Desliming of dense minerals in fluidized beds

Minerals Engineering 39 (2012) 9–18

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Desliming of dense minerals in fluidized beds K.P. Galvin ⇑, J. Zhou, J.E. Dickinson, H. Ramadhani Centre for Advanced Particle Processing and Transport, Newcastle Institute for Energy and Resources, University of Newcastle, Callaghan, NSW 2308, Australia

a r t i c l e

i n f o

Article history: Received 23 March 2012 Accepted 25 June 2012 Available online 2 October 2012 Keywords: Fluidization Desliming Classification Inclined sedimentation Iron ore Drift flux

a b s t r a c t This paper applies Drift Flux theory for the first time to explore the relationship between the liquid and solids flow through fluidized beds in the context of desliming of dense minerals. At low solids fluxes the process is not flux curve constrained. Here continuity considerations indicate the net liquid flux through the lower bed is upwards, ideal for the removal of slimes from the underflow product. Moreover, the liquid split to the underflow is also lower, further reducing the slimes content of the underflow. At higher solids fluxes the net liquid flux becomes downwards and eventually the system becomes flux curve constrained. Under these circumstances slimes entrainment to the underflow increases significantly. In order to operate at these higher solids fluxes, and achieve efficient desliming, Split Fluidization should be used. Here additional fluidization liquid is added at a higher elevation, producing a net liquid flux in the upwards direction through the zone above the Split Fluidization entrance level. Desliming experiments covering a range of solids fluxes were conducted to investigate the effects of increasing solids flux and Split Fluidization. This study shows that a system of parallel inclined channels, a key feature of the Reflux Classifier, permits the introduction of this additional liquid while preventing the ultra fine product from being entrained to the overflow reject stream, thus permitting the efficient removal of slimes. Experiments were conducted using an iron ore feed with particles smaller than 0.50 mm in diameter. Efficient desliming at a relatively high solids feed flux of 20 t/m2 h was achieved, but to efficiently deslime at 40 t/ m2 h significant Split Fluidization was found to be essential. Ó 2012 Elsevier Ltd. All rights reserved.

1. Introduction With unprecedented demand for resources such as iron ore there is increasing need to beneficiate the fine particles. Existing technologies such as spirals produce an iron ore product typically in the size range from 0.1 to 1.0 mm, containing excessive levels of ultrafine particles of a low iron grade, and other material such as clays. These ultrafine particles contribute to the retention of excessive moisture, and in turn, to major problems in handling and transporting the product. Desliming, which is essentially a form of particle size classification, can be achieved using mechanical screens, or hydrodynamically using hydrocylones or fluidized bed hydrosizers. Mechanical screens, which provide a separation dependent on the physical size of the particle, suffer from a low capacity and the effects of wear, which in turn can cause a significant shift in the separation size. At ultrafine sizes, of order 0.050 mm, significant forces are required to overcome the surface tension that resists the flow of water through the fine apertures of the screen, hence the screen needs to be curved to promote a centrifugal force. Hydrodynamic methods produce separations dependent on the particle settling velocity ⇑ Corresponding author. Tel.: +61 2 40339077. E-mail address: [email protected] (K.P. Galvin). 0892-6875/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.mineng.2012.06.013

which depends on both particle size and density. This dependence leads to different separation sizes across the density range of the particles in the feed. In both approaches, ultrafine slimes tend to split with the water and hence to minimize this effect, the coarser solids should be recovered at a high solids concentration. Upward current classifiers, which are increasingly used to deslime fine product, present a number of difficulties. These devices operate as liquid fluidized beds, with water used to both fluidize the particles and wash or elutriate the ultrafine particles from the bed. Hence the ultrafines are transported upwards towards the outer overflow launder. One key concern is the tendency for fine, high grade particles to be entrained to the overflow, resulting in a loss in recovery. Therefore the upwards velocity needs to be kept below a critical level. Fluctuations in feed solids rates and pulp density are, however, unavoidable due to variations in the size distribution of the overall plant feed. With a significant fraction of the feed solids withdrawn in the downwards direction as an underflow stream, there is also a strong tendency for the water that discharges with the solids to entrain ultrafines with the underflow. Clearly, conventional upward current classifiers must operate within very tight constraints in order to limit a loss of the finer iron ore to overflow, and to limit entrainment of ultrafines such as clays into the underflow product. The Reflux Classifier, which is a novel

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fluidized bed system, is usually used in gravity separation. Fig. 1 shows the unit arrangement, with the fluidized bed housing evident in the lower section, and system of parallel inclined plates above. The feed slurry enters from above, plunging downwards, and then dispersing. Fluidization water enters via a distributor located below the lower conical section. When the bed density exceeds a set point value the solids discharge through the lower valve. A PID controller then adjusts the underflow discharge in order to maintain a fixed bed suspension density. The balance of the flow passes up through the system of inclined channels, reporting as overflow. Boycott (1920) was the first to report on the significance of inclining a tube containing a suspension. The particles segregate from the suspension at a much higher rate than those within a vertically aligned tube. The lamellae thickener, which achieves solid– liquid separation, and clarification, at a high rate relative to the vessel area is based on this principle. Nakamura and Kuroda (1937) and Ponder (1925) are credited with the first explanation for the observed increase in the rate of sedimentation, in terms of an increase in the effective sedimentation area. Zhang and Davis (1990) investigated the classification of particles from an inclined elutriator, incorporating a form of underflow recycle to achieve sharper separations. Galvin and Nguyentranlam (2002) produced a similar recycle effect using a system of inclined channels above a fluidized bed, referred to as the Reflux Classifier. The study of inclined sedimentation has led to a number of important advances in suspension mechanics. Lubrication theory predicted that smooth spheres should achieve a vanishingly small velocity during their approach towards an inclined surface. Smart et al. (1993) provided a definitive explanation for the finite translational and rotational velocities first observed by Carty (1957) of particles in motion down inclined surfaces, incorporating the effects of surface roughness and friction into their analysis. Leighton and Acrivos (1987) accounted for the influence of the shear rate on the diffusional transport of particles from the vessel wall using a Couette device. This work was also significant in explaining the complex particle–particle and particle–wall interactions that develop during inclined settling. Laskovski et al. (2006) later used Dimensional Analysis to arrive at an empirical expression for quantifying the tendency of a particle to either segregate or convey up through inclined channels of high aspect ratio (the ratio of the channel length to nor-

mal channel spacing). They quantified the limitations of the PNK theory, arriving at an asymptotic result for describing the effective increase in the sedimentation rate. King and Leighton (1997) developed a criterion for determining the inertial lift of a moving sphere in a shear field. Galvin et al. (2009) conducted elutriation experiments using closely spaced inclined channels, with a gap of 1.77 mm and channel length of 1.0 m, demonstrating how to elutriate particles according to their density. The inertial lift was sufficient to suspend and convey lower density particles, while permitting higher density particles to segregate and slide downwards. The importance of the lift force was quantified later by Galvin and Liu (2011) using the theory of King and Leighton (1997). This work led to a significant advance in achieving highly efficient gravity separation of fine particles (Galvin et al., 2010), with a significant reduction in the Ecarte Probable (Ep). The Reflux Classifier also provides a basis for effecting sharp size classification, as demonstrated in a previous study involving particles of fine silica (Doroodchi et al., 2006), with particle separations at typically 0.160 mm in diameter. They achieved an Imperfection, I, defined later in this paper, of only 0.07, equivalent to an error of only 15 lm. In order to achieve a given size separation additional water was added to the system to help convey the finer particles to the overflow. This water was initially added as part of the fluidization required through the base of the vessel. It was shown, however, that the fluidization superficial velocity should always be kept at just above the minimum in order to promote steady conditions at the underflow discharge. Excessive fluidization generated mixing, and hence poorer separation. Where additional water was required, it was concluded the water should be added in at a higher level in the system. They used the term Split Fluidization to describe this approach. The discussion applied to this work was based on a relatively simple bed washing analysis. In this paper, a rational approach to understanding the benefits of Split Fluidization is introduced for the first time, based on Drift Flux theory. The corresponding solids flux curve analysis is used to describe the constraints that link the settling particles and the fluid flow. Of specific interests are the direction of the interstitial water through the fluidized bed, and the fraction of the total water fed to the system that reports with the finite size particles to the underflow. The present study used the Split Fluidization approach, focusing on the benefits of introducing additional water either as part of the feed or separately at a lower position between the feed and fluidization entry positions. This additional water increases the tendency of valuable ultrafine particles to entrain and become lost via the overflow reject stream. In order to address this concern, a Reflux Classifier incorporating inclined channels, was used with the objective to retain these valuable particles as part of the product while promoting the process of desliming. The basic hypothesis of this applied study was that the quality of the desliming achieved would depend on the overall water split between the underflow and the overflow, and on the washing of the slimes from the feed.

2. Theory

Fig. 1. Schematic representation of the Reflux Classifier showing relevant liquid and solids fluxes. Note that the Split Fluidization is not always used. The ‘‘extra water’’ is introduced via an additional pump to simply aid the transport of the underflow.

In order to understand the effect of the fluidization on the process of desliming, it is useful to consider the feed as a binary system consisting firstly of settling particles of one finite size and density, and secondly non-settling particles, of a negligible size, which simply flow with the water. Thus in effect this notional binary feed can be considered in terms of the settling of a mono-component feed. Hence, any reference to the feed solids in this analysis corresponds to the particles of finite size only, and not to the slimes; the slimes are simply treated as being part of the fluid, and to split in accordance with the fluid. The following analysis is applicable to any upward current classifier including the Reflux Classifier shown in

K.P. Galvin et al. / Minerals Engineering 39 (2012) 9–18

Fig. 1. It is noted that the term ‘‘j’’ is used to denote the flux or superficial velocity of a given phase while ‘‘V’’ is used to denote key velocities related to the movement of the solid phase. Feed slurry enters the separator continuously at a volumetric solids flux, jp, with the finite size particles settling downwards to join the lower bed. The liquid feed flux is jf. The lower bed is fluidized via an upwards liquid flux, jw, which enters through a distributor at the base of the vessel. Underflow is withdrawn continuously from the lower bed at the same solids flux, jp, in order to preserve a fixed fluidized bed height. The corresponding liquid flux 0 reporting to the underflow is jL ; while the excess liquid reports to the system overflow at a liquid flux, jo. It is assumed that all of the entering solids of finite particle size discharge as part of the underflow, a useful first order approximation to the process of desliming. Our interest here is in the direction of the interstitial flow of the fluid through the bed, as this directly influences the process of desliming. If the net flux of liquid is in the upwards direction, then there are good prospects for reducing the slimes content of the underflow below that predicted by the fluid split, S, to underflow. We are interested here in establishing an understanding of the system constraints that govern the direction of this interstitial flow. Clearly, as the volumetric solids feed flux increases, the volumetric underflow rate must also increase in order to preserve a 0 fixed bed height. The liquid flux, jL ; reporting to the underflow must increase, as additional liquid must be associated with the 0 additional particles. If this underflow liquid flux, jL , becomes larger than the input fluidization flux, jw, then the net liquid flux, L, and interstitial flow through the bed, is downwards, and hence more slimes will reach the underflow, impacting on the quality of the underflow product. Conversely, if the net liquid flux, L, is upwards, then the slimes that report to the underflow must be those either attached to the particles or those in relatively close proximity to the surface of the coarser particles which are conveyed by association with the surrounding fluid. The particle settling velocity, Vs, relative to the container in a batch system can be described using a conventional hindered settling model of the Richardson and Zaki (1954) form,

V s ¼ V t ð1  /Þn

ð2Þ

Clearly, V 0p , is the slope of the modified flux curve. Note that, ¼ V p þ jw , where Vp is the propagation velocity in the absence of a fluidization flux, jw. In the case of batch settling there is no

V 0p

Fig. 2A. Conventional flux curve with no fluidization. The internal bed concentration, ub, is given by the tangent concentration, and the steady state underflow concentration is given by the intersection between the tangent and the horizontal axis. The slopes of construction lines, Vp, and Vs, are shown. The graph is effectively normalized by setting the terminal velocity, Vt = 1 m/s.

ð1Þ

where Vt is the terminal velocity, / the volume fraction of the particles, and n an exponent typically in the range 2.3–4.65, with the value increasing as the particle Reynolds number decreases. Drift Flux theory, which is based entirely on continuity considerations and the assumption that Vs is a function of / only, provides a one-dimensional description of the propagation velocity of a concentration wave in two phase flow (Kynch, 1952; Wallis, 1969). The flux curve, applicable to batch settling, is obtained using a plot of w ¼ /ðV s Þ versus /, as shown in Fig. 2A. Thus Fig. 2A shows the system operation in the absence of any fluidization, as occurs in a thickener. The plot is effectively normalized by setting the terminal velocity, Vt = 1 m/s. When fluidized through the base of the vessel at the flux, jw, the modified flux curve is produced using a plot of w ¼ /ðV s  jw Þ versus / as shown in Fig. 2B. In this study the Yoshioka flux curve construction has been used (Fitch, 1979), allowing the one batch settling flux curve, or modified flux curve, to be used to cover continuous operation at different underflow fluxes. In accordance with continuity considerations (Kynch, 1952) the propagation velocity, V 0p , of a given concentration wave within the fluidized bed is given by,

@/ðV s  jw Þ V 0p ¼ @/

11

Fig. 2B. Modified flux curve adjusted to include the effect of a fluidization velocity, jw = 0.02 m/s. The internal bed concentration, ub, is the same as in Fig. 2A. The slopes of construction lines, V 0p , Vs–jw, and jw, are shown. The graph is effectively normalized by setting the terminal velocity, Vt = 1 m/s.

underflow removal, and the propagation wave of interest is the upwardly moving bed surface at concentration, /b, which rises at the rate given by the slope of the tangent to the flux curve at /b as shown in Fig. 2B. Extension of the tangent to the vertical axis intersects at the feed flux, jp, necessary to maintain /b. The graphical construction shown in Fig. 2B consists of two parts defined by the tangent. That is,

jp ¼ V 0p /b þ ðV s  jw Þ/b ¼ ðV p þ V s Þ/b

ð3Þ

where V 0p ¼ V p þ jw is the slope at concentration /b. It is evident the right hand-side of Eq. (3), in particular the steady state bed concentration, /b, is independent of the fluidization flux, jw. Combining Eqs. (1)–(3) leads to the result defining the tangent concentration, /b,

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K.P. Galvin et al. / Minerals Engineering 39 (2012) 9–18 

V s ð1  /b Þ : ¼ jp /2b n

ð4Þ

Fig. 3 shows a flux curve with a number of feed operating lines  extending from different feed flux values, jp , to the point of intersection between the modified flux curve and the horizontal axis. The superscript, , is used here for circumstances unconstrained by the flux curve, i.e., for operating lines which are not tangents. The logic is explained by firstly considering the analogous problem of a batch system of particles, fluidized at the velocity, jw. In the first instance we assume there is no feed addition or underflow removal. The point of intersection between the modified flux curve and the horizontal axis governs the volume fraction of the particles, u = 0.73. The concentration, u, has been used here to distinguish the concentration from the tangent value, ub. Note that in this example interpretations should not be placed on the actual numerical values, given the concentration actually exceeds the maximum random packing value for spheres. Rather the focus is on establishing a simple theoretical framework. Here, the key point is that the fluidization velocity, jw, matches the opposing particle velocity, Vs, as obtained by Richardson and Zaki (1954). Thus, using Eq. (1), the volume fraction of the particles in the bed is, 1=n

/ ¼ 1  jw =V t . The gradual volume addition of particles to the batch system at  a feed flux, jp , will simply result in the height of the fluidized sus pension rising with time at the velocity jp =/ . The suspension within the system will remain at the same volume fraction, u = 0.73.  For example, as jp increases to 0.04 m/s or 0.08 m/s, the system concentration remains fixed at u. In order to preserve the suspension height, the suspension must discharge to the underflow according to the operating lines shown in Fig. 3 at the volumetric flux, 0





jL þ jp ¼ jp =/ 0

ð5Þ

attention to the critical underflow rate, jp =/ . The net liquid flux up through the system is, 0



L ¼ jw  jL ¼ jw  jp

  1  / /

ð6Þ

Clearly, the condition (L = 0) defining the onset of an upwards liquid flux is, 

jp / ¼ jw 1  /

ð7Þ

Ultimately, as the volumetric feed flux of particles increases, the system becomes flux curve constrained. This first occurs when the tangent formed from the feed flux, jp, on the vertical axis meets the flux curve as a tangent on the horizontal axis. Here jp = 0.12 m/s and the volume fraction u = ub = 0.73. The point at which the system first becomes flux curve constrained is obtained by substituting the condition jw ¼ V s ¼ V t ð1  /b Þn into Eq. (4). In general, for a flux curve constrained system Eqs. (1) and (4) give (Wallis, 1969),

jp ¼ /2b nV t ð1  /b Þn1

ð8Þ 0

Incorporating Eqs. (1) and (8) into Eq. (3) while noting, jp þ jl ¼ V 0p , gives

jp þ jl ¼ V 0p ¼ V t ð1  /b Þn 0



 n/b  1 þ jw ð1  /b Þ

ð9Þ

Rearranging Eq. (9) gives the net liquid flux passing up through the bed,

L ¼ jw  jl ¼ V t ð1  /b Þn1 ðn/2b  ðn þ 1Þ/b þ 1Þ 0

ð10Þ

When the system is flux curve constrained, the minimum allowed bed concentration is umin = 2/(n + 1) while the maximum al1=n lowed bed concentration is /max ¼ / ¼ 1  jw =V t . Thus,



where jL is the liquid flux and jp the solids flux reporting to under flow. Discharge at a rate greater than jp =/ will result in the loss of the entire bed, and in turn a lower underflow concentration that can be calculated using a simple volume balance. Thus we restrict our

Fig. 3. Operating lines unconstrained by the flux curve, with jp values of 0.04 m/s, 0.08 m/s, up to the maximum at the tangent, 0.12 m/s. The fluidization velocity is jw = 0.02 m/s. The graph is effectively normalized by setting the terminal velocity, Vt = 1 m/s.

2 1=n < /b < 1  jw =V t nþ1

ð11Þ

The minimum concentration, umin, occurs at the flux curve inflection (Wallis, 1969), corresponding to the maximum possible propagation velocity. This conditions defines the maximum possible feed flux beyond which ‘‘flooding’’ occurs. The maximum concentration is governed by the value of the fluidization flux, jw. It can be shown that for n = 3, L is always negative when the system is flux curve constrained.

Fig. 4. Upward flux of liquid passing through the fluidized bed with jw = 0.02 m/s versus the solids feed flux for unconstrained and constrained sedimentation. The graph is effectively normalized by setting the terminal velocity, Vt = 1 m/s.

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Fig. 4 shows the upward liquid flux, L, versus the volumetric flux of particles, jp. Both parameters are effectively normalized by setting the terminal velocity, Vt = 1 m/s. At low feed fluxes the system is unconstrained by the flux curve. The fluidization flux, jw, effectively normalized by setting the terminal velocity, Vt = 1 m/s, is constant at 0.02 m/s. The first part of the curve corresponds to conditions unconstrained by the flux curve, and was produced using Eq. (6). It is evident the liquid flux up through the bed is positive at first and eventually becomes negative. At the higher feed fluxes the system is constrained by the flux curve, given by Eq. 10. Here the liquid flux up through the bed is always negative and thus the actual flow is downwards relative to the vessel. The water split, S, to underflow is given by, 0

0

 ð1/ Þ

jp / jL jL S¼ ¼  jw þ jf þ js j0L þ jO jO

ð12Þ

where jf is the liquid flux in the feed, js is the Split Fluidization flux, and jo is the overflow liquid flux. Fig. 5 shows the corresponding values of S covering the regimes where the system is (i) unconstrained and (ii) constrained by the flux curve. The feed liquid flux, effectively normalized by setting Vt = 1 m/s, was set at a constant level of 0.5 m/s. Clearly, the liquid split to underflow increases monotonically as the feed solids flux increases. Application of a lower fluidization flux, jw, or a higher feed liquid flux, jf, or the introduction of Split Fluidization, js, lowers the liquid split values. This analysis shows that high quality desliming should be possible using a relatively low solids feed flux, ideally unconstrained by the flux curve. Some of the experiments conducted in this study address this hypothesis, with the interstitial liquid moving upwards through the bed. In industry, however, the objective is to operate at the highest possible feed flux in order to reduce the capital investment. This analysis shows that a high feed flux leads to a downwards flow of interstitial liquid through the bed and a high liquid split to underflow. In this study we examine the potential to overcome this problem. Consider a desliming operation at a relatively high feed flux, constrained by the flux curve. Here, the interstitial liquid flux proceeds in the downwards direction. This problem can be solved by introducing additional fluidization, referred to as Split Fluidization, js, at an elevated position well above the base of the vessel as shown in Fig. 1. Then, in the zone between the source of the Split Fluidization and the source of the feed an upwards liquid flux, L0 , can be achieved. The feed particles then settle through this dilute

zone, experiencing a counter current washing of the slimes. In the zone below the source of the Split Fluidization, the sedimentation process resumes, and the system behaves in accordance with the flux curve theory, with a negative liquid flux, L, through this lower zone. Here, the fluidization entering via the distributor, jw, simply serves to establish a state of suspension to help transport the solids to underflow, but little else. 3. Experimental Fig. 1 shows a schematic representation of a laboratory-scale Reflux Classifier used in the experiments. The device consists of a 1 m high fluidized bed with a cross section of 80 mm  100 mm and a 1 m long system of inclined channels, 70° to the horizontal. Eleven thin plates were inserted to form 12 channels each with a perpendicular spacing of 5.7 mm. The plates protruded above the overflow weir, breaking the flow communication (Fan et al., 2010) between the channels, hence insuring even flow up through each channel. Fine iron ore feed, less than 0.50 mm, was prepared in two portions. The first was as a semi-dry solids portion nominally larger than 0.10 mm, and the second was a dilute slimes portion nominally less than 0.10 mm. The semi-dry iron ore solids were well mixed and then sampled into typically 5 kg portions. These solid portions were added manually in a uniform fashion to the feed tube over set periods of time. This approach was used to guarantee a consistent feed composition and solids rate. The typical overall feed size distribution is shown in Table 1. A flow of water at 2 L/ min was added to the feed tube to assist the transport of the feed solids. The slimes were prepared separately using a 300 L mixing tank, with baffles and stirrer. During the experiment, the slimes were pumped up to the feed tube at a pre-determined rate. This approach insured the proportion of feed solids and slimes remained constant throughout the experiments. The fluidization water, jw, was introduced via a distributor located at the base of the Reflux Classifier at 2 L/min in all experiments. Split fluidization water, js, was introduced from the side of the fluidized bed at typically 6 L/min via a peristaltic pump. The underflow was discharged via the base of the fluidized bed using a peristaltic pump. In order to facilitate the smooth transport of the concentrated underflow, a ‘‘T’’ piece arrangement was used. Hence a flow of ‘‘extra water’’ at a rate of 1 L/min was used to assist the flow of the concentrated underflow. Note that underflow pulp density values from the device were re-calculated to remove the effect of this added water. The overflow from the unit, which consisted predominantly of the slimes, reported to the overflow launder, and then into holding tanks. Once steady state was reached, the underflow and the overflow were sampled for specific time periods. The feed sample was obtained by firstly riffling one of the bags of semi-dry feed and com-

Table 1 Typical size distribution of iron ore feed used in the experiments.

Fig. 5. Underflow split of fluid versus the solids feed flux for unconstrained and constrained sedimentation. The fluidization flux jw = 0.02 m/s and feed liquid flux jf = 0.2 m/s. The graph is effectively normalized by setting the terminal velocity, Vt = 1 m/s.

Screen aperture retained (mm)

Geometric mean (mm)

Mass (%)

0.355–0.500 0.250–0.355 0.180–0.250 0.150–0.180 0.125–0.150 0.106–0.125 0.090–0.106 0.075–0.090 0.063–0.075 0.038–0.063 0.038

0.421 0.298 0.212 0.164 0.137 0.115 0.098 0.082 0.069 0.049 0.019

31.66 25.09 16.23 4.35 3.71 2.08 0.95 1.27 1.07 1.37 12.22 100.00

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bining with a portion of the slimes, matching the proportion used in the experiment. All three samples were weighed and then deslimed using laboratory sieves at 0.038 mm. The portion of material above 0.038 mm was dried. The portion below 0.038 mm was allowed to settle prior to decanting and drying. All dry solids were weighed to obtain the pulp density of each flow stream. The dry solids above 0.038 mm were sieved at 0.355, 0.250, 0.180, 0.150, 0.125, 0.106, 0.090, 0.075, 0.063 and 0.038 mm for each stream. Then, the partition curve describing the desliming process was determined. A series of experiments was conducted as shown in Table 2. Note that in order to arrive at a balanced data set, feed data were determined using the sum of the measured output flows. It is also noted that the yield of the solids reporting to the underflow was determined using the underflow and overflow solids rates. Feed samples were nevertheless obtained in order to secure data on the feed size distribution. The partition curves were established using a simple mass balance approach involving the size distributions of the feed, underflow, and overflow, incorporating a function minimization technique to minimize the adjustments in the size distributions (Galvin et al., 1995). This approach led to a second yield value. The difference between these two yield values provided a measure of the quality of the data. On average the magnitude of the yield difference (error) was 1.6%, and when the yield difference exceeded 3%, the data were deemed unacceptable. A consistent yield value of 90% was targeted in each of the experiments however a range of yields was inevitably produced given the difficulty of setting the underflow rate to achieve a specific yield. Experiments with a yield outside the target range of 85–95% were also excluded from the analysis in this paper.

The initial two experiments, Runs A, and B, were conducted at a solids feed flux of about 20 t/m2 h, normally regarded as significant for iron ore particles less than 0.5 mm in diameter. For the purposes of this paper, however, this level will be referred to as being low. The feed pulp density in Run A was significantly higher than in Run B, however additional water was introduced at the Split Fluidization position, located between the feed entry and base of the vessel, in Run B. The net result was that the underflow and overflow solids and liquid fluxes were similar in both runs. Thus these experiments provided a basis to assess the effects of adding water either with the feed or separately from the feed. In the next three experiments, Runs C–E, the solids feed flux was increased significantly to a solids loading of about 40 t/m2 h. This loading is considered very high for a feed less than 0.50 mm in size. In Run E the feed pulp density was notably higher than in Run C. In Run E the Split Fluidization was supplied well below the feed, while in Run C the Split Fluidization was supplied with the feed. In one of the experiments, Run D, there was no Split Fluidization, hence the overall water addition to the device was very low in Run D, resulting in a higher water split to underflow. It should be noted that additional feed water helps to reduce the water split, S, to underflow. However, additional water will cause entrainment of fine iron ore towards the overflow, a major problem in a more conventional upward current classifier. Thus it was important to examine the effects of this additional water. The next experiment, Run F, was conducted at an intermediate solids flux of about 30 t/m2 h in order to further examine the effects of the solids flux on desliming performance. A final experiment, Run G, was again conducted at a very high solids loading of more than 40 t/ m2 h. This time a more extreme level of Split Fluidization was used

Table 2 Summary of desliming runs. Run number

Flow stream

A

Fluidization water Split Fluidization Feed Product Overflow

B

C

D

E

F

G

Fluidization water Split Fluidization Feed Product Overflow Fluidization water Split Fluidization Feed Product Overflow Fluidization water Split Fluidization Feed Product Overflow Fluidization water Split Fluidization Feed Product Overflow Fluidized water Split Fluidization Feed Product Overflow Fluidized water Split Fluidization Feed Product Overflow

Pulp density (%wt)

Solid flux (t/m2 h)

Solid flux (m3/m2 h)

Liquid flux (m3/m2 h)

27.49 63.90 2.30

0.0 0.0 17.8 15.3 2.5

0.0 0.0 4.6 3.8 0.7

15.0 51.0 46.8 8.6 104.2

18.53 66.30 2.40

0.0 0.0 20.9 18.5 2.3

0.0 0.0 5.3 4.6 0.6

15.0 0.0 91.7 9.2 97.4

43.83 67.70 4.60

0.0 0.0 42.1 37.5 4.6

0.0 0.0 10.8 9.4 1.2

15.0 45.0 53.9 18.0 95.9

53.54 71.00 6.90

0.0 0.0 39.8 37.2 2.6

0.0 0.0 10.2 9.3 0.7

15.0 0.0 34.5 15.3 34.2

56.00 70.20 2.10

0.0 0.0 39.9 38.3 1.7

0.0 0.0 10.2 9.6 0.4

15.0 45.0 31.4 16.5 74.9

44.62 72.90 3.51

0.0 0.0 31.1 27.9 3.2

0.0 0.0 8.0 7.0 0.9

15.0 45.0 38.6 10.4 88.3

47.18 69.00 4.20

0.0 0.0 43.5 37.5 6.0

0.0 0.0 11.2 9.4 1.6

15.0 90.0 48.7 16.8 136.9

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K.P. Galvin et al. / Minerals Engineering 39 (2012) 9–18

d75  d25 I¼ 2d50

ð13Þ

where d75 is the particle diameter corresponding to a partition number of 0.75, and d25 the particle diameter corresponding to a partition number of 0.25. The error in the separation was equal to the product, Id50, of the imperfection and the separation diameter. 4. Results and discussion

Partition Number (%)

100

80

60

40

20 Run C 0 0.0

0.1

0.2

Run E 0.3

Run D 0.4

0.5

Particle Size (mm) Fig. 7. Partition curves produced for runs conducted at very high solids throughputs of order 40 t/m2 h. It is evident that at this very high solids throughput separation performance is poor. However, it is also evident Split Fluidization in Runs C and E led to improved performance over Run D which involved no Split Fluidization.

100

Partition Number (%)

in order to further reduce the water split to underflow and in turn improve desliming. It is useful to compare the settling characteristics for these experiments with the solids flux curves shown earlier in this paper. Clearly the feed used has a broad particle size range while the flux curve theory is based on particles of one size. The geometric mean for a particle diameter range of 0.1–0.5 mm is 0.224 mm. Assuming the Intermediate regime applies (Rep > 0.1), a particle of this diameter and density of 4000 kg/m3 has a terminal velocity, Vt, in water of 0.043 m/s. A solids flux of 40 t/m2 h corresponds to a volumetric flux of 10 m3/m2 h (or 0.0028 m/s). If we scale the solids feed flux to a level consistent with the scaling of the terminal velocity from 0.043 m/s to 1 m/s (by a factor of 23), then jp = 0.065 m/s. Similarly the fluidization flux of 15 m3/m2 h (0.0042 m/s) would be scaled to jw = 0.1 m/s. These values, jp = 0.065 m/s and jw = 0.1 m/s, help to place the work conducted in this study into an appropriate context with reference to the earlier solids flux curve analysis. Separation performance was established by determining the partition curve as a function of the particle size. The partition number is the probability that a particle of diameter d reports to the underflow. The d50 is the separation diameter with equal probability of exiting via the underflow or overflow. The sharpness of the separation can be represented by the imperfection, I, defined here by,

Run F Medium Throughput Run E High Throughput Run A Low Throughput

80

60

40

20

The performance of the desliming was quantified using partition curves. The curves are shown in Figs. 6–8 and the characteristic parameters and Imperfections given in Table 3. The objective was to achieve a sharp particle size separation at about 0.080 mm. The feed particles used had already undergone beneficiation hence there was relatively little low density particles in the feed. It is noted that as a concentrate, the density of the +90 micron feed was 3640 kg/m3 while the density of the 90 lm feed was 3936 kg/m3, hence the challenge of desliming here was more difficult than would be the case if the slimes were of a more normal density of 2650 kg/m3. Moreover, the density of the 90 lm

0.10

0.20

0.30

0.40

0.50

Particle Size (mm) Fig. 8. The effect of the solids throughput on desliming performance. The partition curve for Run F, produced at an intermediate solids throughput of order 30 t/m2 h is shown. The results from Run A at a lower solids throughput of order 20 t/m2 h and Run E at the very high solids throughput of order 40 t/m2 h are also shown. In each case significant Split Fluidization was provided below the feed entry in order to maximize performance. Clearly, the separation performance decreases as the solids throughput increases, with the partition number at 0.050 mm increasing.

Table 3 Analysis of separations and calculation of Imperfection.

100

Partition Number (%)

0 0.00

80

60

40

Run

d25

d50

d75

Imperfection I

A B C D E F G

87 63 53 – 48 65 90

93 73 65 – 57 75 106

100 79 78 – 63 83 122

0.070 0.11 0.19 – 0.13 0.12 0.15

20 Run B 0 0.00

0.10

0.20

0.30

0.40

Run A 0.50

Particle Size (mm) Fig. 6. Partition curves produced for desliming runs conducted at relatively low solids throughputs of order 20 t/m2 h. The desliming was clearly very efficient in both cases with the partition number less than 5% at 0.050 mm.

particles reporting to the underflow product was typically 4300 kg/m3, while the density of the 90 lm particles reporting to the overflow reject was typically 3800 kg/m3. So, invariably misplaced fines reporting to the underflow product were in fact high grade particles of iron ore. Given the broad range of experimental conditions used in this study it was difficult to judge the required underflow rate needed

16

K.P. Galvin et al. / Minerals Engineering 39 (2012) 9–18

to target a separation at 0.080 mm. The underflow rate was set at a level needed to withdraw a fixed fraction of the feed solids. Some judgement was necessary here during an experiment and hence it was inevitable that different underflow solids splits would be produced. In order to insure a consistent basis for investigating the desliming, the analysis in this paper was limited to a solids split (yield) to underflow in the range 0.85–0.95. Fig. 6 shows partition curves obtained at a significant solids feed flux of about 20 t/m2 h. This level is described in the paper as low, relative to the other solids feed fluxes. The separations are very sharp with Imperfections of 0.07–0.11 and exhibit a negligible slimes entrainment arising from the water split to underflow. Significant Split Fluidization at a flux of 51 m3/m2 h was used in Run A, opposing the downwards transport of the slimes towards the underflow. In Run B, the extra water was added as part of the feed. It is evident Run A produced the better separation, with a lower Imperfection and smaller partition number at 0.050 mm. Given the very sharp separations achieved at this solids feed flux, it was necessary to examine separations at very much higher solids feed fluxes in order to assess the merit of Split Fluidization. Fig. 7 shows the partition curves produced at a very high solids feed flux of about 40 t/m2 h. The partition curves are clearly relatively poor, with a ‘‘fish-hook’’ shape, and a significant partition number at about 0.050 mm. Clearly, this solids throughput is too high. Run D, conducted with no Split Fluidization, produced by far the worst separation, with the water split to the underflow significant at 0.3. The other two experiments were conducted using Split Fluidization at a flux of 45 m3/m2 h. In Run C the Split Fluidization was introduced at a level below the feed, causing a fish-hook effect, while in Run E the extra water was introduced directly with the

Partition Number (%)

100

80

60

40

20 Run G High Throughput and High Split Fluidization

0 0.0

0.1

0.2

0.3

0.4

0.5

Particle Size (mm) Fig. 9. Partition curve obtained in Run G conducted at a very high solids feed flux in excess of 40 t/m2 h. Here a more extreme level of Split Fluidization of 90 m3/m2 h was used to reduce the water split to underflow. The effectiveness of the Split Fluidization is evident given the relatively low partition number at 0.050 mm.

feed. It is evident there was better performance in Run C, with the partition number lower at 0.050 mm, though the ‘‘fish-hook’’ effect was more pronounced, perhaps due to less effective washing of the feed solids. The Imperfections were reasonable despite the high solids throughput with I = 0.19 for Run C and I = 0.13 for Run E. In order to establish the solids throughput limit, Run F was conducted using an intermediate solids flux of about 30 t/m2 h. The partition curve is shown in Fig. 8 together with partition curves produced using the lower and higher solids flux levels. The Imperfection is low at 0.12. It is clear the partition curves are all relatively sharp. However the curves reach a minimum partition number at a particle diameter of about 0.050 mm, with the lower limit increasing as the solids throughput increases. This shift in the lower limit is common in a range of devices including hydrocyclones, and is attributed to the ultrafine particles partitioning with the water split to the underflow. Some effort has been given to introducing water at a location near the cyclone underflow to reduce this effect (Honaker et al., 2001; Dueck et al., 2010; Farghaly et al., 2010). Fluidized bed devices have a clear advantage due to the potential to introduce both fluidization and Split Fluidization water. Run G was conducted using the high solids feed flux level of 40 t/m2 h, this time using a much more extreme level of Split Fluidisation delivered at 90 m3/m2 h. The result, shown in Fig. 9, achieves a satisfactory Imperfection of 0.15, but importantly the partition number at 0.050 is relatively low at 0.07. This is a remarkable result given the previous difficulty at the high solids feed flux. It is clear that the Split Fluidization is effective and could readily be extended to improve the results obtained at the lower solids feed fluxes. This effectiveness, however, was only possible due to the strong segregation produced by the system of inclined channels. Table 4 provides a summary of the volumetric fluxes of the solid and liquid components used in the series of experiments. The upward liquid flux, L, between the base of the vessel and the Split Fluidization position is shown, as well as the upward liquid flux, L0 , between the Split Fluidization and feed positions. It is evident the upward liquid flux, L, was only positive in the experiments conducted at 20 t/m2 h and 30 t/m2 h, however, it was possible to achieve an upward liquid flux, L0 , at the very high solids feed fluxes by introducing the Split Fluidization. Fig. 10 shows the upwards liquid flux, L, versus the volumetric solids feed flux. The results are similar to those shown earlier in Fig. 4 based on the Drift Flux analysis. At a zero solids feed flux steady state demands no underflow removal and hence the liquid flux is equal to the superficial fluidization velocity, 15 m3/m2 h. As the solids feed flux increases, the higher underflow rate required to achieve steady state causes a reduction in the net upwards liquid flux through the bed, leading to negative values. Under these conditions Split Fluidization becomes essential in order to reduce the water split to underflow and carry slimes to the overflow. Fig. 11 shows the partition number (%) obtained for a particle size of 0.050 mm versus the liquid split, S, reporting to the underflow. The partition curves in Figs. 6–9 exhibit a minimum at a particle size of 0.050 mm, and thus provide a useful measure

Table 4 Summary of desliming fluxes. Run

A B C D E F G

Feed Solids

Under-flow Solids

Net Liquid Flux L

Net Liquid Flux L0

Over-flow Liquid jo

(m3/m2 h)

Under-flow Liquid j0L (m3/m2 h)

(t/m2 h)

(m3/m2 h)

(m3/m2 h)

(m3/m2 h)

17.8 20.9 42.1 39.8 39.9 31.1 43.5

3.8 4.6 10.8 9.3 9.6 7.0 9.4

8.6 9.2 18.0 15.3 16.5 10.4 16.8

6.4 5.8 3.0 0.3 1.5 4.7 1.8

57.4 5.8 42.0 0.3 43.5 49.7 88.2

104.2 97.4 95.9 34.2 74.9 88.3 136.9

Partition Number at 0.05 mm (%)

2.6 5.0 22.1 72.3 29.8 13.5 7.0

Liquid Split

Ratio

Under-flow

jp/jo

0.077 0.086 0.158 0.309 0.181 0.105 0.110

0.044 0.055 0.112 0.298 0.137 0.090 0.081

K.P. Galvin et al. / Minerals Engineering 39 (2012) 9–18

Upward Liquid Flux L (m/h)

15

10

5

0 0

5

10

15

-5

Feed Flux, j p, m3/m2.h Fig. 10. The liquid flux passing up through the fluidized bed versus the solids feed flux. In the absence of continuous feed, steady state demands no underflow removal, and hence the upwards liquid flux matches the superficial fluidization velocity of 15 m3/m2 h. At high solids feed flux values, the high underflow rates produces a net downwards flux of liquid.

Partition Number (%) at 0.050 mm

100

80

60

40

20

0 0.00

0.05

0.10

0.15

0.20

0.25

0.30

0.35

Water Split to Underflow Fig. 11. The partition number at 0.050 mm was obtained for each run and used as a measure of separation performance. It is evident there is clear correlation between this partition number and the water split, S, reporting to the underflow. The water split to underflow tends to increase with the solids throughput, and is also influenced by the addition of the Split Fluidization below the feed.

of the desliming performance of the process. It is evident that the partition number increases linearly with the fraction, S, of the water reporting to the underflow. Based on Eqs. (6) and (12), this water split is approximated by, 

S

jP



1/ /

jo

 ð14Þ

The numerator, which is based on the solids flux and solids volume fraction, provides a measure of the liquid flux reporting to the underflow, while the denominator is based on the overflow flux of liquid. Reasonable agreement with the actual water split values in Table 4 is obtained using a solids volume fraction of 0.4. Eq. (14) is

17

significant as it shows the water split to the underflow is strongly  governed by the solids flux, jp , reporting to the underflow, relative to the liquid flux reporting to the overflow. The coarser particles in the solids flux that report to the underflow will entrain other finer particles including slimes. These finer particles will achieve a degree of association with the coarser particles by locating behind the falling particles and via lubrication forces and adhesion. Particles 0.050 mm in diameter are large enough to dissociate from the coarser particles, especially when subjected to the Split Fluidization. The ultrafine particles less than 0.038 mm retain a stronger level of association hence we observe the classic ‘‘fish hook’’ shape in the partition curve. It is noted however that the significant absence of particles from the underflow, in the size range 0.038–0.085 mm, creates a very permeable underflow, improving the dewatering characteristics. This means the ultrafine particles less than 0.038 mm will tend to readily drain from the underflow product and not necessarily pose a concern. While much of the focus of this study was on the lower zone of the Reflux Classifier, the system of parallel inclined channels played a very significant role, insuring the retention of the fine high density particles, especially those near the separation size of 0.10 mm. We have previously investigated the retention of these particles, developing a rigorous empirical relationship using Dimensional Analysis (Laskovski et al., 2006). It is evident in Table 2 that the Split Fluidization leads to a high overflow flux of liquid (75–137 m3/m2 h). A value of 95 m3/m2 h or 0.027 m/s is used in the following analysis. This velocity is equivalent to the terminal velocity of a 0.16 mm particle of density 4000 kg/m3. Thus in a conventional upward current classifier this level of Split Fluidization would cause a significant loss of particles up to 0.160 mm in diameter, much larger than the desliming size of about 0.085 mm. The Split Fluidization also leads to a net upwards liquid flux, L0 , in the zone just below the feed entry of about 45 m3/ m2 h or 0.013 m/s. This flux counters the strong downwards solids flux. This liquid superficial velocity of 0.013 m/s is equal to the terminal velocity of a 0.1 mm particle of density 4000 kg/m3. Thus the superficial velocity impinging on the falling feed particles is sufficient to oppose the downward movement of particles less than 0.1 mm, providing an ideal barrier to the movement of slimes. The inclined channels of the Reflux Classifier provide substantial segregation area and hence a mechanism for the retention of relatively fine particles. The asymptotic throughput advantage, applicable to inclined channels of infinite aspect ratio (Laskovski et al., 2006), is

U 0 =v t ¼ 7:5Re0:33 p

ð15Þ

where U0 is the superficial velocity of the liquid through an inclined channel, Vt the terminal velocity of the particle, and Rep the particle Reynolds number. Consider a particle 0.040 mm in diameter and 4000 kg/m3 in density. The particle Reynolds number, Rep, is 0.1 and hence the asymptotic throughput advantage, U 0 =V t , is 16.0 within a single inclined channel. Based on the superficial velocity through the vertical section, the throughput advantage reduces to 15.1, and then reduces to 11.0 when the plate thickness and finite plate length are considered (Laskovski et al., 2006). The particle terminal velocity, Vt, is 0.0025 m/s and hence the corresponding superficial velocity is U = 11  0.0025 = 0.027 m/s, equivalent to the liquid flux reporting to the overflow. Thus particles smaller than 0.100 mm in diameter were swept upwards from the zone below the feed entry while particles smaller than 0.160 mm in diameter were swept upwards from the zone above the feed entry. However, particles larger than 0.040 mm were prevented from elutriating and leaving the system. These conflicting conditions lead to an accumulation of particles in the

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K.P. Galvin et al. / Minerals Engineering 39 (2012) 9–18

size range 0.040–0.160 mm in diameter in the zone above the feed entrance. The internal build up in the concentration of fine particles within this size range provides the foundation for a sharp separation. Clearly, the particles in the size range 0.100–0.160 mm will ‘‘leak’’ through the feed entry zone, via dispersion, and readily segregate towards the underflow. The particles in the size range 0.040–0.100 mm however, need to accumulate to a far greater extent before ‘‘leakage’’ to the underflow becomes possible. With this accumulation there is an increased tendency for the finer particles of about 0.040 mm to elutriate via the overflow due to increasing level of hindered settling. The overall separation produced is governed, in the end, via the setting of the underflow pump which controls the precise fraction of the feed solids that are permitted to report to the underflow. Thus the internal processes within the Reflux Classifier simply provide a sorting mechanism, delivering the fastest settling particle to the underflow ahead of progressively slower settling particles. Once the required solids-split to the underflow has been reached, the balance of the feed must ultimately pass to the overflow. Finally, it is noted that in the alternative application of gravity separation, where the objective is to separate particles primarily on the basis of density, Split Fluidization should not be applied, given the water injection will tend to create a zone of lower bed density, permitting the larger, lower density, particles to settle downwards. 5. Conclusions A series of experiments was conducted on the desliming of a feed of fine iron ore, nominally less than 0.50 mm in diameter, using the Reflux Classifier. Very sharp separations were produced at a solids throughput of about 20 t/m2 h, as defined by the partition curves. Split Fluidization was found to enhance the quality of the separation in a number of ways. Firstly, the extra liquid flux opposed the downwards transport of the finer particles. Secondly, the extra liquid reduced the fraction of the water, S, reporting to the underflow and hence the probability of ultrafine entrainment towards the underflow. These benefits were especially evident in the experiments conducted at a very high solids throughput of about 40 t/m2 h. Here the solids flux reporting to the underflow produced negative liquid fluxes in the lower section. The Split Fluidization, however, led to a significant upwards liquid flux in the zone above the level of water addition, which in turn produced improved separation. When there was no Split Fluidization, and a high solids flux was used, very poor separation was produced. Satisfactory performance at the intermediate solids flux of 30 t/m2 h provided a measure of the throughput limit for this feed. Disclosure statement The University of Newcastle holds international patents on the Reflux Classifier and has a Research and Development Agreement with the company Ludowici Australia.

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