vigorous conditions associated with normal steel degassing. Nevertheless, there does seem to be plenty of scope for reducing the mass transfer intensity whilst still retaining realistic zinc partial pressures to effect distillation.

Evaporative Distillation for Lead Recovery

If the matte is at a temperature of 1623K (1350oc), for example, and say the steady level of lead in this matte is 2 wt%, then for almost copper saturated conditions (say a Cu = 0.95), the equilibrium partial pressures of elemental lead and lead sulphide, based on published free energy data [13] and activity coefficient reported by Sohn et al. [14] are PPb = 3.2 mm Hg and PPbS = 1.5 mm Ng. On the basis of these figures the evaporative distillation option indicates that for every mole of total lead evaporated 0.32 mol would have to be reverted back as lead sulphide, unless zinc vapour is present in sufficient quantity to react according to equation (3). Enough zinc is certainly present for the normal Zn/Pb ratios employed in the zlnc-lead blast furnace but with higher lead levels reverting such a large fraction of the metal production may be unacceptable and an alternative strategy based on liquid-llquid extraction concepts has been devised [15]. For McArthur River ore smelting the evaporative route would be applicable.

IRON SLAGGING AND CONVERTING

Normal submerged tuyere operation in the iron slagging region of the matte circuit would undoubtedly produce excessive zinc and lead losses to slag, but calculations suggest that the top blow configuration with the cellular arrangement may overcome the problem. Provided the oxygen lances are operating in the non-splash mode, the maximum loss of zinc and lead to slag cannot exceed the rate of liquid phase transport of these species from the bulk of the matte to the matte surface. What is required is a method of enhancing the gas phase mass transfer process to supply oxygen for the reaction stoichiometry whilst at the same time inhibiting liquid diffusion. This is one of the main reasons for proposing a pure oxygen top blow at present but clearly there could well be an optimum degree of oxygen enrichment in practice. If it is assumed that the iron slagging and similar metal oxidation reactions proceed at a rate controlled mainly by liquid phase

14 N . A . WARNER

diffusion, estimates of the local mass transfer coefficients can be made which then allows concentration driving forces required to satisfy metal balances to be established. This in turn allows one to estimate the steady state composition of the circulating matte.

In the limit of exclusive liquid phase mass transfer control, the effective concentration of iron sulphide immediately under the jets can be presumed zero and the surface would then be white metal (Cu2S) undergoing the normal converting reaction to copper, whilst at the same time iron sulphide is being oxidised to slag. A complex series of kinetic interactions are likely to occur, the copper formed on the surface eventually sinking through the bulk matte with its iron content dictated by its liquid phase transport requirement.

A pool of copper would become established under the top blow region and the circulating matte would tend towards copper saturation by virtue of the high matte to zinc circulation ratio. Depending on the copper content of the charge and the need to recover silver and other precious metals, some copper may need to be periodically tapped from the system. The amount of copper being taken into solution by the circulating matte would be self adjusting to replace that consumed by the zinc and lead reactions in the vacuum vessel, but this copper is eventually regenerated by the converting reaction. For a low copper feed some copper must be introduced into the system, probably as a low grade sulphide concentrate to compensate for copper losses in the slag.

MATTE DE-OXIDATION

For oxygen blowing under probably gas diffusion control, the surface layers of matte in the immediate impact area under the top blow lances can be expected to approach equilibrium in terms of dissolved oxygen, corresponding to about 1.0 wt percent oxygen right on the surface. The matte can thus be expected to pick up dissolved oxygen corresponding to the liquid phase diffusion rate. A major portion of this dissolved oxygen has to be removed before the matte is exposed to the reduced pressure inside the RH vacuum vessel because every mol of SO 2 released here has the potential of drossing 3 mol zinc by oxidation and sulphidisation.

In theory immediately after the oxygen diffuses into the bulk matte the following reaction should occur with &G o of around - 30 kcal/mol.

Cu20 + FeS = Cu2S + FeO (8)

Assuming that the above reaction can proceed homogeneously and that the participating chemical species are all in solution in the matte phase, an excess of FeS in the bulk melt will push the reaction from left to right, effectively increasing the activity of FeO, aFeo, until a steady state level is reached in the recirculating system. The ultimate level of aFe O is thus determined by other rate processes in the circuit. For example, if the solid charge is exposed to gaseous sulphidising conditions in the preheater before it is assimilated into the bath so that very little iron oxide originally present in the ore remains unconverted to ferrous sulphide by the time the charge is melted, the fused ore slag so produced will be characterised by a very low iron oxide activity and as such is capable of absorbing FeO from the matte into the slag. Under these conditions equilibration of the circulating matte and the fused ore slag can be seen to offer an effective means for achieving matte deoxidation.

There is also another important factor to be taken into account in considering the permissible steady state level of aFe O. Reaction can occur between ZnS and FeO both dissolved in the matte resulting in loss of zinc to slag as represented by the following equation.

ZnS + FeO = ZnO + FeS (9)

At 1700K the equilibrium constant for this reaction is about 0.23 and again if circulating matte and fused ore slag are equilibrated, a knowledge of the activity coefficients of the various dissolved species both in the matte and slag phases permits evaluation of the likely level of zinc loss under specified operating conditions.


Besides matte/fused ore slag equilibration, matte can also be deoxidised by chemical reaction with appropriate reductants such as coal char, coke or natural gas. The approach to be used will clearly be dictated by the particular circumstances.

There is also the possibility that appropriate design of the zinc condenser will permit zinc recovery at high recovery even in the presence of substantial amounts of SO2, making matte deoxidation of academic interest only.

IN-SITU SLAG CLEANING

Rottmann and Wuth [16] have demonstrated that the force of a non-splash mode top blowing jet is sufficient to keep the depression or impact zone free of slag in a static crucible experiment. Because forced matte circulation continuously removes slag, it would seem reasonable, therefore, to envisage the slagging region to consist of a network of surface depressions all essentially free of slag with direct contact between the oxidizing gas and the matte.

Forced matte circulation coupled with top blowing at an intensity just below the splashing mode, should permit adequate slag penetration without excessive over oxidation and magnetite formation. The natural circulation induced in the matte immediately below the area of impact removes the slag as it is formed away from the region of highest temperature and oxygen potential, whilst the bulk motion of the circulating matte ensures the continuing removal of slag away from the converting region. Once the slag is moving away in a thin layer it can then be contacted with solid reductants to effect magnetite reduction and reversion of copper, from slag back into the matte phase.

Slag cleaning is not difficult because of the chemistry involved but rather getting all the slag and reductant together in close physical association is made difficult by the transport properties of slag. Contacting a thin continuous layer of slag with reductant as proposed in the process should be ideal for conducting the chemistry of slag cleaning in a rapid manner. Subsequent settling in the slag compartment is then a question of providing residence time after the chemistry is completed. In-situ slag cleaning splits the overall operation logically into a sequence not unlike the familiar mixer- settler concept of continuous solvent extraction.

SMELTING INTENSITY

Themelis and Kellogg [17] distinguish two basic types of modern smelting process: flash smelting and bath smelting. The process in this paper is bath smelting. However, it is fundamentally different from the others in that it does not rely on injection of an oxidising gas into the melt but rather the turbulence required to sustain effective heat and mass transfer is established by forced convection resulting from matte circulation. The smelting intensity is thus very much under the control of the process designer/operator, probably even more so that in comparable technology.

Assuming that the charge solids after preheating are fully supported by the liquid bath, the clearance between the underside of the charge column and the floor of the furnace hearth is predictable from the well established theory and practice of silos. By appropriate engineering design this clearance can in principle be varied whilst operating as dictated by the particular smelting requirements. Similarly, the matte circulation rate can be varied by control of the amount of lift gas admitted to the RH upleg.

The rate at which solids are assimilated into the melt reflects smelting intensity for a given physical size of plant. This in turn is determined principally by rates of heat and mass transfer between the solid and liquid phases. For the configuration under discussion, the appropriate dimensionless correlations for fully developed turbulent flow are

Nu = 0.023 Re 0-8 Pr 0-33 (lO)

Sh = 0.023 Re 0-8 Sc 0-33 (11)

16 N . A . WARNER

where Nu = Nusselt Number = hde/k Sh = Sherwood Number = kcde/D Re Reynolds Number deVP/~ Sc = Schmidt Number = u/pD

and Cp = specific heat of liquid d e = flow channel equivalent diameter D = interdiffusivity in liquid h = heat transfer coefficient k = thermal conductivity of liquid k c = mass transfer coefficient

Beneath the charge preheater the free passage of matte is constrained and all the circulation is forced to flow under the charge through a channel or conduit with an equivalent diameter d e equal to four times the ratio of cross section area to wetted perimeter. To a first approximation, the transport processes can be regarded as taking place at a planar solid/liquid interface corresponding to the top surface of this conduit, thereby allowing calculation of smelting intensity based on physical properties either available in the literature or estimated. However, to ensure that the system is actually controlled by mass transfer in the liquid phase, it is desirable to conduct a simple laboratory test.

For example, the initial phase of the MIRO project was concerned with the production of zinc from high grade concentrate. Single pre- sintered pellets were immersed in copper-saturated cuprous sulphide under isothermal static conditions. At 1200oc an immersion of 20 seconds resulted in a diameter change from 1.8 cm to 1.0 cm, whilst after 60s no pellet remained. The time for complete dissolution was thus somewhere between 20s and 60s. Mass transfer is presumed to take place by natural convection in a similar fashion to dissolution of magnetite in matte as reported by Kaiura and Toguri [18]. Using procedures given in standard chemical engineering texts, and noting that interphase mass transfer is taking place both by diffusion and a bulk flow component, the observed dissolution rate can be shown to be consisted with the data listed in Table 3 and the Ranz-Marshall equation for single spheres.

Sh = 2.0 ÷ 0.6 Gr0.25Sc0-33 (12)

where Gr = Grashof Number = gd 3 (Pi-Pb)/92pi

and d = sphere diameter Pi = liquid density at solid interface Pb = liquid density in bulk

= average kinematic viscosity

TABLE 3 Mass Transfer Parameters for ZnS

Dissolution in Matte at 1200oc

sintered pellet density

Matte viscosity

Bulk matte density

Matte Interfacial composition

Matte Interfacial density

Interdiffusivity DZnS_Cu S 2

2.84g/cm 3

10 centipoise

5.2g/cm 3

56 wt percent ZnS

4.4g/cm 3

8 x 10-5cm2/s


Consider now a full-scale plant with a pelletised charge preheater configuration as shown schematically in cross-section in Figure 7. The solids are floating on the matte, rather like an iceberg in the ocean. There is a pool of copper to protect the refractories from erosion and for the same reason the side-walls are shielded from the scouring action for the rapid matte flow by the solid charge. To ensure free movement of the solid charge column into the melt, a condition unlikely to be maintained if solids adhere to the side walls, high thermal conductivity side walls would be required in the preheater vicinity to prevent matte freeze-up impeding passage of solids into the bath.

I

ill i;:

Fig.7

~ c o m p o s i t e o r e c h a r g e

- ~ M a l n G a s O f f t a K e

~ ~ - ~ - - S e c o n o a r y A l r

M a t t e F low C h a n n e l s

Schematic cross-section of proposed furnace showing solid charge profile and matte flow passages

TREATMENT OF McARTHUR RIVER ORE

Laboratory Tests on Ore Fusion

Some preliminary tests on a consignment of McArthur River ore were undertaken to measure non-ferrous metal elimination from the ore when it is fused by being submerged in a bath of molten copper sulphide matte. In each experiment five pieces of lump ore as received (typically 2-3 cm) were thermally pretreated in a horizontal tube furnace by slow heating up to 1050oc in an atmosphere that simulated process conditions with some elemental sulphur gas as well as carbon dioxide and nitrogen. The material was then allowed to cool to room temperature under nitrogen before being plunged beneath the surface of molten matte in a Salamander clay - graphite crucible. Typically, an original ore mass of 30 g and 1200 g matte were used in each experiment. The ore pieces were held beneath the surface with a graphite push rod for 20 min before the rod was lifted to allow the slag or clinker to separate on the surface. Table 4 shows the very marked effect of temperature on metal elimination. The 1400oc data refer to one experiment, different portions of the slag mass being subjected to analysis. The results are indicative of potential very high zinc and lead elimination from fused ore.

The slag that resulted from the 1300oc immersion test was examined by electron microscopy. SEM/EDAX revealed spherical matte particles rich in S, Fe, Cu, Mn and Zn dispersed throughout a slag matrix that itself was almost entirely free of Cu and Zn and also contained very little iron. With appropriate modification of slag viscosity and provision of retention time in the slag pool it seems reasonable to assume that the matte particles, typically 300-750 um in diameter, would settle out.

ME 2/I--B

18 N . A . W~NER

TABLE 4 MR ore dissolution experiments

Temperature % oc Fe Pb Zn S

1 200 Clinker n.a. 0.38 7.78 n.a. 1300 Slag 4.45 0.41 I .41 n.a. 1400 Slag (a) 1.03 0.12 0.08 0.34

Slag (b) 0.14 0.09 0.06 0.44

If the laboratory performance can be repeated in practice, direct MR ore smelting should yield a fused ore slag with only about 0.6% of the input zinc. Provided that all drosses and fumes are reverted to the matte circuit, the only other source of metal loss is the slag produced in iron slagging and copper conversion.

Blended Composite Ore Feed

After allowance for dilution on mining, the ore to be treated by a concentrator, for example, would have the following typical analysis in weight percent: 32 SiO2; 6.5 A1203; 4.5 Mg0; 6.5 Ca0; 10.9 Fe; 3.0 K20; 0.1 Na20; 4.0 Pb; 9.0 Zn; 0.2 Cu; 14.0 S; 0.7 C; 9.5 C02; 0.02 Cd. For direct ore smelting, however, ore of this composition would probably need blending with a flux material to avoid the formation of highly viscous ore slag or alternatively abnormally high smelting temperatures would be required.

One of the key natural beneficial attributes of MR ore is that it occurs in association with layers of dolomitic sediments, a material ideally suited for fluxing the MR ore deposit itself.

For the purposes of initial calculations, it is assumed that either by consciously changing the mining pattern or by bringing in dolomitic material from nearby, blended composite ore of the following analysis could be made available for a direct ore smelter: 25.9 Si02; 5.2 A1203; 7.3 Mg0; 12.3 CaO; 2.3 K20; 16.1 CO2; 0.5 C; 10.6 S; 9.8 Fe; 6.8 Zn; 3.0 Pb.

The calculated viscosity of the fused ore slag produced from this blended composite is 14 poise at 1400oc and 9 poise at 1450oc. In effect, the equivalent of 15g (CaO + MgO) for 100 g MR ore has produced a composite feed with about the right fluidity for ore smelting at a reasonable temperature level. There is of course an energy penalty in bringing additional carbonates to be thermally decomposed, but as subsequent calculations show this really is not a major a problem.

To avoid confusion in subsequent discussion throughout this paper, when reference is made to ore production rates and energy requirements, these are related back to the MR ore analysis first given, although it is tacitly assumed that,in fact, an ore smelting operation would probably need a blended composite ore feed, perhaps approximating that given.

Sulphur Disposal Considerations

Run-of-mine MR ore contains 13 to 15% sulphide sulphur and given today's price of sulphur as approximately US$140/t and the availability in the Northern Territory of relatively inexpensive natural gas, serious consideration must be given to elemental sulphur production. In fact, the value of sulphur in the ore at the current sulphur price represents about one eighth of its total contained value, so on this basis a pyrometallurgical process producing initially a high strength continuous stream of sulphur dioxide is a prime candidate for natural gas reduction. Allied Chemical Corporation technology commercially demonstrated at the 500 tpd level based on SO 2 from roaster gas at Falconbridge's nickel operation in Canada would appear directly applicable to McArthur River, where a production capacity of 1400 tpd would be required.

Whereas elemental sulphur can be shipped across the world, finding readily accessible markets for the equivalent amount of sulphuric acid produced at McArthur River is at present not a realistic option.


Leaving sulphur in ore residues either as elemental sulphur or sulphide sulphur is favoured by preponents of hydrometallurgical processing but on strict environmental grounds this is clearly unacceptable. The long term environmental consequences of this practice would probably be worse than the comparable problem now acknowledged with jarosite residues, for example, in zinc hydrometallurgy.

In total energy conservation terms, fixation of sulphur as anhydrite by high temperature reaction of SO 2 and oxygen with limestone or dolomite and the subsequent recovery of the exothermic heat of reaction has a great deal to offer. The anhydrite is formed at a high temperature well above the stability temperatures of other deleterious nonferrous sulphates when in the presence of excess lime so there is no question of adventitious leaching of harmful metals from the calcine when this material is dumped.

Even the most stable of the basic lead sulphates PbSO4.4PbO is unstable in the presence of excess CaO under the conditions envisaged for MR ore treatment. Given the opportunity to reach thermodynamic equilibrium, the calcine discharging from the desulphurization reactor will contain only the oxides of metals such as zinc, lead, cadmium and copper along with excess lime and environmentally safe CaSO 4 as anhydrite. Furthermore, provided there is an excess of lime, magnesium sulphate will not be formed but rather magnesium oxide and magnesium carbonate or mixtures of both will emerge depending on the particular conditions employed.

The chemical reactions of high temperature desulphurization are well known and there is current widespread interest in their application in the so-called second generation of coal-fired fluidized bed combustors (FBCs) in which sulphur in coal is fixed by incorporation of limestone as a desulphurization agent into the solid feed of a FBC. On the other hand, the metallurgical industry to date has not adopted high temperature desulphurisation of its strong sulphur dioxide streams as a treatment means for what would appear to be the following sound reasons:

(a) Process gas streams in sulphide smelting are very much more concentrated in SO 2 than those derived from coal combustion. Large tonnages of limestone, dolomite or the equivalent would have to be mined and then ultimately dumped for disposal.

(b) Because of the high SO 2 concentrations referred to in (a), prime consideration is normally given to the manufacture of sulphuric acid as a marketable byproduct of the smelting operation.

A number of factors including the geological makeup of the orebody and its surrounding mineralisation, the remoteness of the site from markets for sulphuric acid and the probable need to establish expensive port and infrastructure facilities for shipping sulphur, a relatively low bulk density product, all point towards a reassessment of the traditional viewpoint when considering the specific case of McArthur River.

Although elemental sulphur production by reduction of SO 2 with natural gas should not be dismissed, the author has a growing conviction that fixation of sulphur at high temperature and its disposal locally as anhydrite could in the long run turn out to be the most favourable in economic terms. If around half a million tonnes of bulky elemental sulphur per year does not have to be shipped to the outside world and if only compact metal products result from MR ore treatment, then clearly the viability of port development and its infrastructure must be suspect. Road haulage of the metals to Darwin using existing roads and port facilities becomes increasingly attractive once natural gas rather than coal is the preferred fuel and mineral concentrates no longer feature in the development.

If one is really looking for natural beneficial attributes of the McArthur River deposit, again the availability of dolomitic material in the same proximity as the orebody is a very significant positive factor in terms of energy recovery and sulphur fixation.

Oxidation of sulphide sulphur to SO 2 and then its capture at elevated temperature as CaSO 4 gives considerably more energy to the process than oxidation to SO 2 by itself. Effective use of this fact is a key component in the scheme that has been devised and is currently the subject of a UK patent application. Both zinc and lead are recovered as metals from MR ore in a

20 N.A. WARNER

single smelting furnace without significant external fuel input other than that required for air separation to satisfy the oxygen requirements for ore smelting. The outline flowsheet for this configuration is given in Fig. 8.

McARTHUR RIVER ORE

• DOLOMITIC =- MATERIAL

D ! o < 1 D O O D D D D O O D O 0 0 0 O O O O O 0 0 0 O O O O O O O O O O O O

• e f f l u e n t g a s 4 , o . - , i compos i te c lean-up & stack ._ ore feed i~

.- OXYGEN COPPER (SILVER) i ~

• ~L__ l . . I~rHOT GAS L. I ',~.'' . PRECIPITAT.J- l iquid sulphurt in sit u - : M A T T E

: . . . . . . . . . . . 7 slag : ~ . ,:.:" : | c l e a n •CIRCULATION ~ i JJ ~ .~Z INC/LEAD I : L ~ I ~ -~ - - - - -~ -~- - - - i'-~- :." I 7 C O N D E N S E R I : I - I ' 1 ' CHARGE ~" ~ R H ~ : ~ " - - - - - " J " - - PREHEATER ~ " ~ . i | - - , - - "~ - LEAD ZINC

r-SULPH UR t • . . _ • o I P ~ u n c M ~ o [ • i ron ~ vuseo • o I ~ " " ~ " v " ' l • c o n v e r t e r o r e • o

| : s lag s lag : | e n e r g y r e c o v e r y °° 0

l , , , , , , , 1 = . , , , , , , " t c a r r i e r g a s o t r e c y c l e loop

~- RECYCLE---~ ," t o °° IGAS BOOST. J " t o

O

0 0 0 0 0 0 0 0 D O 0 D O O O D D O 0 D D O O D D O D O O 0

t < !

I SLAG GRANULES & CALCINE

Fig.8 Outline Flow sheet for MR Ore Treatment

In the proposed process sulphur dioxide is being generated at high temperature by reaction of mineral sulphides with oxygen, air or oxygen-enriched air. After removal of entrained dust and fume from the smelter, the hot gases are passed continuously to a high temperature gas-solid contactor in which preferably some degree of countercurrent contact is achieved. A conventional long rotary kiln or staged bubbling fluidized beds are suitable in this regard or alternatively, circulating fluidized bed technology employing what is known as "fast fluidization" can be adapted. The product of the highly exothermic reaction between the hot gases containing sulphur dioxide and the now calcined limestone or dolomite generates a solid product or calcine at a substantially elevated temperature depending on the particular conditions but typically in the region of 1100K - 1300K or possibly even higher. It is this material that can be advantageously transported to the slag energy recovery plant for recovery of its sensible heat into the carrier gas circuit being employed for the recovery of thermal energy from slag. The overall chemistry involved in the desulphurization reactions at high temperature can be described as:

For limestone

CaCO 3 + SO 2 + I/2 02 = CaSO4(anhydrite) + CO 2 (13)

and for dolomite

CaCO 3 MgCO 3 + SO 2 ÷ I/2 02 = CaSO4(anhydrite) + MgO + CO 2 (14)


By bringing together droplets of partly solidified slag or even liquid slag drops and the hot calcine from the desulphurization reactor into the one fluidised bed, it is possible to secure very substantial process benefits. These can be assessed in laboratory tests simulating the conditions likely to pertain on the industrial scale. To understand this more fully the following background information is presented.

There are at least two distinctly different approaches to quenching slag droplets in a fluidised bed. As in work already being undertaken at Birmingham University, stabilisation of partly solidified droplets can be pursued by allowing large drops to rain through a bed of finer particles. The fine particles may not necessarily be of the same chemical composition as the slag drops, but more probably would be fine slag "sand". To retain fine particles in the bed implies relatively low gas velocities.

An alternative approach using much higher gas velocities with fine material totally entrained and recirculated to the bed via a cyclone has much to commend itself and would appear to offer better prospects for energy recovery in systems related to direct ore smelting in which the highest possible gas temperature from the energy recovery system is the prime objective.

A paper published in 1982 by Japanese workers, Yoshinga, et al. [19] claims successful quenching of a 5mm dia slag droplet into a fluidised bed of slag sand at 850oc. Their calculations indicate that it takes about 3 sec to cool a droplet from 1300oc if only the gas convective cooling is taken into account. They also point to the additional cooling due to slag sand adhering to the surface of slag granules and show a photomicrograph of a cross section of a cooled granule. The outer shell remains amorphous by virtue of the sudden quenching resulting from the slag powders adhered to the surface, whilst inside a large amount of crystallization is apparent indicating a very much slower cooling within. They also note the highly beneficial effect of slag powder coating preventing mutual adhesion and agglomeration of the slag granules.

Whilst dry granulation of slag by a fluidized bed of fine slag sand is a great leap forward compared with existing technology employing water quenching of slag without due regard to energy recovery, detailed engineering calculations have revealed already its limitations as a means for returning really high temperature gas back to an ore smelting operation. There is an inevitable degradation of available gas temperature in a system which has a primary quench not employing the full gas flow. If it were possible to effect the initial gas/solid contact in reasonably sized equipment, whilst accommodating a gas mass rate of flow of the same order as the solid, this would add considerably to the potential of the method as a means of transferring heat from fused ore slag back to the ore preheater of a direct smelter. The establishment of a stable fluidized bed of fine material under such high gas throughputs requires an excessively large cross-sectional area but by operating at very high space velocities, well above the terminal velocity of slag sand, it should be possible to design a relatively compact energy recovery plant. Rather than having a fluidizing air velocity of around Im/s as used by the Japanese workers, the slag granules themselves are fluidized at space velocities around 10 m/s and powdered fine material is entrained into the fluidizing gas to effect the initial coating of non-sticky material on the slag granules as they are assimilated into the main fluidized bed, which in the present case is a composite bed of slag granules and relatively coarse granular calcine from the desulphurization reactor.

A certain amount of surface abrasion of the calcine material occurs and by the process of attrition fine lime and magnesia along with anhydrite is held up within the coarse material of the bed for long enough to provide the coating of non-sticky material essential to prevent the slag granules from agglomerating with each other and becoming defluidised.

Above the turbulent bed of coarse composite material, the fines generated from the calcine are swept upward fully entrained in the fluidizing gas. As the molten slag drops must fall through the rising fine solids on their free-fall into the bed of coarser material, the quenching process and the surface adhesion of fine inert material is initiated during free-fall by interaction with fine entrained solids, again materially assisting in preventing agglomeration of drops once they enter the turbulent bed of coarser material.

22 N .A . WARNER

A further very worthwhile adjunct to the process being described is to use carbon dioxide as the principal gaseous component of the carrier gas recycle stream in the energy recovery circuit. This CO 2 can be derived from combustion processes occurring elsewhere in the plant particularly if oxy- fuel combustion techniques are being used, or alternatively by using well established near ambient temperature technology, CO 2 can be separated out from the gas effluent from the limestone or dolomite reactor circuit used for sulphur fixation.

Using CO 2 as the carrier gas rather than say the obvious choice of nitrogen from the air-separation plant required to satisfy the demands of oxygen smelting, results in smaller sized fluidized contactors and lower energy consumption for gas recycling in the energy recovery loop. More importantly, however, is that by staging the gas/solid contact using, for example, fluidized beds in series, the system can be designed so that the CO 2 reacts with MgO and unreacted CaO in the calcine at an intermediate temperature to revert these compounds back to carbonate.

CaO + CO 2 + CaCO 3 (15)

MgO + CO 2 ÷ MgCO 3 (16)

The reactions are kinetically rather slow but in energy terms are strongly exothermic. Recovery of at least part of this chemical reaction energy is worthwhile and the extent to which the reversion is pursued depends on the reactivity of the material available locally. In this regard it is worth noting that dolomite is known to be more reactive to SO 2 than limestone and by analogy one would expect calcined dolomite to be more reactive to CO 2 than calcined limestone. In economic terms, it becomes a balance between providing the necessary retention time and the value to the process of the energy recovered.

Recycling nitrogen as the carrier gas for energy recovery, in theory, does not require any adjustment of the oxygen potential before contacting the hot gases with the solid ore charge in the preheater, necessary to prevent oxidation of the contained zinc, lead and iron sulphides. However, any admission of air by leakage or accumulation of gaseous oxidising impurities would need counteracting as indeed the oxygen potential of CO 2 may need adjusting prior to admission of the hot recycle gas to the base of the charge preheater. In the McArthur River example being discussed, this is best done by recycling back to the hot gas stream at least part of the sulphur condensed out from the gaseous effluent from the preheater, in which pyrite decomposition provides the initial source of elemental sulphur, or alternatively, recycling the carrier gas hot so that it retains a small but adequate vapour pressure of elemental sulphur for the purpose.

The balance of the sulphur released by thermal dissociation of the pyrite contained in the ore feed is condensed to liquid. The crude liquid sulphur is pumped back to the oxidising arm of the furnace where it reacts with excess oxygen from the top blowing region (the oxygen utilization efficiency of the design considered in the paper is 72%) to provide a valuable heat input to the system.

The service requirements of the recycle gas booster in the energy recovery carrier gas recycle loop shown in Figure 8 are rather similar to those already in commercial operation at the Falconbridge Dominicana ferronickel plant. Their shaft furnaces recycle spent reduction gas at temperatures of 180 - 240oc from just below the stockline via cyclone dust separators and two blowers in tandem each rated for a gauge pressure of 17.5k Pa (2.5 psi) giving a total rated capability of twice this pressure. In the present flowsheet this recycle carrier gas stream is a closed loop with make-up or bleed as required but with no major gas effluent. This is important because the gas in this circuit could contain reduced sulphur compounds necessitating incineration before being released to the atmosphere. The bleed in the MR flowsheet would be added to the oxidising arm of the furnace and would leave the system via the one gas effluent from the process. This single gas effluent results from oxygen smelting plus a small air addition for supplementation of the total oxygen requirements for anhydrite formation and for gas temperature moderation. The gaseous effluent is thus relatively small in total volume in view of the total tonnage of metal being produced and its clean-up prior to stack discharge should accordingly be cost effective.


First estimates of the operating conditions and plant requirements in the desulphurization and energy recovery section are summarised in Table 5. One possible plant configuration for contacting the liquid slag droplets, hot desulphurisation calcine and the energy recovery carrier gas is given in Figure 9. This depicts three circulating fluidized beds (CFBs) arranged for three stage countercurrent contact. The space velocities in each CFB would be in the region of 10ms -I.

TABLE 5 Indicative Process Requirements for Treating 10,000 tpd MR Ore

(Basis: Ratio Dolomite to Rock in Dolomitic Material = 5/1)

I . Total solids to energy recovery

Total fused ore plus "converter slag" CaSO4/MgO/rock calcine

TOTAL SOLIDS TO ENERGY RECOVERY

= 8,700 tpd ffi 9,780 tpd

= 18,480 tpd

2. Countercurrent Multiple Fluidized Bed Contact For Enerqy Recovery

(a) Number of stages = 3

(b) Efficiency of heat utilization {after Kunii and Levenspiel)

= 75 percent

(c) At 1350K and 2 psig, terminal velocities dp = 2mm ; U t = 15.5 m/s dp = 3mm ; U t = 22.5 m/s

(d) Bed diameter for U o = 10 m/s (i) for carbon dioxide, diam. = 7.7 m (ii) for nitrogen, diam. = 10.0m

3.

4.

Impinqing Jets for Slag Disintegration

(a) Slag head in tundish = 600 mm

(b) jet diameters = 50 mm

(c) Number of pairs required = 4

Rotary Kiln Dolomite Reactor for Desulphurization

Maximum gas rate = 42m3/s at 0oc for MR Plant

= 38m3/s at 0oc (a) Calculated for Exmibal Reduction Kiln (Laterite Ni) 5.5m dia x 100m

(b) Calculated for Long Rotary Lime Kiln - US Practice = 32m3/s at 0°C 4.27m dia x 122m

Based on (b), One kiln 5.5m dia for MR Smelting

Desulphurization service.

24 N . A . WARNER

G O S t O 1 Ntt

M o l t e n s l o g

H o t C~rrte~ TunOtsn

I e t c ,

O II i ® G O S f r o m 2

®

G O S f r o m 3

Fig.9

S l O g G r o n u l e s R e c y c l e C o r r t e r o n O C o l c l n e G O S

Energy recovery from slag and hot desulphurisation calcine using staged circulating fluidised beds

THE TOP BLOWING REGION

Flow and mass transfer in the liquid

With reference to Figure 10, when a gas jet strikes a liquid surface, part of the momentum in the jet is transferred to the liquid due to shear forces between the two phases. For the impact of a single jet, this results in the induction of a radial surface flow. Measurements made by Wakelin [20] using a water model have shown the radial velocity away from the point of jet impact, Ur, to vary inversely with the radial distance, r, from the 'source'. When this is the case, the liquid phase mass transfer model developed by Davenport [21] which is based on the Higbie-Danckwerts theories of surface renewal, predicts the mean liquid phase mass transfer coefficient over a round bath of radius r b to be:

kL(bath) = 2(2/~) 1/2 ((Urr)DL)I/2 (1/r b) (17)

With the mean t~ansfer coefficient over the bath surface varying with the root of the diffus~vity. This is typical of systems where transfer can be described by surface renewal theory.

This simple picture of liquid phase mass transfer is of little practical use in predicting the top-blowing requirements for direct polymetallic smelting. In a real system, the inherent complexities are such that further progress can only be made by experiment. Empirical data can then be correlated by the use of the appropriate dimensionless groups as reported recently by Jones and Warner [22]. The essence of this approach is that while the temperature difference between room temperature experiments and the high temperature


application is large, the fundamental mechanisms of transfer operating in the cold systems should be similar to those at high temperature. Data from the various systems are correlated by the application of suitable similarity criteria, the selection of these being rationalised by the use of dimensional analysis. It was assumed that:

k L = f(M,H,do,db,DL,g,QL,~L,O )

where

(18)

kL

M :

H : d O : d b = D L = g = IlL = O =

= liquid phase mass transfer coefficient based on area with diameter db, ms -I jet momentum per unit time or flux, N lance height above liquid surface, m nozzle diameter, m bath diameter, m diffusivity in liquid phase, m2s -I acceleration due to gravity, ms -2 liquid viscosity, kg m-ls -I surface tension, Nm -I

Gas

Liquid

Fig.10 Gas jet impinging onto a liquid surface

A series of experiments involving variation of the nozzle diameter do, holding all other parameters constant, suggested that d o could be left out of the analysis. The system can then be completely described using the following groups:

DL PLDL PL g UL

(19)

i.e.

NSh = f (NSc, Nmom, NGa, H/d b, NMorton)

where

NSh Sherwood number ) NSc Schmidt number ) NGa Galileo number ) N M Morton number ) NRe Reynolds number ) Nmo m Momentum number )

As defined in the above equation.

The final form of the correlation is illustrated in Figure 11 which incorporates oxygen desorption data from the recent study together with data from previous laboratory top-blowing studies involving carbon dioxide absorption into water and oxygen absorption into molten silver.

26 N . A . WARNER

NshNsc-O.SNG -~J3(db/H)2NMO.1 s / 10 .2 log/log ~ /

1 0 " 3

1 0 - 4 I I I

10 .4 M / H3P Lg 10 "1 Liquid Phase Mass Transfer Correlation

Fig.t1 Dimensionless correlation of liquid phase transfer data

Given the widely differing physical properties between the hot and cold systems, the correlation is considered acceptable. With reference to the smelter application, a prediction of lead and zinc losses in the converting zone, and of oxygen transfer to the bulk matte, is now possible for a given set of blowing conditions.

Mass Transfer in the Gas Phase

The relatively gentle blowing conditions envisaged, combined with the high density of the liquid matte, suggest that for the purposes of calculating oxygen transfer in the gas phase, the impacted surface may be regarded as being approximately flat. In this case, the Holger Martin [23] model for gas phase transfer is the most appropriate for the calculation of oxygen transfer rates. With dimensionless parameters based on the nozzle, the basic transfer equation takes the form:

NshNsh0-42 do [ I - 1"I (do/r) 1 -- f(NRe) (20) --r I + 0.1[ (H/d O )]-6]do/r

where NSh = kGdo/D , NRe = doUoPG/UG , and k G is the gas phase transfer coefficient over the surface of radius r.

For the present design, the nozzle diameter is 0.1m, the nozzle velocity is 8 m/s, which for oxygen gives a Reynolds number of 2.2 x 104 if the transport properties in our non-isothermal case are evaluated at mean film conditions.

Function f(NRe) then takes the form:

f(NRe) = 1.36 NRe 0-574

Basing gas phase calculations on these equations and assuming a nozzle spacing in a multiple array of say 1.5 m with H = 0.6 m, the gas phase transfer coefficient is determined as 0.104 m s -I . Transfer is calculated using a log mean oxygen driving force over the impacted surface, allowance having been made for the entrainment of furnace gas into the jet. With a mean driving force of 0.33 kg m -3, an oxygen transfer rate of 6.02 x 10 -I per lance is achievable. This represents an overall oxygen efficiency of 72 percent.

Oxygen Requirement and Nozzle Arrangement

Treatment of I0,000 tpd of run-of-mine MR ore is equivalent to a zinc feed of 10.42 kg/s and from the chemical analysis and reaction stoichiometry, the


oxygen required to react with the iron sulphide to produce a ferrous oxide containing slag and SO 2 along with that required for direct oxygen smelting of ZnS and PbS in the feed works out at 18.2 kg/s. Given that the oxygen utilization efficiency for the configuration being considered is 72 percent, this is equivalent to an actual oxygen flow of 25.3 kg/s. One nozzle delivers 0.083 kg/s so the total number of nozzles required is 305.

No claim is being made that the above number, pattern and nozzle size are optimum, but the figures given are to be regarded as indicative. It should be pointed out that heating or cooling of large flat surfaces is often carried out in devices consisting of arrays of round or slot nozzles through which air or other gaseous medium impinges vertically upon the product surface. The annealing of metal and plastic sheets are examples. The configuration of nozzles in mind in the present case is analogous to these examples but bears little resemblance to the single or multiple lances associated with LD steelmaking, Isasmelt technology or indeed the Mitsubishi continuous copper process.

The nozzles in the present case operate well below the onset of liquid splashing and their soft blowing characteristics can only be expected to penetrate the very thin slag layers produced as a result of forced melt circulation which carries the slag away immediately from the zone of impact. Structurally, one can envisage oxygen being piped to a number of header box sections with the underneath side containing many orifices all impinging downwards onto the relatively quiescent melt surface maintained in bulk flow by forced circulation.

Simple geometry indicates that a furnace about 10 m wide by 60 m length would be required to accommodate the configuration being discussed. This is somewhat longer than a large copper reverberatory furnace of say 10 m wide by 35 m length, but the aspect ratio is more appropriate for a melt circulation process. It is more akin to a float glass chamber, which in commercial practice can be around 8 m wide by 61 m in length.

Finally, to put the calculated furnace size of 10 m by 60 m in perspective, it must be remembered that 10,000 tpd MR ore is equivalent to an annual zinc production of 324,000 tonnes per year and as such is larger than three modern zinc smelters put together.

Metal losses : predicted liquid phase transfer

Potential sources of metal losses in the overall smelting circuit are shown schematically in Figure 12.

Zn & Pb revert to sulphide

Vacuum Vessel

I% I% I I t I % I

I t I | I I | I % I

I t I t I %

I t i t s l a l t o g Zn & Pb oxidised

Pb,Zn

Oxygen Oxygen Oxygen

Top-Blow

M e t a l Losses

Fig.12 Potential sources of metal losses

Given the specified blowing configuration, prediction of~the liquid phase transfer rates of oxygen, lead, and zinc is possible using the

28 N.A. WARNER

correlation depicted in Figure 11. Basing calculations on pure oxygen leaving the nozzle at room temperature, the jet momentum flux for each lance is calculated as the product of nozzle velocity and mass flowrate. In this case, M takes the value 0.67 N for each jet, this being the same at the nozzle and the point of impact. Using a matte density of 4520 kg.m -3 the momentum group M/H3pL is found to be 7 x 10 -D. When the other relevant dimensionless groups are calculated, and reference made to Figure 11, the value of the liquid phase mass transfer coefficient k L is found to be 1.81 x 10 -5 m s -I . If pure oxygen is blown, the interface gas composition will be SO 2 at effectively one atmosphere pressure, the corresponding equilibrium dissolved oxygen content of the Cu2S being calculable as 55 kg m -~. The mass transfer rate of oxygen into the bulk matte is then calculated as the product of the transfer coefficient, the area on which this is based, and the concentration driving force, WC.

Oxygen transfer = k L A WC

= 1.81 x 10 -5 x 1.77 x 55

so for 305 jets

= 1.76 x 10 -3 kg s -I

= 305 x 1.76 x 10 -3

= 0.537kg s -I

This being the rate of oxygen dissolution into the bulk matte.

Every mole of dissolved oxygen will appear as one mole of SO 2 in the R-H vessel, this in turn possibly causing 3 moles of zinc to revert to the oxide and sulphide. In the present case, this means that up to a maximum of 3.3 kg s -I zinc could revert, representing 32 percent of the total zinc input. To reduce this figure it will clearly be necessary to deoxidise the matte prior to arrival at the vacuum vessel. However, it must be stressed that these are not zinc losses but rather indicate a possible loss in efficiency in once through recovery.

For a given blowing configuration, lead and zinc losses in the converting region will be determined by the steady state concentration of these species in the circulating matte, this being controlled by the rate at which metal is removed from the vacuum vessel. Expected zinc losses, calculated for the proposed design, are based on steady state zinc concentrations of 0.25 and 0.5 percent. Expressed as a percentage of the total proposed zinc make, the losses represent 1.1 and 2.2 percent, respectively. With the actual steady state concentrations expected to lie in this range, metal losses in the converting zone are thus not considered to be serious.

Melt Circulation and Vacuum Pumping Requirements

For the direct smelting of 10,000 tpd MR to metallic zinc and lead, about 85t/h of metallic copper has to be transported to the RH vacuum vessel from the top blow region. The preferred way of doing this is to make use of the mutual solubility of copper and copper sulphide matte and in this way no external handling of molten copper is required. The matte acts as a solvent for both copper and the zinc and lead sulphides.

In the present case, various theoretical considerations indicate a required mass ratio of matte circulation rate to total zinc plus lead metal production of about 250 to I. To put this figure in context it is of interest to note that a zinc-lead blast furnace circulates lead through its condenser system at a rate of some 380 or so times the zinc production rate.

At the above melt circulation ratio, the various changes in solute concentration and temperature can be evaluated. Thus there would be a change in copper content of the matte in passing through the RH vessel of about 0.63 wt percent, the zinc level would drop by 0.28 wt percent and the melt temperature would drop by at least 21oc. In the oxygen top blow region the temperature differential between inlet and outlet matte could be up to about 75°C if the post combustion of excess oxygen with elemental sulphur is also included.

Considering the liquid phase mass transfer taking place inside the RH vessel with an inlet zinc level of 0.78 wt percent, the outlet zinc concentration would be 0.5 percent, whilst at the gas/liquid interface, assuming complete


backmixing of liquid within the vessel, would be in the region of 0.27 wt percent, which at a temperature of say 1380oc would generate an equilibrium zinc partial pressure of about 44 mm Hg.

At this stage it must be stressed that the various temperature and concentration changes are self regulating. They will find their own absolute levels without sophisticated control. On the other hand, if zinc losses to slag, for example, are to be limited, the zinc concentration base level in the circulating matte cannot be allowed to rise too far.

A first order estimate of the vacuum pumping required to sustain melt circulation can be derived from published information on vacuum degassing of steel. In volumetric terms the melt circulation duty required is at least some five times that of a commercial steel RH installation so the lift gas requirements have to be extrapolated. On the other hand it seems likely that very much higher residual gas pressures can be tolerated in the zinc smelting case and for the present purpose it is assumed that vacuum pumping at 10mm Hg is adequate.

The lift gas volumes are likely to be relatively small compared with the volume of SO 2 released from the matte, particularly if technology is developed to condense out zinc in the presence of substantial SO 2. The highest energy consumption for vacuum pumping will correspond to zero matte deoxidation elsewhere in the matte circuit, in which case all the oxygen absorbed by the melt in the top blow region will appear as evolved SO 2 in the RH vessel. In the present case this corresponds to a maximum volume of gas to be exhausted back at the pumping station of about 158,000 m3/h at 10mm Hg and I00OC, which is equivalent to an actual electrical power requirement of about 740 kW based on a vacuum pump overall efficiency of 67 percent. Mechanical pump sets with capacities of 230,000 m3/h at I mm Hg pressure are known to have been used for steel degassing in European practice, so in this respect the technology may be considered as proven.

Area Requirements for Ore Fusion

The thermal power to be transmitted to the basal area of the preheated solid charge column by forced convection heat transfer in the matte bath works out at 115 MW. Assuming that a local melt velocity 0.4 m/s is maintained in the matte flow channels between the solid charge and the underlying hearth protection layer of molten copper and the equivalent diameter of these flow channels is sa Z 0.2 m, the local heat transfer coefficient can be shown to be about 3860 W/mZK. Under these conditions a total area of 180 m 2 is required for ore fusion. If the furnace width is 10m, this means that about 18m of the furnace length would be committed to assimilating the preheated ore into the bath.

Downstream of the ore melting section of the non-oxidising arm of the melt loop, provision for phase disengagement is required as matte droplets dispersed in molten slag settle out. It may also be necessary to carry out matte deoxidation in this non-oxidising arm of the furnace.

The basic assumption in the size estimate for ore fusion is that the process is controlled by liquid phase heat transfer. This is indeed the case when pellets of lime plus copper-nickel ore are fused by immersion in copper-nickel matte. Recent work published by Mwansa and Warner [24] indicates that the rate of fusion of composite ore pellets into matte can be predicted by a model based on liquid phase heat transfer as the rate controlling step. There is no obvious reason why McArthur River ore lumps should not behave similarly.

The real area of uncertainty in the proposed design stems from a lack of information on the extent of silicate compound formation in the charge preheater as opposed to delayed reaction until the solids are immersed in the melt. The formation of silicates from the free lime and magnesia resulting from decomposition of dolomite in the composite ore charge is highly exothermic. Whilst not making any difference to the overall heat balance, the location of the occurrence of the reaction is significant in determining the thermal load to be supplied by the circulating matte. The figures given are conservative in that it has been assumed that the exothermic heat terms were not available to assist ore fusion in the bath.

30 N A WARNER

Overall Energy Considerations

In broad brush terms the treatment of 10,000 tonnes per day of McArthur River Ore by the proposed advanced smelting process requires very substantial rates of thermal energy transfer from exothermic or heat producing zones to areas where heat is absorbed.

The thermal power on the debit side is made up of two major terms. Firstly there are the requirements for decomposition of carbonates, dissociation of pyrite and melting of both sulphide and oxide components of the composite ore charge. This can be considered as the rate of energy dissipation in the overall process of ore fusion and in the present case this amounts to about 300 MW. The second major energy consumer is metal production in the RH reactor which works out at about 44 MW.

The exothermic terms on the credit side of the account are comprised of the energy transferred back to the smelter by the carrier gas of the energy recovery circuit together with the heat generated on the oxidising side of the matte circulation loop by iron sulphide oxidation to slag, copper sulphide conversion to copper metal and combustion of reverted liquid sulphur in excess oxygen from the top blow region. The rate of energy recovery term is either 225 MW or 290 MW depending on whether magnesia and unreacted lime are converted back to carbonates. Energy is transmitted to the circulating matte in the top blow and associated region at a rate of 160 MW.

The figures given are for bone-dry feed materials and with zero radiation and convection heat losses and in these terms, clearly the available input of 450 or 385 MW exceeds the total requirement of 344 MW. When realistic heat losses and moisture contents are incorporated into the overall heat balance, it is seen that the MR ore smelting system is approximately autogenous or thermally self-balanced. However, to sustain this very favourable position, electrical energy is consumed as summarised in Table 6.

TABLE 6 Electrical Power Requirements for MR Ore Smelting

Treatment rate 10,000 t ore/day MW

Air Separation Plant Gas Recycle for Energy Recovery RH Vacuum Pumping Requirement Rotary Kiln & Ancillaries Hot Gas Handling & Dust Collection Low Pressure Air Sulphur Pumps Molten Metal Pumps

27.9 11 .0 2.5 2.0 1.3 1.4 0.3 0.2

TOTAL ELECTRICAL POWER 46.6 MW

To arrive at a figure for the gross energy requirement for metal production from MR ore, information is required on the electrical loads involved in the mine, water supply, town supply, power station auxiliaries, mine crushing and pumping and mine site infrastructure. Such data are available in a major feasibility and environmental report submitted to the N.T. Government by MIM. An indicative figure for this associated electrical load is about 30 MW [25].

The remaining major energy consideration relates to the provision of around 14,000 tpd dolomitic material for the combined usage of ore fluxing and sulphur fixation. Very much influenced by statements made by Croxford and Jephcott [3] in relation to the stratigraphic succession in the McArthur area and the absence of further geological input at this time, it is assumed that dolomitic material is available nearby or at least within a 20 km radius of the plant site. The gross energy requirements of the associated dolomite mining operation involved including transportation to the smelter site has been estimated using the approach detailed by Chapman and Roberts [26]. A figure of 78 MJ/t is obtained and to this a further energy input of say 35 MJ/t at the open-pit dolomite mine is added for crushing the material to


-15mm. This gives a gross energy rate for the provision of dolomitic material of 18.3 MS( to which a further say 1.6 MS( power at the smelter site would be consumed in fine crushing, fine screening and storage prior to incorporation into the smelting process. These gross energy requirements in going from ore to metal are summarised in Table 7.

TABLE 7 Gross Energy Requirements for Zinc and Lead Metal Production

(10,000 tpd MR ore treatment rate) MW

I . Electrical Power

(a) Direct Ore Smelting 46.6 (from Table 6)

(b) Associated Electrical Load, covering the following: (Mine, water supply, town supply, power ) (station auxiliaries, mine crushing and ) (pumping, mine site infrastructure )

(c) Dolomitic material, fine crushing, screening and storage

30.3

1.6

2.

Sub-total Electrical Power

Thermal Power Equivalents

(a) Electrical Power at 30% conversion efficiency (from above)

(b) Dolomitic Material

78.5 MW

261 .7

18.3

Sub-total thermal power equivalent 280 MW

The 280 MW (thermal) figure in Table 7 can be translated into energy requirements for zinc and lead metal production, but to do so requires, as pointed out by Chapman and Roberts [26], some convention on how to divide the energy between the two products. In the present case a weight basis for apportioning the energy requirements seems logical and this gives 19.6 GJ/tonne as the gross energy requirement for both metals. This can be compared with published figures of 61.3 GJ/tonne for electrolytic zinc and 48.7 GJ/tonne for thermal zinc via the zinc-lead blast furnace, in all cases going from ore to finished metal product. The same literature source gives the gross energy requirement for refined lead production as 28.4 GJ/tonne.

The MR ore smelting figures relate to zinc 4 grade and unrefined lead bullion. In the zinc case, reference has already been made to the potential for autogenous further vacuum refining to Zlnc I specification with respect to lead content and only a small additional energy input to reach Zinc 1 for cadmium content, so if this approach is adopted the 19.6 GJ/tonne figure would not alter dramatically, perhaps by I or 2 GJ/tonne. Lead bullion refining, on the other hand, consumes around 8.3 GJ/tonne according to Chapman and Roberts [26] so the gross energy requirement for lead produced from MR ore becomes (19.6 + 8.3) = 27.9 GJ/tonne, apparently comparable to the 28.4 GJ/t figure already quoted for current technology.

ACKNOWLEDGEMENT

The author is indebted to the Minister of Mines and Energy, the Hon. Mr. Barry Coulter, for providing the opportunity to present the original version of this paper to the Mining, Petroleum and Suppliers Conference, Darwin, Northern Territory of Australia, May 10-13, 1988

32 N.A. WARNER

REFERENCES

I. D. Royston, F.S. Wong and E. Yigit, Selective chloride leach process for McArthur River semi-bulk concentrates in Complex Sulfides Processing of Ores. Concentrates and By-products A.D. Zunkel, R.S. Borman, A.E. Morris and R.J. Wesley, ed; 397-417, TMS-AIME Warrendale Pa. (1985)

2. D.T. Buchanan, The McArthur River Project. The Aus. I.M.M. Conference Darwin, N.T., 49-57 (1984)

3. N.J.W. Croxford and S. Jephcott, The McArthur lead-zinc-silver deposit, N.T. Proc. Aus. I.M.M. No. 243, 1-26 (1972)

4. C.P. Broadbent and N.A. Warner, Review of the current status on energy recovery from slag Energy and the Process Industries Inst. Mech. Engrs., London, 23-29 (1985)

5. C.P. Broadbent and N.A. Warner, A novel approach to energy recovery from slag Iron and Steel Intl. 57, 101-105 (1984)

6. C.P. Broadbent and N.A. Warner, Self-impinging jet disintegration of slag for energy recovery Trans. Instn. Min. & Metall. Sect. C 93, 30-133 (1984)

7. D. Johnson, The RH degassing system at ATH Ruhrort Works Iron Steel 46, 174-177 (1973)

8. O. Kubaschewski, E.LI. Evans and C.B. Alcock, Metallurgical Thermochemistry 4th edn. Pergamon, Oxford 495 p. (1967)

9. E.G. King, A.D. Mah and L.B. Pankratz, Thermodynamic Properties of Copper and its Inorganic Compounds Cambridge, Mass.: The International Copper Research Association (1973)

10. N.A. Warner, Kinetics of continuous vacuum dezincing of lead in Advances in Extractive Metallurgy 317-32, 421-2 London: I.M.M. (1968)

11. J.G. Herbertson and N.A.Warner, Liquid-metal mass transfer in vacuum distillation of the circulating lead of a zinc blast- furnace Trans. Instn. Min. Metall. Section. C 82, C-16-20 (1973), 83, C126-9 (1974)

12. S. Goto, Equilibrium calculations between matte, slag and gaseous phases in copper smelting in Copper Metallurgy: Practice and Theory. M.J. Jones etc. 23-24 London: I.M.M. (1975)

13. I. Barin and O. Knacke, Thermochemical Properties of Inorganic Substances Springer-Verlag, Berlin and New York, (1983)

14. S.N. Sinha, and H.Y. Sohn, Activities of PbS and A-S 0 5 in High- Grade Copper Mattes in Advances in Sulfide Smelting H.Y. ~ohn, D.B. George and A.D. Zunkel, ed; 171-196, AIME, New York, N.Y. (1983)

15. N.A. Warner, Towards polymetallic sulfide smelting in Complex Sulfides - Processing of Ores Concentrates and By-products, A.D. Zunkel, R.S. Boorman, A.E. Morris and R.J. Wesley, ed; 847-865 T.M.S. AIME Warrendale Pa. (1985)

16. G. Rottmann and W. Wuth, Conversion of copper matte by use of the top-blowing technique in Copper Metallurgy: Practice and Theory M.J. Jones ed., 49-52, London: IMM (1975)

17. N.J. Themelis and H.H. Kellogg, Principles of sulfide smelting, in Advances in Sulfide Smeltin 8 H.Y. Sohn, D.B. George and A.D. Zunkel, ed; 1-29, AIME, New York, N.Y. (1983)

18. G.H. Kaiura and J.M. Toguri, Natural convective mass transfer rates between solid magnetite and molten mattes Metal1. Trans. B. 10B, 595-606 (1979)

19. Yoshinga, et al. Dry granulation and solidification of molten blast furnace slag Transactions Iron and Steel Instn. of Japan 22, 823-829 (1982)

20. D.H. Wakelin, PhD thesis, University of London, 1966. 21. W.G. Davenport, A.V.B. Bradshaw, D.H. Wakelin, Interaction of both bubbles

and gas jets with liquids in Heat and Mass Transfer in Process Metallurgy Hills, A.W.D. ed. 207 London: IMM (1967)

22. T. Jones and N.A. Warner, Top-blowing requirements for direct polymetallic smelting in Pyrometallurgy '87 605-626, IMM, London (1987)

23. H. Martin, Heat and mass transfer between impinging gas jets and solid surfaces Adv. Heat Transl. 13, 1-60 (1977)

24. J.C. Mwansa and N.A. Warner, Natural convective heat transfer between single ore pellets and molten copper nickel mattes in African Mining 253-266, IMM London (1987)

25. Personal Communication, C.W. Hoffmann, M.I.M. Holdings Limited, 19 February 1988

26. P.F. Chapman and F. Roberts, Metal Resources and Energy Butterworths, London, pi09, 207, 208, 215, 216 (1983)

advanced technology for smelting mcarthur river ore

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