Uploaded by mikrametalurgi

ContentServer (2)

Smelting of difficult laterite ores
Maurice Y. Solar and Sina Mostaghel*
The favourable physicomechanical characteristics of nickel have resulted in its wide application in
various products, around 2m t year21. Nickel is an essential alloying element in austenitic
stainless steel and other speciality alloys. The drastic increase in production of such alloys over
the last decade, mostly because of economical growth of Asian countries, particularly China, has
resulted in a considerably increased demand for ferronickel. To fulfil such an expanding industrial
demand, newer and more efficient routes for extraction of nickel from poorer and more difficult
ores are required. The current paper summarises years of experience in designing, operating,
and optimising ferronickel smelters, and describes how this knowledge can be used to develop
the mitigation measures necessary to address some of the foreseeable smelting challenges
associated with metal extraction from low-grade and previously undesirable laterite ores.
Keywords: Laterite, Ferronickel, RKEF, Furnace stability, Smelting challenges, Phosphorous
The abundance and geology of nickel in the earth’s crust,
its metallurgical history and economic importance have
been discussed by numerous authors over the years (Bolt,
1967; Elias, 2002; Halikia et al., 2002; Mudd, 2009;
Makinen and Taskinen, 2008; Rhamdhani et al., 2009a,
2009b; Kobayashi et al., 2011) and more recently by
Bunjaku and coworkers (2011, 2012a, 2012b), and Swinbourne (2014). However, the current paper focuses on
smelting of laterites; hence, nickel production from other
sources and/or by other methods will not be discussed.
Laterite deposits are formed by weathering of ultramafic
rocks, which are complex mixtures of ferromagnesian
minerals such as olivine (Fe,Mg)2SiO4, pyroxene (Fe,Mg)2
Si2O6 and amphibole (Fe,Mg)7Si8O22(OH)2 (Dalvi et al.,
2004). Consequently, the deposits are predominantly
found in wet tropical areas such as Cuba, Indonesia,
New Caledonia, the Philippines and South America.
Exceptions are the ‘dry’ laterites found in Australia.
Most laterite deposits can be divided into three layers of
potential economic interest, namely:
1. Limonite: Goethite (FeOOH, also written as
Fe2O3?H2O), haematite and sometimes magnetite
are the main minerals. Nickel grades are usually
between 1?0 and 1?5 wt-%. This layer is rather
2. Saprolite: Garnierite (Ni,Mg)6Si4O10(OH)8 is the
main nickel-bearing mineral with grades of up to
10 wt-%Ni. Saprolites are quite heterogeneous and
contain many other minerals with no nickel or less
nickel than garnierite. Their overall nickel grades are
around 1?8–2?5 wt-%.
3. Transition: Lies between saprolite and limonite.
Conventional wisdom is that limonites must be leached
and saprolites must be smelted (Crundwell et al., 2011).
But these are economic rather than metallurgical
imperatives. For example, Larco in Greece is (and the
ferronickel plants in Macedonia and Ukraine were)
smelting nickeliferous haematite with a composition
very similar to limonite, and Goro in New Caledonia
was designed to pressure acid leach the totality of the ore
body, the bottom third of which is saprolite.
One characteristic of these deposits is that nickel is
distributed on an atomic scale within the nickel-bearing
minerals and that, unlike sulphides, these minerals
cannot be concentrated by conventional mineral dressing techniques. Some upgrading can be achieved by
screening out the coarser fractions or, in a few rare
cases, by heavy media separation. The ores may contain
up to 40 wt-% free H2O but are typically of 25–35 wt-%
moisture and up to 12–13 wt-% crystalline water.
Pyrometallurgical processing of such wet ores requires
drying, calcination and smelting of all of the upgraded
material, and refining of the resulting ferroalloy, each
step being quite energy intensive. Moreover, the
increased demand for nickel (mostly Chinese) over the
past decade has meant that the deposits being brought
into operation are lower grade, often with undesirable
SiO2/MgO ratios, high goethite and thus high Fe/Ni
ratios, and high alumina concentrations. These ores pose
new practical challenges to operators. A few of these
involve the kilns; for example, the presence of more and
more free silica and goethite limits the calcine temperature that can be achieved without sintering. Most of the
challenges, however, involve operation of the electric
furnaces. The current paper will therefore focus on some
of the difficulties faced in the smelting process.
Major challenges
Hatch Ltd, 2800 Speakman Drive, Mississauga, Ont., Canada L5K 2R7
*Corresponding author, email [email protected]
ß 2015 Institute of Materials, Minerals and Mining and The AusIMM
Published by Maney on behalf of the Institute and The AusIMM
Received 11 April 2014; accepted 13 August 2014
DOI 10.1179/1743285514Y.0000000075
The first important consideration is that not all potential
ores can (or should) be smelted to an arbitrarily pre-selected
Mineral Processing and Extractive
Metallurgy (Trans. Inst. Min Metall. C)
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
26.5 (matte)
Nickeliferrous haematite
Dominican Republic
New Caledonia
Larco GMMC
Not known
Aneka Tambang
Solway Group
PT Inco
Cerro Matoso S.A. (CMSA)
Loma de Niquel (LdN)
Buruktal Metallurgical Plant (BMZ)
Doniambo (SLN)
PT Antam
Pobuzhsky Ferronickel
Combine (PFK)
Pacific Metal Co. Ltd (PAMCO)
Current owner
Table 1 Selected laterite smelters – Essential information
Operators are aware that lower ferronickel grades result
in lower nickel loss to the slag phase through both lower
nickel contents in the slag and lower slag volumes
(because more iron reports to the ferronickel phase).
However, higher reductions require more reductant
and more power per tonne of calcine smelted. Lower
ferronickel grades thus typically result in lower ore
throughputs, as most smelters are furnace limited.
Therefore, for any given ore and smelter, there is a
maximum nickel production level, where increasing
recovery is counterbalanced by decreasing ore throughput. Figure 2 summarises these correlations as function
of the extent of iron reduction for an ore at 1?6 wt-%Ni,
16 wt-%Fe, and 2?2 SiO2/MgO. Note that the sudden
change in the slope of the slag mass curve, at around
45% iron reduction, is because of dissolution of silicon in
the ferronickel produced, the behaviour of which has
been discussed earlier by Solar et al. (2008). Operating
costs also vary widely from location to location. The
costs specific to a given location must be determined,
taking into account the cost of the ore itself (whether it is
mined on site or purchased from third parties), labour,
power, fuel, reductant and other commodities costs,
closeness to market, etc. Revenues must then be
estimated. That is relatively easy in the case of nickel,
but what about iron? Ferronickel producers typically do
not receive any credit for the iron contained in their final
product, unless the nickel market is very tight and
customers are anxious to ensure supplies. However some
producers, particularly in China, are affiliated with
stainless steel plants and highly value the ‘free’ iron
Combining these three sets of calculations, i.e. mass
balances, operating costs, and revenues, into a net
present value for a series of potential ferronickel grades
yields curves with maxima, which depend mostly on the
iron credits assumed. The grade corresponding to the
Ore type
Determination of the economically optimum
ferronickel grade
Ore composition
Reduction to ferronickel/ %
ferronickel grade. Table 1 and Fig. 1 show the wide variety
of ferronickel grades that operators have chosen to
produce. It is evident that ferronickel smelters fall within
two groups: low reduction smelters that reduce about 15–
30% of the iron in their ores, and high reduction smelters
that reduce about 45–65% of their iron. Nickel pig iron
(NPI) is simply an extension of the latter case, with NPI
producers pushing the extent of iron reduction to the 70–
80% range. Larco and Buruktal are examples for which the
limitation is physical: ferronickel at 20 wt-%Ni cannot be
produced directly from ores about 1 wt-%Ni and 30 : 1 Fe/
Ni ratios (Zevgolis, 2004), at least not without excessive
nickel losses to the slag phase. Larco produces a ferronickel
at about 12 wt-%Ni in its furnaces and upgrades this to
about 18–20 wt-%Ni, by oxidising iron in an oxygen
converter. Cerro Matoso S.A. (CMSA) is an example for
which the limitation is economical. It is in a relatively
remote location, far from nickel users; therefore, it is
advantageous to produce a high-grade ferronickel and
thereby reduce smelting, refining and product transportation costs.
Every new project should therefore start with a
determination of its optimum ferronickel grade, as
described by Solar et al. in 2008. This determination
requires the following steps.
Laterite Smelting Challenges
Solar and Mostaghel
Solar and Mostaghel
Laterite Smelting Challenges
1 Ferronickel grades v. extent of iron reduction for different laterite smelters
maximum calculated for the most likely iron credit is the
‘economically optimum’ ferronickel grade.
Assessment of metallurgical feasibility of this
optimum grade
This step involves estimating a multitude of operating
parameters such as the liquidus temperatures of the
ferronickel and slag phases, the required ferronickel and
slag superheats and operating temperatures, the potential stability of the operation, the likelihood of silicon
reversals and carbon boils, etc. The selection of the
target ferronickel grade for a new project is thus a
compromise between the economic optimum for a given
ore/smelter and the likely stability and longevity of
operation at such a grade. Therefore, this selection
cannot be made arbitrarily based on a preconceived idea
of what the market prefers.
Control of ferronickel grade and impact
on furnace operating stability
Ferronickel grades in actual operation can typically be
controlled to only within ¡10% relative of the target.
This means that a high reduction smelter aiming at a
grade of 20 wt-%Ni will actually achieve final products
in the range 18–22 wt-% Ni, while a low reduction
smelter targetting 35 wt-%Ni will have products varying
from 32 to 38 wt-%Ni. These fluctuations can have
many causes, e.g. variations in ore and coal composition
and granulometry, batch addition of recycled streams
2 Nickel grade, recovery, production and masses of slag and ferronickel are correlated with the extent of iron reduction
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
Solar and Mostaghel
Laterite Smelting Challenges
3 Carbon content v. extent of iron reduction for commercial ferronickel smelters
4 Impact of carbon and silicon on the liquidus temperature of ferronickel at 15 and 20 wt-%Ni. For each set of lines, the
upper one is for 15 wt-%Ni and the lower one for 20 wt-%Ni (FactSage 6?4)
such as dust, or even the ability of the feeding systems
and weightometers to maintain constant proportions for
all these streams. One question is then ‘What are the
Table 2 Range of impurities in crude ferronickels and
impact on ferronickel liquidus temperature
(FactSage 6?4)
Typical range/wt-% 0.01–3 0.01–2 ,0.1 0.2–0.6 0.01–3
DT/uC per 1
wt-% increase
consequences of such variations?’ The impact of
the extent of iron reduction on the levels of carbon,
chromium and silicon in the ferronickel produced has
been discussed at length in previous papers by the
current authors (Solar et al., 2008, 2014). The relationship for carbon is represented in Fig. 3 for ease of
reference. This relationship is quite important as carbon
significantly decreases the liquidus temperature of ferronickel. Silicon has a similar effect, but to a lesser extent
as summarised in Table 2 and Fig. 4. The liquidus temperatures were calculated using the FactSage 6?4 thermodynamic package (Bale et al., 2002; Pelton and
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
Solar and Mostaghel
Laterite Smelting Challenges
5 Overall impact of extent of iron reduction on liquidus temperature of ferronickel. Ferronickel grades and liquidus temperatures of 18, 20, and 22 wt-%Ni ferronickels are marked
Chartland, 2000). Note that at the nickel grades
illustrated in Fig. 4 (15 and 20 wt-%Ni), a 5 wt-%
increase in nickel concentration lowers the liquidus
temperature of the ferronickel by 10–15uC, as addition
of nickel to pure iron depresses its liquidus temperature
from 1538uC down to a minimum of about 1440uC at
60–80 wt-%Ni (Sachs, 1960; Swartzendruber et al., 1991;
Caccianami et al., 2006).
The overall impact of the extent of iron reduction on
the liquidus temperature and grade of the ferronickel
produced is summarised in Fig. 5. The curves shown are
calculated for an ore at 1?8 wt-%Ni, 16?5 wt-%Fe and a
SiO2/MgO ratio of 2?3 (referred to as Project X
hereafter), typical of many lower quality laterites that
are now under consideration for development. As seen,
the ferronickel liquidus temperature is reasonably
constant at around 1430–1440uC in the low-iron
reduction range of 15–30% while it drops slightly from
1290 to 1260uC in the high iron reduction range of 45–
65%. However, between 30 and 45% iron reduction, the
liquidus temperature of the ferronickel produced drops
very rapidly from 1440 to 1290uC. If a ferronickel grade
of 20 wt-%Ni was targeted for such an ore (requiring an
iron reduction of 40%), the expected fluctuations in
grade, from 18 to 22 wt-%Ni, could be caused by iron
reductions fluctuating between 36 and 45% and, hence,
ferronickel liquidus temperatures could vary from 1300
to 1425uC. If the furnace power, secondary voltage and
electrode current were selected to yield a ferronickel
operating temperature of 1440uC (53uC superheat for
ferronickel at 20 wt-%Ni), the actual ferronickel superheat would vary from 140uC for ferronickel at 18 wt%Ni down to 15uC for ferronickel at 22 wt-%Ni. These
variations result mainly from the changes in carbon
content of the ferronickel produced, which increases
from 0?3 wt-% at 22 wt-%Ni to 1?8 wt-% at 18 wt-%Ni.
Such drastic changes in ferronickel liquidus temperature
would have significant impacts on operations, from
increased build-up on the tapping launders at lower
superheats to shortened tap block life at higher superheats. Downstream operations will also be affected; for
example, shotting of ferronickel after refining requires a
superheat of about 160uC to be carried out safely
without steam explosions (Zamallos et al., 2009). The
18 wt-%Ni ferronickel would already have a superheat
of 140uC when leaving the furnace, but the 22 wt-%Ni
ferronickel would require extensive heating because
its superheat would be only 15uC. Of course, these
fluctuations would be dampened somewhat by the large
metal inventory in the furnace. However, the drastic
drop in ferronickel liquidus temperature that could be
experienced when operating in the 30–45% iron reduction range may explain why no commercial smelter
currently operates in this region on a steady-state basis.
Protection of sidewalls
Another consideration in selecting the ferronickel grade
for a new project is the conditions that will be faced by
the furnace sidewalls, particularly in the tidal zone at the
slag/ferronickel interface. Modern ferronickel furnaces
are equipped with extensive sidewall cooling (e.g. Nelson
et al., 2004; Walker et al., 2009) and depend on frozen
slag to protect the refractory that remains once thermal
equilibrium has been reached on the walls. The tidal
zone is defined as the area that sequentially faces
ferronickel and then slag as ferronickel is tapped
intermittently. Protection of this area is essential as it
is typically where the sidewalls suffer the fastest
refractory wear. This problem is especially serious when
the operating temperature of the ferronickel is higher
than the liquidus temperature of the slag: any slag frozen
on the sidewall would then be re-melted when the metal
inventory is rebuilt after a tap. Figure 6 illustrates the
conditions experienced at two well-established smelters,
CMSA and Société Le Nickel (SLN), and those possible
for the hypothetical Project X. These conditions are
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
Solar and Mostaghel
Laterite Smelting Challenges
6 Liquidus temperatures and operating temperatures under various conditions. Ferronickel temperature on left side of
each column and slag temperature on right side
detailed in Table 3 together with ferronickel and slag
superheats, temperature differences and the freeze lining
SLN is well known for smelting the best saprolite
ore available: high nickel grade (.2?6 wt-%), low iron
(12?5 wt-%) and excellent Fe/Ni and SiO2/MgO ratios
(4?7 and 1?7, respectively). It is a high-reduction smelter
with typical slag and ferronickel temperatures of 1600
and 1500uC and liquidus temperature of 1536 and
1310uC, respectively (Solar and Davenport, 2014). Its
freeze lining margin, defined as the difference between
the slag liquidus temperature and the ferronickel
operating temperature is 36uC, assuring that slag frozen
on the tidal zone of the sidewalls remains frozen when
submerged in ferronickel. Cerro Matoso S.A. also had
a high nickel grade ore of .2?2 wt-%, medium iron
concentration of around 15 wt-%, and a medium Fe/Ni
ratio of 6?8, but the highest SiO2/MgO ratio of all
saprolites at 2?8, before this ore ran out almost 2 years
ago and CMSA started to smelt a lower grade ore with
an SiO2/MgO ratio of 1?8. It was a low reduction smelter
with typical slag and ferronickel temperatures of 1550
and 1460uC, and liquidus temperature of 1418 and
Table 3 Liquidus temperatures and operating temperatures (uC) under various conditions
Iron reduction (%)
FeNi grade/wt-%Ni
Operating temperature
Liquidus temperature
Operating temperature
Liquidus temperature
Slag – ferronickel
Freeze lining margin
Project X
1435uC, respectively. Its freeze lining margin was
negative at 242uC. Therefore, special attention had to
be paid to cooling of the tidal zone (Voermann et al.,
2004). Two cases are shown for Project X, for low and
high iron reductions of 25 and 55% as achieved at
CMSA and SLN, respectively. Project X has a SiO2/
MgO ratio of 2?3, meaning that its slag composition falls
right in the middle of the eutectic trough in the ‘FeO’–
SiO2/MgO pseudo-binary phase diagram (see Fig. 7).
Several points are worth noting at this stage:
1. The hypothetical Project X would yield very
different ferronickel grades than CMSA and SLN
at the levels of iron reduction targeted in these two
plants as a consequence of the different Fe/Ni
ratios in the various ores. Operators are well aware
of this dependence, but many project developers
nevertheless select their ferronickel grade based on
perceived market demand or a desire for ‘free’ iron,
whether or not this choice results in a proven extent
of iron reduction for the deposit under consideration. As explained above, a 20 wt-% Ni target for
Project X would put the electric furnace in a
potentially unstable operating window.
2. The difference between the operating slag and
ferronickel temperatures is about the same for these
two very different operations: 100uC for SLN and
90uC for CMSA, which is typical of a steady-state
operation. This is the reason why this parameter was
maintained at 90uC for the Project X calculations.
3. Cerro Matoso S.A. has a high slag superheat
(132uC) and a low ferronickel superheat (26uC).
The reverse is true for SLN (64 and 190uC,
respectively). Both these trends were maintained
for the Project X calculations.
4. The freeze lining margin for SLN is 36uC, but the
ferronickel superheat is 190uC. With such a high
ferronickel superheat, the refractory would be
quickly penetrated by ferronickel if it was not
protected by a frozen slag layer. On the other hand,
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
Solar and Mostaghel
Laterite Smelting Challenges
7 Estimate of slag liquidus temperatures for ‘FeO’–SiO2–MgO system (FactSage 6?4)
CMSA has a negative freeze lining margin but its
ferronickel superheat is only 25uC. At such a low
superheat, ferronickel does not penetrate very far
into the refractory and the tidal zone can survive a
long time without being covered all the time with a
frozen slag layer, especially when protected with
powerful waffle coolers (Voermann et al., 2004).
The hypothetical Project X developer, therefore, has to
choose between targetting ferronickels at 28 or 15 wt%Ni in order to fall within the 15–30% or 45–65% iron
reduction ranges. Considering that the SiO2/MgO ratio
of the deposit under study falls within the eutectic
trough of the ‘FeO’–SiO2/MgO pseudo-binary system,
this may be a case where stability and longevity of
furnace operation would lead to selection of a high iron
reduction, irrespective of what the economically optimum grade might be. The consequences of such a
selection would be:
a. limit the amount of FeO in the slag, thus increasing
its liquidus temperature
b. increase the amount of carbon in the ferronickel
produced, hence decreasing its liquidus temperature
and the required operating temperature
c. create a positive freeze lining margin, allowing a
permanent frozen slag layer in the slag/ferronickel
tidal zone.
A commonly used method to adjust the liquidus
temperature of many different types of slag is fluxing.
Figure 7 appears to point in two different directions:
add MgO to move up to the left-hand side of the trough
or add SiO2 to move to the right-hand side of the
eutectic. However, magnesite (or dolomite) is expensive,
and no commercial ferronickel smelter adds magnesium
carbonate or magnesia as a flux to its electric furnaces.
On the other hand, high SiO2 acidic slags are aggressive
on unprotected refractories (Daenuwy et al., 1992) and
no commercial laterite operation adds silica deliberately.
That is not to say that high MgO and high SiO2 slags
cannot be processed effectively on a commercial scale:
Loma de Niquel (LdN) and CMSA are examples of
successful operations at these two extremes, with SiO2/
MgO ratios of 1?3 and 2?8, respectively (Warner et al.,
2006). However, the high MgO or SiO2 contents in these
operations occur naturally in the ores smelted. This
leaves limestone as the only practical flux for ferronickel
operations. This practice was quite common in earlier
times, but the development of advanced sidewall cooling
has eliminated this approach in most modern smelters.
PT Aneka Tambang, on the island of Sulawesi in
Indonesia, is a good example of this development:
limestone is still used on their FeNi-1 furnace, but is no
longer added to the two modern FeNi-2 and FeNi-3
furnaces designed by Hatch (Bergman, 2003; Nelson
et al., 2007). Pobuzhsky Ferronickel Combine (PFK) in
Ukraine still uses limestone, and this practice is
particularly favoured by Chinese ferronickel and NPI
producers (Guo, 2009). Larco in Greece also has high
levels of CaO in its slags, but these high levels come from
the ores, which assay from 2 to 6 wt-%CaO, compared
to ,1 wt-% for normal laterites (Zevgolis, 2004).
The influence of CaO on slag viscosity and liquidus
temperature has been previously studied by several
authors (Chen et al., 2004, 2005; Somerville et al., 2004;
Jak and Hayes, 2010). Figure 8 shows the impact of
Al2O3 and CaO on the liquidus temperatures of magnesium silicate slags containing 20 wt-%FeO, calculated
using FactSage 6?4. The dramatic effect of Al2O3 on slag
is evident: the liquidus temperature is lowered significantly, especially on the right-hand side of the eutectic,
and the eutectic itself is moved to higher SiO2/MgO
ratios. Addition of CaO to the system further accentuates these trends; however, it increases the liquidus
temperature of the system at intermediate SiO2/MgO
ratios. As an example, the effect of limestone addition to
the calcine of Project X is shown in Fig. 9. As seen,
addition of 10 wt-% limestone increases slag liquidus
temperature by 30–40uC depending on the grade of
ferronickel assumed (which defines the FeO content of
the resulting slag). The main question is: ‘Is it worth it?’
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
Solar and Mostaghel
Laterite Smelting Challenges
8 A Influence of Al2O3 on liquidus temperature of slag at 20 wt-% FeO; B Influence of CaO on liquidus temperature of
slag at 20 wt-%FeO and 6 wt-%Al2O3 (FactSage 6?4)
There are several reasons to avoid limestone additions,
besides the obvious extra operating costs. Addition of,
for example, 10% flux dilutes the feed by 10% and
reduces the nickel throughput in a given smelter by
about 6%, as most smelters are furnace limited (Solar et
al., 2008). This 6% reduction is the minimum experienced. If the limestone is not completely calcined in the
reduction kiln, the decrease in throughput is greater
than 6%. However, limestone requires a temperature
of at least 900uC to be calcined effectively within a
reasonable length of time. That is the temperature at which DGu for the reaction CaCO3 5 CaO z
CO2 is zero. Further, the residence time required at
1000uC is a minimum of 30 min depending on the source
of the limestone (Moffat and Walmsley, 2006; Muazu
et al., 2011). Consequently, industrial lime producers
use temperatures in excess of 1000uC (Metso Minerals
Industries, 2014). The composition of the difficult
laterite ores discussed here typically limits the reduction
kilns to calcine temperatures of around 750uC. This is
because of the presence of significant amount of goethite
FeOOH and free silica in these ores. When such a
material is reduced, the resulting FeO combines with the
free silica to form end-members of the olivine solid
solution, mainly fayalite (Fe2SiO4). These compounds
have a minimum melting point of 1177uC, sintering
of which starts at about 800uC (Daenuwy and Dalvi,
1997). Bulk calcine temperatures are, therefore, limited to
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
Solar and Mostaghel
Laterite Smelting Challenges
9 The effect of limestone additions on slag liquidus temperature of Project X, producing 15 and 20 wt-%Ni ferronickel
(FactSage 6?4)
about 750uC. Consequently, limestone is not decomposed
in these kilns, and the calcines produced have high losson-ignition (LOI) contents. Such calcines can lead to
serious slag foaming episodes when fed to electric
furnaces in which insufficient charge banks do not allow
for complete calcination of the limestone before it enters
the slag phase. Therefore, the smelters that are still using
limestone have to run their kilns with an oxidising
atmosphere (which permits higher temperatures) and
increased residence times to allow for sufficient decomposition of the calcium carbonate. These requirements
decrease the ore throughput that would otherwise be
possible and increase the energy requirements on both
kilns and furnaces.
Considering that the furnace sidewalls can effectively
be protected with modern cooling systems, it is not
surprising that most smelters have abandoned this
fluxing practice.
Phosphorus distribution
Phosphorus is an impurity that has recently become a
concern for smelters, as some of the new deposits under
consideration contain significant amounts of this element. Traditional ores typically assay ,0?005 wt-%
phosphorous and are not analysed for this element on
a routine basis. Nor are the slags tapped from the
electric furnaces. The ferronickels tapped are analysed
for phosphorous, and a de-phosphorisation step is
included in the refining process if they contain more
than the 0?02–0?03 wt-% phosphorous quoted as maximum in the ISO ferronickel specifications (ISO 6501,
1988). To the best knowledge of the current authors, no
detailed information has ever been published on the
distribution of phosphorus in commercial ferronickel
plants. Bergman (2003) quoted phosphorous assays for
some electric furnace ferronickels before refining, but
not the corresponding slag and ore (or calcine) assays.
Otherwise, the information published refers to the
refining process (e.g. Zamallos et al., 2009) or refined
ferronickels (e.g. Warner et al., 2006). Pagador et al.
(1999) have published data on the results of laboratory
equilibration experiments between nickel alloys and
MgO-saturated or CaO-fluxed iron–silicate slags. These
authors reported their results in terms of distribution
ratios, defined as Lx5(% X)/[% X] (where parentheses
represent concentrations in slag and brackets represent
that in ferronickel); whereas, the industry usually refers
to partition coefficients, the inverse of the distribution
ratios. The phosphorous data of Pagador et al. (1999)
show an increase in partition coefficients from 0?001 to
0?2, as the oxygen potential in their systems was
decreased from 1027 to 1029 atm. However, Solar and
Davenport (2014) have shown that commercial ferronickel furnaces are not at thermodynamic equilibrium.
Table 4 quotes industrial data that have recently been
summarised at Hatch, Canada. The figures shown are
monthly averages except in the case of the fourth set for
Plant B, which gives the averages for eight individualrefining heats. It is evident that the partition coefficients
for phosphorous in commercial plants range from 8 to
14, several orders of magnitude larger than the 0?001–
0?2 values measured by Pagador et al. (1999) in
laboratory equilibration tests. Therefore, a different
approach must be developed to explain the behaviour of
phosphorus in commercial furnaces.
The approach proposed here is similar to that used by
Jones et al. (2009) to explain the recovery of nickel,
cobalt, chromium and platinum group metals (PGM)
from laterites and other feeds. This mechanistic model is
based on thermodynamic principles, correlating the
reduction of the element under study to that of iron in
the same system. The reactions can be written as
FeO~Feo z0:5O2
PO2:5 ~Po z1:25O2
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
Solar and Mostaghel
Laterite Smelting Challenges
10 Nickel and phosphorus yields v. that of iron. Calibrated mechanistic model v. actual
from which it can be derived
~KFe cFe
RP ~
1z KPcP
where the Ks are the equilibrium constants for the
reduction reactions as written, the Rs are the proportions of iron and phosphorous reduction in the system,
and cFe ~cFeO =cFe0 , while cP ~cPO2:5 =cP0 :
Combining equations (3) and (4) yields
RP ~
(KFe cFe )2 5
K P cP
Equation (5) can then be calibrated against commercial
data by determination of the A-factor using the leastsquares method. Figure 10 shows the results of such a
calibration, using all of the data point quoted in Table 4.
The best fit was obtained for A-factor equal to 0?2.
Many of the new nickel laterite resources that are
currently under consideration for development present
serious challenges because of factors such as high iron
contents, high SiO2/MgO ratios, and high concentrations of Al2O3, Cr2O3 and even phosphorous. This is
why the authors referred to them as ‘difficult’. Selection
of the ferronickel grade that will be targeted by such
projects cannot be based on perceived market preferences or a desire for ‘free’ iron units. A holistic approach
must be used that balances the preferences of the
developer with the economical optimum for the project
and what is metallurgically feasible.
Particular attention must be paid to the level of iron
reduction and the difference between the slag liquidus
Table 4 Phosphorus distribution in commercial ferronickel plants
Plant A
Plant B
Data span
1 month
1 month
1 month
1 month
8 heats
3 months
Wt-%Ni in crude ferronickel
Percent yield to ferronickel
in electric furnace
Partition coefficient
Crude FeNi
P on inputs
P on outputs
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
Plant C
Solar and Mostaghel
temperature and ferronickel operating temperature. The
level of iron reduction is important as it controls the
amount of carbon reporting to the ferronickel phase
and, consequently, its liquidus temperature. There is a
drastic drop in this liquidus temperature from y1440uC
to less than 1290uC in the 30–45% iron reduction range
because of an increase in carbon content from y0?1 to
y2 wt-%. Such a significant dependence of the liquidus
temperature on the extent of iron reduction could lead
to serious operating instabilities, which may explain why
no commercial plant operates in that reduction range.
The difference between slag liquidus temperature and
the ferronickel operating temperature dictates whether
the tidal zone at the slag/ferronickel interface is
protected by freeze lining or not and, consequently,
the design of the cooling systems in that area of the
furnace sidewalls. Some nickel smelters, especially exSoviet and Chinese operations, flux with limestone as
remediation for the difficult slags that may be experienced. However, fluxing with limestone presents its own
challenges: it reduces the nickel concentration of the feed
material (already low-grade ores), is difficult to calcine
properly at the kiln temperatures that must be maintained, and its use may hence lead to serious foaming
problems in the electric furnace. Modern ferronickel
smelters have abandoned this practice, as the sidewall
cooling systems now available have made it redundant.
In addition, these ‘difficult’ ores often also contain
significant quantities of phosphorus, the distribution of
which between ferronickel and slag can be calculated
using a calibrated mechanistic model.
The authors wish to thank Hatch Ltd for their support
and collaboration during the performance of this review.
The comments of Dr Bert Wasmund on the manuscript
are gratefully appreciated.
Bale, C. W., Chartrand, P., Degterov, S. A., Eriksson, G., Hack, K.,
BenMahfoud, R., Melanc, J., Pelton, A. D. and Petersen, S. 2002.
FactSage thermodynamical software and databases, Calphad., 26,
(2), 189–228.
Bergman, R. A. 2003. Nickel production from low-iron laterite ores:
process descriptions, CIM Bull., 95, (1072), 127–138.
Bolt, J. R. Jr. 1967. The winning of nickel – its geology, mining and
extractive metallurgy, (Technical Editor P. Queneau), Toronto,
Longmans Canada Ltd.
Bunjaku, A., Kekkonen, M., Taskinen, P. and Holappa, L. 2011.
Thermal behaviour of hydrous nickel-magnesium silicates when
heating up to 750uC, Trans. Inst. Min. Metall. C, 120, (3), 139–
Bunjaku, A., Kekkonen, M., Pietilä, K. and Taskinen, P. 2012a. Effect
of mineralogy on reducibility of calcined nickel saprolite ore by
hydrogen, Trans. Inst. Min. Metall. C, 121, (1), 16–22.
Bunjaku, A., Kekkonen, M., Pietilä, K. and Taskinen, P. 2012b. Effect
of mineralogy and reducing agent on reduction of saprolitic nickel
ores, Trans. Inst. Min. Metall. C, 121, (3), 156–165.
Caccianami, G., De Keyzer, J., Ferro, R., Klotz, U. E., Lacaze, J. and
Wollants, P. 2006. Critical evaluation of the Fe–Ni, Fe–Ti and
Fe–Ni–Ti alloy systems, Intermetallics, 14, 1312–1325.
Chen, S., Erasmus, L., Barnett, S. C. C., Jak, E. and Hayes, P. C. 2004.
Liquidus temperatures of ferronickel smelting slags, International
Laterite Nickel Symposium-2004, North Carolina, USA, (eds.
W. P. Imrie and D. M. Lane), 503–518, Warrendale, TMS.
Chen, S., Jak, E. and Hayes, P. C. 2005. Phase equilibria in the
cristobalite, tridymite and pyroxene primary phase fields in the
MgO-‘FeO’-SiO2 system in equilibrium with metallic iron, ISIJ
Int., 445, (6), 791–797.
Laterite Smelting Challenges
Crundwell, F. K., Moats, M. S., Ramachandran, V., Robinson, T. G.
and Davenport, W. G. 2011. Extractive metallurgy of nickel,
cobalt and platinum-group metals, 21–93, Oxford, Elsevier.
Daenuwy, A. and Dalvi, A. D. 1997. Development of reduction kiln
design and operation at PT Inco (Indonesia), Proc. Nickel-Cobalt
97 International Symposium, Sudbury, ON, Canada, (eds. W. C.
Cooper, I. Mihaylov and C.A. Levac), 93–113, Metallurgical
Daenuwy, A., Dalvi, A. D., Solar, M. Y. and Wasmund, B. 1992.
Development of electric furnace design and operation at PT Inco
(Indonesia), Proc. Int. Symp. on Non-Ferrous Pyrometallurgy:
Trace Metals, Furnace Practices and Energy Efficiency, 31st
Conference of Metallurgists Metallurgical Society, Edmonton,
Alberta, 223–244.
Dalvi, A. D., Bacon, W. G. and Osborne, R. C. 2004. The past and
the future of nickel laterites, International Laterite Nickel Symposium-2004, (eds. W. P. Imrie and D. M. Lane), TMS, 2004, 23,
full text available on line at: http://www.pdac.ca/pdf-viewer?doc5/
Elias, M. 2002. Nickel laterite deposits–geological overview, resources
and exploitation, Giant Ore Deposits: Characteristics, Genesis
and Exploration, (eds. D. R. Cooke and J. Pongratz), 205–220,
Hobart, University of Tasmania, CODES Special Publication No.
Guo, X. J. 2009. Chinese nickel industry–projects, production and
technology, Pyrometallurgy of Nickel and Cobalt 2009, (eds J.
Liu, J. Peacy, M. Barati, S. Kashani-Nejad and B. Davis),
Metallurgical Society, 3–21, Canadian Institute of Mining,
Metallurgy and Petroleum (CIM); Montreal, Canada
Halikia, I, Skartados, K. and Neou-Syngouna, P. 2002. Effect of
reductive roasting on smelting characteristics of Greek nickel
laterites, Trans. Inst. Min. Metall. C, 111, (3), 135–142.
International Standard. Ferronickel – specification and delivery requirements, ISO 6501, 1988.
Jak, E. and Hayes, P. C. 2010. Slag phase equilibria and viscosities in
ferronickel smelting slags, The 12th International Ferroalloy
Congress – Sustainable Future, Helsinki, Finland, June 2010,
631–639, The South African Institute of Mining and Metallurgy
Jones, R. T., Geldenhuys, L. J. and Reynolds, Q. G. 2009. Recovery of
base metals and PGMs in a DC alloy-smelting furnace, J. S. Afr.
Inst. Min. Metall., 109, 587–592.
Kobayashi, Y., Todoroki, H. and Tsuji, H. 2011. Melting behavior of
siliceous nickel ore in a rotary kiln to produce ferronickel alloy,
ISIJ Int., 51, (1), 35–40.
Makinen, T. and Taskinen, P. 2008. State of the art in nickel smelting:
direct Outokumpu nickel technology, Trans. Inst. Min. Metall. C,
117, (2), 86–94.
Metso Minerals Industries. 2014. Preheater-Kiln Lime Calcining
Systems, Brochure 1330-03-02 MPR/Danville, Available: http://
Moffat, W. and Walmsley, M. R. W. 2006. Understanding lime
calcination kinetics for energy cost reduction, Presented at the 59th
Appita Conference, Auckland, New Zealand, May 2006, Available:
Mostaghel, S., Matsushita, T., Samuelsson C., Björkman, B.,
Seetharaman, S. 2013. ‘‘Influence of alumina on physical
properties of an industrial zinc-copper smelting slag-Part 3:
Melting behaviour’’, 122, (1), 56–62.
Muazu, K., Abdullahi, M. and Akuso, A. S. 2011. Kinetic study of
calcination of Jakura limestone using power rate law model,
Niger. J. Basic Appl. Sci., 19, (1), 116–120.
Mudd, G. M. 2009. Nickel sulfide versus laterite; the hard sustainability
challenge remains, 48th Annual Conference of Metallurgists
(COM2009), Sudbury, 1–10, Canadian Institute of Mining,
Metallurgy and Petroleum (CIM); Sudbury, Canada.
Nelson, L. R., Geldenhuis, J. M. A., Miraza, T., Badrujaman, T.,
Hidayat, T., Jauhari, I., Stober, F. A., Voermann, N., Wasmund,
B.O. and Jahnsen, E. J. M. 2007. Role of operational support in
ramp-up of the FeNi-II furnace at PT Antam in Pomalaa, 11th
International FerroAlloy Congress (Infacon XI), February 2007,
New Delhi, India.
Nelson, L. R., Sullivan, R., Jacobs, P., Munnik, E., Lewrance, P, Roos,
E, Uys, M.J.N., Salt, B., de Varies, M., McKenna, K., Voermann,
N. and Wasmund, B.O. 2004. Application of a high-intensity
cooling system to DC-arc furnace production of ferrocobalt at
Chambishi, 10th International Ferroalloys Congress, Cape Town,
South Africa.
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
Solar and Mostaghel
Laterite Smelting Challenges
Pagador, R. U., Hino, M. and Itagaki, K. 1999. Distribution of minor
elements between MgO saturated FeOx–MgO–SiO2 or FeOx–
CaO–MgO–SiO2 slag and nickel alloy, Mater. Trans., JIM, 40,
(3), 225–232.
Pelton, A. and Chartland, P. 2000. The modified quasi-chemical model part
II. Multicomponent solutions, Metall. Trans. B, 33B, 1355–1360.
Rhamdhani, M. A., Hayes, P. C. and Jak, E. 2009a. Nickel laterite part 2 –
thermodynamic analysis of phase transformations occurring during
reduction roasting, Trans. Inst. Min. Metall. C, 118, (3), 146–155.
Rhamdhani, M. A., Hayes, P. C. and Jak, E. 2009b. Nickel laterite;
part1: microstructure and phase characterisations during reduction roasting and leaching, Trans. Inst. Min. Metall. C, 118, (3),
Sachs, G. 1960. Fe–Ni iron-nickel, in Constitution of binary alloys,
metals handbook, (ed. T. Lyman), 1211, American Society for
Metals, Ohio, the USA.
Solar, M. Y., Candy, I. and Wasmund, B. 2008. Selection of the
optimum ferronickel grade for smelting nickel laterites, Executive
Summary in CIM Magazine, 3, 2, 2008, 74, Full paper in CIM
Bulletin, Vol. 11, No 1107.
Solar, M. Y. and Davenport, W. G. 2014. Are ferronickel furnaces at
thermodynamic equilibrium, Proc. 53rd Conf. of Metallurgists –
COM 2014, Vancouver, BC, Canada, 2014, Accepted for
Solar, M. Y., Mostaghel, S. and Nicol, S. 2014. Some insights into the
process design of ferronickel smelters, Proc. 53rd Conf. of
Metallurgists – COM 2014, Vancouver, BC, Canada, 2014,
Accepted for publication.
Somerville, M., Wright, S., Sun, S. and Jahanshahi, S. 2004. Liquidus
temperatures and viscosities of melter slags, VII International
Conference on Molten Slags, Fluxes and Salts, 219–224,
South African Institute of Mining and Metallurgy (SAAIMM),
Johannesburg, South Africa.
Swartzendruber, L. J., Itkin, V. P. and Alcock, C. B. 1991. The Fe–Ni
(iron–nickel) system, J. Phase Equilib., 12, (3), 288–312.
Swinbourne, D. R. 2014. Understanding ferronickel smelting from
laterites through computational thermodynamics modelling,
Trans. Inst. Min. Metall. C, Submitted, 123(3), 127–140.
Voermann, N., Gerritsen, T., Candy, I., Stober, F. and Matyas, A.
2004. Developments in furnace technology for ferronickel
production, INFACON X, 455–465, The South African
Institute of Mining and Metallurgy (SAAIMM), Cape Town,
South Africa.
Walker, C., Kashani-Nejad, S., Dalvi, A. D., Voermann, N., Candy,
I. M. and Wasmund, B. 2009. Future of rotary kiln – electric
furnace (RKEF) processing of nickel laterite, Proc. European
Metallurgical Conference (EMC2009), 1–30, the German Society
for Mining, Metallurgy, Resource and Environmental
Technology (GDMB); Innsbruck, Austria.
Warner, A. E. M., Diaz, C. M., Dalvi, A. D., Mackey, P. J. and
Tarasov, A. V. 2006. World nonferrous smelter survey, part III:
nickel: laterite, J. Met., 3, 11–20.
Zamallos, M., Reyes, A., Fañas, J. J. and Frias, J. R. 2009. New
developments in ferronickel refining at Xstrata Nickel’s Falcondo
Smelter, pyrometallurgy of Nickel and Cobalt 2009, (eds. J. Liu,
J. Peacy, M. Barati, S. Kashani-Nejad and B. Davis),
Metallurgical Society, 195–208, Canadian Institute of Mining,
Metallurgy and Petroleum (CIM); Montreal, Canada.
Zevgolis, E. N. 2004. The evolution of the Greek ferronickel production
process, International Laterite Nickel Symposium-2004, (eds.
W. P. Imrie and D. M. Lane), 619–632, TMS, Charlotte, VI.
Mineral Processing and Extractive Metallurgy (Trans. Inst. Min. Metall. C)
Copyright of Mineral Processing & Extractive Metallurgy: Transactions of the Institution of
Mining & Metallurgy, Section C is the property of Maney Publishing and its content may not
be copied or emailed to multiple sites or posted to a listserv without the copyright holder's
express written permission. However, users may print, download, or email articles for
individual use.
Random flashcards
Rekening Agen Resmi De Nature Indonesia

9 Cards denaturerumahsehat

Rekening Agen Resmi De Nature Indonesia

9 Cards denaturerumahsehat

sport and healty

2 Cards Nova Aulia Rahman


2 Cards oauth2_google_3524bbcd-25bd-4334-b775-0f11ad568091

Create flashcards