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IN THIS ISSUE

• Improved nickel and cobalt recovery from laterites

• Concentration of iron ore flotation tailings

• Solvent extraction of zirconium using triphenylphosphine oxide

• Full-scale continuous centrifugal concentrator

• Copper losses in smelting slag dependent on slag composition

• Organic binder for the pelletization of a magnetite concentrate

• Ammonia separation of Ni from spent fly ash leach liquor

• Flotation of zinc oxide minerals from low-grade tailings

• Design and operation of a teeter-bed hydroseparator

• Reverse flotation of diaspore from aluminosilicates

Volume 27, Number 3August 2010

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Volume 27, No. 3

August 2010

Improved nickel and cobalt recovery from nickeliferous laterites 117 in acidic fluoride media

Wei Luo and Jiankang Wen: General Research Institute for Nonferrous Metals, Beijing, ChinaDianzuo Wang and Qiming Feng: Central South University, Beijing and Changsha, China

Pellet feed production via the concentration of flotation tailings 124 M.G. Vieira: Samarco Mineração SA, Ouro Preto, MG, Brazil A.E.C. Peres: UFMG, Belo Horizonte, MG, Brazil

Solvent extraction of zirconium from zircon leach liquor using 129 triphenylphosphine oxide K. Saberyan and E. Zolfonoun: Nuclear Science and Technology Research Institute, Iran P. Vahedian-Donyaparast: Islamic Azad University, Tehran, Iran M. Noparast and A. Nemati: Tehran University, Tehran, Iran

Performance of a full-scale continuous centrifugal concentrator 133 in reconcentrating fine hematite from tailings Luzheng Chen, Siqing Liu and Shuming Wen: Kunming University of Science and Technology , Kunming, Yunnan Province, China

Copper losses in sulfide concentrate smelting slag are dependent 141 on slag composition Ž. Živkovic, I. Mihajlovic and Đ. Nikolic: University of Belgrade, Bor, Serbia N. Mitevska: Mining and Metallurgy Institute, Bor, Serbia

Laboratory study of an organic binder for pelletization of a magnetite 148 concentrate

S. L. de Moraes: Institute for Technological Research, Sao Paulo, BrazilS. K. Kawatra: Michigan Technological University, Houghton, Michigan

Ammonia separation of Ni from spent fly ash leach liquor 154N.-S. Kim: University of Texas at El Paso, El Paso, TexasS.-Y. Hwang: KEN Research Center, Seokyeong University, Seoul, Republic of Korea

Flotation of zinc oxide minerals from low-grade tailings by oxine 158 and dithizone using the Taguchi approach

A. Hajati: Iran University of Science and Technology, Arak, IranA. Khodadadi and S. M. Koleini: Tarbiat Modares University, Tehran, Iran

Design, operation and control of a teeter-bed hydroseparator for 166 classification J.N. Kohmuench, E.S. Yan and M.J. Mankosa: Eriez Manufacturing, Erie, Pennsylvania G.H. Luttrell and R.C. Bratton: Virginia Tech, Blacksburg, Virginia

Reverse flotation of diaspore from aluminosilicates by a new 173 cationic organosilicon quaternary ammonium collector

Xinyang Yu, Hong Zhong and Guangyi Liu: Central South University, Changsha, China

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MINERALS&METALLURGICALPROCESSING(ISSN0747-9182)ispublishedquarterlybytheSocietyforMining,Metallurgy,andExploration,Inc.,at8307ShafferParkway,Littleton,CO,80127–4102. Phone: 1–800–763–3132 or 303–948–4200. Fax:303–973–3845 or e-mail: [email protected]. Website: www.smenet.org.PostagepaidatLittleton,Colorado,andadditionalmailingoffices.Canadianpost:publicationsmailagreementnumber 0689688. POSTMASTER: Send address changes toMINERALS & METALLURGICAL PROCESSING, 8307 ShafferPkwy,Littleton,CO,80127.Disclaimer:SMEisnotresponsibleforanystatementmadeoropinionexpressedinitspublications.Subscriptions:Annualmembersubscriptionrates:print+onlineUS$119,onlineonly$US99.Annualnonmemberindividualrates:print + online US$149 domestic and US$159 foreign; onlineonlyUS$129.Library/institutionalrate:US$599printandprint+online(whereavailable).Single copies:US$35.00(memberandnonmember individualsubscribersonly).Reproduction:MorethanonephotocopyofanitemfromSMEmaybemadefor internal use, provided that fees are paid directly to theCopyrightClearanceCenter,222RosewoodDr.,Danvers,MA,01923, USA (Phone: 978-750-8400; Fax: 978-646-8600). AnyotherformofreproductionrequiresspecialpermissionfromSMEandmaybesubjecttofees.Copies of articles: CopiesofpreviouslypublishedpapersareavailablefromtheLindaHallLibrary,KansasCity,MO(www.lhl.lib.mo.us)Olderpapersmayalsobeavailableatwww.OneMine.org.

August 2010 • Vol. 27, No. 3 ISSN 0747–9182

MINERALS & METALLURGICAL PROCESSING Vol. 27 No. 3 • August 2010117

Improved nickel and cobalt recovery from nickeliferous laterites in acidic fluoride mediaWei Luo National Engineering Lab of Biohydrometallurgy, General Research Institute for Nonferrous Metals, Beijing, P.R.China, E-mail: [email protected]

Dianzuo WangChinese Academy of Engineering, Central South University and General Research Institute for Nonferrous Metals, Beijing, P.R.China

Qiming FengProfessor, Department of Mineral Processing, Central South University, Changsha, China

Jiankang WenProfessor, National Engineering Lab of Biohydrometallurgy, General Research Institute for Nonferrous Metals, Beijing P.R.China

AbstractNickeliferous laterites were characterized by particle size analysis, X-ray fluorescence (XRF), X-ray diffraction (XRD), scanning electron microscope (SEM) and energy dispersive X-ray analysis (EDX), Fourier transform infrared (FTIR) spectroscopy and thermal analysis. Results showed that the laterites consist mainly of nickel-substituted lizardite with cobble-like morphology and traces amounts of magnetite and phlogopite. As fluoride ions can react with silicon in the lizardite crystal structure and form metal-fluoride complexes, as well as increase the equilibrium constant of the lizardite dissolution in acid solution dramatically, the addition of minor fluoride salts (0.2wt.% NaF) gave a markedly enhanced metal extraction rate (Ni 81.7% and Co 52.3%) under the experi-mental conditions of particle size d50= 25 μm, 10% (v/v) H2SO4, reaction time 2 h, reaction temperature 90°C, liquid-to-solid ratio 3 mL g-1 and stirring at 500 rpm compared with H2SO4 alone (Ni 69.6% and Co 30.8%).

Paper number MMP-09-020. Original manuscript submitted May 2009. Revised manuscript accepted for publication De-cember 2009. Discussion of this peer-reviewed and approved paper is invited and must be submitted to SME Publications Dept. prior to February 28, 2011. Copyright 2010, Society for Mining, Metallurgy, and Exploration, Inc.

IntroductionNickeliferous laterites host most of the world’s ter-

restrial nickel resources, comprising about 60% of the total 130 million tons of nickel reserves (Kempthorne and Myers, 2007). With the continuous depletion of high-grade nickel sulfide ores (Tang and Valix, 2006), there is the need to win metals from the abundant low-grade nickel laterite ores. However, commercial application of nickel laterite processing methods has been minimally successful, due to a number of political, technical, mineralogical, geographical and economic reasons, such as poor metal recovery (Agatzini-Leonardou and Zafiratos, 2004; Canterford, 1975; Onodera et al., 1987; Rubisov et al., 2000).

Nickeliferous laterites experience poor recovery, as they are difficult to dissolve in ordinary mineral acid solutions at atmospheric pressure. This difficulty is a result of nickeliferous laterites being residual

products that derive from a wide variety of rocks through intensive chemical weathering under strongly oxidizing and leaching conditions, which leads to the incorporation of nickel into the crystal structure of laterites (Rueda et al., 1996; Valix and Cheung, 2002).

Although metal extraction from laterites is an important step in the hydrometallurgical processing of nickel and cobalt and various aspects of the dissolution behavior of laterites have been investigated in recent decades (Briceno and Osseo-Asare, 1995; Caron, 1950; Curlook, 2004; Deepatana et al., 2005; Pickles, 2005; Whittington et al., 2003), only a limited number of studies have reported on metal recovery from nickeliferous laterites at atmospheric pressure in acidic fluoride media (Hi-rasawa, 1987; Hirasawa et al., 1983; Lloyd and Turner, 1984).

Many studies demonstrate that most oxide minerals can be decomposed by hot concentrated hydrofluoric acid (Comba et al., 1991; Hirasawa, 1987; Hirasawa, 1989; Hirasawa et al., 1983; Lei et al., 1991; Lloyd and Turner, 1984; Majima et al.,

Key words: Oxide minerals, Mineralogy, Leaching, Nickel/ nickel ores

August 2010 • Vol. 27 No. 3 MINERALS & METALLURGICAL PROCESSING118

1988). These studies indicate that the dissolution rates of metal oxides were greatly enhanced by the increase of hydrogen ion activity in the leachant. The presence of both H+ and F- in the leachant could break the crystal lattice of laterites. Meanwhile, nickel and cobalt could react with F- to form fluoro complexes, which are soluble in acid solution as stable species. However, hydrofluoric acid used as a leaching agent is a weak acid, with an ionization constant of 10-3.2 at 298.15K (Dean, 1998); thus, the H+ activity in the leachant is controlled by the ionization equilibrium of HF as long as the leaching solution contains HF alone. Therefore, it is hypothesized that the addition of concentrated sulfuric acid and soluble fluoride salts, which are necessary to form fluoro complexes, will achieve better efficiency compared with adding HF or H2SO4 alone (Lei et al., 1991; Singh and Mendenhall, 2004).

A large number of researchers have demonstrated that the addition of fluoride salts, which are necessary for the formation of soluble fluoro complexes and react with silicon in laterites’ crystal structure, together with sulfuric acid increase effective-ness in leaching metals from silicates (Comba et al., 1991;

Hirasawa, 1989; Hirasawa et al., 1983; Lei et al., 1991; Lloyd and Turner, 1984; Majima et al., 1988; Singh and Mendenhall, 2004). Fluoride ions prove to be the most active additives for nickel and cobalt extraction from laterites, because leaching by fluoride ions together with sulfuric acid is accomplished mainly by lattice destruction by hydrogen and fluoride ions, as well as the formation of metal-fluoride complexes.

As part of oxide dissolution studies, the leaching mechanism of complex oxide minerals in weak acid solution containing fluoride ions was examined. In this study, the mineralogical analysis of the nickeliferous laterites and the dissolution of nickel and cobalt from laterites at atmospheric pressure in acidic fluoride solution were investigated. In particular, the effect of different fluoride salts, such as NaF, NH4F and CaF2, on nickel and cobalt dissolution was studied.

ExperimentalMaterials. A low-grade nickeliferous laterite ore (from

Yunnan, China) was used. After drying at 105°C, the repre-sentative sample was crushed primarily using a jaw crusher followed by a roller crusher to -2 mm. To prepare the feed for leaching experiments, samples were attrition-scrubbed using vibration milling for 10 min.

Methods. A number of experimental conditions were evaluated in order to investigate the impact of different fluoride salts on nickel and cobalt dissolution. These conditions were:

- Sulfuric acid concentration- Reaction temperature- Liquid-to-solid ratio- Reaction time and stirring speed

The aqueous solutions of H2SO4 alone, H2SO4+NaF, H2SO4+NH4F and H2SO4+CaF2 mixtures were used as leachants for leaching experiments. For studying the leach-ing behavior of laterites in acidic fluoride media, 0.3 L of the leachant was put into a 0.5-L Teflon separatory flask immersed in an electric-heated thermostatic water bath. When the tem-perature of the leachant reached a predetermined constant level, a mineral sample of 100 g was charged into the leachant to commence the leaching experiment. Mechanical agitation during the experiment was done using a Teflon blade. Solution samples were periodically withdrawn for chemical analysis of nickel and cobalt by atomic adsorption spectroscopy (AAS) using standard procedures. Chemicals used in this study were all of reagent grade and all aqueous solutions were prepared with deionized water.

Characterization. Particle size analysis was performed using a Mastersizer 2000 particle size analyzer. The X-ray dif-fraction (XRD) pattern was detected on a D/Max-2550 X-ray diffractometer with CuKα radiation from 10 to 85°. Differential scanning calorimetry (DSC) and thermogravimetric (TG) curves were obtained using a NETZSCH STA 449C thermoanalyzer heated at 10°C min-1 in an Ar atmosphere. Calcined Al2O3 was used as an inert substance. Leica Cambridge Stereoscan S-440 Scanning Electron Microscope (SEM) by M/S Leica Cambridge Ltd UK was used for EDX analysis. Fourier trans-form infrared (FTIR) spectra were measured on a Nicolet 5700 spectrophotometer with KBr discs in the transmittance mode in the range 4000-400 cm-1. The chemical composition of the

Figure 1 — Effect of sulfuric acid concentration on the leaching of nickel and cobalt from laterites.

Figure 2 — Effect of reaction temperature on the leaching of nickel and cobalt from laterites.

Sulfuric acid concentration (% v/v)


sample was determined by X-ray fluorescence (XRF) on a Philips PW2424 spectrometer.

Results and discussionDerivation of chemical leaching methodologyEffect of sulfuric acid concentration. An important factor

to be studied is the effect of sulfuric acid concentration. The results plotted in Fig. 1 show that in general, acid concentra-tion has a pronounced effect on metal extraction. Increasing the concentration up to 10% (v/v) appears to increase nickel dissolution dramatically. Further increase in the concentration results in less effective nickel dissolution. Cobalt dissolution also supports this observation.

Effect of reaction temperature. In order to investigate the effect of leaching temperature, a series of experiments were car-ried out with the leaching temperature varied from 30 to 90°C. Results included in Fig. 2 indicate that reaction temperature has a noticeable effect on metal extraction. Keeping other reaction conditions constant, with increasing reaction temperature, the leaching rate improves. A leaching temperature of 90°C was used in all subsequent experiments.

Effect of liquid-to-solid ratio. The results of the liquid-to-solid ratio (varying from 1 to 3 mL g-1) on nickel and cobalt dissolution from laterites presented in Fig. 3 suggest that metal dissolutions are dependent on the liquid-to-solid ratio. About 69.6% of the nickel and 30.8% of the cobalt in the laterites can be leached at a liquid-to-solid ratio of 3 mL g-1.

Effect of reaction time and stirring speed. After several leaching experiments, it was indicated that the optimal metal extraction could be obtained after 2 h. The diffusion of reac-tants and products between the solution and the surface of the particle was fast and thus did not control the leaching within the range of stirring speeds investigated. A stirring speed of 500 rpm was used in all subsequent experiments.

Effect of fluoride salts on leaching laterites. A series of leaching experiments were performed to study the effect of fluoride salts on the dissolution of nickel and cobalt from lat-erites in dilute H2SO4 solution containing different amounts of fluoride salts. Experimental conditions for the investigation were those derived from previous chemical leaching results:

particle size d50=25 μm, 10% (v/v) H2SO4, reaction time 2 h, reaction temperature 90°C, liquid to solid ratio 3 mL g-1 and stirring at 500 rpm.

Hirasawa (Hirasawa et al., 1983) examined the effects of the addition of NaF into H2SO4 leaching solution on the extraction of Ni, Fe, and MgO from garnierite ores. The highest Ni extrac-tion (94.5%) was attained at an NaF addition of 0.02g/g ore. Results included in Table 1 indicate that the dissolution rates of nickel and cobalt increased considerably, especially for cobalt, even at NaF additions as low as 0.1%. An additional increase in NaF concentration to 0.2% results in the increase in the dis-solution rates of nickel and cobalt, although the total H2SO4 concentration in the solution is the same. This result suggests that increasing fluoride ion concentration in the leachants is an important way to improve dissolution. However, when the amount of NaF was higher than 0.2%, the dissolution of both nickel and cobalt decreased dramatically. From the yield of the leaching residues with the addition of different amounts of fluoride salts, the reason for the significant drop in leach recovery of Ni and Co beyond a certain fluoride concentration may be due to the coprecipitation of Ni2+ and Co2+ with MgF2 in the residues, as well as the adsorption of silicic acid formed in the leach solution.

Figure 3 — Effect of the liquid-to-solid ratio on the leaching of nickel and cobalt from laterites.

Table 1 — Effect of fluoride salts on the leaching of nickel and cobalt from laterites.

aContent, wt.%

%Ni leached %Co leached % Residue yield

NaF NH4F CaF2 NaF NH4F CaF2 NaF NH4F CaF2

0 69.6 69.6 69.6 30.8 30.8 30.8 71.8 71.8 71.8

0.1 81.6 77.1 71.7 50.4 41.9 51.3 70.1 73.2 73.6

0.2 81.7 77.4 72.6 52.3 43.8 51.2 71.4 73.0 73.5

0.3 81.7 78.0 74.9 51.2 44.2 52.8 71.8 72.5 73.1

0.4 72.7 77.0 74.4 47.9 42.3 50.0 70.6 73.4 73.8

0.5 67.7 76.9 69.8 37.2 41.3 41.4 68.8 72.7 68.8

1 57.9 65.1 62.6 27.0 30.4 29.4 69.1 74.9 78.5

5 52.4 59.7 60.8 25.8 28.8 28.6 78.1 82.7 84.1

10 39.5 40.2 58.4 18.0 21.2 24.7 83.1 85.6 87.0a Content refers to the mass ratio of fluoride salts to sample


Based on the equilibrium constants (Baes and Mesmer, 1976), Si(OH)4 is the predominant species in the acidic pH range. The silicic acid particles in the solution are character-ized by a very large surface area and are highly dispersed in the solution. On the surface are highly reactive groups, silicon, oxygen and hydrogen (Si-O-H). This large surface area and surface energy means that colloidal silicic acid is an excellent adsorbent, which takes up ions or molecules of gas or liquid. Whittington (Whittington and Johnson, 2005; Whittington et al., 2003) examined nickel losses in the leach residue by pressure acid leaching of arid-region laterite ore and found that cations like Ni2+ are adsorbed on amorphous silica, which has a strongly negative surface charge, even in the presence of Mg2+ cations. Hence, when the leach solution is cooled, silicic acid is precipitated from the solution. Moreover, the presence of magnesium sulfate in the solution as well as the low pH of the solution lowers the equilibrium concentration of silicic acid (Piryutko, 1959).

Leaching behavior of the laterites in acidic fluoride solution indeed indicated that, besides the free hydrogen ion action, fluoride ions provide an additional leaching agent, possibly through lattice destruction by fluoride ions, as well as the formation of metal-fluoride complexes. Mg3Si2O5(OH)4(S) + 6H+ =

3Mg2+ + 2H4SiO4(S) + H2O(I) (1)The Gibbs free energy change of lizardite dissolution in

sulfuric acid solution Eq. (1) at 363.15K is -173.88kJ mol-1 and its equilibrium constant is 1025.0, as shown in Table 2, so this reaction should be spontaneous. While the dissolution reaction occurs in acidic fluoride solution, the intermediate product hexafluorosilicic acid forms: Mg3Si2O5(OH)4(S) + 18H+ + 18F- =

3MgF2(S) + 2H2SiF6 + 9H2O(I) (2)Hexafluorosilicic acid then hydrolyzes to orthosilicic acid

and hydrofluoric acid: H2SiF6 + 4H20 = H4SiO4 + 6HF (3)

The Gibbs free energy change of the overall reaction with the precipitation of MgF2, Eq. (4), at 363.15K is -382.96kJ mol-1 and its equilibrium constant is 1055.1: Mg3Si2O5(OH)4(S) + 6H+ + 6F- =

3MgF2(S) + 2H4SiO4 + H2O(I) (4) This result means that the addition of fluoride ions has a

significant influence on the lizardite dissolution reaction in acid solution. Moreover, nickel and cobalt could react with fluoride ions to form fluoride complexes, which are soluble in

acid solution as stable species, while MgF2 will remain inside the residues (Majima et al., 1988) and its solubility product equilibrium constant (pKsp) at 25°C is 10.29 (Dean, 1998). Therefore, it is reasonable to assume that fluoride ions not only give a markedly enhanced metal extraction, but also improve selectivity for nickel and cobalt dissolution, particularly over that of magnesium.

To investigate more precisely the effect of fluoride ions on the dissolution of nickel and cobalt, some additional leach-ing experiments on laterites were done in NH4F+H2SO4 or CaF2+H2SO4 solution under the same conditions as in NaF+H2SO4 solution. As in the case of the use of NaF+H2SO4 for leaching laterites, the use of other fluoride salts results in increased rates of nickel and cobalt dissolution, a finding consistent with Hirasawa (Hirasawa, 1987; Hirasawa et al., 1983), who found that the effects of CaF2 addition on the extraction of Ni, Fe and MgO were similar to those of NaF addition. Further, the use of NH4F+H2SO4 and CaF2+H2SO4 also exhibit optima; however, the optimum varies a little from that of NaF+H2SO4 (i.e., 0.3% in both cases versus 0.2% in the case of NaF+H2SO4).

Laterite characterization and comparison of derived products. Particle size analysis of the laterites reveals that 50% of the ore contains material smaller than 25μm (Fig. 4). Results of the chemical composition analysis included in Table 3 and the spot EDX of the laterite samples in Table 4 show that the laterites are typical of silicate laterite. The main mineral, which is evident from the X-ray diffraction pattern (Fig. 5), is lizardite ((Mg,Al)3[(Si,Fe)2O5](OH)4, referred to in JCPDF file

Figure 4 — Particle size analysis of laterites.

Table 2 — Gibbs free energy change and the equilibrium constant of the lizardite dissolution reaction in acid solution

with and without fluoride ions at 298.15K and 363.15K

T(K) Reaction ΔGT (kJ mol-1) KT

298.15 Mg3Si2O5(OH)4(S) + 6H+ = 3Mg2+ + 2H4SiO4(S) + H2O(I) -183.49 1032.1

Mg3Si2O5(OH)4(S) + 6H+ + 6F- = 3MgF2(S) + 2H4SiO4(S) + H2O(I) -341.62 1059.9

363.15Mg3Si2O5(OH)4(S) + 6H+ = 3Mg2+ + 2H4SiO4(S) + H2O(I) -173.88 1025.0

Mg3Si2O5(OH)4(S) + 6H+ + 6F- = 3MgF2(S) + 2H4SiO4(S) + H2O(I) -382.96 1055.1


No.50-1625) and the minor phases are magnetite (Fe3O4) and phlogopite (KMg3(Si3Al)O10(OH)2). DSC analysis included in Fig. 6 showed two endothermic peaks at about 66 and 607°C due to the release of adsorbed water and lizardite dehydrox-ylation, respectively (Tartaj et al., 2000). The weight loss corresponding to these two processes are 2.45% and 10.65%. On heating, DSC displayed an additional exothermic peak at 807°C associated with forsterite (Mg2SiO4) recrystallization. As is clear from the result, the nickeliferous laterites consist mainly of nickel-substituted lizardite revealing the cobble-like morphology as shown in Fig. 7.

FTIR spectroscopy is a simple and rapid technique and has been used successfully by many investigators in the study of the structure and composition of phyllosilicates. From the FTIR spectrum of laterites included in Fig. 8a, the band at 3,687 cm-1 can be assigned to the outer -OH stretching vibra-tions coordinated to three magnesium of the octahedral layer. The shoulder at 3,649 cm-1 is due to the inner -OH stretching vibrations (Scholtzov et al., 2003). Moreover, the frequency value of the inner -OH stretching band at 3,649 cm-1 and its intensity are directly proportional to the percentage of nickel cations in the trioctahedral sheet replacing the magnesium cation (Scholtzov et al., 2003). Water in laterites gives a broad band at 3,435 cm-1, corresponding to the -OH stretching vibrations and the band at 1,631 cm-1 is attributed to the -OH bending vibration (Madejov et al., 2002). The bands at 1,081 and 966 cm-1 display the asymmetric stretching modes of the Si-Ot- and Si-Ob-Si bond, respectively (Rinaudo et al., 2003). The 618 cm-1 peak is an -OH vibration mode specifically related

to the presence of magnesium ions in adjacent octahedral sites (Golightly and Arancibia, 1979). The shoulder at 579 cm-1 is associated with the Mg-O bending mode out of the plane. The intense band at 445 cm-1 is associated with the Si-O bending mode. From FTIR spectra (Fig.8b-8e) of solid residues leached with H2SO4 alone and the same solution containing different fluoride salts, the frequency value of the stretching vibration of the inner -OH group at 3,649 cm-1 increased about 11 cm-1

and the intensity of this band decreased in comparison with

Figure 5 — Power XRD pattern of laterites: -lizardite, ●-magnetite, □-phlogopite

Figure 6 — DSC-TG curves of laterites.

Table 3 — Chemical analysis of nickeliferous laterite sample.

Constituent O Mg Al Si S Cl Ca

Content, wt.% 50.2 20.4 0.88 18.9 0.011 0.061 0.021

Constituent Cr Mn Fe Co Ni Zn

Content, wt.% 0.27 0.14 7.48 0.023 1.38 0.011

Figure 7 — SEM image of laterites.

Table 4 — EDX data of nickeliferous laterite sample.

Constituent O Mg Al Si Mn Fe Ni

Content, wt.% 45.8 23.0 0.93 20.6 0.42 7.31 1.94


the inner -OH stretching vibration of the laterites which means that most of the nickel has been extracted from laterites. The relative increase in the intensity of υasSi-Ot- at 1,081 cm-1 and the decrease in the intensity of υasSi-Ob-Si at 966 cm-1 also indicate that the phyllosilicate structure has been decomposed by both hydrogen ions and fluoride ions in the leachant.

These results clearly show the importance of fluoride ions together with a strong acid, such as H2SO4, in breaking the silicate crystal lattice of laterites and improving the metal dissolution reaction in acid solution. However, for minerals containing significant amounts of silica, a judicious choice of fluoride ion concentration is required to optimize the overall nickel and cobalt recovery and to prevent metal loss due to the coprecipitation of Ni2+ and Co2+ with MgF2 in the residues and the adsorption of silicic acid formed in the leach solution.

Figure 8 — FTIR spectra of (a) laterites, (b) solid residues leached with H2SO4, (c) solid residues leached with H2SO4+0.2% NaF, (d) solid residues leached with H2SO4+0.3% NH4F and (e) solid residues leached with H2SO4+0.3% CaF2.

ConclusionsAn investigation was conducted to study the metal dissolu-

tion behavior of nickel-substituted lizardite in acidic solution containing different fluoride salts, including NaF, NH4F and CaF2. The presence of both hydrogen ions and fluoride ions in the leachant was found to be necessary for the improved dissolution of laterites, as fluoride ions could significantly increase the equilibrium constant of the dissolution reaction in acid solution. Furthermore, regardless of originating salt, the presence of F- was the same, thus indicating that the other ions in solution had no significant impact on the manner in which F- interacts with the system to produce enhanced nickel and cobalt dissolution. The experiment results were in accordance with the FTIR observations. As much as 81.7% Ni and 52.3% Co could be leached with the addition of 0.2 wt.% NaF under


the experimental conditions of particle size d50=25 μm, 10% (v/v) H2SO4, reaction time 2 h, reaction temperature 90°C, liquid-to-solid ratio 3 mL g-1 and stirring at 500 rpm.

AcknowledgmentsThis work is supported by a grant from the Major State

Basic Research Development Program of China (973 Pro-gram) (No.2007CB613602), the National High Technology Research and Development Program of China (863 Program) (No.2007AA060903) and Key Projects in the National Science & Technology Pillar Program during the Eleventh Five-Year Plan Period (No.2006BAB02A10). We thank Dr. Y. Chen for enlightening discussions and Dr. C. H. Deng for assistance with the samples.

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serpentinic nickeliferous ore - Part II. Sulfuric acid heap and agitation leach-ing,” Hydrometallurgy, Vol. 74, No.3-4, pp. 267-275.

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Briceno, A., and Osseo-Asare, K., 1995, “Particles in hydrometallurgy. Part II. Dewatering behavior of unflocculated laterite acid leach residues,” Metall. Mater. Trans. B, Vol. 26B, No.6, pp. 1133-1138.

Canterford, J. H., 1975, “Treatment of nickeliferous laterites,” Miner. Sci. Eng., Vol. 7, No.1, pp. 3-17.

Caron, M. H., 1950, “Fundamental and practical factors in ammonia leaching of nickel and cobalt ores,” JOM, Vol. 188, No.1, Trans., pp. 67-90.

Comba, P., Lei, K.P.V., and Carnahan, T.G., 1991, “Calcium fluoride-enhanced hydrochloric acid leaching of a manganese-bearing silicate ore,” U.S. Bureau of Mines Report of Investigations RI 9372, pp. 7.

Curlook, W., 2004, “Improvement to the acid pressure leaching of nickel laterite ores,” International Laterite Nickel Symposium, Proceedings of [a] Sympo-sium held during the TMS Annual Meeting, Charlotte, NC, United States, Mar. 14-18, pp. 325-334.

Dean, J. A., 1998, Lange’s Handbook of Chemistry, McGraw-Hill Professional, New York.

Deepatana, A., Tang, J., and Valix, M., 2005, “Kinetics of limonite and nontronite ore leaching by fungi metabolic acids,” Proceedings of the International Conference on Environmental Degradation of Materials in Nuclear Power Systems--Water Reactors, 12th, Salt Lake City, UT, United States, Aug. 14-18, pp. 1069-1078.

Golightly, J. P., and Arancibia, O. N., 1979, “The chemical composition and infra-red spectrum of nickel- and iron-substituted serpentine from a nickeliferous laterite profile, Soroako, Indonesia,” Can. Mineral., Vol. 17, No.4, pp. 719-28.

Hirasawa, R., 1987, “Effects of additives on the sulfuric acid leaching from the garnierite ore,” Toyama Kogyo Koto Senmon Gakko Kiyo, Vol. 21, pp. 15-19.

Hirasawa, R., 1989, “Effects of additives and roasting on the acid leaching of

garnierite ore,” Shigen to Sozai, Vol. 105, No.3, pp. 255-260.Hirasawa, R., Matsuoka, J., and Wakabayashi, K., 1983, “Studies on the acid

leaching of garnierite ore - effects of sodium fluoride addition on the extrac-tion of nickel, iron and magnesium oxide in sulfuric acid,” Nippon Kogyo Kaishi, Vol. 99, No.1139, pp. 37-41.

Kempthorne, D., and Myers, D. M., 2007, Mineral commodity summaries 2007: U.S. Geological Survey, Washington.

Lei, K. P. V., Droes, S. R., and Carnahan, T. G., 1991, “Calcium fluoride-sulfuric acid leaching of a manganese silicate ore,” EPD Congr. 91, Proc. Symp. TMS Annu. Meet., pp. 63-68.

Lloyd, R., and Turner, M. J., 1984, “Recovery of metals from ores,” WO, Aus-tralian Pat. No. 8404759, pp. 15.

Madejov, J., Janek, M., Komadel, P., Herbert, H. J., and Moog, H. C., 2002, “FTIR analyses of water in MX-80 bentonite compacted from high salinary salt solution systems,” Appl. Clay Sci., Vol. 20, No. 6, pp. 255-271.

Majima, H., Awakura, Y., Mashima, M., and Hirato, T., 1988, “Dissolution of columbite and tantalite in acidic fluoride media,” Metall. Trans. B, Vol. 19B, No.3, pp. 355-363.

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Rinaudo, C., Gastaldi, D., and Belluso, E., 2003, “Characterization of chrysotile, antigorite, and lizardite by FT-Raman spectroscopy,” Can. Mineral., Vol. 41, No.4, pp. 883-890.

Rubisov, D. H., Krowinkel, J. M., and Papangelakis, V. G., 2000, “Sulfuric acid pres-sure leaching of laterites - universal kinetics of nickel dissolution for limonites and limonitic/saprolitic blends,” Hydrometallurgy, Vol. 58, No. 1, pp. 1-11.

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Tang, J. A., and Valix, M., 2006, “Leaching of low grade limonite and nontronite ores by fungi metabolic acids,” Miner. Eng., Vol. 19, No.12, pp. 1274-1279.

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Whittington, B. I., and Johnson, J. A., 2005, “Pressure acid leaching of arid-region nickel laterite ore. Part III: Effect of process water on nickel losses in the residue,” Hydrometallurgy, Vol. 78, No.3-4, pp. 256-263.

Whittington, B. I., McDonald, R. G., Johnson, J. A., and Muir, D. M., 2003, “Pres-sure acid leaching of arid-region nickel laterite ore - Part I: effect of water quality,” Hydrometallurgy, Vol. 70, No. 1-3, pp. 31-46.


Pellet feed production via the concentration of flotation tailingsM.G. VieiraResearch engineer, Samarco Mineração SA, Ouro Preto, MG, Brazil

A.E.C. Peres Associate professor, Department of Materials and Materials Engineering, UFMG, Belo Horizonte, MG, Brazil

ABSTRACTThis investigation addressed the possibility of producing pellet feed at a silica content of ≤ 1.0% using a repre-sentative sample of Samarco´s mechanical flotation tailings. This sample contained 11.0% Fe and presented an F80 = 116 µm and a d50 = 63 µm. The higher silica content was verified in the coarser fractions and the higher iron content in the fraction -37 µm. Specular hematite is the most abundant mineralogical phase in the fraction -37 µm. The technological tests used one of four methods: magnetic concentration using a wet high-intensity magnetic separator, magnetic concentration using a SLon high-gradient magnetic separator, reverse flotation or direct flotation. The methods using the magnetic concentrator SLon, reverse flotation and direct flotation did not present results meeting the target of the study: iron content in the concentrate (rougher stage) greater than 34% or iron recovery greater than 80%. Cleaner stages were not achievable with the resulting concentrates. In contrast, the most successful method consisted of using wet high-intensity magnetic separators (WHIMS) as a rougher stage and reverse flotation as a cleaner stage. Using this route, it was possible to produce a pellet feed with a silica grade of 1.0% and high specular hematite content.

Paper number MMP-09-019. Original manuscript submitted May 2009. Revised manuscript accepted for publication October 2009. Discussion of this peer-reviewed and approved paper is invited and must be submitted to SME Publications Dept. prior to February 28, 2011. Copyright 2010, Society for Mining, Metallurgy, and Exploration, Inc.

IntroductionSamarco, a joint venture between Vale (50%)

and BHP (50%), operates an iron ore mine and the Germano concentrator in the Iron Quadrangle, Minais Gerais, Brazil. The pellet feed fines produced in the concentrators (24 million metric tons per year) are transported via a 396-km pipeline to the Ponta Ubu pelletizing plant, in the state of Espirito Santo. The blast furnace and direct reduction pellets are exported worldwide. The purpose of the present investigation was to recover iron oxides from a stream that is cur-rently emptied into a tailings pond.

Figure 1 shows a simplified Germano’s concentra-tor flowsheet, illustrating the points where flotation tailings are generated: mechanical (Wemco, 14.16 m3 or 500 ft3, 4 lines, 14 cells per line), tank cell (Outotec, 160 m3, 4 units) and column circuits. The mechanical flotation tailings present, in average, 10.5% Fe and represent 78% of the weight and 75% of the total iron content of the combined flotation tailings. This significant participation in the tailings stream was the reason for the selection of the mechanical cells circuit tailings for the present investigation.

A brief description of the flowsheet follows: the ROM containing 46 % iron and particle size -12.5 mm feeds the pre-primary and primary grinding, where the particle size is reduced from 44 % +105 µm to 10 % +105 µm. The product from this circuit

feeds the desliming (three stages, hydrocyclones of 38, 25.4 and 10 cm) (15, 10 and 4 in.). The overflow (-10 µm) feeds an iron fines recovery plant and the underflow feeds the me-chanical flotation circuit. The concentrate from the mechanical cell circuit feeds the regrinding mills #4, 5 and 6 operated in an open circuit and the reground product feeds the tank cell circuit. The concentrate from the tank cells feeds the column flotation circuit, where final silica specification is achieved (1.0 % direct reduction concentrate and 1.7 % blast furnace concentrate). The column flotation concentrate is classified in hydrocyclones of 25.4 cm (10 in.), where the overflow feeds the concentrate thickeners and the underflow feeds the mills #1, 2 and 3, operated in a closed circuit. The reground product (minimum 87.2 % -44 µm) feeds the concentrate thickeners, and then the concentrate is pumped via a 396 km pipeline to the pelletizing facilities.

The concentration of iron minerals contained in flotation tailings will maximize the production and minimize the flow of tailings disposed to the environment, improving the sustain-ability of the operation.

Literature reviewThe concentration of iron minerals from flotation tailings

was investigated by Brazilian researchers (Santos, 2003; Oliveira, 2006; Rocha, 2008) using magnetic concentration and/or reverse flotation.

Magnetic concentration in a wet high-intensity magnetic separator (WHIMS) machines is a traditional iron ore ben-

Key words: Pellet feed, Magnetic concentration, Flotation, Iron/iron ores


eficiation technique used in Brazil at Vale since 1972 (Jones separators). The pulp flows vertically, feeding a rotating cir-cular matrix (carousel) at two opposite points located under the action of a high-intensity magnetic field provided by per-manent magnets. The diamagnetic quartz particles flow past the matrix, while the paramagnetic hematite particles stick to the matrix and are washed down when the segment is free of the magnetic field action.

The SLon magnetic separator was specifically designed to make up for deficiencies of wet high-intensity magnetic sepa-rators, such as reduced capacity for fines, matrix obstruction and nonmagnetic particle entrapment. SLon was developed in China and the first industrial application occurred in 1988. In addition to the horizontal movement, the matrix pulsates vertically, preventing obstruction and entrapment. According to Hearn and Dobbins (2007), it is possible to produce cleaner concentrates from feed consisting of particles greater than 100 µm. The motivation for developing this technology was to concentrate Chinese ores with iron grades of < 30%.

Zeng and Dahe (2003) compared the performances of WHIMS-2000 and SLon-1500, installed in parallel at Qidashan Mineral’s concentrator. The iron feed grade was 15.78%. Slon-1500 yielded higher iron contents in the concentrate (+ 3.79%) at enhanced weight (+ 3.57%) and metallurgical (+13.60%) recoveries. Matrix obstruction was a more severe problem in the case of WHIMS, resulting in a lower iron grade in the concentrate. The increased iron recovery was attributed to the SLon pulsating mechanism. The fact that the matrix is always clean results in enhanced iron recovery.

Quartz reverse cationic flotation is the most widely used itabirite concentration method for pellet feed production. Quartz is floated with etheramines partially neutralized with acetic acid and iron minerals are depressed with nonmodified

starches. Vieira and Peres (2007) point out the major challenges in this system: coarse quartz particles do not respond well to the amine collecting action and fine iron oxide particles do not respond well to the starch depressant action.

The anionic direct flotation of iron oxides is potentially attractive regarding:

i. concentration of low grade ores;ii. concentration of marginal ores aiming at reducing the

waste/ore ratio;iii. iron recovery from material deposited in tailings ponds.

Nevertheless, most results from laboratory investigations indicated that iron oxide flotation with anionic collectors (fatty acids) and amphoteric collectors (sarcosinates and sulfosuc-cinamates) yields concentrates with high silica contents. Sodium silicate was not an effective quartz depressant in this system (Araujo et al., 2006).

Materials and methodsSampling and sample preparation. The flotation tailings

sample was formed by increments collected over 20 days in the four lines of mechanical flotation. After this time a 10,000-kg sample was generated, dewatered, homogenized in a conical pile and stored in 200-L drums. The splitting for technological tests was done using a riffle sampler.

Characterization of the sample. Quantitative chemical analyses of the head sample and fractions from -297 µm +37 µm and -37 µm were performed using ICP-OES, titration and loss on ignition. The size distribution of the fraction -297 µm +37 µm was determined by Tyler sieves and the size distribution of the fraction -37 μm was done by a laser diffraction particle size

Figure 1 – Germano´s concentrator simplified flowsheet.

(1) Mechanical flotation tailing

(2) Tank cell flotation tailing

(3) Column flotation tailing

(1)

(1)

(1)

(1)

(2)(3)


analyzer (Mastersizer Micro). Mineralogical analyses were performed by the counting grains method, using a reflected light optical microscope.

Technological tests. The tests involved four tech-nological options for the rougher stage:

1. magnetic concentration using a wet high-intensity magnetic separator (Minimag / GAUSTEC);

2. magnetic concentration using a SLon high-gradient magnetic separator (OUTOTEC);

3. reverse flotation; 4. direct flotation.

The collector and depressant used in the cationic reverse flotation were, respectively, decyl ether amine and gelatinized manioc starch, both dosed at 1% w/v solution. The amine was a blend of 75% monoethera-

mine and 25% dietheramine. The manioc starch depressant was gelatinized with caustic soda solution at 5:1 w/w starch : NaOH ratio. The reagent dosages used in the experiments are presented in the tables and figures. Flotation pH was set at 10.5.

Direct flotation was performed at pH = 7, the collector being oleic acid saponified with caustic soda solution at 5:1 w/w acid : NaOH ratio. The depressant used was sodium silicate.

The cleaner stage was performed only if the iron content of the rougher concentrate was > 34%, at an iron recovery level > 80%. If this condition was not achieved, the route was discarded.

The cleaner stage was performed using reverse flotation, with the flotation tests done in a bench scale mechanical machine and a pilot flotation column.

Results and discussionSize distribution and chemical

analyses by size range. Figure 2 presents the sample size distribution. The values of d50, F80 and the slimes content (fraction < 10 µm) were, respectively, 63 µm, 116 µm and 2.69 %.

Table 1 shows the chemical analyses of cumulative retained fractions. There is a high silica distribution in the coarse frac-tions and a higher iron distribution in the fine fractions, a consequence of the difference in grindability between the harder quartz particles and the softer iron oxides. The phosphorus is concentrated predominantly in the -37 μm fraction and alumina and LOI in fractions above 210 μm. The iron content in the -37 μm single fraction is 44.02 %.

Figure 3 shows the main mineralogical phases present in the tailings sample from the mechanical flotation. It is noted that quartz is the most abundant mineral phase, representing about 51 % by weight. The most representative iron mineralogical phases are specular hematite and martite,

Table 1 – Size distribution and chemical analyses of cumulative retained fractions.

Aperture µm

Cum. % passing

Fe % SiO2 % Al2O3 % P % LOI %

297 98.73 0.59 98.19 0.68 0.001 0.25

210 93.41 1.22 97.35 0.65 0.001 0.22

149 75.44 2.60 95.96 0.18 0.001 0.09

105 63.42 3.01 95.45 0.13 0.001 0.08

74 36.86 3.47 94.82 0.09 0.001 0.10

53 26.26 3.90 94.22 0.08 0.001 0.10

44 18.98 4.44 93.37 0.13 0.005 0.12

37 15.43 5.06 92.43 0.15 0.007 0.13

-37 0.00 11.10 83.68 0.27 0.014 0.13

Calculated sample 11.06 83.74 0.27 0.014 0.13

Analyzed sample 11.01 83.68 0.21 0.014 0.33

Figure 2 – Sample size distribution.

Figure 3 – Mineralogical phases distribution.


with 29 % and 14 % by weight, respectively. The percentage of nonliberated quartz is around 2 % by weight and the quartz liberation degree is 96 %. In general, the larger the particle size, the greater the percentage of quartz and the lower the percentage of specular hematite. Specular hematite is the most abundant mineral phase in the -37 µm fraction, at 49 % by weight, followed by liberated quartz (24 % by weight) and martite (22 % by weight).

Technological tests. Magnetic concentration

using WHIMS (Minimag). Table 2 summarizes the results of tests of high-intensity magnetic concen-tration using Minimag. The variables tested were: magnetic field intensity, washing pressure, solids percentage in the feed and carousel speed. The best results were achieved using a field of 12,000 G, a pressure of 2.5 kfg/cm2, solids percentage of 35 % and rotation of 4 rpm.

Table 3 shows the results of two flotation tests (bench scale) performed on a concentrate obtained by the Minimag. With 100 g/t amine and 600 g/t starch it was possible to achieve a pellet feed within the 1.0 % silica specification. The percentage -44 μm was also satisfactory.

Magnetic concentration using a SLon. Figure 4 shows the iron recovery and iron content curves ver-sus magnetic field intensity. None of the tests yielded 34 % iron in the concentrate, at an iron recovery > 80 %; thus, cleaner tests were not performed. The best results were achieved using 2.5 rpm rotation and 200 Hz pulse rate. The change of stroke from 0 to 15 mm was not significant.

Reverse flotation. Figure 5 shows concentrate iron recovery and iron content curves as a function of increased amine dosage and starch dosage. It can be observed that above 100 g/t amine, the iron content does not increase significantly (except at a 200 g/t starch dosage). Increased starch dosage provided higher iron recoveries and higher iron content in the concentrate and the results for the starch dosages 150 and 200 g/t were similar, indicating that higher starch dosages do not affect significantly the concentrate

Table 2 – Compilation of 12 preconcentration tests using Minimag (Fe% minimum to maxi-

mum; average).

Flow Fe % Weight recovery % Iron recovery %

Feed 9.35 – 10.69 20.36 – 28.35

24.82

82.57 – 88.79

86.02Concentrate 32.38 – 38.38

Tailings 1.65 – 2.16

Table 3 – Minimag’s concentrate flotation tests.

Amine starch

g/t

Concentrate Recovery

Fe % SiO2 % P % Al2O3 % -44µm % BSA cm2/g Weight % Iron %

80 / 600 67.77 1.52 0.026 0.23 85.47 855 24.56 46.02

100 / 600 67.70 1.07 0.026 0.23 86.43 831 23.68 45.44

Figure 4 – Curves of iron recovery and iron content obtained with the flotation tailings concentration using SLon.

Figure 5 – Iron content and recovery curves versus starch and amine dosage.


iron grade and recovery.

Direct flotation. Figure 6 shows the curves of iron recovery and iron content depending on the oleic acid and sodium silicate dosages. The iron recovery curves are consistent, increasing with higher oleic acid dosage. Increasing the oleic acid dosage decreased the concentrate iron content. Due to the oleic acid low selectivity, the collected quartz particles contaminate the concentrate.

Increasing the sodium silicate dosage increased the iron content of the concentrate, but decreased the iron recovery, also indicating a low selectivity.

Conclusions Among the four methods selected for the production of

rougher concentrates from iron ore flotation tailings, WHIMS, SLon high-gradient magnetic separator, reverse flotation and direct flotation, WHIMS was the only one to meet the target of an iron content greater than 34% and iron recovery greater than 80%.

The WHIMS concentrate was submitted to reverse flota-

tion cleaning stages yielding 1.0% silica grade in the pellet feed product.

AcknowledgmentsThe authors are grateful to Samarco, Fapemig and CNPq

for financial support.

ReferencesAraujo, A.C., Viana, P.R.M., and Peres, A.E.C., 2005, “Reagents in iron ores

flotation,” Minerals Engineering, Vol. 18, pp. 219-224. Hearn, S.B., and Dobbins, M.N., 2007, SLon magnetic separator: a new approach

for recovering and concentrating iron ore fines, Montreal Energy & Mines, 8 pp.Oliveira, P.S., 2006, Routes for recovery of iron oxides fines contained in the

Conceição´s concentrator slimes thickener underflow, M.Sc. thesis, CPGEM UFMG, 113 pp. (in Portuguese)

Rocha, L., 2008, Study on the economic recovery of slimes from an iron ore concentrator via reverse cationic flotation, M.Sc. thesis, CPGEM UFMG, 121 pp. (in Portuguese)

Santos, I.J., 2003, Study on the production of iron ore concentrates from Germano´s pond tailings, M.Sc. thesis, PPGEM UFOP, 107 pp. (in Portuguese)

Vieira, A.M., and Peres, A.E.C., 2007, “The effect of amine type, pH, and size range in the flotation of quartz,” Minerals Engineering, Vol. 20, p. 1008-1013.

Zeng, W., and Dahe, V., 2003, “The latest application of SLon vertical ring and pulsating high-gradient magnetic separator,” Minerals Engineering, Vol. 16, pp. 219-224.

Figure 6 - Iron content and recovery curves versus oleic acid and sodium silicate dosages.


Solvent extraction of zirconium from zircon leach liquor using triphenylphosphine oxideK. SaberyanProfessor, Nuclear Fuel Cycle Research School, Nuclear Science and Technology Research Institute (NSTRI), Tehran, Iran. E-mail: [email protected]

P. Vahedian-DonyaparastGraduate student, Department of Mining Engineering, Islamic Azad University, Science and Research Branch, Poonak, Hesarak, Tehran, Iran

M. NoparastProfessor, Department of Mining Engineering, University of Tehran, Tehran, Iran

E. ZolfonounChemical engineer, Nuclear Fuel Cycle Research School, NSTRI, Tehran, Iran

A. NematiProfessor, Department of Inorganic Chemistry, Faculty of Chemistry, Tehran University, Tehran, Iran

AbstractA technique is described for the solvent extraction of zirconium. In this method, zirconium has been extracted quantitatively from 0.005 M sodium salicylate solution at pH 2.5 - 3 using 0.01 M triphenylphosphine oxide dissolved in toluene as an extractant. Zirconium has also been recovered quantitatively from 0.2 M potassium thiocyanate solution, but extractions in nitric and hydrochloric acids (as media) were negligible. The resulting extracted metal ion has been stripped with sulfuric acid (0.5 M) and determined with inductively coupled plasma atomic emission spectrometry. This method has been applied successfully to achieve 95% recovery of zirconium from zircon concentrate.

Paper number MMP-09-003. Original manuscript submitted January 2009. Revised manuscript accepted for publication October 2009. Discussion of this peer-reviewed and approved paper is invited and must be submitted to SME Publications Dept. prior to February 28, 2011. Copyright 2010, Society for Mining, Metallurgy, and Exploration, Inc.

IntroductionZircon is the most important commercial source

of zirconium and its compounds and alloys. Zircon occurs as a common mineral in pegmatites, particu-larly in syenitic and granitic pegmatites and granites, diorites and gneisses. Zircon, as a heavy mineral very resistant to chemical decomposition and erosion, is concentrated with other heavy resistant minerals in placer deposits, principally in river and beach sands. Large commercial detrital deposits are found in Aus-tralia, South Africa, Brazil, Ceylon, India, Madagascar, USA and Egypt. Zircon is commonly associated with other heavy minerals, such as illmenite, magnetite, monazite and rutile (Abdel-Rehim, 1974 a,b; Abdel-Rehim, 2002 a,b; Abdel-Rehim, 2005; Stanley and Lafond, 1975).

Several methods have been used for the industrial processing of zircon to extract zirconium and produce zirconium dioxide and tetrachloride (Bidaye and Venkatachalan, 1999; Doyle and Duyvesteyn, 1982; Habashi, 1999; Welham, 2000).

Zirconium is used in the nuclear industry, due to its

excellent corrosion resistance and low neutron cross-section. Organophosphorus compounds, namely TBP and D2EHPA, have been extensively employed in the solvent extraction studies of Zr (IV). Levitt and Freund (1956) have studied the extraction of Zr (IV) from hydrochloric acid solutions by TBP and suggested that ZrCl4.2S is the extracting species. Solvent extraction studies of Zr (IV) from chloride solutions by TOPO have indicated that the species are extracted into the organic phase as disolvates (Sato, 1982). The extraction behavior of Zr (IV) from aged chloride solutions by di-(2-ethyl hexyl) phosphoric acid (D2EHPA) has been investigated by Biswas and Hayat (2002).

A literature survey in this subject reveals that trioctylamine (TOA) has been used to extract zirconium (Rajmane and Sargar, 2006; Chatterjee and Basu, 1990; Das and Lahiri, 1994; Kutyrev et al., 2000). The method involved extracting zirconium from a hydrochloric acid solution with TOA in benzene, in which the extraction efficiency was subjected to the length of amine chain (Sato and Watanabe, 1971). It has been reported that higher reagent concentrations are required (Seidl et al., 1983). The trioctylamine reagent is less effective for mutual separation of elements when they are present as

Key words: Zirconium, Solvent extraction, Triphenylphosphine oxide, Thiocyanate, Salicylate


congeneric pairs in aqueous solutions. A solution of Alamine 336 in an appropriate diluent has been used to extract zirconium from 1 M sulfuric acid and 4 M malonic acid media (Kutyrev et al., 2000). However, these methods suffer mainly from being multistage extractions and from emulsion formation. Extraction of zirconium from hydrochloric acid solution with di-n-octylamine in benzene has also been investigated under different conditions (Nekovar et al., 1997).

Solvent extractions of zirconium with Cyanex 272 (Kent and Corinne, 1985), Cyanex 302 and Cyanex 301 (Wang and Li, 1996) from sulfuric acid solutions have been investigated systematically. The behavior of a third phase formation in solvent extraction of zirconium by TBP has also been studied by Kim et al. (1993). Some other phosphorus-containing ex-tractants have been reported for the extraction of zirconium from acid media, such as tritolyl phosphate (TTP) (Curtui and Soran, 2000), dibutyl hydrogen phosphate (Fedorov et al., 2000; Zilberman et al., 2000), di-o-tolylphosphoric acid (Biswas and Habib, 1994), tri-isoamyl phosphate (TAP) (Hasan and Rupainwar, 1990), tetraphenylalkylenediphosphine dioxides (Turanov et al., 2000). Moreover, the solvent extractions of the tetravalent metal ions zirconium (IV), hafnium (IV) and thorium (IV) have been investigated using a recently synthe-sized naphthol-derivative Schiff base (HAPMN) (Saberyan et al., 2008).

The extraction of zirconium decreased markedly with increasing sulfate ion concentration in the aqueous phase; the extraction was found to be incomplete, as well. Selective stripping of zirconium has never been reported. Reddy and Kumar (2004) and Taghizadeh et al. (2007) have shown that sulfuric acid is the proper reagent for zirconium stripping from loaded Cyanex 272, although the addition of some hydrogen peroxide to sulfuric acid can improve stripping efficiency. However, hafnium stripping has not yet been studied in any literature review.

Even though extensive work has been performed on the extraction and separation of zirconium (IV), in many instances the available methods are found to be inadequate for effective extraction. It is therefore felt worthwhile to evolve an entirely new procedure, one which facilitates the extraction and separa-tion of zirconium (IV) from associated elements with a mini-mum amount of manipulation. The present research attempts to determine the solvent extraction behavior of Zr (IV) using triphenylphosphine oxide (TPPO) in toluene.

ExperimentalApparatus. The determination of zirconium was performed

on a Varian model Liberty 150 AX inductively coupled plasma optical emission spectrometer (ICP–OES). A Metrohm model 744 digital pH-meter, equipped with a combined glass-calomel electrode, was employed for the pH adjustments.

Reagents. The extractant triphenylphosphine oxide was obtained from Fluka. All other reagents were pro-analysis grade, from Merck Company. The zirconium stock solutions

of 100 µg mL-1 were prepared by dissolving appropriate amounts of ZrOCl2.8H2O (from Merck) in 100 ml of 0.1 M hydrochloric acid.

Extraction procedure. An aliquot of solution (20 ml) con-taining zirconium (10 μg mL-1) and sodium salicylate (0.005 M) was adjusted to pH 2.5 with dilute hydrochloric acid and sodium hydroxide. The organic phase was prepared by dis-solving triphenylphosphine oxide (0.05 M) in toluene. Batch contacts with phase ratio O/A=1 were performed at 25 °C us-ing a mechanical shaker. The contact time was 30 min, which was about twice that required to attain apparent equilibrium. Organic and aqueous phases were separated by a separatory funnel. Following separation, zirconium was then stripped from the organic phase by 20 mL of 0.5 M sulfuric acid. The metal concentrations in the aqueous phase and stripped phase were determined by ICP-OES.

Results and discussionEffect of hydrochloric and nitric acids. Preliminary ex-

periments on zirconium extraction with this extractant have shown that zirconium extraction efficiencies in nitric and hydrochloric acids (as media) without adding complexing agents are negligible, and the maximum extraction has been achieved, about 7% in 5 M hydrochloric acid media. The results are shown in Table 1.

Effect of addition of complexing agents. The effect of the addition of various complexing agents, such as acetate, thiocyanate and salicylate, into the aqueous phase was studied. The results are summarized in Table 2. As the results indicate, the extraction of zirconium by triphenylphosphine oxide is maximized by using salicylate as a complexing agent.

Effect of pH of the aqueous phase. In this study, the effect of aqueous phase pH containing salicylate ion on zirconium extraction was examined. The results are shown in Fig. 1 and show that the maximum extraction efficiency can be achieved at a pH over 2.5 and higher, whereas at a lower pH, extraction is decreased, probably due to the reaction of salicylate anions with H+ ions preventing the formation of a neutral complex

Table 1 — The effect of hydrochloric and nitric acid on

zirconium extraction.

Zirconium extraction, %

HNO3 HCl Concentration (M)

1.5 2.3 0.1

3.2 4.6 3

5.3 7.4 5

Table 2 — Effect of adding complexing agents into the

aqueous phase.

Complexing agents

(0.01 M) Extraction, %

pH=1 pH=2 pH=3

Acetate 7.7 18.2 37.4

Thiocyanate 16.5 22.6 40.2

Salicylate 60.4 85.3 97.1

Figure 1 — The effect of the pH of the aqueous phase containing salicylate anion on extraction efficiency.


ionpair with zirconium (IV).

Effect of sodium salicylate concentration. In order to ob-tain a quantitative extraction of Zr (IV), the sodium salicylate concentration has been varied from 5×10-4 to 5×10-2 M at pH 2.5. The results are shown in Fig. 2. The sodium salicylate concentration required for quantitative extraction of Zr (IV) is found to be 5×10-3 M. In Fig. 3, a log–log plot of the distribution ratio versus log [salicylate] (at fixed pH and triphenylphosphine oxide concentration) gave a slope of 1.29. The mechanism for such reaction can be represented as follows: ZrO2+

(aq.) + Sal2-(aq.) → ZrO(Sal)(aq.) (1)

ZrO (Sal)(aq.) + nTPPO(org.) → ZrO(Sal) . nTPPO(org.) (2)

The effect of diluents. The suitability of several diluents, such as toluene, xylene, benzene and n-hexane, for the ex-traction of zirconium was investigated. The extraction was found to be almost total with 0.005 M salicylate and 0.05 M solution triphenylphosphine oxide in toluene (97%). However, the extraction of Zr (IV) is incomplete with xylene (93.2%), benzene (95.3%) and n-hexane (92.1%).

Effect of TPPO concentration. Zr (IV) has been extracted with different concentrations of triphenylphosphine oxide in toluene (10-3-10-1 M) and the results are shown in Fig. 4. Quantitative extraction has been achieved from 10-2 M triphenylphosphine oxide solution. Therefore, the optimum concentration of extractant selected is 10-2 M triphenylphos-phine oxide in toluene.

Effect of metal loading. The loading capacity of 0.01 M TPPO in toluene has been determined by contacting the organic phase repeatedly with the same volume of aqueous phase containing 10 – 2000 μg mL-1 of Zr(IV) and 0.05 M of salicylate. The results are shown in Fig. 5. Zirconium (IV)

has been extracted quantitatively in the concentration range of 10 – 500 μg mL-1. It is possible to quantitatively extract higher concentrations of zirconium (IV) using a larger volume and a higher concentration of TPPO.

Choice of stripping agents. In this step, different stripping agents have been tested to strip the zirconium from the organic phase and the obtained results are represented in Table 3. It was found that 0.5 M sulfuric acid back-extracts zirconium (IV) quantitatively from the organic phase.

Comparison of extraction efficiency of Zr (IV) with other associated metal ions. The extraction behaviors of hafnium (IV), titanium (IV), aluminum (III) and iron (III) (10 and 100 μg mL-1 each) have been studied under the optimum condi-tions for zirconium extraction. The results are summarized in Table 4 and show that the extraction efficiency of zirconium in the presence of other metal ions is nearly constant. The extraction of Hf (IV), because of chemical similarities with Zr (IV), is higher than other metal ions. The lowest extraction belongs to Ti (IV).

ApplicationThe studied system has been implemented to determine the

zirconium in zircon concentrate. Zircon concentrate (procured from China) was treated as described in Gilbert (1954) and Beyer et al. (1954). 20 g of zircon concentrate was mixed with NaOH (50%) at a phase ratio of 1/1 g.mL-1 in a nickel crucible and heated at 600 °C for about four hours. The mass was then mixed with 100 mL water for about five hours and filtered through medium-speed filter paper. Then the residue was leached with 5 M hydrochloric acid at an S/L ratio of 1/5 g.mL-1 and 90 °C for about one hour (leaching recovery 97%). An aliquot (2 mL) of this solution was diluted to the proper volume and extracted for zirconium by this technique. The sodium salicylate and TPPO concentrations were 0.05 and 0.01, respectively; phase ratio O/A = 1 and contact time 30 min. The result shows that about 95% of zirconium was recovered

Figure 2 — The effect of sodium salicylate concentration on extraction efficiency.

Figure 3 — Plot of log (D) (distribution ratio) versus log sodium salicylate concentration.

Figure 4 — The effect of triphenylphosphine oxide (TPPO)concentration on extraction efficiency.

Figure 5 — The effect of Zr (IV) concentration on extraction efficiency.


from leach liquor. The final concentration of zirconium in the leach liquor was 500 μg mL-1 and the amount of zirconium recovered by the proposed method was 475 μg mL-1.

Conclusions A technique has been developed to extract the zirconium

from zircon leach liquor.The advantages of this method are:• The method is simple, selective and rapid.• The method needs no pre-equilibration and is free from

interference resulting from commonly associated ele-ments.

• Using this method to extract zirconium (IV) from as-sociated elements of zircon leach liquor led to 95% zirconium recovery.

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Table 3 — The effect of stripping agents on stripping efficiency.

Stripping, % Stripping agents

2M 1M 0.5M 0.1M

88.8 98.6 98. 6 94.1 Sulfuric acid

64.6 83.5 87.9 69.2 Nitric acid

10.9 36.6 41.6 9.9 Hydrochloric acid

Table 4 — Comparison of extraction efficiency of Zr (IV) with other associated metal ions

Extraction, % Zr Extraction,% Concentration (μg.mL-1) Foreign ions

76.5 96.35 10Hf (IV)

69.9 89.63 100

51.8 93.23 10Fe (III)

28.9 88.16 100

32.8 95.43 10Al (III)

17.5 90.58 100

2.08 96.82 10Ti (IV)

2.06 96.71 100


Performance of a full-scale continuous centrifugal concentrator in reconcentrating fine hematite from tailings

Luzheng Chen, Siqing Liu and Shuming WenAssociate professors and professor, respectively, Faculty of Land Resource Engineering, Kunming University of Science and Technology, Kunming, Yunnan Province, China

AbstractThe study aimed to evaluate the separation performance of a full-scale continuous centrifugal concentrator in the reconcentration of fine hematite from high-gradient magnetic separation (HGMS) concentrate in a tailings recov-ery plant, and to investigate the usefulness of the technology in industrial applications. The effect of changing the operating variables on concentrate grade and recovery was studied. The results indicate that among the variables considered in the study (water spray pressure, centrifugal acceleration, feed volume flow rate, feed percent solids and the reciprocation and spacing of water sprays), the change in centrifugal acceleration affects the recovery the most and continuous concentration can be achieved only when the water spray has a sufficient pressure. In addition, the reciprocation and spacing of the water sprays have a significant influence on the concentrate grade. Although performance of the full-scale concentrator is slightly inferior to pilot-scale trials, due to limitations in the scale-up design and the fluctuations of feed material characteristics, this concentrator provides a potential way to achieve continuous concentration and obtain a high-grade concentrate from tailings.

Paper number MMP-09-031. Original manuscript submitted July 2009. Revised manuscript accepted for publication No-vember 2009. Discussion of this peer-reviewed and approved paper is invited and must be submitted to SME Publications Dept. prior to February 28, 2011. Copyright 2010, Society for Mining, Metallurgy, and Exploration, Inc.

IntroductionCurrently, reverse flotation is widely used to obtain

a high-grade concentrate in iron ore processing flow-sheets, in such applications as the concentration of hematite from quartz (dos Santos and Oliveira, 2007; Pavlovic and Brandao, 2003), the removal of hematite from silica (Mowla et al., 2008) and the concentration of magnetite in a flotation column (Zhang Haijun et al., 2008). However, flotation always presents a relatively high energy cost, as well as having a high reagent consumption and causing environmental pollution. Moreover, where low capacity and cost, simple operation and flexibility are needed, flotation

may not be preferred, due to its high operating expenses and complexity compared with physical separation methods.

There is increasing interest in using physical separation devices to clean tailings material and obtain high-grade iron concentrates. This interest is due to increasing energy and re-agent costs and stringent environmental restrictions on the one hand (Burt, 1999) and to the increasing demand for high-grade materials in the iron and steel industries on the other (Arol and Aydogan, 2004; Yu Yongfu, 2004). It should be noted that numer-ous pilot-scale trials to clean various magnetite minerals using magnetic aggregation techniques have been conducted during the past decade in China. For example, Jin Wenjie et al. (2000) achieved an improved concentrate product assaying 53.40 %

Key words: Gravity separation, Hematite, Tailings, Iron/ iron ores


Fe by cleaning a magnetite concentrate assaying 46.89 % Fe using a magnetic separation column. This magnetic separator usually achieves a 1 - 4% grade improvement when used to clean relatively high-grade magnetite concentrates, e.g., 58 - 62% Fe or higher; exceeding this grade, quality improvement becomes increasingly difficult, even negligible (Liu Yang et al., 2009). A magnetic gravity concentrator using the technique was also reported effective in obtaining a high-grade magnetite concentrate (Yuan Zhitao et al., 2001).

Although magnetic aggregation is used to obtain high-grade magnetite concentrate, it cannot be used to treat weakly magnetic minerals, such as hematite. High-gradient magnetic separation (HGMS) is now widely used to preconcentrate weakly magnetic minerals (Xiong Dahe, 1997), but generally it does not produce a high-grade marketable concentrate. This outcome is due to the fact that the strong capturing effect of the magnetic force upon magnetic particles results in the high recovery of partially liberated magnetic particles, thereby deteriorating the quality of magnetic concentrate.

Centrifugal concentration provides a means to reconcen-trate weakly magnetic minerals and produces a high-grade concentrate. Knelson concentrators are widely applied in the gold industry (Coulter and Subasinghe, 2005) and Falcon concentrators are used for fine coal cleaning (Honaker et al., 1996), but not many centrifugal concentrators are reported to be effective in concentrating iron minerals, because the concentration of iron minerals always involves a tremendous

concentrate mass compared with that of precious minerals, such as gold. It was reported by Chen Luzheng et al. (2007; 2008) that a continuous centrifugal concentra-tor was able to obtain a high-grade concentrate by reconcentrating HGMS concentrate from a refrac-tory fine hematite ore and tailings. Further theoretical analysis and pilot-scale trials have demonstrated the effectiveness of this concentra-tor, notably in reconcentrating fine hematite from HGMS concentrate.

Therefore, a full-scale continu-ous centrifugal concentrator was installed in a tailings recovery plant at the Hainan Iron and Steel Company in southernmost China in 2007, to investigate using the technology to reconcentrate fine hematite from an HGMS concen-trator in industrial applications. The purpose of this paper is to evaluate the separation performance of the concentrator, the main focus being the grade and recovery of the con-centrate product and the response of the concentrator to changes in feed and operating conditions.

B rie f in t ro d u c t io n t o c o n t i n u o u s c e n t r i f u g a l concentration. The concentrator achieves separation due to the

centrifugal acceleration of a thin flowing film, in which particles separate depending on differences in density and the characteristics of the flowing film (Ergün and Ersayin, 2002; Majumder et al., 2006). As illustrated in Fig.1, slurry is evenly fed onto the inner surface of the drum. As the drum rotates, the axial velocity of the flowing film increases due to the drum’s angle of inclination. While traveling over the drum, heavy and light minerals stratify and leave it in separated layers. Light minerals (hollow circle in Fig. 1) obtain a higher axial acceleration over the drum, but a lower radial settling velocity, due to being subjected to a weaker centrifugal force; thus, they are in the upper moving layers of the film and flow quickly out of the drum. Heavy minerals (solid circle in Fig. 1) in the film settle down into the drum and form a settled bed, due to being subjected to the stronger centrifugal force.

Continuous concentration is achieved using reciprocating water sprays to remove the settled bed, similar to the cutting of rocklike material with a moving water jet (Momber and Kova-cevic, 1997). The reciprocation of the sprays is much lower than the axial velocity of the flowing film, so that the light minerals in the film escape being flushed into the concentrate; instead, these materials travel through the sprays in spiral curves via the spacings between the sprays (as schematically presented in Fig. 1) and flow out of the drum. Some of the lighter minerals will inevitably be flushed down into the concentrate, as the sprays flush down the settled bed through the whole depth of the film, resulting in deterioration of the concentrate grade. The

Figure 1 — Schematic demonstration of a continuous centrifugal concentration.

Figure 2 — Full-scale continuous centrifugal concentrator: 1 = support frame, 2 = drum, 3 = feed device, 4 = guard, 5 = concentrate collector, 6 = water sprays, 7 = reciprocating mechanism, 8 = drum driver, F = feed, C = concentrate, T = tailings, W = water inlet.


quantity of these lost minerals, however, can be minimized by optimizing the reciprocating sprays.

ExperimentalFull-scale continuous centrifugal concentrator. The

1600 × 900 mm (drum diameter × length) full-scale continu-ous centrifugal concentrator as shown in Fig. 2 consists of a conical drum made of polyester alkyd with a wearable layer in the inner surface and a unique reciprocating mechanism. In the upper chamber of the drum is installed a concentrate collector below the reciprocating water sprays.

While the concentrator is being operated, the feed material enters the rotating drum as a slurry through a feeding device. The slurry is forced onto the drum wall and rotates with the drum in the form of flowing film. Heavy particles in the film settle down onto the drum surface, due to centrifugal force, and light particles are driven outward to become tailings. The reciprocating mechanism ejects high-pressure water sprays against the rotating drum, causing the settled bed of heavy particles on the drum surface to be flushed down into the concentrate collector.

This full-scale continuous concentrator is modeled on the pilot-scale concentrator (Chen Luzheng et al., 2008) and the conventional intermittent centrifugal concentrator. Two scale-up conditions are used in the full-scale concentrator:

1. Operating parameters are the same as the pilot-scale con-centrator; i.e., centrifugal acceleration is 0 - 60 g (g is gravity acceleration), water spray reciprocation is 20 - 60 mm/s and water spray pressure is 0 - 0.6 MPa. In addition, the spacing of water sprays is designed to be adjustable from 50 to 100 mm.

2. Drum size is the same as the conventional intermittent concentrator developed in the 1960s in China; i.e., 1600 × 900 mm. However, the continuous concentrator has an decreasing three-inclination drum configuration of 10°- 5°- 2.5°, instead of the single- or two-inclination configuration of 3 - 5° in a conventional concentrator.

It should be noted that, due to the limitations in precision and the slight elasticity of the polyester alkyd material of the

drum, a vibration at least 7 - 10 mm in a radial direction and a vibration at least 6 - 8 mm in an axial direction are detectable while the drum rotates at a speed of 200 r/min (around 48 g). And as shown in Fig. 2, compared with other centrifugal separa-tors, one of the essential features of this full-scale concentrator is a horizontal cone drum similar to that of the multi-gravity separator, instead of a vertical one as characteristic of the Knelson and Falcon concentrators.

Description of feed. The fine hematite material fed to the full-scale concentrator was from the magnetic preconcentration product stream of a SLon-2500 pulsating HGMS separator operating at a background magnetic induction of 0.9 T at the tailings recovery plant of the Hainan Iron and Steel Company (Chen Luzheng et al., 2009). The tailings treated at the plant are from a rich hematite ore processing plant, two lean hema-tite ore processing plants and a sulfur-rich lean hematite ore processing plant. Thus, hematite is the dominant mineral in the feed material, with the primary gangue minerals being quartz, muscovite, epidote, tremolite, garnet and barite.

The material is characterized by fluctuations in feed condi-tions, notably iron grade and particle size. Normally, iron grade varies in the 48 - 54 % region, and 91 - 97 % and 68 - 76% of the material is respectively less than 74 μm and 37 μm, accounting for more than 96 % and 70 % of the iron values, respectively. This fluctuation is mainly due to the frequently disrupted operations of one or two of the four hematite ore processing plants itemized above.

Methods. As shown in Fig. 3, feed volume flow rate was controlled by an adjustable valve located below the feed box with a volume of 0.50 × 0.45 × 0.45 m3. Feed volume flow rates were measured using a sample can and a chronometer. The reciprocating mechanism of the concentrator was connected to a nearby clean water pipe via a high-pressure water tube. A pressure pump was installed in the pipeline to produce high-pressure water sprays. The required water spray pressure was achieved through another adjustable valve located between the

Figure 3 — Experimental setup of full-scale continuous cen-trifugal concentration: 1 = full-scale continuous centrifugal concentrator, 2 = pressure pump, 3 = adjustable valve, 4 = pressure gauge, 5 = feed box, 6 = sample can.

Figure 4 — Effect of water spray pressure on concentration continuity. Conditions: feed volume flow rate = 10 m3/h, feed % solids = 30%, feed grade = 49.56% Fe, feed size = 92.60% (-74 μm).


pressure pump and the reciprocating mechanism and a pres-sure gauge was used to indicate the operating pressures of the sprays. For all tests in this investigation, the impact angle of the water sprays with respect to the drum surface was fixed at 45° (Ren Nanqi et al., 2009) and the impact energy of the sprays was kept sufficient to flush down the bed using nozzles with different diameters. Centrifugal acceleration is adjustable through an AC transducer connected to the drum driver and the reciprocation of the sprays is adjusted through another AC transducer connected to the reciprocating mechanism of the concentrator.

After sufficient time was allowed for the concentrator to achieve steady operation, a typical sample was collected from the concentrate stream for analysis. The feed volume flow rate and feed % solids were measured and examined periodically to keep a steady feed flow to the concentrator.

Due to the fluctuations in feed material, the effect of chang-ing the operating variables on the separation performance of the concentrator was assessed only when the feed to the concentrator was found to be basically steady.

Results and discussionEffect of water spray pressure on concentration

continuity. The effect of the impact pressure of the water sprays on the concentration continuity was first studied with the concentrator at different centrifugal accelerations. It can be seen from Fig. 4 that when all the other variables are kept constant, the water spray pressure has a dominant influence over the concentrate weight. For each acceleration, concentrate weight increased drastically as the water spray pressure increased, until it reached a maximum; this maximum increased as the acceleration increased. The separation results indicate that the maximum concentrate weights are approximately 44 %, 61.5% and 70.5 % when the centrifugal acceleration increases from 22 g to 40 g and 60 g, respectively.

Two main factors contributed to this result. First, the ag-gregating force between heavy particles in the settled bed on the drum surface increases as the centrifugal acceleration increases, as a result of their exposure to a stronger centrifugal force. Second, the settled bed obtains a higher rotating energy as the acceleration (i.e., drum rotation speed) increases. The combination of these two factors requires a higher impact pressure in order for the water sprays to flush down the settled bed of heavy particles completely. Continuous concentration is achieved as the concentrate weight approaches an invariable value with the increase of water spray pressure.

Figure 5 — Effect of centrifugal acceleration on performance. Conditions: water spray pressure = 0.48 MPa, feed volume flow rate = 9.8 m3/h, feed % solids = 32%, feed grade = 52.17% Fe, feed size = 91.68% (-74 μm), water spray reciprocation = 50 mm/s, water spray spacing = 100 mm.

Table 1 — Effects of feed volume flow rate and feed % solids on performance.

Operating parametersFeed grade,

% Fe

Concentrate

Weight,%

Grade,% Fe

Iron recovery,%

Feed volume flow

rate (m3/h)

(Feed % solids,

31.30 - 31.78%)

4.56 52.45 78.36 55.29 81.06

6.34 52.18 70.83 56.07 74.68

8.52 52.75 63.61 58.08 68.73

9.53 52.70 61.85 58.77 67.69

10.60 52.56 61.81 58.36 67.35

11.48 52.54 61.29 58.22 66.65

12.66 52.39 58.89 58.36 64.49

14.41 52.67 53.50 58.27 58.09

Feed % solids (%)

(Feed volume

flow rate, 9.80 -

9.85 m3/h)

16.89 52.25 48.36 58.99 52.59

20.56 52.38 56.51 58.41 61.15

25.28 52.15 59.61 58.17 64.04

29.73 52.40 60.88 58.00 64.91

33.56 52.62 61.63 58.22 65.69

35.13 52.45 62.21 57.35 65.52

40.37 52.35 67.07 56.61 69.86

Conditions: feed size = 90% (-74 μm), centrifugal acceleration = 48 g, water spray reciprocation = 50 mm/s, water spray spacing = 100 mm, water spray pressure = 0.45 MPa.


Therefore, continuous concentration is possible only when the water sprays have a sufficiently high pressure. Furthermore, the critical spray pressure for continuous concentration is related to the iron grade and particle size of the feed material, because these two variables also contribute to the centrifugal force acting on particles in the flowing film.

Effect of centrifugal acceleration on performance. Centrifugal acceleration has the most significant effect on the separation performance of an enhanced gravity concentrator, because it determines the centrifugal force on the particles to be separated. Therefore, the effect of centrifugal acceleration on the separation performance of the concentrator was studied during the industrial test, with feed volume flow rate and feed % solids separately fixed at around 10 m3/h and 32%. As seen in Fig. 5, the centrifugal acceleration has a very significant effect on the performance of the concentrator, particularly on iron recovery. As the acceleration increases from 20 g to 60 g, concentrate grade increases, reaches a plateau and then begins to decrease, while iron recovery increases throughout; the effect of acceleration on recovery is more pronounced at low levels. The highest performance is achieved when the concentrator operates at an acceleration of 40 - 48 g, where the concentrate grade and recovery vary in the narrow regions of 58.67 - 58.86 % Fe and 67.73 - 68.44 %, respectively; beyond the 40 - 48 g range, the grade deteriorates markedly.

As explained elsewhere (Chen Luzheng et al., 2008), ac-celeration is the dominant force determining the number of particles going into the concentrate. When operated at an insufficient acceleration, the concentrator mainly concentrates partially liberated larger hematite minerals, which results in a low concentrate grade and recovery. The centrifugal force acting on particles increases as the acceleration increases, and more fully liberated finer hematite minerals go into the concen-trate and the performance improves tremendously. However, when the concentrator is operated at an excessive accelera-tion, as illustrated in Fig. 5, the concentrate grade decreases, as many lighter gangue minerals go into the concentrate, due to their being subjected to an excessively strong centrifugal force. Combining the results of the pilot-scale and full-scale concentrators in reconcentrating this fine hematite, the effect of centrifugal acceleration on the performance matches well.

Effect of feed volume flow rate and feed % solids on performance. The effect of feed conditions, i.e., feed vol-ume flow rate and feed % solids, on the performance of the concentrator operating at an optimum acceleration of 48 g and at a sufficient water spray pressure of 0.45 MPa was studied. From the separation results in Table 1, feed volume flow rate and feed % solids both have significant effects on performance. An increase in feed volume flow rate increases the concentrate grade but reduces the recovery rapidly for a steady feed % sol-ids of around 31.50% and the highest performance is achieved when the feed volume flow rate is in the narrow region of 9.53 - 10.6 m3/h; increasing feed % solids obviously decreases the concentrate grade but increases the recovery, for an optimized feed volume flow rate of around 9.8 m3/h.

Increasing the feed volume flow rate over an inclined deck increases the flow depth and the axial velocity of the film (Ergün and Ersayin, 2002), thereby affecting the settling trajectories of particles in the film. An insufficient feed volume flow rate decreases the flow depth and axial velocity of the film, result-

ing in the settling of partially liberated hematite and gangue minerals onto the drum, which decreases the concentrate grade and increases the recovery. As the feed volume flow rate increases, the flow depth and axial velocity of the film increase, resulting in less time for the settling of well-liberated fine hematite minerals. This effect reduces the recovery but increases concentrate grade. Thus, a particular feed volume flow rate is needed to achieve selective concentrations of iron values from slurry.

It was demonstrated during the industrial test that the op-timum feed volume flow rate is closely correlated with feed % solids. In a limited range, an increase in feed volume flow rate allows for an increase in feed % solids, thus improving processing capacity. As illustrated in Table 1, when the feed volume flow rate is restricted within the range of around 9.8 m3/h, feed % solids may vary from 25.28 to 33.56%, with almost no negative effect on separation results. This action allows for a high flexibility in operating the concentrator. In addition, it should be noted that the optimum feed volume flow rate and feed % solids are dependent on such factors as the characteristics of the feed (e.g., iron grade and particle size) and the centrifugal acceleration. For example, increasing the centrifugal acceleration increases the settling velocity of very fine particles in the flowing film and thus the concentrator achieves a higher processing capacity.

Analyzing the results in Table 1, it is concluded that a com-bination of a feed volume flow rate of 8.5 - 11.5 m3/h and feed % solids of 25.5% - 33.5% is suitable for the materials tested in the present investigation, with the processing capacity of the concentrator varying from 2.68 to 5.16 t/h.

Effect of water spray reciprocation on performance. As noted above, continuous concentration is achieved using reciprocating water sprays to remove the settled bed of heavy particles on the drum surface. As the reciprocation and spacing

Figure 6 — Effect of water spray reciprocation on perfor-mance. Conditions: feed volume flow rate = 9.8 m3/h, feed % solids = 32%, feed grade = 51.23 % Fe, feed size = 91.5 % (-74 μm), centrifugal acceleration = 48 g, water spray spac-ing = 100 mm, water spray pressure = 0.45 MPa.


(or number) of water sprays are the two dominant operating parameters affecting the performance of the concentrator in terms of the reciprocation, their effects on the performance were studied in detail.

As seen in Fig. 6, a suitable reciprocation improves the concentrate grade significantly but slightly reduces the recov-ery. The grade increased and reached a high point of 58.66% Fe with 68.82% recovery at a reciprocation of 40 mm/s and then decreased as the reciprocation progressed. Recovery decreased continually with the increase of the reciprocation from 20 to 60 mm/s.

The reciprocation has a profound effect on the thickness distribution of the settled bed on the drum surface and hence generates a significant influence on the characteristics of flowing film; this in turn has a significant effect on the performance of the concentrator. A more detailed description on the effect of the reciprocation of water sprays on the separation performance of a continuous concentrator was reported by Chen Luzheng et al. in 2008.

Effect of water spray spacing on performance. When the reciprocation of water sprays is fixed at the optimum value of 40 mm/s, the effect of the spacing of the sprays on separa-tion performance is shown in Fig. 7. The spacing is changed by altering the number of nozzles fixed on the reciprocating mechanism of the concentrator, as schematically presented in Fig. 2. It should be noted that, in order to flush down the settled bed of heavy particles on the drum surface completely, the nozzles on the mechanism are designed to be capable of traveling through the whole separation width of the drum.

As can be seen from Fig.7, the spacing of water sprays has a significant effect on the separation performance of the con-centrator. An increase in the spacing increases the concentrate grade but reduces the recovery; this result is due to the fact that the spacing determines the quantity of light particles to be flushed into the concentrate, as the water sprays reciprocating at

a fixed reciprocation remove the settled bed on the drum surface.As mentioned above, continuous concentration is achieved

using reciprocating water sprays to remove the settled bed of heavy particles on the drum surface. Most light and partially liberated particles escape being flushed into the concentrate via the spacing between the sprays (schematically illustrated in Fig. 1) and instead flow out as tailings, with a very few such particles inevitably flushed down into the concentrate. In other words, as shown in Fig. 8, each spray of water creates a barrier; these barriers decrease as the spacing of water sprays increases, because the number of sprays decreases with increasing spac-ing. Therefore, it is apparent that the quantity of light particles flushed down by the sprays decreases with the increasing spacing of water sprays, resulting in the improvement of concentrate grade and the reduction of iron recovery (Fig. 7).

The test results indicate that when the spacing is adjusted from a minimum of 50 mm to a maximum of 100 mm, the concentrate grade increases from 55.69 % Fe to 59.48 % Fe, while the recovery drops from 75.38 to 67.86 %. This test confirmed that the negative flushing effect of the water sprays can be minimized by optimizing the water spray reciprocation.

From Fig. 7, it is predicted that a higher concentrate grade may be achieved if a spray spacing larger than 100 mm is used, with a corresponding reduction of recovery. Also, a larger spac-ing of the sprays necessitates the redesign of the reciprocating mechanism in the concentrator.

Evaluation of full-scale continuous centrifugal concentrator

In summary, when all variables were set at optimum lev-els, this full-scale concentrator achieved a concentrate grade improvement of 6 - 7% Fe with 66 - 69% recovery, which is slightly inferior to the best separation result (concentrate grade improvement of 9.9 % Fe with 65.02 % recovery) found with the pilot-scale concentrator in reconcentrating this fine hematite, as illustrated in Table 2.

Figure 7 — Effect of water spray spacing on performance. Conditions: feed volume flow rate = 9.82 m3/h, feed % solids = 32.3%, feed grade = 51.63% Fe, feed size = 90.7% (-74 μm), centrifugal acceleration = 48 g, water spray reciprocation = 40 mm/s, water spray pressure = 0.45 MPa.

Figure 8 — Photographic diagram of barriers generated by the water sprays (reciprocating at an excessively low rate of 10 mm/s) in the film.


This inferiority is the result of the fluctuations involved in the feed material during the industrial test, notably iron grade and particle size as described above. As such, it is the result of the limitations in the scale-up design of the concentrator. As shown in Fig. 7, a higher concentrate grade may be achieved with a larger water spray spacing. Fig. 9 compares the separation results of two full-scale concentrators of different radial and axial vibrations at the tailings recovery plant; they are operated under the same conditions, i.e., feed volume flow rate around 9.6 m3/h, feed % solids around 31.0 %, feed grade around 51.0% Fe, feed size around 90.0 % below 74 μm, water spray reciprocation at 40 mm/s, water spray spacing at 100 mm and water spray pressure of 0.45 MPa. Obviously, drum vibration has a significant influence on the iron recovery; however, it has no marked effect on concentrate grade.

Another reason for the inferiority is that the tailings pumped to the recovery plant have become more complicated and difficult to treat than those tested in the pilot-scale concentra-tor, with a sulfur-rich lean hematite ore processing plant of HGMS-reverse flotation flowsheet going into operation in late 2007. This plant contributes about 40% by weight of the total tailings to the recovery plant, reducing iron grade and particle size of the tailings fed to the recovery plant.

This full-scale concentrator was also tested for its abil-ity to reconcentrate the fine hematite minerals from HGMS concentrate from a refractory hematite ore discarded by the Donganshan processing plant (Anshan Iron and Steel Com-pany, northeast China). The feed to the concentrator is also characterized by wide fluctuations in iron grade and particle size (Chen Luzheng et al., 2007). A typical separation result for this hematite ore is illustrated in Table 2. As can be seen, the concentrator achieved a better performance (concentrate grade improvement of 12.02% Fe with 79.41% recovery) compared with those of the concentrators in reconcentrating the fine hematite tailings. Apparently, the separation performance of this concentrator is in close correlation with the characteristics of the feed material.

The performance of the full-scale concentrator cannot be compared to the pilot-scale trials and to its performance in reconcentrating the refractory hematite ore; however, it produces a high-grade marketable concentrate in the recovery plant. Although the concentrator has not reached its optimum scale-up design and the performance may be improved as discussed above, it provides a potential way to achieve con-tinuous concentration and to obtain a high-grade concentrate

from tailings.

ConclusionsIt can be seen from the discussion above that:- Continuous separation is achieved only when the recipro-

cating water sprays have a sufficiently high pressure to flush down the settled bed of heavy particles on the drum surface. The maximum critical water spray pressure necessary for achieving continuous separation increases with increasing centrifugal acceleration, as the aggregating force between heavy particles in the settled bed and the rotating energy of the bed increase with the increase of centrifugal acceleration.

- As with the pilot-scale concentrator, the centrifugal ac-celeration has a very significant effect on the performance of the full-scale concentrator, particularly on iron recovery, because it determines the centrifugal force on particles to be separated and thus presents the dominant force determining the number of particles going into the concentrate. Also, the highest performance was achieved when the concentrator operates at an acceleration of 48 g.

- Feed volume flow rate and feed % solids affect the charac-teristics of the flowing film significantly, generating remarkable

Table 2 — Performance of pilot-scale and full-scale.

ConcentratorGrade, % Fe

Weight, % Iron recovery, % Efficiency, % Feed typeFeed Concentrate Tailings

Full-scale a 49.05 61.07 27.88 63.78 79.41 52.22Refractory

hematite ore

Pilot-scale 52.42 62.32 40.47 54.69 65.02 41.13 Fine hematite

tailingsFull-scale b 51.37 58.64 41.22 58.27 66.52 30.99a Typical separation achieved by a full-scale concentrator with refractory hematite ore (Anshan Iron and Steel Company) b Average separation achieved by ten full-scale continuous concentrators over 72 hours with fine hematite tailings (Hainan Iron

and Steel Company).

Figure 9 — Effect of drum vibration on performance: 1) Vibration: 1#, radial 7-10 mm, axial 6-8 mm; 5#, radial 14-18 mm, axial 10- 15 mm.


influences on the performance of the concentrator. In a limited range, the optimum feed volume flow rate is closely correlated with feed % solids; an increase in feed volume flow rate al-lows for an increase in feed % solids and thus an improved processing capacity.

- The reciprocation and spacing of water sprays have signifi-cant effects on the concentrate grade. The reciprocation affects the thickness distribution of the settled bed on the drum surface and the characteristics of flowing film, which in turn have a significant effect on the performance of the concentrator. An optimized reciprocation improves the concentrate grade sig-nificantly but slightly reduces the recovery. Also, an increase in the spacing increases the concentrate grade but reduces the recovery, as the spacing determines the quantity of light par-ticles to be flushed into the concentrate; a higher concentrate grade may be achievable if a spacing larger than 100 mm (the maximum of the concentrator) is used, although there will be a corresponding reduction in the recovery.

- Due to the limitations in the scale-up design, the fluctua-tions of the feed material, it was not possible for the concen-trator to obtain a separation performance comparable to the pilot-scale concentrator in reconcentrating fine hematite or to its own performance in reconcentrating refractory hematite ore in earlier studies. However, the continuous concentrator produces a high-grade marketable concentrate from the tailings in the recovery plant.

Although the performance of the full-scale concentrator is slightly inferior to pilot-scale trials in reconcentrating the fine hematite, it is still significant compared with other separation methods (such as flotation), due to its low cost, simple operation and flexibility. As the concentrator is not in its optimum scale-up design and the performance may be improved as discussed above, it is thus concluded that it provides a potential way to achieve continuous concentration and to obtain a high-grade concentrate from tailings.

AcknowledgmentsThe project was supported by the Educational Key Program

of Yunnan Province of China (Grant No. 09Z0014), the Yunnan Provincial Foundation for Applied Basic Research (Grant No. 2009ZC007X) and the key program of the National Natural Science Foundation of China (grant No. u0837602.)

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Copper losses in sulfide concentrate smelting slag are dependent on slag compositionŽ. ŽivkovicProfessor, University of Belgrade, Technical faculty in Bor, Bor, Serbia

N. MitevskaResearch engineer, Mining and Metallurgy Institute, Bor, Serbia

I. MihajlovicProfessor, University of Belgrade, Technical faculty in Bor,Bor, Serbia; E-mail: [email protected]

Đ. NikolicProfessor, University of Belgrade, Technical faculty in Bor, Bor, Serbia

AbstractThe silicate slag and copper matte that are the byproducts of smelting copper in a reverberatory furnace contain SiO2, FeO, Fe3O4, CaO and Al2O3, and Cu, Fe and S, respectively. These components influence the oxide (Cuox), sulfide (Cusul) and total (Cutotal) copper content lost in the waste slag. This paper uses multiple linear regression analysis (MLRA) to determine the degree of influence. The results of calculations obtained using MLRA were compared to actual measured results and found to have a large degree of correspondence, with the coefficient of determination R2 = 0.975. These results indicate that slag composition and copper content in the matte influence the copper losses in the waste slag with a probability of above 97%. It was determined that, of the total copper content of the slag, 74% is in sulfide and 26% in oxide form, and that this ratio is not dependent on the slag or matte composition.

Paper number MMP-09-047. Original manuscript submitted October 2009. Revised manuscript accepted for publication March 2010. Discussion of this peer-reviewed and approved paper is invited and must be submitted to SME Publications Dept. prior to February 28, 2011. Copyright 2010, Society for Mining, Metallurgy, and Exploration, Inc.

IntroductionDuring pyrometallurgical copper extraction, smelt-

ing follows the oxidative roasting of the concentrates. In the modern pyrometallurgical processes of copper extraction, both operations are performed in the same unit, resulting in the development of two main prod-ucts: silicate slag and copper matte. These products are differentiated along the melting zone length by gravitational settling, according to their specific gravities. Partially dissolved copper and unsettled copper matte droplets present in the waste slag result in copper losses (Biswas and Davenport, 2003). With stationary smelting processes, such as smelting in a reverberatory furnace, this loss is about 0.5% Cu and is permanent, because such slag is disposed of in the waste yard. With dynamic smelting processes such as Flash, INCO, etc., the copper content in the silicate slag can reach 2%, an amount that demands further processing of the slag. Processing of such copper-rich slag is facilitated using flotation or electroprecipitation of remaining sulfides, resulting in about 0.5% Cu in

the remaining waste slag (Živković and Savović, 1996; Sarrafi et al., 2003; Habashi, 2007). Global annual slag production resulting from copper smelting operations is above 25 tons (Gorai and Jana, 2003). Thus, copper loss to waste slag strongly influences the worldwide rise of the copper extraction process (Maweja et al., 2009).

Numerous investigations have revealed that slag composi-tion has a dominant influence on its copper content (Matousek, 1991; Imriš et al., 1997; Jalkanen et al., 2003; Goni and Sanchez, 2009). The slag is a silicate solution, with SiO2 and FeO as dominant constituents. Accordingly, the prevailing compound in the slag is fayalite (2FeO.SiO2). Copper losses in silicate smelting slag depend on a large number of factors, ranging from objective ones, such as the composition of the input materials or the overall physicochemical structure of the smelting products, to subjective factors, such as those depending on the management of the process (Jalkanen et al., 2003). In recent investigations it was determined that there is a connec-tion between individual components of the slag composition, copper content in the matte and the copper content in the slag (Živković et al., 2009). A relationship between the copper in

Key words: Smelting, Slag, Copper/ copper ores, Copper losses, Multiple linear regression analysis

´

´

´


the slag, as a function of the input materials composition, and the copper in the matte during copper sulfide smelting has also been derived by Sridhar et al. (1997).

The motive for the investigations presented in this paper was to perform statistical analysis of the slag constituents that influence total copper losses to the smelting slag and to develop a mathematical model describing these influences. With the aim of obtaining the most reliable results possible about the conditions of an industrial furnace with stable operation, the complex sampling operation was performed at the industrial level. The results obtained should be helpful for copper smelter management related to the input charge composition that will enable better management of the copper losses in the smelting slag, due to the fact that regulation of input charge composition can be used to control copper losses in the smelting slag. Also, the degree of desulfurization obtained during oxidative roasting can be used to control the copper content in the matte (Reuter et al., 1992; Aldrich et al., 1994; Magaeva et al., 2000; Cocić et al., 2007; Liu et al., 2009; Živković et al., 2009).

ExperimentalFor the analysis of the influence of slag composition on

copper losses, we used data obtained from the Copper Smelter Company – RTB Bor, (Serbia). The smelting process in this company is done in a reverberatory furnace with an annual capacity of 60,000 tons of anodic copper. The sampling of the slag and matte was performed during a tapping operation and the analysis of the composition was carried out on a daily basis for several years (Živković and Savović, 1996). To obtain a clear picture of the real conditions in the reverbera-tory furnace, a special experiment was designed that included sampling of the slag and matte from the furnace roof using so-called deep sounding during stable operation of the furnace (Mitevska, 2000).

The position of the measuring points (MM) (which are the openings for sampling formed at the furnace roof) and the sampling probe shape are presented in Fig 1. The sampling probe is a pipe made of steel (1), having an inner diameter of 12.7 mm and a length of 6,000 mm. The plug (2) is positioned at the lower band of the pipe, which is connected to the steel

Figure 1 — Schematic presentation of the sampling places (MM=measuring points) in the reverberatory furnace (a) and schematic presentation of the sampling probe (b).

Table 1 — Concentrations of the main slag and matte constituents in the samples collected along the furnace length (mea-

suring points MM3-MM6) and at the different distances from the slag surface (samples I – IV).

Sample

MM3 MM4 MM5 MM6

I II III IV I II III IV I II III IV I II III IV

Smelting slag

Cusulfide [%] 0.32 0.34 0.40 0.42 0.35 0.38 0.48 0.54 0.46 0.47 0.55 0.62 0.50 0.51 0.59 0.67

Cuoxide [%] 0.12 0.13 0.14 0.16 0.13 0.14 0.15 0.17 0.14 0.15 0.17 0.18 0.15 0.16 0.18 0.20

CuΣ [%] 0.44 0.47 0.54 0.58 0.48 0.52 0.63 0.71 0.60 0.62 0.72 0.80 0.65 0.67 0.77 0.87

SiO2[%] 33.62 33.38 32.78 32.62 33.10 32.70 32.20 32.06 31.94 31.80 31.48 31.02 31.08 30.96 30.78 30.40

FeO [%] 42.26 42.73 43.28 43.56 42.86 43.18 43.69 44.68 44.08 44.28 44.84 46.06 44.75 45.03 45.91 46.69

Fe3O4 [%] 7.14 7.29 7.38 7.42 7.31 7.38 7.42 7.49 7.43 7.47 7.53 7.61 7.55 7.57 7.66 7.68

CaO [%] 4.95 4.92 4.87 4.76 4.90 4.84 4.76 4.67 4.73 4.70 4.67 4.63 4.63 4.62 4.60 4.56

Al2O3 [%] 3.65 3.59 3.51 3.46 3.55 3.51 3.44 3.42 3.44 3.41 3.39 3.35 3.35 3.31 3.27 3.24

Copper matte

Cu [%] 36.31 39.66 40.84 42.12

Fe [%] 34.91 31.58 30.65 29.44

S [%] 26.13 25.80 25.67 25.45


wire that goes through the pipe. The probe is slowly lowered from the furnace roof and sunk in the melt at the location of each measuring point (MM) until it reaches the bottom of the furnace. During this process, the probe is filled with the mol-ten slag and the matte, which enters through the gap between the pipe and the plug. It takes 30 seconds for the probe to reach the bottom of the furnace, the plug to close under pres-sure and the probe to lift through the furnace roof. Because the temperature was too high at measuring points MM1 and MM2, they were omitted and measuring was performed only at points MM3-MM6.

After cooling in the air, the probe was sliced transversely to determine the border between the slag and the matte. From each probe, four slag samples were obtained. Slag samples were collected from the surface of the slag (I); at the distance of 60 mm below the surface of the slag (II); at the distance of 120 mm below the surface of the slag (III) and at the distance of 180 mm below the slag surface (IV). The matte sample was obtained at 10 mm below the border between the slag and the matte. The thickness of each sample for chemical analysis was 10 mm.

The technological parameters of the furnace operation, for the day in which the sampling was performed, were:

- 570 tons of produced copper matte with a copper content of 42.12% Cu in the cumulative sample,

- 940 tons of slag with 0.49% Cu in the cumulative sample. These parameters indicate stable furnace operation.For a quantitative determination of the elements with a con-

centration above 0.1%, an optical emission spectrograph OES JARRELL -Ash 70.000 (USA) was used, while for the elements with a concentration below 0.1%, a mass spectrograph JOEL JMS-0.1 MB (Japan) with the ionic optical system Matauh- Herzog was used. The concentration of the main constituents of the slag and the matte is presented in Table 1.

In accordance with the results presented in Table 1, it is clear that the composition of the slag and the copper matte influences the copper losses in the slag. Also, copper losses in the slag are related to the position of the measuring point (MM) in the furnace. The measuring point MM3 is located closer to the front wall of the furnace, compared to the other measuring points. The burners are located in the front wall of the furnace, near the MM1 position. At this location, the gas temperatures in the furnace are decreasing from the measuring point MM3 to the measuring point MM6. Average gas temperatures in the reverberatory furnace measured during sampling were: 1,510° C at MM3, 1,485°C at MM4, 1,420°C at MM5 and 1,405°C at MM6. According to the copper contents in the slag detected at different measuring points (see Table 1), it can be stated that cooler molten material affects the melt’s viscosity, thereby influencing slag losses with respect to copper content. Thus, the highest copper content is detected at MM6 and the lowest at MM3. For processing of the obtained results, presented in Table 1, with the aim of determining the analytical depen-dence: (Cu) = f (slag composition, matte composition); for each particular component and for whole set, multiple linear regression analysis (MLRA) was used.

Results and discussionResults presented in the literature indicate that the chemical

composition of the slag has a strong influence on its copper content (Gorai and Jana, 2003). The copper smelter plant in Bor has had the same balance of the main slag constituents for many years: SiO2: 30 – 35%; FeO: 40- 50 %; Fe3O4: 5 – 8

%; CaO : 4 – 6 % and Al2O3 : 3 – 6 %. Copper content in the matte ranges from 30 to 45% (Živković and Savović, 1996). Results obtained in the experiments and analyzed in this paper are within these ranges, indicating that they can be regarded as representative for analyses described in the following text.

According to the results presented in Table 1, copper in the slag is detected in the oxide (Cu2O) and the sulfide (Cu2S) forms. Metallic copper wasn’t detected in the slag, since the reaction for Cu2O reduction requires mattes with copper content above 60% (Jalkanen et al., 2003).

The influence of each individual constituent on the slag’s copper content was analyzed using linear regression method-ology. This way, the dependence obtained is in the form of a linear equation:

(Cu) losses = A · (% compound) + B (1)

Where (Cu) losses represents the copper losses in the slag (oxide, sulfide and total), % compound is the concentration of the individual slag constituent (SiO2, FeO, Fe3O4, CaO and Al2O3) and A and B are the constants of the linear equation.

Equation (1) was also used to define the influence of the copper matte composition on copper losses in the slag. In this case, % compound represents the concentration of the individual matte constituent (Cu, Fe and S).

The influence of the SiO2 content in the slag. The SiO2 and FeO represent more than 70% of the slag content and are bound in the form of fayalite (2FeO.SiO2). Figure 2 presents the dependence of the copper content in the slag on SiO2.

Results presented in Fig. 2 suggest that an increase in the SiO2 content leads to a decrease in the copper content of the slag. This kind of dependence results from moving the position of the equilibrium of the reaction, which happens at the phase border between slag and matte, to the right:

[FeS] + 3(Fe3O4) + 5(SiO2) → 5(2FeO.SiO2) + SO2 (2)

Thus, a decrease in the oxidation potential of the melt is occurring that leads to a decrease in the oxidation rate of

Figure 2 — Dependence between the copper content in the smelting slag and its SiO2 content.


the copper sulfide oxidizing to Cu2O and, accordingly, to a decrease in the amount of chemically dissolved copper in the slag (Gorai and Kana, 2003).

The influence of the FeO content in the slag. An increase

in the FeO content in the slag moves the equilibrium of Eq. (3) to the right, which is in the direction of increased copper solubility in the slag (Biswas and Davenport, 2003):

[Cu2S] + (FeO) → (Cu2O) + [FeS] (3)

Figure 3 presents the dependence between the copper content in the slag and the FeO content.

The influence of the magnetite content in the slag. Mag-netite is an unwanted component in the copper slag, formed during smelting process, but is also added with the returned converter slag (Živković and Savović, 1996). As Fe3O4 increases in the slag, the equilibrium of Eq. (4) moves to the right, which increases the oxidation potential of the slag, and in this way the chemical solubility of the copper in the slag becomes higher.

[Cu2S] + 3(Fe3O4) → (Cu2O) + 9FeO + SO2 (4)

The dependence between copper and magnetite content in the slag, which indicates an increase of the Cu content resulting from the increase in Fe3O4, is presented in Fig. 4.

Figure 3 — Dependence between copper loss and the FeO content in the slag.

Figure 4 — Influence of Fe3O4 content on copper loss in the slag.

Figure 5 — Influence of the CaO content on copper loss in the slag.

Figure 6 — Influence of Al2O3 on copper loss in the slag.


The increase in the Fe3O4 content of the slag, besides mov-ing Eq. (4)’s equilibrium to the right, leads to the separation of the magnetite crystals, which causes heterogenization of the silicate melt and a sudden increase in its viscosity, and in this way increases the mechanical copper losses to the slag.

Influence of the CaO content in the slag. An increase of the CaO content in the slag leads to the destruction of its silicate network (Živković and Savović, 1996; Biswas and Davenport, 2003), which results in a decrease in the slag’s viscosity and copper content. In the case of the system described in this study, Fig. 5 presents the dependence between the copper losses and the CaO content in the slag.

Influence of the Al2O3 content in the slag. Figure 6 presents the dependence between copper losses and the Al2O3 content in the slag.

In principle, an increase in the Al2O3 content of the slag leads to formation of Al-O complexes because of the high ionic potential of the Al3+ ions. This formation amplifies the effect of the Si-O complexes, which is why Al2O3 can be defined as a network-building compound (Kim and Sohn, 1998).

The Al-O complexes formed under the influence of Al2O3 decrease the oxidation potential of the copper melt and accord-ingly the amount of the copper oxides that are easily dissolved in the slag. At the same time, with development of the Al-O

complex, interphasial tension increases between the matte and the slag, which leads to a decrease in the mechanical copper losses in the slag.

Using the linear regression analysis, the dependence between the content of each individual slag constituent and its oxide, sulfide and total copper content is obtained in the linear form (Eq. 1). The values of corresponding constants A and B, as well as the values of the coefficient of determination (R2) for obtained linear dependences, are presented in Table 2.

Calculated values of the coefficient of determination (R2) (Table 2) indicate a good fit to the obtained results for all slag compounds. Accordingly, it is clear that the content of each individual slag constituent has a significant influence on the copper content in the slag.

Influence of the matte composition on the copper losses to the slag. Part of the copper matte is kept in the slag in the form of unprecipitated droplets. If these drops have a larger copper content, the copper loss to the slag also increases. The content of iron and copper in the copper matte (system Cu2S – FeS) are in inverse proportion. For that reason, the influence of iron in the matte on copper content in the silicate slag is the opposite of the influence of copper content in the matte. The sulfur content in the matte depends on its Cu/Fe ratio, since one mole of Fe is attached to one mole of sulfur, while two moles of Cu are attached to one mole of sulfur.

Table 2 — Calculated values of the constants in the linear dependences between the copper content in the slag and

the slag and matte constituent concentration.

Component Coefficients Cuox Cusulf Cutot

Co

mp

on

ents

in t

he

slag

SiO2 A - 0.018 - 0.097 - 0.116

B 0.752 3.604 4.357

R2 0.725 0.869 0.855

FeO A 0.015 0.077 0.092

B - 0.530 - 2.934 - 3.464

R2 0.885 0.960 0.960

Fe3O4 A 0.137 0.676 0.814

B - 0.869 - 4.573 - 5.443

R2 0.828 0.882 0.884

CaO A - 0.153 - 0.779 0.932

B 0.879 4.166 5.046

R2 0.775 0.880 0.874

Al2O3 A - 0.168 - 0.842 - 1.010

B 0.732 3.365 4.097

R2 0.783 0.858 0.856

Co

mp

on

ents

in t

he

cop

per

mat

te

Cu A 0.005 0.034 0.04

B - 0.075 - 0.885 - 0.960

R2 0.356 0.547 0.518

Fe A - 0.006 - 0.036 0.042

B 0.0346 1.613 1.959

R2 0.347 0.534 0.506

S A - 0.052 - 0.303 - 0.355

B 1.496 8.28 9.777

R2 0.376 0.557 0.531


The dependences between the individual matte constitu-ents (Cu, Fe and S) and the copper losses in the slag are also calculated using linear regression (Eq.1) applied to the data from Table 1. The values of corresponding constants A and B, as well as the values of the coefficient of determination for obtained linear dependences, are presented in Table 2.

According to the results presented in Table 2, the degree of fitting is lower (lower values of the R2 coefficients), in all cases, for the matte compounds compared to the slag compounds. The reason for such behavior could be in the different influences of the slag and the matte composition on the copper losses in the slag. The slag’s constituents determine its copper losses, mainly through their influence on slag viscosity. On the other hand, copper matte constituents influence copper loss in the slag via their relation toward the copper concentration in the part of the matte that remains dissolved or entrained in the slag. Furthermore, it should be noted that the amount of data avail-able for analysis was smaller for the copper matte constituents than for the constituents of the slag (Table 1).

Determination of the copper content in the slag, as the function of the slag and the matte composition, using the multiple linear regression analysis (MLRA). Calculated dependences for the copper content in the slag as a function of the proportion of individual slag components, while the content of all the other compounds is regarded as constant, can result only in approximate information on each compounds’ influence. However, because the slag is composed of complex compounds and its composition changes constantly, even outside the narrow borders defined by the technology of the smelting operation,

important information can be obtained using statistical analysis of the content of all compounds in the slag as well as copper, iron and sulfur content in the matte, considering that the Eqs. (2), (3) and (4) happen at the slag-matte phase border. For determination of the quantitative dependence of the copper content in the slag on slag composition and content of the Cu, Fe and S in the matte, multiple linear regression analysis (MLRA) was used as part of the SPSS software application. The MLRA defines the output of the linear correlation of one variable (copper content in the slag) depending on the slag composition (content of slag constituents) and the content of the Cu, Fe and S in the matte. This dependence can be presented in the following form:

(Cu) = a + b(SiO2) + c( FeO ) + d(Fe3O4) + e(CaO) +

f(Al2O3) + g[Cu] + h[Fe] + k[S] (5)

Where: a, b, c, d, e, f, g, h and k are the coefficients in the linear regression equation, and ( ),[ ] are the % concentrations of the slag and matte components.

Obtained values for the coefficients in Eq. (5), according to the MLRA, are presented in Table 3.

Those dependences enable prediction of the copper content in the slag according to its composition and according to the composition of the matte. Using MLRA in the SPSS software, testing of the above model’s validity (Eq. (5)) was performed. For the testing operation, calculated values were compared to those experimentally obtained. Dependence between mea-sured and calculated values of the copper content in the slag is presented in Fig. 7.

Calculating the values of the coefficients of determina-tion (R2) indicate a high degree of fit in all three cases. Also, the copper content in the waste slag can be predicted with a probability of above 97%, if the composition of the slag and the matte is known. Considering that the composition of the charge can be controlled, as well as the degree of desulfurization during oxidative roasting (which determines the composition of the matte), copper losses in the slag could be managed us-ing by exploiting these dependent relationships (Eq. 5) with an accuracy of 99.1%. This step would improve the overall management of the pyrometallurgical copper extraction process at the smelting stage. The remaining 0.9% can be regarded as under the influence of other parameters, mostly related to the dynamics of smelting operations in the furnace, which strongly rely on human factors.

Conclusions1) After examination of the influence of individual slag

constituents in the reverberatory furnace smelting slag (SiO2; FeO; Fe3O4; CaO; Al2O3) and Cu, Fe and S content in the copper matte, on copper content in the waste slag (Cu) using the linear regression method, corresponding coefficients of determination (R2) were obtained.

2) Obtained values of the coefficients of determination

Table 3 — Calculated values of the constants in the linear regression equation (Eq. 5), obtained by MLRA methodology.

Coefficient a b c d e f g h k R 2

Cuox 0.486 0.016 0.019 - 0.004 - 0.045 - 0.126 - 0.027 - 0.030 0.040 0.99

Cusulf 3.208 - 0.083 0.067 - 0.237 - 0.119 0.068 - 0.148 - 0.189 0.424 0.975

Cutotal 3.694 - 0.067 0.085 - 0.241 - 0.163 - 0.058 - 0.175 - 0.219 0.464 0.991

Figure 7 — Dependences between measured and calcu-lated values for the copper content in the slag using MLRA methodology ( — ideal; ○ - oxide; × - sulfide; □ - total )


indicated that the prediction of the influence of each slag com-ponent’s concentration on its copper content can be accurate (above 80%).

3) The prediction of the copper losses in the slag as the function of the copper matte composition has an accuracy above 50%.

4) When considering the influence of all slag and matte constituents, at the same time, using MLRA methodology, correlation among measured and calculated values of copper content in the waste slag is obtained with the determination coefficients (R2) equal to 0.99 for the (Cu)ox, 0.97 for the (Cu)sul and the 0.991 for the (Cu)total. This result indicates that management and control of copper loss to the slag could be achieved at an accuracy of over 99.1% through the control of the input charge composition and the degree of desulfurization during oxidative roasting.

5) Considering that results presented in this paper are ob-tained during the relatively stationary mode of furnace opera-tion, under standard industrial conditions, it can be stated with a probability of 99.1% that the slag and copper matte compo-sition influence the copper losses in the slag. The remaining influences mostly include human factors in the management of the technological process.

6) It was also determined that the ratio of sulfide to oxide copper content in the slag is 2.8:1. This ratio is not dependent on the slag or matte composition during the smelting of sulfide concentrates.

7) Copper losses in the slag are also influenced by the temperatures in the furnace. Cooler molten material affects the melt’s viscosity, thereby influencing slag losses with respect to copper content.

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Laboratory study of an organic binder for pelletization of a magnetite concentrate S. L. de MoraesResearch engineer, Environmental and Energy Technologies Center, Institute for Technological Research, Sao Paulo, Sao Paulo, Brazil

S. K. KawatraProfessor, Department of Chemical Engineering, Michigan Technological University, Houghton, Michigan

AbstractThis study aimed to identify a way to reduce the use of bentonite in the pelletization of magnetite. With this goal, different combinations of binders were compared to bentonite by examining the quality of pellets obtained by balling drum agglomeration. Pellets were subjected to routine tests that included simple compression of the wet and dry 105°C pellets, dropping wet pellets to determine their ability to survive handling and submitting pellets to thermal shock to determine how they tolerate drying and preheating. The best results use sodium silicate (1.5%) as a binder and show that it is possible to pelletize iron ore without using bentonite.

Paper number MMP-09-029. Original manuscript submitted July 2009. Revised manuscript accepted for publication De-cember 2009. Discussion of this peer-reviewed and approved paper is invited and must be submitted to SME Publications Dept. prior to February 28, 2011. Copyright 2010, Society for Mining, Metallurgy, and Exploration, Inc.

IntroductionPelletization is a method of agglomeration used

to turn fine fractions of iron ore into an adequate product (pellet) to be fed to blast furnaces and direct reduction reactors, where it will be reduced to pig iron or sponge iron.

Binder must be added to the pelletization mixture to raise the liquid phase viscosity inside the capillary spaces, to maintain the cohesion of “green” pellets and to improve the compressive strength of the thermally treated pellets, by contributing to the formation of ceramic or iron ore oxide bridges or to scorify discrete points during thermal treatment.

The classic binder for iron ore pelletization is bentonite, which is characterized by its high swell-ing power, great superficial area, cation exchange capacities around 60 to 170 meq/100g and property of thixotropy. In contact with water, the bentonite platelets separate, forming a colloidal gel.

Bentonite enhances the strength of the iron ore pellets by the following mechanism: the presence of colloidal material shortens the distance between particles, raising the intensity of the Van der Waals forces. The disposition of bentonite plates in the point of contact between ore particles also enhances the pellet’s compressive strength. The typical dosage of bentonite as a binder for iron ore pelletization is 0.5 to 0.7% of the ore mass (dry basis). A disadvantage of bentonite is that it adds the undesirable contaminants alumina and silica to the pellet.

Binders that would not leave a residue after ther-

mal treatment would be extremely interesting. They would be advantageous for minimizing the variations in quality of the pellets, lowering the dosage to be used to around 0.05%.

Materials and methodsSample ore and binders. A sample of magnetite concentrate

from the Empire Mine (Palmer, MI) and received by Michi-gan Tech on 16/03/2009 was utilized. The average chemical composition of the sample is shown in Table 1.

The work was developed using the following binders: benton-ite, sodium hydroxide, sodium carbonate, sodium metasilicate, sodium tripolyphosphate and carboxymethyl cellulose. The typical chemical composition of bentonite is shown in Table 2 and the specification of the binders in Table 3.

The choice of binders to be tested was based on the author’s previous experience using these binders in the agglomeration of hematite concentrate (Moraes and Cassola, 2008).

The sodium hydroxide and sodium silicate were used in the form of a 50% (w/w) solution in water.

The moisture of bentonite was determined by drying it in an oven at a temperature of 105° C until constant weight. The moisture content of bentonite was 13% (dry basis).

Homogenization and reduction of sample mass. The sample of iron ore concentrate was received in three contain-ers containing approximately 15 kg (wet weight) each. It was dried at a temperature of 100°C for 24 hours.

The dry sample was disaggregated using a 28-mesh (0.71 mm) sieve, blended and split by a rotatory splitter. 24 sub-samples of approximately 1.7 kg each were produced. Figure 1 shows the process of homogenization and reduction of mass

Key words: Pelletization, Magnetite, Iron/ iron ores, Binders, Bentonite


of the sample.

Pelletization tests. For each test the ore, binder and moisture were added to a laboratory mixer (Fig. 2a) for two minutes. The tests were carried out using a laboratory-scale balling drum (Fig. 2b) with aliquots of 1.7 kg (dry basis) ore.

The balling procedure consisted of:

1) Delump mixed binder and concentrate through an eight-mesh screen;

2) Add a small amount of ore mixture to drum rotating at 25 rpm; moisten it with a water spray to produce pellet “seeds;”

3) Enlarge seeds by adding more concentrate while spray-ing with water, periodically removing and screening to keep uniform size. Pellets are finished when they are between 1.6 and 12.7 mm (0.06 and 0.5 in.) in diameter.

The composition of each mixture is given in percentage by mass concentration based on dry components. Moisture content and binder dosages are shown in Table 4.

Characterization of the pellet properties. Two standard

tests are used to measure the strength of pellets, whether the pellets are green pellets or fired pellets. These tests are the “drop” test and the “compression” test.

The drop test requires dropping a random sample of pel-lets from a height of 46 cm (18 in.) onto a steel plate. Pellets are dropped repeatedly until the pellets crack. The number of drops needed to crack each pellet is recorded and averaged.

Compression strength is measured by compressing or apply-ing pressure to a random sampling of pellets until the pellets crumble. The pounds of force required to crush the pellets is recorded and averaged.

These two tests are used to measure the strength of both wet and fired pellets. The drop and compressive test measurements are important because pellets, proceeding through the balling drum and subsequent conveyor belts, experience frequent drops as well as compressive forces from the weight of other pellets

Table 1 – Iron ore concentrate chemical composition

average.

Compound Value Unit

Fe 58.59 %

SiO2 4.59 %

CaO 6.46 %

MgO 1.83 %

Mn 0.04 %

Al2O3 0.25 %

P 0.014 %

CaO/SiO2 1.41 None

Total oxides 96.98 %

Table 3 – Binder specifications.

Product FormulaProperties

Grade Assay Form Impurites

Sodium hydroxide NaOH Reagent 98.8% White pellets ≤0.001%

N; ≤0.02%

NH4OH;

≤1.0%Na2CO3

Sodium carbonate Na2CO3 Reagent

anhydrous

≥99.5% Granular ≤0.001% N;

≤0.005%

silica; ≤0.01%

NH4OH;

≤0.01% in-

solubles

Sodium metasilicate pen-

tahydrate

Na2SiO3.5H2O Pure 95.7% (Titration) White small

beads

-----

Sodium tripolyphosphate

(TPP)

Na5P3O10 Technical

grade

85% White granu-

lar powder

-----

Sodium carboxymethyl

cellulose (CMC)

Average Mw~

90,000

Viscosity (C=4%,

H2O at 25 °C) =

115 cp

Off-white

powder

-----

Table 2 – Typical chemical analysis of the bentonite.

Analyte Concentration, %

SiO2 35.6

Al2O3 23.2

Fe2O3 5.5

CaO 9.6

MgO 2.1

Na2O 1.0

K2O 0.4

TiO2 1.3

MnO2 0.0

LOI 13.2

LOI: Loss of ignition.


of “fines” or coarse particles in the pelletized mineral ore. A binder which increases the pores formed in a pellet improves the pellet’s ability to resist thermal shock.

The pellets obtained were subjected to the tests for charac-terization described as follows.

Determination of moisture. After the pelletizing, about 300 g pellets were randomly sampled and dried in an oven at temperature of 105ºC until constant weight.

Drop number green pellet. From each batch of green pel-lets obtained, 20 pellets were randomly selected and dropped freely from a height of 46 cm (18 in.) onto a platform with a steel shank. Each pellet was dropped repeatedly until the first fractures appeared. The number of drops incurred by each pellet was recorded. The result is the average of values given by number of falls per pellet.

Wet compression strength. From each batch of pellets obtained, 20 pellets were randomly subjected to automatic compression until breaking. At this time the load was recorded and the arithmetic mean of the values represents the test result in kgf or N/pellet.

Dry compression strength. This test is similar to that done with green pellets, differentiated an additional step: the 20 pel-lets were dried in an oven at 105° C for 24 hours. After cooling for 15 minutes, the pellets were subjected to a compression test and the result was expressed in kgf or N/pellet.

Thermal shock resistance test. From the batch of green pellets, groups of 15 pellets were selected at random. Then

Figure 2 – Pelletization tests: (2a) Laboratory mixer (2b) Laboratory-scale balling drum.

Table 4 – Pelletization experiment results.

Binder dosage Moisture%

Wet drop

(18 in)±P95

Wetcrush

(N/pel-let) ±P95

Drycrush

(N/pel-let) ±P95

Crush after thermal shock

(N/pellet) ±P95

300°C 500°C 700°C 900°C

Bentonite (0.66%) 9.5 3.2±0.2 19.0±1.0 38.3±2.0 19.9±3.3 43.0±12.8 101.9±17.3 119.9±22.8

Bentonite

(0.40%)+NaOH(0.02%)

9.0 2.9±0.2 18.1±0.9 20.7±0.7 16.8±1.0 40.3±8.3 58.1±31.2 141.78±5.6

CMC (0.04%) 8.9 2.4±0.2 16.9±0.6 15.9±0.7 10.9±1.0 19.0±5.6 67.7±20.3 51.2±3.8

CMC (0.04%)+NaOH

(0.02%)

9.6 2.4±0.2 14.8±0.8 19.1±1.0 10.7±1.9 22.9±5.9 123.4±22.4 54.3±4.5

CMC (0.04%)+TPP

(0.02%)

9.1 2.9±0.2 19.6±0.7 21.5±1.3 14.7±2.4 26.7±8.5 89.8±15.1 89.3±11.5

Sodium carbonate

(1.5%)

8.5 2.8±0.2 19.7±1.4 44.3±2.4 33.2±10.7 36.5±8.8 167.6±30.5 180.8±21.9

Sodium silicate (1.5%) 9.0 1.9±0.1 11.2±0.6 52.6±4.6 36.6±7.8 74.7±6.5 120.2±16.8 151.4±21.2

Sodium carbonate

(0.75%)+ sodium

silicate (0.75%)

9.8 3.0±0.2 17.5±0.7 51.9±4.9 24.2±3.9 32.2±4.9 99.0±18.3 108.0±12.0

Figure 1 – Sample preparation.

traveling on top of them.Thermal shock resistance is a factor which must be taken

into consideration in any process for agglomerating mineral ore concentrate. Increases in a pellet’s thermal shock resistance improve the pellet’s ability to resist internal pressures created by the sudden evaporation of water when the pellet is heated in a kiln. If the pellet has numerous pores through which the water vapor can escape, thermal shock resistance is improved. If the surface of the pellet is smooth and continuous, without pores, the pellet has an increased tendency to shatter upon rapid heating. This causes a concurrent increase in the amount

a b


each group was placed in a muffle furnace at temperatures of 300°C, 500ºC, 700°C and 900°C for 10 minutes at each temperature. After this time, the groups were left to cool in the air and subjected to a test of compressive strength. This step also evaluates the percentage of pellets that crack and/or explode.

Results and discussionTable 4 shows the results obtained in the pelletization tests

with iron ore concentrate and different binders. For each value reported, the mean and standard deviation were determined for 20 pellets. The error bars shown on the graphs represent the 95% confidence intervals calculated using the t-distribuition.

The variation of the moisture content in each test was from 8.5% to 9.8%. The biggest moisture value was observed in the test CMC + NaOH and the smallest in the test with sodium carbonate. The wet drop results are shown in Fig. 3.

The highest value observed was that of the tests bonded with bentonite, which survived 3.2 drops. This value is not the minimum industrially acceptable, but in this case will be the value of reference. The worst result was observed for pel-lets bonded with sodium silicate, which only lasted1.9 drops. Pellets bonded with sodium carbonate, bentonite+NaOH and CMC (sodium carbomethyl cellulose) +TPP (sodium tripoly-

phosphate) show the same result of 2.9 drops each.

Effect of bentonite. Figure 4 shows the compressive strength of pellets bonded with bentonite. Two tests were carried out with ben-tonite at the dose of 0.66%. They showed wet and dry compressive strength results of 19 N/pellet and 38.3 N/pellet, respectively, higher than the minimum industrially acceptable (9 N/pellet, wet and 22 N/pellet, dry). Pellets bonded with bentonite at 0.4% and NaOH at 0.02% showed less wet and dry compressive strength compared to plain bentonite at 0.66%. However, the result of the wet compression (18.1 N/pellet) is still greater than

the minimum industrially acceptable standard. The dry com-pressive strength result (20.7 N/pellet) is almost the minimum industrially acceptable that is 22 N/pellet. These results suggest that the dose of bentonite and NaOH can be adjusted to reach the target strength.

The behavior of pellets bonded with bentonite (0.66%) and bentonite (0.4%)+NaOH (0.02%) during thermal shock can be seen in Fig. 4. As shown in this figure there are no significant differences between the results at 300°C and 500°C. At 700°C the performance of pellets with bentonite (0.4%)+NaOH (0.02%) falls and at 900°C the compressive strength improves again. These results show that is possible to reduce the amount of bentonite without reducing the quality of the pellets.

Pellets bonded with bentonite presented cracks in all tem-peratures during thermal shock resistance tests (Fig. 5). The cracks increase the formation of dust. There were no explosions during thermal shock resistance tests in pellets bonded with bentonite and bentonite + NaOH.

Effect of CMC. The use of CMC in iron ore pelletization is not new. Studies carried out by Lima (1991) and Lima and Chaves (1992) with hematite concentrate showed the results using CMC and CMC + TPP as binder. The authors found that CMC alone as a binder is not effective, because the pellets do not

resist heat. CMC and the additive TPP present the greatest potential for use as binders. However, it should be noted that TPP adds phosphorus to the pellets, which is an undesirable contaminant.

In 1998, Cassola and Chaves found that the dispersion of limo-nite from the surface of particles of hematite due to the dispersant effect of TPP was the fundamental mechanism for the process.

The aim of using CMC and CMC+NaOH and CMC+TPP in this work was to check if the same dispersant effect happens when pellets are made using a magnetite concentrate.

Figure 6 shows the compressive

Figure 3 – Strength of magnetite concentrate pellets bonded with different binders.

Figure 4 – Compressive strength of magnetite concentrate pellets bonded with bentonite.


strength of magnetite concentrate pellets bonded with CMC and the additives NaOH and TPP.

Magnetite pellets made using 0.04% CMC alone as a binder had insufficient dry compressive strength. The behavior of the pellets during the thermal shock resistance test did not provide enough compressive strength to survive the drying and firing steps.

When NaOH was added together with CMC to the magnetite concentrate pellets, the dry compressive strength increased, but was still insufficient. Similar results were found with the thermal shock resistance test.

Pellets bonded with CMC and TPP showed better results than pellets bonded with CMC and NaOH. The addition of 0.02% TPP improved the pellet dry compressive strengths to the minimum industrially acceptable (22 N/pellet). Similar results were found with the thermal shock resistance test.

These results suggest that dispersants are beneficial in mak-ing pellets with magnetite concentrate.

There were no cracks or explosions during thermal shock resistance during all tests.

Effect of sodium carbonate and sodium silicate. The compressive strength results for pellets bonded with sodium carbonate and sodium silicate are shown in Fig. 7.

The choice of these binders was made primarily to check the effect of the binder in the firing step. Specifically, it was determined whether it would be possible to decrease the tem-perature of sintering.

Pellets using 1.5% sodium carbonate as a binder showed overall better results than pellets bonded with 0.66% bentonite. However, 27% of these pellets exploded during the thermal shock resistance test at each temperature (300°C, 500°C and 900°C). The explosions may have occurred because the sodium carbonate should be acting as a dispersant over the fines par-ticles contained in the magnetite concentrate. This dispersion is very intense and saturates the pores of the pellet with water and solids, preventing the passage of gases of the evaporating water during the drying step. However, the compressive strength of pellets bound with sodium carbonate is higher at 700°C and 900°C than that of pellets bonded with bentonite (0.66%).

Pellets made with sodium carbonate as a binder can be seen after explosion during the thermal shock resistance test (Fig. 8).

The best results obtained were those of the pellets made with 1.5% sodium silicate as a binder. The addition of 1.5% sodium silicate improved the compressive strength, to well above the minimum industrially acceptable values. This result demonstrates that the dose of sodium silicate can be reduced.

The behavior of sodium-silicate-bonded pellets during the thermal shock experiments was excellent. There was no forma-tion of cracks and dust and no explosion occurred.

The compressive strength of the pellets made with sodium silicate as a binder was higher at 700°C and 900°C than that of the pellets bonded with bentonite (0.66%).

The pellets bonded with sodium carbonate (0.75%) and sodium silicate (0.75%) were made in a effort to eliminate explosions during thermal shock resistance test. This test showed that by reducing the dosage and using a combination of sodium carbonate and sodium silicate, it is possible to get the same results as that of the pellets bonded with bentonite (0.66%). In this case there were no cracks or explosions during

Figure 5 – Cracks in a pellet bonded with bentonite after the thermal shock resistance test (900°C).

Figure 6 – Compressive strength of pellets bonded with CMC and the additives NaOH and tripolyphosphate (TPP).


the thermal shock resistance test.

Conclusion• The results of magnetite concentrate pellets bonded

with bentonite (0.4%) + NaOH (0.02%) compared with pel-lets bonded with bentonite (0.66%) showed that is possible to reduce the bentonite dosage without reducing the quality of the pellets.

• Pellets bonded with bentonite presented cracks at all temperatures during thermal shock resistance tests. Cracks increase the formation of dust. There were no explosions dur-ing the thermal shock resistance tests of pellets bonded with bentonite and bentonite + NaOH.

• Magnetite pellets made using 0.04% CMC alone as a binder had insufficient dry compressive strength. The thermal shock resistance test gave insufficient compressive strength to survive the drying and firing steps.

• When NaOH was added together with CMC, the dry compressive strength increased, but not enough. Similar results were obtained in the thermal shock resistance test.

• The addition of 0.02% TPP improved the pellet dry

compressive strength to the minimum industrially acceptable (22 N/pellet). Similar results were obtained in the thermal shock resistance test. These results suggest that dispersants are beneficial in making pellets with magnetite concentrate.

• Pellets made with 1.5% sodium carbonate showed better results than pellets bonded with 0.66% bentonite. However, 27% pellets exploded during thermal shock resistance test at 300°C, 500°C and 900°C.

• The best results were achieved with pellets made with 1.5% sodium silicate as a binder. The addition of 1.5% sodium silicate improved the compressive strength to over the minimum industrially acceptable value. This result demonstrates that the dosage of sodium silicate can be reduced.

• The compressive strength at 700°C and 900°C of pellets made with sodium carbonate and sodium silicate is higher than that of the pellets bonded with bentonite (0.66%). These preliminary results suggest that is possible to optimize the sintering temperature of the pellets by adjusting the amounts and types of binders.

• The pellets bonded using sodium carbonate (0.75%) and sodium silicate (0.75%) showed that reducing the dose and us-ing the combination of sodium carbonate and sodium silicate, it is possible to get the same results as that of pellets bonded with bentonite (0.66%). In this case there were no cracks or explosions during the thermal shock resistance test.

ReferencesLima, J. R. B., 1991, Estudo da carboxi-metil-celulose como aglomerante para

pelotização, PhD dissertation, Escola Politécnica, Universidade São Paulo, São Paulo, 145 pp.

Cassola, M. S., and Chaves, A. P., 1998, “Effect of the addition of organic bind-ers on the behavior of iron ore pellets,” KONA: Powder and Particle, Osaka, No. 16, pp. 136-142.

Lima, J. R. B., and Chaves, A. P., 1992, “Estudo da carboxi-metil-celulose como aglomerante para pelotização,” São Paulo: EPUSP, 1992. (Boletim Técnico da Escola Politécnica da USP. Departamento de Engenharia de Minas, BT/PMI/014).

Moraes, S. L., 2004, Comparação de desempenho de aglomerante orgânico em relação à bentonita na operação de pelotização de concentrados de minério de ferro brasileiros de diversas procedências, 2004, Master’s thesis, Escola Politécnica, Universidade São Paulo, São Paulo, 80 pp.

Moraes, S. L., and Cassola, M. S., 2008, “Microstructure of iron ore pellets – organic and inorganic binders,” 3rd International Meeting on Ironmaking and 2nd International Symposium on Iron Ore. September 22-26th, 2008, Sao Luis, MA, Brazil, pp. 464-471.

Ripke, S. J., 2002, Advances in iron ore pelletization by understanding bonding and strengthening mechanisms, PhD Dissertation, Michigan Technological University, Michigan, 186 pp.

Figure 7 – Compressive strength of pellets bonded with sodium carbonate and sodium silicate.

Figure 8 – Pellets made with sodium carbonate as binder after exploding during the thermal shock test.


Ammonia separation of Ni from spent fly ash leach liquorS.-Y. Hwang Dept. of Metallurgical and Materials Engineering, University of Texas, El Paso, Texas

N.-S. KimKEN Research Center, Seokyeong University, Seoul, Republic of Korea

AbstractAn ongoing problem for industry in Korea is effective removal of heavy metals from and safe disposal of spent power plant fly ash. This research endeavor attempts to develop selective separation of Ni from leach liquor using a selective chemical precipitation method. The research was conducted by changing two variables, the pH and the concentration of ions in solution. The effect of other ions present in the system on the separation efficiency of Ni species was also investigated by adjusting OH and NH3 concentrations. In the pH range of 7-9, the chemical precipitation method produced a high yield of Ni by precipitation from the fly ash leach liquor.

Paper number MMP-09-054. Original manuscript submitted November 2009. Revised manuscript accepted for publication April 2010. Discussion of this peer-reviewed and approved paper is invited and must be submitted to SME Publications Dept. prior to February 28, 2011. Copyright 2010, Society for Mining, Metallurgy, and Exploration, Inc.

IntroductionIn South Korea, fly ash is an ever-increasing

industrial waste that is a functional byproduct of power plants accumulating 25,000 tons annually (MKE Technical Report, 2008). Fly ash was widely used in the past as an additional fuel in the cement industry (Korea Cements Industrial Association, 2007; Ssangyong Cement Industrial Co. Ltd., 2009). Currently, environmental regulations in South Korea prevents the use of fly ash as a fuel. For this reason, numerous studies have been conducted seeking to remove heavy metals from fly ash and to produce an environmentally safe material. These studies include the extraction of V and Ni (Abdel-latif, 2002; Tsai and Tsai, 1998; Khorfan et al., 2001; Vitolo et al., 2001; Vitolo et al., 2000; Murase et al., 1998; Stas et al., 2007) and the production of activated carbon from fly ash (Caramuscio et al., 2003; Uddin et al., 2007).

The purpose of the current research is to recover Ni from a fly ash leach liquor (FALL) in order to develop an economical approach to recover Ni from FALL. The Ni can then be reused as a valuable metal. This process may be economically viable in view of the recent price of Ni at $22,675 US$/ton (London Metal Exchange, March 1, 2010). A FALL that contained 2,300 ppm of Ni and 1,600 ppm of V along with trace amounts of Fe, Mn and Zn has been prepared for the study. The main focus of this study is to selectively separate Ni from the FALL by chemical precipitation.

Thermodynamic observationsThe distribution of Ni species in ammonia solution

as a function of pH was analyzed using thermodynamic data (see Tables 1 and 2). The activity of Ni in the NH3-H2O and H2O systems was calculated to determine the yield of Ni by precipitation in each system. The

activity of Ni(II) in the H2O system and the NH3-H2O system as a function of pH was analyzed to compare the empirical data of FALL with the theoretical data from the Eh-pH diagrams of the Ni-NH3-H2O system.

The precipitation of Ni(II) ion forms Ni(OH)2(s) and NiO(s) as the pH of the solution increases (Chang et al., 2007; Peacock and Sherman, 2007; Galvez et al., 2009). Figure 1 shows the activity of Ni(II) as a function of pH; this activity decreased with an increase in the pH of the solution. The activities were calculated using MathCAD software from the thermodynamic data in Table 1.

The concentration of NH3 in the NH3-H2O system was calculated using MathCAD and shows an increase in the concentration of NH3 at a pH above 9.3. Table 2 shows the Gibbs free energy of formation for various Ni(II) ammine complexes. The activity of Ni(II) and Ni ion complexes in the NH3-H2O system as a function of pH was analyzed with the thermodynamic data and the predominant species at various free ammonia concentrations were established in the Eh-pH diagrams of the Ni-NH3-H2O system (Meng and Han, 1996).

ExperimentalIn order to examine the surface morphology and particle

size distribution of the fly ash, a scanning electron micro-scope (SEM; JSM-6390) and a particle size analyzer (PSA; CILAS-1064) were used. Elemental analysis was also carried out using SEM-EDX.

In this experiment, a chemical precipitation method was used to selectively precipitate Ni(II) from the FALL. In order to selectively separate Ni(II), the FALL was diluted to a ratio of one part fly ash to ten parts water. The pH was controlled by adding either NH3 or NaOH per 100 ml of FALL, with an initial pH of 2.3-2.6. The pH was then increased from 3 to 11 to see the effect of pH on separating Ni from V in the FALL.

The precipitation reaction was continued for one hour and then the supernatant was separated from the precipitation

Key words: Chemical precipitation; Ni species; Selective separation; Fly ash


Figure 3 — Surface structure of fly ash.

compound using a centrifuge. The metal ion concentration was measured using an inductively coupled plasma (PerkinElmer-5300DV).

Results and discussionThe chemical composition of the dried fly ash was determined

by wet chemical analysis and a ThermoGravimetric Analysis (TGA) and is shown in Table 3. It is believed that the carbon in the dried fly ash started to oxidize at 400°C, which resulted in decreasing weight loss of 80% until 620°C (see Fig. 2).

As seen in Fig. 3, the surface morphology of fly ash was characterized using SEM (JSM-6390). There are many mi-

cro- and nano- sized pores on the surface of the particles. It is believed that these mesoscale pores are channels for the reactant and product ions mass transfer.

Fig. 4 shows the particle size distribution of the fly ash. The volumetric mean particle size (d50) was 65 micrometers. As seen in Fig. 4, fly ash particles were distributed from 1 to 200 micrometers.

Figure 1 — Activity of Ni(II) in Ni-H2O systems as a func-tion of pH at 25°C.

Table 1 — Thermodynamic properties (Stumm and Morgan,

1996).

Species ∆Go, J/mol

NH3 -26,570

NH4+ -79,370

Ni2+ -45,600

NiO -211,600

Figure 2 — TG analysis data of fly ash (Hwang et al., 2009).

Table 3 — Chemical composition of the tested fly ash (unit: wt.%)

C V Ni Fe Mn Zn Ash

80.4 1.9 0.7 0.4 0.01 0.02 16.5

Figure 4 — Particle size distribution of the fly ash.

Table 2 — Thermodynamic data of Ni(II) for various am-

mines (Meng and Han, 1996).

Species ∆Go, J/mol

Ni(NH3)2+ -87,814

-126,842

-162,981

-196,022

-22,647

-252,931


As seen in Fig. 1, the activity of Ni(II) in the Ni-H2O system continuously decreases with the increase in pH of the system. It is believed that the amount of Ni precipitation increased as the pH increased, which is expected in view of Fig. 1. Without any complexing agent in the FALL, the precipitation yield of Ni in the H2O system increased by increasing the pH, possibly due to a large quantity of NiO and Ni(OH)2 being formed as the OH concentration increased. Figs. 5 and 6 show most Ni, V, Mn, Fe and Zn ions were precipitated at a pH above 9-10 without a complexing agent when NaOH was used to adjust the pH. It is noted that some of these metal oxides are soluble at high pH and vanadium is soluble over the wide pH range of oxides. However, in the FALL, without complexing agents, vanadium oxide redissolved at a pH higher than 11.

In this study, for economic reasons, Ni and V separation from FALL were highlighted and less attention was focused on Fe, Mn and Zn separation from the FALL.

Figure 7 shows that most Fe was precipitated at a pH above 7 in the NH3-H2O system. It is noted that Mn and Zn are soluble in the wide range of pH 9-11 in the presence of ammonia as a

lixiviant. At pH 9, most Fe ions in the FALL precipitated out of the NH3-H2O system (see Fig. 7). However, at a pH above 9, Mn and Zn precipitates started to dissolve back into the FALL solution. It is reported that Mn and Zn are soluble in the range of pH 6-10 in the presence of NH3, but in the FALL the relative amount of Zn and Mn was too small in comparison to Ni, V and Fe (see Table 3). The concentration of free ammonia in the range of pH 3-9.3 was believed to be too little to make ammonia complexes in the FALL.

From an economic point of view, Ni can be selectively separated from FALL using the NH3 chemical precipitation method. It is confirmed that NH3 is more effective than NaOH at selectively separating Ni from other metals. This is due to a predominance of NH3 at the elevated pH. Likewise, NH4(I) is predominant at pH below 9.3, which explains why the Ni(II) precipitation rate decreased from 68% to 13% at pH 10 (Fig. 8). In this system, the precipitation of Ni increased to 68% at pH 9, while it decreased to 13% when the pH was increased above 10.

Chemical precipitation could be interpreted in view of the

Figure 6 — Precipitation yield of Fe, Mn and Zn in the H2O system.

Figure 5 — Precipitation yield of Ni(II) and V in the H2O system.

Figure 7 — Precipitation rate of Fe, Mn and Zn in the NH3-H2O system.

Figure 8 — Ni and V precipitation in the NH3-H2O system.


Eh-pH diagrams of the Ni-NH3-H2O system. Figure 9 shows the relationship between the NH3 species and the Ni activity, based on thermodynamic calculation and compared to empirical data. NH3 is predominant at a pH above 9.3, while NH4(I) is predominant at a pH below 9.3. Therefore, Ni was formed as ammine complexes with various ammines at a pH above 9.3. The Ni species in the NH3-H2O system were calculated by MathCAD software. Figure 9 shows Ni activity as a function of pH by changing concentration of NH3. As seen in Fig. 9, in the presence of NH3, the predominant species was Ni(NH3) 62+ (Deng and Chen, 2004). The activity of total Ni(II) complexes was increased with the increase of the concentration of free ammonia, a finding that agrees well with the empirical data at pH 9-10.

ConclusionsThe adjustment of the pH of the solution at about pH 10 us-

ing NH3 proved to be effective and economical in the selective separation of Ni from V in FALL. The Ni precipitation yield was increased by increasing the pH from 3 to 9; however, Ni returned to solution when the pH was increased above 9.3 in the presence of ammonia. The activity of Ni decreased with increasing pH in the range of pH 3-11, due to the precipitation of Ni by increasing pH without complexing agents in the H2O system. It has been shown that by using the NH3 chemical precipitation method, Ni separation from V can be 32-51% achieved in the FALL at pH 7-9.

AcknowledgmentsThis work was supported by the Energy Efficiency & Re-

sources of the Lorea Institute of Energy Technology Evaluation and Planning (2007-R-RU-11-P-14-3-010-2007) grant funded by the Korea government Ministry of Knowledge Economy.

ReferencesAbdel-latif, M.A., 2002, “Recovery of vanadium and nickel from petroleum

flyash,” Minerals Engineering, Vol. 15, pp. 953-961.

Caramuscio, P., Stefano, L.D., Seggiani, M., Vitolo, S., and Narducci, P., 2003, “Preparation of activated carbons from heavy-oil fly ashes,” Waste Manage-ment, Vol. 23, pp. 345-351.

Chang, P., Wang, X., Yu, S., Wu, W., 2007, “Sorption of Ni(II) on Na-rectorite from aqueous solution: Effect of pH, ionic strength and temperature,” Colloids and Surfaces A: Physicochem. Eng., Vol. 302, pp. 75-81.

Deng, X., and Chen, Z., 2004, “Preparation of nano-NiO by ammonia precipita-tion and reaction in solution and competitive balance,” Materials Letters, Vol. 58, pp. 276-280.

Galvez, J.L., Dufour, J., Negro, C., and Lopez-Mateos, F., 2009, “Routine to estimate composition of concentrated metal-nitric-hydrofluoric acid pickle liquors,” Hydrometallurgy, Vol. 96, pp. 88-94.

Hwang, S.Y., Kim, S.R., and Kim, N.S., 2009, “Recycling of power plant fly ash for construction materials by removal of heavy metals,” Korea Society of Waste Management, Vol. 26, No. 3, pp. 206-212.

Khorfan, S., Wahoud, A., and Reda, Y., 2001, “Recovery of vanadium pentoxide from spent catalyst used in the manufacture of sulfuric acid,” Periodica Polytechnica Ser. Chem. Eng., Vol. 45, No. 2, pp. 131-137.

Korea Cements Industrial Association, 2007, http://www.cement.or.kr/tech/resource.asp.

London Metal Exchange, March 19, 2010, http://www.lme.com/nickel.asp.Meng, X., and Han, K.N., 1996, “The principles and applications of ammonia

leaching of metals-A review,” Min. Proc. & Ext. Rev., Vol. 16, pp. 23-61. MKE Technical Report, 2008, “A study on fabrication of functional deodorized

panel and commercialization that use for carbonaceous circulate materials,” Ministry of Economy Knowledge, Republic of Korea.

Murase, K., Nishikawa, K., Ozaki, T., Machida, K., Adachi, G., and Suda, T., 1998, “Recovery of vanadium, nickel and magnesium from a fly ash of bitumenin-water emulsion by chlorination and chemical transport,” Journal of Alloys and Compounds, Vol. 264, pp.151-156.

Peacock, C.L, and Sherman, D.M., 2007, “Sorption of Ni by birnessite: Equilib-rium controls on Ni in seawater,” Chemical Geology, Vol. 238, pp. 94-106.

Ssangyong Cement Industrial Co. Ltd., 2009, http://www.ssangyongcement.co.kr/kor/new/environment/environment03_1.htm.

Stas, J., Dahdouh, A., and Al-chayah, O., 2007, “Recovery of vanadium, nickel and molybdemum from fly ash of heavy oil-fired electrical power stration,” Periodica Polytechnica Ser. Chem. Eng., Vol. 51, pp. 67-70.

Stumm, W., and Morgan, J.J., 1996, Aquatic Chemistry, third edition. John Wiley & Sons, Inc., New York, USA.

Tsai, L.S, and Tsai, M.S., 1998, “A study of the extraction of vanadium and nickel in oil-fired fly ash,” Resources, Conservation and Recycling, Vol. 22, pp.163-176.

Uddin, M.A, Shinozaki, Y., Furusawa, N., Yamada, T., Yamaji, Y., and Sasaoka, E., 2007, “Preparation of activated carbon from asphalt and heavy oil fly ash and coal fly ash by pyrolysis,” J. Anal. Appl. Pyrolysis, Vol. 78, pp. 337-342.

Vitolo, S., Seggiani, M., Filippi, S., and Brocchini, C., 2000, “Recovery of vanadium from heavy oil and orimulsion fly ashes,” Hydrometallurgy, Vol. 57, pp.141-149.

Vitolo, S., Seggiani, M., and Falaschi, F., 2001, “Recovery of vanadium from a previously burned heavy oil fly ash,” Hydrometallurgy, Vol. 62, pp.145-150.

Figure 9 — The Ni species in the NH3-H2O system as a function of pH by changing the concentration of NH3.


Flotation of zinc oxide minerals from low-grade tailings by oxine and dithizone using the Taguchi approachA. HajatiInstructor, Department of Mining, Iran University of Science and Technology, Arak, IranE-mail: [email protected]

A. Khodadadi and S. M. KoleiniProfessors, Department of Mining-Mineral Processing, Tarbiat Modares University, Tehran, Iran

AbstractThe main purpose of this study was to find the optimum values for factors affecting the Zn recovery rate using oxine or dithizone as a chelating agent collector during flotation in a Hallimond tube. Eight control factors, including four levels of pH, oxine/dithizone, corn starch, sodium thiosulfate, sodium citrate, sodium silicate, conditioning and floating durations were considered in a Taguchi experimental design technique. An L32 orthogonal array was applied to determine the signal-to-noise (S/N) ratio. Analysis of variance (ANOVA) was used to determine the optimum conditions and the most significant parameters affecting the reaction rate. Analysis of the experiments using the Taguchi approach indicated that pH had the highest contribution to the recovery rate of zinc oxide particles. The results showed that the recovery rate of zinc increased by 43.4 ± 3.0% and 24.9 ± 3.2% using oxine and dithizone, respectively, under optimum conditions with a confidence level of 90%.

Paper number MMP-09-002. Original manuscript submitted January 2009. Revised manuscript accepted for publication September 2009. Discussion of this peer-reviewed and approved paper is invited and must be submitted to SME Publica-tions Dept. prior to February 28, 2011. Copyright 2010, Society for Mining, Metallurgy, and Exploration, Inc.

IntroductionThere are two million tons of mine tailings depos-

ited around the zinc processing plant at the Goshfil Mine (BAMA Co.) in Esfahan, Iran, due to the lack of feasible techniques for the recovery of zinc and other valuable metals. The zinc residues are stockpiled until the recovery of valuable metals in the residues becomes economical. The stockpiled residues cause serious environmental problems, such as groundwater pollution due to the release of zinc and toxic elements (including such heavy metals as Pb, Cd, Cr, Hg and Co). Consequently, recovery of zinc oxide from tailings will help reduce environmental hazards and create a new source of raw minerals (Kerfoot et al., 1994; Freeman and Nightingale, 1980).

Different technologies, such as leaching, bioleach-ing and flotation, have been employed to extract metals (such as Pb and Zn) from oxide ores, depending on the type and concentration of minerals and also the type of impurities involved (Hoffmann et al., 1989; Haddadin et al., 1993; Gomez et al., 1999; Mulligan et al., 1999; Witne and Philips, 2001; Wang et al., 2005).

However, it is generally known that oxidized miner-als of zinc are more difficult to float than their sulfide

counterparts. This fact is due to the higher solubility of oxide minerals, as well as the extensive hydration of oxide surfaces (Fuerstenau et al., 1985; Önal and Abramov, 2002; Yianatos et al., 2005). Oxidized minerals are frequently treated with sulfidizing agents prior to flotation. When more sulfide ions are added, the system is depressed once more due to sulfide ion adsorption in preference to collector adsorption (Rastas et al., 1990; Önal et al., 2005; Rashchi et al., 2005)

Direct flotation of zinc oxide minerals by applying the ap-propriate collectors eases the above problems. However, few attempts have been made to float zinc oxide minerals using chelating agents (Marabini and Rinelli, 1973; Marabini et al., 1994; Barbari, 1998; Önal and Abramov, 2002; Pereira and Peres, 2005; Önal et al., 2005).

This paper reports an attempt to recover zinc oxide mineral from mining waste through flotation without surface sulfida-tion, using oxine and dithizone as chelating collectors. The aim was to evaluate the optimum experimental conditions that provide the highest flotation recoveries of zinc metal from a solid matrix, such as the ore wastes around the openpit lead and zinc mine at Goshfil in Esfahan, Iran.

The recovery conditions were optimized using a Taguchi orthogonal array L32 (48) design. The Taguchi approach has been shown to be an effective means for the improvement of

Key words: Tailing, Zinc oxide, Flotation, Oxine, Dithizone, pH


productivity at the research and development stage, so that high quality items can be produced quickly at low cost. It has been applied to a wide range of industrial fields worldwide (Opur et al., 2003; Moghaddam et al., 2006). The Taguchi approach was used in the present work to determine the optimum recovery conditions for maximizing the zinc content of the flotation concentrate using oxine and dithizone as chelating agent col-lectors without surface particle sulfidation.

ExperimentalSampling and treatment. Samples were obtained from

oxide dam tailings of the Goshfil mine of BAMA Co. (Esfahan, Iran). The samples were crushed by a jaw crusher and ground with a ball mill (Bond model) to obtain a homogenized final sample that included 5.3% zinc oxide (smithsonite and sphal-erite), 2.0% lead oxide (galena and cerussite) and 13.3% iron oxide (hematite and goethite) (Oliazadeh and Abdoulzadeh, 1997; Hajati et al., 2003). The results are shown in Table 1. In order to determine the effect of different particle size distribu-tions on the flotation recovery, 1 kg of the homogenized sample was sieved and divided into several limit sizes in different time durations of 0, 5, 10, 20, 30, 40, 50 and 60 minutes using 1000, 297, 149, 74, 53, 44 and 37-micron ASTM standard sieves. Eight different curves were plotted as mass cumulative pass-ing percentages versus particle size (microns). The grinding time with the least production of fine particles in the sample feed flotation and the lowest degree of liberation of zinc oxide (smithsonite) and iron oxide minerals was 10 minutes (Figs. 1 and 2). The samples were fixed and the surfaces of the particles were polished to determine a suitable size limit according to the degree of liberation of zinc oxide (smithsonite) and iron oxide minerals.

Microscopic tests to define degree of liberation. Mi-croscopic analysis enabled the counting of the number and percentage of liberated minerals. Fig. 2 shows the liberation degree of smithsonite and iron oxide minerals. Liberation of minerals can significantly affect the flotation process (Trahar, 1976; Weller et al., 1995; Hajati et al., 2003). At -95 microns, more than 80% smithsonite and 60% iron oxide minerals were liberated. At this size, the liberation of smithsonite was greater than that of iron oxide. The limit of -95 microns was deemed most suitable for further study.

Desliming and magnetic tests. The presence of fine par-ticles of iron oxide minerals is one of the main problems in the flotation process. A comparison of the liberation degree of smithsonite and iron oxide minerals was necessary to select a suitable size limit. By grinding the samples at -95 microns and desliming at 37 microns, -95, +37 microns the size limit was achieved in the feed flotation tests. In order to reduce the iron content, the samples were passed through a wet magnetic field with an intensity of 2 Tesla.

Micro flotation cell. For the flotation tests, a Hallimond cell was used. The Hallimond cell consists of a Pyrex tube with a fixed diameter of nearly 20 mm, a height of 300 mm and two

feeding parts, one for pH control and the other for injection of the materials. There is a sinter glass at the bottom of the tube for air bubble flow. The upper part of the tube is inclined 30 degrees to collect the floated minerals.

Reagents. The particles were suspended by a gently rotat-ing magnetic stirrer. Oxine and dithizone were used as che-lating agent collectors (Merck, Darmsdat, Germany). Oxine (C9H7NO) is a white crystal with the molecular weight of 145.16 g/mole (Fig. 3a) and dithizone (C6H5N or H2Dz) is a purplish black solid soluble in both chloroform (20 mg/ml) and CCl4 (0.5 mg/ml) (Fig. 3b). Since chloroform and CCl4 are insoluble in water, acetone was employed as a solvent. The oxine concentration was kept below 2%, because of the solubility of acetone in water, which may remove oxine from the system. Fuel oil (gasoline) was also added. Pine oil with a range of 120 g/ton was selected as frother. The frothing time was 12 minutes. The main parameters used for flotation tests were pH, conditioning time, floating time and concentration of oxine and/or dithizone, corn starch, sodium thiosulfate, sodium

Table 1 — Results of feed analysis by X-ray fluorescence and X-ray diffraction.

XRF:ZnO: 5.3 % Fe2O3: 13.3 % SiO2: 27.4 % Al2O3: 5.4 % SO3: 4.6 %

PbO: 2.0 % TiO2: 0.23 % CaO: 15.4 % MgO: 6.8 % L.O.I: 3.9 %

XRD: barite, dolomite, quartz, smithsonite, cerussite, goethite, hematite, sphalerite, galena

Figure 1 — The percent cumulative passed material as a function of particle size for eight grinding times (t in minutes).

Figure 2 — Degree of liberation of iron oxide and smith-sonite based on the size of the particles.


citrate and sodium silicate. Preliminary tests were carried out in order to evaluate the effectiveness of various concentrations of the key reagents. Four different concentrations were chosen as shown in Table 2. Temperature was kept at 20 °C during the experiments.

AAB analysis. Concentrations of Zn in the feed, concen-trate and tailing flotation tests were analyzed at BAMA Co. laboratories by an atomic absorption spectrophotometer (AAB-Perkin Elmer, Model 100, USA) using the standard addition technique. For metal analyses, ~0.2 g of the sample (dry wt.) was digested in a Teflon bomb for four hours at 140 °C with 2 ml concentrated nitric acid.

Orthogonal arrays and experimental parameters (Taguchi method). The technique of defining and investigating all the possible conditions in an experiment involving multiple factors is known as the design of experiments (Roy, 1990). Basically, classical parameter design, developed by Fisher, is complicated and not easy to use. A large number of experiments must be conducted when the number of process parameters increases. To solve this problem, the Taguchi method uses a special design of orthogonal arrays to study the entire parameter space with only a small number of experiments (Nian, 1999).

The Taguchi method was developed by Genichi Taguchi between 1950 and 1960 to improve the implementation of total quality control in Japan (Taguchi and Yokoyama, 1993). The goal is to find out the optimal and robust product or process characteristics that have a minimized sensitivity to noise. Taguchi design can determine the effect of flotation factors on characteristic properties and the optimal conditions of the factors. This method is a simple and systematic approach to optimize the design for performance, quality and cost (Roy, 1990; Park, 1996; Cox and Reid, 2000; Montgomery, 2001; Mousavi et al., 2007). The Taguchi approach uses orthogonal

arrays and analysis of variance (ANOVA) as the tools of analysis. ANOVA can estimate the effect of a fac-tor on the characteristic properties and the experiment can be performed with minimal replication using the orthogonal arrays. Conventional statistical experimen-tal designs can determine the optimal conditions on the basis of the measured values of the characteristic properties, while the Taguchi method can determine the experimental conditions having the least variability as the optimal conditions. The variability is expressed by the signal-to-noise (S/N) ratio. The terms ‘signal’ and ‘noise’ represent the desirable and undesirable values for the output characteristics, respectively. The Taguchi

method uses the S/N ratio to measure the quality characteristic deviating from the desired value. The experimental condition having the maximum S/N ratio is considered to be the opti-mal condition, as the variability characteristics are inversely proportional to the S/N ratio (Roy, 1990).

For the Taguchi design and subsequent analysis, the Qualitek-4 software (version 7.5+) was used. The appropriate orthogonal array for the experiment was determined by the software. A well-designed experiment can substantially reduce the number of experimental tests required. The Taguchi tech-nique applies fractional factorial experimental designs, called orthogonal arrays (OA), to reduce the number of experiments, meanwhile obtaining statistically meaningful results. Through observation and review of literature, eight control factors at four different levels were examined in this study (Table 2). These control factors included pH, oxine/dithizone (g/t), corn starch (g/t), sodium thiosulfate (g/t), sodium citrate (g/t), so-dium silicate (g/t), time conditioning (min) and time floating (min). All control factors have four levels. By obtaining the L32 orthogonal array of the mentioned eight parameters at four levels, the number of experiments required was reduced to 32 (Table 3). In other words, 32 experiments with differ-ent combinations of factors should be conducted in order to study the main effects and interactions, which in the classical combination method using full factorial experimentation means we will need 48=65,536 experiments to capture the influential parameters.

Results and discussionMineralogical and magnetic test results. The results of

mineralogical analysis through X-ray fluorescence demon-strated that nearly 80% of the zinc minerals in the Goshfil tailings are smithsonite (Table 1). The plot of accumulated weight vs. micron size for different grinding times is shown in Fig. 1. Increasing the grinding time causes the percentage of

Figure 3 — The chelating of zinc by oxine (a) and dithizone (b).

a. b.

Table 2 — Experimental parameters and the levels tested.

ParametersLevel

1 2 3 4

(A) pH 3 6 9 12

(B) Oxine/dithizone as a collector (g/t) 20 60 180 540

(C) Corn starch (g/t) 500 1000 1500 2000

(D) Sodium thiosulfate (g/t) 20 60 180 540

(E) Sodium citrate (g/t) 20 60 180 540

(F) Sodium silicate (g/t) 20 60 180 540

(G) Time conditioning (min) 5 10 15 30

(H) Time floating (min) 10 20 30 40


coarse particles converting to fine particles to increase. Also, reducing the slope of the curve or increasing the grinding time makes smithsonite, hematite and goethite more likely to be ground into fine particles. The results of counting the number of minerals and the percentage of liberated minerals can be found in our previous work (Hajati et al., 2003). However, at -95 microns, more than 80% of the smithsonite and 60% of the iron oxide minerals are liberated. By desliming and submitting

the samples to magnetic tension, under the best conditions 35% of the weight percentage of iron oxide is removed and the grade of Zn, Pb and Fe changes from 7.6, 3.5 and 12.2% to 7.8, 5.0 and 3.0%, respectively, in the part to be used as a flotation feed.

Experimental design and data analysis. Eight-factor, four-level factorial designs were used for evaluating the effects of the

Table 3 — Results of an orthogonal array L32 (48) test of Zn flotation recovery using oxine and dithizone.

Trial No.

Parameters and their levelRecovery (%)

Oxine Dithizone

A B C D E F G H R1 R2 AvgS/N ratio R1 R2 Avg

S/N ratio

1 1 1 1 1 1 1 1 1 4.3 4.7 4.0 12.7 5.0 5.3 4.7 13.9

2 2 2 2 2 2 2 2 1 5.5 6.9 4.2 14.3 8.3 7.6 9.0 18.3

3 3 3 3 3 3 3 3 1 7.5 9.5 5.6 16.9 12.2 11.0 13.3 21.6

4 4 4 4 4 4 4 4 1 14.9 14.6 15.3 23.5 9.3 6.9 11.7 18.7

5 4 3 3 2 2 1 1 2 10.0 9.3 10.8 20.0 8.1 3.7 12.5 15.0

6 3 4 4 1 1 2 2 2 30.5 28.9 32.1 29.7 9.5 5.0 14.0 17.3

7 2 1 1 4 4 3 3 2 12.1 11.1 13.2 21.6 9.1 6.5 11.6 18.4

8 1 2 2 3 3 4 4 2 26.3 19.9 32.7 27.9 25.5 15.5 35.5 26.6

9 3 2 1 4 3 2 1 3 22.7 24.9 20.4 27.0 21.6 15.0 28.2 25.8

10 4 1 2 3 4 1 2 3 33.9 29.6 38.3 30.5 14.8 11.5 18.1 23.0

11 1 4 3 2 1 4 3 3 16.1 20.3 11.9 23.5 7.3 4.8 9.7 16.1

12 2 3 4 1 2 3 4 3 24.2 22.5 25.9 27.6 17.7 15.1 20.2 24.8

13 2 4 3 3 4 2 1 4 7.8 7.5 8.2 17.9 3.5 4.5 2.5 10.1

14 1 3 4 4 3 1 2 4 16.1 18.6 13.6 23.9 3.4 4.4 2.3 9.6

15 4 2 1 1 2 4 3 4 14.1 12.3 16.0 22.9 5.7 4.8 6.5 14.9

16 3 1 2 2 1 3 4 4 13.1 12.0 14.2 22.3 7.5 10.8 4.3 15.7

17 2 3 2 4 1 4 1 1 7.9 8.3 7.4 17.9 6.4 6.3 6.6 16.1

18 1 4 1 3 2 3 2 1 5.0 6.7 3.3 12.9 11.8 8.9 14.8 20.9

19 4 1 4 2 3 2 3 1 7.7 7.6 7.9 17.8 6.5 5.7 7.3 16.1

20 3 2 3 1 4 1 4 1 26.9 25.4 28.4 28.6 8.5 9.6 7.3 18.4

21 3 1 4 3 2 4 1 2 11.6 9.8 13.3 21.0 2.8 3.0 2.7 9.0

22 4 2 3 4 1 3 2 2 9.6 10.3 8.8 19.6 4.5 4.6 4.3 13.0

23 1 3 2 1 4 2 3 2 10.9 10.8 11.0 20.7 13.4 10.5 16.3 22.1

24 2 4 1 2 3 1 4 2 20.8 20.4 21.2 26.3 8.2 7.4 8.9 18.1

25 4 4 2 1 3 3 1 3 23.6 22.3 25.0 27.4 12.3 12.7 11.9 21.8

26 3 3 1 2 4 4 2 3 22.8 22.5 23.0 27.1 11.1 12.4 9.9 20.1

27 2 2 4 3 1 1 3 3 32.4 25.7 39.2 29.8 8.7 9.3 8.2 18.8

28 1 1 3 4 2 2 4 3 11.6 13.3 9.8 21.1 5.4 6.3 4.4 14.3

29 1 2 4 2 4 3 1 4 14.2 12.1 16.3 22.8 2.6 3.1 2.1 7.9

30 2 1 3 1 3 4 2 4 16.2 13.0 19.3 23.8 5.2 4.6 5.9 14.2

31 3 4 2 4 2 1 3 4 9.7 10.4 9.0 19.7 4.7 5.5 3.9 13.2

32 4 3 1 3 1 2 4 4 22.6 18.4 26.9 26.8 5.3 5.2 5.5 14.5

Average recovery = 16.0 %

StDev = 8.5

Average recovery = 8.9%

StDev = 5.7

Rj: Recovery


following factors on zinc recovery: pH, oxine/dithizone (g/t), corn starch (g/t), sodium thiosulfate (g/t), sodium citrate (g/t), sodium silicate (g/t), time conditioning (min) and time floating (min). In order to estimate the best condition for the recovery of zinc metal from the oxide minerals, 32 experiments were performed. In this study, the focus was on the main effects of the eight most important factors on Zn recovery. The results of these experiments are given in Table 3. The mean values of each recovery yield for the corresponding factors and at each level are presented in Fig. 4. In all cases, each reported recovery is the average of four measurements, in each of which the parameter of interest was kept constant, while the other

parameters were changed (Table 3). The optimum values of the selected factors (pH, concentration of oxine/dithizone, corn starch, sodium thiosulfate, sodium citrate, sodium silicate, dura-tion of conditioning and duration of flotation) for Zn recovery using the oxine collector were 9, 540, 500, 540, 20, 540, 10, 30, respectively. For the dithizone collector they were 9, 540, 1500, 180, 180, 60, 10 and 30, respectively. The ANOVA results for the calculated models (Table 4) indicated that the levels of pH, oxine (g/t) and sodium thiosulfate (g/t) play the most important roles in Zn recovery using the oxine collector, while pH, dithizone (g/t) and sodium citrate (g/t) were most important when using the dithizone collector. The remaining

Figure 4 — Response graph illustrating the zinc recovery ratios plotted against various flotation parameters by oxine and dithizone chelating agents as a collector flotation.

Table 4 — ANOVA Results of Zn recovery by oxine and dithizone and S/N ratio analysis.

Parameters Degree of freedom (f)

Sum of squares (S)

Variance (V) F-ratio (F) Percentage contribution (%)

1. Oxine S/N Ratio

(A) pH 3 307.7 102.6 12.0 37.7

(B) Oxine/ dithizone (g/t) 3 99.6 33.2 3.9 9.9

(C) Corn starch (g/t) (3) 35.3 11.8 1.4 (pooled) 1.3

(D) Sodium thio-sulfate (g/t) 3 87.9 29.3 3.4 8.3

(E) Sodium citrate (g/t) (3) 31.6 10.5 1.2 (pooled) 0.8

(F) Sodium silicate (g/t) (3) 42.2 14.1 1.6 (pooled) 2.2

(G) Time conditioning (min) (3) 30.7 10.2 1.2 (pooled) 0.7

(H) Time floating (min) (3) 53.0 17.7 2.1 (pooled) 3.7

Error 7 59.8 8.5 35.4

Total 31 747.9 100%

2. Dithizone S/N Ratio

(A) pH 3 276.8 92.3 6.8 33.7

(B) Oxine/ dithizone (g/t) 3 64.5 21.5 1.6 3.4

(C) Corn starch (g/t) (3) 12.7 4.2 0.3 (pooled) 0

(D) Sodium thio-sulfate (g/t) (3) 58.9 19.6 1.4 (pooled) 2.6

(E) Sodium citrate (g/t) (3) 37.6 12.5 0.9 (pooled) 0

(F) Sodium silicate (g/t) 3 115.1 38.4 2.8 10.6

(G) Time conditioning (min) (3) 31.8 10.6 0.7 (pooled) 0

(H) Time floating (min) (3) 6.9 2.3 0.2 (pooled) 0

Error 7 95.1 13.6 49.6

Total 31 699.5 100%

floating

ery


parameters, including corn starch, sodium citrate, sodium silicate, time conditioning and float time using oxine and corn starch, sodium thiosulfate, sodium citrate, time conditioning and float time using dithizone had the least significant effects on zinc recovery from oxide matrix minerals. These factors are listed in Table 4. Further experiments were performed under the proposed conditions and the results obtained from the L32 (48) matrix (Table 5) showed that under the optimized condi-tions, the recoveries were similar to the optimum performance calculated using the following expression:

(1)

Where T is the grand total of all results, N is the total number of results, Yopt is the performance under optimum conditions and are the average performances of the pH and concentration of collectors at optimum levels, respectively. Accordingly, the performance was estimated using only the significant fac-tors for oxine and dithizone (Roy, 1990). Table 4 shows that there are satisfactory agreements between the results for zinc recovery from the solid matrices estimated based on Eq. (1) and those obtained under optimum conditions. In the optimum conditions, the confidence interval (C.I.) of performance was calculated using the following expression:

(2)

Where, F (1, n2) is the F value from the F table at a required confidence level at a degree of freedom (DOF) 1 and DOF of error, n2; Ve is the variance of error term (from ANOVA) and Ne is the effective number of replications (Yamini et al., 2008). Table 5 shows that the obtained recovery for Zn under optimum conditions using oxine (43.4%) is larger than that obtained using dithizone (24.9%). Under this condithon, oxine is more effective than dithizone in achieving zinc oxide recovery. The confidence intervals of zinc recoveries by oxine and dithizone are ± 3.0 and ± 3.2, respectively (Table 5).

Analysis of variance. According to the ANOVA results, the initial pH had the largest variance for oxine and dithizone. The initial oxine and sodium thiosulfate concentrations for the

oxine collector and sodium silicate and dithizone concentrations took the second and third places in the recovery flotation tests (Table 4). Therefore, it can be concluded that pH was the most influential factor. On the other hand, the degree of freedom (DOF) for each factor was 3 and the total DOF was 31, so the DOF for error was 7. Finally the variance for the error Ve, obtained by calculating the error sum of squares and dividing by the error degrees of freedom, was calculated. Therefore, it was possible to calculate the F-ratio, defined as the variance of each factor divided by Ve. In order to calculate the DOF from the error, a pooled ANOVA was applied. The percentage con-tribution of each factor to Zn recovery, which is calculated by the ratio of the variance for each factor to the total variance, is shown in Table 4. The percentage contribution of the initial pH by oxine flotation was the greatest, 37.7, as compared to those of oxine and sodium thiosulfate concentrations, which were 9.9 and 8.3%, respectively. Further, the percentage contribution of the initial pH by dithizone flotation was the greatest, 33.7 %, compared to those of dithizone and sodium silicate concentra-tions, which were 3.4 and 10.6%, respectively.

S/N ratio analysis. The S/N ratio analysis was used to compute the signal-to-noise ratio for each level of the process parameters. Regardless of the category of the quality (the-lower-the-better, the-higher-the better and the-more-nominal-the-better), a larger S/N ratio corresponds to better quality characteristics (Yoon et al., 2002). Since the current study took the percentage recovery of zinc (w/w) as the quality characteris-tic, the higher-the-better criterion was applied when evaluating the S/N ratios of various extraction parameters. The impact of all parameters was calculated from the results of ANOVA flotation tests after conducting the S/N ratio calculations (Fig. 4; Tables 3 and 4). Estimations of all performance parameters under optimum conditions using raw data and the results of S/N ratio analysis of zinc recovery are shown in Table 5. The most important parameters affecting the ratio of Zn flotation recovery were analysed as follows:

Effect of pH. The floatability of particles in Zn oxide minerals using oxine and dithizone depends on the pH level, since both of the chelating agent collectors are controlled by this factor. As can be inferred from the results, the pH of zinc oxide flotation plays an important role in the floating of these particles from the solid matrix samples. This means that Zn

Table 5 — Estimation of Zn recovery with two chelating agents under optimum conditions.

ParametersOxine Dithizone

Optimized level Level Optimized level Level

(A) pH 9 3 9 3

(B) Oxine/ dithizone (g/t) 540 3 540 3

(C) Corn starch (g/t) 500 1 1500 3

(D) Sodium thiosulfate (g/t) 540 4 180 3

(E) Sodium citrate (g/t) 20 1 180 3

(F) Sodium silicate (g/t) 540 4 60 2

(G) Time conditioning (min) 10 2 10 2

(H) Time floating (min) 30 3 30 3

Current grand average of performance (%) 16.0 8.9

Expected results under optimum conditions (%) 43.4 24.9

Confidence level: 90% 90%

Expected results: 43.4 ± 3.0 24.9 ± 3.2


flotation recovery is usually enhanced as the pH rises from 3 to 9; above pH 9, Zn recovery decreases (Fig. 4). An increase in pH causes an increase in the concentration of electrical charge on the surface of the particles and an increase in collector se-lectivity to higher recovery. Thus, pH affects the flotation of zinc oxide minerals for both oxine and dithizone. As shown in Fig. 4, the optimized pH for Zn recovery is 9.

Effect of oxine/dithizone concentration. The effect of oxine/dithizone concentration on the recovery of Zn (oxide phase) was investigated by the addition of various concentrations of oxine/dithizone in acetone solvent to the Hallimond flotation cell and the results are shown in Fig. 4. As can be seen, when the concentration of oxine/dithizone increases, Zn recovery increases. Spread over all concentrations of the chelating agents, the recovery of Zn for oxine is better than that for dithizone. At a low concentration of either chelating agent, the recovery of Zn with oxine fluctuates widely, but increases steadily with the use of dithizone. At higher concentrations of the chelating agents, Zn recovery using both collectors is similar. This result means that the Zn-oxine interaction is stronger than that of Zn-dithizone. Fig. 4 also shows that the optimum concentration of agents for the maximum Zn recovery is more than 540 g/t.

Confirmation and comparison of Zn recovery by oxine and dithizone. The confirmation test is a crucial and highly

recommended step to verify the experimental conclusions. In fact, running the confirmation test is necessary to show the optimum conditions and compare the results with those of the expected performance. If the new design does not meet the specified requirements, the process must be reiterated using new systems until the criteria are met. In this study, increasing the oxine/dithizone concentration to more than 540 g/t caused the best recovery. Several confirmation experiments were car-ried out at the optimum levels of flotation parameters for zinc recovery (Table 6). Figure 5 compares the grade and recovery of zinc oxide flotation recovery by variation in oxine and dithi-zon concentrations. As shown, the grade and recovery of zinc increase with the increase of both collectors up to an optimal condition around 1500 g/t at a pH level of 9, after which they decrease. However, the dithizone curves tend to decrease com-pared to those of oxine. This result could be due to the lower solubility of the collector. Similar results have been found by Natarajan and Nirdosh, 2006. The final step is to predict and verify the improvement of performance characteristics. The confirmation test indicated that the zinc recovery rates were maximized by increasing oxine and dithizone concentrations to 1000-2000 g/t and 1250-1500 g/t, respectively. In this con-dition, recoveries and grades of zinc were 38%, 53.2% and 12.6%,11.2% respectively. These results indicate that zinc recovery using dithizone is higher compared to recovery using oxine, while its grade is lower. Flotation using dithizone also

Table 6 — The effect of oxine and dithizone concentration on the grade and recovery of Zn under optimum conditions.

Concentration (g/t)Oxine Dithizone

Grade (%) Recovery (%) Grade (%) Recovery (%)

20 7.3 16.6 6.7 7.7

60 7.86 15.6 6.9 7.1

180 7.97 15.9 6.7 8.6

540 8.86 21.1 7.4 10.6

750 10 22.9 10 32.4

1000 11.5 33.7 11 36.7

1250 12.6 37.8 11.2 55.4

1500 12 37.6 10.5 53.2

1750 9.5 38 8.65 47.1

2000 10 36 8.5 24

2500 9.75 29.9 8.5 30.7

3000 9.5 28.5 9.1 23.9

Figure 5 — Comparison of the recovery and grade of zinc oxide by varying oxine and dithizone concentration under optimum conditions (pH 9).

Figure 6 — The relation between Zn recovery and grade using oxine and dithizone under optimum flotation condi-tions (pH 9).


seems to produce a higher error rate. The experimental results confirmed the validity of the applied technique (Taguchi ap-proach) for optimizing Zn recovery parameters, showing that it is possible to increase the Zn recovery rate significantly. The relation between Zn flotation recovery and its grade was inter-plotted using oxine and dithizone results (Table 6 and Fig. 6). The relationship between Zn grade and recovery is direct. As Fig. 6 shows, to attain the highest zinc recovery, dithizone is better than oxine, but to attain the highest zinc grade, oxine is better than dithizone.To achieve 100% recovery, the maximum zinc grade will be 20.1% for oxine and 12.9% for dithizone.

ConclusionsThis study attempted to optimize Zn flotation recovery from

low-grade tailings. Using the Taguchi method of experimental design and analysis of the S/N ratio, the relative contributions of different parameters for maximum Zn recovery were inves-tigated using oxine and dithizone. The maximum Zn recovery rate for oxine was obtained at a pH of 9, an oxine concentration of 1000-2000 g/t, a sodium thiosulfate concentration of 540g/t, a flotation time of 30 min, a conditioning time of 10 min and the following concentrations of reagents: sodium silicate, 540g/t, corn starch, 500g/t and sodium citrate, 20g/t. The maximum Zn recovery rate for dithizone was obtained at a pH of 9, a dithizone concentration of 1250-1500 g/t, a sodium silicate concentration of 540g/t and a sodium thiosulfate concentra-tion of 180g/t. The results also showed that the recovery rate of zinc increased to 43.4 ± 3.0% using oxine and to 24.9 ± 3.2 using dithizone at optimum conditions with a confidence level of 90%.

AcknowledgmentsThe authors acknowledge the Department of Mining-

Mineral Processing, Tarbiat Modares University and BAMA Co. (Esfahan, Iran) for providing the necessary facilities for carrying out this work.

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Design, operation and control of a teeter-bed hydroseparator for classificationJ.N. Kohmuench, E.S. Yan and M.J. MankosaManager, R&D process group, senior research scientist and vice president of operations, respectively, Eriez Manufacturing, Erie, Pennsylvania

G.H. Luttrell and R.C. BrattonProfessor and senior research associate, Dept. of Mining and Minerals Engineering, Virginia Tech, Blacksburg, Virginia

AbstractTeeter-bed (i.e., hindered-bed) separators are used throughout the mineral processing industry for both classifica-tion and density separation. The high capacity and sizing characteristics of these units make them ideal for feed preparation prior to coarse flotation (+0.100 mm) circuits. Teeter-bed separators are typically easy to control with two basic operating parameters, including fluidization water rate and bed level. However, data show that these two parameters greatly interact with one another. Given this finding, a high-level automatic control scheme was developed and implemented on a full-scale separator. In addition, dialogue with plant operators has led to sev-eral simple changes to an already established design that improves operational and maintenance characteristics.

Paper number MMP-09-040. Original manuscript submitted October 2009. Revised manuscript accepted for publication December 2009. Discussion of this peer-reviewed and approved paper is invited and must be submitted to SME Publications Dept. prior to February 28, 2011. Copyright 2010, Society for Mining, Metallurgy, and Exploration, Inc.

IntroductionHydraulic separators are frequently used in mineral

processing for sizing applications. Of these devices, the teeter-bed or hindered-bed separator (TBS) con-tinues to be a mainstay within the mining industry. Typically, these devices are open-top vessels through which elutriation water is injected. This teeter water is commonly distributed across the base of the cell through a network of distribution pipes or perforated plating.

During operation, feed solids are introduced into the upper section of the separator and are influenced by the teeter water. While the smallest/lightest par-ticles are hydraulically carried over into the collection launder, the coarser/denser particles have sufficient mass to settle against the fluidization water and cre-ate a “teeter bed” of suspended particles. The small interstices within the bed create high interstitial liquid velocities that resist the penetration of the slower settling particles. As a result, intermediate fine/ light particles continue to accumulate in the upper section of the separator and are eventually carried over the top of the device into a collection launder. Large or heavy particles, which settle at a rate faster than the upward current of rising water, eventually pass through the fluidized bed and are discharged out one or more restricted ports through the bottom of the separator.

The separation provided by a teeter-bed separator is governed by the following:

(1)

where Ut is the hindered-settling velocity of a particle (m/sec), g is the acceleration due to gravity (m/sec2), d is the par-ticle size (m), ρs is the density of the solid particles (kg/m3), ρf is the density of the fluidizing medium (kg/m3), η is the apparent viscosity of the fluid (kg·m-1·s-1), φ is the volumetric concentration of solids, φmax is the maximum concentration of solids obtainable for a given material, and β is a function of Reynolds number (Re). It should be noted that Eq. (1) was derived from an expression advocated by Masliyah (1979) and modified utilizing relationships offered by Richardson and Zaki (1954) and Swanson (1989). An in-depth derivation of Eq. (1) is not reviewed here but is offered elsewhere (Kohmuench et al. 2002 and 2006).

By inspection of Eq. (1), it can be determined that both the size of a particle and its density greatly influence how that particle will settle within a hindered state. As such, teeter-bed separators are frequently used in the minerals processing industry to segregate fine particles according to size, shape or density (Wills, 1997). Typically, if a size distribution is very narrow and there is a large variation in particle specific gravity (e.g., mineral sands), then a teeter-bed separator can be used to segregate material based on density. Likewise, if a size distribution is very wide and the density distribution is relatively tight (e.g., silica sands), a teeter-bed separator can be successfully utilized as a classifier.

Key words: Teeter-bed separators, Hydroseparation, Separation


Typical control parameters and responseTeeter-bed separators are considered operator-friendly and

simple to control. Traditionally, there are only two primary control variables, which include i) the fluidization rate and ii) the teeter-bed level. The fluidization rate is the amount of water or other fluidizing medium that is injected into the cell against which the material must settle. This parameter is typically reported as the volume flow of liquid required for a given cross-sectional area (m3/hr/m2); however, it can also be reduced further and reported as a liquid rise velocity (cm/sec). The teeter-bed level is essentially an indication of the mass of the settled material held up within the device and is typically reported as a relative level reading, or as an average bed pressure.

This basic control scheme is illustrated in Fig. 1. In this ap-proach, feed is presented to the separator and allowed to settle against the upward current of the fluidization medium, which in most cases is process water. The water rate can be controlled manually or automatically using a simple PID control loop. In this latter approach, a flow meter (FIT) provides feedback information to a flow-indicating controller (FIC) that makes adjustments to an automatic flow control valve (FCV) in order to maintain the fluidization flow constant.

The bed level is controlled in a similar manner, where the weight of material within the separator is measured using a pressure-sensing device. The pressure-sensing device is typi-cally a pressure or level transducer (LT) that can be mounted to the side of the unit or suspended from above. It should be noted that this indication of level is commonly, but incorrectly, referred to as a density reading. Regardless, the pressure sen-sor provides a feedback signal to a level-indicating controller (LIC) that makes adjustments to an automatic underflow control valve (LCV) to maintain a constant bed level.

Fluidization (teeter-water) rate. Given that the particles must settle against the upward current of fluidization flow within a teeter-bed separator, it is clear that this operating parameter has the most direct influence on separation cut size (d50) which is defined as the particle size having a 50% probability of reporting to either the overflow or underflow product stream as determined from a Tromp separation curve. As a result, the teeter-water rate is considered a coarse tuning parameter and must be sufficient to maintain the bed fluid. If the teeter-water rate is insufficient, the material within the separator will sand out and no separation will occur.

Data show that there is generally a linear relationship be-tween the fluidization rate and the separation cut-point (d50). The relationships for two different mineral applications are shown in Fig. 2. Included in this figure are data for classifica-tion of 1.00 x 0.106-mm phosphate ore. As seen, a fluidization rate between 0.5 and 1.5 cm/s provides a range of cut points between 0.25 and 0.65 mm. Also shown in Fig. 2 are data for a 3.00 x 0.85-mm potash application. In this case, cut points range from 0.4 to 1.2 mm for fluidization rates of 0.6 to 2.5 cm/s.

As expected, the denser phosphate requires a higher fluidi-zation velocity to affect the same cut-point as the less dense potash (2.7 vs. 2.0 SG). However, there is only a moderate difference in the medium-rise velocity for a given cut-point. This result can be attributed to the influence of other process variables, including the feed particle size distribution and the liquid medium characteristics. In this example, the potash is significantly coarser than the phosphate. In addition, it should be noted that the potash classification is carried out using saturated brine as the liquid medium (ρf = 1.24 SG) which influences the

hindered-settling velocity as seen in Eq. (1). The higher liquid specific gravity effectively raises the apparent density within the separator and increases the cut point when compared to process water (ρf = 1.00 SG).

The effect of mineral density is more clearly seen in Fig. 3. In this figure, cut size (d50) is shown as a function of upward current velocity for various mineral densities. As seen, denser material requires a greater rise in velocity in order to achieve the same separation cut point. In other words, it takes a greater amount of flow to adequately fluidize denser particles and maintain them in a hindered suspension.

Teeter-bed level (bed pressure). Unlike the fluidization rate, which expands or contracts the teeter bed, the bed-level adjustment is used to modify the proportion of coarse/heavy material retained inside the separator. As such, bed level can be considered a fine-tuning parameter that essentially modifies the height or accumulation of these particles within the teeter

Figure 1 — Typical control system for TBS units.

Figure 2 — Cut point vs. fluidization rate.


bed. A typical response shows that the relative reading provided by the level transducer rises as the amount of coarse/heavy particles increases. In other words, as more coarse material is held up within the separation chamber, the apparent weight of the suspension is increased. As a result, a specific level reading can be maintained as a control set point that regulates the discharge from the underflow valve, thereby keeping this parameter constant.

A typical response of separation cut point (d50) versus bed pressure is provided in Fig. 4. Two cases are provided that show the effect of bed pressure on the cut-point for both a relatively low (1.70 cm/s) and high (2.55 cm/s) fluidization rate. It can be seen in each case that as the bed pressure in-creases, the separation cut point also increases. This trend is not surprising, given that a teeter bed in a classification operation consists of particles that are being segregated based on size

and hindered-settling velocities. As such, the finest material with the slowest settling velocity will accumulate near the top of the teeter bed and the coarser particles with the high-est settling rate will penetrate the teeter bed and accumulate at the bottom of the separation chamber. Changes in the bed pressure set point effectively dictate the overflow weir level with respect to particle size.

Design and control enhancementsChanges in equipment design and control are derived from

a demand to enhance the separation that is currently achieved through existing means. In most cases, this implies improving the overall efficiency of the unit operation. For classification ap-plications, the sizing efficiency is typically determined through inspection of the partition curve generated for the separation. An example of two partition curves is provided in Fig. 5. The partition factor represents the recovery of dry solids from the feed to the underflow (oversize) product for each size class.

The partition curve is used to determine the separation ac-curacy by approximating the slope of the line at the cut point (d50). This value is typically reported as the Ecart probable (Ep) error, i.e.:

(2)

where d75, d50, and d25 represent the size at which 75%, 50%, and 25% of the feed mass reports to the underflow of the separator, respectively. It should be noted that a lower Ep value reflects a steeper curve and thus a better separation. A vertical line represents a perfect separation. Because Ep values vary with d50, a more useful term is imperfection (I), which allows for the direct comparison of separations that occur at different cut points. As seen in Eq. (3), imperfection is simply the Ep normalized to the d50:

(3)

As shown by Lynch (1977), an alternative approach to determining the relative sharpness of the separation is to fit

Figure 3 — Cut point vs. teeter rate using water as a fluidizing medium for various particle densities.

Figure 4 — Effect of bed level on separation cut point (d50).

Figure 5 — Separation curves for industrial sizers.


measured sizing data using an empirical partition function such as:

(4)

in which P is the partition factor, d is the particle size, d50 is the particle size cut point (defined at P = 50%), and α is a parameter that reflects the sharpness of the size separation (defined as the slope at P = 50%). Note that unlike Ep or Imperfection, a larger value of α indicates a sharper (more efficient) particle size separation.

Figure 5 illustrates a case study, in which two units were compared side-by-side with respect to sizing efficiency. In this case, one of two existing hydraulic classifiers was retrofitted with an Eriez CrossFlow feed presentation system. The Cross-Flow design uses an improved feed delivery system that gently introduces the feed slurry across the top of the separator as opposed to injecting the slurry at a high velocity directly into the teeter bed, as seen in Fig. 6. Feed is effectively presented horizontally across the top of the cell towards the overflow launder. Compared to conventional systems, the tangential feed introduction ensures that variations in feed slurry characteristics (e.g., solids content) do not impact separator performance. In the CrossFlow, the teeter-water velocity remains constant through-out the separation chamber and the overall design minimizes turbulent mixing, which can be detrimental to performance.

The resultant data from the full-scale, side-by-side evaluation are presented in Table 1. As seen in this table, the CrossFlow significantly reduced the particle cut size from 729 to 362 mi-crons, with only the feed presentation system altered. In other words, all other operating parameters were the same for each of the separators except for the feed arrangement. Tonnages to both units were equal and left unchanged; however, the sizing efficiency of the CrossFlow improved substantially as indicated by the increase in alpha from 3.4 to 8.1. Furthermore, the amount of misplaced coarse solids (+0.425 mm) was reduced from 9.0% to only 1.7%.

A second round of testing was completed in order to compare the two classifiers at a similar particle size cut point. Since the separation cut point (d50) had been reduced simply by changing the feed arrangement, the teeter water and bed level of the retrofitted unit were adjusted until the cut size between the two units was similar. The separation curves for the two units normalized to their respective cut points are those shown in Fig. 5. It is easily seen that the retrofitted classifier offers a

much sharper separation curve when compared to the existing unit. A comparison of the separation curves indicates that the retrofitted separator operated with a 33% higher efficiency.

The design of the CrossFlow Separator is based on maintain-ing a constant, quiescent and precise upward velocity within the separation chamber of the unit. Hydrodynamic studies indicate that quiescent flow/non-turbulent conditions must exist in a teeter-bed separator to maintain a high efficiency (Heiskanen, 1993). Excessive turbulence or changes in flow conditions can result in the unwanted misplacement of particles and a corresponding reduction in separation efficiency.

Unfortunately, conventional hydraulic separators typically utilize a feed injection system that discharges directly into the main separation chamber, which creates excessive turbulence. A larger issue with this type of feed injection system is the discontinuity in flow velocity created by the additional water that enters with the feed solids and reports to the overflow launder. The conventional approach greatly increases the rising water velocity at or above the feed entry point. As a result, the volume of water entering with the feed slurry at higher feed rates is likely greater than the volume flow of teeter water required for proper particle fluidization. The discontinuity created by the feed water often results in a secondary interface of fluidized solids, which varies uncontrollably as the solids content of the feed varies (Fig. 6).

Some classifiers utilize a flat-bottom arrangement. In this approach, a bottom plate is installed directly beneath the tee-ter zone, to which one or more underflow discharge nozzles are fitted. These nozzles are placed in close proximity to the fluidization system. Given this arrangement, there is no clear

Figure 6 — Water flow velocities in classifiers.

Table 1 — Comparison of full-scale classifiers.

Test variable Existing CrossFlow separator

Test 1 –

Particle cut size 729 μm 362 μm

Alpha value 3.4 8.1

Misplaced +0.425-mm 9.0% 1.7%

Test 2 –

Particle cut size 490 μm 420 μm

Alpha value 3.2 7.2

Imperfection 0.162 0.109


delineation between the fluidization zone and the underflow discharge area. As a result, the underflow stream will be rela-tively dilute, given that the material is being pulled from an area of low percent solids. More importantly, the underflow solids content will vary significantly as the underflow valve opens and closes, which can greatly vary the upward fluidiza-tion velocity within the separation chamber.

To better manage these fluctuations and improve overall control, a dewatering cone is used to maintain a constant wa-ter split within the separator. In essence, this cone is located underneath the separation chamber. As particles settle past the teeter pipes, they drop into this sanded zone and move via mass-action to the underflow valve. The purpose of the dewater-ing cone is to force the maximum amount of liquid out of the coarse stream, thereby creating a highly dewatered product. Under normal operating conditions, the solids content of this underflow stream will remain very consistent (+/- 2%). As a result, the upward water flow rate also remains very consistent, which improves separation efficiency.

Effective (calculated) teeter-bed density. The importance of proper fluidization control and how this parameter can greatly influence particle setting rates through changes in the effective teeter-bed density is seen in Fig. 7. In this illustration, a known amount of silica sand is fluidized in a laboratory-scale (50 x 200-mm) CrossFlow separator. Two cases were investigated that included the use of both high and low fluidization rates. In Fig. 7a, the sand is fluidized using 5 L/min of process water. As seen by the relative position of the sand-water interface, the bed is expanded to within 50 mm of the overflow weir. In the second example (Fig. 7b), the sand is fluidized with only 3 L/min of process water. As a result, the sand-water interface has dropped significantly. It should be noted that no sand has been added or removed from the system during this exercise.

More importantly, a free-floating, small glass vial partially filled with sand, water, and air was positioned in the teeter zone to act as a visual reference for density. In the 5 L/min case, the bed is highly fluidized and expanded. While the top of the teeter bed is high, the relative density of the teeter bed is quite low, with the glass jar nearly submerged. In contrast, at the lower fluidization rate, the bed contracts and the voids between the suspended particles become smaller. As a result, the relative density is quite high, with the glass jar less than half submerged. The higher density within the teeter zone provided additional buoyancy to the vial.

The complex interaction between fluidization flow, bed level and the true effective bed density is also revealed through further inspection of Fig. 4. In this figure, bed level readings are shown for the same application (i.e., the same solids-specific gravity) while using two fluidization rates. It is seen that the higher fluidization rate resulted in lower bed level readings (at the same cut point), which is a clear indication of bed expan-sion. However, it can also be deduced from the discussion above that a greater degree of expansion will result in lower calculated bed densities. While the interaction between these parameters can be challenging to quantify, each are measurable and therefore susceptible to automatic control.

Based on these observations, a new process control strategy was implemented by Eriez for managing the vari-ous operating parameters of teeter-bed-style separators. In this approach (Fig. 8), two transducers are used instead of a traditional single-pressure sensor. These sensors are installed at different elevations but in close proximity to one another. The lower sensor is used to control the bed level as previously described. This sensor provides the process input for a level indicating control (LIC) which proportionally opens and closes the underflow valve (LCV) to maintain the mass above the sensor constant.

The upper pressure transducer is used in conjunction with the bottom sensor in order to calculate the true density of the teeter zone located between the sensors. This bed density (ρb) is calculated using the following equation derived from first principles:

(5)

where ΔP is the differential pressure reading calculated from the two pressure transducers, A is the cross-sectional area of the separator, Vz is the volume of the zone between the two sensors, and H is the elevation difference between the sensors. This calculated parameter is monitored by a den-

Figure 7 — Effect of water velocity on bed density.

Figure 8 — Updated control system for TBS units.


sity indicating controller (DIC) which then finely adjusts the fluidization flow in order to maintain the bed density stable. This control scheme can be advantageous in situations where feed or product characteristics change on a regular basis (e.g., particle size distribution or solids content).

Full-scale case study – enhanced fluidization control. A teeter-bed-based separator was installed at a phosphate beneficiation plant. The original control scheme for this separator included the use of manual control valves for the adjustment of the fluidization water rate. Unfortunately, the feed characteristics were constantly changing, due to vari-able operator settings and natural matrix inconsistencies. For example, when the feed to the separator became significantly coarser, the teeter bed would be insufficiently fluidized (i.e., sanded). As a result, operators hedged against this occurrence by operating at relatively high fluidization rates, which resulted in misplacement of material during normal operation or when the feed size distribution became much finer.

To combat these issues, the existing separator was outfit-ted with an automatic process water flow control valve and a separate differential pressure transducer. This allowed the separator to be controlled, as shown in Fig. 8. Prior to automat-ing the teeter-water addition rate, the response of the calculated bed density with varying teeter-water rate was investigated. A clear correlation between these two parameters was established as seen in Fig. 9. In addition, it was found that with all other operating parameters equal, coarser feed will cause an increase in bed density due to its greater mass and required fluidization rate. In contrast, fine feed will result in a lower bed density and fluidization rate.

Given the correlation seen in Fig. 9, the circuit was placed under automatic control. For the purpose of proving the concept, a teeter-bed density of 1.660 SG was chosen as the operating set point. As such, the density-indicating controller supplied a continuously updated set point for the teeter-water control in order to maintain a constant bed density. This approach provided for an extremely stable process, as demonstrated by the flat line trending produced by the plant’s distributed

control system (DCS) as seen in Fig. 10. This figure illustrates the controlled response during a period when the feed particle size distribution became significantly finer. As such, the teeter-water rate was reduced in order to maintain the calculated bed density of 1.660 SG.

Accurate control of the teeter water requires a trouble-free fluidization system, as characterized by constant and uninter-rupted fluid flow. As such, the design of the water delivery system is an important issue in the development of a hydraulic separator. Conventional teeter-bed designs use a series of lat-eral pipes or a steel plate located at the base of the separation zone. These pipes and plates are perforated at regular intervals with large numbers of small-diameter holes. Elutriation water is injected through these holes over the entire cross-section of the separator. The large water flow rates, combined with the small injection hole diameters, leave the device susceptible to blockage/plugging due to contaminants in the process water. When several orifices become blocked, a dead zone occurs in the fluidization chamber, resulting in a loss of performance in this area. As a result, conventional teeter-bed separators have an inherent design flaw that limits both the capacity and efficiency of the unit.

In contrast, the CrossFlow separator incorporates another design feature that improves water distribution. A novel ap-proach has been developed that incorporates a slotted plate to disperse the elutriation water across the base of the separa-tor. In this design, a horizontal slotted plate is located at the base of the separation chamber. Water is introduced beneath the plate through a series of large diameter holes (>1.25 cm) which are significantly larger than the largest particle. Unlike existing separators, these orifices are located at distant intervals (typically >15 cm) and serve simply to introduce the water, while water dispersion is achieved by the baffle plate. This modification essentially eliminates problems associated with plugging of distributor plates or pipes. In addition, the water distribution system incorporates a check valve to prohibit backflow and manifold valves that facilitate clean-out in the unlikely event of a plugging issue. This arrangement includes manually actuated valves that are located on each end of the

Figure 9 — Teeter water and density correlation. Figure 10 — Example of bed density control.


water supply manifolds. This approach allows each manifold to be easily flushed by diverting flows from one side of the unit to the other.

SummaryTeeter-bed separators are efficient classification devices that

provide a means to separate coarse and fine material according to hindered-settling relationships. Material characteristics and operating variables, such as fluidization water flow and bed level, can greatly affect the separation. As such, it is important to understand how these parameters interact with one another.

In general:

1. The fluidization rate is considered a coarse tuning pa-rameter. Data show that there is a linear relationship between teeter-water addition rate and the separation cut point (d50).

2. Bed level is considered a fine-tuning parameter, allowing for subtle changes in separation cut point (d50). An increase in bed level increases the d50 by allowing more coarse/dense material to accumulate within the teeter zone of the separator.

3. The classification efficiency of teeter-bed separators is commonly quantified through examination of partition curves. The Ecart probability (Ep), imperfection (I), and alpha (α) are common terms used to define the steepness (i.e, sharpness) of the slope of these separation curves at the process cut point.

4. Through the examination of partition curves generated from a full-scale investigation, an Eriez CrossFlow Separator was shown to provide improved separation when compared to a traditional classifier. This improvement is attributed to the more efficient management of the water that arrives with the feed solids.

5. An improved control strategy was implemented based on laboratory and field observations. This control method involves adjusting teeter-water rates to maintain a constant calculated

bed density. A second pressure transducer is required for this approach, which provides additional control stability for vari-able processes.

6. This improved control strategy was implemented on a production scale at an industrial phosphate plant. By monitor-ing and correcting for the calculated bed density, operating parameters were automatically adjusted to combat the changes in feed characteristics.

7. Given that ease of maintenance is an important aspect for reliable service, additional improvements were made to the teeter pipe and manifold arrangement to facilitate clean-out of the fluidization system in the unlikely event of fouling.

AcknowledgmentsThe authors would like to acknowledge and thank the mul-

tiple industrial representatives who participated in the various test campaigns described in this work. Their contributions in terms of time, manpower, expertise and direct financial support are gratefully acknowledged.

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Reverse flotation of diaspore from aluminosilicates by a new cationic organosilicon quaternary ammonium collectorXinyang YuPh.D candidate, School of Chemistry and Chemical Engineering, Central South University, Changsha, Hunan Province and School of Resource and Environmental Engineering, Jiangxi University of Science and Technol-ogy, Ganzhou, Jiangxi Province, China

Hong ZhongProfessor, School of Chemistry and Chemical Engineering, Central South University, Changsha, Hunan Prov-ince, China

Guangyi LiuAssociate professor, School of Chemistry and Chemical Engineering, Central South University, Changsha, Hunan Province, China

AbstractThe flotation of diaspore, kaolinite, pyrophyllite, illite and the diaspore-kaolinite mixture were studied with a new cationic organosilicon quaternary ammonium collector entitled QAS222. The acting mechanism of the collector with the four minerals was analyzed by Zeta potential measurement and FT-IR spectrum analysis. The results indicate that the QAS222 was effective in the reverse flotation of the diaspore-bauxite mixture of different Al2O3 -to-SiO2 mass ratios (A/S ratio) at an optimum pulp pH of 11. The potentials of the four minerals increased when they were conditioned in the QAS222 solution; however, the potentials of the three aluminosilicate miner-als increased faster than that of the diaspore, especially at pH>9. Furthermore, they were positive in a wide pH range, indicating that QAS222 has a strong adsorption on the surface of the aluminosilicate minerals and thus that these minerals can be floated. The FT-IR spectra analysis further indicates that, besides electrostatic adsorp-tions, formation of the hydrogen bond and ammonium absorptions, chemical adsorptions occur between QAS222 and the aluminosilicate minerals, which allow for the QAS222 to be strongly adsorbed on the aluminosilicate minerals. However, these reactions between QAS222 and diaspore at pulp pH=11 are weaker than those of the aluminosilicate minerals. Thus, surface adsorptions cannot be formed effectively, resulting in reduced floatability.

Paper number MMP-09-036. Original manuscript submitted October 2009. Revised manuscript accepted for publication May 2010. Discussion of this peer-reviewed and approved paper is invited and must be submitted to SME Publications Dept. prior to February 28, 2011. Copyright 2010, Society for Mining, Metallurgy, and Exploration, Inc.

IntroductionIn China there is abundant bauxite; however,

around 98% of the resource is diasporic bauxite and more than 70% of the resource has an Al2O3-to-SiO2 mass ratio (A/S ratio) lower than 7. It is therefore necessary to desiliconize bauxite and to increase its A/S ratio if the low-cost Bayer process will be used to produce Al2O3 from the bauxite. Though direct flotation is used in the bauxite processing industry, its disadvantages, such as the difficulty in dewatering concentrate and its high reagent consumption, have restricted its wide application in industry. Therefore, the study of bauxite processing with a reverse flota-tion process, especially the search for high-effective flotation collectors for aluminosilicate minerals, is important to the Chinese bauxite processing industry.

In recent years, many flotation collectors for aluminosilicate minerals in the reverse flotation process, notably alkylamines (Jiang et al., 2001), quaternary ammonium salts (Zhao et al., 2007), N-(2-aminoethyl)-dodecanamide (Zhao et al., 2003a), N-(2-aminoethyl)-1-naphthalene-acetamide (Zhao et al, 2003b), N-(3-aminopropyl)-dodecanamide (Zhao et al., 2003c), N-dodecy1-1, 3-diaminopropanes (Hu et al., 2003), γ-alkyl-propylamines (Cao et al., 2004), methylnaphthalene amine (Du et al., 2003), dedecylguannidine sulfate (Zhong et al., 2008), Gemini quaternary ammonium (Xia et al., 2009), dodecyl tertiary amines (Liu et al., 2009) and alkylguanidine (Guan et al., 2009) have been reported effective in collecting aluminosilicate minerals. However, most of these collectors are not used in industry due to their deficiencies, such as the excessive amount of foams, high cost, high reagent consump-tion, low selectivity, etc.

Key words: Flotation, Bauxite, Collectors


Cationic organosilicon quaternary ammonium, a new sur-factant with bioactivity, is innocuous, effective, multifunctional and antibacterial (Liu et al., 2007); it is now widely used in the fields of fabric finishing, chemical engineering, pharmaceuticals and agricultural chemicals (Su et al., 2008). However, there is little information regarding the use of this surfactant for mineral flotation. Therefore, a new cationic organosilicon quaternary ammonium collector, entitled QAS222, was synthesized and its performance in the reverse flotation of diaspore, kaolinite, pyrophyllite, illite and the diaspore-kaolinite mixture was investigated. The acting mechanism of this collector with the four above minerals was analyzed by Zeta potential measure-ment and Fourier transform infrared (FT-IR) spectrum analysis.

ExperimentalMaterials and reagents. The diaspore, kaolinite, pyro-

phyllite and illite materials used for the present study were obtained from Xiaoguan of Henan Province, Jiaxian of Henan Province, Qingtian of Zhejiang Province and Ouhai of Zhejiang Province of China, respectively. They were handpicked and then crushed and ground to less than 0.074 mm in a porcelain mill. The purities of the materials are higher than 95%, with chemical compositions as shown in Table 1.

NaOH (0.1 M) and HCl (0.1 M) solutions were used as pH modifiers and distilled water was used in all experiments. The new collector of organosilicon quaternary ammonium has a microstructure as illustrated in Eq.(1):

(1)

Flotation tests. Pure minerals (3.0 g) or artificially mixed

bauxite materials (3.0 g) were placed in a plexiglass cell (40 ml), which was then filled with distilled water. After adding the desired amount of reagents, the suspension was agitated for three minutes and the pH was measured before flotation. The flotation was conducted for four minutes. The products were collected, dried and weighed. The recovery was calculated based on the dry weights of the products obtained.

Zeta potential measurements. Zeta potentials were mea-sured using a Delsa−440SX Zeta potential instrument from the Beckman Coulter (Brea, CA). All measurements were conducted in a 0.1-M KNO3 background electrolyte solution. Materials were ground to less than 5 μm in diameter. 0.05 g of material was placed in a 100-ml breaker for five min with 80 ml distilled water and the pH was adjusted and measured. The results presented in this paper are the average of three

independent measurements with a typical variation of ±2 mv. Repeated tests showed that this conditioning procedure was capable of producing mineral surfaces suitable for studying the effect of various treatments.

FT-IR spectrum analysis. Diffuse reflectance infrared spectroscopy was used to characterize the surface species on the materials treated. Materials were ground to a diameter of less than 5 μm as they were prepared for the microflotation tests. FT-IR spectrum analysis was performed with an AVATAR360 FT-IR from Nicolet Corporation of USA, and was presented without any baseline correction.

Results and discussionFlotation performance of QAS222 with minerals. Mi-

croflotation tests were conducted to show flotation behaviors of diaspore, kaolinite, pyrophyllite and illite as functions of pulp pH and QAS222 dosage, to show the collecting ability of the QAS222.

Figure 1 plots the impact of pH on the floatability of the four pure minerals when 2.0×10-4 M QAS222 was used. As seen in Fig.1, the pulp pH has different degrees of influence on the floatability of the four minerals; however, it affects the recovery of diaspore the most. The recovery of diaspore increased with the increase of pulp pH and reached its highest point (>95%) at around pH=4.5; then it decreased drastically with further increase of pulp pH until it approached nearly zero above pH 12. Under acidic conditions, i.e., pH<7, the recoveries of kaolinite, pyrophyllite and illite were found to be practically independent of pulp pH and all three achieved relatively high recoveries. Specifically, kaolinite and pyrophyl-lite were higher than 80% and illite was higher than 60%; under

Table 1 — Chemical composition of minerals tested (w/%).

Mineral Al2O3 SiO2 Fe2O3 TiO2 CaO MgO K2O Na2O Water Loss

Diaspore 80.98 0.78 0.29 2.84 0.01 0.046 0.007 0.025 14.06 14.50

Kaolinite 39.2 43.67 0.32 1.98 0.01 0.068 0.094 0.028 13.65 13.98

Pyrophyllite 32.16 59.65 0.16 0.20 0.01 0.019 0.62 0.17 6.03 6.33

Illite 37.71 45.97 0.14 0.21 0.01 0.068 10.2 0.22 4.51 4.85

Figure 1 — Effect of pulp pH on the performance of QAS222.


alkaline conditions, the recoveries of the three aluminosilicate minerals decreased at different rates with the increase of pulp pH, but they were still much higher than that of the diaspore. According to the results shown in Fig.1, the diaspore achieved a much lower recovery at a pulp pH>9, compared with the recoveries of the other aluminosilicate minerals, indicating a significant tendency for reverse flotation of the diaspore from the aluminosilicate minerals.

The flotation of diaspore, kaolinite, pyrophyllite and illite were further conducted at pulp pHs of 9, 11 and 13, respectively, to study the impact of QAS222 dosage on their floatability.

As shown in Fig.2, when the pulp pH is fixed at 9, the recoveries of the four minerals increased with the increase of QAS222 concentration; the recoveries of diaspore and kaolinite are higher than 85% and the recoveries of pyrophyllite and illite are round 80%, when the QAS222 concentration was 4.0×10−4 M. Thus, it is difficult to separate diaspore from kaolinite, pyrophyllite and illite when pulp pH is around 9.

When the pulp pH is adjusted to 11, as shown in Fig. 3, the recoveries of kaolinite, pyrophyllite and illite increased rap-idly with the increase of QAS222 concentration, and reached 90.12%, 72.47% and 68.74%, respectively, as the QAS222 concentration reached the maximum of 4.0×10−4 M. However, the recovery of diaspore was very low, less than 10%, and did not change with the increase of QAS222 concentration. There-fore, the diaspore may be effectively separated from kaolinite, pyrophyllite and illite under a higher pulp pH condition.

The pulp pH was further increased to 13 in order to further investigate the flotation behaviors of the four minerals with the QAS222 as collector. As seen in Fig. 4, the recoveries of kaolinite, pyrophyllite and illite are much lower than those achieved at pulp pH=11; this is due to the fact that the three aluminosilicate minerals are depressed under an excessively high alkaline pulp condition.

Thus, QAS222 may be an effective collector for the three aluminosilicate minerals. In addition, it seems that the pulp pH should be higher than 9 if these minerals are to be effectively removed from diasporic-bauxite with QAS222 as a collector in the reverse flotation process. Although the three alumino-

silicate minerals showed good floatability at a pulp pH<9, their recoveries were similar to that of the diaspore as the QAS222 concentration increased and there was no favourable condi-tion to perform reverse flotation. It can be concluded from the results above that the optimum pulp pH is around 11, where the aluminosilicate minerals showed good floatability. Such a pulp pH level is also favourable for a reverse flotation manipulation.

Reverse flotation of the diaspore-kaolinite mixture. Reverse flotations of the artificially mixed bauxite of different diaspore-to-kaolinite mass ratios were carried out to investigate the possibility of separating diaspore from aluminosilicate min-erals with QAS222 as collector. The concentrate and tailings products were weighed separately after filtration and drying and the A/S ratio was determined by silicon-molybdenum blue colorimetry. The recovery was calculated based on the dry

Figure 2 — Effect of QAS222 on the floatability of minerals at pulp pH=9.




weights of products obtained. The flotations were performed at a pulp pH of 11 and at a QAS222 concentration of 4.0×10−4 M, with the results illustrated in Table 2.

It can be seen from Table 2 that the QAS222 has good selectivity and a strong collecting affinity for kaolinite. It achieves an effective reverse flotation at a pulp pH=11. When

the A/S ratio of feed is 2.7, the A/S ratio, Al2O3 grade and Al2O3 recovery of concentrate product are 12.82, 77.79% and 69.91%, respectively. As the A/S ratio of the feed increases, the A/S ratio, grade and recovery of the concentrate product also increases. It is therefore concluded that the QAS222 is an effective collector in the reverse flotation of diaspore from

aluminosilicate minerals.

Zeta potential measurements of minerals. The Zeta potentials of the four minerals conditioned with and without the QAS222 (4.0×10−4 M) were measured as a function of pulp pH. It should be noted that the iso-electric points (IEPs) of the diaspore, kaolinite, illite and pyrophyllite in distilled water are 6.32, 3.4, 3.0 and 2.5, respectively, which are in accordance with those reported by Zhao et al. (2003c).

As shown in Fig. 5, the IEP of diaspore increased from 6.32 in distilled water to 10.9 when conditioned with the QAS222 solution; obviously, the QAS222 has an effect on the Zeta potential of diaspore, indicating that QAS222 was adsorbed on the diaspore surface and thus may increase its floatability. However, under alkaline conditions, the Zeta potential of the diaspore begins to decrease, especially at pulp pH>9, indicating that the adsorption of QAS222 on the surface of the diaspore has decreased and thus may lessen its floatability. The above analysis is in accordance with the results of flota-tion tests above.

As shown in Fig. 5, under acid conditions the Zeta potentials of kaolinite, pyrophyllite and illite increased with the increase of pulp pH, as there are steady ad-sorptions of the QAS222 on their surfaces, and then these aluminosilicate minerals have good floatability. However, the potentials of the aluminosilicate minerals decreased at different rates at pulp pH>8, indicating that the adsorption of QAS222 on these minerals has decreased and thus their floatability has deteriorated. Measuring the Zeta potential of the aluminosilicate minerals further confirmed the conclusions above. In addition, as shown in Fig. 5, the IEPs of kaolinite, pyrophyllite and illite are 3.4, 2.5 and 3.0, respectively, in the absence of QAS222. After they were conditioned with 4.0×10−4 M QAS22 solution, their Zeta potentials are all positive, in the range of pulp pH 2-12 and their increased values of Zeta potential are always bigger than that of diaspore. It is clear that QAS222 was more strongly adsorbed on the aluminosilicate minerals and may achieve higher selectivity for these minerals.

Table 2 — Reverse flotation results of the diaspore-kaolinite mixture.

Diaspore-to-kaolinite

mass ratio of mixture

A/S ratio of mixture

Concentrate grade (%) Concentrate recovery (%) A/S ratio of concentrateAl2O3 SiO2 Al2O3 SiO2

1:1 2.7 77.79 6.07 69.91 14.75 12.82

3:2 3.58 77.45 5.53 73.75 18.86 14.01

2:1 4.45 78.27 5.02 77.74 22.17 15.59

3:1 6.13 79.41 4.36 81.07 27.30 18.21

Figure 6 — FT-IR spectra analysis of diaspore conditioned with and without QAS222.

Figure 5 — Zeta potentials of minerals with 4.0×10-4 M QAS222 as a function of pH.


FT-IR spectra analysis of minerals. To further reveal the acting mechanism of QAS222 with the four minerals, the FT-IR spectra of the minerals conditioned with and without 4.0×10−4 M QAS222 solution at a pulp pH of 11were measured and are presented in Figs. 6, 7, 8 and 9, respectively.

As shown in Fig. 6, there are no new spectrum adsorbing peaks on the diaspore surface when it was conditioned with QAS222, indicating that little QAS222 was adsorbed on the diaspore surface. Such a phenomenon is in accordance with the flotation results above.

However, as shown in Fig. 7, the spectra of kaolinite affected by QAS222 exhibit weak peaks at 2926.50 cm−1 and 2855.01 cm−1 respectively, as a result of the stretching bonds of CH3 and -CH2, indicating that QAS222 is adsorbed on the kaolinite surface by physical electrostatic adsorption. The peak at 1637.27 cm−1 is attributed to the bending mode of H-O-H on the kaolinite spectrum. This peak shifts to the higher wavenumber when it was affected by the QAS222, indicating that there exist hydrogen bonds between the QAS222 molecule and the kaolinite (Liu et al., 2007). Another two peaks at 1309.43 cm-1 and 1267.76 cm-1, respectively, are due to the stretching vibration of the C-N of qua-ternary ammonium; these peaks disappeared when the kaolinite was affected by QAS222, indicating that ammonium is adsorbed on the kaolinite surface.

Though there are no new strong adsorbing peaks in the spectra of kaolinite when it was exposed to QAS222, chemical adsorptions may exist between kaolinite and QAS222. As shown in Fig. 7, the peaks in the 1000-1110 cm-1 region may be attributed to the Si-O stretching vibration and the peak at 912.69 cm-1 is caused by the Al-O stretching vibration. The peaks at 793.64 cm-1, 753.90cm-1 and 695.47 cm-1, respectively, result from the Si-O-Si symmetrical stretching vibration and the peak at 538.18 cm-1 is due to the Si-O-Al stretching vibration. These peaks broaden distinctly when the kaolinite is affected by QAS222. As there are Si-O bonds in the QAS222 molecule lattice, the Si and O atoms may link up with other Si and O atoms on the kaolinite surface and thus produce more bonds of Si-O-Si, Si-O and Si-O-Al on the surface, leading to broadened spectra peaks, being a representatively chemical adsorp-tion. From the FT-IR spectrum analysis above, it is clear that the increased floatability of kaolinite is due to the affinity of QAS222.

As shown in Figs. 8 and 9, the spectra of py-rophyllite and illite are similar to that of kaolinite when they were affected by QAS222; it is thus obvious that the acting mechanisms of QAS222 with pyrophyllite and illite are similar to that of QAS222 with kaolinite.

In summary, besides physical electrostatic ad-sorptions between QAS222 and the three alumino-silicate minerals, chemical adsorptions that link up bonds of Si-O-Si, Si-O and Si-O-Al allow for the QAS222 to be strongly adsorbed on the surfaces

Figure 7 — FT-IR spectra analysis of kaolinite conditioned with and without QAS222.

Figure 8 — FT-IR spectra analysis of pyrophyllite conditioned with and without QAS222.

Figure 9 — FT-IR spectra analysis of illite conditioned with and without QAS222.


of the three minerals, so that they show good floatability under alkaline conditions. As for diaspore, it is difficult for QAS222 to be adsorbed on its surface, resulting in lessened floatability.

ConclusionsIt can be seen from the results and discussion above that:- QAS222 has strong selectivity and collecting ability for

aluminosilicate minerals in a wide pH range and by using it as collector the diaspore may be effectively separated from the three aluminosilicate minerals under a strong alkaline condi-tion of pulp pH=11.

- The Zeta potentials of the four minerals increased when they were conditioned with the QAS222 solution; the poten-tials of the three aluminosilicate minerals were positive in the wide range of pH=2~12, as a result of the strong adsorption of QAS222 on their surfaces, thus increasing their floatability. However, the adsorption of QAS222 on the diaspore surface was weak and it was difficult for the mineral to be floated.

- The FT-IR spectrum analysis further confirmed the above analysis. At a pulp pH of 11, chemical and physical electrostatic adsorptions exist between QAS222 and the three alumino-silicate minerals, which allows for QAS222 to be steadily absorbed on the surfaces of the three minerals and results in their increased floatability. However, at the same pulp pH of 11, the diaspore presents no new spectrum-adsorbing peaks on the surface, indicating that little QAS222 was adsorbed. Accordingly, floatability was weak.

It is thus concluded that QAS222 is effective in collecting the aluminosilicate minerals and provides an effective collector in the reverse flotation of diaspore from aluminosilicate minerals.

AcknowledgmentsThe work is supported by the National Program for Key

Basic Research Projects of China (Grant No. 2005CB623701), the National Natural Science Foundation of China (Grant No. 50874118) and the Foundation for the Author of National Ex-

cellent Doctoral Dissertations of China (Grant No. 2007B52).

ReferencesCao Xuefeng, Hu Yuehua and Xu Jing, 2004, “Synthesis of γ-alkoxy-propylamines

and their collecting properties on aluminosilicate mineral,” Journal of Central South University of Technology, Vol. 11, No. 3, pp. 280-286.

Du Ping, Cao Xuefeng, Hu Yuehua, Jiang Yuren and Li Haipu, 2003, “Study of structure and property of amine collectors,” Light Metals, No. 1, pp. 27-31 (in Chinese).

Guan Feng, Zhong Hong and Liu Guangyi, 2009, “Flotation of aluminosilicate minerals using alkylguanidine collectors,” Transactions of Nonferrous Metals Society of China, Vol. 19, No. 1, pp. 228-234.

Hu Yuehua, Cao Xuefeng, Li Haipu, Jiang Yuren and Du Ping, 2003, “Synthesis of N-decyl-1, 3-diaminopropanes and its flotation properties on aluminium silicate minerals,” The Chinese Journal of Nonferrous Metals, Vol. 13, No. 2, pp. 417-420 (in Chinese).

Jiang Hao, Hu Yuehua, Qin Wenqing and Wang Dianzuo, 2001, “Mechanism of flotation for diaspore and aluminium-silicate minerals with alkylamine collectors,” The Chinese Journal of Nonferrous Metals, Vol. 11, No. 4, pp. 688-692 (in Chinese).

Liu Changmiao, Hu Yuehua, and Cao Xuefeng, 2009, “Substituent effects in kaolinite flotation using dodecyl tertiary amines,” Minerals Engineering, Vol. 22, No. 9-10, pp. 849-852.

Liu Guangyi, Zhong Hong, Hu Yuehua, Zhao Shenggui and Xia Liuyin, 2007, “The role of cationic polyacrylamide in the reverse flotation of diasporic bauxite,” Minerals Engineering, Vol. 20, No. 13, pp. 1191-1199.

Su Xiaoming, Zhong Hong and Yu Xinyang, 2008, “Synthesis and application of Quaternary Ammonium Silane,” Journal of Chengdu Textile College, No. 25, pp. 1-4 (in Chinese).

Xia Liuyin, Zhong Hong and Liu Guangyi, 2009, “Flotation separation of the aluminosilicates from diaspore by a Gemini cationic collector,” International Journal of Mineral Processing, Vol. 92, No. 1-2, pp. 74-83.

Zhao Shenggui, Zhong Hong and Liu Guangyi, 2007, “Effect of quaternary am-monium salts on flotation behavior of aluminosilicate minerals,” Journal of Central South University of Technology, Vol. 14, No. 4, pp. 500-503.

Zhao Shimin, Hu Yuehua, Wang Dianzuo and Xu Jin, 2003a, “Investigation on the flotation of Aluminosilicates Using N-(2-aminoethyl)-dodecanamide,” Acta Physico-chimica Sinica, Vol. 19, No. 6, pp. 573-576.

Zhao Shimin, Wang Dianzuo, Hu Yuehua and Xu Jin, 2003b, “Flotation of alu-minosilicates using N-(2-aminoethyl)-1-naphthalene-acetamide,” Minerals Engineering, Vol. 16, No. 6, pp. 1031-1033.

Zhao Shimin, Wang Dianzuo, Hu Yuehua and Xu Jin, 2003c, “Flotation of alumi-nosilicates using N-(3-aminopropyl)-dodecanamide,” The Chinese Journal of Nonferrous Metals, Vol. 13, No. 5, pp. 1273-1277.

Zhong Hong, Xia Liuyin and Liu Guangyi, 2008, “Flotation separation of diaspore from kaolinite, pyrophyllite and illite using three cationic collector,” Minerals Engineering, Vol. 21, No. 12-14, pp.1055-1061.

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